Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019): Volume II [1st ed. 2020] 978-3-030-22062-4, 978-3-030-22063-1

Table of contents :
Front Matter ....Pages i-xix
Modeling of Roll Roughness Transfer Process to Strip During Skin-Pass Rolling (N. N. Ogarkov, S. I. Platov, E. U. Zvyagina)....Pages 1-7
Working Surface Calculation of Teeth Bevel Gear Helical-Bevel Gearing at Milling with Hob (E. A. Poluektov, B. A. Lopatin, S. V. Plotnikova)....Pages 9-16
Effectiveness of Application of Additional Strengthening Processing of Surface Plastic Deformation on Increase in Fatigue Life of Parts (V. A. Lebedev, A. V. Kirichek, L. V. Chunakhova)....Pages 17-25
Determination of Rational Technological Parameters of Cold Bending with Rolling of Pipes from Corrosion-Resistant Steels (E. V. Khaliulin, A. V. Bobylev, A. V. Kozlov)....Pages 27-39
Recomposition Procedure of Automatic Replacement Laser Modules for CNC Machines (P. A. Ogin, D. G. Levashkin)....Pages 41-49
Tooling Backup of Cutting and Deforming Processing of Non-rigid Shafts (E. I. Yatsun, N. P. Anikeyeva, I. S. Karnaukhov)....Pages 51-58
Determination of Flanging Parameters and Length of Screwing in Producing Holes by the Method of Thermal Drilling in Thin-Sheet Metal (P. V. Shalamov, I. A. Kulygina, A. N. Yasnitsky)....Pages 59-68
Practice of Dimensional Modeling in the Implementation of the Methodology of Group Interchangeability (M. G. Galkin, A. S. Smagin)....Pages 69-77
Determination of Settling Efficiency of Solid Finely Dispersed Particles Within Devices with Rectangular Separators (I. N. Madyshev, O. S. Dmitrieva, A. V. Dmitriev)....Pages 79-84
Method for Predicting Thermal Characteristics of Machine Tools Based on Experimental Modal Analysis (A. N. Polyakov, A. N. Goncharov, I. V. Parfenov)....Pages 85-93
Development of New Types of Contact Devices for Heat-Mass Transfer Apparatuses, Used at Petrochemical Enterprises (I. N. Madyshev, O. S. Dmitrieva, A. V. Dmitriev)....Pages 95-101
Stiffness Maximization on the Basis of Layout Characteristics of the Elastic Machine System and Milling Process (R. M. Khusainov, A. R. Sabirov)....Pages 103-109
Formation Automation of Geometric Configuration of Real Machine Parts (V. E. Lelyukhin, O. V. Kolesnikova, O. M. Ponkratova)....Pages 111-119
Information Model for Machine’s Electronic Structure Storage (A. Loginov)....Pages 121-129
Software Spindle Speed Variation as Method for Chatter Suppression in Drilling (V. M. Svinin, A. V. Savilov, A. V. Shutenkov)....Pages 131-139
Experimental Studies of Steady-State Sources of Vibrations of Machinery Production Process Equipment to Substantiate Choice of Vibration Protection Methods (S. I. Gvozdkova, L. E. Shvartsburg)....Pages 141-149
On Issue of Verifying New Method for Studying Dynamics of Deep Hole Machining (L. Mironova, L. Kondratenko, V. Terekhov)....Pages 151-162
Rational Provision of Robustness Properties of Bolted Joints of Assembly with Implementation of Anaerobic Materials (I. I. Voyachek, D. V. Kochetkov, S. G. Mityasov)....Pages 163-170
Machining Accuracy Improving with the Use of Mobile Mechatronic Systems as Industrial Robot End Effectors (E. I. Shchurova, P. G. Mazein)....Pages 171-180
Voxel and Finite Element Modeling of Twist Drill (E. I. Shchurova)....Pages 181-190
3D Modeling of Turbine Rotor Journal Machining with Location on a Bearing Bottom Half (A. V. Shchurova)....Pages 191-198
Geometric Modeling of Macro-defects of Parts Surfaces Based on Discrete Solid-State Modeling (I. A. Shchurov)....Pages 199-207
Selecting Criterion of Long-Term Strength in Assessing Durability of Constructions Operating Under Non-isothermal Loading Processes (A. V. Belov, A. A. Polivanov, N. G. Neumoina)....Pages 209-216
Virtual Prediction of Accuracy of Processing on Example of External Circular Grinding (P. P. Pereverzev, A. V. Akintseva, M. K. Alsigar)....Pages 217-224
Designing of Optimal Grinding Cycles, Sustainable to Unstable Mechanical Processing on the Basis of Synthesis of Digital Double Technology, and Dynamic Programming Method (P. P. Pereverzev, A. V. Akintseva, M. K. Alsigar)....Pages 225-232
Controlled Assembly of Rotors (S. M. Beloborodov, V. F. Makarov, M. L. Tselmer)....Pages 233-240
Study of Energy Intensity of Shaped Holes Broaching Process (V. V. Kuts, M. S. Razumov, V. S. Kochergin)....Pages 241-248
Methods of Quality Control Manufacturing Gears of the Differential Satellites (D. T. Safarov, A. G. Kondrashov)....Pages 249-259
Computational and Experimental Studies of Deformations of Air-Cooled Diesel Cylinders at Its Assembling (I. E. Agureev, K. Yu. Platonov, R. N. Khmelev)....Pages 261-270
Feasibility Study of Applying Group Interchangeability Method for Assembly Components of Machines (I. I. Voyachek, E. S. Slashchev, D. M. Malikova)....Pages 271-280
Development of Integrated Criterion to Select Environmentally Sound Cutting Fluids and Relevant Application Systems in Shape-Forming Processes (L. E. Shvartsburg, O. V. Yagolnitser, E. V. Butrimova)....Pages 281-288
Gear-Grinding Temperature Modeling and Simulation (N. V. Lishchenko, V. P. Larshin)....Pages 289-297
How Parameters of Agricultural Machine and Tractor Unit Affects Effectively Used Mean Indicated Power (M. Ya. Durmanov, B. G. Martynov, S. V. Spiridonov)....Pages 299-312
Method to Reduce Oil Burning in Diesel Engine of Agricultural Machine and Tractor Unit (V. P. Antipin, M. Ya. Durmanov, O. A. Mikhailov)....Pages 313-323
Mill Conditions Effect on Roughness of Injection Molds’ Forming Surfaces (S. I. Kozhevnikov, V. F. Makarov)....Pages 325-335
Effect of Magnetic Processing on Mass Transfer in a Frictional Pair “Alloyed Steel–Carbon Steel” (A. N. Gots, V. V. Zelinskiy, E. A. Borisova)....Pages 337-345
Strength Parameters of Hardening Cylindrical Workpieces by Tapered Roller (A. A. Udalov, A. V. Udalov, S. L. Vasilevykh)....Pages 347-354
Formation of Schemes Generating Geometric Structure of Machine Parts (O. V. Kolesnikova, V. E. Lelyukhin, F. Yu. Ignatev)....Pages 355-363
Optimization of Cutting Parameters in Milling by Means of System Nyquist Plot (R. M. Khusainov, P. N. Krestyaninov, D. D. Safin)....Pages 365-371
Comprehensive Evaluation of Shaft Manufacturability: Mathematical and Information Models (A. Sychugov, Yu. Frantsuzova, V. Salnikov)....Pages 373-379
Simulation of Electrical Discharge Machining of Micro-holes (T. D. Nguyen, V. M. Volgin, V. V. Lyubimov)....Pages 381-390
Investigation of Kinematic–Geometric Characteristics of Electrochemical Machining (V. V. Lyubimov, V. M. Volgin, V. P. Krasilnikov)....Pages 391-398
Model for Numerical Simulation of Temperature Field and Bead Profile in Hybrid Laser-Arc Welding of T-Joint (S. Ivanov, E. Valdaytseva, I. Udin)....Pages 399-409
QMS as Tool for Improving Maintenance and Repair Processes of Traction Rolling Stock (A. V. Muratov, V. V. Lyashenko, S. A. Petukhov)....Pages 411-418
Heat Treatment Effect on the Structural and Elastic Characteristics of a Single-Component Abrasive Tool (M. Yu. Polyanchikova)....Pages 419-426
Approach for Modeling and Situational Management of Industrial Product Efficiency (E. V. Orlova)....Pages 427-437
Modeling of Try-Out and Reliability Estimation of Technological Processes of Machine-Building Products (V. M. Trukhanov, M. P. Kukhtik, A. M. Makarov)....Pages 439-446
Methods of Cutting Cost Minimizing in Problem of Tool Route Optimization for CNC Laser Machines (A. F. Tavaeva, A. A. Petunin, E. G. Polishchuk)....Pages 447-455
Development of Hardware-Algorithmic System for ICE Diagnostics (L. A. Galiullin, R. A. Valiev)....Pages 457-467
Information Support of Gas-Turbine Engine Life Cycle Based on Agent-Oriented Technology (A. Zagitova, N. Kondratyeva, S. Valeev)....Pages 469-476
Automatic Print Job Scheduling and Management Over Multiple 3D Printers (I. A. Gushchin, I. V. Martynovich, I. S. Torubarov)....Pages 477-487
CNC Processing Equipment’s Technical Operability Evaluation by Developing Mathematical Models Based on Continuous Logic of Antonyms (E. Krylov, N. Kozlovtseva, A. Kapitanov)....Pages 489-499
Digital In-Line Moisture Meter of Thin Sheet Materials (S. V. Makartichyan, L. V. Khoperskova, V. E. Avvakumov)....Pages 501-509
Modeling Physical Operating Principles During Search Design of Cooling and Refrigerating Systems (A. A. Yakovlev, V. S. Sorokin, S. G. Postupaeva)....Pages 511-520
Analysis of Measurement System Accuracy Based on 2D Laser Triangulation Scanner When Measuring Soiled Pipe Thread Pitch and Height (D. S. Lavrinov)....Pages 521-529
Mechanics of Elastic Wheel Rolling on Rigid Drum (M. Yu. Karelina, T. A. Balabina, A. N. Mamaev)....Pages 531-540
Technical Audit of Rotary Aggregates (S. G. Mogilny, A. A. Sholomitskii, A. L. Sotnikov)....Pages 541-549
Hydroblow as Mechanism of Additional Intensification of Liquid Forages Preparation in Centrifugal-Rotor Dismembrators (Yu. N. Kamishov, N. A. Makarova, A. A. Sitnikov)....Pages 551-558
Change of Structure of Dispersion Material Under Dynamic Loads (E. K. Chabutkin)....Pages 559-565
Increasing Efficiency of Vibratory Rollers Through Adjusting Magnitude of Disturbing Force (Y. G. Popov, E. K. Chabutkin)....Pages 567-576
Energy Component of Properties of Material Crushability Layer (Yu. A. Lagunova, V. S. Bochkov)....Pages 577-584
New Criterion for Continuous Compaction Control Systems by Soil Vibratory Rollers (I. S. Tyuremnov, A. S. Morev)....Pages 585-593
Reducing Amplitude of Load Swinging During Operation of Hydraulic Manipulators of Forest Transport Machines (P. Popikov, M. Drapalyuk, D. Druchinin)....Pages 595-608
Increase in Operational Reliability of Locking Device of Hydraulic Actuator of Vehicle (N. A. Fomenko, S. V. Aleksikov, S. G. Artemova)....Pages 609-616
Influence of Hydrogen Additives on Cycle-to-Cycle Variability of Working Process of Rotary Engine (Y. V. Levin, K. V. Prikhodkov, E. A. Fedyanov)....Pages 617-624
Efficiency of Usage of Transport and Technological Machines (G. V. Redreev, G. A. Okunev, S. A. Voinash)....Pages 625-631
Differential System of Crane Braking (S. V. Streltsov, V. A. Ryzhikov)....Pages 633-642
Thermal Loading Estimation of the Friction Pairs of a Vehicle Automated Brake System (V. Dygalo, I. Zhukov)....Pages 643-650
Problem of Increasing Tractive Effort of Railway Locomotives in Conditions of Arctic and Continental Shelf Regions (A. Keropyan, S. Albul, A. Zarapin)....Pages 651-658
New Approach for Experimental Identification of Internal Combustion Engine Power and Performance Characteristics (A. Egorov, N. Syutov, V. Belogusev)....Pages 659-666
Development of Hardware and Software Complex for Increase of Technical Readiness Transport-Technological Machines in Forestry (S. V. Lyakhov, S. V. Budalin)....Pages 667-678
Study of Characteristics of Engine Operation in Stress-Testing Mode of Electric Gasoline Pump (A. Vozmilov, D. Vlasov, K. Glemba)....Pages 679-686
Technical Level Analysis of Structures of Quarried Excavators (H. N. Sultonov, K. Z. Tilloev)....Pages 687-694
Reliability Analysis of Bus Steering System (A. N. Savkin, K. A. Badikov, A. A. Sedov)....Pages 695-701
Process Research of Wheel-Rail Mining Machines Traction (A. M. Keropyan, D. A. Kuziev, A. E. Krivenko)....Pages 703-709
Development of Double-Sided Summer–Winter Pneumatic Tires (I. Voiku, I. Komissarov)....Pages 711-718
Asphalt Concrete as Object of Destruction by Operating Units of Milling Machines (D. V. Furmanov, V. A. Nikolayev, N. N. Klochko)....Pages 719-726
Improving Automotive Torque Converter Quality (N. N. Trushin, V. Y. Antsev, A. A. Obozov)....Pages 727-735
Stabilization of Biaxial Trailer Motion (Yu. Stroganov, A. Popova, D. Zhelev)....Pages 737-745
Development of Optimization Algorithm to Control Open-Pit Excavator Operation (O. Lukashuk, A. Komissarov, K. Letnev)....Pages 747-753
Complex for Inspection of Crane Rails Design (V. Yu. Antsev, P. V. Vitchuk, K. Yu. Krylov)....Pages 755-763
Study of Impact of Amount of Shock Absorbers on Parameters of Vibrations of Drum and Frame of Vibrating Roller (I. S. Tyuremnov, D. V. Fyodorova, A. S. Morev)....Pages 765-773
Optimization of High-Power Belt Conveyor Parameters (G. G. Kozhushko, O. A. Lukashuk, T. A. Roscheva)....Pages 775-782
Intelligent Control, Correction, and Adaptability of Output Parameters of Vehicles’ Intake System (A. V. Gritsenko, I. V. Makarova, G. N. Salimonenko)....Pages 783-793
Duration of Ignition Delay of Fuel–Air Mixture in Diesel Engines (V. F. Guskov, A. N. Gots)....Pages 795-803
Modeling of Maximum Cycle Pressure Based on Engine External Speed Performance (A. N. Gots, V. S. Klevtsov)....Pages 805-812
Effectiveness of Road Transport Technology in Modern Housing Systems (A. V. Kulikov, S. Y. Firsova)....Pages 813-821
Robot Manipulator Control with Efforts Stabilization in Capture of Object with Fuzzy Geometrical Characteristic (V. I. Chizhikov, E. V. Kurnasov, A. B. Petrov)....Pages 823-837
Minimization of Energy Costs for Movement Resistance of Ground for Walking Device by the Control of Support Points Motion (V. V. Chernyshev, V. V. Arykantsev, I. P. Vershinina)....Pages 839-848
How Different Autonomous Cutter Cooling Methods Affect Machining Performance (D. Yu. Dubrov)....Pages 849-855
Calculation of Thermodynamic Parameters of Geometrically Complex Parts at Abrasive Globoid Gear Machining (V. A. Spirin, V. F. Makarov, O. A. Khalturin)....Pages 857-864
Providing of Surfaces’ Geometry at the Design Stage of Profile Milling Operation of Off-Grade Workpiece (A. A. Fomin, V. G. Gusev, N. F. Timerbaev)....Pages 865-873
The Method of the Combined Flat Peripheral Grinding (A. V. Morozov, V. G. Gusev)....Pages 875-883
Defect Analysis of Operating Hydro-Gasified Piping System (Min Ko Hlaing, Phone Htet Kyaw, B. N. Maryn)....Pages 885-894
Contact Zone Effect Analysis onto the Deforming Rollers Geometrics During the Surface Plastic Deformation Treatment (Y. N. Oteniy, O. V. Martynenko, N. I. Nikiforov)....Pages 895-900
Increasing Resistance of Cutting Tool with Diamond Burnishing (N. Papsheva, O. Akushskaya)....Pages 901-909
Determining Coordinates of Cutting Force Application Point in Grinding Zone (V. A. Nosenko, M. V. Danilenko)....Pages 911-919
Improving Efficiency of Machining of Grooves on Shafts of Increased Hardness Structural Steel (S. V. Grubyi, P. A. Chaevskiy)....Pages 921-929
General Patterns in Formation of Surface Layer of Machine Parts Treated by Combined Electro-technical Methods (S. V. Usov, P. A. Davydenko, D. S. Sviridenko)....Pages 931-941
Evaluation of Tool Life Equation of Single-Point Cutting Tool by Accumulation Model (A. V. Antsev, N. I. Pasko, A. V. Khandozhko)....Pages 943-951
Ultra-precision Machining of Surfaces of Elements of Devices from Optical Materials (S. V. Grubiy, M. A. Shavva, V. V. Lapshin)....Pages 953-961
Identification of Deformations and Errors in Flat Thin Workparts During Grinding (T. N. Ivanova)....Pages 963-972
Novel Method of Single-Pass Threading by Cutter (N. N. Zubkov)....Pages 973-982
Chip-Forming Processes at High-Speed Grinding (V. M. Shumyacher, O. G. Kulik, S. A. Kryukov)....Pages 983-994
Express Control of Abrasive Tool Operational Characteristics (V. M. Shumyacher, S. A. Kryukov, O. G. Kulik)....Pages 995-1001
Study of Physical and Chemical Processes Occurring During Polycondensation of Bakelite Binder in Order to Adjust Technological Process of Abrasive Tool Production (I. Yu. Orlov, T. N. Orlova, I. V. Bashkirtseva)....Pages 1003-1010
Optimal Sequencing at Selection of Abrasive Tools’ Characteristics for Their Improvement (I. Yu. Orlov, N. V. Baidakova, P. Yu. Bochkarev)....Pages 1011-1017
Change in Microhardness of Metal Depending on Wetting Ability of Lubricating Coolant (I. V. Bashkirtseva, T. N. Orlova)....Pages 1019-1026
Studies on Titanium Alloy Turning Rate Improvement (A. V. Savilov, V. M. Svinin, S. A. Timofeev)....Pages 1027-1033
Ultrasonic Impact Study on Strain Hardening of Thread Profile Surface Layer (V. Golovkin, O. Batishcheva, V. Papshev)....Pages 1035-1041
Dissipative Structure of Contact Interaction When Cutting Metals (V. A. Kim, B. Ya. Mokritsky, A. V. Morozova)....Pages 1043-1051
Cutting Temperature by Polymer-Abrasive End Brushes for Machining Planes (D. B. Podashev, Yu. V. Dimov)....Pages 1053-1060
Researching the Influence of the Location Tool on the Treatment of Large Shafts Requiring Surface Shaping (Y. A. Bondarenko, N. A. Maslennikov, A. A. Mamchenkova)....Pages 1061-1069
Numerical Modeling of the Material Layer Upset Forging with Extrusion Under the Stiffening Rib into the Forging Cavity (O. A. Nikitina, T. M. Slobodyanik)....Pages 1071-1077
Centrifugal Rolling of Flexible Shafts for Achieving Best Possible Roughness of the Surface (N. V. Vulykh)....Pages 1079-1088
Thread Milling Cutter Flute Production Possibility Research by Using Typical Profiles Grinding Wheels (O. V. Malkov, I. A. Pavlyuchenkov)....Pages 1089-1096
Thermo-Emf as Method for Testing Properties of Replaceable Contact Pairs (Z. Tikhonova, D. Kraynev, E. Frolov)....Pages 1097-1105
Mathematical Apparatus for Predicting Cutting Tool Life in Turning Process After Prior Plastic Deformation (D. Kraynev, A. Bondarev, Z. Tikhonova)....Pages 1107-1114
Numerical Modeling of Heat Transfer and Material Flow During Wire-Based Electron-Beam Additive Manufacturing (A. V. Shcherbakov, D. A. Gaponova, R. V. Rodyakina)....Pages 1115-1125
Control of Weld Bead Position in Additive Manufacturing Process with Using Backscattered Electron Collector Signal (A. V. Shcherbakov, D. A. Gaponova, R. V. Rodyakina)....Pages 1127-1135
Operating Efficiency of Worm Gears Under Ultrasonic Vibration Imposition in the Cutting Region (S. I. Agapov, Yu. I. Sidyakin, A. F. Tolstyakov)....Pages 1137-1147
Features of Formation of Surface Layer Properties in Multistage Processing of Cr–Ni Steel (I. V. Firsov, Ju. L. Tchigirinskiy, N. V. Chigirinskaya)....Pages 1149-1157
Contactless Monitoring of Processed Surface Microrelief at Manufacturing Environment (A. P. Gontar, S. V. Mednikov, N. V. Chigirinskaya)....Pages 1159-1166
Characteristics, Composition, Mechanisms of Function and Modern Aspects of Implementation of Digital Production Systems in Mechanical Engineering Industry (A. R. Ingemansson)....Pages 1167-1174
Abrasive Machining of Low-Carbon Steels: Ways to Improve the Surface Quality (O. A. Kursin, S. B. Fam, N. A. Fedotov)....Pages 1175-1184
Features of Contact Interaction in Cutting High-Alloyed Steels with Carbide Tool (A. A. Lipatov, J. L. Tchigirinsky, Hoang Trung Pham)....Pages 1185-1191
Analysis of Influence of Strains of Technological System Elements on Machining Accuracy Under Turning of Non-rigid Shafts Based Between Centers (P. S. Nesterenko, J. L. Tchigirinsky, E. N. Nesterenko)....Pages 1193-1201
Peculiarities of Application of Th20 Hard Alloy for Turning Processing of Various Steels with Advanced Plastic Deformation (P. A. Norchenko, V. A. Solodkov, S. I. Kormilitsyn)....Pages 1203-1211
Application of Smoothing Rollers in Processes Finishing–Strengthening Treatment of Shafts’ SPD (Yu. I. Sidyakin, S. N. Olshtynsky, S. Y. Abakumova)....Pages 1213-1222
Preparing Automated of Software Complex for Technological Processes with Imposition of Electric Field (O. V. Skrygin, V. P. Smolentsev, E. A. Saltanaeva)....Pages 1223-1231
Technology of Combined Chemical–Mechanical Processing (V. P. Smolentsev, V. V. Ivanov, E. V. Panichev)....Pages 1233-1239
Technology of Combined Treatment of Engine Cooling Elements (V. P. Smolensev, A. V. Shchednov, J. S. Smolenseva)....Pages 1241-1247
Temperature in Intermittent Cutting (V. A. Solodkov, S. I. Kormilitsin, P. A. Norchenko)....Pages 1249-1259
Increasing Calorific Value of Biogas by Steam Explosion Activation of Renewable Raw Materials (D. B. Prosvirnikov, A. R. Sadrtdinov, Z. G. Sattarova)....Pages 1261-1270
Application of Statistical Modeling Methods to Assess Decontamination Effect of Electromagnetic Field on Raw Materials for Food Industry (N. N. Ovchinnikova, G. S. Kochetkova, T. A. Tolmacheva)....Pages 1271-1283
Energy Consumption Modeling of Machining Processes (V. Salnikov, Yu. Frantsuzova)....Pages 1285-1294
Design and Calculation Method of Composite Housings for New Generation Magnetorheological Devices (K. V. Naigert, V. A. Tselischev)....Pages 1295-1303
Methodology and Constructive Implementation of Active Vibration Protection of Large-Scale Structures (K. V. Naigert, V. A. Tselischev)....Pages 1305-1313
Substantiation of Parameters of Machine with Volumetric Hydraulic Drive for Formation of Wells in Ground (A. V. Gorin, N. V. Tokmakov, I. S. Kyznetsov)....Pages 1315-1323
Research of Pneumodrive with Energy Recovery into Additional Volume (A. N. Sirotenko, S. A. Partko, S. A. Voinash)....Pages 1325-1333
Model of Airflow Process Through Throttling Sections of Automated Deadweight Absolute Pressure Measurement System (A. Markov)....Pages 1335-1347
Investigation of Radial Gas Bearings with Longitudinal Micro-Grooves of Various Transverse Profiles (I. V. Vishtak, V. A. Fedotov, A. N. Solomon)....Pages 1349-1356
Flexible Composite Diversion Water Conduits of Small HPPs for Recreational Facilities in the Republic of North Ossetia (D. V. Kasharin, S. A. Kalmikov, O. A. Surzhko)....Pages 1357-1367
Experimental Research on Reducing Hydraulic Resistance When Transporting High-Viscosity Fluids by Pipeline (L. Ilina, N. Goncharov, A. Shagarova)....Pages 1369-1376
Modeling Three-Dimensional Liquid Flows in Computer-Controlled Vibrojet Mixer Using FlowVision (Yu. S. Sergeev, S. V. Sergeev, G. E. Karpov)....Pages 1377-1385
Improving Efficiency of Boiler in Case of Coal Hydrotransport (K. V. Osintsev, M. M. Dudkin, Iu. S. Prikhodko)....Pages 1387-1394
Finding Flow of Non-Newtonian Fluids in Circular Pipe with Wall-Adjacent Gas Layer (L. Ilina, P. Vasilyev, M. Krasnodubrovsky)....Pages 1395-1403
Mathematical Model of Gas-Dynamic Temperature Transducer (V. V. Korzin, D. B. Melekhov)....Pages 1405-1418
Results of Determining Optimal Correlation Between Components of Biodiesel Fuel on Basis of Rape (D. V. Varnakov, V. V. Varnakov, S. A. Simachkov)....Pages 1419-1426
Decrease in Destructive Environmental Impact and Fuel Consumption in Internal Combustion Engine of Vehicles as Result of Using Aluminium and Hydrogen Technologies (I. K. Andronchev, D. Ya. Nosyrev, A. A. Mishkin)....Pages 1427-1434
Automated Resource-Saving System for the Use and Regeneration of Epilam-Based Lubricating-Cooling Technological Liquid (N. R. Bukeikhanov, S. I. Gvozdkova, E. V. Butrimova)....Pages 1435-1442
Use of Wastes from Metalworking Machining for Packings in Contact Heat-and-Mass Exchange Devices (N. A. Merentsov, A. V. Persidskiy, V. N. Lebedev)....Pages 1443-1454
Natural and Energy Resource Saving Based on the Development of Technology for Profile Milling of Wood Waste (A. A. Fomin, R. V. Yudin, A. R. Sadrtdinov)....Pages 1455-1464
Performance Evaluation of Static Mixers in the Urea Injection Pipe for SCR Systems (A. Blinov, N. Malastowski, A. Bykov)....Pages 1465-1473
Comparison of Lifting Mechanisms for Raising Wind Wheel in Mobile Power Complex Based on Renewable Energy Sources (A. Kulganatov, Ahmed Ibrahim, A. Miroshnichenko)....Pages 1475-1482
Study of Steam Injection Effect on Course of Combustion Processes in Combustion Chamber of Gas Turbine Unit (D. A. Akhmedzaynov, A. E. Kishalov, V. D. Lipatov)....Pages 1483-1492
Prediction of Road Accidents’ Severity on Russian Roads Using Machine Learning Techniques (D. Donchenko, N. Sadovnikova, D. Parygin)....Pages 1493-1501
Computer-Aided Ecological and Profitable Scheduling of the Oil Depot Reservoirs Filling Process (E. Krushel, A. Panfilov, I. Stepanchenko)....Pages 1503-1512

Citation preview

Lecture Notes in Mechanical Engineering

Andrey A. Radionov Oleg A. Kravchenko Victor I. Guzeev Yurij V. Rozhdestvenskiy Editors

Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019) Volume II

Lecture Notes in Mechanical Engineering

Lecture Notes in Mechanical Engineering (LNME) publishes the latest developments in Mechanical Engineering—quickly, informally and with high quality. Original research reported in proceedings and post-proceedings represents the core of LNME. Volumes published in LNME embrace all aspects, subfields and new challenges of mechanical engineering. Topics in the series include: • • • • • • • • • • • • • • • • •

Engineering Design Machinery and Machine Elements Mechanical Structures and Stress Analysis Automotive Engineering Engine Technology Aerospace Technology and Astronautics Nanotechnology and Microengineering Control, Robotics, Mechatronics MEMS Theoretical and Applied Mechanics Dynamical Systems, Control Fluid Mechanics Engineering Thermodynamics, Heat and Mass Transfer Manufacturing Precision Engineering, Instrumentation, Measurement Materials Engineering Tribology and Surface Technology

To submit a proposal or request further information, please contact the Springer Editor in your country: China: Li Shen at [email protected] India: Dr. Akash Chakraborty at [email protected] Rest of Asia, Australia, New Zealand: Swati Meherishi at [email protected] All other countries: Dr. Leontina Di Cecco at [email protected] To submit a proposal for a monograph, please check our Springer Tracts in Mechanical Engineering at http://www.springer.com/series/11693 or contact [email protected] Indexed by SCOPUS. The books of the series are submitted for indexing to Web of Science.

More information about this series at http://www.springer.com/series/11236

Andrey A. Radionov Oleg A. Kravchenko Victor I. Guzeev Yurij V. Rozhdestvenskiy •

•

Editors

Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019) Volume II

123

•

Editors Andrey A. Radionov South Ural State University Chelyabinsk, Russia Victor I. Guzeev South Ural State University Chelyabinsk, Russia

Oleg A. Kravchenko Platov South Russian State Polytechnic University Novocherkassk, Russia Yurij V. Rozhdestvenskiy South Ural State University Chelyabinsk, Russia

ISSN 2195-4356 ISSN 2195-4364 (electronic) Lecture Notes in Mechanical Engineering ISBN 978-3-030-22062-4 ISBN 978-3-030-22063-1 (eBook) https://doi.org/10.1007/978-3-030-22063-1 © Springer Nature Switzerland AG 2020 This work is subject to copyright. All rights are reserved by the Publisher, whether the whole or part of the material is concerned, specifically the rights of translation, reprinting, reuse of illustrations, recitation, broadcasting, reproduction on microfilms or in any other physical way, and transmission or information storage and retrieval, electronic adaptation, computer software, or by similar or dissimilar methodology now known or hereafter developed. The use of general descriptive names, registered names, trademarks, service marks, etc. in this publication does not imply, even in the absence of a specific statement, that such names are exempt from the relevant protective laws and regulations and therefore free for general use. The publisher, the authors and the editors are safe to assume that the advice and information in this book are believed to be true and accurate at the date of publication. Neither the publisher nor the authors or the editors give a warranty, expressed or implied, with respect to the material contained herein or for any errors or omissions that may have been made. The publisher remains neutral with regard to jurisdictional claims in published maps and institutional affiliations. This Springer imprint is published by the registered company Springer Nature Switzerland AG The registered company address is: Gewerbestrasse 11, 6330 Cham, Switzerland

Preface

International Conference on Industrial Engineering took place on March 25–29, 2019, in Sochi, Russian Federation. The conference was organized by four universities—South Ural State University (National Research University), Moscow Polytechnic University, Platov South-Russian State Polytechnic University, and Volgograd State Technical University. The conference was carried out under financial support of the South Ural State University (National Research University). The conference was really large-scaled and international. The international program committee has selected more than 500 reports. The conferees represented 63 Russian cities from the western and central parts to the Far East regions. International participants represented such countries as China, Germany, Kazakhstan, Kyrgyzstan, Portugal, Saudi Arabia, Tajikistan, Ukraine, USA, Uzbekistan. The conference participants submitted papers reflecting recent advances in the field of Industrial Engineering, in Russian and English. The conference was organized into 13 sections, including Part 1 “Mechanical Engineering” (Machinery and Mechanism Design; Dynamics of Machines and Working Processes; Friction, Wear, and Lubrication in Machines; Design and Manufacturing Engineering of Industrial Facilities; Transport and Technological Machines; Mechanical Treatment of Materials; Industrial Hydraulic Systems; Green Manufacturing) and Part 2 “Materials Engineering and Technologies for Production and Processing” (Polymers, Composites and Ceramics; Steels and Alloys, Metallurgical and Metalworking Technologies; Chemical and Hydrometallurgical Technologies; Surface Engineering and Coatings; Processing and Controlling Technologies). The international program committee selected 294 papers from Part 1 of the conference technical sections for publication in book series “Lecture Notes in Mechanical Engineering.”

v

vi

Preface

The organizing committee would like to express our sincere appreciation to everybody who has contributed to the conference. Heartfelt thanks are due to authors, reviewers, participants and to all the team of organizers for their support and enthusiasm which granted success to the conference. Chelyabinsk, Russia

Andrey A. Radionov

Contents

Part 1 Modeling of Roll Roughness Transfer Process to Strip During Skin-Pass Rolling . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . N. N. Ogarkov, S. I. Platov and E. U. Zvyagina

1

Working Surface Calculation of Teeth Bevel Gear Helical-Bevel Gearing at Milling with Hob . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . E. A. Poluektov, B. A. Lopatin and S. V. Plotnikova

9

Effectiveness of Application of Additional Strengthening Processing of Surface Plastic Deformation on Increase in Fatigue Life of Parts . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . V. A. Lebedev, A. V. Kirichek and L. V. Chunakhova

17

Determination of Rational Technological Parameters of Cold Bending with Rolling of Pipes from Corrosion-Resistant Steels . . . . . . E. V. Khaliulin, A. V. Bobylev and A. V. Kozlov

27

Recomposition Procedure of Automatic Replacement Laser Modules for CNC Machines . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . P. A. Ogin and D. G. Levashkin

41

Tooling Backup of Cutting and Deforming Processing of Non-rigid Shafts . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . E. I. Yatsun, N. P. Anikeyeva and I. S. Karnaukhov

51

Determination of Flanging Parameters and Length of Screwing in Producing Holes by the Method of Thermal Drilling in Thin-Sheet Metal . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . P. V. Shalamov, I. A. Kulygina and A. N. Yasnitsky Practice of Dimensional Modeling in the Implementation of the Methodology of Group Interchangeability . . . . . . . . . . . . . . . . . M. G. Galkin and A. S. Smagin

59

69

vii

viii

Contents

Determination of Settling Efficiency of Solid Finely Dispersed Particles Within Devices with Rectangular Separators . . . . . . . . . . . . . I. N. Madyshev, O. S. Dmitrieva and A. V. Dmitriev

79

Method for Predicting Thermal Characteristics of Machine Tools Based on Experimental Modal Analysis . . . . . . . . . . . . . . . . . . . . . . . . A. N. Polyakov, A. N. Goncharov and I. V. Parfenov

85

Development of New Types of Contact Devices for Heat-Mass Transfer Apparatuses, Used at Petrochemical Enterprises . . . . . . . . . . I. N. Madyshev, O. S. Dmitrieva and A. V. Dmitriev

95

Stiffness Maximization on the Basis of Layout Characteristics of the Elastic Machine System and Milling Process . . . . . . . . . . . . . . . R. M. Khusainov and A. R. Sabirov

103

Formation Automation of Geometric Configuration of Real Machine Parts . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . V. E. Lelyukhin, O. V. Kolesnikova and O. M. Ponkratova

111

Information Model for Machine’s Electronic Structure Storage . . . . . . A. Loginov Software Spindle Speed Variation as Method for Chatter Suppression in Drilling . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . V. M. Svinin, A. V. Savilov and A. V. Shutenkov Experimental Studies of Steady-State Sources of Vibrations of Machinery Production Process Equipment to Substantiate Choice of Vibration Protection Methods . . . . . . . . . . . . . . . . . . . . . . . . S. I. Gvozdkova and L. E. Shvartsburg

121

131

141

On Issue of Verifying New Method for Studying Dynamics of Deep Hole Machining . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . L. Mironova, L. Kondratenko and V. Terekhov

151

Rational Provision of Robustness Properties of Bolted Joints of Assembly with Implementation of Anaerobic Materials . . . . . . . . . . I. I. Voyachek, D. V. Kochetkov and S. G. Mityasov

163

Machining Accuracy Improving with the Use of Mobile Mechatronic Systems as Industrial Robot End Effectors . . . . . . . . . . . . . . . . . . . . . . E. I. Shchurova and P. G. Mazein

171

Voxel and Finite Element Modeling of Twist Drill . . . . . . . . . . . . . . . . E. I. Shchurova 3D Modeling of Turbine Rotor Journal Machining with Location on a Bearing Bottom Half . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . A. V. Shchurova

181

191

Contents

Geometric Modeling of Macro-defects of Parts Surfaces Based on Discrete Solid-State Modeling . . . . . . . . . . . . . . . . . . . . . . . . . . . . . I. A. Shchurov Selecting Criterion of Long-Term Strength in Assessing Durability of Constructions Operating Under Non-isothermal Loading Processes . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . A. V. Belov, A. A. Polivanov and N. G. Neumoina Virtual Prediction of Accuracy of Processing on Example of External Circular Grinding . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . P. P. Pereverzev, A. V. Akintseva and M. K. Alsigar Designing of Optimal Grinding Cycles, Sustainable to Unstable Mechanical Processing on the Basis of Synthesis of Digital Double Technology, and Dynamic Programming Method . . . . . . . . . . . . . . . . . P. P. Pereverzev, A. V. Akintseva and M. K. Alsigar

ix

199

209

217

225

Controlled Assembly of Rotors . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . S. M. Beloborodov, V. F. Makarov and M. L. Tselmer

233

Study of Energy Intensity of Shaped Holes Broaching Process . . . . . . . V. V. Kuts, M. S. Razumov and V. S. Kochergin

241

Methods of Quality Control Manufacturing Gears of the Differential Satellites . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . D. T. Safarov and A. G. Kondrashov

249

Computational and Experimental Studies of Deformations of Air-Cooled Diesel Cylinders at Its Assembling . . . . . . . . . . . . . . . . . I. E. Agureev, K. Yu. Platonov and R. N. Khmelev

261

Feasibility Study of Applying Group Interchangeability Method for Assembly Components of Machines . . . . . . . . . . . . . . . . . . . . . . . . I. I. Voyachek, E. S. Slashchev and D. M. Malikova

271

Development of Integrated Criterion to Select Environmentally Sound Cutting Fluids and Relevant Application Systems in Shape-Forming Processes . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . L. E. Shvartsburg, O. V. Yagolnitser and E. V. Butrimova Gear-Grinding Temperature Modeling and Simulation . . . . . . . . . . . . N. V. Lishchenko and V. P. Larshin How Parameters of Agricultural Machine and Tractor Unit Affects Effectively Used Mean Indicated Power . . . . . . . . . . . . . . . . . . . . . . . . M. Ya. Durmanov, B. G. Martynov and S. V. Spiridonov

281 289

299

x

Contents

Method to Reduce Oil Burning in Diesel Engine of Agricultural Machine and Tractor Unit . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . V. P. Antipin, M. Ya. Durmanov and O. A. Mikhailov

313

Mill Conditions Effect on Roughness of Injection Molds’ Forming Surfaces . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . S. I. Kozhevnikov and V. F. Makarov

325

Effect of Magnetic Processing on Mass Transfer in a Frictional Pair “Alloyed Steel–Carbon Steel” . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . A. N. Gots, V. V. Zelinskiy and E. A. Borisova

337

Strength Parameters of Hardening Cylindrical Workpieces by Tapered Roller . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . A. A. Udalov, A. V. Udalov and S. L. Vasilevykh

347

Formation of Schemes Generating Geometric Structure of Machine Parts . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . O. V. Kolesnikova, V. E. Lelyukhin and F. Yu. Ignatev

355

Optimization of Cutting Parameters in Milling by Means of System Nyquist Plot . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . R. M. Khusainov, P. N. Krestyaninov and D. D. Safin

365

Comprehensive Evaluation of Shaft Manufacturability: Mathematical and Information Models . . . . . . . . . . . . . . . . . . . . . . . . . A. Sychugov, Yu. Frantsuzova and V. Salnikov

373

Simulation of Electrical Discharge Machining of Micro-holes . . . . . . . . T. D. Nguyen, V. M. Volgin and V. V. Lyubimov

381

Investigation of Kinematic–Geometric Characteristics of Electrochemical Machining . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . V. V. Lyubimov, V. M. Volgin and V. P. Krasilnikov

391

Model for Numerical Simulation of Temperature Field and Bead Profile in Hybrid Laser-Arc Welding of T-Joint . . . . . . . . . . . . . . . . . . S. Ivanov, E. Valdaytseva and I. Udin

399

QMS as Tool for Improving Maintenance and Repair Processes of Traction Rolling Stock . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . A. V. Muratov, V. V. Lyashenko and S. A. Petukhov

411

Heat Treatment Effect on the Structural and Elastic Characteristics of a Single-Component Abrasive Tool . . . . . . . . . . . . . . . . . . . . . . . . . M. Yu. Polyanchikova

419

Approach for Modeling and Situational Management of Industrial Product Efficiency . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . E. V. Orlova

427

Contents

xi

Modeling of Try-Out and Reliability Estimation of Technological Processes of Machine-Building Products . . . . . . . . . . . . . . . . . . . . . . . . V. M. Trukhanov, M. P. Kukhtik and A. M. Makarov

439

Methods of Cutting Cost Minimizing in Problem of Tool Route Optimization for CNC Laser Machines . . . . . . . . . . . . . . . . . . . . . . . . A. F. Tavaeva, A. A. Petunin and E. G. Polishchuk

447

Development of Hardware-Algorithmic System for ICE Diagnostics . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . L. A. Galiullin and R. A. Valiev

457

Information Support of Gas-Turbine Engine Life Cycle Based on Agent-Oriented Technology . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . A. Zagitova, N. Kondratyeva and S. Valeev

469

Automatic Print Job Scheduling and Management Over Multiple 3D Printers . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . I. A. Gushchin, I. V. Martynovich and I. S. Torubarov

477

CNC Processing Equipment’s Technical Operability Evaluation by Developing Mathematical Models Based on Continuous Logic of Antonyms . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . E. Krylov, N. Kozlovtseva and A. Kapitanov Digital In-Line Moisture Meter of Thin Sheet Materials . . . . . . . . . . . S. V. Makartichyan, L. V. Khoperskova and V. E. Avvakumov Modeling Physical Operating Principles During Search Design of Cooling and Refrigerating Systems . . . . . . . . . . . . . . . . . . . . . . . . . . A. A. Yakovlev, V. S. Sorokin and S. G. Postupaeva Analysis of Measurement System Accuracy Based on 2D Laser Triangulation Scanner When Measuring Soiled Pipe Thread Pitch and Height . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . D. S. Lavrinov

489 501

511

521

Mechanics of Elastic Wheel Rolling on Rigid Drum . . . . . . . . . . . . . . . M. Yu. Karelina, T. A. Balabina and A. N. Mamaev

531

Technical Audit of Rotary Aggregates . . . . . . . . . . . . . . . . . . . . . . . . . S. G. Mogilny, A. A. Sholomitskii and A. L. Sotnikov

541

Hydroblow as Mechanism of Additional Intensification of Liquid Forages Preparation in Centrifugal-Rotor Dismembrators . . . . . . . . . . Yu. N. Kamishov, N. A. Makarova and A. A. Sitnikov

551

Change of Structure of Dispersion Material Under Dynamic Loads . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . E. K. Chabutkin

559

xii

Contents

Increasing Efficiency of Vibratory Rollers Through Adjusting Magnitude of Disturbing Force . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Y. G. Popov and E. K. Chabutkin Energy Component of Properties of Material Crushability Layer . . . . Yu. A. Lagunova and V. S. Bochkov

567 577

New Criterion for Continuous Compaction Control Systems by Soil Vibratory Rollers . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . I. S. Tyuremnov and A. S. Morev

585

Reducing Amplitude of Load Swinging During Operation of Hydraulic Manipulators of Forest Transport Machines . . . . . . . . . . P. Popikov, M. Drapalyuk and D. Druchinin

595

Increase in Operational Reliability of Locking Device of Hydraulic Actuator of Vehicle . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . N. A. Fomenko, S. V. Aleksikov and S. G. Artemova

609

Influence of Hydrogen Additives on Cycle-to-Cycle Variability of Working Process of Rotary Engine . . . . . . . . . . . . . . . . . . . . . . . . . Y. V. Levin, K. V. Prikhodkov and E. A. Fedyanov

617

Efficiency of Usage of Transport and Technological Machines . . . . . . . G. V. Redreev, G. A. Okunev and S. A. Voinash

625

Differential System of Crane Braking . . . . . . . . . . . . . . . . . . . . . . . . . . S. V. Streltsov and V. A. Ryzhikov

633

Thermal Loading Estimation of the Friction Pairs of a Vehicle Automated Brake System . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . V. Dygalo and I. Zhukov

643

Problem of Increasing Tractive Effort of Railway Locomotives in Conditions of Arctic and Continental Shelf Regions . . . . . . . . . . . . . A. Keropyan, S. Albul and A. Zarapin

651

New Approach for Experimental Identification of Internal Combustion Engine Power and Performance Characteristics . . . . . . . . A. Egorov, N. Syutov and V. Belogusev

659

Development of Hardware and Software Complex for Increase of Technical Readiness Transport-Technological Machines in Forestry . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . S. V. Lyakhov and S. V. Budalin Study of Characteristics of Engine Operation in Stress-Testing Mode of Electric Gasoline Pump . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . A. Vozmilov, D. Vlasov and K. Glemba

667

679

Contents

xiii

Technical Level Analysis of Structures of Quarried Excavators . . . . . . H. N. Sultonov and K. Z. Tilloev

687

Reliability Analysis of Bus Steering System . . . . . . . . . . . . . . . . . . . . . A. N. Savkin, K. A. Badikov and A. A. Sedov

695

Process Research of Wheel-Rail Mining Machines Traction . . . . . . . . . A. M. Keropyan, D. A. Kuziev and A. E. Krivenko

703

Development of Double-Sided Summer–Winter Pneumatic Tires . . . . . I. Voiku and I. Komissarov

711

Asphalt Concrete as Object of Destruction by Operating Units of Milling Machines . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . D. V. Furmanov, V. A. Nikolayev and N. N. Klochko

719

Improving Automotive Torque Converter Quality . . . . . . . . . . . . . . . . N. N. Trushin, V. Y. Antsev and A. A. Obozov

727

Stabilization of Biaxial Trailer Motion . . . . . . . . . . . . . . . . . . . . . . . . . Yu. Stroganov, A. Popova and D. Zhelev

737

Part 2 Development of Optimization Algorithm to Control Open-Pit Excavator Operation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . O. Lukashuk, A. Komissarov and K. Letnev Complex for Inspection of Crane Rails Design . . . . . . . . . . . . . . . . . . . V. Yu. Antsev, P. V. Vitchuk and K. Yu. Krylov Study of Impact of Amount of Shock Absorbers on Parameters of Vibrations of Drum and Frame of Vibrating Roller . . . . . . . . . . . . . I. S. Tyuremnov, D. V. Fyodorova and A. S. Morev Optimization of High-Power Belt Conveyor Parameters . . . . . . . . . . . . G. G. Kozhushko, O. A. Lukashuk and T. A. Roscheva Intelligent Control, Correction, and Adaptability of Output Parameters of Vehicles’ Intake System . . . . . . . . . . . . . . . . . . . . . . . . . A. V. Gritsenko, I. V. Makarova and G. N. Salimonenko Duration of Ignition Delay of Fuel–Air Mixture in Diesel Engines . . . . V. F. Guskov and A. N. Gots Modeling of Maximum Cycle Pressure Based on Engine External Speed Performance . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . A. N. Gots and V. S. Klevtsov

747 755

765 775

783 795

805

xiv

Contents

Effectiveness of Road Transport Technology in Modern Housing Systems . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . A. V. Kulikov and S. Y. Firsova

813

Robot Manipulator Control with Efforts Stabilization in Capture of Object with Fuzzy Geometrical Characteristic . . . . . . . . . . . . . . . . . V. I. Chizhikov, E. V. Kurnasov and A. B. Petrov

823

Minimization of Energy Costs for Movement Resistance of Ground for Walking Device by the Control of Support Points Motion . . . . . . . V. V. Chernyshev, V. V. Arykantsev and I. P. Vershinina

839

How Different Autonomous Cutter Cooling Methods Affect Machining Performance . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . D. Yu. Dubrov

849

Calculation of Thermodynamic Parameters of Geometrically Complex Parts at Abrasive Globoid Gear Machining . . . . . . . . . . . . . . V. A. Spirin, V. F. Makarov and O. A. Khalturin

857

Providing of Surfaces’ Geometry at the Design Stage of Profile Milling Operation of Off-Grade Workpiece . . . . . . . . . . . . . . . . . . . . . A. A. Fomin, V. G. Gusev and N. F. Timerbaev

865

The Method of the Combined Flat Peripheral Grinding . . . . . . . . . . . . A. V. Morozov and V. G. Gusev

875

Defect Analysis of Operating Hydro-Gasified Piping System . . . . . . . . Min Ko Hlaing, Phone Htet Kyaw and B. N. Maryn

885

Contact Zone Effect Analysis onto the Deforming Rollers Geometrics During the Surface Plastic Deformation Treatment . . . . . . . . . . . . . . . Y. N. Oteniy, O. V. Martynenko and N. I. Nikiforov Increasing Resistance of Cutting Tool with Diamond Burnishing . . . . . N. Papsheva and O. Akushskaya

895 901

Determining Coordinates of Cutting Force Application Point in Grinding Zone . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . V. A. Nosenko and M. V. Danilenko

911

Improving Efficiency of Machining of Grooves on Shafts of Increased Hardness Structural Steel . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . S. V. Grubyi and P. A. Chaevskiy

921

General Patterns in Formation of Surface Layer of Machine Parts Treated by Combined Electro-technical Methods . . . . . . . . . . . . . . . . . S. V. Usov, P. A. Davydenko and D. S. Sviridenko

931

Contents

xv

Evaluation of Tool Life Equation of Single-Point Cutting Tool by Accumulation Model . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . A. V. Antsev, N. I. Pasko and A. V. Khandozhko

943

Ultra-precision Machining of Surfaces of Elements of Devices from Optical Materials . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . S. V. Grubiy, M. A. Shavva and V. V. Lapshin

953

Identification of Deformations and Errors in Flat Thin Workparts During Grinding . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . T. N. Ivanova

963

Novel Method of Single-Pass Threading by Cutter . . . . . . . . . . . . . . . . N. N. Zubkov

973

Chip-Forming Processes at High-Speed Grinding . . . . . . . . . . . . . . . . . V. M. Shumyacher, O. G. Kulik and S. A. Kryukov

983

Express Control of Abrasive Tool Operational Characteristics . . . . . . V. M. Shumyacher, S. A. Kryukov and O. G. Kulik

995

Study of Physical and Chemical Processes Occurring During Polycondensation of Bakelite Binder in Order to Adjust Technological Process of Abrasive Tool Production . . . . . . . . . . . . . . . 1003 I. Yu. Orlov, T. N. Orlova and I. V. Bashkirtseva Optimal Sequencing at Selection of Abrasive Tools’ Characteristics for Their Improvement . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1011 I. Yu. Orlov, N. V. Baidakova and P. Yu. Bochkarev Change in Microhardness of Metal Depending on Wetting Ability of Lubricating Coolant . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1019 I. V. Bashkirtseva and T. N. Orlova Studies on Titanium Alloy Turning Rate Improvement . . . . . . . . . . . . 1027 A. V. Savilov, V. M. Svinin and S. A. Timofeev Ultrasonic Impact Study on Strain Hardening of Thread Profile Surface Layer . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1035 V. Golovkin, O. Batishcheva and V. Papshev Dissipative Structure of Contact Interaction When Cutting Metals . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1043 V. A. Kim, B. Ya. Mokritsky and A. V. Morozova Cutting Temperature by Polymer-Abrasive End Brushes for Machining Planes . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1053 D. B. Podashev and Yu. V. Dimov

xvi

Contents

Researching the Influence of the Location Tool on the Treatment of Large Shafts Requiring Surface Shaping . . . . . . . . . . . . . . . . . . . . . 1061 Y. A. Bondarenko, N. A. Maslennikov and A. A. Mamchenkova Numerical Modeling of the Material Layer Upset Forging with Extrusion Under the Stiffening Rib into the Forging Cavity . . . . . 1071 O. A. Nikitina and T. M. Slobodyanik Centrifugal Rolling of Flexible Shafts for Achieving Best Possible Roughness of the Surface . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1079 N. V. Vulykh Thread Milling Cutter Flute Production Possibility Research by Using Typical Profiles Grinding Wheels . . . . . . . . . . . . . . . . . . . . . 1089 O. V. Malkov and I. A. Pavlyuchenkov Thermo-Emf as Method for Testing Properties of Replaceable Contact Pairs . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1097 Z. Tikhonova, D. Kraynev and E. Frolov Mathematical Apparatus for Predicting Cutting Tool Life in Turning Process After Prior Plastic Deformation . . . . . . . . . . . . . . . . . . . . . . . . 1107 D. Kraynev, A. Bondarev and Z. Tikhonova Numerical Modeling of Heat Transfer and Material Flow During Wire-Based Electron-Beam Additive Manufacturing . . . . . . . . . . . . . . 1115 A. V. Shcherbakov, D. A. Gaponova and R. V. Rodyakina Control of Weld Bead Position in Additive Manufacturing Process with Using Backscattered Electron Collector Signal . . . . . . . . . . . . . . . 1127 A. V. Shcherbakov, D. A. Gaponova and R. V. Rodyakina Operating Efficiency of Worm Gears Under Ultrasonic Vibration Imposition in the Cutting Region . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1137 S. I. Agapov, Yu. I. Sidyakin and A. F. Tolstyakov Features of Formation of Surface Layer Properties in Multistage Processing of Cr–Ni Steel . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1149 I. V. Firsov, Ju. L. Tchigirinskiy and N. V. Chigirinskaya Contactless Monitoring of Processed Surface Microrelief at Manufacturing Environment . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1159 A. P. Gontar, S. V. Mednikov and N. V. Chigirinskaya Characteristics, Composition, Mechanisms of Function and Modern Aspects of Implementation of Digital Production Systems in Mechanical Engineering Industry . . . . . . . . . . . . . . . . . . . . . . . . . . . 1167 A. R. Ingemansson

Contents

xvii

Abrasive Machining of Low-Carbon Steels: Ways to Improve the Surface Quality . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1175 O. A. Kursin, S. B. Fam and N. A. Fedotov Features of Contact Interaction in Cutting High-Alloyed Steels with Carbide Tool . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1185 A. A. Lipatov, J. L. Tchigirinsky and Hoang Trung Pham Analysis of Influence of Strains of Technological System Elements on Machining Accuracy Under Turning of Non-rigid Shafts Based Between Centers . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1193 P. S. Nesterenko, J. L. Tchigirinsky and E. N. Nesterenko Peculiarities of Application of Th20 Hard Alloy for Turning Processing of Various Steels with Advanced Plastic Deformation . . . . . 1203 P. A. Norchenko, V. A. Solodkov and S. I. Kormilitsyn Application of Smoothing Rollers in Processes Finishing–Strengthening Treatment of Shafts’ SPD . . . . . . . . . . . . . . . 1213 Yu. I. Sidyakin, S. N. Olshtynsky and S. Y. Abakumova Preparing Automated of Software Complex for Technological Processes with Imposition of Electric Field . . . . . . . . . . . . . . . . . . . . . . 1223 O. V. Skrygin, V. P. Smolentsev and E. A. Saltanaeva Technology of Combined Chemical–Mechanical Processing . . . . . . . . . 1233 V. P. Smolentsev, V. V. Ivanov and E. V. Panichev Technology of Combined Treatment of Engine Cooling Elements . . . . 1241 V. P. Smolensev, A. V. Shchednov and J. S. Smolenseva Temperature in Intermittent Cutting . . . . . . . . . . . . . . . . . . . . . . . . . . 1249 V. A. Solodkov, S. I. Kormilitsin and P. A. Norchenko Increasing Calorific Value of Biogas by Steam Explosion Activation of Renewable Raw Materials . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1261 D. B. Prosvirnikov, A. R. Sadrtdinov and Z. G. Sattarova Application of Statistical Modeling Methods to Assess Decontamination Effect of Electromagnetic Field on Raw Materials for Food Industry . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1271 N. N. Ovchinnikova, G. S. Kochetkova and T. A. Tolmacheva Energy Consumption Modeling of Machining Processes . . . . . . . . . . . . 1285 V. Salnikov and Yu. Frantsuzova Design and Calculation Method of Composite Housings for New Generation Magnetorheological Devices . . . . . . . . . . . . . . . . . 1295 K. V. Naigert and V. A. Tselischev

xviii

Contents

Methodology and Constructive Implementation of Active Vibration Protection of Large-Scale Structures . . . . . . . . . . . . . . . . . . . . . . . . . . 1305 K. V. Naigert and V. A. Tselischev Substantiation of Parameters of Machine with Volumetric Hydraulic Drive for Formation of Wells in Ground . . . . . . . . . . . . . . . . . . . . . . . 1315 A. V. Gorin, N. V. Tokmakov and I. S. Kyznetsov Research of Pneumodrive with Energy Recovery into Additional Volume . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1325 A. N. Sirotenko, S. A. Partko and S. A. Voinash Model of Airflow Process Through Throttling Sections of Automated Deadweight Absolute Pressure Measurement System . . . . . . . . . . . . . . 1335 A. Markov Investigation of Radial Gas Bearings with Longitudinal Micro-Grooves of Various Transverse Profiles . . . . . . . . . . . . . . . . . . . 1349 I. V. Vishtak, V. A. Fedotov and A. N. Solomon Flexible Composite Diversion Water Conduits of Small HPPs for Recreational Facilities in the Republic of North Ossetia . . . . . . . . . 1357 D. V. Kasharin, S. A. Kalmikov and O. A. Surzhko Experimental Research on Reducing Hydraulic Resistance When Transporting High-Viscosity Fluids by Pipeline . . . . . . . . . . . . . . . . . . 1369 L. Ilina, N. Goncharov and A. Shagarova Modeling Three-Dimensional Liquid Flows in Computer-Controlled Vibrojet Mixer Using FlowVision . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1377 Yu. S. Sergeev, S. V. Sergeev and G. E. Karpov Improving Efficiency of Boiler in Case of Coal Hydrotransport . . . . . . 1387 K. V. Osintsev, M. M. Dudkin and Iu. S. Prikhodko Finding Flow of Non-Newtonian Fluids in Circular Pipe with Wall-Adjacent Gas Layer . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1395 L. Ilina, P. Vasilyev and M. Krasnodubrovsky Mathematical Model of Gas-Dynamic Temperature Transducer . . . . . 1405 V. V. Korzin and D. B. Melekhov Results of Determining Optimal Correlation Between Components of Biodiesel Fuel on Basis of Rape . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1419 D. V. Varnakov, V. V. Varnakov and S. A. Simachkov Decrease in Destructive Environmental Impact and Fuel Consumption in Internal Combustion Engine of Vehicles as Result of Using Aluminium and Hydrogen Technologies . . . . . . . . . . . . . . . . . 1427 I. K. Andronchev, D. Ya. Nosyrev and A. A. Mishkin

Contents

xix

Automated Resource-Saving System for the Use and Regeneration of Epilam-Based Lubricating-Cooling Technological Liquid . . . . . . . . . 1435 N. R. Bukeikhanov, S. I. Gvozdkova and E. V. Butrimova Use of Wastes from Metalworking Machining for Packings in Contact Heat-and-Mass Exchange Devices . . . . . . . . . . . . . . . . . . . . 1443 N. A. Merentsov, A. V. Persidskiy and V. N. Lebedev Natural and Energy Resource Saving Based on the Development of Technology for Profile Milling of Wood Waste . . . . . . . . . . . . . . . . 1455 A. A. Fomin, R. V. Yudin and A. R. Sadrtdinov Performance Evaluation of Static Mixers in the Urea Injection Pipe for SCR Systems . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1465 A. Blinov, N. Malastowski and A. Bykov Comparison of Lifting Mechanisms for Raising Wind Wheel in Mobile Power Complex Based on Renewable Energy Sources . . . . . 1475 A. Kulganatov, Ahmed Ibrahim and A. Miroshnichenko Study of Steam Injection Effect on Course of Combustion Processes in Combustion Chamber of Gas Turbine Unit . . . . . . . . . . . . . . . . . . . 1483 D. A. Akhmedzaynov, A. E. Kishalov and V. D. Lipatov Prediction of Road Accidents’ Severity on Russian Roads Using Machine Learning Techniques . . . . . . . . . . . . . . . . . . . . . . . . . . 1493 D. Donchenko, N. Sadovnikova and D. Parygin Computer-Aided Ecological and Profitable Scheduling of the Oil Depot Reservoirs Filling Process . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1503 E. Krushel, A. Panfilov and I. Stepanchenko

Modeling of Roll Roughness Transfer Process to Strip During Skin-Pass Rolling N. N. Ogarkov, S. I. Platov and E. U. Zvyagina

Abstract In this paper, the modeling of a roll roughness transfer process to the strip in terms of skin-passing conditions was carried out. In the present paper, the analysis of the strain–stress state of material imprinted into microcavities of roll face is made, for which reason the latter is presented in the form of V-shaped grooves. The model of roll roughness transfer to the strip was developed by taking into account the real distribution of material in the rough layer of the roll and the strip, as well as according to the type of roll processing and skin-pass conditions that allows evaluating the degree of filling a singular microcavity under known pressure values of the strip on contact with a roll, friction coefficient, roll roughness parameters, and skin-pass conditions. The findings afford to forecast a reproduction of roll microrelief of the surface of deformed strip.







Keywords Microrelief Microrelief transfer mechanism Reproduction Microcavity Temper-rolled strip Coefficient of imprinting










1 Introduction One of the understudied key problems to be solved relates to relating the rolling conditions and roll roughness parameters and cold-rolled strip to the necessary required surface microgeometry of the finished product with account of the consumer performance [1–8]. The formation of the prescriptive strip microrelief directly near the roll pass is carried out by means of transfer mechanism of roll face microrelief to the strip and transformation of the initial microrelief [9–14].

N. N. Ogarkov
S. I. Platov
E. U. Zvyagina (&) Magnitogorsk State Technical University named after G.I. Nosov, 38, Lenin Av, Magnitogorsk 455000, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_1

1

2

N. N. Ogarkov et al.

2 Theoretical Research We define the quantitative evaluation of imprinting process of roll roughness to the skin-passed strip by the coefficient of imprinting K which is represented by ratio of material flow value of the skin-rolled strip zf to the size of roll microcavity characterized by a parameter Rz (Fig. 1). The problem will be solved for triangular roll microcavity the stress condition in which while filling while skin-rolled steel is shown in Fig. 1. Please do not alter the formatting and style layouts which have been set up in this template document. As indicated in the template, papers should be prepared in single column format suitable for direct printing onto paper. Do not number pages on the front, as page numbers will be added separately for the preprints and the proceedings. Leave a line clear between paragraphs. All the required style templates are provided in the file “MS Word Template” with the appropriate name supplied, e.g. choose “heading1” for your first-order heading text, “abstract” for the abstract text, etc. The determination of dimensionless stress necessary for filling the material into V-shaped cavities will be fulfilled by using Bocharov formula [15], which relative to adjacent toward one other roughness cavities takes the following form: " # ! P 4 1 pS ¼ pffiffiffiffiffiffi 3 ln ; þ z rS h 1  RfZ 3p

ð1Þ

where P—rolls pressure during skin-pass rolling; rs—yield strength of skin-rolled steel; Rz and S—height and stepwise roll face roughness parameter [16, 17]; h— thickness of outlet skin-pass strip.

Fig. 1 Stress condition diagram by filling V-shaped roll microcavity with skin-rolled steel

Modeling of Roll Roughness Transfer Process to Strip During …

3

z

Considering that Rfz ¼ K and solving an Eq. (1) in relation to K we shall obtain: 1

exp 13

1 hpffiffiffiffi

3pPcp 4rS

i:  pS h

ð2Þ

To evaluate the rolls’ pressure, we shall rely on the features of skin-pass rolling process and admit the following assumptions: • working rolls of the skin-rolling mill are drive, of the same diameter with the same roll face roughness; • form of roll contact with skin-rolled strip is flat; • dry skin-pass rolling is carried out with the friction coefficient at the deformation area … l ¼ 0; 25. . .0; 35. Due to higher friction coefficient values, the neutral point at the deformation area is located near the center of contact line. With due account of the assumptions, the pressure at the deformation area without regard to roughness contacting will be determined by the following equation:  P ¼ rd

h0 ð 1  e Þ lL



 exp

  lL 1 ; h0 ð 1  e Þ

ð3Þ

where e—skin-roll strip draft; L—contact length a roll—a skin-roll strip; µ— coefficient of friction at the roll pass; h0—thickness of skin-roll strip in input. Relating to the contact of roll deformed strip with large friction coefficient, it was shown in the papers [3, 18, 19] that a contact length a roll—a strip can be determined on the following formula: L¼

 qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 Del þ ðDelÞ2 þ 8Dh0 e ; 4

ð4Þ

where D—work roll diameter. The minimum pressure of skin-roll passing required for strip deformation shall be determined from the following formula [20]: rd ¼ 1:15ðrS þ d lg 1000_eÞ  rt ;

ð5Þ

where rs—yield stress determined by tension testing at a standard strain rate; d— dynamic coefficient considering the velocity effect when magnified ten times; e_ — strain rate during skin-roll passing; rt—tensile stress at the roll pass determined by strip tension stresses between a uncoiler and mill stand and a mill stand and coiler. The average strain rate e_ , at which a skin-roll passing is carried out, can be determined by approximation:

4

N. N. Ogarkov et al.

e_ ¼

2V ; Dl

ð6Þ

where V—velocity of roll periphery. The shortcoming of the Eq. (3) is that it includes the friction coefficient at roll pass, but it does not consider the fact that an imprinting of work roll surface roughness to skin-rolled strip is carried out within the limits of their rough layers and depends significantly on the fact to what extent the deformation processes of the skin-roll strip at roughness contact zone interact against each other and with the «base», on which they are located. During skin-roll passing, the pressure that defines the bearing power of rough layers shall be determined by their bearing surfaces. The regularity of change in bearing surfaces when approaching of roughness is determined by their supporting curves. Considering that the degree of involvement of roll face micro-irregularities and the strip is different, it is reasonable to use their medium integral values [21] that determine the relative quantity of the steel in rough layers. The approach of roll rough surfaces and skin-rolled strip will lead to the occurrence of total supporting surface of rough layers. We suppose that the total supporting surface is to be determined by a ratio: q¼

q1 þ q2 ; 2

ð7Þ

where q1 and q2—the percentage of material in rough layers of the roll and the strip. The parameter points q1 are presented in Table 1. The parameter q2 while strip rolling by sizing rolls shall be 0.43–0.59, depending on roll surfaces condition. In consideration of the foregoing, the actual contact pressure of rough surfaces at the roll pass shall be Psh ¼ P=q:

Table 1 Parameter points q1 for different cold roll surface finishing

ð8Þ

Processing type

Rmax/lm

q1

Grinding

4.7/3.2 2.4/1.6 1.2/1.0 8.6/5.8 4.8/3.2 1.2/0.8 12.1/7.2 9.6/4.9 4.2/1.2

0.402/0.485 0.423/0.519 0.550/0.566 0.482/0.538 0.503/0.564 0.548/0.576 0.489/0.543 0.533/0.564 0.567/0.576

Wheel-blasting

Electroerosion texturizing

Modeling of Roll Roughness Transfer Process to Strip During …

5

The dimensionless steel pressure of the skin-pass mill rolls with account of formulas (3), (4), (5), and (6, 7 and 8) shall be written as follows: " ! #     2V d lg 1000 Dl Psh 1 rt h0 ð1  eÞ lL 1:15 1 þ ¼ exp   1 : ð9Þ q lL h0 ð 1  e Þ rs rs rs The formula used for calculation the coefficient of imprinting will be finally written as follows: K ¼1

npffiffi h exp 6qp 1 þ

2V d lg 1000Dl rs





rt rs

ih

1 h0 ð1eÞ lL

in

h exp

lL h0 ð1eÞ

i

o o : ð10Þ  1  pS 3h

The application of the formula (10) allows to determinate the coefficient of imprinting of the roll roughness on the strip including the real distribution of material in a rough layer of the roll and the strip depending on the original roughness of roll face, the thickness of roll skin strip, rolling speed, tension stress, and drafting. The target values of the coefficient of imprinting by variances of skin-pass rolling are given in Fig. 2.

Fig. 2 Dependence of coefficient of imprinting K compared to experimental data—Δ □ ○ by skin-pass rolling of a strip of different thickness with original data D = 500 mm, e = 0.02, q = 0.5, d = 12 on: a stage of microroughness; b velocity of roll periphery; c tension capacity; d coefficient of friction

6

N. N. Ogarkov et al.

The resulting dependence allows concluding that the best reproduction of the roll microrelief on the skin-rolled strip is evidenced during “dry” skin-pass rolling of thin strip with little microroughness width. The resulting data are recommended for dry skin-pass rolling of the strips and for wet skin-pass rolling to a limited extent [22, 23]. The increase in strip tension dramatically reduces the imprinting of roll microrelief at the surface of skin-rolled strip. By increasing in the roll velocity, there is a minor effect on change in coefficient of imprinting especially for thick strips.

3 Conclusions The roll roughness transfer model to the strip was developed with account of the roll processing type and skin-pass rolling conditions that allow estimating the degree of filling of microrelief of singular microcavity at the known pressure values at the contact of the strip with the roll, friction coefficient, roll roughness parameters, and skin-pass rolling conditions. The resulting data can be used in amendments of the existing and developments of new skin-pass rolling conditions for cold-rolled strips to obtain the required microrelief according to the consumer properties by taking into account the coefficient of microgeometry of roll face of the skin-pass rolling mill.

References 1. Polukhin PI (1974) Contact interaction of steel and work tool during rolling. Metallurgy, Moscow 2. Ogarkov NN, Khalin SV (2003) Regulation of rolled products surface roughness depending on its consumer performance. In: Processes and machinery of metallurgic production: interregional collection of studies. Edition of Nosov Magnitogorsk State Technical University, Magnitogorsk, p 131 3. Roberts VL Cold rolling of steel: translated from English. Metallurgy, Moscow 4. Zvyagina EU (2017) Equipment upgrading and technological advancement of roll ragging for improvement of automotive sheet surface quality. Dissertation, Nosov Magnitogorsk State Technical University 5. Ogarkov NN, Zvyagina EU, Zaletov UD (2006) Effect of roll face processing technique on the quality of cold-rolled sheet. In: Modern methods of design and technology of mechanic engineering: international collection of studies. Edition of Nosov Magnitogorsk State Technical University, Magnitogorsk, p 39 6. Belov BK, Begletsov DO Modelling of generation process of surface microtopography during skin-pass rolling of automotive sheet. In: Salganik VM (ed) Modelling and development of process of metal treatment under pressure: interregional collection of studies. Edition of Magnitogorsk State Technical University, Magnitogorsk, p 58 7. Utsch M, Vinke P (2004) Roll texturing technology as a base of modern surfaces in automotive cold mill flat products. In: MS&T: Conference Proceeding, p 599

Modeling of Roll Roughness Transfer Process to Strip During …

7

8. Ogarkov NN, Zaletov UD, Laskov SA, Zvyagina EU, Pozhidaev UA (2010) Improvement of wheel blasting of mill rolls for production of automotive sheet, vol 2(30). Vestnik of Nosov Magnitogorsk State Technical University, p 41 9. Rasp W, Wichern CM (2002) Effects of surface-topography directionality and lubrication condition on frictional behavior during plastic deformation. J Mater Process Technol 125:379 10. Kuznetsov LA, Mamyshev AB (1989) Theoretical determination of coefficient of imprinting of roll microrelief into the strip during cold rolling, vol 6. News of Higher Educational Institutions: Iron and Steel Industry, p 38 11. Gorbunov AB, Radionov AF, Belov BK et al (2007) Production of automotive sheets with specified surface microtopography, vol 4. Rolling, p 15 12. Marigue C, Bragard A (1983) Surface roughness and user properties of cold rolled steel sheets. American Society for Metals, London, 11–13 May 1983, p 242 13. Kuznetsov LA, Mamyshev AB (1989) Theoretical determination of coefficient of imprinting (impression) of roll microrelief into the strip during cold rolling, vol 6. News of Higher Educational Institutions: Iron and Steel Industry, p 38 14. Mukhin UA, Ryblov AB, Bobkov EB, Tcherny BA (2014) Formation of surface microgeometry of cold rolled strips: study guide. Edition Lipetsk Technical University Press, Lipetsk 15. Bocharov Y, Kobayashi S, Thomsen EG (1962) The mechanics of the coining process. Trans ASME Series B J Eng Ind 84:491 16. Ogarkov NN, Zvyagina EYu, Zaletov YuD, Khomenko NN, Kerimova LF (2016) Improvements in shot blasting to increase the surface peak density of auto-industry steel sheet. Steel Transl 46(12):847 17. Tang Jing Gang (2004) Determination and analysis of surface roughness of cold rolled steel sheets, vol 2. Iron Steel Vanadium Titanium, p 66 18. Aliev IC (2002) Study of contact plastic friction factor. In: Upgrading of processes and machinery for pressure shaping in steel industry and mechanic engineering: collection of studies. DGMA, Kramatorsk, p 112 19. Vasiliev JD, Dementienko AB (2001) Study of contact line with a roll during cold rolling. In: Ferrous Metals, vol 7. News of Higher Educational Institutions, p 21 20. Roberts WL (1972) An approximate theory of temper rolling. In: Iron and steel engineer year book, p 530 21. Ogarkov NN (1996) Roughness formation of the rolled products with high-quality surface finishing by means of regulation control of roll surface coating condition. Dissertation, Nosov Magnitogorsk State Technical University 22. Terentiev DV, Ogarkov NN, Platov SI, Kozlov AB (2018) Effect of operational conditions and oil consumption of contact areas on the thickness of lubricating film in heavy duty friction assemblies of melting facilities. In: Ferrous metals, p 60 23. Ogarkov NN, Platov SI, Shemetova ES, Tepentev DV, Nekit VA, Samodurova MN (2017) Oil absorption capacity of the contact surfaces in metal-forming processes. Metallurgist 61(1– 2):58

Working Surface Calculation of Teeth Bevel Gear Helical-Bevel Gearing at Milling with Hob E. A. Poluektov, B. A. Lopatin and S. V. Plotnikova

Abstract The paper considers an internal gearing of a bevel gear and a spur gear, shaped according to a conventional technology. There is a technological problem of obtaining the working profile of the bevel gear teeth in the internal spur–bevel gearing when the working profiles are being shaped. Obtaining a theoretically accurate surface of the gear teeth is troublesome due to the difficulty of making cutters with internal teeth and practical realization of machine gearing with the axis inclination of the workpiece or the cutter axis. This is related to the need of producing a shaping cutter with internal teeth for each gear train so that its geometry would be identical to that of the spur gear in the gear train. The paper describes a method for shaping an approximate teeth profile of a bevel gear in a spur–bevel gearing with a rack-type tool. The proposed method provides a sufficient degree of approximation of the shaped surface to the theoretically accurate one, which makes it applicable. In the paper, the equations are presented describing the teeth surface, as a result of a two-parameter bending with a hob of the surface with conical billet for evaluation. This makes it possible to evaluate the degree of approximation received for the working surface of the teeth to the theoretically accurate surface. Keywords Spur–bevel gearing Gear milling


Internal meshing
Non-involute gear


E. A. Poluektov
B. A. Lopatin (&)
S. V. Plotnikova Zlatoust Branch of South Ural State University (National Research University), 16, Turgenev Street, Zlatoust 456209, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_2

9

10

E. A. Poluektov et al.

1 Introduction A spur–bevel gearing is a gear train, in which one of the gears has a cylindrical blank, and the other—a conical one [1–5]. To provide a linear teeth contact in the internal spur–bevel gearing, the generating gear in the cutter-blank meshing has to be an involute spur gear identical to the spur gear of the gear train [6–9].

2 Main Part Figure 1 shows a diagram of the internal spur–bevel meshing with a generating involute spur gear. The teeth flank of the spur gear is an involute cylinder. This surface (see Fig. 1) in the moving reference frame X2Y2W2, associated with the spur gear, is described by the equations: X2 ¼ rb2 ½sinðmy2  wb2 Þ  my2
cosðmy2  wb2 Þ Y2 ¼ rb2 ½cosðmy2  wb2 Þ þ my2
sinðmy2  wb2 Þ W2 ¼ u;

ð1Þ

where rb2 is the radius of the base spur gear cylinder, my2 is the roll angle of the involute curve, wb2 is half of the angular thickness of the tooth dedendum on the rolling circle of the spur gear, u is the z-coordinate of the face section of the spur gear.

Fig. 1 Meshing of the generating: 1—spur gear, 2—bevel gear

Working Surface Calculation of Teeth Bevel Gear …

11

The angle my2 for the given radius ry2 is determined by the expression my2 ¼ tg
arccos

  rb2 : ry2

ð2Þ

The angle wb2 is calculated according to the dependence wb2 ¼

p 2
x2
tg a þ þ inv a 2
z2 z2

ð3Þ

where x2 is the profile shift coefficient of the spur gear. The flank surface of the bevel gear teeth is an envelope of the generating surface and is non-involute [2]. This surface in the moving reference system X1Y1W1, associated with the bevel gear is described by the equations: X1 ¼ rb1 ½cos /2 ðsin atw  my1 cos atw Þ  sin /2 cos Rðcos atw þ my1 sin atw Þ þ u sin /2 sin R Y1 ¼ rb1 ½sin /2 ðsin atw  my1 cos atw Þ  cos /2 cos Rðcos atw þ my1 sin atw Þ

ð4Þ  u cos /2 sin R W1 ¼ rb1 sin Rðcos atw þ my1 sin atw Þ þ u cos R cos atw ¼ cosðmy1  wb1  /1 Þ rb1 ¼ ; / ¼ /2
i12 u
tg d1 1 where R is the shaft angle, atw is the gearing angle in the face section of the spur gear, u1, u2 are the turning angles of the bevel gear and the spur gear. Theoretically, exact flank surface of the bevel gear teeth can be formed by a shaping cutter with internal teeth. In this case, a geometrical gearing diagram of the internal conjugate spur–bevel meshing is implemented in the cutter-blank meshing. However, due to the complex manufacture of shaping cutters with internal teeth, this tool is not produced on the industrial level. Thus, currently it seems impossible to form a theoretically exact profile using this method. An approximate profile can be shaped with gear-cutting equipment using a rack-type cutting tool [10]. In this case, the rack-type cutting tool (a milling cutter, a grinding wheel) moves along the blank gear axis according to a certain law. Such cutting is carried out by standard gear milling machines equipped with a follow-up device or by CNC machines. Figure 2 shows the cutting diagram. By selecting the cutter path, we can obtain a tooth similar in shape to the exact non-involute gear tooth [11, 12]. To calculate the coordinates of the envelope curve points, we used the mathematical apparatus of the involute bevel gearing developed by Bezrukov [13] and took the tool angle d0i as variable in each section. The tool shifting coefficient in an arbitrary face section of the bevel gear in the midpoint of the tooth depth is determined according to the expression.

12

E. A. Poluektov et al.

Fig. 2 Meshing of the non-involute bevel gear blank and a hobbing cutter: 1—billet of the gear, 2— cutter path, 3—milling cutter

 xti ¼

 sti p z   inv ati þ inv atymid i 2tg ati 2rmid i 2z

ð5Þ

where sti is the tooth thickness in the depth midpoint of the theoretically exact profile; atymid i is the face angle on the radius rmid i ; z is the number of teeth of the bevel gear to be cut; rmid i is the radius of the bevel gear corresponding to the midpoint of the tooth depth. Having determined the value of the profile shift coefficient xti for different face sections of the gear ring, we obtained the following equation by approximation: xti ¼ aw2 þ bw þ c

ð6Þ

where a, b, c are the coefficients that were obtained by approximation; w is the zcoordinate of the face section of the bevel gear. Multiplying this equation by the module, we obtained the equation of the reference surface envelope of the tool y ¼ xti
m ¼ ðaw2 þ bw þ cÞm:

ð7Þ

The current value of the angle d0i between the tangent to the curve, described by Eq. (7), and the y-axis (see Fig. 2) is determined by the formula d0i ¼ arc tgð2awi þ bÞ:

ð8Þ

The motion trajectory of the milling cutter axis is equidistant to the envelope of the tool reference surface. The distance between these curves is equal to the pitch radius r0 of the milling cutter.

Working Surface Calculation of Teeth Bevel Gear …

13

Figure 2 allows us to determine the current coordinates of the points Oi on the axis trajectory of the milling cutter: y0i ¼ aw2 þ bw þ c þ r0 cos d0i ; w0i ¼ w þ r0 sin d0i

ð9Þ

where the ordinate y0i is equal to the current inter-axial distance ayi in the cutter-blank gearing. Finally, we obtain y0i ¼ ayi ¼ a0 w2i þ b0 wi þ c0

ð10Þ

The coefficients of this equation are found similar to those of expression (6). The proposed method allows us to obtain the cutting surface of the bevel gear teeth approximate to the theoretically exact non-involute surface described by Eq. (4). To estimate the deviations of the cut profile from the theoretically exact one, it is necessary to have a mathematical description of the working surface of the bevel gear teeth, when the gear is two-parameter enveloped with a rack-type tool moving along a curvilinear toolpath.

Fig. 3 Diagram of the cutter-blank meshing when cutting a non-involute bevel gear

14

E. A. Poluektov et al.

When deriving the equation of the flank surface of the non-involute bevel gear teeth, we used a mathematical apparatus developed by N. N. Krylov and M. L. Erihov for two-parameter enveloping [14]. Figure 3 shows the relative position of the pitch surfaces and the applied reference frames. To obtain the equation of the surface under study, the equation of the surface of the milling cutter teeth was revised in line with the system associated with the gear blank, and the obtained dependences were supplemented by two equations of meshing. Having made all the transformations, we obtained: x1 ¼ x0 cos /1 þ y0 sin /1 y1 ¼ x0 cos /1 þ y0 sin /1 w1 ¼ ðx þ r0 Þ sin d0i þ g cos d0i þ wp

ð11Þ

where

ð12Þ The parameters used in the calculated dependences have the following notations (index 0 refers to the tool, index 1—to the spur gear to be cut): m is the normal module (standard); z0 is the number of cutter entries; k0 and kb0 are the lead angles of the helical curve on the pitch cylinder and the base cylinder; a and at are the pressure angles of the tool in the normal section and face section (a = 20°); p is the helix parameter; r0 and rb0 are the radii of the reference cylinder and the main cylinder; t and u are the curvilinear coordinates (parameters) of the involute helical surface; u0 (u1) and wp are the enveloping parameters (the angle of rotation of the milling cutter or the blank and z-coordinate of the milling cutter point).

Working Surface Calculation of Teeth Bevel Gear …

15

Expressions (11), (12) are the equations of the flank surface of a bevel gear tooth, cut by a hobbing cutter, where (12h and i) are the transformed equations of meshing with wp ¼ const and / ¼ const, respectively.

3 Conclusion The obtained equations allow us to estimate the profile shift of the surface cut by a rack-type tool and the theoretically exact profile. A comparative analysis of the profiles showed that maximum deviations are arise in some gearings in the farthest face sections of the gear and obtain to hundredths fraction of a millimeter. When larger ends have these deviations, the resulting tooth profile is wider than the theoretical one, which can disrupt the gearing due to interference or misalignment in the gear meshing. The profile deviations can be eliminated by setting teeth modifications when the bevel gear is being cut [15]. The modification consists in adjusting addendum modification coefficients which were used to calculate the cutter trajectory in those face sections of the bevel gear where the greatest profile deviations are observed. The modification degree for each gearing is specific [16]. The method for cutting bevel gear teeth was used in the design and manufacture of planetary reduction drives with spur–bevel gearings [17–19]. Such reduction drives have a number of advantages; in particular, their load capacity is comparable to that of strain wave gearings, and their service life is much longer. Moreover, they are self-braking, which makes the use of additional braking devices unnecessary and makes it possible to select gear backlashes [20, 21].

References 1. Davydov YS (1950) Non-involute gearing. Mashgiz, Moscow 2. Lopatin BA (1998) Development of theoretical bases of design, manufacture and testing of helical bevel gears with small interaxle corners. Dissertation, Kalashnikov ISTU 3. Lopatin BA, Plotnikova SV (2017) Helical-bevel gearing with small wheel axles angles. In: International conference on industrial engineering, ICIE 2017. Procedia Eng Ser 1189–1194 4. Lopatin BA, Tsukanov ON, Plotnikova SV (2003) Cylinder-conical toothed gears in car drives. Bull Mach Build 8:7–9 5. Lopatin BA, Tsukanov ON (2005) Cylinder-conical gears. Monograph. SUSU, Chelyabinsk 6. Lopatin BA, Zaynetdinov RI, Plotnikova SV (2014) The control of the teeth of the non-involute gear of the cylinder-conical transmission. Bull SUSU Ser Mech Eng 14(1): 52–58 7. Lopatin BA, Plotnikova SV, Khaustov SA (2015) Involute helical-bevel gearing. In: International Conference on Industrial Engineering, ICIE 2015. Procedia Eng Ser 891–895 8. Lopatin BA, Plotnikova SV (2016) Finishing of the helical-bevel gear teeth flanks. Procedia Eng 150:889–893

16

E. A. Poluektov et al.

9. Lopatin BA, Plotnikova SV (2017) Helical-bevel gearing with small wheel axles angles Procedia Eng 206:1189–1194 10. Lopatin BA, Poluektov EA, Lopatin DB, Zaynetdinov RI, Rublev VM (2009) Method for cutting non-involute gear teeth of an internal spur-bevel gearing. RU Patent 2364480 C1, IPC B 23 5/24/–No. 2008117944; decl. May 2008 11. Plotnikova SV, Polouektov EA (2013) The method of teeth profiling of cylindro-bevel gears of internal engagement. Science SUSU, Chelyabinsk, pp 329–332 12. Lopatin BA, Zainetdinov RI (2018) Forming of tooth profiles of non-involute bevel gear on CNC machines. Intell Syst Prod 16:53–57 13. Bezrukov VI (1986) Gear trains with involute bevel gears. Reference for geometric calculation of involute and worm gearings. Mechanical engineering, Moscow, pp 254–261 14. Litvin FL (1968) Theory of gearings. Science, Moscow 15. Plotnikova SV, Poluektov EA (2009) Ensuring the required profile accuracy of a non-involute gear through longitudinal teeth modification. Sci SUSU 1:292–294 16. Polouektov EA, Lopatin SD (2011) Modification of tooth profiles in cylindro-bevel gear internal engagement. Science SUSU, Chelyabinsk, pp 160–164 17. Lopatin BA, Khaustov SA, Poluektov EA, Burnazyan SR, Zaynetdinov RI (2009) Software package for calculating and analyzing the geometry of spur-bevel gearings. Certificate of state registration of the software 200961002/–No. 2009614819; decl. April 2009 18. Bruzhas VV, Lopatin BA (2015) Development of solid-state models for the gears of different geometry. Procedia Eng 129:369–373 19. Eremin VP, Eremin NV, Kirillin AN et al (2015) Creation of a new generation of electromechanical drives of trans-formed spacecraft systems. FSUE “GNPRKC” TsSKB-Progress, Samara, pp 23–61 20. Tsukanov ON, Lopatin BA, Bruzhas VV (2015) Planetary gear. Patent for utility model RUS 159017, May 2015 21. Lopatin BA, Plotnikova SV (2015) Cilingrical gear formed by helical-bevel gearing. Sci Educ Mod Soc 8:120–121

Effectiveness of Application of Additional Strengthening Processing of Surface Plastic Deformation on Increase in Fatigue Life of Parts V. A. Lebedev, A. V. Kirichek and L. V. Chunakhova

Abstract The paper presents the results of experimental studies of the effect of additional surface plastic deformation hardening on increasing the fatigue life of parts. Hardening by the surface plastic deformation method leads to surface hardening, to the formation of details of residual stresses of compression in the surface layers, and to the favorable change of microgeometry of surfaces, which result in the increase in their fatigue and contact resistance, wear resistance, and corrosion resistance depending on the functional purpose and an external environment of details. The conducted pilot studies provided reasons for the modes of the vibroshock hardening processing of samples and determination of fatigue longevity of the samples strengthened according to five regulations. It is shown that the application of a single additional vibroshock hardening after intermediate cyclic loading (training) contributes to an increase in the fatigue life of the sample approximately by 1.2 times; double additional vibroshock hardening of samples in combination with a double intermediate cyclic loading increases fatigue durability of samples by 1.1 times; post-deformation rest of the samples has no significant effect on increase in a fatigue longevity that is caused by a stress relief or a softening in the course of aging. It has been specified that additional hardening of parts during operation will have effect under the condition that the degree of hardening of the surface layer provided by the surface plastic deformation does not exceed the maximum permissible value peculiar to a particular material. Keywords Hardening


Plastic deformation
Fatigue life

V. A. Lebedev (&)
L. V. Chunakhova Don State Technical University, 1, Gagarin Sq., Rostov-on-Don 344000, Russia e-mail: [email protected] A. V. Kirichek Bryansk State Technical University, 17, 50-letiya Oktyabrya Blvd., Bryansk 241035, Russia © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_3

17

18

V. A. Lebedev et al.

1 Introduction Reliability and service life of machine parts mainly depend on the qualitative state of their surface layer, which is the carrier of structural, technological, and operational stress concentrators, which value and predetermine the fatigue failure of the construction in operation. The most significant properties of the part surface layer in terms of operation, determining the fatigue strength, are roughness, microhardness, and the level of residual stresses [1–19]. The required parameters of surface quality and almost the majority of the most important operational properties of machine parts can be provided by their surface plastic deformation hardening, showing the maximum potential of the material. Surface plastic deformation hardening leads to an increase in surface hardness, the formation of residual compression stresses in the surface layers of parts and a favorable change in the microgeometry of surfaces. As a result, depending on the functional area and operating conditions of parts their fatigue and contact strength, wear resistance, corrosion resistance increases [3, 4, 9, 10, 12, 13, 15–18, 20, 21]. The aim of the research was to study the effectiveness of additional surface plastic deformation hardening on increasing the fatigue life of parts during their operation.

2 Research Methods The methods of experimental studies included: • Justification of modes of shock-vibrating hardening of samples; • Determination of fatigue life of samples hardened according to the following regulations: Regulation 1. Vibroshock processing and cyclic hardening of samples is carried out in two stages. At the first stage the samples are hardened during the time (t), at which the greatest effect of hardening Y(t) is reached, and, at the second stage, the samples are exposed to their subsequent cyclic loading Tp(N) to destruction, where N—is a number of cycles before destruction. According to this study plan, the influence of effective extent of hardening of samples is established, which mainly depends on the hardening time, on their fatigue life. Regulation 2. Vibroshock processing and cyclic hardening of samples are carried out in two stages. At the first stage, the samples are hardened during the time t1 (which makes 50–60% of t) providing their hardening degree Yt1), then, at the second stage, which is carried out after cyclic loading of samples Tp(N1), the samples are additionally hardened at the same modes during time t2 (which makes 50–40% of t). After additional hardening Y(t2), the samples are subjected to further cyclic loading Tp(N2) to destruction. According to this study plan, the effect of additional shock-vibrating hardening on the increase in fatigue life is established, which is estimated by total quantity of cycles of loading on the first N1 and the second N2 stages.

Effectiveness of Application of Additional Strengthening …

19

Regulation 3. Vibroshock processing and cyclic hardening of samples is carried out in three stages. At the first stage, the samples are hardened during the time t1 (which makes 30% of t) providing their hardening degree Y(t1), then, at the second stage the samples are exposed to the intermediate cyclic loading Tp(N2), where N2 = twenty thousand cycles of loading. At the third stage, the samples are additionally processed on the same modes during time t3 (which makes 30–40% of t) providing their hardening degree Y(t3). After additional hardening Y(t3), the samples are subjected to further cyclic loading Tp(N3) to destruction. According to this study plan, the effect of double additional shock-vibrating hardening on the increase in fatigue life is established, which is estimated by total quantity of cycles of loading on the first N1, the second N2, and the third N3 stages. Regulation 4. Analyzes the effect of recovery (aging) O after hardening Y(t) on the fatigue life of samples hardened according to regulations 1. Regulation 5. Analyzes the effect of recovery (aging) O after hardening Y(t) on the fatigue life of samples hardened according to regulations 2. Surface plastic deformation hardening was performed by shock-vibrating method on a laboratory vibratory machine UVG 4  10. Hardened steel balls with diameter 5–8 mm were used as an operation environment for shock-vibrating machining. Operating conditions of the samples were imitated on flat samples made of aluminum alloy D16T with stress concentrator (Fig. 1a), subjected to transverse bending in one plane on installation for fatigue testing UI–20 (Fig. 1b). Post deformation rest of the strengthened samples was carried out by keeping them at a temperature of 18–20 °C during 3–18 days. As a criterion for justification of the optimal modes of shock-vibrating machining (frequency f, amplitude A, machining time t) was chosen physicomechanical characteristic of the surface layer quality, formed in processing, the size of the squeezing residual stresses estimated in size of a sagging deflection of flat samples.

Fig. 1 a Flat sample for fatigue testing; b fatigue machine UI–2

20

V. A. Lebedev et al.

3 The Results of Studying of Impact of the Additional Vibroshock Strengthening Processing on Fatigue Life Figure 2 illustrates the results of researches of influence of duration of vibroshock processing on change of physicomechanical characteristics of the surface layer quality. It is established that at purity of fluctuations—33 Hz, a vibration amplitude of 2.5–3.5 mm the greatest increase of size of a sample deflection—the witness and as a result the greatest effect of hardening caused by increase of physicomechanical characteristics of the surface layer quality (without destruction) is reached at processing time t = 90 min. Further increase in time of processing, as it follows from results of researches, doesn’t stimulate the increase in effectiveness of hardening. To confirm this experimentally proven fact, a series of researches had been carried out to investigate the impact of duration of vibroshock hardening on a fatigue life of the samples, which demonstrate that with the increase in the degree of sample hardening due to the duration of hardening, the fatigue life rises until the surface layer reaches the maximum degree of hardening corresponding to machining time of 90 min. At the same time, the indicators of fatigue life of samples increased compared to the initial by 9 times (Fig. 3). The reduction in fatigue life of samples which had been hardened for more than 90 min is caused by over hardening of the surface layer samples. Above given researches allowed to determine efficient, in terms of increase in fatigue life of the samples, conditions of the additional surface plastic deformation hardening by a vibroshock method. These processing conditions were taken as a basis while carrying out further pilot studies on estimation of effectiveness of influence of regulations of hardening on increase in fatigue life of the samples. The results of these studies are given in Table 1.

Fig. 2 Effect of machining time on the sample deflection value

Effectiveness of Application of Additional Strengthening …

21

Fig. 3 Impact of hardening degree on fatigue life of samples

Table 1 Indicators of fatigue life of the samples depending on the regulations of vibroshock hardening No. of regulation

Hardening regulation

Initial sample 1 2 3

Without hardening

4 5

Y(90 min) + Tp Y(60 min) + Tp(20 thousand of cycles) + Y(30 min) + Tp Y(30 min) + Tp(20 thousand of cycles) + Y(30 min) + Tp(20 thousand of cycles) + Y(30 min) + Tp Y(90 min) + O(3 days) + Tp Y(60 min) + O(18 days) + Tp(60 thousand of cycles) + Y(30 min) + O(3 days) + Tp

Fatigue life of the samples, N  103 44 452 547 500 219 204

The comparative study of the data represented in Table 1 has shown that: • The use of single additional shock-vibrating hardening after intermediate cyclic loading (training) increases the fatigue life of the sample up to 1.2 times; • Double additional shock-vibrating hardening of the samples in combination with double intermediate cyclic loading increases the fatigue life of the samples up to 1.1 times; • Post-deformation recovery of the samples does not significantly affect the increase in fatigue life, which is caused by stress relaxation or loss of strength. As it is seen from results of researches, the fatigue life of the samples subjected to double additional shock-vibrating hardening has decreased in comparison with a fatigue life of the samples subjected to single additional shock-vibrating hardening.

22

V. A. Lebedev et al.

For understanding this phenomenon, graphical analysis of processes proceeding in the surface layer at the strengthening processing and their cyclic loading was carried out. It was based on the theses described below. At the strengthening processing of parts by methods of surface plastic deformation, as it is shown in work [3], qualitatively new surface layer is formed which is characterized by high physico-mechanical specifications without stress raisers and synergistically balanced evenly crushed structure. Real materials initially contain a certain concentration of defects and non-zero initial energy E0 stored while processing of the material. Especially a lot of defects are formed in the surface layers by hardening processing of parts. In accordance with modern concepts, the process of hardening of the surface layer parts by means of surface plastic deformation is a kinetic process taking place in time. In general, the energy state of a deformable element of the surface layer during the process of surface plastic deformation can be represented as: E = E0 + Ey, where Ey is the growth of the stored energy density of various kinds of defects, and damages accumulated in the material due to surface plastic deformation. According to the kinetic representations of the surface plastic deformation process, the surface level is considered to be maximally hardened when the density E of the internal energy inside the deformable elements located on the surface of the part reaches the limit value, Eopt. This state of the surface level for the specific conditions of surface plastic deformation is characterized by the highest strength properties, such as hardness and residual stress, and provides maximum performance parts. This thesis has got its scientific proof in the way of experimental studies on the establishment of effective conditions of vibroshock processing which were represented in Figs. 2 and 3, thus illustrating that the marginal density of the internal energy of the hardened surface layer Eopt of the samples at which the greatest increase in fatigue life is reached is ensured by the duration of processing t = 90 min. Under the process of exploitation, materials get additional damage, leading to extra increment of the internal energy by the value Ee. Therefore, the condition of upper surface layer destruction in the process of operation can be represented as: Ecr = E0 + Ey + Ee. It is particularly worth mentioning that the representation of the process of damage and destruction of the upper surface layer at the stage of surface plastic deformation and operating as a kinetic arises out of two experimentally established facts, the first of which indicates its statistical nature, while the second one reveals its phenomenological side. In the base of the phenomenological nature of the process, there is a dependence on the power parameters of loading and the time (cycles) of their impact, from which it follows that if two processes take place both at higher external force loading parameters as well as at lower ones, then the difference in the behavior of the deformed element of surface layer will depend only on the speed (degree) of damage.

Effectiveness of Application of Additional Strengthening …

23

Fig. 4 Graphical interpretation of the experimental study results of the effect of hardening processing on the fatigue life of samples

Statistically the process is as follows. At each time point (cycle) of the deformation of the surface layer element, a certain degree of its damageability is corresponding to the specified loading conditions. As soon as the damage of the material in the local microvolume reaches its limiting (critical) value, the process of its destruction in the form of microscopic and macroscopic disorders begins. Figure 4 represents graphical illustration of the effect of regulations (1, 2, 3) of hardening processing on the fatigue life of samples, built on the assumption that the processes of elastic-plastic deformation and damage during cyclic loading of samples proceed in a linear relationship. The solid lines 0, 1, 2, 3 show the change in the density of the internal energy of the surface layer of the original and extremely strengthened samples according to the regulations 1, 2, 3 in the process of their cyclic loading until the moment of destruction, recorded by the results of experimental studies. Dotted lines 4, 5, 6 determine the change in the density of the internal energy of the surface layer of the samples in the process of their intermediate cyclic loading. Segments 2.0–2.1, 3.0–3.1, 3.2–3.3 and 0– 3.0, 3.1– 3.2, 3.3–3.4 show the change in the density of the internal energy of the surface layer in the process of vibroshock hardening of the surface plastic deformation in accordance with the regulations 1, 2, 3.

4 Conclusion The analysis of graphical interpretation of the study results allows us to draw the following conclusions: • Additional hardening processing provides an increase in fatigue life, provided that the degree of the surface layer hardening does not exceed the maximum permissible value peculiar to a particular material; • Double additional shock-vibrating hardening of the samples in combination with double intermediate cyclic loading increases the fatigue life in case the

24

V. A. Lebedev et al.

degree of the surface layer hardening does not exceed the critical value for a particular material; • Exceeding of the maximum permissible value of the surface layer hardening degree leads to a decrease in fatigue life due to overhardening of the surface, which was allowed during the hardening of the samples according to the regulation 3; • To ensure an increase in fatigue life, the modes of additional hardening should be coordinated with the duration (number of cycles) of the intermediate cyclic loading, which is also accompanied by the development of elastic-plastic deformation in the surface layer and, consequently, by an increase in its stress state.

References 1. Babichev AP, Babichev EA (1999) Basic concepts of vibration technology. Don State Technical University, Rostov-on-Don, 620 p 2. Boitsov VB (2005) Technological methods to increase strength and service life. Mashinostroenie, Moscow, 128 p 3. Chuchukalov AP (2006) Application of vibration technologies to increase the surface quality and service properties of parts. Don state Technical University, Rostov-on-Don, 215 p 4. Kirichek AV, Solovyev DL, Lazutkin AG (2004) Technology and equipment of staticopulse machining by surface plastic deformation. Mashinostroenie, Moscow, 287 p 5. Kopilov YR (1999) Shock-vibrating hardening. Institute of Russian Ministry of Home Affairs, Voronezh, 386 p 6. Koval NS (2011) Increasing fatigue limit of long complex parts. Problems of vibration technology. In: Proceedings of international seminar application of low-frequency vibrations for technological purposes, Rostov-on-Don 7. Lebedev VA (2006) Technology of dynamic methods of surface plastic deformation. Don State Technical University, Rostov-on-Don, 183 p 8. Motrenko PD, Prokopets GA, Babichev AP (2003) Shock processes of part vibrating hardening, by surface plastic deformation. Voprosi Vibratsionnoy Tekhologii, Rostov-on-Don, pp 9–11 9. Motrenko PD, Pastukhov FA, Maksimov DV, Vovchenko SV (2018) Premises of using additional surface plastic strain hardening to increase fatigue life of parts. In: Knowledge-intensive and vibtowaving technologies of processing Hi-tech parts. Collection of papers. Don State Technical University, pp 97–99 10. Suslov AG (2006) Engineering support and increasing service properties of parts and their assemblies. Mashinostroenie, Moscow, 447 p 11. Lebedev VA (2007) Energetic aspects of parts strengthening by dynamic methods of surface plastic deformation. Don State Technical University, Rostov-on-Don, 156 p 12. Ivanova VS (1975) Nature of metal fatigue. Metallurgy, Moscow, 272 p 13. Fedorov VV (1985) Kinetics of damageability and destruction of solids. Publishing House FAN, Tashkent, 168 p 14. Kirichek AV, Solovyev DL (2001) Kuznechno-Shtampovochnoe Proizvodstvo (Obrabotka Metallov Davleniem). N Engl J Med 7:28–32 15. Kirichek AV, Solovyev DL (2002) Kuznechno-Shtampovochnoe Proizvodstvo (Obrabotka Metallov Davleniem). N Engl J Med 10:35–40

Effectiveness of Application of Additional Strengthening …

25

16. Kirichek AV, Solovyev DL (2004) Kuznechno-shtampovochnoe proizvodstvo (obrabotka metallov davleniem). N Engl J Med 2:13–17 17. Kirichek AV, Solovyev DL (2008) Russ Eng Res 28. J Mol Med 3:277–279. https://doi.org/ 10.1007/s11980-008-3020-6 18. Kirichek AV, Soloviev DL (2013) J Nano Electron Phys 5. J Mol Med 4:04009 (4p) 19. Kirichek AV, Soloviev DL (2013) J Nano Electron Phys 5. J Mol Med 4:04010 (5p) 20. Kirichek AV, Soloviev DL (2014) J Nano Electron Phys 6. J Mol Med 3:03069 (4p) 21. Kirichek AV, Soloviev DL, Lazutkin AG, Silantiev SA (2002) STIN. N Engl J Med 5:13–15

Determination of Rational Technological Parameters of Cold Bending with Rolling of Pipes from Corrosion-Resistant Steels E. V. Khaliulin, A. V. Bobylev and A. V. Kozlov

Abstract The article describes the technology of cold bending of pipes with rolling from corrosion-resistant steels. It is shown that there are certain features arising in the process of cold bending of pipes made of corrosion-resistant steels, associated with their physical and mechanical properties. At the same time, due to a higher plasticity of these steels, an increase in the internal diameter of the pipe occurs during bending with rolling, which leads to a decrease in the actual tension, as a result, corrugations may appear on the inner surface of the pipe. In this regard, the assessment of the values of tension is made from the condition of the absence of corrugated pipes. In addition, the values of tensions were clarified on the basis of a number of restrictions, namely the shape and size of the deforming elements of the distributor, the limiting values of the pipe material elongations, ensuring the accuracy of the pipe shape in longitudinal section, and also taking into account the allowance for the internal diameter of the pipe. Based on the above, the values of rational tension and feed rates for bending pipes with a diameter of 57 mm with a wall thickness of 5 mm from steel grade 12X18H10T were determined. The results of theoretical calculations are verified by field experiments on a special machine. The results of the research will improve the accuracy of curved pipeline elements made of corrosion-resistant steels.










Keywords Bending of pipes Cold bending Bending with rolling Corrosion-resistant steels Deforming elements Tension Special machine










1 Introduction Cold bending of pipes with rolling [1] allows fabricating curved sections of pipelines [2] of high quality while maintaining the cross-sectional profile, eliminating corrugation, and strengthening the internal surface of the pipe. But providing E. V. Khaliulin (&)
A. V. Bobylev
A. V. Kozlov South Ural State University, 76, Lenin Avenue, Chelyabinsk 454080, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_4

27

28

E. V. Khaliulin et al.

high quality is possible only with the correct purpose of the main technological parameters of cold pipe bending with rolling—tension and feed speed of the pipe. The carried out theoretical and experimental studies [3] made it possible to develop designs of special machines covering the range of flexible pipes with a diameter of 20–219 mm, which were mainly intended for bending pipes from low carbon and high-quality carbon steels (steel 20). However, in many industries, gas and oil, chemical and nuclear, aircraft and rocket production, automobile and tractor construction are increasingly used trunk and local pipelines of corrosion-resistant steel. The most commonly used steel is 12X18H10T, since it has high anti-corrosion properties, has a very long service life at high temperatures (up to 600 °C), and allows to obtain durable and hermetic welded joints.

2 Research Methodology The possibility of using cold pipe bending with rolling largely depends on the mechanical properties of the material of the pipe being bent—tensile strength rv, yield strength rt, and relative elongation of the material dl [4, 5]. The yield strength rt of steel grades 20 and 12X18H10T [4–6] varies slightly, but the temporary resistance rv of steel 20–410 MPa, and steel 12X18H10T—529 MPa. The plastic properties of steel 12X18H10T are higher than similar properties of steel 20 from 17 to 37%— relative elongation of the material dl upon reaching temporary resistance rv for steel 20–0.25, for 12X18H10T—0.4. It is known that when rolling a bent pipe due to the introduction of deforming elements into the surface of the pipe and elongation of its cross section, the inner diameter will increase. Since the preload is created due to the difference between the diameter of the roller and the inner diameter of the pipe, the actual preload will decrease. In this regard, there is the problem of maintaining a stable value of the tension—the main technological parameter of flexible pipes with rolling. Due to the high plasticity of steel 12X18H10T, deforming elements (balls) will go deeper into the inner surface of the pipe wall by a large amount, and since its plastic properties are higher than steel 20 from 17 to 37%, the probability of corrugation increases. The condition for the onset of corrugation may be [7]: bending Mmoment ¼

0:605pRh2 ðrB  rT Þ ; eB

ð1Þ

bending —bending moment at which corrugation starts, kN/m; R—radius of where Mmoment the pipe shell, mm; h—wall thickness of the shell of the pipe, mm; ev—relative elongation of the pipe material when it reaches the tensile strength rv, %. According to formula (1), when bending a pipe made of steel 12X19H10T, the calculated value of the bending moment at which the corrugation starts will be less; hence, the loss of stability and the appearance of bumps on the inner surface of the pipe will occur with less effort.

Determination of Rational Technological Parameters …

2.1

29

Evaluation of Stability Criteria

We will evaluate the stability criterion on the example of a pipe with a diameter of 57 mm and a wall thickness of 5 mm. Calculations show that for steel 12X18H10T, the bending moment at which the corrugation occurs is 15% more than that of steel 20. Based on this, we will calculate the rational tension and the corresponding allowable feeds [8]. The method of calculating the tensions set out in [9] allows to determine only their limiting values, in particular the minimum preload Hmin (mm) at which the formation of plastic zone begins and the maximum possible pressure Hmax (mm) at which the destruction of the pipe begins. However, there are a number of other factors that lead to changes in these limit values.

2.2

Calculation of Rational Tensions

(1) Shape and size of deforming elements. The creation of a plastic zone is accompanied by the pressing of the deforming elements into the surface of the pipe wall [10, 11]. The greater the depth of indentation, the wider the plastic zone. However, the size of this zone is limited. This is due to the laws of the propagation of stresses within the metal and the diameter of the deforming element (ball). It is obvious that the width of the plastic zone cannot exceed the size of the deforming element in the direction of the axial flow of the tool bmax = Dsphere. The distance between two deforming elements (balls) is equal to S/n (mm/pc.). Therefore, the maximum possible tension Hmax is limited by the diameter of the deforming element (ball) Dsphere (see Fig. 1). (2) Ultimate elongation of the pipe material [12–14]. When rolling, there is some elongation of the pipe walls in cross section. In the limiting case, with an increase in tension, the cross section will take the form (for a 3-ball roller), see Fig. 2. Consider the cross section of the pipe brought up to the state of ultimate elongation. The value of the maximum possible tension to limit the associated plasticity of the pipe material Hdl is H

l

2 ОК

internal Dpipe

2 (ОР1

internal КР1 ) Dpipe

2ОР1

Dsphere

internal Dpipe

ð2Þ

30

E. V. Khaliulin et al.

Fig. 1 Scheme for determining the thickness of the crumpled layer in the radial direction

Fig. 2 Diagram of the limiting deformations of the cross section of the pipe

The circumference of the outer surface of the pipe louter is louter ¼ p
Dexternal pipe

ð3Þ

In the extremely stretched state, the length lextreme is more by the amount of relative elongation of the material dl (%)

Determination of Rational Technological Parameters …

31

lextreme ¼ louter þ dl
louter =100 ¼ ð1 þ dl =100Þ
louter ¼ ð1 þ dl =100Þ
p
Dexternal pipe ð4Þ According to the diagram in Fig. 2, this value is

 lextreme ¼ 3
PPl þ p Dsphere þ 2
h ¼ 6
AP þ p Dsphere þ 2
h

ð5Þ

For an arbitrary number of balls in the roller

 lextreme ¼ n
PP1 þ p Dsphere þ 2
h ¼ 2n
AP þ p Dsphere þ 2
h

ð6Þ

From Eqs. 4 and 6 AP ¼


 ð1 þ dl =100Þ
pDexternal  p Dsphere þ 2h pipe 2n

ð7Þ

From the triangle OAP
 ð1 þ dl =100Þ
pDexternal  p Dsphere þ 2h AP pipe OP ¼ ¼ sin pn 2n
sin pn

ð8Þ

Substituting (8) into Eq. (2) and considering that OP = OP1 we obtain Hdl ¼


 pDexternal ð1 þ dl =100Þ  p Dsphere þ 2h pipe n
sin pn

þ Dsphere  Dinternal pipe

ð9Þ

or for the 3-ball roller Hdl ¼

ð1 þ dl Þ
p
Dexternal  p
Dsphere Dsphere pipe  Dinternal þ pipe p 3
cos 2n 2

ð10Þ

—outer diameter of the pipe, mm; Dinternal where Dexternal pipe pipe —internal diameter of the pipe, mm; dl—relative elongation of the pipe material, %; Dsphere—diameter of the deforming element (ball), mm; n—number of deforming elements (balls) in the roller, pcs. With the increase in the number of deforming elements in the rolling mill, a smaller amount of tension will be required to achieve the ultimate deformation of the material when it is rolled out (Fig. 3).

32

E. V. Khaliulin et al.

Fig. 3 Dependence of the tension on the number of deforming elements in the rolling mill

(3) Preservation of the shape of the longitudinal section. As was shown above, with significant preloads, an increase in the outer diameter of the pipe may occur, which will lead to a deviation of the product shape in longitudinal section [15]. Therefore, the maximum possible tension Hmax should be limited based on the conditions of relative elongation of the material dl and it will dl be Hmax : dl Hmax ¼

2
dl
h ; 100

ð11Þ

where h—pipe wall thickness, mm. (4) Restriction on the allowance of the size of the internal diameter of the pipe. In the production of seamless pipes, due to uneven heating of the workpiece, incorrect machine settings, insufficient compression in front of the toe of the mandrel when piercing, as a rule, thin-walling is formed [16–18]. The limit deviations in thickness are governed by GOST 8734–75, so it is necessary to increase the size of the minimum tension Hmin by the value of the maximum deviation of the h : varying wall thickness h, and then, it will be Hmin h Hmin ¼ Hmin þ 2
0; 1
h ¼ Hmin þ 0; 2
h

ð12Þ

3 Research Results Taking into account all the restrictions, we obtain the refined maximum permissible values of (the maximum possible Hmax and minimal Hmin) tensions.

Determination of Rational Technological Parameters …

3.1

33

Determination of Allowable Feeds

Knowing the range of extreme tensions, it is possible to determine the permissible feed. Suppose that n balls are installed in the rolling mill. Consider the zone of deformations of two adjacent deforming elements. As was shown in [9, 19] if the thickness of the crumpled layer reaches Hmin/2 (see Fig. 1), then at this point the stresses reach the yield strength and the material of the pipe wall becomes so plastic that it has almost no resistance to bending. This allows to get a real scheme of stress distribution in the pipe wall from the impact of a deforming element on it. Let the coordinate axes X and Y pass through point 0; then, the coordinates of the beginning of the plasticity zone of the first ball will be xH1, and the end of this zone will have the coordinate xK1. In the system of equations describing the position of the lines of contact of the first ball—denoted by the index 01, and the position of the line of contact of the second ball by the index 02 (see Fig. 4): 8   Dsphere 2 > 2 2 > > þ y ¼ x < 1 1 2   > Dsphere 2 > > : x22 þ y22 ¼ 2

Fig. 4 Overlap of plastic zones of the first and third deforming elements

ð13Þ

34

E. V. Khaliulin et al.

In the required coordinate x1 = xH1, the following conditions are observed 8   S > > x ¼ x  < 2 1 n > H > : y2 ¼ y1 þ min 2

ð14Þ

 h  nHmin n Hmin S2
þ þ 2 2S 2S 4 n

ð15Þ

hence x1 ¼ y1
with (

a ¼ nH2Smin

n b ¼ 2S


h Hmin 4

þ

S2 n2



ð16Þ

so x1 ¼ ay1 þ b

ð17Þ

Substituting the acquired value x0 into (13), we have ðay1 þ bÞ2 þ y21 ¼

  Dsphere 2 2

ð18Þ

or y21




  ! Dsphere 2 a þ 1 þ 2ab
y1 þ b  ¼0 2 2



2

ð19Þ

With 8 2 0. From the scheme (Fig. 4) for a three-ball rolling mill, it follows that when selecting a feed, the plastic zones of the first and third deforming elements (balls) should be covered.

36

3.2

E. V. Khaliulin et al.

Determination of the Total Plastic Zone Width

The beginning of the total plastic zone corresponds to the coordinate xHn—coordinate of the beginning of the zone of plasticity of the last ball and the end of the total plasticity zone will correspond to the coordinate xK1 and then the width of the total plastic zone will be bcommon. By changing the values of tension within the limits of the pipe thin-wall thickness and the accuracy of the shape of the manufactured products, it is possible to determine the allowable feed rates and choose the nearest smaller one on a special machine. The tension range calculated for the Ø57  5 mm pipe made of steel 12X18H10T is thus in the range from 1.2 to 1.7 mm, which corresponds to the feed range from 0.7 to 3.2 mm/rev. Since the ball receives a planetary motion, its rotational speed relative to the center of the distributor is less (see Fig. 5) when the rolling mill rotates. The actual spinner speed of the separator of the rolling mill nseparator is equal to  nseparator ¼ ncone


Dsphere 2

 Dinternal  Dsphere pipe Hmax  Hmax 
ð 1 þ cos l Þ 
2 2 2 2 Dinternal pipe 2




Dsphere 2


cos l

;

ð28Þ

where ncone—rotational speed of boring bar of the machine, rpm; l—separator cone angle, °. So for the pipe Ø57  5 mm, according to the accuracy of the shape, the maximum possible tension is Hmax = 1.5 mm. Then, with the diameter of the ball Dsphere = 14.1 mm of the support cones with the angle l = 45° with the frequency of rotation of the machine boring bar ncone = 500 rpm the actual spinner speed of the separator of the rolling mill is nseparator = 410 rpm. This should be taken into

Fig. 5 Planetary rotation of the rolling tool

Determination of Rational Technological Parameters …

37

account when assigning the speed of rotation of the bending roller for the machine of a particular model. Based on the above, the tension and the feed for bending pipes Ø57  5 mm from steel 12X18H10T with a bend radius 2Dpipe were determined using a rolling mill with 3 deforming elements with a frequency of rotation of the bending roller in the range from 0.7 to 1.4 rpm with a deforming element—a ball Ø14.1 mm. The preload is 1.5 mm, and the feed is 0.7 to 3.2 mm/rev.

4 Control Experiments During the control experiments on a special machine, 65 pipes made of steel 12X18H10T with a diameter of 57 mm with a wall thickness of 5 mm were bent. Adjusted preload—1.5 mm. The rotational speed of the bending roller is 1.0 rpm. Bending was carried out at an angle of 180º (Fig. 6). Corrugation was absent on all samples. All samples were subjected to control measurements in three sections of five points N to assess the value of ovality (Fig. 7). The average value of Q ovality was 4.88% [3, 9, 20]. In addition to the above, it should be noted that during the tests, it became clear that when bending pipes made of steel 12X18H10T, only 2–3 bends are enough for one set of rolling mill. For comparison, when bending pipes made of steel 20, one set of rolling mill lasts from 20 to 25 similar bends. A significant role here is played by the heating of the tool (rolling mill). Steel 12X18H10T has three times lower thermal conductivity [thermal conductivity coefficient for steel is 20–51 W/(m °C), and for steel 12X18H10T—16 W/(m °C)]. The temperature in the unrolling zone in this case can reach from 650 to 700 °C, which, in the absence of cooling, causes tempering of the deforming elements—balls made of steel SHKh15. Usually, when bending with the rolling of pipes made of carbon steels, water with a flow rate of

Fig. 6 Curved pipes Ø57  5 mm made of steel 12X18H10T

38

E. V. Khaliulin et al.

Fig. 7 Results of control measurements: 1—bend with minimal ovality; 2—average ovality of the party; 3—bend with maximum ovality

2.5 to 5 L per minute is used as a coolant. When bending pipes made of steel 12X18H10T, this was not enough, so the water was replaced by an oil-cooled coolant based on VOLTES MG (GOST 17216-2001) and the flow rate was increased to 10 L per minute. Tool durability increased to four to five bends per one rolling mill [21, 22].

5 Conclusion Thus, the values of rational tensions and feeds were determined for bending pipes made of corrosion-resistant steels. It is shown that in this case a more precise adjustment of the machine in the much narrower ranges of the technological parameters of tension and feed is required, and it requires equipment with smooth control of rotation frequencies of the rolling mill and bending roller.

References 1. Khaliulin EV, Kozlov AV, Gerasimov MV (2015) Installation for bending small-sized pipes with rolling. Patent of the Russian Federation 157963, Bul. 35 2. Chinyaev IR, Fominykh AV, Ilinykh EA (2016) The valve is a shutoff for the passive protection systems of pipelines. In: International conference on industrial engineering, ICIE 2016. Pub office. “Procedia Engineering”, Chelyabinsk, p 220–224 3. Kozlov AV, Bobylev AV (2007) Technology and equipment for cold bending thin-walled pipes: monograph. SUSU Publishing House, Chelyabinsk, p 169 4. Dreval AE et al (2005) Brief reference of a metalist, 4th edn. Mashinostroenie, Moscow, p 959 5. Zhestkova IN (ed) (2001) Reference for a designer-mechanical engineer, 3 vol, 8th edn (revised and added, vol 1). Mashinostroenie, Moscow, 920 p 6. Gontar IN, Volchikhina NI (2008) Resistance materials: study guide. Publishing house Penza State University, Penza, p 263

Determination of Rational Technological Parameters …

39

7. Birger IA, Shorr BF, Iosilevich GB (1993) Calculation of the strength of machine parts. Mashinostroenie, Moscow, p 640 8. Sherkunov V, Korsakov A (2015) The usage of 156 mm diameter continuous cast billets on “140” pipe-rolling plant for pipe production. In: International conference on industrial engineering, ICIE 2015. Pub office. “Procedia Engineering”, Chelyabinsk, pp 886–890 9. Kozlov AV (1999) Bending pipes, rolled out with large preloads: a tutorial. Publishing house SUSU, Chelyabinsk, p 129 10. Kozlov AV, Sherkunov VG, YaM Kyilcevich (2009) Stress state in pipe on flexure with internal shairing. Russ Eng Res 29(8):809–812 11. Kozlov AV, Khilkevich Y (2003) The technology and equipment for cold bending of pipes. The 29th international conference on mechanical engineering: book of conference lecturer. Technion, Haifa, Israel, pp 190–192 12. Kolmogorov GL, Kuznetsova EV, Khabarova D (2018) Relaxation of residual stresses and precision of steel pipes. Vestnik Magnitogorskogo Gosudarstvennogo Tekhnicheskogo Universiteta im. G.I. Nosova (Vestnik of Nosov Magnitogorsk State Technical University) 16 (3):103–108. https://doi.org/10.18503/1995-2732-2018-16-3-103-108 13. Poletskov PP, Gushchina MS, Alekseev DY, Emaleeva DG, Kuznetsova AS, Nikitenko OA (2018) Understanding the effect of controlled rolling regimes for pipe steel on the structural condition of hot-deformed austenite. Vestnik Magnitogorskogo Gosudarstvennogo Tekhnicheskogo Universiteta im. G.I. Nosova (Vestnik of Nosov Magnitogorsk State Technical University) 16(3):67–77. https://doi.org/10.18503/1995-2732-2018-16-3-67-77 14. Kurmachev YF (2016) Gauge idle zone profile in cold rolling pipe mills. Vestnik Magnitogorskogo Gosudarstvennogo Tekhnicheskogo Universiteta im. G.I. Nosova (Vestnik of Nosov Magnitogorsk State Technical University) 14(4):61–65. https://doi.org/ 10.18503/1995-2732-2016-14-4-61-65 15. Plotnikov LV, Zhilkin BP, Brodov YM (2016) The influence of cross-profiling of inlet and exhaust pipes on the gas exchange processes in piston engines. In: International conference on industrial engineering, ICIE 2016. Pub office. “Procedia Engineering”, Chelyabinsk, pp 111–116 16. Lopatin BA, Khaziev TR (2015) Improved design of machine for cold cutting of oil pipes and gas pipelines. In: International conference on industrial engineering, ICIE 2015. Pub office. “Procedia Engineering”, Chelyabinsk, pp 563–570 17. Minatsevich SP, Sharonov AA, Borisov SS (2015) The design of safety control systems for unattended points of technological communication on oil and gas pipelines. In: International conference on industrial engineering, ICIE 2015. Pub office. “Procedia Engineering”, Chelyabinsk, pp 266–273 18. Moskvicheva EV, Sidyakin PA, Shitov DV (2016) Method of corrosion prevention in steel pressure pipelines in sewerage systems. In: International conference on industrial engineering, ICIE 2016. Pub office. “Procedia Engineering”, Chelyabinsk, pp 2381–2386 19. Korolev AV, Korolev AA (2016) Microstrip mechanism in raceways and rolling elements of roller bearings. In: International conference on industrial engineering, ICIE 2016. Pub office. “Procedia Engineering”, Chelyabinsk, pp 497–501 20. Zakhezina AM, Pryadko YG (2016) Vibration diagnostics of gas pipelines technological equipment using wavelet analysis: In: International conference on industrial engineering, ICIE 2016. Pub office. “Procedia Engineering”, Chelyabinsk, pp 300–306 21. Kulagin VA, Sokolov NY (2017) Improving the thermal characteristics of heat pipes. J Siberian Fed Univ. SFU Publishing House, Krasnoyarsk 10(3):372–376. https://doi.org/10. 17516/1999-494X-2017-10-3-372-376 22. Lipovka YL, Belilovets VI (2016) Some theoretical aspects of calculating the temperature deformations of underground channel-free heat pipelines. J Siberian Fed Univ. Publishing House of Siberian Federal University, Krasnoyarsk 9(4):546–562. https://doi.org/10.17516/ 1999-494X-2016-9-4-546-562

Recomposition Procedure of Automatic Replacement Laser Modules for CNC Machines P. A. Ogin and D. G. Levashkin

Abstract The article is devoted to the development of the mechanism of systematization of possible variants of assembling automatically replaceable modules for varieties realization of technologies of laser processing of materials on CNC machines. It is proposed to consider the laser module as a set of assembled separate blocks, the parameters of each of which are determined for the purpose of the laser technology being realized. The proposed systematization mechanism makes it possible to alternate blocks of a laser module design to customer requirements with minimal time losses for assembly and design work. At the same time, the element base of the blocks remains unchanged, which leaves the possibility to create different design versions of devices based on the design of the CNC machine tool and the complexity of the geometry of the part. The paper shows the application of a technical solution, assembled on the basis of the proposed systematization mechanism, as a separate device for the MILLSTAR MV660 machine. The proposed device variant is selected from the set of blocks outside the labeling working cycles in the tool magazine of the machine tool. In the operating mode, the module is automatically installed in the machine spindle by the command of the CNC system. Using the proposed methodology, it is shown that the design of the devices can be worked out both at the stage of analysis and modernization of existing laser processing modules and at the stage of their development in the future. Keywords Energy-efficient technologies linkage Modular principle





Laser
Plug-in module
Kinematic

P. A. Ogin
D. G. Levashkin (&) Togliatti State University, 14, Belorusskaya St, Togliatti 445020, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_5

41

42

P. A. Ogin and D. G. Levashkin

1 Introduction The current stage of development of laser technology is due to the expansion of element base of Photonics. Currently, transmission of laser radiation through a flexible fiber optic cable over long distances is possible. The basis of this concept is implemented on laser technological complexes (LTK) modular design. This allows you to combine several technological functions in one equipment, in particular, mechanical and laser processing. There was a so-called hybrid complex LTK. One of the first such equipment on the market in the company’s DMG MORI (Germany) in the form of a family of machining centers LASERTEC with the use of modern laser in the working area of the machine is mounted on a separate rotary module. However, this does not reduce the cost of laser processing, due to the engineering costs and subsequent equipment upgrades, the need for the improvement of its management system, which is a limitation in the further development of this direction. In addition, this technical solution in the form of individual items of equipment has a number of technical limitations associated with the size of the working area of the table and the rigidity of the machine capacity installed on the machine drive systems. All of this combined with the high cost of the equipment imposes significant restrictions when implementing laser technology for a wide range of parts. Then, according to (1), the implementation of open kinematic relations is provided by the installation U1(Nm) element. Closed kinematic connection is realized via a connecting element U2(Nm), which is associated with an optical system U3(Nm). Also in the optical system U3(Nm) mounted beam delivery unit U4(Nm) and the beam focusing device U5(Nm). Open kinematic linkage ensures, for example, the possibility of axial rotation of the device (one degree of freedom). Therefore, the next stage of development of the technological sector of the market of Photonics is the development and creation of complex technical devices based on the automatically replaceable modules for the implementation of energy-efficient laser processing techniques applicable to any modern CNC milling and boring group [1–17]. According to a modular scheme (Fig. 1) constructively module provides implementation, both open and closed kinematic linkages are through kinematic units—blocks [18]. Closed kinematic constraints are implemented without the possibility of moving the blocks U3(Nm)–U5(Nm) of the module, and it is in the composed state of the device. The modular scheme of the module for block U2(Nm) provides for the possibility of axial rotation, linear movement or hinge performance. Let us define the degree of freedom of the movable elements W device based on movements in the module’s kinematic chain: W ¼ 6n  L where L ¼ 5P5 þ 4P4 þ 3P3 þ 2P2 þ P1 .

ð1Þ

Recomposition Procedure of Automatic Replacement Laser …

43

Fig. 1 Schematic of modular performance of N-types laser modules

n—number of links in the chain; L—number of constraints imposed on the unit blocks; P5, P4, P3, P2, P1—kinematic pairs corresponding number of class (5, 4, 3, 2 and 1) included in the kinematic chain module. W ¼ 6n  U1 ðNm Þ
P5 þ U2 ðNm Þ
P4 þ U3 ðNm Þ
P3 þ U4 ðNm Þ
P2 þ U5 ðNm Þ
P1 ð2Þ Given that in the composed state of the module provides n = 0° of mobility, the expression (2) takes the form: W ¼ U1 ðNm Þ
P5 þ U2 ðNm Þ
P4 þ U3 ðNm Þ
P3 þ U4 ðNm Þ
P2 þ U5 ðNm Þ
P1 ð3Þ This is expressed in a general form to determine kinematic characteristics of the laser module according to the applicable set of blocks needed to address the problems of laser processing. In this case, we assume that the module is designed as a predetermined (from the position and the type of problem being solved laser processing techniques) kinematic chain composed of a given set of blocks. In the case of the use of rotary, telescopic or rotating block device obtains common to all units of a preset (fixed) the degree of mobility of m = 1. According to such devices include mechanisms to the first order, then (3) takes the form: W ¼ 5n  U1 ðNm Þ
P4 þ U2 ðNm Þ
P3 þ U3 ðNm Þ
P2 þ U4 ðNm Þ
P1 þ U1 ðNm Þ ð4Þ

44

P. A. Ogin and D. G. Levashkin

The expression in general terms to determine the kinematic characteristics of the module in the composed state, depending on the set of mobile or stationary units needs to solve the current laser processing tasks. Also, using the expression (3 and 4) can solve the inverse problem, where on the basis of the kinematic analysis is determined by the degree of mobility of the equation W composable laser module designed for the task, then define a set of blocks needed for the composition of module. When expanding the functional module capabilities needed in expressions (3) and (4), further consider the appearance in the structural formula of the module according to two parameters Q and F, where Q—the number of additional kinematic linkages apparatus, F—the number of excess (more) degrees of mobility. The expansion of the functionality associated with the modules recomposing, described by the expressions (5, 6): W ¼ U1 ðNm Þ
P5 þ U2 ðNm Þ
P4 þ U3 ðNm Þ
P3 þ U4 ðNm Þ
P2 þ U5 ðNm Þ
P1 þ Q þ F W ¼ 5n  U1 ðNm Þ
P4 þ U2 ðNm Þ
P3 þ U3 ðNm Þ
P2 þ U4 ðNm Þ
P1 þ U1 ðNm Þ þ Q þ F

ð5Þ ð6Þ

To determine the parameters F and Q in (5), (6), consider a few examples of laser modules recomposing (see Fig. 2). Given a U1i module design (see Fig. 1), its structural formula is given by (2) and (3): WU1 ¼ U1 ðNm Þ
P5 þ U2 ðNm Þ
P4 þ U3 ðNm Þ
P3 þ U4 ðNm Þ
P2 þ U5 ðNm Þ
P1 ð7Þ From (7), it follows that in the initial state value of Q = 0 and F = 0 (Fig. 2).

Fig. 2 Option (A) of the laser module

Recomposition Procedure of Automatic Replacement Laser …

45

In the process of a laser module recomposing and extend its functionality assumed extra or replacement of existing units for different tasks. Extra blocks lead to a number of additional kinematic constraints module and a number of excess (more) degrees of mobility of its design. For Option (A) of laser module U1i, its structural formula is WU1A ¼ U1 ðNm Þ
P5 þ U2 ðNm Þ
P4 þ U3 ðNm Þ
P3 þ U4 ðNm Þ
P2 þ U5 ðNm Þ
P1 þ 0 þ 0 ð8Þ Here, it is assumed that as a result of recomposing of a laser module of its structural formula of the device has not changed. For Option (B) U1i module (Fig. 3), its structural formula is (9) WU1B ¼ U1 ðNm Þ
P5 þ U2 ðNm Þ
P4 þ U3 ðNm Þ
P3 þ U4 ðNm Þ
P2 þ U5 ðNm Þ
P1 þ 1 þ 1 ð9Þ In the structural formula (9), it is assumed that the number of additional communications module kinematic Q = 1 and the number of excess (more) degrees of mobility by installing additional rotation unit for an optical system F = 1. For Option (C) of the laser module U1i, (Fig. 4) its structural formula is (10) WU1C ¼ U1 ðNm Þ
P5 þ U2 ðNm Þ
P4 þ U3 ðNm Þ
P3 þ U4 ðNm Þ
P2 þ U5 ðNm Þ
P1 þ 1 þ 2 ð10Þ In addressing other sought after, as well as advanced laser processing tasks in the manufacturing process of the variable product nomenclature, other performances are possible for the laser module.

Fig. 3 Option (B) of the laser module

46

P. A. Ogin and D. G. Levashkin

Fig. 4 Option (C) of the laser module

2 Testing Solutions In general terms, the structural analysis of the laser module design and the elaboration of this composition scheme are proposed to use the following sequence: 1. Statement of the problem on the organization of the process of laser processing and analysis of the official appointment of the developed device. At this stage, set technical limitations, and working on the issues related to the kind of laser treatment, the desired mode and processing conditions. According to the expressions (2) and (3), it performed a preliminary analysis of the laser module of the kinematic scheme. 2. Analysis of the necessary kinematic relationships needed to implement laser processing technology and performance of the service module destination. Running update laser module of the kinematic scheme. 3. Determination of the structural formula of the laser module. Analyzing module kinematic scheme and determined the final version of the structural formula according to (4). 4. The construction of the block diagram (model) of the laser module. Using the modular schemes (see Figs. 2 and 3) is determined by the amount, type and form of execution of the blocks that make up the module. 5. Analysis of the mobility of structural elements of the laser module and the identification of redundant kinematic constraints. Using expressions (5) and (6) are defined and Q values of parameters F. 6. Write the formula of building a laser module based on open and closed kinematic linkages, using the expression (8, 9 and 10). 7. Choosing a set of blocks needed for a meaningful performance laser module. Search of the final module composition.

Recomposition Procedure of Automatic Replacement Laser …

47

Fig. 5 Machine MILLSTAR MV660 with the installed spindle automatically replaceable module laser marking. 1—spindle machine assembly; 2—tool magazine; 3—laser processing module; 4— working machine table CNC

With regard to steps 4–7, it is possible to pre-assemble the module in a prepared environment of computer simulation. Figure 5 shows the implemented solution composited developed a module for the laser marking of parts mounted on the machine table MILLSTAR MV660. The module performed according to expressions (2) and (7) of this work. The choice of the laser radiation source is used in solving concrete technological problems, determined by production needs and economic possibilities of the technology consumer. Turning on and off the laser source is carried out on CNC machines to comply with the management team, program. On MILLSTAR MV660 machine were worked out laser marking cycles, together with the cycles of the machining housing part. The sequence of the combined technologies comprising preparing control programs for CNC machining (cutting process) and laser processing, setting the workpiece on the machine table, placement tool blocks and module laser treatment in the cells of the tool of the machine shop and the refinement of mechanical cycles (blade) and laser treatment are according to the control NC program.

3 The Main Conclusions and Results 1. In paper proposed recomposition methodology of modules for laser processing, which integrated into existing systems machining based on the machining centers with the CNC, the fundamentals of the principle of modular recomposition are based on the analysis of kinematic links. 2. The paper proposed and justified the construction of the laser module, intended to the solution of specific technological problems of laser processing of materials.

48

P. A. Ogin and D. G. Levashkin

3. It is shown and proved the possibility of designing various functional purpose and complexity of the device allowing for the automated equipment, the type, layout execution, the working area size, direction of the main and auxiliary motions of knots, positioning and control systems. 4. The use of the block-modular principle of the laser module construction can significantly reduce design time and creating new technical solutions and design module for the implementation of variant technological laser processing tasks. 5. The effectiveness of the proposed in the composition principles of automatically plug-in modules laser processing is confirmed by the results of the processing of the workpiece MILLSTAR MV660 with the use of technical solutions linked module for laser marking.

References 1. Braier Z, Šidlof P, Čejka V, Žďárek P (2017) New method of contactless measurement and analysis of CNC machine spindle lopping. https://doi.org/10.1007/978-3-319-44087-3_49 2. Porta JM, Ros L, Bohigas O, Manubens M, Rosales C, Jaillet L (2014) The CUIK suite: analyzing the motion closed-chain multibody systems. IEEE Robot Autom Mag 21(3):105– 114. https://doi.org/10.1109/MRA.2013.2287462 3. Aggarwal L, Urbanic R, Aggarwal K (2014) A reconfigurable algorithm for identifying and validating functional workspace of industrial manipulators. SAE Technical Papers, 1. https:// doi.org/10.4271/2014-01-0734 4. Guo ET, Wu T, Zhang LB, Huang FL (2014) Study on the path optimized method based on an improved clustering ant colony algorithm for CNC laser drilling. www.scientific.net/ AMM.556-562.4439 5. Velotti C, Astarita A, Leone C, Genna S, Minutolo FMC, Squillace A (2016) Laser marking of titanium coating for aerospace applications. Paper presented at the Procedia CIRP 41:975– 980. https://doi.org/10.1016/j.procir.2016.01.006 6. Lei Z, Bi J, Li P, Guo T, Zhao Y, Zhang D (2018) Analysis on welding characteristics of ultrasonic assisted laser welding of AZ31B magnesium alloy. Opt Laser Technol 105:15–22. https://doi.org/10.1016/j.optlastec.2018.02.050 7. Jia Q, Guo W, Wan Z, Peng Y, Zou G, Tian Z, Zhou YN (2018) Microstructure and mechanical properties of laser welded dissimilar joints between QP and boron alloyed martensitic steels. J Mater Process Technol 259:58–67. https://doi.org/10.1016/j.jmatprotec. 2018.04.020 8. Wu J, Wei H, Yuan F, Zhao P, Zhang Y (2018) Effect of beam profile on heat and mass transfer in filler powder laser welding. J Mater Process Technol 258:47–57. https://doi.org/10. 1016/j.jmatprotec.2018.03.011 9. Wang L, Wei Y, Chen J, Zhao W (2018) Macro-micro modeling and simulation on columnar grains growth in the laser welding pool of aluminum alloy. Int J Heat Mass Transf 123:826– 838. https://doi.org/10.1016/j.ijheatmasstransfer.2018.03.037 10. Li S, Deng Z, Deng H, Xu W (2017) Microstructure and properties of weld joint during 10 kW laser welding with surface-active element sulfur. Appl Surf Sci 426:704–713. https:// doi.org/10.1016/j.apsusc.2017.07.262 11. Ye Z-P, Zhang Z-J, Jin X, Xiao M-Z, Su J-Z (2017) Study of hybrid additive manufacturing based on pulse laser wire depositing and milling. Int J Adv Manuf Technol 88(5–8):2237– 2248. https://doi.org/10.1007/s00170-016-8894-8

Recomposition Procedure of Automatic Replacement Laser …

49

12. Brecher C, Bleck W, Feldhusen J et al (2017) Multi-technology platforms (MTPs). Integrative production technology: theory and applications, pp 369–513. https://doi.org/10.1007/978-3319-47452-6_6 13. González J, Rodríguez I, Prado-Cerqueira J-L, Diéguez JL, Pereira A (2017) Additive manufacturing with GMAW welding and CMT technology. Procedia Manufact 13:84–847. https://doi.org/10.1016/j.promfg.2017.09.189 14. Ambrizal NHB, Farooqi A, Alsultan OI, Yusoff NB (2017) Design and development of CNC robotic machine integrate-able with Nd-Yag laser device. Procedia Eng 184:145–155. https:// doi.org/10.1016/j.proeng.2017.04.079 15. Yang Y, Cao L, Wang C, Zhou Q, Jiang P (2018) Multi-objective process parameters optimization of hot-wire laser welding using ensemble of metamodels and NSGA-II. Robot Comput-Integr Manufact 53:141–152. https://doi.org/10.1016/j.rcim.2018.03.007 16. Mueller R, Vette M, Ginschel A, Mailahn O (2014) Innovative production technologies for large components. SAE Technical Papers, 2014-September. https://doi.org/10.4271/2014-012237 17. Appendino D (2011) La saldatura laser robotizzata per la fabbricazione di componenti in acciaio inossidabile (Robotic laser welding for the manufacture of components in stainless steel). Rivista Italiana Della Saldatura 63(4):487–501 18. Ogin PA, Levashkin DG (2016) Solutions based on automatic changeable modules used for implementation of the laser technology on machining CNC. Bulletin of South Ural State University. Series “Mechanical Engineering Industry” 16 (3):29–35. https://doi.org/10.14529/ engin160304

Tooling Backup of Cutting and Deforming Processing of Non-rigid Shafts E. I. Yatsun, N. P. Anikeyeva and I. S. Karnaukhov

Abstract The article presents the conditions for ensuring the quality parameters of the machining of surfaces of hydraulic cylinder rod during machining and using the methods of surface plastic deformation. Modern machines and equipment should have a good performance in various conditions of their use. Improvement of performance characteristics can be achieved both through the constructive improvement of machines, the use of new structural materials and the use of advanced technology for manufacturing parts and through the development of new tool designs and improvement of existing ones. The article analyzes the static and dynamic characteristics of the tool system for combined cutting and deforming processing, which affect the accuracy parameters of the machining quality. The processes of combined machining and surface plastic deformation of non-rigid parts of the shaft class are considered, and the results of the study of technological parameters of processing for cutting and deforming rod treatment are presented which are obtained during production tests. Traditionally, non-rigid parts are machined by turning, followed by grinding and polishing. However, turning and abrasive machining often provide the required quality parameters of these parts (JT7 … 8, Ra < 0.16 .., 0.32 micron) as unstable and have low productivity, and as a result these processes are not used for manufacturing these parts. The use of tools and tool systems for the combined cutting and deforming machining can improve the accuracy of the shape and size of the processed non-rigid part, to ensure a small amount of surface roughness and to improve the physical and mechanical properties of the surface layer of the part.




Keywords Non-rigid part Tooling system Anti-vibration holder Dampening cutter





Roughness
Vibration


E. I. Yatsun (&)
N. P. Anikeyeva
I. S. Karnaukhov Southwest State University, 94, St. 50 Let Oktyabrya, Kursk 305004, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_6

51

52

E. I. Yatsun et al.

1 Introduction Purpose of work—increased productivity while ensuring the required quality parameters of machining of non-rigid parts. Subject of research—the processes of combined machining and surface plastic deformation of non-rigid parts, as well as the study of technological parameters of processing in the tool system for cutting and deforming processing of the said parts and the influence of structural elements of the tool system on the machining error in the system. Object of research—tool system for combined cutting and deforming treatment of non-rigid parts. Urgency—practical value it consists in the development of the design of the tool system for cutting and deforming processing of non-rigid parts and recommendations for the selection of technological modes of cutting and deforming processing of non-rigid parts necessary to obtain the required parameters of the product.

2 Problems of Achieving Quality in the Processing of Long Shafts Rod is a part of the hydraulic cylinder (Fig. 1). The development of new high-performance and improvement of existing designs of tools and tool systems in the manufacture of non-rigid parts with the possibility of using them for various types of production is a task of great scientific and practical importance. The solution of the problem of stabilizing the machining process during cutting will ensure the specified resistance of the cutting too [1–18]. Rod refers to the type of long non-rigid shafts, the condition l/d = 10 … 20 is satisfied: diameter 60d9; length 4050 mm; surface roughness Ramax 0.32 microns; hard chrome coating thickness hmin 20 microns. Part material—Steel 45 GOST 23270-89. Hardness is HB269 in a condition of delivery. For a tight connection with the working cylinder by means of bushings with seals it is necessary the high quality of the guide surface of the rod. The main task in

Fig. 1 Rod design

Tooling Backup of Cutting and Deforming Processing …

53

the finishing treatment of hydraulic cylinder rods is to achieve the stability in obtaining the required dimensions and surface roughness. We analyze the processing of part “Rod” (Fig. 1). Rough Stock—Pipe 73X15-35 GOST 23270-89. Since the outer surface of the rod works in an aggressive environment and must withstand heavy loads, then running, polishing and chroming the outer surface are used to harden the surface layer of the part, to increase its wear resistance and to provide the required surface roughness. Problems with processing part “Rod”: (1) using a fixed steady to reduce the spinning of the rough stock during the turning leads to stop the machine and rearrange the steady; (2) pressing leads to a barrel shape in the middle of the shaft and causes the tool to vibrate; (3) using a soldered tool does not give a stable receipt of the roughness parameter in the finishing turning operation; (4) after the operation of running, the fiber delamination of the surface is observed, which has to be removed with a sandpaper.

3 Tasks that Need to Be Solved to Achieve Quality in the Processing of Long Shafts To improve the processing efficiency of the guide surfaces of the rod it is necessary to solve the following tasks: 1. To ensure the specified roughness parameter Ramax 2.5 microns for the finishing turning operation, it is necessary to choose a new design of the turning tool. 2. To reduce vibrations that impair the quality of processing, it is recommended to use anti-vibration cutting tool holders. 3. It is proposed to turn the turning at the turning center PЦM8000 with a through-slide and a moving steady (Fig. 2).

Fig. 2 The scheme of the operation

54

E. I. Yatsun et al.

Under the action of cutting forces and its own weight, the stock bends, vibrates. Consequence: the destruction of the cutting plate of the cutter, the deterioration of the quality of processing, reduced performance. To avoid these shortcomings, the treatment with a moving rest is applied [19–24]. In the process of processing the movable lunette moves with the cutter, which allows you to place the Cams of the lunette in close proximity to the cutting zone and thus almost eliminate the deflection of the workpiece. Chips under this condition moves to the left, which is required. The lunette is exposed on the processed surface by means of a short mandrel. One end of the mandrel is fixed in the cartridge, and on the other end, pre-verified with the indicator, set the Cams of the lunette.

4 Theoretical Studies of Options to Achieve Quality in the Processing of the Shaft The surface under the Cams should be periodically lubricated to reduce friction. Since in the process of processing the part is heated and it is necessary to weaken the compression of the rear center, it is proposed to conduct processing in the direction from the cartridge to the tailstock, that is, from left to right (Fig. 2). Two variants of the tool are considered: (1) Lamina STAR-LINE cutter: holder ST-SXJBL 2525 M06, carbide inserts ST-VBMT 060404L NN LT 10 and ST-DBMT 060404L NN LT 10 with a cutting angle of 55° and 35°, respectively. (2) Mitsubishi cutter: holder CTGPL 2525 with plates TPGR 160308L NX 2525 and TPGR 160304L NX 2525, plate material is metal-ceramic NX 2525 [25]. To select a tool with the best performance, an analysis of the solid model of cutters has been carried out using the Solidworks software (Fig. 3) [26]. As a result of static analysis, the following values are obtained: stress, displacement and deformation of the composite cutter. For specified machining conditions when turning the outer surface of the Ø60 stem (−0.12; −0.18), Mitshubishi cutter (holder CTGPL2525, plate TPGR 160308L NX 2525) is recommended. Fig. 3 Solid-state cutter model CTGPL2525 after generation of finite element mesh

Tooling Backup of Cutting and Deforming Processing …

55

Experimental industrial processing has been carried out to check the performance and stability of the resulting dimensions and surface roughness, to determine the period of resistance of the plates. The roughness of the surface of the rough stock after finishing has been measured. The full-factor experiment 23 has been set up. It has been revealed the dependence of the roughness value of the machined surface after finishing turning on three factors—the rotational speed of the rough stock n, the depth of cut t and the feed of the cutting tool S. Following experimental processing conditions: assembly part—rod; rough stock- Trumpet 73X15-35 GOST 23270-78; number of parts—48; diameter of treated surface ∅60d9; length of the treated surface/part length,—3769/3942 mm; tolerance of the treated surface cylindrical rod along the entire length of 0.01 … 0.02 mm. The cutting modes are n = 214/250 rpm, t = 2/1.5/1.2 mm, S = 0.26/0.3/ 0.33 mm. When increasing or decreasing the feed value, a decrease in the resistance period of the cutting edge of the plate has been observed. Reducing the supply leads to wear on the back surface of the tool and to reduce its durability. The increase in feed leads to higher temperatures in the cutting zone and to wear on the front surface. To prevent strong heating of the cutting plate abundant cooling of the coolant “Cimstar 536” is used. However, the impact of the feed on tool life is disproportionately small compared to the cutting speed. Cutting speed recommended by the manufacturer is 200 … 300 m/min. Cutting speed when processing rods ∅60d9 is
40 m/min. The vibration has not arisen during the experimental processing. Based on the processing carried out, it has been established: The roughness of the guide surface of the rod is ∅d60d9 Ra 9.43 … 9.47 microns; cylindrical tolerance over the entire length of 0.01 … 0.02 mm. Resistance period is T = 470 min on one cutting edge of the plate TPGR 160308L NX 2525 with a radius of the tip R = 0.8 mm, which corresponds to 8 processed parts. The plate has 3 cutting edges.

5 Experimental Studies of Options to Achieve Quality in the Processing of the Shaft Experimental studies were carried out to determine the dependence of the roughness on the technological factors of the surface plastic deformation treatment process. For the experimental study, a special knurling with a damping element was used. Two “Rod” parts have been tested, outer diameter ∅60d9. Surface roughness corresponds to the drawing, there is no fiber delamination. For further theoretical substantiation of ways to reduce vibrations during cutting and the development of devices for their implementation it is required the development of effective methods for studying the processing of long non-rigid shafts

56

E. I. Yatsun et al.

and hardware. The solution of the problem of stabilization the machining process during cutting will ensure the specified resistance of the cutting tool with high productivity. We have proposed several new designs of damping cutters, qualitatively improving the processing of materials [27–33]. Experimental studies have been conducted to determine the dependence of the roughness value when machining with a dampening cutter. The results obtained suggest that when turning such tools of long non-rigid shafts, it is possible to ensure the roughness of the processed surface and increase the productivity of the process due to the cutting speed. To prevent the appearance of fiber delamination during the running process, the design of a rolling tool with a dampening head has been developed and laboratory tests have been carried out. We have proposed several designs of cutters that reduce vibration levels, including those associated with shock loads. A feature of the design of the cutter is the presence of grooves along the perimeter of the holder for better adhesion to the composite material on a polymer basis. The uniformity of the gaps between the tool holder and the inner walls of the glass is selected technological screws. The gaps are filled with a mixture of self-vulcanizing liquid rubber with a filler of different size fractions. Tests of cutters of standard and modified design at various modes of processing of a shaft for the purpose of identification of parameters of roughness of the processed surface and determination of value of natural frequencies of fluctuations of a turning cutter are carried out. Experimental studies have been conducted to determine the dependence of the roughness in the treatment of a damping cutter. The roughness of the treated surface is achieved in accordance with the requirements. The productivity of the process is increased due to the cutting speed.

6 Conclusion Experimental production and laboratory studies of turning and rolling of the guide surface of the rod were carried Ø60d9. During production tests revealed the following. In the lathe machining of the guide surface, the cutters were used: (1) Lamina STAR-LINE cutter: holder ST-SXJBL 2525 M06, carbide inserts ST-VBMT 060404L NN LT 10 and ST-DBMT 060404L NN LT 10. (2) Mitsubishi CTGPL 2525 with plates TPGR 160308L NX 2525 and TPGR 160304L NX 2525, plate material is metal-ceramic NX 2525. The roughness of the workpiece surface after finishing was measured. The full-factorial experiment 23 was put. The dependence of the roughness of the machined surface after the turning on three factors—the workpiece rotation frequency n, the cutting depth t, and the feed of the cutting tool is determined.

Tooling Backup of Cutting and Deforming Processing …

57

The conducted production studies have revealed that: (1) with an increase in the feed value of the cutting tool above the interval 0,20 … 0,40 mm/rev and the cutting depth of more than 2 mm there is an increase in the roughness of the treated surface; (2) when speed increase above 47 m/min, there was vibration of the cutter. Resistance of the cutter №1 T = 210 min, which corresponds to the processing of 3 parts. Resistance of the cutter №2 T = 470 min, which corresponds to the processing of 8 parts. Resistance of the cutter №3 T = 140 min, which corresponds to the processing of 1–2 parts. Recommended cutter №2 Mitsubishi CTGPL 2525 with plates TPGR 160308L NX 2525 and TPGR 160304L NX 2525, plate material is metal-ceramic NX 2525. In the processing of surface plastic deformation applied special plastic deformation tool with a damping element. Roughness according to requirements. Experimental laboratory studies of the turning treatment of the guide surface of the rod with a damping cutter were carried out. The roughness of the machined surface in accordance with the requirements and increase the productivity of the process due to the cutting speed.

References 1. Kudryashov YA, Yatsun EI, Pavlov YV et al (2010) Methods for achieving the reliability of high-pressure hydraulic cylinders of drilling rigs. In: Proceedings of the Samara scientific center of the Russian Academy of Sciences, pp 401–403 2. Yemel’yanov SG, Remnev AI, Yatsun EI et al (2011) The influence of the formation of growths during turning on the tool life and the quality of the treated surface. STIN, Moscow, pp 30–34 3. Malykhin VV, Pavlov YV, Yatsun EI et al (2011) Improving the stability of the turning process damping cutter. News SWSU, Kursk, pp 122–126 4. Kudryashov YA, Remnev AI, Yatsun EI et al (2011) Modeling and calculation of surface roughness when turning with cutters with composite inserts. In: Proceedings of the Samara scientific center of the Russian Academy of Sciences, Samara, pp 975–979 5. Yemel’yanov SG, Remnev AI, Yatsun EI et al (2011) The mechanism of chip formation when cutting metals. J Mech Eng:73–76 (Moscow) 6. Kudryashov YA, Pavlov YV, Yatsun EI et al (2012) Ensuring the accuracy and quality of products with coatings. Problems of increasing the efficiency of metal working in industry at the present stage, Novosibirsk, pp 14–16 7. YA Kudryashov, YV Pavlov, YI Yatsun et al (2012) Features of calculation of parameters of roughness at processing by the blade tool. News SWSU, Kursk, pp 158–163 8. Yatsun EI, Kudryashov YA, Pavlov YV et al (2012) Technological assurance of reliability of high-pressure hydraulic cylinders of drilling rigs, Novosibirsk, pp 181–184 9. Kudryashov YA, Pavlov YV, Yatsun EI (2012) Improving the efficiency of machining of structurally complex parts with tools from superhard metals. Materials digest of the XXXII International research and practice conference London, 20–25 Sept 2012, pp 25–29

58

E. I. Yatsun et al.

10. Seleznev YN, Yatsun EI, Khomutov RN (2013) Methods of obtaining mathematical models of the process of turning products on machines with numerical control. News SWSU, Kursk, pp 23–28 11. YA Kudryashov, Smirnov IM, Yatsun EI (2014) Selection of tools for finishing processes of structurally complex surfaces of parts. High technologies in mechanical engineering, Moscow, pp 10–14 12. Yatsun EI, Kotov IS, Anikeyeva NP (2016) Study of the resistance of replaceable carbide inserts in the processing of high-alloy steels and wear-resistant cast iron. Chemical and oil and gas engineering, Moscow, pp 21–24 13. YA Kudryashov, Smirnov IM, Yatsun EI et al (2017) Improvement of tool support of turning processes of structurally complex parts. STEEN, Moscow, pp 23–28 14. Yatsun EI, Anikeyeva NP, Kotov IS (2017) Study of the life of replaseable inserts during high-speed steel and wear-resistant cast iron machining. Chem Petrol Eng 52(11–12):758–762 15. Malyhin VV, Yatsun EI, Selesnev YN et al (2017) Development of desing of damping cutting tools. Chem Petrol Eng 52(11–12):763–768 16. YA Kudryashov, Smirnov IM, Yatsun EI et al (2018) Impruved tools for the turning jf complex surface. Russ Eng Res 38(3):223–228 17. Kudryashov YA, Yemel’yanov SG, Yatsun EI (2013) Technological support of manufacturing processes of structurally complex parts. Monograph, TNT. Staryy Oskol, p 90 18. Uchayev PN, Yemel’yanov SG, Yatsun EI (2015) Optimization of application tasks. Staryy Oskol, TNT, p 100 19. Novikov SG, Malykhin VV, Yatsun EI et al (2016) Damping cutters with adjustable stiffness. International scientific and practical conference “New and non-traditional technologies in resource and energy saving” 23 sent. Odessa, pp 136–139 20. Novikov SG, Malykhin VV, Yatsun EI (2014) The design of the damping cutter with controlled rigidity. International scientific-practical conference “Physical and computer technologies” 23–24 dek. Khar’kov, pp 75-82 21. Yemel’yanov SG, Kudryashov YA, Yatsun EI (2012) Precision rationing in mechanical engineering. TNT, Staryy Oskol, p 153 22. Yatsun EI, Malykhin VV, Zubkova OS (2016) Tool maintenance processes of mechanical processing of hard alloys and composites: monograph. ZAO University book, Kursk, p 78 23. Kudryashov YA, Smirnov IM, Yatsun EI (2017) Fundamentals of engineering technology. Prometheus, Moscow, p 280 24. Kudryashov YA, Smirnov IM, Yatsun EI (2017) Fundamentals of Assembly technology of engineering products. Prometheus, Moscow, p 288 25. Mitsubishi materials corporation. http://www.mitsubishicarbide.com/EU/ru…catalog/catalog. html 26. Yatsun EI, Malykhin VV, Novikov SG (2014) Vibroacoustic diagnostics of the cutting tool condition and micro-roughness of the treated surface. Handbook, Engineering journal, pp 31–35 27. YA Kudryashov, Yatsun EI (2015) Cutting of materials. Alpha M, Moscow, p 86 28. Yatsun EI, Malykhin VV et al (2012) Damping cutter. RF Patent №2457077 S1 MPK51 V23V27/00 29. Yatsun EI, Malykhin VV et al (2012) Universal damping cutter. RF Patent 2457078 30. Yatsun EI, Malykhin VV et al (2013) Damping cutter with controlled rigidity. RF Patent 2479385 31. Yatsun EI, Malykhin VV et al (2014) Damping cutter with controlled stiffness. RF Patent 2535196 32. Yatsun EI, Malykhin VV et al (2016) Damping cutter. Patent RF 2582403 33. Yatsun EI, Malykhin VV et al (2016) Universal damping cutter with controlled rigidity. RF Patent 2511193

Determination of Flanging Parameters and Length of Screwing in Producing Holes by the Method of Thermal Drilling in Thin-Sheet Metal P. V. Shalamov, I. A. Kulygina and A. N. Yasnitsky

Abstract In mechanical engineering, products from sheet blanks are widely used, which provide for threaded fastening of various components to them. The formation of holes for thread in such blanks has a number of difficulties. The main problem is to ensure the strength of threaded joints. The parameters that determine the strength of a threaded joint are the diameter and pitch of the thread, the ratio of the mechanical characteristics of the material of the bolt and nut, and the length of screwing. In case of insufficient screwing length, the threaded joint is destroyed. To increase the length of screwing in thin-sheet blanks, sheet bending, welding of sleeves, preliminary punching of holes and other methods are used. However, the existing methods do not have sufficient adaptability. With regard to thin-sheet blanks, the most rational use of the method of forming holes is the method of thermal drilling. This eliminates the use of additional elements to increase the length of screwing. This article presents experimental and theoretical studies of the relationship between the shape and geometrical dimensions of flanging, made by thermal drilling in thin-sheet metal, with a screwing length. The influence of the shape and geometrical dimensions of flanging a hole made by thermal drilling in thin-sheet metal on the length of screwing and the strength of a threaded joint for cutting is considered. Keywords Thermal drilling blanks Threaded joint





Hole flanging
Screwing length
Thin-sheet

1 Introduction A method of flanging in sheet metal by thermal drilling for cutting threads is known [1–4]. Flanging is used to obtain threaded joints in various assemblies and mechanisms. One of the parameters that ensure the strength of the threaded joint to cut is P. V. Shalamov (&)
I. A. Kulygina
A. N. Yasnitsky South Ural State University, 76, Lenin Avenue, Chelyabinsk 454080, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_7

59

60

P. V. Shalamov et al.

the length of screwing [5–8]. Currently in the literature, there is practically no information about the definition of this parameter in flanging made by the method of thermal drilling in sheet metal.

2 Experimental Part Studies of the process of thermal drilling were performed on a vertical drilling machine using a punch moving under a given load [9]. The punch is made of hard alloy VK6 with a diameter of 4.2 mm with a grinding angle of the lead-in part 20°. Samples for flanging were made of steel 0.8 kp with a thickness of 0.8 … 2 mm. Measurement of cutting forces was carried out using a three-component dynamometer model 9257B Kistler (Switzerland) with built-in amplifier of type 5070A01110 [10]. Temperature was measured with a chromel-alumel thermocouple [11]. By the nature of the change in the axial force (see Fig. 1a), the whole cycle of the formation of the flared hole can be divided into four stages. At the first stage (stage I), the billet is heated by frictional forces by thermal drilling, it is inserted into the billet, the plastic flow of metal into the upper and lower parts of the billet with the formation of upper and lower flanging. In this case, the axial force does not change. At the second stage of the process (stage II), an axial force oscillates, which can be associated with the destruction of the lower surface of the workpiece. At the third stage (stage III), as the punch is inserted into the workpiece, the axial force decreases slightly. The formation of the upper and lower flanging continues. At the fourth stage of the process (stage IV), when the cylindrical part of the punch is inserted into the flanging hole, the contact area of the surface of the punch cone with the surface of the hole decreases, the resistance to movement of the punch decreases, and the axial force decreases [12–14]. At the time the punch cone leaves the hole, the axial force decreases to zero. The total formation time of the flared hole is 3.8 s. Figure 1a, b show the stages of the formation of flanging in thin-sheet metal in the process of plastic deformation of the workpiece by the method of thermal drilling (Table 1).

3 Results and Discussions When analyzing Figs. 1 and 2, it was assumed that the shape of the lower and upper flangings is truncated paraboloids with holes, and the outer section lines of the flanging walls are limited by a parabola. This form of flanging is produced in the process of plastic deformation of a metal in a heated state [15].

Determination of Flanging Parameters and Length of Screwing …

61

Fig. 1 Stages of formation of a flanging hole made by a rotating punch in thin-sheet metal: a dependence of the axial effort RAF on the time of the process; b I-II-III-IV—stages of the process

Table 1 Geometrical flanging dimensions Axial force, N

Punch rotation speed, rpm

Time of the process, t, s

Flanging dimensions, mm Height of the Height of the lower ring, upper ring, mm mm

Thickness of the walls of the flanging base, mm

500

2000

3.8

2.5

1.0

1.5

62

P. V. Shalamov et al.

Fig. 2 Geometric parameters of flanging: d—thickness of the workpiece, mm; R— radius of the hole, mm; h— height of the lower flanging, mm; h1—height of the upper flanging, mm; D—thickness of the bases of the upper and lower flangings, mm; 1, 2, 1′, 2′—temperature measurement points; …—calculated points of parabola coordinates

In [16], it was established that in the process of friction of one metal on another, a sub-layer flow of metal occurs in a softer layer. Therefore, the shape of the curves of the outer walls can be associated with different flow rates of individual metal layers due to their different mechanical properties during heating. To confirm this assumption, an experimental determination was made of the temperature by the thickness of the bases of the upper and lower flangings at points 1, 2 and 1′, 2′ (see Fig. 2) in the final stage of the formation of flanging. The results of temperature measurements are given in Table 2. Table 2 Experimental values of temperatures at the upper and lower bases of flanging rings at measurement points 1, 2 and 1′, 2′ (Fig. 2) Ring base Upper

Temperature at the base of the rings, T, °C 577 (measuring point 1) 406 (measuring point 2)

Lower

570 (measuring point 1′)

430 (measuring point 2′)

Determination of Flanging Parameters and Length of Screwing …

63

From Table 2, it follows that the temperature along the base thickness of the upper and lower flanging decreases. It is known [17] that, for steel 08KP, the tensile strength and yield strength with decreasing metal temperature within the above limits (see Table 2) increase from 150 to 300 and from 70 to 90 N/m2, respectively. This confirms the assumption that the shape of the curves of the outer surfaces of the upper and lower (see Fig. 2) flangings in the form of a parabola is formed as a result of different flow rates of individual metal layers (subsurface plastic flow) due to changes in its mechanical characteristics with temperature. Determine the shape of the upper and lower flangings, substituting in the equation of the parabola y ¼ ax2

ð1Þ

experimental values: y = h; x = D (see Fig. 2). Then the parabola coefficient is a ¼ Dh2 (a = 2.5; a1 = 1.3, respectively, for the upper and lower flangings). It has been established (see Fig. 2) that the experimental and calculated forms of parabolas coincide. This may confirm the conclusion that the flanging shapes represent truncated paraboloids. Knowing the shape of the outer flanging surfaces, consider the relationship between the parameters of forming a hole with flanging using the method of thermal drilling and its geometrical dimensions. From Figs. 1 and 2, it follows that the lower flanging volume is defined as Vl:f : ¼ bVw

ð2Þ

where b—proportion of deformable metal billet participating in the formation of the lower; Vw—volume of the deformable metal of the workpiece is (Fig. 2). VD ¼ pR2 d

ð3Þ

The volume of the lower flanging, taking into account Fig. 2 is defined as Vl:f: ¼ Vlp  Vhlp

ð4Þ

where Vl.p—volume of the lower truncated paraboloid of the flanging; Vhlp—volume of the hole of the lower truncated paraboloid of the flanging. The volume of the lower paraboloid flanging is determined by the formula [18] i p h Vlp ¼ h ðR þ DÞ2 þ R2 2

ð5Þ

and the volume of the hole of the lower truncated paraboloid can be determined as Vhlp ¼ pR2 h

ð6Þ

64

P. V. Shalamov et al.

Then substituting Eqs. (3), (4), (5), (6) into Eq. (2), we obtain i p h h ðR þ DÞ2 þ R2  pR2 h ¼ bpR2 d 2

ð7Þ

the height of the lower flanging is h¼

2bR2 d ð2R þ DÞD

ð8Þ

The height of the upper flanging with regard to Eq. (8) is h1 ¼

2ð1  bÞR2 d ð2R þ DÞD

ð9Þ

From Eqs. (5), (8) and (9), it follows that the geometrical dimensions of flanging, truncated paraboloids depend not only on the amount of deformable metal, determined by the hole diameter and the thickness of the workpiece, but also on the share of the deformable metal in the formation of flanging. For example, with the above-mentioned geometrical dimensions of flanging (see Table 1), the share of participation of the deformable metal in the formation of the lower flanging, determined by formula (8), is 65%. It was established experimentally that with a decrease in the thickness of the workpiece, the coefficient “b” increases, and this will lead to a decrease in the share of participation of the deformable metal in the formation of the upper flanging and to an increase in the size of the lower flanging (Fig. 3). The obtained data on the shape and geometrical dimensions of flanging make it possible to determine the length of screwing Lscr. When determining the length of screwing in flanging, obtained by thermal drilling, it is necessary to take into

Fig. 3 Influence of the thickness of the workpiece on the share of the participation of the deformed metal in the formation of the lower flanging

Determination of Flanging Parameters and Length of Screwing …

65

Fig. 4 Scheme for calculating the effective height of the lower flanging. x = Hp—outer diameter of the threaded joint

account the shape of the curves of the outer surfaces of the rings in the form of a parabola. In this case Lscr, given the geometrical dimensions of the flanging, will be determined by the effective height of rings hef, depending on the height of the thread (pitch) and the thickness of the workpiece. Figure 4 shows a scheme for calculating the effective height of the lower flanging ring. The equation of a straight line passing through the outer diameter of a threaded hole is written as x ¼ Hp ¼ 0:866P

ð10Þ

where Hp—metric thread height; P—thread pitch [19]. The coordinate of point B (see Fig. 4), taking into account (1), (10), can be written as yB ¼

h
Hp2 D2

ð11Þ

66

P. V. Shalamov et al.

and the effective height of the lower flanging, taking into account Eq. (11) and Fig. 3 is hef ¼ h  B ¼ h 

h
Hp2 D2

ð12Þ

and after transformation (12), taking into account Eq. (10), we obtain that "

ð0; 866PÞ2 hef ¼ h 1  D2

# ð13Þ

Then the length of screwing Lscr taking into account the thickness of the part and the effective height of the lower flanging is "

Lscr

# ð0; 866PÞ2 ¼h 1 þd D2

ð14Þ

Knowing the length of screwing Lscr, it is possible to determine the force [20], which determines the cut of the threads in the flanging according to Eq. (15) Fsh ¼ p
d
Lscr
Kg
Km
sc:s

ð15Þ

where Кg—thread completeness coefficient [21]; Кm—coefficient taking into account the unevenness of deformation; sc.s—strength of the material of the nut (flanging) on the cut [21]. The results of the calculation of the length of screwing, taking into account the thickness of the workpiece, the effective height of the bottom flange, as well as the force of cutting the coils according to Eqs. (14) and (15) are given in Table 3. It also shows the experimental values of the force of cutting threads. To eliminate the deflection of the workpiece during the test, it was installed in a special device. The appearance of the threaded joint is shown in Fig. 5.

Table 3 Calculated and experimental values of the force of cutting threads Type

Length of screwing Lscr, mm

Cutting force of threads, N Calculation Experiment

Without flanging 2.0 2800 3040 With flanging 4.05 5200 6080 Note RAF = 500 N; d = 2 mm; h = 2.5 mm; D = 1.0 mm. M5x0.8 thread; Kg = 0.87 [21]; Km = 0.55 [21]; sc:s = 200 MPa [17]

Determination of Flanging Parameters and Length of Screwing …

67

Fig. 5 Appearance of the threaded joint

4 Conclusions From Table 3, it follows that the flanging, made with the use of the method of thermal drilling in sheet metal with a thickness of 2 mm for threading, allows to increase the length of screwing and the strength of the threaded joint of the thread two times. If according to the design requirements the upper flanging is not removed, then to increase the length of screwing to the Eq. (14) the effective height of the upper flanging is added, calculated by Eq. (13) (see Fig. 2) taking into account the data in Table 1. Thus, the length of screwing in flanging, by thermal drilling in thin-sheet metal, depends on the thickness of the workpiece, the effective height, determined by the shape and geometrical dimensions of flanging and thread pitch. With the decrease in the thickness of the workpiece, the volume of the deformed metal involved in the formation of the lower flanging increases.

References 1. FormDrill (2017) Information on http://formdrill.com/english/formdrill.htm 2. Centredrill (2017) Information on http://www.centerdrill.de 3. Khomenko V (2005) Forming of attachment points in thin-walled parts using the method of plastic drilling. Thesis, Altai State Technical University of I.I. Polzunov 4. Zolotov O (2007) Improvement of processing technology for thin-walled parts using plastic drilling. Thesis, Altai State Technical University. I.I. Polzunova 5. Shalamov P (2012) Formation of holes for threads by a rotating punch in thin-sheet billets. Thesis, South Ural State University

68

P. V. Shalamov et al.

6. Shalamov P (2011) Formation of holes by a rotating punch in a thin-sheet billet. Bulletin of SUSU 18:81–84 7. Usachev V (2012) Technology of thermoplastic drilling. Actual Probl Mech Eng 10:181–184 8. Usachev V (2011) Technology and tool for thermoplastic formation of holes. In: Proceedings of the Samara scientific center of the Russian academy of sciences, vol 13, pp 429–432 9. Shalamov P, Kazantseva Ju (2017) Thermal drilling with force-feed tool. Procedia Eng 206:985–990 10. Utenkov V, Bykov P (2012) Possibilities of using a dynamometer Kistler for testing metal-cutting machines. Eng Bull 10:1–23 11. Nikonov N (2015) Thermocouples. Types, characteristics, designs. Moscow 12. Miller S, Blau P, Shih A (2007) Tool wear in friction drilling. Int J Mach Tools Manuf 47:1636–1645 13. Shalamov P, Kulygina I, Yaroslavova E (2016) ANSYS software based on the study thermal drilling process. Procedia Eng 150:746–752 14. Taureza M, Song H, Castan S (2014) The effect of workpiece material on friction in Microforming and lubrication efficiency. J Mater Process Technol 214:998–1007 15. Guzeev V, Shalamov P, Shultc E (2006) Calculation geometry of heading made by technical welding in thin sheet metal. Progress Technol Mech Eng 63:175–178 16. Ilyushin A Plastic. Elastic-plastic deformations. Moscow 17. Serebrenitsky P (2004) General technical reference. St. Petersburg 18. Bronstein I (2010) Handbook of mathematics for engineers and students of high schools. St. Petersburg 19. Anurev V (2001) Reference book of a designer-mechanical engineer, vol. 1. Mechanical Engineering, Moscow, 920 p 20. Shalamov P, Cherepkov I, Kulygin V (2018) Strenght of a threaded joint on cut in holes with flanges formed by a rotating punch in thin-sheet workpieces. Solid State Phenom 284: 441–446 21. Tyunyaev A (2013) Machine parts. St. Petersburg

Practice of Dimensional Modeling in the Implementation of the Methodology of Group Interchangeability M. G. Galkin and A. S. Smagin

Abstract This paper is devoted to the modeling of dimension links during assembly using single-loop graph dimensional schemes, which describe the connection of separate parts during the implementation at an assembly site of group interchangeability methods, which reflect the specific conditions of the current production. Within the context of dimension modeling, the calculations based on the developed mathematical model which could be applied to both single-loop and multi-loop dimensional schemes with dependent dimension chains are evaluated. In particular, an algorithm for large-scale production, which consists in determining the dimension scatter range of the assembled parts by means of sorting into dimension groups is evaluated. The design procedure considered on a specific example, which describes one of the variants of interaction of the dimensional relationships of the parts. The calculation procedures result in the formation of a common information model in the form of a numeric matrix representing the analog of the dimension of the scheme. In the final form, it accumulates the corrected values of tolerances, group tolerances, mean and limit deviations of each of the links within a given number of size groups. This form of representation of calculation results allows to simplify the mechanism of automation of calculation procedures to accelerate the process of modeling of assembly operations by this method. Keywords Dimension modeling interchangeability Size group





Graph model
Dimension links
Group

1 Introduction Dimension modeling of processing and assembly stages of production helps to achieve in the most cost-effective manner the necessary precision of machine parts during design, production, and use [1–5]. M. G. Galkin
A. S. Smagin (&) Ural Federal University, 19, Mira Street, Yekaterinburg 620000, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_8

69

70

M. G. Galkin and A. S. Smagin

It is obvious that, when manufacturing specific parts forming an assembly unit, it is not always possible to achieve complete interchangeability during the assembly stage of production with the amount of money allocated [6, 7]. The likelihood of sticking to the complete interchangeability principle may be determined based on the average allocation tolerance which is calculated according to the following Eq. (1) [8–10]: Tc ¼

½T  n

ð1Þ

where [T]—the tolerance of dimension chain’s master link, n—the number of components of dimension chain’s links. Tc is evaluated based on the accuracy of processing of parts forming an assembly chain. Important aspects such as the complexity of processing operations and dimensions of parts are also taken into consideration. If the Tc value does not comply with the conditions of cost-effective production, then, in order to achieve the needed precision of a master link, methods of incomplete interchangeability should be used, and it is often done in current production conditions [11, 12]. This paper describes the dimension modeling of the assembly using the group interchangeability method. As a mathematical model, a dimensional scheme is used, which is formed on the basis of a linear graph structure [13]. As an example, Fig. 1 shows an assembly unit of a mechanical transmission, where the idle gear 2 is mounted as a cantilever on the intermediate axis 1. In order to maintain the specified operational requirements, it is necessary during the installation to ensure Fig. 1 Assembly model unit

Practice of Dimensional Modeling in the Implementation …

71

that the axle play between the gear and the retaining ring 3 is within the 0.2  0.6 mm range to ensure the free rotation of the gear [14]. At the same time, the retaining ring is anchored to an axis using the adjustment screw 4.

2 The Calculation Algorithm 2.1

Problem Statement

To solve this problem, it is proposed to use the method of group interchangeability since there are no obvious adjusting elements in the design that would allow the specified accuracy of the assembly process. It is well-known that the use of group methods is typical for the objectives of large-scale and mass production. When this method of selective assembly is used during the design stage, it is known that complete interchangeability will be provided not for all the mating parts, but only for those that fall into a certain dimension group [10, 15, 16]. Moreover, each part may be manufactured according to the cost-effective tolerances. The range of scatter range of the master link forming during the assembly will be subsequently eliminated via preliminary sorting of all parts into dimension groups [17]. The mathematical model to describe this assembly process is built along the X-axis and forms four matings plus the location of the adjustment screw axis (see Fig. 1). The working clearance between the gear and the retaining ring, which is a master link with a nominal value of zero, is also taken into consideration. The mating numbers of an assembly unit should increase along the direction of this axis. The proposed model of an assembly unit (Fig. 1) may be described using a linear graph. The apexes of this graph model will represent the sites where matings of parts are formed during the assembly process. Links which connect the apexes will represent nominal dimensions of design components interacting with one another during the assembly operations. The developed model is a dimension scheme of the entire assembly process in the direction of the respective reference axis. This scheme may contain dimension chains of various assembly units forming a joint. Therefore, this method of modeling assembly operations allows for analysis of interdependence chains. This interdependence will be implemented through common links of these chains. The design element in Fig. 1 contains in its dimension scheme the only dimension chain (see Fig. 2). Using the model in Fig. 2, we get the Eq. 2 for this dimension chain to determine the nominal value of the master link: ½K  ¼

n X i¼1

Ui
Ki

ð2Þ

72

M. G. Galkin and A. S. Smagin

Fig. 2 Dimension scheme of the assembly process

where Ki is the nominal size of a component link of a dimension chain, Ui is the gear ration of a component link of a dimension chain, n is the number of components of dimension chain’s links. For each link, the gear ratio Ui is calculated using the developed dimension scheme (see Fig. 2). In order to do so, one should assign a master link to a model and start from the smaller apex to the bigger one (in this case, from apex 2 to apex 1). If, during this operation, the number of apexes of a component link decreases, the gear ratio of this link will equal −1; if the opposite happens, the Ui will equal 1. In technical literature, component links with positive values are loosely called increasing, and those with negative values are called decreasing. [18] Therefore, the K1 link will have a positive gear ratio, and the K2, K3, and K4 will have the negative ones. [17]. In accordance with the requirements listed above, the record of the master link 0:6 will be as follows: ½K  ¼ 0 þ þ 0:2 . Then, according to these data, [K] = 0 mm, [ES] = 0.6 mm, [EI] = 0.2 mm, [T] = 0.4 mm, and [C] = 0.4 mm. Subsequently, guided by the economic accuracy of the methods of processing parts, the tolerances for each of the component links of the chain are controlled. In this case, one may suppose that, for each link, the accuracy does not exceed the 12th grade, which points to the cost-effective nature of the processing operations. Taking these conditions into consideration, the accuracy parameters of links (Fig. 1) will take values listed in Table 1 [4, 19]. In accordance with the data in Table 1, the total dimension tolerance for this assembly method, when calculated for the max-min values [10, 13], will be as follows: TR ¼ 0:46 þ 0:14 þ 0:22 þ 0:4 ¼ 1:22 mm: It is obvious that this tolerance value exceeds the similar value which complies with requirements of an assembly unit (Fig. 1), i.e., TR > [T], or 1.22 > 0.4 mm. The same conclusion can be made from (1) based on which the average accuracy of chain links should not exceed 0.15 mm.

Table 1 Dimensional accuracy of the assembly, mm

T1

T2

T3

T4

0.46

0.14

0.22

0.4

Practice of Dimensional Modeling in the Implementation …

2.2

73

Application of Group Interchangeability Methodology to Solve the Problem

According to the group interchangeability methods previously used for calculations, it is necessary to determine the total number of dimension groups for parts to be sorted before assembly. The equation will be as follows [8, 9]: ngr ¼

TR : ½T 

ð3Þ

whence ngr ¼ 1:22=0:4 ¼ 3:05: Three groups may be accepted. The next step is to correct the tolerances of the parts in accordance with the accepted number of dimension groups. In the course of this correction, it is necessary to ensure the homogeneity of the parts mated in order for the tolerance of the master links in each group to take the same values. Mathematically, this may be represented by the following equation: k X i¼1

TðKiþ Þ ¼

m X

TðKj Þ ¼ 0:5
ngr
½T

ð4Þ

j¼1

where Pk TðKiþ Þ are tolerance values of links with positive gear ratios, Pi¼1 m  are tolerance values of links with negative gear ratios. j¼1 TðKj Þ In accordance with (4), the dependency for links with positive gear ratios will be as follows: TðK1þ Þ ¼ 0:5
ngr
½T;

ð5Þ

whence TðK1þ Þ ¼ 0:5
3
0:4 ¼ 0:6 mm. The dependency for links with negative gear ratios will be as follows: TðK2 Þ þ TðK3 Þ þ TðK4 Þ ¼ 0:5
ngr
½T;

ð6Þ

whence TðK2 Þ þ TðK3 Þ þ TðK4 Þ ¼ 0:6 mm. Structurally, the following can be taken for each link: TðK2 Þ ¼ 0:1 mm, TðK3 Þ ¼ 0:2 mm, and TðK4 Þ ¼ 0:3 mm. The next step is to calculate the required tolerances for each link under consideration within the group. To do this, we use the following equation [8]: Tigr ¼

Ti ; ngr

ð7Þ

74

M. G. Galkin and A. S. Smagin

whence T1gr ¼ 0:6=3 ¼ 0:2 mm; T2gr ¼ 0:1=3 ¼ 0:03 mm; T3gr ¼ 0:2=3 ¼ gr 0:07 mm; T4 ¼ 0:3=3 ¼ 0:1 mm: Then, it is necessary to assign average deviations for each link in Cn (Ki) groups. In this case, n is the dimension group number, i is the number of a link in a dimension chain. Furthermore, the condition of equality of the average deviation of master links in each dimension group (CnR) and the average deviation of the master link of the [C] dimension chain Eq. (8). CnR ¼ ½C

ð8Þ

In this case, C1R ¼ C2R ¼ C3R ¼ C4R ¼ ½C, whence C1R ¼ C2R ¼ C3R ¼ C4R ¼ 0:4 mm: For the first group, the dependence, according to which it is necessary to select the average deviations of the links, is as follows [8]: C1R ¼ ½C1 ðK1þ Þ  ½C1 ðK2 Þ þ C1 ðK3 Þ þ C1 ðK4 Þ:

ð9Þ

Further on, the following values may be assigned to the links of this group C1 ðK1þ Þ ¼ 0:2 mm; C1 ðK2þ Þ ¼ 0:03 mm; C1 ðK3þ Þ ¼ 0:07 mm; C1 ðK4þ Þ ¼ 0:1 mm. If we insert these values into Eq. (9), we get the following value C1R ¼ 0:2  ½0:03 þ ð0:07Þ þ ð0:1Þ ¼ 0:4 mm, which corresponds to Eq. (8). In the next step, we need to determine the maximum deviations for the links that will fall into the corresponding group during sorting. Dependencies will be as follows: ESn ðKi Þ ¼ Cn ðKi Þ þ

Tngr 2

ð10Þ

EIn ðKi Þ ¼ Cn ðKi Þ 

Tngr 2

ð11Þ

whence the tolerance values for links in group 1 will equal to the ones listed in Table 2. Let us now move on to dimension group 2. To do this, we add the value of group tolerance to the average deviations of each of the links. The dependency will be as follows: Cn þ 1 ðKi Þ ¼ Cn ðKi Þ þ Tigr Overall, the values for dimension group 2 are listed in Table 2.

ð12Þ

Practice of Dimensional Modeling in the Implementation …

75

Table 2 Parameters of links in dimension group 1 Dimension number

Group 1 ES1 C1 (Ki) (Ki) (mm) (mm)

EI1 (Ki) (mm)

Group C2 (Ki) (mm)

[K] K1 K2 K3 K4

0.4 0.2 −0.03 −0.07 −0.1

0.2 0.1 −0.045 −0.105 −0.15

0.4 0.4 0 0 0

0.6 0.3 −0.015 −0.035 −0.05

2 ES2 (Ki) (mm)

EI2 (Ki) (mm)

Group C3 (Ki) (mm)

0.6 0.5 0.015 0.035 0.05

0.2 0.3 −0.015 −0.035 −0.05

0.4 0.6 0.03 0.07 0.1

3 ES3 (Ki) (mm)

EI3 (Ki) (mm)

0.6 0.7 0.045 0.105 0.15

0.2 0.5 0.015 0.035 0.05

The same calculations are made for dimension group 3. The values for this group are also listed in Table 2.

3 Conclusion In conclusion, according to the calculation results, one can develop a general information model that will describe in numerical form the dimensional scheme in Fig. 2. This model will in fact be a table containing link parameters for all dimension groups (see Table 3). It contains the adjusted values of tolerances, group tolerances, and average and maximum tolerance extremes of each of the links within the dimension groups. A table like this will also significantly simplify the automation of computational procedures to speed up the technological design. From Table 3, it is possible to determine the tolerances for each dimension of a part of an assembly unit under consideration. Obviously, the smallest tolerances for all chain links are obtained in dimension group 1, since the calculation of the tolerance extremes begins with it. Therefore, the lower tolerance extreme of the corresponding link (EI) should be taken from the first group, and the upper tolerance extreme (ES) from the last one. Then, the dimensions in question must be manufactured with the following tolerances: 0:7 þ 0:045 þ 0:105 þ 0:15 K1 ¼ 60 þ þ 0:1 ; K2 ¼ 30:045 ; K3 ¼ 70:105 ; K4 ¼ 500:15

Nominal dimensions

0 60 3 7 50

Dimension number

[K] K1 K2 K3 K4

1.2 0.6 0.1 0.2 0.3

Ti 0.4 0.2 0.03 0.07 0.1

Tigr

Tolerance

0.4 0.2 −0.03 −0.07 −0.1

0.6 0.3 −0.015 −0.035 −0.05

0.2 0.1 −0.045 −0.105 −0.15

0.4 0.4 0 0 0

0.6 0.5 0.015 0.035 0.05

0.2 0.3 −0.015 −0.035 −0.05

Deviations of dimension links by groups Dimension group 1 Dimension group 2 C1 (Ki) ES1 (Ki) EI1 (Ki) C2 (Ki) ES2 (Ki) EI2 (Ki)

Table 3 Information model used to implement the group interchangeability method, mm

0.4 0.6 0.03 0.07 0.1

0.6 0.7 0.045 0.105 0.15

0.2 0.5 0.015 0.035 0.05

Dimension group 3 C3(Ki) ES3 (Ki) EI3 (Ki)

76 M. G. Galkin and A. S. Smagin

Practice of Dimensional Modeling in the Implementation …

77

References 1. Kosilova A, Meshcheryakov R (1986) Spravochnik tekhnologa-mashinostroitelya (Handbook of technology-machine builder). Mashinostroenie Publ, Moscow 2. Kosilova A, Meshcheryakov R (2001) Spravochnik tekhnologa-mashinostroitelya (Handbook of technology-machine builder). Mashinostroenie Publ, Moscow 3. Dunaev P, Lelikov O (2000) Konstruirovanie uzlov i detaley mashin (Construction of units and machine parts). Vysshaya shkola, Moscow 4. Ashikhmin V, Zakuraev V (2005) Razmernyy analiz pri tekhnologicheskom proektirovanii (Dimensional analysis in design process). UGTU-UPI, Yekaterinburg 5. Novikov M (1980) Fundamentals of assembly of machinery. Mashinostroenie Publ, Moscow 6. Kondakov A (2012) Technological design in machinery. Knorus, Moscow 7. Myagkov V (1982) Tolerance values and fit. Reference book. Mashinostroenie Publ, Leningrad 8. Paley M, Romanov A, Braginsky V (2009) Tolerance values and fit. Reference book. Politekhnika, Saint-Petersburg 9. Matalin A (2008) Machine-building technology. Mashinostroenie Publ, Moscow 10. Gusev A, Kovalchuk Y, Kolesov I (1986) Machine-building technology: handbook for university students specializing in machine building. Mashinostroenie Publ, Moscow 11. Zholobov A (2005) Design of technological process of machinery assembly. Novoye Znaniye, Minsk 12. Balakshin B (1982) Theory and practice of machine-building technology. Mashinostroenie Publ, Moscow 13. Korsakov V (1983) Assembly and installation of machine-building products. Mashinostroenie Publ, Moscow 14. Markov N (1992) Standardization of accuracy in machine-building industry. Stankin publishing house, Moscow 15. Mostalygin G, Tolmachevsky N (1990) Machine-building technology. Mashinostroenie Publ, Moscow 16. Agafontsev V (2016) Dimension chains. Practical calculation: handbook. Yug publishing house, Krasnodar 17. GOST 16320–80 Dimension chains. Methods of calculation of flat chains 18. GOST 25346–89 Basis terms of interchangeability. Unified system of tolerances and fit. General provisions, various tolerances and basic deviations 19. Solonin I, Solonin S (1980) Calculation of assembly and technological dimension chains. Mashinostroenie Publ, Moscow

Determination of Settling Efficiency of Solid Finely Dispersed Particles Within Devices with Rectangular Separators I. N. Madyshev, O. S. Dmitrieva and A. V. Dmitriev

Abstract Design of dust-collecting device, characterized by low hydraulic resistance and high efficiency of collecting the fine particles, has been developed. It has a simple construction. It lacks small parts and small channels, which increases the reliability of its operation. It can be disassembled and restored. Separators are using for solid particles and droplets capturing from gas flow by means of gravity, inertia of centrifugal force. They can be used in industrial vacuum cleaners and can allow increasing the cleaning efficiency. The influence of design and technological parameters on collecting efficiency has been researched. The availability of several rows of separators leads to an increase in the settling efficiency of solid dispersed particles due to a more ordered structure of the gas flow. The studies of the influence of the geometrical constant of the proposed device on the settling efficiency show that in order to collect the relatively large particles, the dimensions of settling elements should be increased up to 50 mm. Keywords Separator


Sedimentation of particles
Gas cleaning

1 Introduction Intensification of technological processes and creation of high-performance devices in chemical, petrochemical, energy, metallurgical, and food industries lead to an inevitable increase in atmospheric emissions, such as significant amount of toxic dust and harmful gaseous impurities [1–6].

I. N. Madyshev
O. S. Dmitrieva Kazan National Research Technological University, 68, Karl Marx Str, Kazan 420015, Russia A. V. Dmitriev (&) Kazan State Power Engineering University, 51, Krasnoselskaya Str, Kazan 420066, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_9

79

80

I. N. Madyshev et al.

Currently, in order to collect the aerosols from exhaust gases, a wide range of gas-purifying equipment is used (cyclonic separators, filters, dust collectors, settling chambers, etc.) [7–11]. However, due to low efficiency of this equipment when collecting the medium and finely dispersed dust, a need for frequent replacement or cleaning of filter elements, the scope of use is limited. In addition, the hydraulic resistance of such devices can reach 3000 Pa that leads to an increase in energy costs for conducting the dust and gas purification processes [12–14]. The solution to the problem of energy-efficient removal of medium and finely dispersed dust (less than 10 µm) from dusty gas flows can be the use of device, developed by the authors. This device is based on gravitational and inertial settling of solid particles during their dynamic interaction with stationary elements of various shapes [15].

2 Study Materials and Methods The proposed device [16] (Fig. 1) consists of n-number of rows of stationary dust-collecting elements, chequerwise arranged. Separators are the plates with the installed angle-shaped reflecting elements on edges. Such design allows to reduce the area of circulating currents, decreasing the efficiency of the dispersed particles settling processes. The operating principle of dust-collecting device is the following: When dusty gas flow moves between the separators, a centrifugal force field appears, providing the coagulation of finely dispersed particles and contributing to their uniform settling on the whole surface of collecting elements. Availability of several rows of separators leads to an increase in settling efficiency of solid dispersed particles due to a more ordered structure of the gas flow. The purpose of numerical studies is to determine the settling efficiency of solid dispersed particles on the surface of separators within the proposed device (Fig. 1). The study was conducted by means of ANSYS Fluent software in order to simulate the interaction of finely dispersed particles, being carried away by air flow, with the elements of dust-collecting device. Fig. 1 Dust-collecting device

Determination of Settling Efficiency of Solid …

81

The studied device with a height of 100 mm consists of 2, 4, and 6 rows of separators, and there are 5 settling elements in each row. In the course of study, geometrical dimensions of the device in a scale were changed proportionally to its characteristic dimensions, namely, to the width of the separator (without taking into account the angle-shaped elements), equal to 50 mm, and to the gap between adjacent rows of separators, equal to 19 mm. In addition, the average air flow rate was changed from 3 up to 7 m/s, and the size of solid dispersed particles was changed in the range of 1–20 µm. In order to simulate the turbulent flows, a modification of two-parametric turbulence model k-x SST, showing satisfactory conformance with experimental data, obtained in earlier studies, was used [17–20]. Boundary conditions: It was assumed that all surfaces of settling elements have absolute elasticity, i.e., the fraction of absorbed inertia from impacting particles was not taken into account. The settling efficiency of solid dispersed particles can be determined by the equation: E ¼1

N N0

ð1Þ

where N0—total number of particles; N—number of particles, carried away by the gas flow.

3 Results of Study The results of the study show that settling efficiency of solid finely dispersed particles increases with an increase in average gas flow rate and the number of rows of separators (Fig. 2). Fig. 2 Dependency of settling efficiency on a diameter of particles at different average gas flow rate Wav, m/s: 1–3; 2–5; 3–7; full line—4 rows of separators, dashed line—2 rows

82

I. N. Madyshev et al.

Fig. 3 Dependency of settling efficiency on a width of separators at different diameters of dispersed particles a, µm: 1–1; 2–5; 3– 10; 4–15; 5–20; Wav = 5 m/s; number of rows of separators n=4

For example, when the number of rows is equal to 4 and average gas flow rate is of 5 m/s, the settling efficiency of particles with a diameter of 10 µm achieves 100%. Studies of influence of the separator’s width on settling efficiency show the following: When collecting the small particles (up to 7–8 µm), an increase in geometrics of settling elements leads to a decrease in efficiency, and, at the same time, the use of elements with a width of 10 mm is the most efficient (Fig. 3). When collecting the dispersed particles with a diameter of 10 µm, the separator’s width of 25 mm is optimal. Studies of the influence of geometrical constant of the proposed device on settling efficiency show that in order to collect the relatively large particles, the dimensions of settling elements should be increased up to 50 mm, at Mb = 1 (Fig. 4).

Fig. 4 Dependency of settling efficiency on a diameter of particles at different geometrical constants of dust-collecting device Mb: 1–0.2; 2–0.5; 3–1; Wav = 5 m/s; number of rows of separators n = 4

Determination of Settling Efficiency of Solid …

83

Fig. 5 Dependency of collecting the dispersed particles with 100% efficiency on a width of settling elements at different average gas flow rate Wav, m/s: 1–3; 2–5; a number of rows of separators n = 4

Figure 5 shows the dependency of collecting the dispersed particles with 100% efficiency on a width of settling elements at different average gas flow rates. It can be seen that the elements with a width of 10 mm provide full collection of dispersed particles with a diameter of 5 µm at a number of rows of separators equal to 4 and average gas flow rate of 5 m/s.

4 Conclusion Thus, the conducted numerical studies show that in order to improve the settling efficiency within the proposed devices, a width of the separators should be chosen correctly, depending on the dispersed composition of the collected particles. For example, when settling of large particles with a diameter of more than 15 µm, it is recommended to choose separators with a width of at least 25 mm. However, in practice, due to the polydispersed composition of the sprayed particles, it is required to design the devices, consisting of successive rows with different width of settling elements, thereby providing primary and fine purification of gas emissions. Acknowledgements The research was conducted with funding from the RF President’s grant project No. MK-417.2019.8.

References 1. Kudrov AI, Kuz’min AV, Rakov YY (2017) Effective fuel temperature of WWER-1000. MATEC Web Conf 141:01030. https://doi.org/10.1051/matecconf/201711001046 2. Abramovskikh AA, Ulyanova UA, Lavrinenko SV, Matveeva AA, Malishev D (2017) Modernisation of boiler unit automation system at acting power station. MATEC Web Conf 91:01002. https://doi.org/10.1051/matecconf/20179101002

84

I. N. Madyshev et al.

3. Samokhvalov NM, Vinogradov VV (2014) Stationary and efficiency of cleaning dust from gases from on slotted filter. Theor Found Chem Eng 48:837–841. https://doi.org/10.1134/ S0040579514060086 4. Ribeiro Rui PPL, Esteves Isabel AAC, Mota José PB (2017) Two-column relay simulated moving-bed process for gas-phase separations. Sep Purif Technol 182:19–28 5. Song J, Hu X (2017) A mathematical model to calculate the separation efficiency of streamlined plate gas-liquid separator. Sep Purif Technol 178:242–252 6. Zaripov SK, Soloveva OV, Skvortsov EV (2015) Analytical model of the transport of aerosol particles in a circular hole inside a porous medium. Transp Porous Media 107:141–151. https://doi.org/10.1007/s11242-014-0429-x 7. Zhao B, Su Y (2018) Particle size cut performance of aerodynamic cyclone separators: Generalized modeling and characterization by correlating global cyclone dimensions. J Aerosol Sc. 120:1–11 8. Lin JC-T, Hsiao T-C, Hsiau S-S, Chen D-R, Chen Y-K, Huang S-H, Chen C-C, Chang M-B (2018) Effects of temperature, dust concentration, and filtration superficial velocity on the loading behavior and dust cakes of ceramic candle filters during hot gas filtration. Sep Purif Technol 198:146–154. https://doi.org/10.1016/j.seppur.2017.06.014 9. Benassi L, Dalipi R, Consigli V, Pasquali M, Borgese L, Depero LE, Clegg F, Bingham PA, Bontempi E (2017) Integrated management of ash from industrial and domestic combustion: a new sustainable approach for reducing greenhouse gas emissions from energy conversion. Environ Sci Pollut Res 24:14834–14846. https://doi.org/10.1007/s11356-017-9037-y 10. Elsayed K (2015) Optimization of the cyclone separator geometry for minimum pressure drop using Co-Kriging. Powder Technol 269:409–424 11. Patra G, Chakraborty S, Meikap BC (2018) Role of vortex finder depth on pressure drop and performance efficiency in a ribbed hydrocyclone. S Afr J Chem Eng 25:103–109. https://doi. org/10.1016/j.sajce.2018.04.001 12. Sheng Y, Fang L, Sun Y (2018) An experimental evaluation on air purification performance of Clean-Air Heat Pump (CAHP) air cleaner. Build Env 127:69–76. https://doi.org/10.1016/j. buildenv.2017.10.039 13. Xiaowen S, Lin Z, Yuxin X (2015) Evaluation of a flue gas cleaning system of a circulating fluidized bed incineration power plant by the analysis of pollutant emissions. Powder Technol 286:9–15 14. Chen SJ, Fu Y, Huang YX, Tao ZC, Zhu M (2016) Experimental investigation of CO2 separation by adsorption methods in natural gas purification. Appl Energy 179:329–337. https://doi.org/10.1016/j.apenergy.2016.06.146 15. Dmitriev AV, Zinurov VE, Dmitrieva OS, Nguyen WL (2017) Collecting particles from flue gases by rectangular separators. Bul of the Tech Un 20:78–80 16. Dmitriev AV, Dmitrieva OS, Madyshev IN, Nikolaev AN (2017) A device for fine dust removal. RU Patent No. 171615, 7 June 2017 17. Zaripov SK, Solov’eva OV, Solov’ev SA (2015) Inertial deposition of aerosol particles in a periodic row of porous cylinders. Aerosol Sci Technol 49:400–408. https://doi.org/10.1080/ 02786826.2015.1036834 18. Dmitrieva OS, Dmitriev AV, Madyshev IN, Nikolaev AN (2017) Flow Dynamics of Mass Exchangers with Jet-Bubbling Contact Devices. Chem Pet Eng 53:130–134. https://doi.org/ 10.1007/s10556-017-0308-8 19. Solovev SA, Soloveva OV, Popkova OS (2018) Numerical simulation of the motion of aerosol particles in open cell foam materials. Russ J Phys Chem A 92:603–606. https://doi. org/10.1134/S0036024418030275 20. Safikhani H, Mehrabian P (2016) Numerical study of flow field in new cyclone separators. Adv Powder Technol 27:379–387. https://doi.org/10.1016/j.apt.2016.01.011

Method for Predicting Thermal Characteristics of Machine Tools Based on Experimental Modal Analysis A. N. Polyakov, A. N. Goncharov and I. V. Parfenov

Abstract The article describes the method for predicting the thermal characteristics of CNC machine tools working at finishing cutting modes with variable cutting speeds. This allows not taking into account the generation of heat in the main sources due to additional loads from cutting without introducing significant distortions in the adequacy of the mathematical model in the construction of thermal characteristics. The thermodeformation model of the machine is represented by a system of thermal characteristics that describe both its thermal and deformation behavior. A feature of the proposed method is the use of the entire set of approximated experimental thermal characteristics for the complex mode of operation of the machine under consideration. Each approximated thermal characteristic used in the model is formed from the results of the full-scale experiment with a continuous operation of the machine tool at a fixed spindle rotational speed. The mathematical description of each thermal characteristic is based on the experimental modal analysis, in which the modal parameters of the thermodeformation model are determined from the experiment. A feature of the approximated thermal characteristics is their multimodal representation. The results of full-scale and computational experiments are presented in the article. Keywords Machine tool Modal analysis


Thermal characteristics
Thermodeformation model


1 Introduction Direct control of the relative displacement between the tool and the workpiece in machining is the subject of the development activity. Today, there is no industrial implementation of this type of development [1, 2]. Thermal processes in the

A. N. Polyakov (&)
A. N. Goncharov
I. V. Parfenov Orenburg State University, 13, Pobeda Avenue, Orenburg 460018, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_10

85

86

A. N. Polyakov et al.

technological system determine more than 50% of the total geometric errors of the machined parts. The constructions of modern CNC machine tools include all known solutions [3] aimed at minimizing the influence of thermal processes on the machining accuracy. Nevertheless, the problem of the temperature error of machine tools is still relevant today. For example, when studying the thermal deformations of the portal frame structure of the machine tool model FC300, the excess temperatures did not exceed 10 K. In this case, the temperature displacements were about 40 lm [4]. In the work [5], experimental data on temperatures and temperature displacements for a three-axis CNC milling machine are presented. Temperature measurements were performed using 80 sensors. The eddy current sensors fixed the coordinate displacements of the moving parts of the machine tool in a non-contact manner. The experiments were carried out with a spindle rotating at a speed of 6000 rpm in combination with various feeds along the X and Z axes (7.5 and 15 m/min). The maximum displacements along the spindle axis were fixed at a level of 60 lm. In the work [6], the temperature error of the Leaderway-V450 machining center was studied. The maximum error was obtained along the Z axis. The experimental temperature field of the machine tool was formed from the readings data of ten temperature sensors which were placed on the head of the machine tool. The experiment was carried out at spindle idling with the spindle rotating at a constant spindle speed. The tests were carried out at the spindle speeds: 1000 and 2000 rpm. The duration of each experiment was four hours; the experimental data were collected every three minutes. To measure the thermal deformation of Z direction, inductive displacement sensors were used (measuring accuracy is ±0.5 lm). Maximum displacements did not exceed 30 lm after two hours of machine operation process at a spindle speed of 2000 rpm. In the article [7], for the spindle error analyzer was used to measure the temperature displacements of the spindle. The experiment was carried out on a modern high-speed machine tool that provides a maximum spindle speed of 40,000 rpm. The experiment was carried out in the range of spindle speeds from 2000 to 24,000 rpm. At the spindle rotation speed of 2000 rpm, the temperature displacements were 12 lm, at a spindle speed of 6,000 rpm–24 lm, and at a spindle speed of 24,000 rpm, the temperature spindle displacements were fixed at a level of 70 lm. The range of excess temperatures level at different rotational speeds was from 10 (for 2000 rpm) to 25 °C (at 24,000 rpm). Minimization of the temperature error of modern machine tools is solved in various ways: both at the design stage and at the stage of their operation [8–11]. At the same time, when trying to ensure high dimensional accuracy of the product, the practice of machining proves that the implementation of only design solutions to ensure the heat stability of the machine tool is not effective enough. This is explained by the action of many factors arising in real production conditions associated with changing operation regimes [12]. Therefore, during the operation phase of CNC machine tools, it is necessary to use various special automated or automatic compensation systems for the resulting temperature error [3]. In this case, the error correction can be realized by correcting the controlled axes of the machine

Method for Predicting Thermal Characteristics …

87

tool or by introducing a correction of the trajectory of motion relative to the workpiece that is entered into the trajectory of the tool [13]. The predictive models of thermal characteristics are the basis of modern automated systems for compensation of temperature errors in CNC machine tools. At present, when developing automated systems for compensating the temperature error of CNC machine tools, methods of artificial intelligence are widely represented. Among the most well-known methods may include the following: methods of evolutionary calculations (genetic, bee and ant algorithms, a method of particle swarm optimization); neural networks; and algorithms of fuzzy systems. Despite all the attractiveness of new approaches to solving the problem of providing a given machining accuracy, they have objective difficulties in practical implementation: high computational complexity, a large amount of experimental information for tuning such systems, lack of guarantee of finding a global optimum, and others [14– 18]. Along with this, in the automated systems for compensating the temperature error of CNC machine tools, methods of experimental modal analysis based on finding modal parameters from experimental characteristics are successfully implemented [19, 20].

2 Mathematical Model 2.1

Basic Model

Analysis of full-scale experiments results leads to the conclusion that when operating the machine tool at one constant spindle rotational speed, only four types of curves of thermal characteristics can be distinguished (Fig. 1a). The first type is curve 1, which is characterized by an increase in temperature or temperature displacements (Fig. 1a). The second type is a combination of two sections of the curves, characterized by a transition from the heating process to the cooling process (curve 2). The third type is a combination of two sections of curves: In the first

Fig. 1 a Typical thermal characteristics; b experimental thermal characteristics for 400 V

88

A. N. Polyakov et al.

section the cooling of the machine occurs, and in the second section, the heating occurs (curve 3). The fourth type is a combination of three sections of curves, characterized by alternating cooling and heating sections (curve 4). When the machine tool is heated, the temperature or temperature displacements along the axes are generally described by the dependencies: YðtÞ ¼ x1
ð1  et=x2 Þ þ x3
et=x2

ð1Þ

where x1 is the maximum level of excessive temperatures (or temperature displacements) established at a given spindle rotation speed (hereinafter, for the sake of brevity, we will use the term «amplitude»), °C (lm); x2 —thermal time constant, min; x3 —the initial temperature level (or temperature displacements; in the future, due to the invariance of temperatures and temperature displacements, we will use the term «initial level»), fixed at the initial moment of time (lm). At the zero level x3 , the relation (1) takes the form: YðtÞ ¼ x1
ð1  et=x2 Þ

ð2Þ

The process of cooling the machine tool is described by the dependency: YðtÞ ¼ x1 þ ðx3  x1 Þ
et=x2

ð3Þ

In Eq. 3, the parameter x1 is taken as the threshold asymptotic value. For example, if Eq. 3 describes a change in temperature, then x1 is taken as the ambient temperature. Analysis of Eqs. 1, 3 shows the invariance of Eq. 1 for both processes: The process of heating or cooling is determined only by the quantitative ratio of modal parameters x1 and x3 . • The first type of curve (curve 1, Fig. 1a) is a heating curve from the zero level; therefore, it can be described by Eq. 2. • The second type of curve (curve 2, Fig. 1a): Eq. 2 is used in the first section of the curve up to time point s1 , and the dependence characterizing the cooling of the machine is used in the second section. Therefore, curve 2 will be described by two modes. The first mode describes the heating process, and the second mode describes the cooling process:

YðtÞ ¼ x11
ð1  et=x21 Þ þ x12
ð1  et=x22 Þ þ x32
et=x22

ð4Þ

where x11 is the amplitude of the first mode; x12 is the amplitude of the second mode; x21 is the thermal time constant of the first mode; x22 is the thermal time constant of the second mode; x32 is the initial level of the second mode.

Method for Predicting Thermal Characteristics …

89

• The third type of curve (curve 3, Fig. 1a): This type of curve is similar to the previous one. The difference lies in the different order of the cooling and heating processes, and the initial level may be non-zero: YðtÞ ¼ x11
ð1  et=x21 Þ þ x31
et=x21 þ x12
ð1  et=x22 Þ þ x32
et=x22

ð5Þ

where x31 is the initial level of the first mode. • The fourth type of curve (curve 4, Fig. 1a). This type of curve is the most complex, since it consists of three sections alternating cooling and heating processes. In this case, the three-modal approximation function has the form: YðtÞ ¼

3   X x1i
ð1  et=x2i Þ þ x3i
et=x2i

ð6Þ

i¼1

2.2

Operating Cycle

In accordance with the technological process of manufacturing a part, each operation performed on a CNC machine tool consists of a number of machining passes. For each machining pass, cutting parameters are assigned. Thus, a complex mode of operation of the cutting machine is formed in accordance with the alternation of spindle speeds. Taking into account the duration of each machining pass, a technological operation can be represented by diagrammatic work. This allows us to describe the thermal characteristics of the machine tool using Eqs. 2, 4–6. The correction of these equations consists in recalculating the modal parameters for the initial levels of each temperature modes and time sampling for each machining pass. As an example, we can consider an operating cycle represented by two spindle speeds n1 and n2 . We will assume that at the first machining pass the second type of curve is realized, and at the second pass, the third type of curve is realized. For the first machining pass, the curve is described by Eq. 4. The second machining pass will use the equation of the form: YðtÞ ¼ xII11
ð1  et 2 ½0; t2 

II

=xII21

Þ þ xII31
et

II

=xII21

þ xII12
ð1  et

II

=xII22

Þ þ xII32
et

II

=xII22

; tII ð7Þ

xII31 ¼ xI11
ð1  et1 =x21 Þ þ xI12
ð1  et1 =x22 Þ þ xI32
et1 =x22 ; xII32 ¼ 0 I

I

I

where t1 ; t2 —the duration of the first and second matching pass, respectively, min; I, II—machining pass indices.

90

A. N. Polyakov et al.

2.3

Defining Modal Parameters

The modal parameters used in Eqs. 1–7 are determined by experimental thermal characteristics by solving an optimization problem: The minimum objective function gðx1i ; x2i ; x3i Þ is achieved by a certain combination of modal parameters used to describe the thermal characteristics: vffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi !2ffi u m u X ðx1i
ð1  et=x2i Þ þ x3i
et=x2i Þ ! min ð8Þ gðx1i ; x2i ; x3i Þ ¼ t Y3 ðtÞ  i¼1

x1i;min  x1i  x1i;max ; 0\x2i  x2i;max ; x3i;min  x3i  x3i;max where m is the number of temperature modes; Y3 ðtÞ—experimental values of thermal characteristics. The experimental thermal characteristics of the machine tool are formed according to experimental studies conducted at different spindle speeds with certain discreteness. The magnitude of the discreteness is determined by the feature of the thermal deformation system of an appropriate machine tool. For example, for the two machine tools used in this study, 400 V (manufactured in Sterlitamak, Russia) and HAAS TM-1P (manufactured by HAAS, USA), noticeable changes in modal parameters were recorded at discreteness of 500 rpm. Figure 1b shows the experimental thermal characteristics as functions for four spindle rotational speeds: 250, 1000, 3000, and 5000 rpm. Experimental curves are marked with an “e” index. According to the modal parameters found from the solution of the optimization

Fig. 2 Screenshot of the program module implementing the experimental modal analysis

Method for Predicting Thermal Characteristics …

91

problem (8), curves with the index “0” were constructed; curves with a five percent uncertainty modal parameters (±5%) are indexed “+5” and “−5”, respectively. To determine the modal parameters, an upgraded version of the software module that implements the experimental modal analysis was developed. Figure 2 shows a screenshot of the module’s working window. In the working window of this module, you can select nine areas with controls.

3 Discussion of the Results For experimental verification of the proposed method for predicting the thermal characteristics of machine tools, experimental studies were carried out for various operating cycles implemented on the 400 V machine tool. As an example, Fig. 3a shows the operating cycle scheme and experimental thermal characteristics for 12 thermal sensors. For the machine tool was set operating cycle of the six sections of time. In the first section, the machine tool was heated at a spindle rotational speed of 250 rpm for 10 min. Then, during four-time intervals of 40 min each, the spindle rotational speeds were changed according to the scheme: 1000, 3000, 5000, and 3000 rpm. At the last sixth time interval, the spindle rotated at a rotational frequency of 1000 rpm for 120 min. Previous studies showed that the greatest correlation was established between the displacements of the spindle nose and the readings of the thermal sensor 1 installed near the front bearing [19]. Figure 3b shows the predicted thermal response for the first thermal sensor and experimental response 1 (Fig. 3b). The prediction error did not exceed 7%. The greatest prediction error was recorded when changing the spindle rotational speed from 5000 to 3000 rpm.

Fig. 3 a Experimental thermal characteristics for operating cycle; b the experimental and predicted thermal characteristics

92

A. N. Polyakov et al.

4 Conclusion The proposed method for predicting the thermal characteristics of machine tools uses an analytical solution of the heat conduction equation and the optimization problem solution for finding modal parameters. The determination of modal parameters was carried out according to the experimental data obtained for continuous modes of operation of the machine tool. The analysis of the accuracy of matching the experimental and predictive characteristics showed that the proposed method can reliably be used to build predictive models. At the same time, the fixed prediction error at the moment of transition from the heating mode to the cooling mode determines the reserve for improving the prediction accuracy. It consists of building more complex prediction models based on the account of transients that occur when changing the heating and cooling modes. Acknowledgements The reported study was funded by RFBR and Orenburg region according to the research project No. 19-48-560001.

References 1. Leun EV, Leun VI, Shahanov AE (2018) The sapphire tips of the active control devices of detail’s dimensions of produced with hollows and ridges, with the possibility of determining the lateral approximation of ridges. J Phys: Conf Ser 1050:012043. https://doi.org/10.1088/ 1742-6596/1050/1/012043 2. Leun EV, Leun VI, Sysoev VK et al (2018) The active control devices of the size of products based on sapphire measuring tips with three degrees of freedom. J Phys: Conf Ser 944:012073. https://doi.org/10.1088/1742-6596/944/1/012073 3. Mayr J, Jedrzejewski J, Uhlmann E et al (2012) Thermal issues in machine tools. CIRP Ann 61:771–791. https://doi.org/10.1016/j.cirp.2012.05.008 4. Suzumura F, Makihara H, Ohtani K et al (2011) Thermal deformation prediction in machine tool model by using transfer functions with time delay. Manuf Process Technol. Adv Mater Res 189:4064–4067. https://doi.org/10.4028/www.scientific.net/AMR.189-193.4064. Switzerland Trans Tech Publications 5. Horejs O, Mares M (2015) Real-time compensation of machine tool thermal errors including cutting process. J Mach Eng 15:5–18 6. Miao E, Liu Y, Xu J et al (2017) Thermal error modeling method with the jamming of temperature-sensitive points’ volatility on CNC machine tools. Chin J Mech Eng 30:566–577. https://doi.org/10.1007/s10033-017-0109-1 7. Lee J, Kim DH, Lee CM (2015) A study on the thermal characteristics and experiments of high-speed spindle for machine tools. Int J Precis Eng Manuf 16:293–299. https://doi.org/10. 1007/sl2541-015-0039-8 8. Bushuev VV, Kuznetsov AP, Sabirov FS et al (2016) Trends in research on metal-cutting machines. Russ Eng Res 36:488–495. https://doi.org/10.3103/S1068798X16060083 9. Kuznetsov AP (2015) Temperature control of metal-cutting machines. Russ Eng Res 35:194– 199. https://doi.org/10.3103/S1068798X15030090 10. Li X (2001) Real-time prediction of workpiece errors for a CNC turning centre, part 2. Modelling and estimation of thermally induced errors. The Int J Adv Manuf Technol 17:654– 658. https://doi.org/10.1007/s001700170129

Method for Predicting Thermal Characteristics …

93

11. Zhou ZD, Gui L, Tan YG et al (2017) Actualities and development of heavy-duty CNC machine tool thermal error monitoring technology. Chin J Mech Eng 30:1262–1281. https:// doi.org/10.1007/s10033-017-0166-5 12. Mares M, Horejs O (2017) Modelling of cutting process impact on machine tool thermal behaviour based on experimental data. Procedia CIRP 58:152–157. https://doi.org/10.1016/j. procir.2017.03.208 13. Martinov GM, Obukhov AI, Kozak NV (2018) The usage of error compensation tools of CNC for vertical milling machines. Russ Eng Res 38:119–122. https://doi.org/10.3103/ S1068798X18020120 14. Hou R, Yan Z, Du H et al (2018) The application of multi-objective genetic algorithm in the modeling of thermal error of nc lathe. Procedia CIRP 67:332–337. https://doi.org/10.1016/j. procir.2017.12.222 15. Guo Q, Xu R, Yang T et al (2016) Application of Gram and AFSACA-BPN to thermal error optimization modeling of CNC machine tools. The Int J Adv Manuf Technol 83(5):995–1002. https://doi.org/10.1007/s00170-015-7660-7 16. Abdulshahed AM, Longsta AP, Fletcher S et al (2016) Thermal error modelling of a gantry-type 5-axis machine tool using a Grey Neural Network model. J Manuf Syst 41:130– 142. https://doi.org/10.1016/j.jmsy.2016.08.006 17. Li Y, Zhao J, Ji S (2018) Thermal positioning error modeling of machine tools using a bat algorithm-based back propagation neural network. The Int J Adv Manuf Technol 97:2575– 2586. https://doi.org/10.1007/s00170-018-1978-x 18. Eskandari S, Arezoo B, Abdullah A (2013) Positional, geometrical, and thermal errors compensation by tool path modi_cation using three methods of regression, neural networks, and fuzzy logic. The Int J Adv Manuf Technol 65:1635–1649. https://doi.org/10.1007/ s00170-012-4285-y 19. Polyakov AN, Goncharov AN, Kamenev SV (2018) Assessing the temperature error in operational machine tools. Russ Eng Res 38:408–410. https://doi.org/10.3103/ S1068798X18050131 20. Zhang C, Gao F, Yan L (2017) Thermal error characteristic analysis and modeling for machine tools due to time-varying environmental temperature. Precis Eng 47:231–238. https://doi.org/10.1016/j.precisioneng.2016.08.008

Development of New Types of Contact Devices for Heat-Mass Transfer Apparatuses, Used at Petrochemical Enterprises I. N. Madyshev, O. S. Dmitrieva and A. V. Dmitriev

Abstract One of the most important ways to increase the efficiency of technological processes is to improve the column heat-mass transfer apparatuses. The design of a jet-film contact device for carrying out the heat-mass transfer processes within gas-liquid systems has been developed. The advantage of this device is a low hydraulic resistance and high mass transfer efficiency with relatively low energy demands. The development of a new technical solution has been carried out with the software module. The separation efficiency of the contact device with its different geometrical dimensions and diameters of the collected drops has been studied. It is established that the greatest separation efficiency of the contact device is provided when its geometrical constant is equal to 0.5, and medium dispersed aerosol particles with a diameter of 20 µm are collected by 99.5%. The high collecting efficiency at small geometrical constant of contact device is explained by the creation of large centrifugal forces due to small radii of gas flow vortices. Keywords Contact device


Heat-mass transfer
Separation efficiency

1 Introduction Column heat-mass transfer apparatuses are widely used in chemical, petrochemical, and oil refining industries when carrying out the processes of absorption, rectification, liquid extraction, and gas separation [1, 2]. One of the most important ways to increase the efficiency of such processes is to improve the column heat-mass transfer apparatuses and, above all, contact devices. I. N. Madyshev
O. S. Dmitrieva Kazan National Research Technological University, 68, Karl Marx Str., Kazan 420015, Russia A. V. Dmitriev (&) Kazan State Power Engineering University, 51, Krasnoselskaya Str., Kazan 420066, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_11

95

96

I. N. Madyshev et al.

The tendency for developing the apparatuses with high unit capacity has led to the development of a large number of new designs of random and regular packings in recent decades, having higher throughput and slightly greater efficiency. As a rule, random packings are characterized by higher hydraulic resistance and are less reliable in operation with contaminated media and uneasy to be maintained [3, 4]. Regular packings have a more ordered structure, which has a positive effect on their hydrodynamic and mass transfer parameters. The most famous random packings are manufactured by the following foreign and home companies: HY-PAK, CASCADE-RINGS, GIPH, GIAP, Ingekhim; as for the regular packings— INTALOX of NORTON company, Mellapak of Sulzer company, VAKUPAK, KEDR, Glitch-grid, Flexypack, Perform-grid, Ingekhim, Koch-glitsch, etc. [5–11]. When developing the new designs of contact devices, the main technological task is to increase the specific surface area of contact of phases by implementation of intensive interaction of two-phase gas-liquid flows and to decrease the hydraulic resistance of the apparatuses. The authors propose to solve this issue by using the developed and experimentally tested contact devices with film interaction of gas and liquid, having different modifications [12].

2 Research Materials and Methods Contact device with film interaction of gas and liquid [13] (Fig. 1) is a nozzle block, consisting of drain cups 1 for collection and distribution of liquid over the cross section of the heat-mass transfer apparatus. Drain cups 1 are arranged on vertical longitudinal and transverse partitions 2. Round holes, executed within partitions 2, lead to a decrease in the metal consumption for the proposed design. Within the cross section, the drain cups 1 are made in the form of a square, while in the bottom there are petals 3 in the form of circular segments, bent down and directed to the surface of the vertical partitions 2. The contact device with the film interaction of gas and liquid, proposed by the authors, operates as follows. The liquid, dispersing through the distribution device of the heat-mass transfer apparatus, enters the drain cups 1, which, through the bent petals 3, forms a film, freely flowing down the surface of the vertical partitions 2. In this case, an ascending gas flow moves toward the flowing down liquid film. Drain cups 1 are chequerwise arranged in the cross section of the device. Moreover, the below and above arranged drain cups 1 are shifted against the considered level. In this regard, the gas in the proposed nozzle moves along the Z-shaped path, which allows to increase the relative rate of the working media, while intensifying the mass, energy, and momentum transfer processes. The geometric dimensions of the contact device with the film interaction of gas and liquid are chosen on the basis of equality of the areas for the passage of gas in the transverse and longitudinal sections of the heat-mass transfer apparatus. Lack of local expansions and narrowings for the ascending gas flow leads to decrease in the

Development of New Types of Contact Devices …

97

Fig. 1 Contact device with gas and liquid film interaction: 1 drain cup; 2 partitions; 3 petals

hydraulic resistance of the contact device, and lack of parts, manufactured with high accuracy, provides low manufacturing cost of the proposed nozzle blocks. The purpose of the numerical studies is to determine the efficiency of aerosol particles’ settling on the liquid film within the proposed contact device. The studies were conducted by means of ANSYS Fluent software program, which simulated the interaction of gas and liquid flows, as exemplified by air-water system at the temperature of 20 °C. The separation efficiency was estimated both within one contact stage and within the apparatus, consisting of 3 stages. At the same time, in the course of studies the geometrical dimensions of device were changed in a scale, proportional to its characteristic dimensions, namely, to the width of drain cup, equal to 100 mm and the wall height, equal to 47.5 mm. Moreover, dimensions of particles (spherical water drops) were changed within the range of 1–30 lm. The liquid level in the drain cup was taken as maximum. Hydrodynamic calculations were carried out on the basis of Navier–Stokes equations, using the finite volume method. The calculation was carried out in accordance with the turbulence model—SST, showing satisfactory conformance with experimental data, obtained in the course of previous studies [14–17]. In order to simplify the numerical calculation, the following assumptions were made: Thickness of contact elements and walls of case were not taken onto account, and the adhesion condition was set for the walls of contact elements. The separation efficiency can be estimated by means of efficiency value of the aerosol particles’ settling:

98

I. N. Madyshev et al.

E¼

N ; N0

ð1Þ

where N0—total number of particles; N—number of particles, settled on the flowing down liquid film.

3 Research Results In the course of numerical studies, the dispersed particles were evenly distributed over the cross section of the contact device, while their total number was 1000. The research results show that the separation efficiency of liquid drops with a diameter of less than 10 lm within one contact stage is almost independent from geometrical constant and is not more than 25% (Fig. 2). When the diameter of particles increases, the separation efficiency also increases. Thus, for example, drops with a diameter of more than 50 µm are collected by 100%. Analyzing the graph, shown in Fig. 3, it can be seen that the greatest separation efficiency of contact device is provided when its geometrical constant is equal to 0.5, and medium dispersed aerosol particles with a diameter of 20 µm are collected by 99.5%. The high collecting efficiency at small geometrical constant of contact device is explained by creation of large centrifugal forces due to small radii of gas flow vortices. Figure 4 shows that the separation efficiency of drops with a diameter of 1 µm is 0.74, at geometrical constant Mb = 0.5. With a further increase in geometrical constant, the separation efficiency tends to a value of 0.53. Therefore, when

Fig. 2 Dependency of separation efficiency within one contact stage on the diameter of drops at different geometrical constants of contact device Mb: 1 0.5; 2 0.6; 3 0.75; 4 0.85; 5 1; 6 2

Development of New Types of Contact Devices …

99

Fig. 3 Dependency of separation efficiency of contact device on the diameter of drops at different geometrical constants Mb: 1 0.5; 2 0.6; 3 0.75; 4 0.85; 5 1; 62

Fig. 4 Dependency of separation efficiency of contact device on geometrical constant at different diameters of drops a, µm: 1 1; 2 5; 3 10; 4 20; 5 30

developing the design of the contact devices with gas and liquid film interaction, the width of drain cup should be reduced; however, in practice, the use of contact elements with a width of less than 50 mm leads to excessive metal consumption for design.

100

I. N. Madyshev et al.

This paper [18] proposes a dependency for the calculation of critical diameter of the water drop: acr ¼

10:7rL qG Ur2

ð2Þ

where rL—surface tension, H/m; qG—gas density, kg/m3; Ur—relative gas flow, m/s.

4 Conclusion According to the estimated calculations by Eq. 2, when the actual gas flow rate, within the narrowing part of the contact device with gas and liquid film interaction, is equal to 20 m/s, the critical diameter of the formed drops will not exceed 1.6 mm. Thus, the conducted numerical studies show that the settling efficiency of such drops will be 100%, which is confirmed by numerous experimental data [19, 20]. Therefore, when designing the heat-mass transfer apparatuses with the proposed contact devices, there is no need for the use of additional separation devices or drop collectors. Acknowledgements The research was conducted with funding from the RF President’s grant project No. MK-4522.2018.8.

References 1. Kulov NN, Gordeev LS (2014) Mathematical modeling in chemical engineering and biotechnology. Theor Found Chem Eng 48:225–229. https://doi.org/10.1134/S0040579514 030099 2. Khafizov FSh, Afanasenko VG, Khafizov IF, Ash Khaibrakhmanov, Boev EV (2008) Use of vortex apparatuses in gas cleaning process. Chem Pet Eng 44:425–428. https://doi.org/10. 1007/s10556-008-9081-z 3. Kolev N, Kralev B, Kolev D (2013) Gas side controlled mass transfer in a new packing with stamped horizontal lamellae operating at extremely low liquid loads. Chem Eng Process 63:44–49. https://doi.org/10.1016/j.cep.2012.07.004 4. Wei ZJ, You ZL, Gui SQ (2010) Gas pressure drop and mass transfer characteristics in a cross-flow rotating packed bed with porous plate packing. Ind Eng Chem Res 49:3732–3740. https://doi.org/10.1021/ie9009777 5. Farakhov MI, Laptev AG, Basharov MM (2015) Modernization of mass-exchange equipment by new packings in chemical engineering. Theor Found Chem Eng 49:233–238. https://doi. org/10.1134/S0040579515030033 6. Maćkowiak J (2011) Model for the prediction of liquid phase mass transfer of random packed columns for gas-liquid systems. Chem Eng Res Des 89:1308–1320. https://doi.org/10.1016/j. cherd.2011.01.021 7. Contact elements for the column apparatuses (System INTALOX®, Packed Tower System). Catalogue of the company “KOCH-GLITSCH”

Development of New Types of Contact Devices …

101

8. “Metal Random Packings” (Catalogue of the company “Sulzer Chemtech”, 22.64.06. 40-v04-50) 9. Boev EV, Ivanov SP, Afanasenko VG, Nikolaev EA (2009) Polymeric drop-film sprinklers for cooling towers. Chem Pet Eng 45:454–459. https://doi.org/10.1007/s10556-009-9209-9 10. Bessou V, Rouzineau D, Prévost M, Abbé F, Dumont C, Maumus J-P, Meyer M (2010) Performance characteristics of a new structured packing. Chem Eng Sci 65:4855–4865. https://doi.org/10.1016/j.ces.2010.05.029 11. Li X, Yang X, Li H, Shi Q, Gao X (2018) Significantly enhanced vapor-liquid mass transfer in distillation process based on carbon foam ring random packing. Chem Eng Process 124:245–254. https://doi.org/10.1016/j.cep.2018.01.005 12. Dmitrieva OS, Dmitriev AV, Madyshev IN, Nikolaev AN (2017) Flow dynamics of mass exchangers with jet-bubbling contact devices. Chem Pet Eng 53:130–134. https://doi.org/10. 1007/s10556-017-0308-8 13. Dmitriev AV, Dmitrieva OS, Madyshev IN, Nikolaev AN, Kruglov LV (2017) Contact device with film flow of liquid for heat and mass transfer apparatus. RU Patent 171022, 17 May 2017 14. Solovev SA, Soloveva OV, Popkova OS (2018) Numerical simulation of the motion of aerosol particles in open cell foam materials. Russ J Phys Chem A 92:603–606. https://doi. org/10.1134/S0036024418030275 15. Borisov BV (2016) Features applications of the approaches when constructing efficient algorithms during the modelling of some intracanal flows. EPJ Web of Conf 110:01012. https://doi.org/10.1051/epjconf/201611001012 16. Ponomarev KO, Orlova EG, Feoktistov DV (2016) Effect of the heat flux density on the evaporation rate of a distilled water drop. EPJ Web of Conf 110:01060. https://doi.org/10. 1051/epjconf/201611001060 17. Zaripov SK, Solov’eva OV, Solov’ev SA (2015) Inertial deposition of aerosol particles in a periodic row of porous cylinders. Aerosol Sci Technol 49:400–408. https://doi.org/10.1080/ 02786826.2015.1036834 18. Klinskiy BM, Kudravtsev AV (2012) Justification of requirement to the vakue of mass concentration and dispersivity of water drops when designing the bench-scale plant for simulation of rain falling. Eng 81:10–12 19. Dmitriev AV, Madyshev IN, Dmitrieva OS, Nikolaev AN (2017) Research dispersing liquid and gas in the contact device with an increased range of stable operation. Ecol and Ind of Russ 21:12–15. https://doi.org/10.18412/1816-0395-2017-3-12-15 20. Madyshev IN, Dmitrieva OS, Dmitriev AV, Nikolaev AN (2015) Assessment of change in torque of stream-bubble contact mass transfer devices. Chem Pet Eng 51:383–387. https://doi. org/10.1007/s10556-015-0056-6

Stiffness Maximization on the Basis of Layout Characteristics of the Elastic Machine System and Milling Process R. M. Khusainov and A. R. Sabirov

Abstract The paper dwells upon improving the efficiency of milling by increasing the stiffness of the technological system. The stiffness is affected by the system layout geometry, which features the non-uniformity of strains in different spatial directions, and is characterized by the axes of stiffness. The paper defines the axes of stiffness by means of finite-element modeling of the technological system, which consists of the machine supports, the tool, the tool fixture, and the workpiece. Finite-element modeling is performed on any computer-aided engineering system. As a result of the calculation, the deformation of the most characteristic element of the technological system is determined. Based on the finite-element computations, the theory of elasticity is used to find the direction of stiffness axes in the machine workspace. It is herein proposed to use such direction in order to find the cutting parameters by reference to the cutting-force components ratio when milling. Using this method during process engineering enables optimizing the system layout as well as the cutting parameters. Keywords Principal axes


Finite elements
Stiffness

1 Introduction Stiffness is one of the key operability indicators for many machines. Speaking of static stiffness, the most common and recognized definition is the one proposed by Sokolovsky [1]. In his view, the stiffness of a machine-fixture-tool-part system is the ratio of the cutting-force component Py (applicably to turning) to the Y-axis displacement (in a commonly accepted coordinate system) as determined when

R. M. Khusainov (&)
A. R. Sabirov Naberezhnye Chelny Institute (Branch) of Kazan Federal University, 68/19 Mira Prospekt, Naberezhnye Chelny 423812, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_12

103

104

R. M. Khusainov and A. R. Sabirov

subjected to full cutting force. Stiffness affects both the machining accuracy [2] and performance [3], as the latter greatly depends on the vibration stability of the machine, which in its turn is determined by stiffness.

1.1

Relevance

Strains in cutting systems do affect the machining accuracy [4]. This effect is especially relevant for machining non-rigid parts, as well as when using non-rigid tooling [5, 6]. When machining precision (critical) parts, it is important to pre-evaluate possible strains so as to program compensating such errors or to eliminate them by optimizing the tooling or machining diagram selection [7, 8]. This evaluation is important, as it helps avoid defects, thus avoiding material and time losses when using equipment. Papers on this topic [9–11] make emphasis on the eigen- and contact stiffness of technological-system elements. Meanwhile, the stiffness of the system in total is also affected by its layout, as it determines how the elements, each having eigenand contact stiffness of its own, will accommodate loads. This topic is not well-elaborated in papers, but it is also extremely important, as it offers a lot of room for improving the accuracy and dynamic stability of machines. The geometric characteristics that can be used to find the stiffness of a technological system are the axes of stiffness, also referred to as principal-strain axes in the elasticity theory [12]. The axis of maximum stiffness refers to such direction in the geometric space of a technological system where static elastic strains are minimal. The minimumstiffness axis is the direction, where static elastic strains are maximum [13].

1.2

Statement of Problem

Analyzing the layout characteristics of modern machines by top foreign manufacturers (especially milling machines) makes it clear that the high stiffness and vibration stability of such machines is mainly due to good design solutions that optimize the load accommodation. It is therefore obvious that there is a need for studies that would provide scientific foundations for reducing the static elastic strains on the basis of layout characteristics manifested in the axes of stiffness.

Stiffness Maximization on the Basis of Layout …

105

2 Research Essentials In the context of computer-integrated “digital” production, the axes of stiffness can be found by finite-element modeling (FEM) of a technological system [14, 15], which consists of the machine supports, the cutting and the auxiliary tools, the tool fixture, and the workpiece [16–18]. Modeling and computation produces the components of the strain tensor for the basic elastic element, which is the most important element according to [19]: 2

e11 Te ¼ 4 e21 e31

e12 e22 e32

3 e13 e23 5 e33

ð1Þ

where eij are linear and angular strains. Principal strains ei can be found by solving the cubic equation [19]: e3 þ I1 e2 þ I2 e þ I3 ¼ 0

ð2Þ

where Ii are the invariants of the strain tensor “(1)”. Solving Eq. “(2)” has produced the following principal-strain values: 2

3 0; 003 e ¼ 4 0; 001 5 0; 003

ð3Þ

The components e1 and e3 correspond to the maximum principal strains directed along the minimum-stiffness axes; the component e2 corresponds to the minimum principal strain directed along the maximum-stiffness axis. Strain-tensor components can be used to find the orientation of the principal-strain axes [19]. The directions of the principal-strain axes can be found by computing the directional cosines li, mi, ni while also solving the system of equations: 8 ðei  e11 Þli  e12 mi  e13 ni ¼ 0 > > < e12 li þ ðei  e22 Þmi  e23 ni ¼ 0 ð4Þ e l þ e m  ðei  e33 Þni ¼ 0 > > : 2 13 i 2 232 i li þ mi þ ni ¼ 1 Solving this problem produces the axes of stiffness for the workpiece (Fig. 1a) and the tool (Fig. 1b) of the milling-machine system. Using the results thereof enables solving the following problems: 1. Optimize the layout, i.e., find such relative positioning of tools and workpieces (by adjusting the fixture design or the axis rotation in five-axial machines) so as to approximate the maximum-stiffness axis to the resultant cutting force. This

106

R. M. Khusainov and A. R. Sabirov

Fig. 1 Stiffness axes of a the tool and b the workpiece

decreases the static elastic strains while improving the dynamic stability. This solves the design problem of altering the spatial position of the stiffness axes. 2. Alter the position of the resultant cutting force. Knowing the position of stiffness axes against the coordinate axes of the machine, one can direct the resultant cutting force along the maximum-stiffness axis. This method can be implemented by altering the tool and cutting parameters. Consider the second problem for the case of stock removal by an end mill. Cutting-force components on the coordinate axes depend on the technological parameters; in this case, there are the following analytical expressions for the components Fx, Fy, Fz [20]. F x ¼ Prez sin x; F y ¼ Prez cos x; F z ¼ Prez sin x0

ð5Þ

where Prez is the resultant cutting force, x0 is the end mill helical-line elevation angle, x¼

p þ q0  b 4

ð6Þ

Stiffness Maximization on the Basis of Layout …

107

where q0 ¼

tgðb  cÞ tgðb  cÞ þ 2

ð7Þ

where pffiffiffi  1:08 4 gEa1 b þ cos c  sin c p b ¼  arctg 2 cos c þ sin c

ð8Þ

where g¼

1  l21 1  l22 þ E1 E2

ð9Þ

E1, E2 are the moduli of elasticity of the machined and the tool material; l1, l2 are Poisson’s ratios; c is the mean rake angle of the tool (edge rounding taken into account); a, b is the chip thickness and the cutting width. Each projection can be expressed in the cutting parameters or the tool parameters. Finding the exact cutting force is unnecessary for solving this problem. It can even be assumed to equal one. The idea is to find the ratios of the resultant cutting-force projections onto the coordinate axes, i.e., set such components of the vector n that the resultant force must be directed along. In this case “(2)”, the minimum strain is e2; thus, the components of the vector n will be written as follows: n1 :¼ 0; n2 :¼ 1; n3 :¼ 0

ð10Þ

The most convenient way to solve the second problem is to use the inclination angle of the helical mill groove x0, the cutting depth a, and the cutting width b as the controlled parameters. Find x0 and b by using the expressions for Fy, Fz from the “(4)” as well as the found values of these components from the “(10)”:   p tgðb  cÞ þ  b ¼ Fy cos 4 tgðb  cÞ þ 2

ð11Þ

sin x0 ¼ Fz

ð12Þ

Solving this equation by numerical methods, e.g., by the Levenberg–Marquardt algorithm, obtain x0 and b. For this system layout, the value x0 equals 46.4140. Round this angle down to the nearest value used in the tool manufacturer’s designs. Use the value b to find the ratio of the values a and b from the “(8)”:

108

R. M. Khusainov and A. R. Sabirov

pffiffiffi  1; 08 4 gEa1 b þ cos c  sin c p ¼b  arctg cos c þ sin c 2

ð13Þ

This equation can be solved by the same method. According to this solution, for this specific cutting setting the ratio must be a : b ¼ 1:283 : 0:658. Static elastic strains affect the machining accuracy in finishing milling the most. In case of such finish, the cutting depth depends on the stock to be removed. The width can be computed accordingly. The cutting width for this system must be approximately half the depth.

3 Practical Significance We can therefore propose the following method to control the elastic displacements in cutting by setting the machining parameters: 1. Finite-element modeling of the technological system consisting of the elastic machine system, the tool fixture, the tool, and the part, with strains being taken into account. 2. Finding the directions of the stiffness axes “(3)”. 3. Setting the direction of the resultant cutting-force unit vector along the maximum-stiffness axis and finding its projections onto the coordinate axes of the machine “(10)”. 4. Generating analytical dependencies of the cutting-force projections on the cutting parameters or the tool parameters “(4)”–“(8)”. 5. Solve these equations to find the necessary technological parameters (x0, a and b in this case) to ensure that the cutting force is directed as required “(11)”, “(12)”.

4 Conclusion The stiffness of a technological system depends not only on the eigen- and contact stiffness of its elements, but also on their spatial positioning, which is described by the axes of rigidity. Finite-element modeling can find the positions of such axes in the machine workspace. Approximating the resultant cutting force to the maximum-stiffness axis reduces the strain and the vibrations in the system. Knowing where the axes of stiffness are positioned, one can configure such cutting parameters that bring the resultant cutting force as close as possible to the maximum-stiffness axis. The second application of this research is solving the design problem of making such layout that brings the maximum-stiffness axis as close as possible to the resultant cutting force.

Stiffness Maximization on the Basis of Layout …

109

References 1. Sokolovsky AP (1946) Stiffness in mechanical engineering (Zhestkost v tekhnologiyakh mashinostroyeniya). Mashgiz, Moscow 2. Kasjanov SV, Kondrashov AG, Safarov DT (2017) Regulation of geometrical parameters deviations of automotive components parts through diagnostic measurements organization. Procedia Eng 206:1508–1514 3. Irzaev GH (2015) Model of control of products manufacturability in an industrial enterprise. Econ Manag Syst control 15:50–57 4. Anikeyeva OV (2017) The analysis of the experience of modern national scientific schools in the field of a metal cutting machines precision Russian Internet. J Ind Eng 5–4:14–29 5. Govorkov AS, Akhatov RH (2011) Analysis of the technological properties of aviation products based on the information image of the product. Review of Samara Scientific Center of the Russian Academy of Sciences 6-1:285–292 6. Balabanov IP, Simonova LA, Balabanova ON (2015) Systematization of accuracy indices variance when modelling the forming external cylindrical turning process. IOP Conf Ser: Mater Sci Eng 86:012010 7. Ziyatdinov RR (2015) The choice of equipment for automation of hazardous production facilities. IOP Conf Ser. Mater Sci Eng 86:012027 8. Lavrentyieva M, Govorkov A (2017) Using a discrete product model to determine the design element junctures. MATEC Web of Conf 129(1):03003 9. Khomyakov VS, Dosko SI, Zuoyi Liu (1988) Identification of elastic machine systems by modal analysis (Identifikatsiya uprugikh sistem stankov na osnove modalnogo analiza). Russ Eng Res 7:11–14 10. Law M, Altintas Y, Phani AS (2013) Rapid evaluation and optimization of machine tools with position-dependent stability. Int J Mach Tools Manuf 68:81–90 11. Whalley R, Ebrahimi M (2000) Analysis, modeling and simulation of stiffness in machine tool drives. Comput Ind Eng 38:93–105 12. Timoshenko SP, Dzh Gud’er (1975) Theory of elasticity. Nauka, Moscow 13. Kudinov VA (1967) Dynamics of machine tools. Mashinostroenie, Moscow 14. Rudakov KN (2011) FEMAP 10.2.0. Geometric and finite-element modeling of structures (Geometricheskoye i konechno-elementnoye modelirovaniye konstruktsy). KPI, Kiev 15. Goncharov PS, Artamonov IA, Khalitov TF et al (2012) NX Advanced simulation. Engineering analysis. DMK Press, Moscow 16. Ryabov EA, Yurasov SYu, Khisamutdinov RM (2016) Internal contour processing by trochoidal milling with end mills. Innovative engineering technologies, equipment and materials—2016 (ISTK “IETEM-2016”). Kazan, p 150 17. Grechishnikov VA, Petukhov YE, Pivkin PM et al (2017) Trochoidal slot milling. Russ Eng Res 37(9):821–823 18. Golovko AN, Kondrashov AG, Yurasov SYu (2017) Improved design of a worm type instrument for final machining of evolvent gear teeth. Procedia Engineering 206:1333–1336 19. Kolesnikov KS, Aleksandrov DA, Astashev VK et all (1994) Dynamics and strength of machines. Theory of mechanisms and machines. Mashinostroenie, Moscow 20. Gruby SV, Zaytsev AM (2013) Research of end mills during milling of body parts made of aluminum alloys. Sci Educ 12:31–54

Formation Automation of Geometric Configuration of Real Machine Parts V. E. Lelyukhin, O. V. Kolesnikova and O. M. Ponkratova

Abstract The automation of design and technological preparation of production with the integration of CAD systems and technological CAD, capable on the basis of 3D models of parts, plays an important role. The article deals with the issues related to the formal description of the geometric configuration of real machine parts. Two problems are not solvable within the existing geometries and CAD systems. The first problem is related to the possibility of describing only ideal objects. The second problem consists of the possibility of describing only existing geometric configurations without taking into account the schemes and technologies of their generation. This excludes the possibility of formal verification and evaluation of the correspondence of the part parameters to functional requirements, which ultimately leads to incorrect or insufficient representations of parts (objects) in the drawings and 3D models. To solve these problems, the authors propose an original approach. First, it is proposed to consider the geometric configuration of a part in the form of a closed subspace bounded by a totality of elementary infinitely extended surfaces. Secondly, instead of the traditional Cartesian coordinate system, it is proposed to use a space with coordinates corresponding to six degrees of freedom. The proposed forms of representation of the geometric configuration of real objects and their transformation methods make it possible to simply represent the shape and location of surfaces, as well as the structure of their relationships when configuring objects. In addition, the proposed concept considers and describes technologies for generating any non-ideal objects in three-dimensional space.










Keywords Engineering Machining simulation Machine tools Technological process Design technology Geometry Dimensional relationships










V. E. Lelyukhin (&)
O. V. Kolesnikova
O. M. Ponkratova Far Eastern Federal University, 8, Sukhanov St., Vladivostok 690091, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_13

111

112

V. E. Lelyukhin et al.

1 Introduction Automation of machine-building production is an integral part of the formation of digital production (CIM), which involves information integration of the entire production preparation system and directly the production process itself. The application of industrial robots in machine building production intends the use not only of a set of technical means, but also of a set of software systems from robot control to goal-supposed systems. The information basis for the formation of control influences is the description of the correct representation of the design of parts and the technological processes of their manufacture. The automation of design and technological preparation of production with the integration of CAD systems and technological CAD, capable on the basis of 3D models of parts plays an important role [1–6]. The main problem in this way is the automatic design of the technological processes for manufacturing the parts-assembly units based on 3D design models. The authors believe that the problem lies, first of all, in the impossibility of describing real objects and details by means of the geometry used today.

2 Integration Issues of CAD–CAM Systems The main problem of integrating CAD–CAM systems is that in modern CAD systems there is not enough information to build a control program for processing a part of a CAM system. To resolve all emerging issues, human intervention is necessary. The essence of human intervention is the formation of the technological process of manufacturing parts. CAPP, relying on the geometric model of the part that is created in the CAD system, compares the manufacturing technologies available in the database with the production equipment at the plant and develops technological sheets or routes necessary for manufacturing the part [2, 5, 7]. There are two approaches to the development and use of CAPP systems: variant and generative. In the case of a variant approach, the task of the CAPP system is to find a typical or similar technological process in the database and present it for modification. The generative approach consists of recognizing and isolating typical structural elements in a part and then applying typical technological operations [1, 8]. It should be noted that in this case there is no question of developing a new technological process for manufacturing a part. CAPP system allows only to select typical technological processes or operations and to compose a similar technological process. The authors propose to consider the possibility of generating a single technological process based on the 3D model of the CAD system.

Formation Automation of Geometric Configuration of Real Machine …

113

In the Russian and foreign engineering industries, part design systems, including CAD/CAM systems, are based on modern mathematics achievements, including projective, descriptive, and differential geometry. However, numerous attempts to formalize the design of technological processes using two and three-dimensional representation of the geometric configuration of the part do not yet yield tangible results [8–10].

3 Problems in Describing the Geometry of Real Parts It is impossible to achieve absolute dimensional accuracy in the manufacture of real parts in engineering production. Therefore, for any manufactured object, the limits of the allowed errors in the shape and relative position of the surfaces are given, shown in the drawings in the form of tolerance fields [8, 10, 11]. The size tolerance is defined as the difference between the largest and the smallest permissible limit values of the geometric parameter. Figure 1a presents the location and dimensioning for three ideal planes. For real planes, as shown in Fig. 1b, when the surface 2 is positioned relative to the surface 1, the actual dimension A1 can take values between the left and right boundaries (indicated by dashed lines). Similarly, the size of A3 is shown (Fig. 1b) [12]. It is characterized that in order to obtain the required arrangement of the three planes with respect to each other, one of them must be chosen as the initial one, relative to which the second plane will be oriented, and so on (Fig. 1b, c). In view of this, for the mutual orientation of the three surfaces, only two dimensions (constituting the dimensional chain) must be “held.” The third dimension is obtained as the result and is called a closed component. It should be noted that the possible deviations of the closing size are always equal to the sum of permissible deviations of the component sizes. The first problem of describing the geometry of real details is that the tools existing in geometry operate with ideal objects and not describe real (non-ideal) objects.

Fig. 1 Assignment of the position of planes in space: a ideal; b and c the real

114

V. E. Lelyukhin et al.

Fig. 2 Mutual arrangement of planes: a ideal; b the real

As an example of a comparison of ideal and real objects, we considered the arrangement of parallel planes 1, 2, and 3 as shown in Fig. 2. Figure 2a presents the projection of the “ideal” model of part. The given sizes describe strictly single-valued (ideal) mutual arrangement of planes. And the sum of the sizes A and B is equal to the size of the size C. Thus, planes 1 and 2 are at a distance A from each other, planes 2 and 3 at a distance B, and 1 and 3 at a distance C = A + B, respectively [9]. In real detail (Fig. 2b), the dimensions are not exact. The second problem is related to the impossibility of describing geometric configurations with the presentation of schemes (technologies) for their generation. This leads to the emergence of a variety of manufacturing technologies, some of which may not correspond to the functional requirements imposed on the manufactured part. As an example, we can consider the process of manufacturing the part shown in Fig. 2b. When the size A is maintained, its value will be in the interval between Amin and Amax. A similar situation is obtained when the size B is maintained. Then the size of C will be between Amin + Bmin and Amax + Bmax. If the dimensions A and B are kept in the process of manufacturing with an accuracy of ±0.1 mm, the error in the dimension of C will be equal to ±0.2 mm. But another way of manufacturing this part can be considered, which requires keeping the dimensions A and C. Then the analogous error of these dimensions ±0.1 mm will lead to the fact that the error of the remaining size B will be within ±0.2 mm.

4 Description of Objects in the Six-Dimensional Space of Degrees of Freedom To solve the above problems, the authors propose to use six degrees of freedom as a basis for describing geometric objects in three-dimensional space [9, 13]. In three-dimensional space, the body has six degrees of freedom. Authors in their works [9, 12, 13] suggest using six degrees of freedom as a basis for describing the position of the body in space. The coordinates of the vector V {Xl, Yl, Zl, Xa, Ya, Za}

Formation Automation of Geometric Configuration of Real Machine …

115

Fig. 3 Representation of surfaces in three-dimensional space: a cylindrical; b flat; c spherical

determine the presence or absence of freedom. The state of complete freedom of an element in three-dimensional space is described by the vector Va{0, 0, 0, 0, 0, 0} and the uniquely determined position −Vb{1, 1, 1, 1, 1, 1} [9, 14]. The overwhelming majority of details in mechanical engineering can be represented by a combination of some simple elements, such as a cylinder, a cube, a ball. From this point of view, the detail of any configuration can be represented as a closed subspace bounded by the composition of a particular set of elementary infinitely extended surfaces (flat, sphere, cylinder) (Fig. 3). An elementary surface is a two-dimensional topological manifold formed using a binary combination of elementary motions. Such motions are rectilinear displacement along the coordinate axis of the three-dimensional space (l), and rotation around the coordinate axis (a). This representation of the surface completely corresponds to subtractive technologies with the use of mechanical processing, since the main movements of machine tools are moving (straight) and rotating (circumference). In Fig. 3, the elementary surfaces (cylindrical, plane, and spherical) and the reflection of their position in the Cartesian coordinate system in the form of six-celled tables are shown. The tables represent a six-dimensional vector V in which the units denote the fixed degrees of freedom, and the zeros denote the residual degrees. Using the proposed representation of elementary surfaces, it becomes possible to describe not only their relative location, but also a variety of possible schemes for generating these surfaces by means of translational and rotational movements of the instrument.

5 Variants of Schemes for Generating Elementary Surfaces In accordance with the kinematic method, the surface formation consists of moving one line (generatrix) along with the other (directing) [15].

116

V. E. Lelyukhin et al.

Fig. 4 Possible variants of the formation of the cylinder

The above representation of the surface in the form of a six-cell table makes it possible to obtain all the existing combinations of pairs of lines for the formation of a surface. To obtain the production lines, it is sufficient to invert the values in the six-cell table. For example, inverting the vector V{1, 0, 1, 1, 0, 1}, which describes the position of the cylindrical surface, allows us to obtain a vector ¬V{0, 1, 0, 0, 1, 0}. In the inverted vector, the unit values have the coordinates Yl and Ya. The combination of these lines (Yl, Ya) makes it possible to obtain a cylindrical surface by possible variants (Fig. 4). This means that this cylinder can be obtained by moving a straight line parallel to the Y-axis along a circle about the Y-axis or vice versa by moving a circle perpendicular to the Y-axis along a straight line corresponding to the Y-axis.

6 Determination of the Mutual Arrangement of Elementary Surfaces The mutual arrangement of surfaces is determined by a system of fixed dimensions and tolerances. Dimensional chains are used to solve the problem of ensuring accuracy in the manufacture of the parts [16, 17]. To formalize the information about the configuration of the part and the established dimensions, the authors profited by the tools of graph theory. The geometric configuration of a part can be represented as a graph G(S, R), where the set of vertices S: = {s1, s2, …, sn} represents the set of surfaces, and the set of edges R: = {r1, r2, …, rn} corresponds to the set of connections between surfaces. The connections between the surfaces indicated in the graph can be used to represent dimensional chains. As shown in Fig. 2a, three surfaces are connected by three dimensions (Fig. 5a). In accordance with the choice of the closing link, three different dimensional chains can be constructed (Fig. 5b–d). Depending on the

Formation Automation of Geometric Configuration of Real Machine …

117

Fig. 5 Representation of the structure of the location of the planes of the part (Fig. 2): a ideal; b, c, d real variants of dimensional chains

functional purpose of the part, one or another variant of the dimensional chain is selected. It should be noted that the dimensional chain in terms of graphs is a spanning tree. Considering the necessity of constructing linear and angular dimensional chains between the surfaces of a part, the authors propose to represent them in the form of graphs of dimensional relationships for every of the six degrees of freedom. The graphs of the dimensional relationships of the geometric object shown in Fig. 6a are represented in Fig. 6b [18]. In Fig. 6b, the linear dimensional relationships in the direction of the X-, Y-, Zaxes, as well as angular dimensional relationships reflecting the rotations of the part surfaces around the X-, Y-, Z-axes are shown. The construction of graphs of dimensional constraints in all six directions Xl, Yl, Zl, Xa, Ya, and Za allows describing the relative positioning of all parts of the workpiece, as well as generating schemes for generating the geometric configuration of the workpiece.

Fig. 6 Detail with coordinated dimensions (a) and corresponding graphs of dimensional constraints (b)

118

V. E. Lelyukhin et al.

7 Schemes for Generating the Geometric Configuration of a Part Any scheme for generating the geometric configuration of a part can be formed from graphs of dimensional constraints in every of the six degrees of freedom. Depending on the indicated sizes and, accordingly, the constructed graphs of dimensional constraints, various schemes of a generation of the geometric configuration of the part can be generated [17, 19, 20]. For example, the graphs of the dimensional constraints depicted in Fig. 5b–d allow the development of three methods for manufacturing the part (Fig. 2b). Depending on the size relationships established, the following schemes can be proposed: • Surface 1 is manufactured, from it, according to a given size, then surface 2 is manufactured, and then, from surface 2 to surface 3; • The surface 3 is manufactured, the surface 2 is formed from it, and then, from the surface 2 to the surface 1; • The manufacturing starts from surface 2, from which the surfaces 1 and 3 are formed. The examples considered show that using sets of graphs of dimensional relationships in all six directions, it is possible to automatically generate all possible ways of generating the geometric configuration of a real part, taking into account the requirements for mutual arrangement of surfaces [17, 18].

8 Conclusion The arguments and examples given in the article show that for a correct and synonymous representation of nonideal geometric objects, to which the details of machines relate, it is expediently to use the space of six degrees of freedom. Based on the results of the research conducted by the authors, it can be concluded that there is a unique correspondence between the set of possible methods for machining parts on metal cutting machines and a set of schemes for generating elementary surfaces. It is also established that the geometric configuration of the mutual arrangement of surfaces and the scheme of its generation is determined by the correctness of dimensional chains. The developments presented by the authors can be used to formalize the process of designing the technology for manufacturing parts. This, in turn, will overcome the gap between CAD and CAM systems, which will expand the possibilities of digital production and application of robotic systems. Acknowledgements Work is performed with financial support of the Ministry of Education and Science of the Russian Federation on the Government contract No. 02.G25.31.0173.

Formation Automation of Geometric Configuration of Real Machine …

119

References 1. Yusof Y, Latif K (2014) Survey on computer-aided process planning. Int J Adv Manuf Technol 75:77. https://doi.org/10.1007/s00170-014-6073-3 2. Halevi G (2014) Industrial management-control and profit. Lecture Notes in Management and Industrial Engineering 1. Springer International Publishing Switzerland. https://doi.org/10. 1007/978-3-319-03470-6_1 3. Koenig DT (1990) Computer-integrated manufacturing: theory and practice. Taylor and Francis, Boca Raton 4. Li WD, Ong SK, Nee AYC, Li WD, Ong SK, Nee AYC (2004) Optimization of process plans using a constraint-based tabu search approach. Int J Prod Res 42(10):1955–1985. https://doi. org/10.1080/00207540310001652897 5. Lian K, Zhang C, Shao X, Gao L (2011) Optimization of process planning with various flexibilities using an imperialist competitive algorithm. Springer, London 6. Ma GH, Zhang YF, Nee AYC (2000) A simulated annealing-based optimization algorithm for process planning. Int J Prod Res 38(12):2671–2687 7. Skande A, Roucoules L, Klein Meyer JS (2008) Design and manufacturing interface modelling for manufacturing processes selection and knowledge synthesis in design. Int J Adv Manuf Technol 37:443. https://doi.org/10.1007/s00170-007-1003-2 8. Villeneuve F, Mathieu L (eds) (2010) Geometric tolerancing of products. Library of Congress Cataloging-in-Publication Data. Wiley-ISTE, 400 p 9. Lelyukhin VE, Kolesnikova OV, Kuzminova TA (2015) Antonenkova TV Formal representation of parts for the automated development of manufacturing technology. Sci Mag High Technol Eng Mech Eng 11:32–36 10. Tolérancement modal De la Métrologie vers les Spécifications (2009) Laboratoire Systémes et Matériaux pour la MEcatronique d’Annecy et du Centre Technique de l’industrie du Décolletage a Cluses, 192 p 11. Desrochers A (2011) Methodologie de conversion des specifications geometriques de tolerance en zones d’incertitude. Sherbrooke (Québec), Canada 12. Lelyukhin VE, Kuzminova TA, Kolesnikova OV (2015) Influence of geometrical configuration details on the technology of its manufacture. Modern Sci Res Innov 7. http://web. snauka.ru/issues/2015/07/56318 13. Lelyukhin VE, Antonenkova TV, Kolesnikova OV (2016) Structural-parametric representation of the relative position of surfaces part. Sci Mag Vestnik Eng School FEFU 1:3–9 14. Lelyukhin V (2015) The theory of synthesis of methods for shaping the surfaces of a part. Publishing house: LAP LAMBERT Academic Publishing, Germany, 80 p 15. Gorelskaya LV, Kostryukov AV, Pavlov SI (2008) Descriptive geometry: tutorial. OSU, Orenburg, 85 p 16. Matveev VV (1980) The dimensional analysis of technological processes. Mechanical Engineering, Moscow, 592 p 17. Tsitsiashvili GSh, Lelukhin VE, Kolesnikova OV, Osipova MA (2015) Definition of tolerance size in some coordinate direction. Appl Math Sci 9(136):6783–6790 18. Lelyukhin VE, Kolesnikova OV, Ignatev FY (2018) Research of influence of structure of the dimensional relations on processing of the parts. Materials of the international scientific-practical conference “Actual issues of contemporary science”. FECIT, Vladivostok, pp 3–27 19. Thimm G, Britton GA, Fok SC (2004) A graph theoretic approach linking design dimensioning and process planning Part 1: designing to process planning. Int J Adv Manuf Technol 24(3–4):261–271. https://doi.org/10.1007/s00170-003-1683-1 20. Legoff O, Hascoet J-Y (2010) Technological form defects identification using discrete cosine transform method. Int J Adv Manuf Technol 52–58

Information Model for Machine’s Electronic Structure Storage A. Loginov

Abstract The article formulates the reasons for the use of a formalized electronic structure of a product at a machine-building enterprise. The article reported similarities and differences of design, technological, and production structure of the engineering product. It described in short the basic requirements for the composition of the product according to GOST ESKD and according to the practical experience. Based on the formulated requirements, a model for storing of a machine-building product’s composition in the form of a directed acyclic graph is proposed. In this case, the nodes of this graph are the elements of the engineering product, and the edges are the facts of the entry of one element of the product into another. The article presents the permissible types of elements that make up a model of a machine-building product: an assembly unit, a assembly part, a material, a purchased item, and a billet. For each type, it is introduced a set of attributes required to fully describe an element of this type. It is given the matrix of available entrance occurrences of various elements types into each other. It is shown how one can imagine the variable composition of the engineering product with the help of the proposed model. The purposed model can be used at machine-building enterprises in PDM and ERP systems of its own design. The main features of the presented model are simplicity, full compliance with ESKD/ESTD requirements. Keywords Information model


Electronic structure
BOM
Where-used list

1 Introduction The processes of planning and production management in a modern machine-building enterprise are carried out with the help of various information systems [1, 2]. Such information systems should use formalized data about all

A. Loginov (&) JSC Novator Design Bureau, 18, Kosmonavtov, Ekaterinburg 620018, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_14

121

122

A. Loginov

aspects of manufacture: from contracts, orders, suppliers to materials, and components used in production. Electronic structure of the manufactured product plays an important role in the planning of production [3–5]. In the simplest case, the machine’s structure can be represented as a linear list of materials and components required for the production. When development and production are going according to ESKD/ESTD standards, then “Component list” [6] and “Materials list” [7] are used. In non-ESKD/ESTD practice, these two documents are combined into a single document called BOM (the Bill of Materials) [8–10]. However, this simple linear list is not suitable for accurate planning and production management. For proper organization of production, it is necessary to have a complete hierarchical structure of the manufactured product. According to [11], the structure of an engineering product is a combination of the component parts of the product and the connections between them, which determine the hierarchy of the component parts. The use of the component parts of the product in the structure of the product and/or its component parts is called component entry. There are several types of the machine-building product structure: • engineering structure—how the product is seen by the engineers who designed it; as a rule, this structure is easily formed from the design specifications • technological structure is the design structure, with addition of some technological features for manufacture [12, 13]. When forming materials usage lists, the technological structure takes into account materials from technological requirements, extra amount of materials for processing on specific equipment, and a number of other factors • manufactured structure—it is usually the structure of a specific instance of the machine, produced on enterprise. For this type of structure, in the case of a variant structure, ones should make decisions on the use of one of the options at all levels. There should be no alternative materials, components, production methods here. The use of manufactured structure in ERP systems allows you to carry out production planning and management most effectively. It follows that the task of properly representing the structure of the product is very relevant when building an enterprise ERP system.

2 Research Problem An information model in computer science is the presentation of concepts, relationships, constraints, rules, and operations, designed to define the semantics of data for a specific problem area [13].

Information Model for Machine’s Electronic Structure Storage

123

The information model for the structure of a machine-building product should: • reflect the structure of the product • store the parameters of the product composite parts. Moreover, the information model for structure of machine-building product should be [14]: • • • •

formalized complete open for modifications no-duplication. Requirements that are more specific include:

• full presentation of variant information (use of various substitutable components) • support of assembling versions (according to GOST 2.113-75 [15]) • support of complex quantities • support of storage information about assemblies and parts produced on the basis of others components • support of composite materials, the composition of which can be represented in mass fractions.

3 Information Model of Structure of Machine-Building Product The structure of the machine-building product is most accurately reflected by the directed acyclic graph [16], i.e. such a directed graph, in which there are no directed cycles, but there may be “parallel” paths, leaving the same node and coming to the destination node in different ways. The representation of electronic structure as the hierarchy of the components is presented in Fig. 1. From Fig. 1, it can be seen that each part in the model has a set of attributes describing it. In addition, it has a collection of relations to the “parent” elements (through the “Included in” relations) and a collection of links to its “descendants” (through the “Contains” relations). Each relation in this collection represents the parent–child relationship and has the set of specific attributes.

124

A. Loginov

Fig. 1 Components and relations

3.1

Types of the Parts in the Model

Various types of components are used in the machine-building product [17]. For the information model, let us highlight the following main types of component parts: • • • • •

assembly unit part purchased item material billet.

At the same time, all purchased items and materials are elements of a MDM [18, 19]. As a rule, at the machine-building enterprise, there is a separate system of MDM data, which stores information about the standard items and materials produced according to GOST and OST. Elements of the type “Material” and “Purchased item” are imported the “Name” attribute from the MDM system. In addition, they are imported a number of other attributes, while the connection with the basic system containing the MDM master data is maintained by key. Table 1 shows the matrix of allowable entrance of basic types into each other. If at the intersection there is the word “Yes,” then this means that the element with the type indicated in the row may include elements of the type indicated in the column. From Table 1, it can be seen that behavior of “Purchased item” is the same as behavior “Assembly Unit.” Only other materials may be included in the material. This allows you to store the structure of composite materials in the information model [20]. In this case, as a rule, the quantities of materials are indicated in mass fractions or as a percentage.

Information Model for Machine’s Electronic Structure Storage

125

Table 1 Entrance permission matrix Assembly unit Part Purchased item Material Billet

Assembly unit

Part

Purchased item

Material

Billet

Yes – Yes – –

Yes

Yes Yes Yes – –

Yes Yes Yes Yes Yes

– Yes – – –

Yes – –

An element of type “Part” may include an element of type “Billet.” It stores information about the attributes of the billet, necessary for the production of this part. In addition, the item may include materials and purchased items. This is necessary to indicate auxiliary materials for the production of this part. If several billets are included in the part, this is a case of variation. On the relations going to these billets, there is information about variants numbers. Only one or several materials are included in the billet. This allows to store the information about the materials from which this part is made. There are two cases: • all materials included in the billet are used to produce parts • when the parts are manufactured, it is used only one of the materials (sets of materials), and the remaining materials (sets) are alternatives.

3.2

Specific Attributes of Component Parts

The attribute set of component parts mostly repeats the set of the columns from the design specification. The text attribute “Format” contains information about the formats of the design document (according to the specifications). This attribute specifies a complete list of formats, without transferring a part of the list to “Notes” column. For “Materials” and “Purchased items,” this attribute is not filled. The attribute “Name” stores the name of the component, according to the specification or another document. Text attribute “Code” indicates the designation of the component according to the specification. This attribute is not filled for “Materials,” as they have no specific codes (only names). The “Code” attribute has strong validation rules to prevent any errors during data entry. Most organizations have regulatory procedures for the formation of these validation rules. Correct input of this attribute is important due to the fact that it is often used as a key attribute of an entity. In addition, this attribute is often used for making connection with other information systems.

126

3.3

A. Loginov

Specific Attributes of Relations

The attribute “Quantity” indicates the amount in which the part is included into parent component. This number can be both integer or fractional. At the same time, the quantity may have a unit of measurement. Examples of quantities are “15”, “2.5 kg”, “3 m.” Instead of a unit of measurement, the concept of mass fractions or percent can be used. This is used for saving information about multicomponent materials. In one level, you can use quantities in percentage or quantities in mass fractions (mixing of them is not allowed). Information model has the concept of a complex quantity, i.e. such a quantity, which is indicated no amount of something for one upper component, but for some number of them. For example, the quantity “3 kg/15 assemblies” for some material means entering that it is necessary to take three kilograms of this material on the 15 higher assemblies. Using such quantities in the information model adds more flexibility. As a rule, such a representation of quantity is used by technological divisions and production workshops. In addition, in the information model has the concept of quantities with an alternative. For example, if a cable can be measured in meters (m) or kilograms (kg), then its quantity can be represented, for example, like “0.5 kg (10 m).” This means that for the manufacture of the product you need a piece of cable with a length of 10 m and a mass of 0.5 kg. The use of quantities with this alternative allows you to combine a look at the quantity from different departments of enterprise. For example, technologists can operate with lengths of cable production, while the purchasing department can receive information on the weight of purchased materials. The attribute “Norm,” from the relation, allows technologists to indicate the exact amount of materials used. In “Norm” complex quantities and quantities with an alternative can be used. The norm may either be equal to amount of material, or be different from it. The differences between “Quantity” and “Norms” are due to the consumption of material for various technological operations: cutting, etc. The attribute “Norma” is set by technologists after the creation of a technological process to produce a part or assembly unit. “Variant number” and “Variant” are very important attributes of the relation. The integer “Variant Number” defines the sequence number of an interchangeable component within a parent component. Variants can be separated on special groups by using the “Variant.” Figure 2 shows an example that “Assembly 1” can be assembled either using “Subassembly 1” or using the whole group of parts “Part 1,” “Part 2.” At the same time, all relations connecting the top-level assembly with its subcomponents refer to the same element of the “Variant” type.

Information Model for Machine’s Electronic Structure Storage

127

Fig. 2 Interchangeable components

For the complete presentation of the information contained in the products design specifications [1], each relation contains: • the attribute “Position,” that indicates the ordinal numbers of the component parts that are directly included in the product being specified, in the sequence of recording them in the specification; • attribute “Zone,” defines the zone of the component in the drawing (by breaking the drawing field into zones according to GOST 2.104); • attribute “Specification Section” indicates the specification section (according to GOST 2.106-96) of the component (this attribute is not filled when for multicomponent materials and complex purchased products). If the relation is used to store information about the any parts not from the ESKD specification, then the attributes “Zone” and “Position” are not filled. In addition, the relation contains information about other features of the component required for the organization of the design and technological preparation of the production: • the sign “Additional material” is active when the component is an additional material that is not directly included into the product specification but is included in product’s composition and is necessary for production (e.g., paints and some coatings). As a rule, such materials are indicated in the technical requirements of the drawing, according to GOST 2.316-2008 [21] • the sign “Auxiliary material” is active, if the component is necessary for the production (according to the technological process), but not directly included into the product (e.g., gloves for the worker, etc.). The presence of these additional signs allows to generate reports on the materials and components in accordance with the ESKD/ESTD, as well as reports for the economic departments that prepare reports the basic and auxiliary materials [22].

128

3.4

A. Loginov

Common Attributes

As any component, the element of the type “Relation” has the attribute “Comment.” This attribute contains non-formalized information of various kinds. At the same time, if the note is entered for the component, then this common is applied to all places where this component is included. If a comment is specified for a relation, then it is valid only for this specific entry. The presence of the attribute “Comment” allows you to save unformalized information from the specification or technological processes on the components and relations. In addition to the comments, the component parts and relations contain the “Route” attribute, which indicates the route card, according to [23] for the itinerary or route-operational description of technological processes. Moreover, if the route is specified for the component, then the route will be common for all applications of this component within the product. If the route is specified for relation, then it replaces the standard route of the component part, and this route will be used for all calculations.

4 Conclusions The proposed information model allows ones to: • formally describe the design and (partially) technological composition of the engineering product according to ESKD/ESTD; • eliminate duplication of information at all levels; • quickly get information of the composition of a particular assembly unit, and the composition all assembly units to the unlimited depth; • quickly receive information on the entrance and applicability of the components of the machine-building production. The model can be used for the machine’s structure information storage in PDM/ PLM/ERP systems. This model, with minimal changes, has been successfully used since 2006 for storing information in the system named “Composition of Products 2.0” in the Novator Design Bureau [24]. Now it stores information about more than 500 products. Many of these products have more than 100 thousand components.

References 1. Thomas DJ (2011) Production planning and control with SAP® ERP. Galileo Press, Boston 2. Polnomoshnova OM (2017) Decisions in strategic planning with the help of economic and mathematical modeling at a high-tech enterprise. Prod Manag 25(2):68–78

Information Model for Machine’s Electronic Structure Storage

129

3. Koblov NN (2013) Creation of unified informational infrastructure. Bull Sib Sci 1(7):62–68 4. Kulga KC (2014) Using of an integrated automated information system to manage the resources of an aviation plant. News Samara Sci Center Russ Acad Sci 16/1(5):1448–1455 5. Mikhailov VG (2016) Approaches to the creation of an integrated information system PDM-ERP. Inf Proc Decis Making 2:17–24 6. GOST 2.106-96. Unified system for design documentation. Text documents—Introduced 01.07.1997 7. GOST 3.1123-84. Unified system for technological documentation. Forms and rules of technological documents used for materials usage norms—Introduced 01.01.1986 8. Shaun S (2012) The bill of materials in excel, ERP, Planning and PLM/BMMS Software. SCM Focus 9. Bill of Materials for Dummies—ETO. https://virtualdutchman.com/2010/01/19/bill-ofmaterials-for-dummies-eto/. Accessed 20 Nov 2018 10. Creating a Bill of Materials. https://www.arenasolutions.com/resources/articles/creating-billof-materials/. Accessed 20 Nov 2018 11. GOST 2.053-2013. Unified system for design documentation. Product electronic structure. Basics—Introduced 01.06.2014 12. Veprev V, Strelyaev S, Bushkov S (2010) Formation of the technological composition of the product and the development of technological processes along the route of manufacturing technological versions of the part. CAD Graph 1:81–83 13. Abrosimov D, Skopintsev A, Chernobyl G (2014) Technological structure—the basis of technological preparation of production and production management. Clever manufacture. http://www.umpro.ru/index.php?page_id=17&art_id_1=540&group_id_4=93&m_id_4=29. Accessed 20 Nov 2018 14. Tina Lee Y (1999) Information modeling from design to implementation. National Institute of Standards and Technology 15. GOST 2.113-76. Unified system for design documentation. Group and basic design documents—Introduced 01.07.1976 16. Oystein O (2009) Theory of graphs. American Mathematical Society 17. Stolyarov AI, Donetskaya YuV, Gatchin YuA (2017) Creating electronic structure of instrument making equipment. Sci Tech J Inf Technol Mech Opt 17(2):312–317 18. Chernikov BV (2015) Information analysis of MDM systems. News of South Federal University. Tech Sci 2(163):70–81 19. Borodin AM, Mirvoda SG, Porshnev SV (2013) Method of managing MDM information in autonomous information systems. Eng News Don 4(27):76 20. Kerber ML (2008) Polymer composition materials. Structure. Properties. Technologies. Profession, Saint Petersburg 21. GOST 2.316-2008. Unified system for design documentation. Legend, technical rules for graphical documents—Introduced 01.07.2009 22. Auxiliary materials. https://profmeter.com.ua/Encyclopedia/detail.php?ID=851. Accessed 20 Nov 2018 23. GOST 3.1102-2011. Unified system for technological documentation. Stages of development and types of documents—Introduced 01.01.2012 24. Loginov AY, Volman DV, Rozenbaum AE (2012) Organization of a unified information space for managing design and technological information about products based on proprietary information systems. News South-Urals State Univ Comp Sci Manag Radio Electron Syst 3 (262):24–26

Software Spindle Speed Variation as Method for Chatter Suppression in Drilling V. M. Svinin, A. V. Savilov and A. V. Shutenkov

Abstract The spindle speed variation (SSV) method for chatter suppression in drilling was studied. The existing chatter suppression methods used at various types of metalworking, including the drilling process, were analyzed. The spindle speed variation method was described. The experimental setup, which consists of the measuring equipment and the cutting tool, was described. Software and constructive ways for implementation of the SSV method were analyzed. The experimental data on CNC machine capabilities to vary spindle speed rates were presented. SSV method implementation constraints for a specific machine were analyzed. The recommendations for the choice of modulation parameters were given. Modal analysis results for a technological system were presented. Constant spindle speed cutting and variable spindle speed cutting were compared. The effects of spindle speed variation parameters on the chatter amplitude and surface roughness were studied. A spindle speed variation depth ensuring a high-quality surface finish was determined. The dependence of machined sample displacement vibration on a spindle speed variation depth was determined. The experimental results were interpreted. Keywords Drilling


Chatter
Suppression
Speed variation

1 Introduction Spiral drills are often used for machining workpieces. Due to a low rigidity of the tool, drilling is often accompanied by regenerative chatter. Modern high-rate equipment and tools aggravate the problem, as they operate at high-speed rates which intensify chatter vibrations. Along with degradation of surface quality, tool, and machine life, vibrations can cause failure of drills to be extracted from the hole.

V. M. Svinin (&)
A. V. Savilov
A. V. Shutenkov Irkutsk National Research Technical University, 83, Lermontov St., Irkutsk 664074, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_15

131

132

V. M. Svinin et al.

When it is impossible to achieve target accuracy and surface roughness due to chatter vibrations, sinking, reaming, and boring methods increasing production costs are used. In edge cutting machining, chatter regeneration is due to the internal resonance between current chatter vibrations of the technological system and chatter mark vibrations on the cutting surface [1]. Various methods have been developed to suppress them [2]. However, only the cutting speed variation method can be used for drilling which decreases performance. Some researches dealt with drilling modes ensuring cutting stability due to resonance suppression [3–6]. The approaches involve preliminary cutting simulation using standard software products or special high-level language programs. Some authors suggest conducting full-scale experiments to determine material machinability [7]. On the one hand, this method increases simulation reliability. On the other hand, it increases labor intensity. One of the new methods for chatter suppression is spindle speed variation (SSV) [8, 9]. The method involves regular variation of spindle speed in relation to the average value at specific cutting depths and frequencies. The SSV method can be implemented in a constructive way or using software. Constructive implementation involves development of mechanical devices installed into the machine spindle, chuck, or cutting tool [10, 11]. This method is mainly used in turning and milling. In drilling, the method increases dimensions of instrumental adjustments which negatively affects hole drilling in hard-to-reach zones. Software implementation of the SSV method involves using built-in SSV functions in the CNC control systems of some modern processing centers [12] or developing special subroutines which are used in control programs. The issues of practical implementation of the SSV method in turning and milling were analyzed in a number of works [13–17]. Some researchers described SSV as a process which compensates for cutting resistance in order to achieve high machining speeds in milling [14]. Other authors described SSV as spindle speed variation using the Chebyshev method [16]. The method was based on vibration prediction at variable spindle speeds. The software varied spindle speed rates described in [17, 18] in a sinusoidal manner. Similar to previous researches, the method was applied only in milling. The following conclusions can be drawn from the brief analysis of the previous researches. Most chatter suppression methods developed for different types of machining cannot be used for drilling. The SSV method is an exception. In involves selecting spindle speed parameters ensuring chatter-free machining and using variable spindle speed rates extending the range of chatter-free conditions. Efficiency of the SSV method for turning and milling was theoretically and experimentally verified. However, for other types of machining, including drilling, efficiency issues have not been studied yet. The issue of selection of cutting speed variation parameters with regard to a spindle lag remains understudied. Therefore, the purpose of the article is to study CNC machine capabilities to vary cutting speed rates and suppress chatter.

Software Spindle Speed Variation as Method for Chatter …

133

2 Materials and Methods For the experiments, a DMG DMC 635 V vertical milling machine (Fig. 1) was used. Due to the lack of a built-in SSV function, the cutting speed was varied using a special software utility developed in [18]. This program changes spindle speed using a timer. Holes with a diameter of 9 mm were drilled in a workpiece 105  145  50 mm in size (Steel 35 with Hardness 170) to a depth of 8 mm at a feed rate of 0.2 mm/rev. Lubricant coolant was not applied. A Sandvik R840-0900-50-A1A carbide drill was used as a cutting tool. To assess machine capabilities by actual depths of spindle speed variation at its allowable power load of 50%, dry experiments were conducted at spindle speed rates of 800, 1200, 1600, 2000, 3000, 4000, and 5000 rpm with depths varying from 0 to 5% in 1% increments, and relative frequency variation per spindle revolution: 0.5, 1.0, and 1.5. The actual spindle speed variation depth was tracked using the servo trace option. Before carrying out the experiments, modal analysis of the technological system was carried out by the method described in [19]. Based on its results, a diagram of dynamically stable cutting areas at a constant speed rate was constructed (Fig. 2). According to this diagram, an adverse spindle speed rate of 3892 rpm corresponding to the beginning of chatter was selected for the experiments. This spindle speed was used in the main series of experiments on chatter suppression by varying cutting speed rates from 0 to 1% in 0.2% increments and 12% in 1% increments. 17 holes were drilled. In accordance with recommendations developed in [20], relative frequency of spindle speed variation was assumed to be fmod = 1. To record workpiece displacement vibrations, an AP85-100 piezoelectric contact vibrator was used. Surface roughness was measured with a Form Talysurf 200 Profilometer.

Fig. 1 Experimental unit

134

V. M. Svinin et al.

Fig. 2 Dynamic stability diagram

3 Results and Discussion Figures 3 and 4 show selected results (for spindle speed rates of 800 and 4000 rpm in the investigated range from 800 to 5000 rpm) of preliminary experiments on actual cutting speed variation depths at different relative frequencies. Due to the high inertia of the main drive, the machine can fully reproduce a desired spindle speed variation depth only at low rates of its angular spindle speed and variation frequency. For example, at a spindle speed rate of n = 800 rpm and a

Fig. 3 Dependence of the actual spindle speed variation depth on the nominal spindle speed modulation depth at relative frequencies in 800 rpm: a 0.5 mod/rev; b 1 mod/rev; c 1.5 mod/rev

Software Spindle Speed Variation as Method for Chatter …

135

Fig. 4 Dependence of the actual spindle speed variation depth on the nominal spindle speed modulation depth at relative frequencies in 4000 rpm: a 0.5 mod/rev; b 1 mod/rev; c 1.5 mod/rev

relative variation frequency fmod = 0.5 the reproduced modulation depth is 30%, and at fmod = 1.0, the reproduced variation depth is only 12%. If the nominal spindle speed variation depth is set above these values, it is not fully reproduced by the machine. For example, at a spindle speed rate n = 4000 rpm, to reproduce a 3.5% actual spindle speed variation depth at fmod = 0.5, the nominal variation depth should be 12%. However, it increases temperature and power spindle loading. The higher the spindle speed and speed variation frequency, the smaller is the spindle speed variation depth reproduced by the machine. All the results obtained in the preliminary series of experiments are summarized in Fig. 5. The area of the diagram is divided into five zones according to the ability of the machine to reproduce desired depths of spindle speed variation. If angular spindle speed parameters correspond to zone 1, the desired depth is fully reproduced. In zone 2, depth reproduction is 50%. In zone 3, depth reproduction is 10% of the speed variation depth, and in zone 4, it is only 3%. Zone 5 is characterized by unstable and incorrect reproduction of spindle speed variation parameters. The diagram facilitates selection of cutting speed variation parameters for chatter suppression in drilling. Figure 6 shows photographs of surfaces taken in the main series of experiments with different spindle speed variation depths. The photographs show that hole machining at a constant spindle speed rate (see Fig. 6a) or at a small speed variation depth (see Fig. 6b) is accompanied by vibrations. The best surface quality corresponds to a 3% speed variation depth (see Fig. 6d). An increase in the speed variation depth up to 5% (see Fig. 6e) and 12% (see Fig. 6f) degrades the surface finish quality. Combined graphics of the influence of spindle speed variation depth on the chatter amplitude and surface roughness presented in Fig. 7 verify the results. In constant cutting speed drilling, the chatter amplitude was 19 lm and surface roughness—Ra = 1.5 lm. Cutting speed variation reduced sharply the chatter amplitude and the surface roughness. At a speed variation depth of 0.6%,

136

V. M. Svinin et al.

Fig. 5 Zones of actual spindle speed variation depths

Fig. 6 Photographs of surfaces machined at different cutting speed variation depths: a 0%; b 0.4%; c 0.6%; d 3%; e 5%; f 12%

the chatter amplitude is reduced to 14 lm, and surface roughness is reduced to Ra = 0.93 lm, i.e., 1.5-fold. At a 3% speed variation depth, the chatter amplitude was 11 lm, and roughness Ra = 0.86 lm. A further increase in the speed variation depth was accompanied by regular chatter intensification and relaxation and surface finish quality degradation. Controversial influence of cutting speed variation on the

Software Spindle Speed Variation as Method for Chatter …

137

Fig. 7 Influence of the cutting speed variation depth on: a chatter amplitude; b surface roughness

chatter amplitude and surface roughness is consistent with modern views on the physical nature of regenerative chatter vibrations [21]. Their energy source is a phase shift relative to the chatter of a vibration mark on the cutting surface spontaneously setting at about 90° at a constant speed rate. Changes in the harmonic cutting speed variation disrupt this phase shift making it gradually change from 90° to a smaller value over the variation period. The regenerative effect decreases the amount of energy entering the chattering system and chatter intensity. With an increase in the spindle speed variation depth, the range of changes in the phase shift increases and decreases regularly which suppresses and intensifies chatter vibrations.

4 Conclusion The following conclusions can be drawn from the study. • The experiments proved that cutting speed variation can suppress chatter and reduce surface roughness in drilling. Variable spindle speed cutting at fvar = 1.0 and a 3% depth increased the chatter amplitude and reduced hole surface roughness Ra twice. • The spindle lag prevents it from reproducing target cutting speed variation frequency and depth in the whole range of technological parameters. For practical application of the SSV method in the machine-building industry, diagrams of zones of actual spindle speed variation parameters should be constructed. • Spindle speed variation is a promising method for chatter suppression, machining, and parts performance improvement.

138

V. M. Svinin et al.

• The chatter suppression method requires further research on speed variation parameters for different cutting modes, dimensions, and materials of workpieces.

References 1. Altintas Y (2012) Manufacturing automation: metal cutting mechanics, machine tool vibrations. Cambridge University Press, Cambridge 2. Quintana G, Ciurana J (2011) Chatter in machining processes: a review. Int J Mach Tool Manuf 51:363–376 3. Roukema J, Altintas Y (2007) Generalized modeling of drilling vibrations, Part I. Time domain model of drilling kinematics, dynamics and hole formation. Int J Mach Tools Manuf 47(9):1455–1473. https://doi.org/10.1016/j.ijmachtools.2006.10.005 4. Roukema J, Altintas Y (2007) Generalized modeling of drilling vibrations. Part II. Chatter stability in frequency domain. Int J Mach Tools Manuf 47/9:1474–1485. https://doi.org/10. 1016/j.ijmachtools.2006.10.006 5. Ahmadi K, Altintas Y (2013) Stability of lateral, torsional and axial vibrations in drilling. Int J Mach Tools Manuf 68:63–74. https://doi.org/10.1016/j.ijmachtools.2013.01.006 6. Huang BW (2005) The drilling vibration behavior of a twisted microdrill. J Manuf Sci Eng Trans ASME 126(4):719–726. https://doi.org/10.1115/1.1813472 7. Ahmadi K, Savilov A (2015) Modeling the mechanics and dynamics of arbitrary edge drills. Int J Mach Tools Manuf 89:208–220. https://doi.org/10.1016/j.ijmachtools.2014.11.012 8. Stoferle T, Grab H (1972) Vermeiden von Ratterschwingungendurch Periodische Drehzahlanderung. Werkstatt und Betrieb 105:727–730 9. Inamura T, Sata T (1974) Stability analysis of cutting under varying spindlespeed. Ann CIRP 23(1):119–120 10. Svinin VM, Peshkova EA, Rybas AA (2015) Investigation of the process of turning non-rigid shafts using a cutting speed modulation device. In: Collection of articles of the XV International Scientific Practical Conference: in 3 parts. Trans-Baikal State University, Russian Federation, Chita, pp 49–54 11. SINUMERIK828D, 840Dsl: Avoiding vibration in slewing gear pieces. https://support. industry.siemens.com. Accessed 18 Nov 2018 12. Chiappini E, Tirelli S, Albertelli P, Strano M, Monno M (2014) On the mechanics of chip formation in Ti–6Al–4V turning with spindle speed variation. Int J Mach Tools Manuf 77:16– 26. https://doi.org/10.1016/j.ijmachtools.2013.10.006 13. Ishibashi T, Fujimoto H, Ishii S, Yamamoto K, Terada Y (2014) High-frequency-variation speed control of spindle motor for chatter vibration suppression in NC machine tools. In: American control conference (ACC) Portland, Oregon, USA, 4–6 July 2014. https://doi.org/ 10.1109/acc.2014.6858869 14. Otto A, Radons G (2013) Application of spindle speed variation for chatter suppression in turning. CIRP J Manufact Sci Technol 6:102–109. https://doi.org/10.1016/j.cirpj.2013.02.002 15. Totis G, Albertelli P, Sortino M, Monno M (2014) Efficient evaluation of process stability in milling with spindle speed variation by using the Chebyshev collocation method. J Sound Vib 333:646–668. https://doi.org/10.1016/j.jsv.2013.09.043 16. Urbikain G, Olvera D, López de Lacalle LN, Elías-Zúñiga A (2016) Spindle speed variation technique in turning operations: modeling and real implementation. J Sound Vib 335:1–13. https://doi.org/10.1016/j.jsv.2016.07.033 17. Vnukov YN, Natalchishin VV, Germashev AI, Kuchugurov MV, Dyadya SI (2014) The use of modulation of the speed of the main movement of the milling machine for damping

Software Spindle Speed Variation as Method for Chatter …

18.

19. 20. 21.

139

regenerative self-oscillations when milling thin-walled parts. Bull NTUU “KPI” Ser Mach-Build 3(72):1–217 Kuchugurov MV, Germashev AI, Dyadya SI, Pirozhok AV (2015) Features of the implementation of the drive control capabilities of machines based on the rack CNC Siemens. High Technol Mach Equip 1(25):80–87 Savilov AV, Pyatykh AS, Timofeev SA (2013) Optimization of machining processes based on modal and dynamometric analysis. Sci Technol Ind 1(2):42–46 Svinin VM (2006) Selection of cutting speed variation parameters for regenerative chatter suppression. Vestnik of Samara State Technical University. Tech Sci 41:135–142 Svinin VA (2011) Regenerative chatter suppression in milling. Cutting speed variation. LAP LAMBERT Academic Publishing GmbH & Co. KG, Saarbrucken (Germany)

Experimental Studies of Steady-State Sources of Vibrations of Machinery Production Process Equipment to Substantiate Choice of Vibration Protection Methods S. I. Gvozdkova and L. E. Shvartsburg

Abstract Vibration is one of the factors accompanying virtually every production process and largely characterizing its quality in terms of its impact on the environment and humans. The adverse effects of industrial vibrations predetermine a set of measures to reduce vibration. These measures are taken by the developers of structures and technologies, as well as experts in the field of occupational safety and labor comfort and environmental protection. The main sources of industrial vibration are considered here. The main factors contributing to the formation of mechanical oscillations in the components of process equipment kinematic chains are vibration effects, including steady-state, nonsteady, random effects. The experimental studies of steady-state sources of the formation of production vibration were carried out. The aim of the study was to determine the relationship between the change in the level of imbalance and the vibration level in frequencies ranging. The reason for the imbalance of rotating parts is the displacement of the center of mass relative to the axis of rotation. The results of the experimental study of determining the relationship between the change in the level of the imbalance and vibration level in frequencies ranging are presented.




Keywords Vibration Sources of vibration production Process equipment





Vibration protection
Machinery

S. I. Gvozdkova (&)
L. E. Shvartsburg Moscow State University of Technology “STANKIN”, 1, Vadkovsky, Moscow 127055, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_16

141

142

S. I. Gvozdkova and L. E. Shvartsburg

1 Introduction Vibration is one of the factors accompanying virtually any production process and largely characterizing its quality in terms of its impact on the environment and humans [1–4]. The implementation of waste minimization, including vibration, is one of the most important directions of the modern production development [5–7]. Vibration has a negative impact on physiological state of a person. Industrial vibration is a mechanical oscillation of machines, mechanisms, and their components. Adverse effects of industrial vibration predetermine a set of measures to reduce vibration [8–10]. These measures are taken by process and structure developers, as well as by HSE experts [11].

2 Formulation of the Problem Methods of protection against industrial vibration are methods that reduce vibration at the source of its formation and in the path of its propagation [1, 12–14]. Vibration reduction at the source is based on the improvement of kinematic chains of the production system process equipment. The sources of mechanical oscillations of components are based on the contact of components of a mechanical system as a result of action of variable forces, such as friction forces, inertial and impact forces, as well as dynamic forces resulting from manufacturing inaccuracy [15, 16]. Some of the most common components of the kinematic chains of process equipment are gear trains. Vibrations of gear trains are a result of oscillations of the wheels and kinematic chain components associated with them. The causes of mechanical oscillations are the mutual collision of the teeth at the time of engagement, the variable deformation of the teeth as a result of the variability of the forces applied to them, the kinematic errors of the gear wheels, and the variable friction forces. Vibration reduction begins with an analysis of the operating conditions of the gear train. A lack or insufficient amount of lubricants contributes to increase in friction. Therefore, vibration reduction is achieved by using the optimal amount of lubricants. Damping capacity of lubricants increases with viscosity. Some of the reasons for the formation of vibrations in the process of the gear train operation are manufacturing errors, including cyclic errors and profile manufacturing errors. Therefore, the achievement of the required accuracy of gear wheels is of great importance.

Experimental Studies of Steady-State Sources of Vibrations …

143

3 Main Part The main factors contributing to the formation of mechanical oscillations in the components of process equipment kinematic chains are vibration effects, including kinematic and force effects [17–19]. Force effects are characterized by the time functions of the component forces F(t) or torques M(t) that act on the object [20, 21]. Kinematic effects are characterized by acceleration of the oscillation source points a(t), their speeds V(t), and displacements S(t). Figure 1 shows the classification of the main sources of the formation of vibration effects. Vibration effects: • Steady-state • Nonsteady • Random. Let us consider the harmonic process as a kind of steady-state vibration effect [22]. The periodic processes described by the following formula are harmonic (1): xðtÞ ¼ A
sinðx0 t þ /Þ;

ð1Þ

where A is the amplitude, x is the angular frequency, and / is the phase. The sources of harmonic effects are unbalanced parts of the mechanisms of process equipment of the production system, which move rotationally or progressively [23, 24]. Unbalance is the state of the parts of the mechanisms, characterized by such mass distribution, which during rotation causes variable loads on supports of the parts and their bending. Imbalance is a degree of the system unbalance. The reason for the imbalance of rotating parts is the displacement of the center of mass against the axis of rotation. The emergence of imbalance during rotation is a result of asymmetry in distribution of the rotating masses, heterogeneous density of the material. As a result of these reasons, an unbalanced centrifugal force occurs, which is proportional to the mass of the system, the square of the angular velocity, and the

Fig. 1 Classification of the main sources of the formation of vibration effects

144

S. I. Gvozdkova and L. E. Shvartsburg

eccentricity of the system mass center against the axis of rotation. The centrifugal force, which is the disturbing force that leads to run-outs, is determined by the formula (2): F ¼ e
m
x2 ;

ð2Þ

where F is the centrifugal force, m is the mass, x is the angular velocity, and e is the eccentricity. Polyharmonic processes are oscillatory processes. Their sources are n sources of vibration effects independent from each other (3): xð t Þ ¼

n X

Ai
sinðxi0 t þ /i Þ;

ð3Þ

i¼1

where Ai is the amplitude, xi is the angular frequency, and / is the phase. The set of harmonic frequencies is the frequency spectrum, and the set of numbers Ai (i = 1, 2,…, n) is the amplitude spectrum of vibration effects. Run-outs are processes resulting from addition of two harmonic oscillations with close periods. If we assume that the amplitudes and initial phases of these oscillations are the same, and the angular (cyclic) frequencies are close to each other, then (4): xðtÞ ¼ A
sin x1 t þ A
sin x2 t ¼ A sin

x1 þ x2 x1  x2 t
cos t 2 2

ð4Þ

Thus, the movement takes the form of sinusoidal with a long period compared with the period of main movement. The steady-state vibration effects also include shock processes. Shock mechanical effects are short-term effects. The shock shape is a function expressing dependence of force, torque, or acceleration during shock on time. The main characteristics of the shock shape are duration of the shock and its amplitude, that is, the maximum value of mechanical effect at the time of shock. As a result of shock processes in a relatively small period of time, the tool momentum is transferred to the workpiece; that is, during collision with the workpiece the kinetic energy of the tool is converted into the energy of elastic oscillations and partially into thermal energy. During shock processes, the law of conservation of momentum is fulfilled (5): P¼

n X

mi
Vi ;

ð5Þ

1

where mi is the mass, Vi is the velocity, and n is the number of interacting elements. Nonsteady vibration effects are caused by the transient processes occurring in the sources. Random vibration effects are generally not predictable, unlike harmonic and polyharmonic effects. For example, vibrations due to the roughness of friction pairs of the process equipment mechanisms can be attributed to random ones. These vibrations are stochastic in nature. It is difficult to describe them using

Experimental Studies of Steady-State Sources of Vibrations …

145

regular functions. The sources of random vibration effects in the components of the mechanisms of production equipment include frictional oscillations, which are formed due to the action of friction forces.

4 Research Results Experimental studies of steady-state sources of the formation of production vibration were carried out. Figure 2 shows a scheme of the experimental installation. The main units of the installation are a disk, three-phase motor, power supply system, and casing. Installation is fixed to the table plane at four points with clamps. The specially manufactured disk has drilled holes designed for mounting test loads of various weights. The disk has eight holes drilled for each of three mounting positions with radii R1, R2,R3. In the course of experimental studies, the system imbalance was analyzed as the main steady-state source of production vibration. The disk is an unbalanced system. The reason for the imbalance of rotating parts is the displacement of the center of mass against the axis of rotation. As a result of these reasons, an unbalanced centrifugal force occurs, which is proportional to the mass of the system, the square of the angular velocity, and the eccentricity of the system mass center against the axis of rotation. The aim of the study was to determine the relationship between the change in the level of imbalance by attaching test loads of different weights and the vibration level in the frequency range. Experimental studies were carried out using a three-axis vibration meter Oktava-101VM. For measurements, one radial straight line and three mounting radii (R1 < R2 < R3) of five test loads of various weights (M1 < M2 < M3 < M4 < M5) were chosen. The level of vibration was estimated by the level of vibration acceleration La, dB along three axes (OX, OY, and OZ).

Fig. 2 Scheme of the experimental installation: a front view; b top view: 1—body, 2—disk, 3—sensor, 4—motor

146

S. I. Gvozdkova and L. E. Shvartsburg

Fig. 3 Experimental dependence of the vibration acceleration La, dB along OZ-axis on the level of imbalance determined by attaching test loads of five different weights along the mounting radius R1 in the frequency range

Figure 3 shows the experimental dependence of the vibration acceleration along OZ-axis on the level of imbalance determined by attaching test loads of five different weights along the mounting radius R1. Figure 4 shows the experimental dependence of imbalance level, as well as vibration acceleration along OZ-axis determined by attaching a test load of M4 weight for three attachment positions in the frequency range. Experimental studies have shown that increase in the weight of test loads in a particular attachment position contributes to increase in the system imbalance and is

Fig. 4 Experimental dependence of imbalance level, as well as vibration acceleration La, dB along OX-axis determined by attaching a test load of M4 weight for three attachment positions in frequency range

Experimental Studies of Steady-State Sources of Vibrations …

147

characterized by increase in the vibration level. This dependence takes place in the entire frequency range, as well as along three axes and for each radial position of attachment of test loads. Also, the experimental study results indicate that a change in the position of attachment of a specific test load along radial straight line consistently on R1, R2, R3 is characterized by increase in the level of imbalance, which was confirmed by increase in the level of vibration acceleration. The cause for increase in the imbalance level was increase in the eccentricity of the system mass center against the axis of rotation due to change in the radial position of attachment of the test loads.

5 Conclusion Vibration is one of the factors accompanying virtually any production process and largely characterizing its quality in terms of its impact on the environment and humans. Methods of protection against industrial vibration are methods that reduce vibration at the source of its formation and in the path of its propagation. The sources of mechanical oscillations of components are based on the contact of components of a mechanical system as a result of action of variable forces, such as friction forces, inertial and impact forces, as well as dynamic forces resulting from manufacturing inaccuracy. Experimental studies of steady-state sources of the formation of production vibration were carried out. Experimental studies have shown that increase in the weight of test loads in a particular attachment position contributes to increase in the system imbalance and is characterized by increase in the vibration level. This dependence takes place in the entire frequency range, as well as along three axes and for each radial position of attachment of test loads. Also, the experimental study results indicate that a change in the position of attachment of a specific test load along radial straight line consistently on R1, R2, R3 is characterized by increase in the level of imbalance, which was confirmed by increase in the level of vibration acceleration. The cause for increase in the imbalance level was increase in the eccentricity of the system mass center against the axis of rotation due to change in the radial position of attachment of the test loads. The results of experimental studies provide prerequisites for the subsequent analysis and justification of the choice of vibration protection methods for process equipment of machinery production.

148

S. I. Gvozdkova and L. E. Shvartsburg

References 1. Shvartsburg LE (2008) Human and environmental protective ensuring of automated engineering. VESTNIK MSTU “STANKIN” 3(3):19–21 2. Shvartsburg LE, Ivanova NA, Ryabov SA, Gvozdkova SI, Zmieva KA (2012) Automation of maintenance of indicators safety for machine-building technologies formation of the form. Sci Pract Educ-Methodical J Life Safety S2:1–24 3. Ryabov SA, Ivanova NA, Shvartsburg LE (2014) Assessment, analysis and managing occupational risks in the industry. Chief Mech Engineer 12:21–26 4. Bukeihanov NR (2009) Renovation of technological processes—tool of resourcekeeping and improving of ecological security. VESTNIK MSTU “STANKIN” 4(8):21–24 5. Shvartsburg LE (2008) Environmental ensuring of forming technology. VESTNIK MSTU “STANKIN” 1:38–43 6. Rodriguez PE, Shvartsburg LE, Artemyeva MS (2017) Methodological design and commissioning of an experimental stand for the study of the spread of harmful substances in the air of work areas during the processing of metals in industry. Procedia Eng 206:588– 593. https://doi.org/10.1016/j.proeng.2017.10.521 7. Shvartsburg L (2015) Ecoenergetics of cutting manufacturing processes. Ecol Ind Russ 19 (3):4–9. https://doi.org/10.18412/1816-0395-2015-3-4-9 8. Bukeihanov NR, Cmir IM, Hairo DA, Sergeev VN (2010) Heuristic methods of modernization of machine-building enterprises manufactures. VESTNIK MSTU “STANKIN” 3:75–79 9. Bukeihanov NR, Cmir IM (2008) Innovative approaches to solving resource problems in engineering. VESTNIK MSTU “STANKIN” 4(4):161–166 10. Bukeihanov NR (2008) Control of innovative resource projects. VESTNIK MSTU “STANKIN” 3(3):66–70 11. Gvozdkova SI, Shvartsburg LE (2012) Reduce of loss of energy by increase the coefficient of power. VESTNIK MSTU “STANKIN” 2(20):32–36 12. Shvartsburg LE, Zvenigorodskij UG, Bukeihanov NR (2001) Methodology of resource saving projects development. VESTNIK MSTU “STANKIN” 2(14):14–17 13. Zmieva KA, Shvartsburg LE (2009) Automated energy and resource saving systems for industrial enterprises. Ecol Ind Russ 11:7 14. Shvartsburg LE (2008) Human and environmental protective ensuring of automated engineering. VESTNIK MSTU “STANKIN” 3(3):19–21 15. Shvartsburg LE, Markin AV (2012) A study of the influence of geometric cutting tool in shaping vibrations in the cutting zone. Sci Pract Educ-Methodical J Life Safety 2:30–32 16. Shvartsburg LE, Butrimova EV, Drozdova NV (2012) Experimental research of distribution vibroacoustics factors in the environment for forecasting of their levels in the certain point of space. Sci Pract Educ-Methodical J Life Safety 2:27–30 17. Egorov SB, Kapitanov AV, Mitrofanov VG, Shvartsburg LE, Ivanova NA, Ryabov SA (2016) Formation of the integral ecological quality index of the technological processes in machine building based on their energy efficiency. Int J Environ Sci Educ 11(11):4065–4078 18. Shvartsburg LE, Butrimova EV, Yagolnitser OV (2017) Energy efficiency and ecological safety of technological processes of form-shaping. Procedia Eng 206:1009–1014. https://doi. org/10.1016/j.proeng.2017.10.586 19. Shvartsburg LE, Butrimova EV, Yagolnitser OV (2017) Quantitative evaluation of the effectiveness of best available technologies of form-shaping. MATEC Web Conf 129(01027). https://doi.org/10.1051/matecconf/201712901027 20. Shvartsburg LE, Butrimova EV, Drozdova NV (2014) Development of an algorithm for automated prediction of vibration and noise in the technological environment. VESTNIK MSTU “STANKIN” 4(31):187–190 21. Ivanova NA, Ryabov SA, Shvartsburg LE (2016) The role of information technology in rotor balancing. Russ Eng Res 36(3):235–238. https://doi.org/10.3103/S1068798X16030096

Experimental Studies of Steady-State Sources of Vibrations …

149

22. Gvozdkova S (2015) Analysis of provision methods of environmental safety by minimization of energy losses by the example of industrial vibration and noise. Ecol Ind Russ 19(3):14–17. https://doi.org/10.18412/1816-0395-2015-3-14-17 23. Gvozdkova SI (2008) Analysis of the functional relationship between the parameters of production noise and vibration. VESTNIK MSTU “STANKIN” 3(3):35–40 24. Gvozdkova SI, Shvartsburg LE (2017) Analysis of sources and methods for reducing noise by minimizing vibrations of engineering technological processes. Procedia Eng 206:958–964. https://doi.org/10.1016/j.proeng.2017.10.578

On Issue of Verifying New Method for Studying Dynamics of Deep Hole Machining L. Mironova, L. Kondratenko and V. Terekhov

Abstract The article outlines the principles for constructing a mathematical model for the study of dynamic phenomena of deep holes machining. A theoretical justification is given for using the differential equation of the angular momentum for torsional vibrations. A solution of a second-order partial differential equation using the Laplace integral transform method is presented. Two ordinary differential equations are introduced, which sufficiently describe the relationship between the angular acceleration and the gradient of the change in tangential stress in the rod and the rate of change in voltage with the gradient of the angular velocity of motion. These equations allow us to calculate the frequency characteristics of the drive as applied to the technology of deep hole machining using the boring and trepanning association (BTA) method. The correctness of the mathematical formulations of the new research method is justified by the verification of the solutions obtained using classical calculation methods and models.
















Keywords Deep hole Drill Rotary drive Rod Motion speed Shear stress Laplace transform Torsional vibrations







1 Introduction Deep hole machining with a blade tool, drills, reamers, etc., should be attributed to the most complex metalworking technological operations.

L. Mironova (&)
V. Terekhov National Research Nuclear University MEPhI (Moscow Engineering Physics Institute), 31, Kashirskoe Sh., Moscow 115409, Russia e-mail: [email protected] L. Kondratenko State Science Center RF of TsNIITMash, 4, Sharikopodshipnikovskaya Str., Moscow 115088, Russia © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_17

151

152

L. Mironova et al.

As the cutting part of the tool is significantly removed from the machine spindle during the drilling process, various adverse effects due to dynamic phenomena are observed. From a variety of adverse factors, it is necessary to highlight uneven spindle rotation, variable elasticity of the line, change in torque and cutting force, precession oscillation of the axis of rotation, and longitudinal oscillation of the tool cutting part. These circumstances result in degradation of manufacturing accuracy, deviations of the shape of the surfaces due to the formation of faces on the surface of the hole, to significant deviations of the axis of the hole from the theoretical axis, etc. [1–3]. In addition, there are such problems as basing the support elements of the cutting tool head and removal of chips from the cutting zone. The task of studying dynamic phenomena in relation to the deep hole machining is similar to the process of drilling wells on oil and gas and is described in [4]. However, the two processes differ by the physical–mechanical properties of the materials used, the cutting tool and the surface layer being processed. Therefore, creation of a mathematical model for the study of dynamic phenomena in deep hole machining operations is urgent for the development of innovative metalworking technologies. The following approach is proposed to study this problem.

1.1

The Essence of Deep Hole Machining

In recent years, the BTA method is used in deep hole machining in large heat and power apparatuses (steam generators, various heaters, etc.) as shown in Fig. 1. This operation is usually carried out at special machines. The tool with cutting part 1, fixed by thread on drill pipe 2, is installed at work spindle 3, which performs both rotation and translational motion, not necessarily interrelated. Before start-up, the movable machine support stand is positioned so that the axis of aligning bushing 4 coincides with the point of machined part 6 where the hole will be made.

Fig. 1 Deep hole machining scheme: 1—cutting part with cutting and support plates, 2—drilling pipe, 3—work spindle, 4—aligning bushing, 5—machine support stand, 6—machined part

On Issue of Verifying New Method for Studying Dynamics …

153

After that, the supply of coolant, the spindle rotation, and its axial flow s0 are switched on in series. Coolant, entering the cutting zone through the channels between the guide sleeve and the cutting head, lubricates, cools, and pushes the chips into the internal channel of the drill pipe, which then falls into the chip catcher. In principle, this way you can get deep holes of sufficiently high quality. However, in the process of drilling, various dynamic phenomena occur, which have a negative effect on both the working capacity of the tool and the quality of the hole. For the mathematical description of dynamic processes as applied to BTA, there is the problem of applying models in calculations that are more accurately adapted to the conditions of the studied processes. On the one hand, by applying more accurate mathematical models, we are forced to solve systems with a large number of equations and variables, to look for conditions for the convergence and stability of solutions. At the same time, the obtained approximate solutions may go beyond the specified accuracy and not correspond to the actual results. Therefore, in many cases it is advisable to carry out theoretical studies using classical models of continuum mechanics [5].

2 Derivation of Equations During drilling, the quality of the cutting process, besides the tool parameters, is greatly influenced by the speed of interaction of the cutting head with the part blank and the stresses in the body of the drill pipe. In mechanics, a partial differential equation is usually used to describe oscillations. @2u G @2u ¼
2 @t2 q @x

ð1Þ

where u is the angular displacement of the section; G is the shear modulus; q is the density of the material; t is time; x is the longitudinal coordinate. Ultimately, this equation relates the displacement of the input and output coordinates. However, it is more convenient to carry out studies of this process using a different method [6], where not only displacement is taken into account, but also stress, which generally depends on displacement, speed of displacement change, viscosity, etc. In this case, from the equation of the moment of momentum in differential form [7], we can obtain the equation rq

@X @s ¼ @t @x

ð2Þ

where X is the angular velocity of rotation of the rod cross section; s is the maximum tangential stress in the cross section of a rod of radius r.

154

L. Mironova et al.

In this case, we can write X¼

@u : @t

ð3Þ

Taking into account the substitution (3) in (2), formula (1) can be rewritten as rG

@2u @s ¼ : @x2 @x

ð4Þ

After integrating expression (4) over the coordinate x and differentiation over time, we get rG

@X @s ¼ : @x @t

ð5Þ

This approach allows us to establish the relationship of shear rates of cross sections of the elementary volume of a straight cylindrical rod with a change in tangential stresses. Without considering the viscous friction losses and assuming that the density and shear modulus are constant, we write Eqs. (4) and (5) in Laplace images dsðsÞ ; dx

ð6Þ

1 dXðsÞ ssðsÞ ¼  rG dx

ð7Þ

rqsXðsÞ ¼ 

where s is a complex variable. Differentiating Eq. (6) with respect to the x coordinate, using (7) eliminating the derivative dX/dx, and introducing a new variable h(s), we obtain a second-order differential equation with constant coefficients relative to the image @ 2 sð s Þ  h2 ðsÞsðsÞ ¼ 0 @x2

ð8Þ

rffiffiffiffi q hðsÞ ¼  : G

ð9Þ

where

The general solution of Eq. (8) is a form sðs; xÞ ¼ C1 exp½hðsÞx þ C2 exp½hðsÞx:

ð10Þ

On Issue of Verifying New Method for Studying Dynamics …

155

The integration constants C1, C2 are determined by the boundary conditions: x ¼ 0; sðs; xÞ ¼ s1 ðs; 0Þ; @sðs; xÞ G ¼  h2 ðsÞX1 ðs; 0Þ @x s

ð11Þ

where X1 is the angular velocity of rotation of the leading link. Substituting (11) into (8), we get   1 G s1 ðs; 0Þ  hðsÞX1 ðs; 0Þ ; 2 s   1 G C1 ¼ s1 ðs; 0Þ þ hðsÞX1 ðs; 0Þ : 2 s

C1 ¼

ð12Þ

After substituting (12) into (10) and introducing hyperbolic functions, the particular solution of the differential Eq. (8) will be sðs; xÞ ¼ s1 ðs; 0Þch½hðsÞx 

Gr hðsÞX1 ðs; 0Þsh½hðsÞx: s

ð13Þ

Similarly, we solve the system of Eqs. (6), (7) with respect to the variable X(s, x). We will get Xðs; xÞ ¼ sðs; 0Þch½hðsÞx 

1 ss1 ðs; 0Þsh½hðsÞx: GhðsÞ

ð14Þ

According to Fig. 1, instead of the x coordinate in the obtained solutions (13) and (14), one should substitute the variable l (pipe length). The coefficient h(s) is the operator coefficient of wave propagation. Assuming that the angular momentum applied to the drill pipe is completely consumed in the load, we supplement the system of these equations with boundary conditions Xðs; lÞ ¼ X2 ðsÞ; Xðs; 0Þ ¼ X1 ðsÞ; sðs; lÞ ¼ s2 ðsÞ; sðs; 0Þ ¼ s1 ðsÞ;  1  s2 ð s Þ ¼ Mr ðsÞ þ hf X2 ðsÞ þ JsX2 ðsÞ Wp2

ð15Þ

where Mr is the moment of resistance at the outlet end of the pipe of length l; X1, X2 are angular speeds of rotation of the input (leading) and output (driven) end of the pipe; s1, s2 are maximum tangential stresses at the inlet and outlet ends of the power section of the drill pipe; hf is the coefficient of friction loss; Wp2 is geometric polar

156

L. Mironova et al.

moment of resistance of the pipe section adjacent to the cutting head; J is the moment of inertia of the cutting head, which is brought to the driven end of the pipe. From the joint decision (13)–(15), we get   1 X 1 ðsÞ  Mr ðsÞ#k s ; wðsÞ ch½hðsÞl " # 1 X1 ðsÞðhf þ JsÞ   s 2 ðsÞ ¼
Mr þ Wp2 wðsÞ ch hðsÞl X2 ðsÞ ¼

ð16Þ

ð17Þ

where wðsÞ ¼ 1 þ hf #k ðsÞs þ J#k ðsÞs2 ; #k ðsÞ ¼

l Z k ðsÞ GrWp

  th hðsÞl : Zk ðsÞ ¼ hðsÞl

ð18Þ

Equations (16), (17) make it possible to calculate the frequency characteristics of the drive, i.e., determine the response of the drive to the harmonic change in the speed of the driving link or the moment of resistance acting on the head with a matched load, when the angular moment applied to the system is completely consumed in the load. A detailed exposition of this question is given in [3, 8].

3 The Study of the Frequency Characteristics of the Drive at Deep Hole Machining Consider the drive work case when X1  0. Assuming in (16)–(17) the complex variable s is: s ¼ jx;

ð19Þ

where x is the frequency of harmonic changes, and we obtain two frequency characteristics of WM(jx) and WMs(jx). These frequency characteristics are determined by the following formulas: X2 ðjxÞ #k ðjxÞ ¼ : Mr ðjxÞ 1 þ hf #k ðjxÞjx þ J#k ðjxÞðjxÞ2

ð20Þ

i sðjxÞ 1 h ¼ 1 þ hf #k ðjxÞjx þ J#k ðjxÞðjxÞ2 : Mr ðjxÞ Wp2

ð21Þ

WM ðjxÞ ¼ WMs ðjxÞ ¼

On Issue of Verifying New Method for Studying Dynamics …

157

Frequency characteristic WM(jx) in (20) shows how the moment of resistance Mr, which occurs at the output link of the pipe, influences X2—the angular velocity of rotation of the output link of the pipe. In expression (21), the frequency response WMs(jx) illustrates the effect of the moment of resistance Mr on shear stresses s, caused by the force effects during drilling. Putting Mr = 0, we obtain two more frequency characteristics WX(jx) and WXs(jx), illustrating the influence of the angular velocity of rotation of the input link X1 on the output parameters X2 and tangential stresses. Formulas have the form WX ðjxÞ ¼ WXs ðjxÞ ¼

X2 ðjxÞ 1 1 ¼
; X1 ðjxÞ chðjak Þ 1 þ hf #k ðjxÞjx þ J#k ðjxÞðjxÞ2

i sðjxÞ 1 J ðjxÞ h ¼ þ 1 þ hf #k ðjxÞjx þ J#k ðjxÞðjxÞ2 ; X1 ðjxÞ chðak Þ Wp2 rffiffiffiffi q ak ¼ lx : G

ð22Þ ð23Þ ð24Þ

From (22) to (24), it can be seen that the changes in the maximum tangential stresses and the speed of rotation of the executive body do not coincide with the changes in the input actions. Using this approach, it is possible to determine the frequency response WX(jx), illustrating the dependence of the oscillation of the angular velocity of rotation of the input link X1 with respect to the oscillations of the speed of the longitudinal movement of the cutting part of the drill. This task is described in [8]. Using the above research method, the frequency characteristics of the volume hydraulic drive were obtained, which characterize the influence of fluctuations of the resistance moment on the output shaft rotation frequency and voltage, caused by force effects [11]. Let us consider in more detail the value of function (18) under condition (19). We will get h pffiffiffiffiffiffiffiffiffii th jxl q=G th½hðjxÞl tg ak pffiffiffiffiffiffiffiffiffi ¼ Zk ðjxÞ ¼ ¼ ; hðjxÞl ak jxl q=G

ð24Þ

chðjak Þ ¼ cos ak ;

ð25Þ

#k ðjxÞ ¼ #k ðak Þ:

ð26Þ

As we see, functions (24)–(26) are not complex functions. The graph of the real function Zk is shown in Fig. 2. In the case when ak ! 0, then Zк ! 1. The function Zк ! 1 becomes negative in the intervals

158

L. Mironova et al.

Fig. 2 Dependence of the real function Zk on the dimensionless parameter ak

p=2 þ kp [ ak [ p þ kp; ðk ¼ 0; 1; . . .; nÞ: If ak ! 0, then Eq. (16) is reduced to the well-known equation describing dynamic processes in a drive with short lines X 2 ðsÞ ¼

1 ½X1 ðsÞ  Mr ðsÞ#k s; wðsÞ

ð27Þ

wðsÞ ¼ 1 þ hf #k ðsÞs þ J#k ðsÞs2 : This conclusion makes it possible to quite accurately determine in which cases the dynamics in a mechanical drive should be analyzed using equations for short or for long lines [9, 10].

4 Checking the Correctness of the Decision From Eqs. (20) and (21), we determine the amplitude of oscillations of the angular velocity of rotation of the output link. Suppose that the angular velocity of the input link is zero, the moment of resistance varies according to the harmonic law X1 ¼ 0; Mr ¼ Mra sinðwtÞ:

On Issue of Verifying New Method for Studying Dynamics …

159

Then, the final amplitude formula will take the form Mra AM ¼ rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi h i: ½1J#k ðak Þx2 2 þ h2f ½# ða Þx2 k

ð28Þ

k

This means that the maximum value of AM at Mra = const will be achieved when the following equation is fulfilled 1  J#k ðak Þx2 ¼ 0: #k ðak Þx

ð29Þ

We will show that condition (29) is similar to the rule for determining the eigenfrequencies formulated by Babakov [12]. Comparison will be carried out on the example of a cargo of mass m suspended on a rod of length l, with a cross section f, previously stretched by a force F. Under the action of a force, the load begins to move downward, while the initial speed of movement is zero. Then, the eigenfrequencies are determined from the equation b tg b ¼ a; a ¼

mr mc

ð30Þ

where b is a parameter; a is a dimensionless quantity equal to the ratio of the mass of the rod to the mass of the load. For torsional vibrations is true rffiffiffiffiffiffiffiffi Jr b ¼ lx GJc

ð31Þ

where Jr, Jc are the moments of inertia of the rod and the load, respectively. We write the formula of the moment of inertia for our case.   Jp ¼ 0:5 r 4  r04 :

ð32Þ

In view of (31) and (32), expression (30) takes the form rffiffiffiffiffiffiffiffi q b ¼ lx ¼ aj : GJc

ð33Þ

Dividing the numerator of the left side of Eq. (29) by the denominator, we get 1 ¼ Jx: #k ðak Þx

ð34Þ

160

L. Mironova et al.

We perform the following transformations: rffiffiffiffi rffiffiffiffi q q ¼ Jx; rWp lx ¼ lJx2 tg ak G G rffiffiffiffi rWp l lx tg ak lx tg ak q ak tg ak ¼ qffiffiffi ; : ¼ q G J q q

GrWp lx lx tg ak

G

Or rWp l ¼ ak tg ak : J

ð35Þ

Denote rWp l ¼ v; J and then formula (35) takes the form ak tg ak ¼ v:

ð36Þ

The quantity v is the ratio of the flywheel moment of inertia of the pipe to the flywheel moment of inertia of the load. Therefore, frequency equations (30) and (36) are similar. To determine the resonant frequencies from (22) and (23), we find the modulus of the function WM (jx), taking into account (28). We will get jWM ðjxÞj ¼ h

1

ð1#k ðak ÞJx2 Þ2 #2k ðak Þx2

þ h2f

i2 :

ð37Þ

It follows from (37) that resonance occurs when disturbing frequencies are as follows 1 xres ¼ pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi : J#k ðak Þ

ð38Þ

In the case of the absence of oscillations of the moment and the harmonic change of the rotational speed of the leading end, the amplitude of the frequency response will be determined by the relation 1 jWX ðjxÞj ¼ qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi  2ffi : ½1  J#k ðak Þx2 2 þ hf #k ðak Þx

ð39Þ

On Issue of Verifying New Method for Studying Dynamics …

161

Obviously, by substituting formula (39) into expression (38), we obtain the maximum amplitude value; i.e., in this case, the resonant (eigen)frequencies should also be determined by relation (38). At resonance, the maximum of the amplitude characteristics is as follows AM ðmaxÞ ¼

1 1 ; AXðmaxÞ ¼ : hf hf #k ðak Þxres

ð40Þ

In [3], as an example, graphs of the frequency characteristics of the deep drilling process were given. The above proofs allow to verify the calculated data using the amplitude–frequency characteristic of the boring and trepanning association (BTA) drill with the following initial data: The diameter of the drill is 16.3 mm with the length of the force section of the drill pipe 1000 mm; the outer diameter of the drill pipe is 15 mm, the wall thickness is 1.5 mm, the torsional elasticity is #0 = 4.26  10−6 [Nmm]−1, and the stroke moment of inertia of the head is J = 3.09  10−3 [Nmmsek2]; friction loss coefficient hf = 1 Nmmsek; Fig. 3. From the graph, it can be seen that the mechanical drive, which consists of a spindle, a drill pipe, a drill, periodically falls into resonant modes at frequencies xres = 4396; 13330; 22510… s−1. The number of such regimes, as in the usual elastic system, is infinite.

Fig. 3 Amplitude–frequency characteristic of the drill BTA

162

L. Mironova et al.

5 Conclusion The described techniques allow to study the fluctuations of the rotation speed of the drill head and stresses in the stem (drill pipe) of different lengths, to evaluate the influence of process parameters in deep drilling technologies, as well as to choose rational operating modes and tool design. The above test shows that the new research method adequately reflects the dynamics of the process of drilling deep holes. This method was generalized by the example of a mathematical model of a rotating rod with a fixed disk and extended to other mechanical engineering technologies [13].

References 1. Kondratenko LA (2005) Vibrations and speed regulation methods of movement of technological objects. MRSU, Moscow 2. Terechov VM, Kondratenko LA, Mironova LI (2016) Some dynamic features of interaction the cutting part of the gun drill with the detail. Eng Autom Probl 4:78–84 3. Kondratenko L, Mironova L, Terekhov V (2017) Investigation of vibrations during deepholes machining. 25th International Conference Vibroengineering, Liberec, May 2017. Vibroengineering Procedia 11:7–11. https://doi.org/10.21595/vp.2017.18285 4. Kondratenko LA (1987) Fluctuations in the bit rotation frequency while simultaneous drilling by the rotor and downhole hydraulic machines. Mach Sci 4:44–52 5. Sedov LI (1970) Continuum mechanics. Nedra, Moscow 6. Kondratenko LA, Terekhov VM, Mironova LI (2017) About one method of research torsional vibrations of the core and his application in technologies of mechanical engineering. Eng Autom Probl 1:133–137 7. Kondratenko LA (2008) Calculation of velocity variations and stresses in machine assemblies and components. Sputnik, Moscow 8. Kondratenko L, Mironova L, Terekhov V (2017) On the question of the relationship between longitudinal and torsional vibrations in the manufacture of holes in the details. 26th Conference in St. Petersburg, June 2017. Vibroengineering Procedia 12:6–11. https://doi.org/ 10.21595/vp.2017.18461 9. Kondratenko L, Dmitriev V, Mironova L (2017) Simulation of a drive with a long connecting link. 26th International Conference Vibroengineering, St. Petersburg, June 2017. Vibroengineering Procedia 12:231–236. https://doi.org/10.21595/vp.2017.18743 10. Kondratenko LA, Mironova LI (2018) Imitation of nonlinear drives with distributed parameters of power lins. Eng Autom Probl 1:92–97 11. Kondratenko L, Mironova L (2018) Features of loss of stability of the work of two-link mechanisms that have an infinite number of degrees of freedom. J Math Eng Manag Sci 3 (4):315–334 12. Babakov IM (1958) Theory of oscillations. GTTL, Moscow 13. Kondratenko L, Terekhov V, Mironova L (2016) The aspects of roll-forming process dynamics. 22nd International Conference Vibroengineering, Moscow, October 2016. Vibroengineering Procedia, 460–465

Rational Provision of Robustness Properties of Bolted Joints of Assembly with Implementation of Anaerobic Materials I. I. Voyachek, D. V. Kochetkov and S. G. Mityasov

Abstract Stress concentration in thread roots and non-uniform load distribution along the thread turn is the major disadvantages of bolted joints. Under changing cyclic loads on areas of high-stress concentration, crack formation and destruction of parts occur. The purpose of the work is to provide rational robustness properties of bolted joints by decreasing stress on thread turns and stress concentration in case of assembly with the implementation of anaerobic materials polymerizing in the contact zone. Modelling of bearing strength of bolted joints is conducted according to the finite element method, with the theory of contact interaction between mating surfaces being used. A rational length zone of bolted joints of the bolt-screw nut type where it is necessary to place an anaerobic material is set out, a necessary volume of anaerobic material is measured, some recommendations on unbolting screw joints concentration in case of assembly with implementation of anaerobic materials are given.













Keywords Bolted joints Anaerobic materials Assembly Load Robustness Finite element method




1 Introduction Bolted joints (BJ) in most cases serve as critical assembly units, which determine the robustness of the whole construction, so they must comply with exploitation requirements such as robustness (static and dynamic), stiffness, tightness and resistance to fretting, corrosion and self-loosening. The robustness of BJ, especially under the action of cyclic loads, significantly depends on load distribution along thread turns [1–22]. I. I. Voyachek (&)
D. V. Kochetkov Penza State University, 40, Krasnaya Street, Penza 440026, Russia e-mail: [email protected] S. G. Mityasov LLC “Stankomashstroy”, 9, Titov Street, Penza 440028, Russia © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_18

163

164

I. I. Voyachek et al.

In the works [1–19], it was identified that the load on traditional BJ of bolt– screw nut type increases significantly to the first thread turns by the side of load application according to the law of hyperbolic cosine. The first thread turn is exposed to 30% of applied load and even more. This leads to stress concentration and reduces static and dynamic robustness of BJ. That is why the development of ways of stress decreasing of bolted joints thread turns is extremely important. Contact interaction control of mating parts thanks to anaerobic material (AM) input into contact zone refers to effective means of providing bolted joints running abilities. Anaerobic materials represent one component polymer compositions, which can stay in liquid state for a long time and polymerize on the metal surfaces relatively fast, given there is no contact with aerial oxygen. For example, when AM gets into closed volumes, including ones in contact zone of threaded parts. After polymerizing, anaerobic material (AM) gets elastic properties and shear strength. It is known that cavity filling with an anaerobic material in bolted joints contact zone of a detail leads to considerable stress decreasing on thread turns (stress is partly taken by elastic interlayer of an anaerobic material, which connect thread turns together, making a solid heterogeneous structure) and consequently leads to a cyclic robustness growth of bolted joints [5, 14, 15, 19, 23–26]. This work investigates the problems of how to provide robustness properties of bolted joints of assembly with the implementation of anaerobic materials in the most effective way. The task of determining the rational length zone of bolted joints of bolt–screw nut type, and the necessary volume of anaerobic material is set.

2 The Study of Influence of Anaerobic Materials Over Stress Concentration in the Mating Zone of Bolted Joints by Finite Element Method Analysis of stress–strain state with the use of finite element method (FEM) starts from the creation of solid model (of assembly) of bolted joint. In order to do this, different CAD programs like SolidWorks or Nastran can be used. The solid models of BJ (with or without AM) are demonstrated in Fig. 1. The assembly of thread connections between parts with helix surfaces is complicated to fulfil. So, with a consideration of rather small lead angle thread is replaced with set of rings with cross section being equivalent to thread profile. It was established that this approximation exerts insignificant influence over the results of robustness test of BJ with stiffness test being exposed to less impact [15, 19, 24, 27]. According to Saint-Venant’s principle, to exclude the impact of fastening and decrease the volume of calculations, the length of the bolt is limited to a distance of one diameter from the ends of the screw nut. To reduce the calculation, one-sixth

Rational Provision of Robustness Properties of Bolted Joints …

165

Fig. 1 Solid models of bolted joints: a the model of bolted joint without anaerobic materials; b the model of bolted joint with anaerobic materials

part of bolted joint is chosen as computation model. It becomes possible due to symmetry of the joint. In his work [27], the author advises to choose rather small angular size of the sector to provide minimum amount of finite elements in circumferential direction. At the same time, we should consider the fact that the smaller it is, the bigger degeneracy factor of elements adjacent to the axis is, and the smaller computational stability of solution technique is provided. The feature of the task to be solved consists in contact of several deformed bodies (bolt and screw nut) and the layer of AM. When AM filled the voids along the length of thread engagement during modelling, we set the conditions for AM connection to surfaces of bolt and screw nut. The use of finite element method (FEM) lets us consider complex volume deformation of mating parts, non-uniform contact pressure, different conditions for contact interaction between mating parts and their slide along surfaces of thread turns. During the study of bolted joints by FEM, the following boundary conditions were used (Fig. 2):

Fig. 2 Computation model of bolted joint

166

I. I. Voyachek et al.

• for surfaces along which we chose one-sixth part of bolted joint condition “symmetry” was set. Under this condition, there is no movement normal to these surfaces, i.e. normal displacement for surfaces, which go through the symmetry axis, is equal to zero; • we put the limitations on the left end of screw nut, which imply the absence of movement along the axis of the joint; • uniformly distributed normal force F was applied to the left end of the bolt; • in contact zone of thread surfaces of bolt and screw nut, we applied boundary conditions in the form of “no penetration, unit to unit” and set coefficient of friction f. Modelling by finite element method was conducted with following initial conditions: bolted joint M10–6H/6g; external diameter d = D = 10 mm; internal diameter d1 = D1 = 8,647 mm; medium diameter d2 = D2 = 9,188 mm; nut width across flats D0 = 17 mm; the pitch of thread P = 1,25 mm; thread angle profile a = 60°; bolt material—steel grade 45X (E1 = 2.06  105 MPa, l1 = 0,32); and screw nut material—steel grade 35X (E2 = 2.14  105MPa, l2 = 0,29), where E1, E2 and l1, l2—moduli of elasticity and Poisson’s ratios of bolt and nut. External load F = 2 kN was applied to the bolted joint. We used AM of HM162 grade (EAM = 2.62  103 MPa—modulus of elasticity of anaerobic material, ssh
AM = 35 MPa—shear strength). Distribution of the load F along thread turns in joint of bolt–nut type was determined. Joints, when an anaerobic material is placed in cavities of contact zone with a restricted thread turn number (not on the whole length of thread engagement), were studied. The following options were considered sequentially (see Fig. 3): 1—an assembly without AM; 2—AM fills the cavities in the contact zone Fig. 3 Distribution of load along thread turns of the joint

Rational Provision of Robustness Properties of Bolted Joints …

167

of the first turn (from the load side F); 3—the first two turns; 4—the first three turns; 5—the first four turns; 6—the first five turns; 7—all six turns (on the entire length of screwing). Figure 3 shows the results of the FEM study to determine the share of axial load per thread windings. Analysis of the results obtained (see Fig. 3) shows that when AM fills the cavities of the threaded contact zone over the entire length of the bolt and nut screwing (dependence 7), the load on the first turn decreases by 32% compared to the PC without AM (dependence 1). Almost the same result is obtained when filling the AM cavities in the first three turns (dependence 4). The result is slightly worse (decrease by 30%) when the cavities of the first two turns are filled (dependence 3). Thus, for the purpose of rational use of AM, it is advisable to place in the cavities of the contact zone only the first two or three turns from the side of application of the load (no more than half of the make-up length). At the same time, the used volume of AM decreases by a factor of 2 and more, and the cost of BJ assembly operation with AM decreases.

3 Determination of the AM Volume Required to Fill the Cavities of the Threaded Contact Zone and the Moment of Unbolting the Bolted Joints One of the main tasks at all design stages of a bolted joint when assembling with AM is to consider issues related to the rational joints quality. In order to save AM, it is necessary to develop a method for determining the optimal AM volume necessary to fill completely the cavities of the threaded contact zone. At the first stage, the AM volume is determined, which is necessary for filling cavities at one thread pitch. To do this, consider the zone of the threaded contact on the length of one step (Fig. 4). It was determined [5, 15, 19] that the required volume of AM can be found by the following formula AM VPn ¼ RApd2 nk

ð1Þ

where n—the number of turns; k ¼ 1; 5. . .2; 0—coefficient taking into account the deviations of the geometry of the turns of the threaded contact, the method and irregularity of application AM; RA—the amount of shaded areas in Fig. 4, determined by the following formula RA ¼ 0:025 P2 þ 0:162PSlan tg

a a þ H1 Slan tg ; 2 2

ð2Þ

where H*—is the height of the original thread triangle (Fig. 4); H1 —the actual working height of the thread profile; Slan —landing clearance.

168

I. I. Voyachek et al.

Fig. 4 Diagram of the threaded contact zone for the one-step length: 1—a zone of the working height of the thread; 2—a zone of a thread turn; 3—a zone of a hollow of a nut turn

If the BJ must be unscrewed during operation, the choice of an anaerobic material brand and the area of its application should be made according to shear strength of the bolt rod when unscrewing. In particular, it is proposed to carry out the selection of anaerobic material according to its shear strength ssh
AM by the formula MUNS ¼ MT þ MAM ¼ MT þ ssh
AM AAM

d2  Wt ssh ¼ 0:2d33 ssh 2

ð3Þ

where MUNS —the moment of unscrewing the bolted joint; MT —tightening torque of the bolted joint; MAM —moment required to cut the intermediate material when unscrewing; AAM —cut-off area of the intermediate material; d2 —average bolt thread diameter; Wt ¼ 0:2 d33 —moment of resistance of the cross section of the bolt rod torsion; d3 —internal diameter of the bolt thread at the bottom of the cavity; ssh —the strength of the material of the bolt shear. The value AAM can be determined by the formula AAM ¼ 1:625 Ppd2 nAM

ð4Þ

where P—is the thread pitch; nAM —the number of threads, in the cavities where AM is used. Thus, with a decrease in the number of turns, the nAM unscrewing moment of a bolted joint decreases. It should be noted that the moment of BJ unscrewing can be significantly reduced by preheating the BJ to the temperature of destruction AM—200 … 250 °C.

Rational Provision of Robustness Properties of Bolted Joints …

169

4 Conclusion 1. An analysis of literature and industrial experience showed that the main disadvantage of bolted joint (BJ) is stress concentration in thread roots and non-uniformity of load distribution along thread turns of bolted joint. Under changing cyclic loads, crack propagation and destruction of parts occur in areas of stress concentration. 2. Among the existing design and technological methods for ensuring the reliability of BJ, BJ assembly using anaerobic materials is effective, which is confirmed by a number of studies. However, there were no studies related to the rational use of AM in the assembly of BJ. 3. A finite-difference BJ model was constructed, and the influence of AM on the stress–strain state in the zone of mating threaded contact of parts was studied. The axial load distribution was determined by thread winding in a bolt–nut joint. The compounds were studied when AM is in the cavities of the contact zone of a limited number of turns (not the entire length of the BJ screwing). 4. It was established that when AM fills the cavities of the threaded contact zone along the entire length of screwing the bolt and nut, the load falling on the first turn decreases by 32% compared to the BJ without AM. Almost the same result is obtained when filling the AM voids in the first three turns. When filling the AM voids in the first two turns, a load reduction of 30% was achieved. Thus, for the purpose of rational use of AM, it is advisable to place in the cavities of the contact zone only the first two or three turns from the side of application of the load (no more than half of the make-up length). At the same time, the used volume of AM decreases by a factor of 2 and more, and the cost of BJ assembly operation with AM decreases. 5. The AM volume is determined, which is necessary to fill the cavities of the threaded contact zone, taking into account the number of turns, the cavities between which are filled with AM. 6. The condition for ensuring the unscrewing of the BJ, taking into account the tightening torque and the torque required to cut AM, has been revealed references.

References 1. Birger IA, Iosilevich GB (1990) Threaded and flanged connections. Mashinostroenie, Moscow 2. Birger IA, Shorr BF, Iosilevich GB (1993) Machine part strength calculation. Mashinostroenie, Moscow 3. Anurev VI (2001) Reference book for designer-mechanical engineer. Mashinostroenie, Moscow 4. Ivanov MN (1991) Machine parts. Vysshaya shkola, Moscow 5. Voyachek II (2006) Integrated design of fixed couplings. Penza State University, Penza

170

I. I. Voyachek et al.

6. Andrienko LA, Baikov BA, Ganulich IK et al (2002) Machine parts. Bauman University Publishing House, Moscow 7. Yakushev AI, Mustaev RKh, Mavlyutov RR (1979) Increasing the strength and reliability of threaded joints. Mashinostroenie, Moscow 8. Berezin SY, Chumakov RE (2005) Quality management of threaded connections. Assembly Mech Eng Instrum Making 11:37–41 9. Dunaev VV, Shirshov AA (2009) Improving the life and sealing of bolt joints in aircraft. Russ Eng Res 12:33–39 10. Ivanov AS, Baikov BA, Popov BA (2009) Check calculation of threaded joints subject to a splitting force and a tipping moment, taking account of contact pliability. Russ Eng Res 1:33–37 11. Iosilevich GB (1981) Concentration of stresses and deformations in machine parts. Mashinostroenie, Moscow 12. Yasin AM (2002) Technological provision of fretting resistance of threaded joints. Dissertation, Penza State University 13. Prokofiev AN (2000) Advanced technological methods to improve the quality of threaded connections. Eng J 2(35):9–12 14. Voyachek II (2003) Assembly of threaded joints using anaerobic materials. Assembly Mech Eng Instrum Making 10:24–26 15. Voyachek II, Kochetkov DV (2009) Improvement of the functional characteristics of threaded joints when assembling with anaerobic materials. Assembly Mech Eng Instrum Making 6:37–40 16. Ivanov AS, Reshetov DN (2001) Improving the method of calculation and design of threaded joints, loaded with tearing force and overturning moment. Russ Eng Res 4:30–36 17. Berezin SY, Leonov VN (2007) Elastic simulators in a solution of a Joukovski problem. Assembly Mech Eng Instrum Making 4:23–27 18. Zorev NN, Safarov YS, Safarov VK (1973) Patterns of distribution of tensile load along thread turns of a bolted joint. Russ Eng Res 12:10–14 19. Voyachek II, Artyomov II, Kochetkov DV, Voyachek LG, Trazanov AV (2011) The method of obtaining a threaded connection. RF Patent 2,413,099, 27 February 2011 20. Suslov AG, Fedorov VP, Gorlenko OA et al (2006) Technological support and improve operational properties of machine elements and their compounds. Mashinostroenie, Moscow 21. Kurnosov NE (2001) Ensuring the quality of fixed connections: monograph. Penza State University, Penza 22. Zakharchenko AD (2004) Fatigue strength of riveted and bolted joints of aircraft structures. Assembly Mech Eng Instrum Making 4:19–20 23. Voyachek II (2003) The application of anaerobic materials in assembling fixed joints type shaft—bushing. Assembly Mech Eng Instrum Making 9:33–37 24. Evstifeeva EA (2009) Technological provision of the strength characteristics of pressure joints in the assembly with anaerobic materials. Dissertation, Penza State University 25. Kostyukov VN, Naumenko AP (2009) Research on exploitation characteristics of anaerobic adhesives and sealants. Assembly Mech Eng Instrum Making 3:9–12 26. Tulinov AB, Goncharov AB (2003) New composite materials for the assembly and repair work. Assembly Mech Eng Instrum Making 7:26–28 27. Alyamovsky AA (2007) SolidWorks/COSMOSWorks. Engineering analysis by the finite element method. DMK, Moscow

Machining Accuracy Improving with the Use of Mobile Mechatronic Systems as Industrial Robot End Effectors E. I. Shchurova and P. G. Mazein

Abstract Industrial robots are widely used for machining due to technological flexibility, relatively small sizes, and a reasonably large working space. Tool spindles that drive rotary cutting tools are often used as robot end effectors. In this case, feed motion is carried out by moving of robot operating mechanism. However, robot relatively low rigidity compared with metal-cutting machine rigidity causes more significant errors of the motion path of the end effector in machining process. For this reason, accuracy improving on large-sized workpiece machining with the use of robots is the issue of the day. This chapter concentrates on the use of a compact mechatronic system, which is applied as a robot end effector and is attached to the workpiece by means of electromagnets. The positioning of the fixed mechatronic mechanism (MM) in the workpiece coordinate system is checked using laser instrumentation system. After machining of the workpiece in the specified area, the mechatronic system is reinstalled in a new area. A preliminary rigidity estimation of the mechatronic device to achieve the required machining accuracy has been done. It has been detected that cutting tool is the least rigid element of mechatronic mechanism.







Keywords Machining Industrial robot Mechatronic system Machining accuracy Attachment rigidity





Mobile system


1 Introduction Industrial robotic manipulators are widely used in machine-building manufacture due to technological flexibility and relatively small size with a reasonably large working space [1]. A wide variety of such robots is known. Among them there are articulated, parallel, rectangular or Cartesian, cylindrical, polar or spherical, pendular, and SCARA robots [2]. Large-sized workpiece machining is one of the E. I. Shchurova (&)
P. G. Mazein South Ural State University, 76, Lenin Avenue, Chelyabinsk 454080, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_19

171

172

E. I. Shchurova and P. G. Mazein

application fields of such robots [3, 4]. Tool spindles that drive drills, core drills, reamers, taps, milling cutters, and other tools are often used as robot end effectors [5–10]. In this case, feed motion is realized by displacement of robot operating mechanism. There are two main groups of robots used in large-sized workpiece machining: robots with sequential kinematics and parallel kinematics [2]. Along with mentioned above advantages, robots have some disadvantages. Compared with CNC machines, robots have relatively low rigidity, large errors of repeatability, and errors of end-effector motion path in machining [11–13]. For example, passport data of small-sized robot KUKA KR 3 R540 specifies that its pose repeatability (ISO 9283) is ±0.02 mm [14]. Studies of such robots under load show that the error may increase to 1 mm [1, 11, 12]. In use of these robots together with errors of repeatability, there exist other kinds of problems. There are problems with dynamics and vibrations, negatively affecting machining accuracy, and machined surface roughness [15]. Similar problems are actually for robots with parallel kinematics too. For example, error value in machining using Fanuc F200iB hexapod robot is about 600 microns [16, 17]. These facts generate a need for control program corrections either before the robot starts working [18] or in the machining [19]. The method of using laser CMMs to check machining accuracy is suggested in [20]. Sensors are installed on robot end effectors. According to sensor readings, the CMM accurately determines end-effector position in the coordinate system of the floor, on which both the robot and the workpieces are fixed. In general, the problem of ensuring high accuracy of robotized machining is not solved still. In metalworking along with robots, mechatronic mechanisms are widely used. They are mounted on the work spindles of CNC machines and increase technological flexibility. As a rule, power tools with traditional sequential kinematics are used. As an example, power tools with rotary primary motion, such as drills, taps, milling cutters [5], and power tools with radial feed motion, such as boring tools [6–10], can be used (Fig. 1a). Similar solutions are available for modification of three-axis milling machines to five-axis by installing model 5414 R3 Programmable Five-Axis Spindle Head Attachment manufactured by Tri-Tech Precision Products, Inc. [21] (Fig. 1b). This head attachment provides two additional rotary coordinates for milling cutter. Installation of two perpendicularly located tool-holding slides for longitudinal and transverse feed motion, an axial feed device, and a motor of main motion (tool rotation) drive seems to be a highly reliable solution. Contouring head system manufactured by Cogsdill is the solution of such type. Especially, ZX200-TC-KM500CM device weighing 107 kg, with the sizes Ø200  307 mm enables displacement of 38 mm along the slides, with an accuracy of 3 lm (7 quality) [6]. A more flexible solution is the use of mentioned above head attachment with the slides. In this case, a mechatronic mechanism (MM) with five controlled coordinates is composed. Another solution is the use of devices with parallel kinematics, such as Newport (Fig. 1c) hexapod [22]. This equipment has the following characteristics: mass of 2.2 kg, sizes Ø200  151 mm and provides the following displacements: X/Y/Z = ± 17/ ± 15/ ± 7 mm and hx/hy/hz = ± 90/ ± 8.50/ ± 180 with the

Machining Accuracy Improving with the Use of Mobile …

173

Fig. 1 a Mechanism of tool programmable transverse motion [6]; b mechanism of tool spindle rotation about two axes [21]; c high precision hexapod, 125 mm diameter platform, 5 kg load, M6; d high load capacity hexapod with 50 kg centered [22]

“Origin Repeatability” accuracy: X/Y/Z = ± 5.0/ ± 5.0/ ± 2.5 lm and U/V/W = ± 2.5, ± 2.5, ± 5.0 mdeg. Hexapod “Rigidity” is for the axes X/Y/Z = 2/2/25 N/l m. “Maximum Speed” is for the axes X/Y/Z = 14/12/5 mm/s. An example of more rigid and massive equipment is shown in Fig. 1d. Its mass is 15.5 kg, and displacements along the axes are three times as much as for the previous equipment. Thus, the reviewed mechatronic systems are rather accurate and rigid. Small end-effector displacements relative to the workpiece are obvious disadvantages of such systems. Therefore, it is necessary to combine mobility and a large workspace of robotic manipulators with high accuracy and rigidity of mobile mechatronic systems. It seems expedient to use small mechatronic systems as end effectors of manipulators. To ensure proper rigidity, it is advisable to use additional adaptive bearing supports, which enable to fix the needful device on the workpiece surface. Finally, the accuracy the mechatronic system displacement, its installation, and attachment on the workpiece can be estimated using laser CMMs [20, 23]. All these issues have not been studied deeply yet. Both layout arrangement and rigidity of mobile mechatronic systems and the rigidity of attachment on the workpiece systems have to be examined. In what follows, some of these issues are considered.

2 Development of Machining Mechatronic System General Layout Arrangement and Its Position Checking in the Workpiece Coordinate System 2.1

Problem Statement

To improve machining accuracy, it is proposed to execute all the feed motions by mechanisms of a small mechatronic system attached to the workpiece by means of electromagnets. Therefore, there are two main problems: (1) required technological capabilities of the mechatronic system guaranteeing and (2) machining errors minimizing using a mobile MM and a robot. The second problem has a direct link with three generalized factors [24, 25]. The first generalized factor (2-1) is linked with the estimation of the errors induced by the cutting tool: errors before

174

E. I. Shchurova and P. G. Mazein

machining—the accuracy of tool manufacture and errors in machining—its elastic and thermal strains, and tool wear. The second generalized factor (2-2) is linked with the estimation of errors emerging because of deviations from the specified motion path of the tool-fixing unit. The third generalized factor (2-3) is linked with the estimation of errors that emerge because of workpiece thermal and elastic strains in machining. The errors linked with the second generalized factor (2-2), related to motion path deviation, in their turn, depend on the guarantees of the following parameters: (2-2-1) specified motion of MM tool-fixing unit accuracy provided with CNC system commands; (2-2-2) specified rigidity of the fitting which attaches MM to the workpiece and (2-2-3) specified accuracy of MM installation in the workpiece coordinate system. Next, we consider a number of issues of listed problems. (1) MM technological capabilities guarantee. This problem includes the following set of issues. First, it is the selection of MM-type system with sequential or parallel kinematics. Secondly, it is fixing of a number of tool motion controlled coordinates. In the third, the ranges of displacements along the indicated controlled coordinates should be prescribed. (2) MM tool-fixing unit motion accuracy guarantee. This problem includes two main issues. First, the errors of tool-fixing unit idling motion have to be estimated. These errors are caused by MM manufacturing accuracy; the accuracy of MM drives; the accuracy of CNC system commands realization in the toolpath. Secondly, there are errors of end-effector motion in chip clearance. These errors are linked with MM end-effector motion errors in cutting process and depend on operating mechanism elastic and thermal strains, displacements in the clearances between mated moving elements. Thus, the problem of the required tool motion accuracy guaranteeing can be decomposed into the following three solution stages. First, it is necessary to rank these factors according to their efficiency on the machining accuracy. Secondly, it is necessary to determine the functional relations between actual motions of tool cutting edges profiling parts, mechatronic system parameters, system control parameters, tool parameters, machining conditions, and workpiece parameters. Thirdly, it is necessary to develop methods of mechatronic system control, which enable to compensate adverse effects of all the factors, as much as possible. (3) The required rigidity of the attachment of the mechatronic system to the workpiece. This problem is the least formalized problem at the present time, since it is related to the type of device, which attaches the mechatronic system base to the workpiece. Thus, the solution to this problem includes the following stages. First, it is necessary to develop a device for the mechatronic system attachment. Secondly, it is necessary to detect the parameters of the accuracy and rigidity, which characterize the operation of the device elements and to calculate their values. Thus, the probable displacement of the MM coordinate system in the workpiece coordinate system is calculated. This displacement may be abnormal if the cutting force breaks the contact between the attachment

Machining Accuracy Improving with the Use of Mobile …

175

device and the workpiece (electromagnets are not able to hold the device). The displacement may be predicted if elastic strains of the fitting take place. Thirdly, it is necessary to develop methods of mechatronic system control, which enable to compensate adverse effects of its predicted displacement in the workpiece coordinate system. (4) The required accuracy of MM positioning in the workpiece coordinate system guaranteeing. This problem is related to the fact that MM displacement and positioning on the workpiece is carried out by mechanism. At present, as noted above, it is possible to position operating mechanism of the robot in the base coordinate system accurately. However, MM installation and its attachment to the workpiece may cause additional errors of the final device location. Therefore, it is advisable to verify this location. The application of laser CMMs is one of the suitable verification methods [23]. This system enables to calculate the location at any specific time. This is convenient for developing an adaptive control system of MM operating mechanism, depending on its displacement caused by the action of cutting forces at present time moment. Thus, there is a need for additional estimation of MM positioning accuracy using laser CMM. Summing up the problems, issues and stages of the examination presented above, the following hierarchical diagram has been developed (Fig. 2). It is highly likely that the presented list of problems, solution stages, and the issues under consideration is not exhaustive and requires a special study. It is clear that many of these issues are separate problems of mechatronics and robotics in mechanical engineering in their turn. Therefore, further we examine only some parts of the problems, which have the prime importance in our opinion. Among these problems, there are the following ones: (1) the selection of mechatronic system type, its kinematics, the number of controlled coordinates, and their operating ranges; (2) the selection of MM attachment to the workpiece system type and it development; (3) the selection of method and device for estimation of MM location on the workpiece.

Fig. 2 Scheme of problems, solution stages, and the issues under consideration to improve machining accuracy

176

2.2

2.2.1

E. I. Shchurova and P. G. Mazein

Development of the Mechatronic Mechanism Scheme, Scheme of Its Attachment to the Workpiece and Selection of the Device for Estimation of MM Location on the Workpiece Development of the Mechatronic Mechanism (MM) Scheme

In the selection of MM kinematics type, it is necessary to identify its basic requirements. Among these are performance reliability associated with the strength of operating mechanisms, MM rigidity, machining accuracy and process flexibility depending on the number of simultaneously controlled coordinates. For providing the required reliability, strength, and rigidity, it is advisable to use the accumulated experience of using mechatronic systems in metal cutting. As noted above, devices with sequential kinematics are widely used due to their reliability. A typical approach is the use of dovetail guides, which hold moving parts of the mechanism at the same time. Wedge assemblies enable to minimize clearance and provide minimal compliance in mates. To increase rigidity and accuracy, twin-screw feed mechanisms, such as applied for Mori Seiki machines, are used [26]. The box-shaped multifaceted feeding device also provides high rigidity of these machines. Based on such approaches, it is proposed to apply this type of kinematics for the mechanism of three linear displacements (Fig. 3a). To increase the flexibility of MM, it is appropriate to install a swivel head, for example, mentioned above attachment head by Tri-Tech Precision Products, Inc. (Fig. 3b). Alternatively, to expand the ranges of displacements a combination of the mechanisms with sequential and parallel kinematics can be used, but this combination is highly cumbersome (Fig. 3c).

Fig. 3 Layout and kinematics of sequential three linear coordinates feed MM (a); five coordinates MM (b); MM with 8 controlled coordinates (a system with parallel kinematics obtained from Newport [22])

Machining Accuracy Improving with the Use of Mobile …

2.2.2

177

Development of the Scheme of MM Attachment to the Workpiece

It should be noted again that the expediency of using robotic manipulators for machining is caused by the need of large-sized workpieces machining. As a rule, the surfaces of such workpieces have long surfaces, often with flat surface sections. Such sections should be used for MM attachment. Such areas are often flat; however, workpieces with surfaces of variable height are available, too. At present, in addition to traditional robot gripping devices, two main types of mechanisms are used for workpiece positioning: vacuum and electromagnetic. Combined vacuum– magnetic grippers are also used, for example Magvacu® Combi Grippers—47 N by Goudsmit Magnetic Systems BV [27–29]. With mass of 0.09 kg and sizes of Ø42  56 mm, this gripper provides the magnetic force of 47 N and the vacuum pressure of 70 N. Another device Magvacu® Combi Grippers—395 N with mass of 0.7 kg and sizes of Ø103  66 mm provides the magnetic force of 370 N and a vacuum pressure of 540 N. Such magnetic grippers are mounted on cylindrical bars and, as a rule, are spring-loaded in the axial direction. Therefore, a similar mechanism to attach MM to the workpiece is offered. As a first approximation, four bars fixed at the corners of a square mounting plate can be used. In turn, four springs and four pneumatic magnetic grippers should be installed on the bars. Their main purpose is to prevent MM displacement along the workpiece surface. The force of the robot has to be directed along the MM axis and the axis of the bars. The force should result in spring compression and, thus, the arrangement of all the four gripper bearing surfaces on the workpiece surface. Next, magnets and/or vacuum devices have to be turned on to fix the MM on the workpiece. In addition, a special device for grips fixation on the bars is required. For this purpose, it is advisable to use pneumatic clamping heads, similar to Pneumerlock® Clamp Head LHA by TAIYO Parker Fluidpower Co., Ltd [30]. The mass of this head seems to be overestimated up to 2.2 kg for sizes of Ø65  88 mm, but it provides a force up to 196 N on a bar with a diameter of 18 mm. Other heads weighing up to 0.5 kg with smaller sizes are designed for bars of smaller diameters. Obviously, another design solution with the use of grippers of smaller diameter and weight for large lengths and diameters of bars is required. Based on the solutions considered, the design of a mobile mechatronic system installed on the robotic arm may look like as shown in Fig. 4.

Fig. 4 Mobile MM design (a); MM with attachment mechanism and sensors for monitoring with the use of laser CMM (b)

178

2.2.3

E. I. Shchurova and P. G. Mazein

Development of a Scheme for Estimation of MM Location on the Workpiece

As noted above, the main disadvantage of robotic arms is the relatively low tool positioning accuracy in the workpiece coordinate system, especially in the case of variable load caused by cutting forces. The need to attach MM to the workpiece surface using four pneumatic magnetic grippers also requires the selection of appropriate areas on the workpiece. Thus, the robot has to deploy and place the MM with allowance for working spaces overlap and for attachment ensuring. The inevitable elastic strains in the course of attachment cause additional errors. Therefore, it is necessary to estimate the location of MM coordinate system in the workpiece coordinate system accurately. Since the workpiece and the robot are mounted on the floor, it is reasonable to install laser RBV Faro Laser Tracker ION by FARO Technologies Inc. in the same place [23]. At the four vertexes of the MM mounting plate on its reverse side, four Magnetic SMR Probe Adapters have to be installed [23]. In this case, the location of the MM coordinate system in the workpiece coordinate system is easily estimated both in the course of attachment and of machining, and the corresponding corrections of the control program can be executed in advance or during machining. Thus, in technical perspective accuracy improving on large-sized workpieces machining by robotic manipulators with mobile MM, devices of MM attachment to the workpiece and of MM location estimation using laser CMMs are realizable. However, in practice, the proposed machining scheme may not furnish the desired result, if in machining process, magnitudes of forces are too large and strains are much higher than the expected errors. Therefore, at the next stage of the study it is necessary to execute a number of calculations, including calculations of strains of the technological system elements. The goal of these calculations is to estimate MM rigidity in advance.

3 Calculation of MM Attachment Point Strains The discussed mobile MM design is an initial development. All the sizes are accepted constructively based on existing experience in small machine tools design. However, already at this stage it is advisable to examine the MM rigidity to determine ways of further improvement. In this connection, 3D models of all MM elements have been imported into ANSYS software and finite element calculation model of MM steel elements and of cutting tools has been developed. The total number of FE model nodes is more than 3.4 million. More than 2.2 million elements of Solid-45 type (tetrahedron) and of CONTAC170-174 contact pairs between all assembly parts have been generated. Clamping on all six degrees of freedom has been executed on magnetic grippers bearing surfaces. The load in the ratio Px/Py/Pz = 1/2/10 has been applied to the end mill face. Springs and control sensors have been removed conditionally. The results of strain calculation are shown in Fig. 5.

Machining Accuracy Improving with the Use of Mobile …

179

Fig. 5 FE model and strains of MM elements and of the tool

As shown in the figure, maximum strains are located in the tool body. The rigidity of the MM elements is several orders of magnitude higher. This fact leads to the conclusion that it is advisable to reduce the sizes of MM elements to reduce its total mass. Optimization of the geometry and materials of the MM should be carried out at the next research stage. However, it is obvious that the operational calculation of the MM set of tools rigidity is also necessary.

4 Conclusions (1) Accuracy improving on large-sized workpiece machining using robotic manipulators is achievable by the use of mobile mechatronic systems as end efforts, which are attached to the workpiece with magnetic and pneumatic grippers. (2) It is necessary to improve the design of the mechatronic device based on the calculation of the rigidity of its elements and the rigidity of the cutting tools.

References 1. Aliev R, Guseynov R (2011) Milling robots—technical state view. Autom Mod Technol 11:11–18 2. ISO 8373:2012. Robot and robotic DEVICES. Vocabulary, 48 p 3. Robotic machining (2015) Stäubli International AG, 7 p 4. Moler D-IC, Schmidt HS, Koch P (2017) Machining of large scaled CFRP-parts with mobile CNC-based robotic system in aerospace industry. Procedia Manuf 14:17–19 5. Driven tools (2018) Motorizzati. Gerardi tooling, 48 p 6. The world of ZX (2015) Cogsdill, 12 p 7. Metal Cutting Solutions (2013) Nickunj Eximp Entp P Ltd, 52 p

180

E. I. Shchurova and P. G. Mazein

8. 9. 10. 11.

Allied criterion boring system. Catalog (2014) Allied Machine & Engineering Cor, 72 p Precision boring systems (2011) Criterion Catalog no. 23, 85 p Contouring head system (2012) ZX TM Cogsdill, 5 p Sirinterlikci A, Tiryakioglu M, Bird A, Harris A, Kweder K (2009) Repeatability and accuracy of an industrial robot: laboratory experience for a design of experiments course. Technol Interface J 9(2):1–10 Zhang T, Du L, Dai X (2014) Test of robot distance error and compensation of kinematic full parameters. Adv Mech Eng, 1–9 Zhang D, Wei B (2017) Adaptive control for robotic manipulators. Taylor & Francis Group, 441 p Kuka KR 3 R540. Technical data. 0000-270-971/V16.3/05.09.2018/en Semyonov EN, Sidorova AV, Pashkov AE, Belomestnykh AS (2016) Accuracy assessment of Kuka KR210 R2700 extra industrial robot. Int J Eng Technol 16(01):19–25 Barnfather JD, Goodfellow MJ, Abram T (2017) Positional capability of a hexapod robot for machining applications. Int J Adv Manuf Technol 89:1103–1111 Barnfather JD, Goodfellow MJ, Abram T (2018) Achievable tolerances in robotic feature machining operations using a low-cost hexapod. Int J Adv Manuf Technol 95:1421–1436 Zhang J, Cai J (2013) Error analysis and compensation method of 6-axis industrial robot. Int J Smart Sens Intell Syst 6(4):1383–1399 Pan Z, Zhang H (2009) Improving robotic machining accuracy by real-time compensation. In: ICCAS-SICE International joint conference, IEEE, USA, pp 4289–4294 Morozov M, Riise J, Summan R, Pierce SG, Mineo C, MacLeod CN, Brown RH (2016) Assessing the accuracy of industrial robots through metrology for the enhancement of automated non-destructive testing. In: IEEE International conference on multisensor fusion and integration for intelligent systems (MFI), pp 1–6 5-Axis machining alternatives (2016) Tri-Tech Precision Products, Inc, 2 p HXP Hexapods. 6-Axis-parallel kinematic positioning systems (2018) Newport BR-041601-EN (22/05/18), 20 p FARO Laser tracker accessories manual (2017) FARO Technologies Inc, 134 p Matveev VV (1978) Narezanie tochnich rezb (Cutting of precision threads) Moscow. Mashinostroenie, 88 p Shchurov IA (2004) Teoriya rascheta tochnosti obrabotki i parametrov instrumentov na osnove diskretnogo tverdotel’nogo modelirovaniya (The theory of precision machinery and tools parameters calculation on the base of discrete solid modelling). SUSU, Chelyabinsk, 320 p. http://lib.susu.ru/ftd?base=SUSU_METHOD&key=000436340&dtype=F&etype=.pdf NMV 5000 DCG High-precision, 5-axis control vertical machining center (2007) MoriSeiki V.0709.CDT.0000, 50 p Magnet grippers (2012) Goudsmit Magnetics, 4 p Magvacu Combi Grippers—47 N. HGC-RO-040-VR-R-M-F (2018) Goudsmit Magnetics, 1 p Magvacu® Combi Grippers—395 N. HGC-RO-100-VR-M-G-F (2018) Goudsmit Magnetics, 1p Air-oil booster. Pneumatic equipment (2004) Taiyo Cat. No. A00421. High power system 1 1(R):36

12. 13. 14. 15. 16. 17. 18. 19. 20.

21. 22. 23. 24. 25.

26. 27. 28. 29. 30.

Voxel and Finite Element Modeling of Twist Drill E. I. Shchurova

Abstract Digitalization of an engineering process necessitates the development and application of digital models of all the technological systems elements, including cutting tools. Such digital models have to include geometric models of tool surfaces and models for physical modeling, for example, finite element or SPH models based on geometry. Digitalization also results in the evolution of universal approaches to model development of entity sets of cutting tools. At present time, two approaches to geometric modeling are used: analytical solid modeling and discrete solid modeling. The latter type of modeling, based on algebra of sets, is more flexible and more computationally stable. Universal models of thread-cutting tools have been developed by this time. However, discrete models of such widely used tools as twist drills have not been worked out still. The objective of the presented paper is to develop voxel and finite element models of standard twist drills with solid body. The developed twist drills model make it possible to obtain sets of tool surface points and to calculate finite element meshes. The model is suitable for twist drills of any design presented by the state standards of the Russian Federation (with the exception of drills with thinned chisel edge). Keywords Twist drill modeling


Voxel
FEA
Parametric model
Discrete solid

1 Introduction Engineering process is a subsystem of a single complex of process automation on any modern enterprise, which is traditionally associated with the conception of product lifecycle management (PLM) [1]. Engineering process includes such traditional systems as computer-aided design (CAD), computer-aided engineering (CAE), and computer-aided manufacturing (CAM) [2]. This situation is caused by E. I. Shchurova (&) South Ural State University, 76, Lenin Avenue, Chelyabinsk 454080, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_20

181

182

E. I. Shchurova

the fact that the industry sector itself requires the development of techniques (CAD for manufacture), the analysis of techniques operation (CAE for manufacture), and techniques production (CAM). It is well known that techniques include machines, attachments, tools, and workpieces [3]. All these objects are involved in the machining process and are affected by factors of various physical phenomena. These impacts cause the generation of workpiece surfaces, which differ from the nominal construction part surfaces specified in the drawings. Minimization of surface inaccuracy enables to get more precise machined parts and, accordingly, quality product. In this connection, the development of geometric and physical models of technological system elements is a topical problem. It is necessary to forecast machining errors and to affect the system for the inaccuracy minimization. At present, such models are developed using three main approaches. The first and the earliest approach is the usage of specific analytic relations [4]. The second approach is the usage of space-analytic geometry, and of classical solid modeling, of finite differences method or finite elements method associated with it [5–8]. The latest, third approach is the usage of numerical methods of geometric modeling, which is at present time associated with voxel modeling [9, 10]. Another name for this type of modeling is discrete solid-state modeling [11]. Geometrical modeling is required for following estimation of tool rigidity and strength and also for calculation of machining accuracy [12, 13]. Meshless calculation methods, for example, the SPH method, are used for physical modeling in increasing frequency [14]. The absolute stability of calculations caused by the use of algebra of sets is the advantage of discrete solid-state modeling. Unlike space-analytic geometry, an approach based on algebra of sets does not require solving systems of equations. Another advantage of discrete modeling is the possibility to use arbitrarily small geometric elements of models with any shape complexity [15, 16]. The minimum size of model element can be equal to one voxel. The shape complexity of the geometric element can be caused, for example, by tool wear, which is related to complex physical phenomena in the cutting zone [17]. Thus, it is not possible to model the surface of wear using the analytical dependence, as a rule. The disadvantage of discrete modeling is obvious—great computer resources are needed. However, the complexity of this problem is reducing in time. In any case, this disadvantage is non-essential and is not associated with scientific problems in mechanical engineering [18]. For this reason, it appears advisable to model elements of the technological system using numerical modeling. At present, mathematical relations of discrete solid-state models for thread-cutting tools have been developed [11]. These relations and appropriate software make it possible to calculate point sets of the surfaces and finite element meshes for almost all known standard taps, and even for special taps, and for a great number of working elements of thread-cutting heads [11]. The application of this approach enables to calculate operating parameters of these tools (rake, cutting edge, flute area, tool strength) and to model workpiece thread machining. Such modeling in its turn enables to predict specific angular, axial, and diametric thread sizes and to determine thread integral characteristics, such as virtual pitch thread diameter. However, specific features of twist drills make

Voxel and Finite Element Modeling of Twist Drill

183

it impossible to use mathematical relations, which have been developed for thread-cutting tools. For example, second-order tools for drill flutes machining have differences in their specified sizes: There are free sizes that need to be determined according to the parameters of the radial drill profile [19]. Drill radial section parameters should be determined in different way. Finally, major flank grinding is performed using different set of surfaces, including a helical surface. There are specific features of the formation of drill land, land thinning, and body clearances. For all these reasons, the purpose of the presented research is the development of analogous discrete geometric models and calculation of finite element meshes for the other commonly used group of tools—twist drills. This type of tool is selected for modeling because drilling and internal thread cutting have similar feature—the accuracy of machined holes significantly depends on the tool geometrical parameters, tool manufacturing errors, tool wear, and tool strains among them. At this stage of the research, all the models of twist drills with the solid body presented in the following state standards of the Russian Federation have been studied: GOST 10902-77, GOST 10903-77, GOST 12121-77, GOST 12122-77, GOST 12273-71, GOST 12274-71, GOST 12275-71, GOST 12276-71, GOST 19543-74, GOST 19544-74, GOST 19545-74, GOST 19546-74, GOST 19547-74, GOST 20694-75, GOST 20695-75, GOST 20696-75, GOST 20697-75, GOST 2092-77, GOST 4010-77, GOST 8034-76, GOST 886-77. The relations obtained enable to register in the universal complex model the design features of all the twist drills, with the exception of drills with thinned chisel edge.

2 Voxel Modeling of the Twist Drills 2.1

Problem Statement

To develop universal model of solid body twist drills, it is necessary to decompose drill body into three main elements: flutes including faces; lands and body clearance; and major flanks of the cutting part. Following the previously developed methodology modeling which is based on the compromise approach with the usage of algebra of sets has been executed. This approach includes derivation of inequalities for the radius vectors of the surface points of the specified three drill body elements in the radial cross sections and assembling these conditions into one integrated condition. Thus, it is necessary to derive inequalities for each of drill body elements, in view of all possible methods of drill manufacturing. It is obvious that such inequalities will have specific form not only for intervals of radial sections, but also for the corresponding intervals of axial sections, each of which may include specific drill design parameters or their combination.

184

2.2 2.2.1

E. I. Shchurova

Theoretical Investigation and Voxel Modeling Flute Surface Modeling

Geometrical parameters of twist drill flutes are presented by two state standards: GOST 19543-74 and GOST 20694-75. By analogy with production process of taps, drill flute surface and sizes are generated by second-order tools, such as single-formed cutters. It is clear that at the time of milling flute surface topography in any section is always different from the profile of second-order tool axial sections. In this connection, it is necessary to solve the problem of helical surface generation by disk blades in the general case. This problem was studied rather well in previous decades. Both analytical and numerical problem solutions are known [4, 20, 21]. Thereby in the presented study, we use the assumption that the profile and the set of geometric parameters of the second-order tool axial section fully correspond to the profile and parameters of the flute. Therefore, it is considered that flute parameters can be calculated according to standard second-order tool axial sections parameters. In this case, calculation scheme in the radial drill section is given in Fig. 1. Parameters presented by the standards (Figs. 1 and 2a) are the design data. The drill parameters are as follows: d—drill diameter; d1 —body clearance diameter; a—half of web thickness (this parameter may vary depending on z—drill axis coordinate); B—width of fluted land; x—helix angle; L—overall length. Parameters of the single-formed cutter used for drill production are as follows: Bm —cutter width; r—cutter profile major circular arc radius; r1 —cutter profile minor circular arc radius; b—distance between cutter left face and minor circle center; c—distance between cutter left face and the mark; w—the angle between cutter right face line and cutter its rectilinear edge.

Fig. 1 a Twist drill radial section including the flute; b Twist drill axial sizes

Voxel and Finite Element Modeling of Twist Drill

185

Relations for drill surface point calculation are determined in the auxiliary coordinate system—the flute coordinate system. Angles of counter-clockwise rotation about abscissa axis are considered as negative value angles, and angles of clockwise rotation are considered as positive value angles. Then, to cover the variation of /f polar angle within three main ranges referred to the flute coordinate system (Fig. 2), the following relations which enable to calculate voxels inside the flute (weak inequality) or on the flute surface (equality) are used: /f ¼ w. . .90  w; R  r1 = cosð/f Þ;

ð1Þ

/f ¼ s. . .  w; R  r1

ð2Þ

/f ¼ 90  w. . .60  s; R  ðr  r1 Þ cosð/f  gÞ þ

ffi qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 ðr  r1 Þ cosð/f  gÞ ðr  r1 Þ2 þ r1 2

ð3Þ

where s ¼ 2p  n is calculated using one of the following formulas, selected by the designer at own discretion: ð1Þ n ¼ arctanððh  r1 Þ=ðBm  b  fe cosðwÞÞÞ; the designer sets fe distance; ð4Þ ð2Þ n ¼ arctanððh  r1 Þ=ðBm  b  eÞÞ; the designer sets e distance

ð5Þ

It is clear that the third option can be used, according to which the designer sets n angle directly. Angle g can also be determined in one of the two ways, depending on which drawing parameter is used by the designer.

Fig. 2 a Flute producing single-formed cutter axial section; b Drill radial section used for land and body clearance parameters calculation

186

E. I. Shchurova

ð1Þ g ¼ n þ w; where n has been calculated above; ð2Þ The designer sets g angle directly:

ð6Þ

Next, it is necessary to calculate the flute coordinate system origin coordinates (af and bf ) referred to the drill coordinate system using the following formulas: pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi G ¼ Bm  c  e; N ¼ r 2  G2 ; Q ¼ r1  hþN;   ð7Þ  B1 ¼ B=cosðxÞ; B1 ¼ B=cosðxÞ; h1 ¼ arccos 0:25d12 þB21 þ0:25d 2 ðB1 d Þ h2 ¼ arctanð2a=d Þ; S ¼ B1 sinð0:5p  h1  h2 Þ; af ¼ r1  a; bf ¼ S1 þ Q=cosðwÞ  0:5d:

ð8Þ

Then for the first and the second flutes (j :¼ 1 and j :¼ 2) ~ qðxfi ; yfi ; zfi Þ vector coordinates referred to the coordinate systems of these flutes can be calculated. Each of such vectors corresponds to the vector of number i current point referred to the drill coordinate system: 0 1 0 0 1 0 11 xfi af xi   @ yfi A ¼ Mz v  xf @Mz ðv þ xf Þ@ bf A  @ yi AA; ð9Þ zfi 0 zi v ¼ ð0; pÞ; xf ¼ 2z tanðxÞ=d:

ð10Þ

Then, /f polar angle and R radius vector of the current point for formulas (1) are calculated as follows: qffiffiffiffiffiffiffiffiffiffiffiffiffiffi    ð11Þ /f ¼ arctan yfi xfi and R ¼ x2fi þ y2fi :

2.2.2

Land and Body Clearance Surface Modeling

According to the standards, land and body clearance surface parameters may be as follows (Fig. 2b): cylindrical land segment ð/1 . . ./2 Þ; plane land segment ð/2 . . ./3 Þ; body clearance ð/4 . . ./5 Þ; additional cylindrical land at the edge of body clearance ð/6 . . ./7 Þ; and plane segment of this land ð/7 . . ./8 Þ. Some of these parameters are presented by the standards: f —width of land; f1 —width of cylindrical land segment; af —clearance of the land; fh —the land width at the body clearance edge; fh1 —the width of the cylindrical land section at the body clearance edge; ahf —clearance of the land at the body clearance edge. To model these surface segments, it is necessary to calculate the following angles.

Voxel and Finite Element Modeling of Twist Drill

187

/1 ¼ arcsinða=d Þ þ p=2; /2 ¼ p=2  arcsinððf1  aÞ=ð0:5dÞÞ;

ð12Þ

/3 ¼ p=2  arctanððf  aÞ=ð0:5d  DÞÞ; D ¼ ðf  f1 Þ tanðaf Þ;

ð13Þ

if D\0:5ðd  d1 Þ than /4 ¼ p=2  arctanð2ðf  aÞ=d Þ else /4 ¼ /3 ; /5 ¼ /7 þ 2 arctanðfh1 =d1 Þ; /6 ¼ /7 þ 2 arctanðfh1 =d Þ; /7 ¼ /8 þ arctan½2ðf  fh1 Þ=d ; /8 ¼ /s ¼ arctanð2ð0:5d  SÞ=d1 Þ; where S is calculated using ð8Þ:

ð14Þ ð15Þ ð16Þ

Then for mentioned ranges of values, we have the following (non-strict inequality corresponds to body voxel coordinates, and equality—to drill surface): ð17Þ / ¼ /1 . . ./2 ; R  0:5d;      / ¼ /2 . . ./3 ; R  ðf1  aÞ tan af þ 0:5d cos af cosð/ þ 0:5p  ah1 Þ; ð18Þ 

2.2.3

/ ¼ /3 . . ./4 ; R  ðf  aÞ=cosð/Þ;

ð19Þ

/ ¼ /4 . . ./5 ; R  0:5d1 ;

ð20Þ

/ ¼ /5 . . ./6 ; R  0:5fh1 =cosð/  wh Þ; wh ¼ 0:5ð/6 þ /7 Þ;

ð21Þ

/ ¼ /6 . . ./7 ; R  0:5d;     / ¼ /7 . . ./8 ; R  0:5d cos /7 þ ahf  wh cos / þ ahf  wh :

ð22Þ ð23Þ

Drill Flank Modeling

In practice, two grinding methods are most often used: facet sharpening and helical sharpening. Further, both these methods are examined. Facet sharpening is characterized by a—clearance of major cutting edge. In this case, the equation of the plane referred to drill coordinate system is: Vx þ Wy þ Fz þ D ¼ 0; where V ¼ sinðaÞ; W ¼ 0:5dcosðaÞ=sinð/Þ; F ¼ 0:5d cosðaÞ  L; D ¼ 0:5ad sinðaÞ

ð24Þ

In view of this, R  D=cosð/ þ arctanðW=VÞÞ:

ð25Þ

Helical sharpening is characterized by such basic data as P—helical parameter, and /d —half of the point angle. In this case, non-strict inequality corresponding to voxels referred to drill coordinate system is:

188

E. I. Shchurova

R  ðz0 þ P/Þ tanð/ks Þ=sinð/Þ:

ð26Þ

where z0 —parameter that determines the position of the section, where helix generating line is specified; /ks —angle between helix generating line and drill radial plane. tanð/ks Þ ¼ ðP arctanð2a=dÞ þ 0:5d cotð/d Þ  0:5pPÞ=ð0:5d  aÞ:

ð27Þ

The obtained relations (1)–(27) make it possible to calculate drill body and surface voxel coordinates. Further, using previously developed methodology it is possible to determine drill operating characteristics and to calculate finite element meshes for drilling process modeling, calculation of drill thermal and elastic strains.

3 Calculation of Drill Surface Voxel Nodal Points and of Finite Element Meshes The following results of voxel nodal points for the surfaces of various drill types calculation and of finite element meshes generation confirm the adequacy of the relations obtained in Sect. 2. Discrete solid-state model which includes a set of inequalities describing flutes, lands, body clearances and major flanks enables to characterize any set of surfaces of standard twist drills with any sizes and various

Fig. 3 Twist drill photos and voxel nodal points of twist drills (x ¼ 60 ) with one a and two b lands; c twist drills (x ¼ 30 ) without lands; d twist drills (x ¼ 25 ) with one land; e core drill; f corresponding meshes of brick finite elements and stress calculation results

Voxel and Finite Element Modeling of Twist Drill

189

design elements. This confirms model versatility. Using the same model, generation of brick finite elements meshes has been done. Thereby, the model is suitable not only for geometrical, but also for following physical modeling (Fig. 3). Thus, the approach based on discrete modeling has confirmed the possibility to generate digital twins of metal-cutting tools for the subsequent calculation of machining accuracy.

4 Conclusions The accomplished research has the following conclusions. (1) The method of discrete solid modeling is sufficient to generate integrated geometric and finite element model of standard twist drills. (2) The method has proved universality of modeling and stability of calculations for generation of digital twins of these tools.

References 1. Vila C, Abellan-Nebot JV, Albinan JC, Hernandez G (2015) An approach to sustainable product lifecycle management (Green PLM). Procedia Eng 132:585–592 2. Xu X (2009) Integrating advanced computer-aided design, manufacturing, and numerical control: principles and implementations. IGI Global, 424 p 3. GOST 27.004-85 Industrial product dependability. Technological systems. Terms and definitions, 13 p 4. Astakhov VP (2010) Geometry of single-point turning tools and drills: fundamentals and practical applications. Springer, 584 p 5. Sambhav K, Tandon P, Dhande SG (2012) Geometric modeling and validation of twist drills with a generic point profile. Appl Math Model 36:2384–2403 6. Sambhav K, Dhande SG, Tandon P (2010) CAD based mechanistic modeling of forces for generic drill point geometry. Comput Aided Des Appl 7(6):809–819 7. Hsieh J-F, Lin PD (2005) Drill point geometry of multi-flute drills. Int J Adv Manuf Technol 26:466–476 8. Jovanovic JD, Spaic O (2012) Geometric modeling of twist drills. In: 16th international research/expert conference “Trends in the development of machinery and associated technology, pp 115–118 9. Jerard RB, Angleton JM, Drysdale RL, Su P (1990) The use of surface points sets for generation, simulation, verification and automatic correction of NC machining programs. In: Proceedings of NSF design and manufacturing systems conference, pp 143–150 10. Lynn R, Dinar M, Huang N et al (2018) Direct digital subtractive manufacturing of a functional assembly using voxel-based models. J Manuf Sci Eng 140:021006–2–021006-14 11. Shchurov IA (2004) Teoriya rascheta tochnosti obrabotki i parametrov instrumentov na osnove diskretnogo tverdotel’nogo modelirovaniya (The theory of precision machinery and tools parameters calculation on the base of discrete solid modelling). SUSU, Chelyabinsk, 320 p. http://lib.susu.ru/ftd?base=SUSU_METHOD&key=000436340&dtype=F&etype=.pdf

190

E. I. Shchurova

12. Makhecha A, Thangaraj A R, Sutherland JW (1994) Prediction of drilling thrust and torque using a mechanistic model calibrated through non-linear optimization. Manuf Sci Eng, ASME Bound Volume—PED 68(1):237–244 13. Abele E, Fujara M (2010) Simulation-based twist drill design and geometry optimization. CIRP Ann 59(1):145–150 14. Heisel U, Zaloga W, Krivoruchko D, Storchak M, Goloborodko L (2013) Modelling of orthogonal cutting processes with the method of smoothed particle hydrodynamics. Prod Eng Res Devel 7:639–645 15. Wu J, Yu G, Wang D, Zhang Y, Wang CCL (2009) Voxel-based interactive haptic simulation of dental drilling. In: Proceedings of the ASME 2009 international design engineering technical conferences & computers and information in engineering conference, pp 1–10 16. Uner G, Konukseven EI (2010) Development of a novel 6DOF multi-contact material cutting model for visiohaptic rendering applications. Int J Des Eng 3(3):260–275 17. Zhang C, Liu X, Fang J, Zhou L (2011) A new tool wear estimation method based on shape mapping in the milling process. Int J Adv Manuf Technol 53:121–130 18. Lynn R, Contis D, Hossain M et al (2017) Voxel model surface offsetting for computer-aided manufacturing using virtualized high-performance computing. J Manuf Syst 43:296–304 19. GOST 15543-74 Twist drills with cylindrical shank for working of light alloys. Middle series. Design, 8 p 20. Oancea N, Popa I, Teodor VG, Oancea VG (2010) Tool profiling for generation of discrete helical surfaces. Int J Adv Manuf Technol 50:37–46 21. Radzevich SP (2014) Generation of surfaces. Kinematic geometry of surface machining. CRC Press, 724 p

3D Modeling of Turbine Rotor Journal Machining with Location on a Bearing Bottom Half A. V. Shchurova

Abstract Restoration of worn turbine rotor journal without removing it from a power unit is carried out with location on cylindrical plain bearing bottom half. Thus, the workpiece is located on the machinable surface. Such location causes difficulties in the required cylindrical accuracy guaranteeing. Our earlier studies were based on the assumption that the worn rotor journal is dominantly a cylindrical surface and has a multi-lobed directional line in its radial cross section. This assumption makes it possible to determine rational parameters of the technological operation, including the variation of tool radial feed. The purpose of the present research is to study the influence of parameters, which characterize the entire surface of a worn rotor journal under the assumption that the surface does not belong to the class of cylindrical surfaces. It is proposed to describe the surface by classifying its inaccuracy into three main types of profile errors: radial, axial, and screw. Each of these error types is described by Fourier trigonometric series. The paper presents preliminary modeling of such complex surface machining with its location on a semi-cylinder and the identification of the relation of surface dominant harmonics and of tool radial feed variation on the machining accuracy. It has been determined that the highest accuracy is obtained using the coefficient of circular frequency in feed variation equal to (or exceeding by 3 or 4) the similar coefficient of the dominant harmonic of the transverse workpiece profile.










Keywords Turbine restoration Rotor journal Location on journal Location on bearing Centerless grinding Circularity deviation Voxel modeling










A. V. Shchurova (&) South Ural State University, 76, Lenin Avenue, Chelyabinsk 454080, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_21

191

192

A. V. Shchurova

1 Introduction During normal operational life of a turbine, its rotor damage may arise due to several factors such as steam erosion, rotor rubbing, etc. [1, 2]. Wear of rotor journals seating surfaces is often in the case that conducts to the increased vibration and stops of the equipment. In this case, repair operations including turning, milling, and grinding of such seating surfaces on the mounting bearing are carried out. In view of rotor large size and weight, one of the most economical ways of its journal restoration is to machine it without removing from the power unit [3, 4]. In this case, the upper half of the plain bearing is removed and the opened upper journal half is machined. Mobile machining device is installed on the half-plane of the unit, and the journal is machined due to the radial feed of a tool, rotating abrasive tool, for example [5]. At the same time, turbine rotor is rotated around its axis with location of its machinable journals on the bottom bearing half. Since the worn journal has irregularities, rotation with the location on it results in a constant displacement of the rotor shaft axis along different directions. All this causes difficulties in ensuring the required circularity deviations [6–12]. Earlier in our studies, it was detected that there are certain rational parameters of the tool installation and operating parameters, including the variation of radial feed. Selection of these parameters enables to minimize deviations during a fewer number of shaft revolutions [13–15]. It has been detected that these parameters are associated with the dominant harmonic of the shaft radial section profile. It has been assumed that shaft section profile is constant over the entire shaft length because shaft journal surface belongs to the class of cylindrical surfaces with multi-lobed (most often elliptical) radial profile. However, such an assumption, based on measurements of real energy turbines rotor journals, may not be valid for all cases. In general, journal surface may be entirely arbitrary. Therefore, it is advisable to obtain a mathematical description of such a surface and find out the relation of its parameters and the parameters of the technological operation and journal circularity deviations after machining. This will enable in practice to choose rational parameters of the technological operation and minimize the errors of the restored journal using shaft journal digital twin obtained with laser CMM, the surface point cloud [16–19].

2 Modeling of the Worn Journal Surface and Calculation of the Restored Journal Circularity Deviations in View of the Technological Operation Parameters 2.1

Problem Statement

A description of the real surface of a worn journal can be executed by developing its mathematical model. Earlier, the author has described the journal supposing, in consideration of practical measurements, that the worn journal radial cross section

3D Modeling of Turbine Rotor Journal Machining with Location …

193

can be represented as an ellipse. All the radial journal cross section profiles have been viewed as identical; the journal has been described as a cylindrical surface. Later, the radial sections have been described as multi-lobed figures with three, four, or more lobes. In this study, it is proposed to describe journal axial section profile in the same way. In this case, the surface has not only lobes in the radial section, but also the same lobes in its axial section. Obviously, the resulting surface will not belong to the class of cylindrical surfaces, nor to the class of surfaces of revolution. However, such a surface can have several planes of symmetry, which is incorrect for real surfaces. In this regard, it is proposed to take into consideration further class of deviations, a screw surface. Obviously, each of these deviations can be described by a Fourier trigonometric series. In addition, the screw surface can have either the right or left helix directions, or both directions can be used simultaneously. Apparently, the use of these three types of errors enables to obtain the most adequate mathematical models of real worn journals. The second important issue of restored journal accuracy improving problem is the determination of rational parameters of the technological operation. As has been established in previous studies, the most important parameters include w angle between the vertical axis and the radius vector directed from the longitudinal axis of the workpiece to the axis of the rotating tool. The other operation parameters are following: the number of revolutions of the workpiece during its machining, the maximum depth of the radial tool motion, and the law of radial tool motion variation. It is advisable to present this variation, as well as in the previous studies, using trigonometric functions. In fact, this variation can be realized using the CNC drive feed. All these issues are examined in more detail further.

2.2

Description of the Worn Journal Surface

The surface of the worn shaft journal can be described using the following formulas. It is proposed to perform the calculations in a cylindrical coordinate system. Then, current zi and /i coordinates can be calculated by the formulas: zi þ 1 ¼ zi þ Dz; zmin  zi  zmax ; /i þ 1 ¼ /i þ D/; /min  /i  /max

ð1Þ

where zmin . . .zmax coordinates of the overall cross sections of the journal; /min . . ./max extreme polar angles of the modeled surface(0…360°). Modeling of DRð/i Þ deviations in the radial section can be made according to the following known relation

194

A. V. Shchurova

DRð/i Þ ¼

1 X

  Dr/k cos x/k /i þ /k :

ð2Þ

k¼1

At this stage of research, it is proposed to examine only single harmonic of the series:   DRð/i Þ ¼ Dr/ cos x/ /i ;

ð3Þ

where Dr/ the amount of radius deviation from the circle in the radial section; x/ the number of journal surface lobes in its radial section. Modeling of deviations in the axial section and the helical shape of such deviations is also done using only single harmonic:   DRðzi Þ ¼ Drz cosðxz zi =½zmax  zmin Þ þ Drz/ cos xz/ ð/i þ zi =½zmax  zmin Þ ; ð4Þ where Drz —the deviation from a straight line—generating cylinder line in its axial section; xz —the number of lobes along the specified line within the length of the cylinder; Drz/ —the value of deviation from a straight line—generating cylinder along its axis associated with the rotation of the surface point along the helix; xz/ — the number of lobes of the specified helix within the length of the cylinder. The total cylindricity deviation of the journal surface is determined by the superposition: DRð/i ; zi Þ ¼ DRð/i Þ þ DRðzi Þ:

ð5Þ

Then, the coordinates of the journal surface points can be determined by the formulas: xi ¼ ½Rimax ð/i Þ þ DRð/i ; zi Þ cosð/i Þ;

ð6Þ

yi ¼ ½Rimax ð/i Þ þ DRð/i ; zi Þ sinð/i Þ;

ð7Þ

where Rimax ð/i Þ—the radius of the middle cylinder of the journal. If necessary, it is possible to represent it as a function of the polar angle, additionally.

2.3

Description of the Law of the Cutting Tool Motion Along the Radial Direction to the Axis of the Shaft Journal

This law has been presented in the previously published paper [20].

3D Modeling of Turbine Rotor Journal Machining with Location …

195

Then, the calculating formula for the tool feed at the specified workpiece rotation is SDi þ 1 ¼ SDi þ ð/max =/cut ÞDS þ A
DS cosðK
lobes
/i  / Þ;

ð8Þ

where SDi —tool feed at present rotation angle /i ; /max —maximum workpiece rotation angle with spark-out period when needed; /cut —workpiece rotation angle with tool feed; DS—average tool feed; A—amplitude of tool feed changing; K and / are frequency and phase of the current harmonic. Thus, all the necessary relations are described. Thus, in view of previously published relations [13–15, 20], it is possible to perform computer modeling and to determine rational parameters of a technological operation.

2.4

3D Modeling of Worn Shaft Journal Machining Neck with Its Location on the Half-Cylinder of a Plain Bearing

The obtained relations enable to develop computer program and to perform modeling of shaft journal machining. Following are the main parameters of the model areas: zmin ¼ 0; zmax ¼ 400 mm; Dz ¼ 1 mm; /min ¼ 0; /max ¼ 359 ; D/ ¼ 1 ; R ¼ 150 mm; x/ ¼ 3; xz ¼ 5; xz/ ¼ 5. Other model parameters have been increased: Dr/ ¼ 3 mm; Drz ¼ 2 mm; Drz/ ¼ 1 mm, for the visualization of modeling illustrations. In Fig. 1, the modeling results of the initial worn neck and neck after machining are presented. Modeling input data are as follows: The cut-in is performed on one revolution (nc ¼ 1), spark-out at zero tool feed is performed on one revolution (nf ¼ 1); angle of the tool installation w ¼ 42 ; maximum tool displacement to the center of the workpiece SDmax ¼ 10:8 mm. For value of K in Eq. (8), the following maximum circularity deviations have been obtained: for K ¼ 2 DR ¼ 3:787 mm; for K ¼ 3 DR ¼ 3:731 mm; for K ¼ 4 DR ¼ 3:939 mm. As shown in the figure, the surface of the shaft journal is rather complicated. Consequently, as the initial approximation of the mathematical model of the shaft journal with specified circularity deviations, it is sufficient to use the first terms of the trigonometric Fourier series for three directions: radial, axial, and screw.

Fig. 1 Nodal points of the journal surface voxels: a the initial surface; surface after machining modeling: b K = 2; c K = 3; d K = 4

196

A. V. Shchurova

It is obvious that processing on one turn does not enable to obtain the high-quality part. In this regard, further processing has been simulated for K ¼ 2 with various numbers of the workpiece revolution nc and with nf ¼ 1. So for nc ¼ 2 DR ¼ 2:787 mm; for nc ¼ 3 DR ¼ 1:571 mm; for nc ¼ 4 DR ¼ 1:363 mm; for nc ¼ 5 DR ¼ 1:501 mm. As can be seen, starting with a certain number of revolutions, the circularity deviation DR begins to increase again. Since for nc ¼ 4 the deviation is minimal, then the effect of increasing the number turns of spark-out at zero tool feed has been examined, nf ¼ 0. . .6. So for DR ¼ 1:865 mm; for nf ¼ 3 DR ¼ 1:262 mm; for nf ¼ 4 DR ¼ 1:224 mm; for nf ¼ 6 DR ¼ 1:113 mm. As can be seen from the results of the calculation, the circularity deviation asymptotically decreases. Finally, an attempt to study the effect of K—the law of the radial tool feed variation on deviation DR has been made. For nc ¼ 4 and nf ¼ 1, all the cases have been examined, from K ¼ 2 to K ¼ 9 (Fig. 2). As can be seen, the influence of K is not monotonous: A decrease is initially observed, and then for K ¼ 5, a certain rise

Fig. 2 Location of surface nodes and surface circular pattern after journal machining modeling

3D Modeling of Turbine Rotor Journal Machining with Location …

197

is observed; then, again there is a decrease till K ¼ 9 and then a rise again. The total minimum is observed at K ¼ 6. From this, we can conclude that the minimum error is not strictly related to one of the dominant harmonics. In this case, the value of K is two times higher than the dominant harmonic. This fact requires further explanation and further study.

3 Results and Discussions The obtained results show that in the case of shaft journal machining modeling with its location on the bottom half of the plain bearing in the three-dimensional formulation, the results of the study differ from those obtained earlier in the flat-simulation, in radial section. The effect of operating parameters is more complex and depends both on the dominant harmonic of one section and on such harmonics in other sections. There are no monotonic dependencies of the circularity deviation on the parameter of the law of variation of tool motion to the workpiece axis. It is necessary to obtain the relations of the trigonometric Fourier series to describe the surface of the shaft journal directly. Such relations should be found using the surface point cloud obtained with laser CMM.

4 Conclusion The accomplished research has the following conclusions. (1) As a preliminary approximation of mathematical model of the shaft journal with specified circularity deviations, it is reasonable to use the first terms of the trigonometric Fourier series for three directions: radial, axial, and screw. (2) For 3D shaft journal model, the relationship between the parameter of the radial tool motion (the coefficient at the circular frequency of the sine law) and the circularity deviation is not monotonic. Several minimum points of the function are observed if the parameter value exceeds the similar coefficient of the dominant harmonic by 3 or 4.

References 1. Mobile Lathe Information Sheet (2008) Alstom, France, Saint-Ouen, 3 p 2. Bloch HP, Geitner FK (2005) Machinery component maintenance and repair: practical machinery management for process plants. Elsevier Inc., USA, 630 p 3. Orbital Tool Technologies (2012) Shaft and journal repair, USA, 2 p 4. Portable machine tools (2011) Solutions for all in situ site machining applications, Acteon, Derby, UK, 13 p

198

A. V. Shchurova

5. Molochek V (1968) Remont parovyh turbin (Steam turbine repair). Energy, Moscow, 376 p 6. Brian RW (2009) Principles of modern grinding technology. William Andrew, USA, 416 p 7. Kang K (2003) Modelling of the centerless infeed (plunge) grinding process. KSME Int J. 17 (7):1026–1035 8. Wu Y, Kondo T, Kato M (2005) A new centerless grinding technique using a surface grinder. J Mater Process Technol 162–163:709–717 9. Xu W, Wu Y, Sato T, Lin W (2010) Effects of process parameters on workpiece roundness in tangential-feed centerless grinding using a surface grinder. J Mater Process Technol 210: 759–766 10. Xu W, Wu Y (2011) A new in-feed centerless grinding technique using a surface grinder. J Mater Process Technol 211:141–149 11. Xu W, Wu Y (2011) A new through-feed centerless grinding technique using a surface grinder. J Mater Process Technol 211:1599–1605 12. Xu W, Wu Y (2012) Simulation investigation of through-feed centerless grinding process performed on a surface grinder. J Mater Process Technol 212:927–935 13. Shchurova AV (2016) Modeling of the turbine rotor journal restoration on horizontal balancing machines. Procedia Eng 150:854–859 14. Shchurova AV (2017) Modeling the turbine rotor journal restoration located on cylindrical surface of the supporting bearer. Procedia Eng 206:1142–1147 15. Shchurova AV (2017) Izmenenie polozheniya osi vala pri ego shlifovanii s bazirovaniem v polucilindricheskoj opore (Axis relocation of a grinded shaft that is located on hemi-cylindrical bearing seat). Bull Tula State Univ Ser Tech Sci 8(1):338–344 16. Franceschini F, Galetto M, Maisano D, Mastrogiacomo L (2016) Combining multiple large volume metrology systems: competitive versus cooperative data fusion. Precis Eng 43:514– 524 17. Besic I, Gestel VN, Kruth J-P, Bleys P, Hodolic J (2011) Accuracy improvement of laser line scanning for feature measurements on CMM. Opt Lasers Eng 49(11):1274–1280 18. Moroni G, Syam WP, Petro S (2014) Performance improvement for optimization of the non-linear geometric fitting problem in manufacturing metrology. Meas Sci Technol 25:1–15 19. Du Z, Wu Z, Yang J (2016) Error ellipsoid analysis for the diameter measurement of cylindroid components using a laser radar measurement system. Meas Sci Technol 27(12):1– 13 20. Shchurova AV (2018) Effect of the tool-operating mode on circularity deviation in multi-lobed turbine rotor journal restoration with location on a bearing bottom half. Lecture Notes in Mechanical Engineering, in press

Geometric Modeling of Macro-defects of Parts Surfaces Based on Discrete Solid-State Modeling I. A. Shchurov

Abstract Traditional solid modeling enables to describe component parts and assemblies in an idealized formulation. It is considered that surfaces of model objects are canonical or spline surfaces, the edges are lines, and the vertices are dots. None of real component parts has such properties. In this regard, subsequent calculations, finite element analysis, for example, are not be completely adequate, especially for contact problems. Meanwhile, often the finite elements are comparable in size with the indicated defects and the description of such defects is realizable for the modern level of computer capacity. Thus, developing of CAD and following CAE models in view of surface defects is an important problem. At the present time, voxel modeling is increasingly used. The presented paper studies the application of voxel approach to develop the required CAD models. Voxel CAD modeling is called discrete solid modeling. The problems of collecting libraries of typical surfaces and libraries of surface defects are considered. Modeling of defects and of their locations on the surfaces is also the subject of this study. The results of modeling prove the validity of the primer approach. These results are presented as finite element meshes of models of component part surfaces and of surface defects.





Keywords Discrete solid modeling Finite element method Surface modeling Surface defects CAD CAE








Voxel modeling


1 Introduction Practically all modern design projects are carried out using CAD/CAM/CAE systems [1]. The basis of CAD/CAM systems is classical solid-state or surface modeling [2, 3]. All objects, such as component parts, assemblies, and devices are modeled using the equations of canonical or spline surfaces. The edges of the models are the intersection lines of such mating surfaces. Lines intersections genI. A. Shchurov (&) South Ural State University, 76, Lenin Avenue, Chelyabinsk 454080, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_22

199

200

I. A. Shchurov

erate vertex points [4]. It is obvious that none of the real objects is bounded with ideal surfaces, does not contain lines as its edges, and has no vertex points. The use of such ideal objects is a necessary compromise. This idealization is generally practical for following physical calculations, which use ideal geometric objects to generate finite element meshes, for example. Moreover, there is a practical need to remove small structural elements, such as chamfers and fillets [5]. However, there some are cases for which macro- and even micro-surface defects are significant. Such cases include, first, modeling of small-sized component parts, which sizes are comparable to the defect sizes [6]. In other cases, the interaction of some parts with others occurs in areas of small size, which are also comparable to defects. The latter include cutting process modeling with the chip formation [7–10]. As is known, in practice the minimum values of cutting depth and feed per tooth are about percent of a millimeter [11]. Obviously, the macro-defects of the workpiece surface are comparable with such sizes. However, the cutting wedge and its geometrical parameters, such as cutting edge radius of rounding and corner radius have the sizes comparable to tool defects [12]. Therefore, cutting operation modeling without notice of cutting edge radius of rounding in cases of orthogonal cutting and modeling of non-free cutting without notice of corner radius is generally not adequate. It is also necessary to consider the fact that the newly sharpened tools, strictly speaking, are not able to cut at all. The tools are worn beginning from the first seconds of cutting operation. Chipping and cratering are often observed near cutting edges; tool faces and flanks change their shape because of tool wear [13]. Specialists studying tapping note that in machining using extremely worn tools an eightfold increase in torque is observed. It is obvious that such defects and wear cannot be ignored in modeling. Therefore, the problem of development CAD models and next CAE models of similar objects with view of the indicated surface defects is of current importance. An obvious solution of this problem is the use of additional canonical and spline surfaces for defects modeling. However, as is well known, traditional CAD modeling has one significant drawback. Expansion in the number of surfaces, which are mated in one space area and the high complexity of surface forms, modeling of surface intersections is very difficult and often becomes impossible. In this case, CAD software sends a message that the design is unrealizable. A large variety of similar problems is not solved [14]. In our opinion, such a problem is generally solvable. The graphics core of each CAD software is constantly being improved. However, there is another more significant problem. It arises because of continuous increase of defects and of wear of the rubbing surfaces. For example, in metal cutting small particles are pulled out from tool surface and edges at every single moment. The sizes of these particles are relatively small, and the description of each of them using canonical or spline surfaces adds great difficulty to the tool modeling and to the cutting process simulation. In our opinion, problem discretization is the effective solution approach. As is known, development of numerical calculation methods like finite differences and finite elements method made greatly easier solution of the problem of elasticity and heat conduction calculation [15]. It seems that it is expedient to perform the

Geometric Modeling of Macro Defects of Parts Surfaces …

201

geometric modeling of complex objects using similar discretization. The basis of this approach is voxel (early terms: cell, receptor) modeling [16, 17]. As applied to CAD systems based primarily on classical solid-state modeling, this discretization will result in the development of discrete solid-state modeling [17, 18]. Voxel sets are presently used to describe various bodies and surfaces, starting from small objects of medicine and up to global objects of cartography and landscapes. In metal cutting, voxel models also widely used [18–21]. However, the modeling of tool chipping and cutting process modeling in combination with wear modeling is still impractical [18]. CAD modeling of construction part surfaces with micro- or macro-defects is hardly early used. The use of voxels for geometric modeling of the wear process as modeling of voxels removal from the object is a trivial task. Since voxels, often represented as cubes, are easily transformed into finite elements, then physical modeling of any process at any stage of wear and chipping is easily realizable. These approaches have been taken in the present studies. The modeling of component parts with surface defects is one of the research stages. Such type of modeling as part of voxel discrete solid modeling is considered further.

2 Theoretical Issues of Component Part Surfaces Modeling in View of Surface Defects 2.1

Determination of Component Part Surface Types and of Surface Defects Types

One of the first issues of modeling surface defects is their systematization. Obviously, it is necessary to systematize and classify both defects and surfaces. The classification of surfaces should be made according to the stages of their production. Such surfaces are primarily divided into surfaces obtained in (1) blank production, (2) in material removal operations-cutting and physical-technical processing methods, and (3) finishing operations. At present, due to the widespread use of relatively new processes (4) of additive technologies, systematization renewing is required. Blanking operations include: (1.1) casting, (1.2) rolling, (1.3) forging, (1.4) stamping, (1.5) cutting, including gas cutting. The operations associated with the chip removal include many common process operations such as turning, milling, drilling, grinding, and others. Using this set of operations, most of part surface types are formed. However, in the first approximation, all operations be selected into: (2.1) rough cutting operations, (2.2) semi-rough cutting operations, and (2.3) finishing cutting operations. Processes (2.4) of electro-discharge machining, (2.5) electrochemical machining, and others can be classified separately. Finishing operations (3.1) honing, polishing; (3.2) roll forming; (3.3) coating, etc. should be placed in a separate group. Accordingly, additive technologies (4) can be classified. If necessary, further systematization according to the degree of surface roughness can be done. Analogous systematization according to types of workpiece materials

202

I. A. Shchurov

can be done. In particular, composite materials should be highlighted. Such a systematization according to the technology of surfaces processing and to materials used, and not according to geometric features of surfaces, is related to the fact that each technology has specific types of defects, defects sizes, amount of defects, and other features (specified, for example, by GOST 21014-88, GOST ISO 6157-1-2015, GOST ISO 6157-2-2015, etc.). Analogous systematization is also accepted for series of roughness standards. Thus, it seems expedient to collect sets of defects for each technology by scanning workpiece surfaces using office scanners. Further processing of raster images and obtaining of voxel sets for each defect has been described in previously published papers [17, 18]. In accordance with the previously proposed method, graphic images of defects like holes and large scratches, as well as abrasive grains or particles of composite fillers are initially represented as sets of pixels on the black-and-white screen of the scanned image. Further, on the basis of statistical data, such two-dimensional images are transformed into three-dimensional images using certain mathematical procedures. These images represented as sets of voxels ffVi gd g, each of which corresponds to one defect d. These collections of defect voxel sets ffVi gd gS can be associated with the types of surfaces noted above. Each of defect voxel sets ffVi gd g forms a representative volume element of this defect RVEd . Each representative volume element (RVE) must have its own local coordinate system and reference point, which is the system origin [17, 18]. Accordingly, RVE location and orientation are uniquely determined by the location of the position of the reference point and the direction of the axes of the local system in some coordinate system of the part surface.

2.2

Modeling of Defects Locations on the Part Surface

It is obvious that the development of construction part geometric model, which describes surface defects, should be simple and automated like ordinary solid modeling using CAD system. Such automation can be executed in the same way that is used by the software for selecting the grade of the part. Similarly, the designer must choose a method for surface processing and, if necessary, specify its roughness class. In this case, sets of RVE defects voxels will be imported from the collection of typical defects for the specified surface and number of defects per unit area N will be determined. Thus, the problem of defects modeling is reduced to determining defects locations on the surface and their orientation along the coordinate axes of the surface. It is obvious that as the first way of defects locations modeling the random arrangement may be accepted. The problem of orientation of the coordinate systems of defects may be solved in the same way. In course of the part modeling using classical solid-state modeling the problem is solved trivially, since the equations of

Geometric Modeling of Macro Defects of Parts Surfaces …

203

the surfaces and their bounding lines are calculated by CAD systems automatically. However, when using discrete solid-state modeling, which is based on algebra of sets, this problem requires its own clarification. As known, when using discrete solid-state modeling, operations with sets are the main mathematical operations: Bv ðtÞ ¼ BSv n

t X

fBTv ðsÞg;

s ¼ 0; . . .; t

ð1Þ

s¼0

where Bv ðtÞ

voxel set ðv ¼ 1; . . .; V Þ of modeled object body at the current stage of modeling t and/or to the current time t; BSv voxel set of modeled object body with the use of canonical surfaces at the initial time (primary blank); BTv ðsÞ voxel set of modeled object which is removed from the initial object at the current stage of modeling s and/or at the current time s; Pt operator of union of sets of objects bodies removed from the initial body s¼0 starting from the beginning of the simulation to the current time (analogous to the operation of sum of sets U); “\” operator of difference of sets. Any combination of the operators U, \ , “\”, and any collection of objects used for the generation of the required body is possible. Obviously, using this approach, the only one set of voxels, which is the body of the ready-made part, is finally obtained. This set does not contain information about belonging of voxels to the specified surface. Therefore, the distribution of defects voxel sets on the part surfaces is difficult. This difficulty can be eliminated in the case of determining not only the sets of voxels of the part’s body, but also the sets of voxels, which belong to each of its surfaces separately. In earlier published papers, the relation, which enables to select a part of object surface voxels from the set of the whole body voxels, has been presented [17, 18]. However, in this case, the entire surface of the object is defined as one complex structure. To isolate the sets of voxels of each elementary surface, the following approach is proposed. Each voxel in accordance with its definition [17] has a state parameter—Pv , ðPv 2 ZÞ. It is assumed that Pv  0 if this voxel does not belong to the object, in this case Pv corresponds to the number of the simulated environment (vacuum Pv ¼ 0, gas Pv ¼ 1, liquid Pv ¼ 2, etc., in general: Pv ¼ PEv  0). It is also assumed that Pv [ 0, if this voxel belongs to the object, in this case Pv corresponds to the number of the modeled surface (Pv 2 N) in this case. In accordance with the fourth principle of discrete solid-state modeling [18, 22], such modeling is carried out in the order and in accordance with the manufacturing technology of the object. In this case, it is assumed that all voxels of the body of the primary blank have Pv ¼ 1 number.

204

I. A. Shchurov

To determine all voxels of the surface of the primary blank, the following relationship is used to calculate their state parameters: Pi;j;k ¼ 1jfPi þ 1;j;k ¼ PEv \ Pi1;j;k ¼ PEv \ Pi;j þ 1;k ¼ PEv \ Pi;j1;k ¼ PEv \ Pi;j;k1 ¼ PEv \ Pi;j;k1 ¼ PEv g

ð2Þ

where i; j; k—integer coordinates of the nodal point of the current voxel [17] in the coordinate system of the modeled body: imin  i  imax ; jmin  j  jmax ; kmin  k  kmax , limited by the corresponding overall coordinates (overall dimensions). To implement the first operation of the algebra of sets for s ¼ 2, for which a set of voxels Bv ðtÞjt ¼ s is generated, all the state parameters of the set BTv ð2Þ are determined. Such parameters may have PTv ¼ 1002 number, for example. Then the state parameters of the surface of the original object Bv ð2Þ, formed by the difference with the set BTv ð2Þ, are determined by the relations: Pi;j;k ¼ 2jfPi þ 1;j;k ¼ PTv \ Pi1;j;k ¼ PTv \ Pi;j þ 1;k ¼ PTv \ Pi;j1;k ¼ PTv \ Pi;j;k1 ¼ PTv \ Pi;j;k1 ¼ PTv g

ð3Þ

After determining the state parameters the generated surface, the operation is performed according to the formula (1). Similarly, the remaining operations of the set difference are performed and, each time, a new number of the state parameter of the generated surface is determined. Thus, by the time of modeling finishing all the state parameters of the body voxels of the original object and the state parameters of the voxels of all its surfaces are determined. Having determined the voxels of the surfaces of the object, one can determine the coordinates of their nodal points. Accordingly, any three points of the surface (of the specified voxel and two adjacent voxels) enable to find the surface normal. Defect location at a specified surface point should be made by equating the coordinates of the defect RVE base point to the coordinates of this point. Z axis of the RVE coordinate system must match the part surface normal. After that, all coordinates of the defect voxel nodal points are recalculated by using the base rotation matrix around the RVE Z axis at a certain angle n. Obviously, the values determined as the result of a random selection of surface voxel are used as the coordinates of the points where defects are located. It is about the coordinates of the nodal point of such a voxel mentioned above. The number of such voxels is determined by the specified density N and by the surface area SQ. Knowing the sizes of the voxels, and, therefore, the area of each voxel face; knowing the number of voxels of the specified surface, it is easy to calculate the required area. Similarly, the angle value n is determined using a random number generator. Thus, the above method enables to produce discrete solid-state modeling of objects using algebra of sets, to determine the surfaces of these objects and to locate the required defects on the surfaces in the required amount in the automated mode. The intersection of

Geometric Modeling of Macro Defects of Parts Surfaces …

205

these defects on the edges and vertexes enables to model tool cratering and chipping at such edges and vertexes. To verify the proposed relations and methods of discrete solid-state modeling, a computer program has been developed and corresponding modeling has been executed. The results of this simulation are presented below.

3 Results of Computer Discrete Solid-State Modeling of Component Parts with Defects on Their Surfaces To obtain sets of defects voxels, the previously developed method for obtaining sets of abrasive material grains voxels and of sets of particles of disperse-reinforced composites voxels using scanning of grains and particles with an office scanner has been applied [17]. Similarly, digital photographs of real parts have been obtained. The contours of defects have been selected and their 3D images have been obtained. Such images are represented by their voxel sets. On an enlarged scale, these defects are shown in Fig. 1a. Using a graphic editor such as Photoshop and its «Magic Wand Tool», the boundaries of surface defects can be identified and their contours can be recorded on the image second layer, Fig. 1b. Next, modeling methods similar to methods of composite structures modeling, which use statistical data to obtain three-dimensional images have to be applied [22]. As a result, voxel sets for each defect are obtained, Fig. 1c, d. By superimposing of voxel sets of the component part and of its defects Fig. 1e, and by using the difference operation of the sets, the voxel mesh with defects is calculated, Fig. 1f. As shown in the figure, the model of the surface with defects is sufficiently adequate.

Fig. 1 Photo of a fragment of the shaft surface with defects (a); photographs of the defects and their contours (b) of a plurality of voxels of defects: front (c) and oblique (d) views; the same sets with part body sets (e) and voxels of a part with surface defects (f)

206

I. A. Shchurov

4 Conclusion The developed mathematical relations of discrete solid-state modeling of component part bodies and surfaces based on algebra of sets and the previously developed method of obtaining voxel sets of volume fragments based on scanning with a flat-bed scanner enable to develop adequate models of surfaces of parts with defects, including finite element models. (1) To automate the process of 3D component parts modeling based on a discrete solid-state type of modeling, it is necessary to have collections of defects voxel sets, which correspond to special types of surfaces, as components of CAD systems. (2) To automate the selection of defects sets, it is necessary to develop the structured tree of surface types. These surfaces have to be systematized according to the stages of production, processing methods, degrees of roughness, and materials of component parts. Each surface must be associated with the corresponding collection of voxel sets of defects.

References 1. Xu X, (2009) Integrating advanced computer-aided design, manufacturing, and numerical control: principles and implementations. IGI Global, p 424 2. Stroud I, Nagy H (2011) Solid modelling and CAD systems. Springer, p 699 3. Cederfeldt M, Sunnersjo S (2003) Solid modelling with dimensional and topological variability. In: International conference on engineering design ICED-03, Stockholm, pp 1–10 4. Congli W (2004) Representation of curves and surfaces in b-rep solid modelers. Symp Ser Mechatron 1:498–507 5. Buric M, Marjanovic D (2018) A tool for idealisation of cad models. In: International design conference—design 2018. https://doi.org/10.21278/idc.2018.0367 6. Voicu R, Vintila S, Vilag V, Mihalache R (2016) CAD modelling: light weight composite centrifugal rotor manufacturing for energy efficiency. Mater Plast 53(4):623–625 7. Maurel-Pantel A, Fontaine M, Thibaud S, Gelin JC (2012) 3D FEM simulations of shoulder milling operations on a 304L stainless steel. Simul Model Pract Theory 22:13–27 8. Abena A, Soo SL, Essa K (2017) Modelling the orthogonal cutting of UD-CFRP composites: development of a novel cohesive zone model. Compos Struct 168:65–83 9. Saffar RJ, Razfar MR, Zarei O, Ghassemieh E (2008) Simulation of three-dimension cutting force and tool deflection in the end milling operation based on finite element method. Simul Model Pract Theory 16:1677–1688 10. Pittala GM, Monno M (2011) A new approach to the prediction of temperature of the workpiece of face milling operations of Ti-6Al-4V. Appl Therm Eng 31:173–180 11. Coromant S (2005) Metal cutting technical guide milling. p 186 12. Coromant S (2017) Turning tools. p 848 13. Coromant S (2017) Metal cutting technology training handbook, p 391 14. Rossignac JR (1986) Offsetting operations in solid modelling. Comput Aided Geom Des 3:129–148 15. Segerlind LJ (1984) Applied finite element analysis. Wiley, p 222

Geometric Modeling of Macro Defects of Parts Surfaces …

207

16. Bouhadja K, Bey M (2014) Classification of simulation methods in machining on multi-axis machines. Lect Notes Eng Comput Sci 2:992–997 17. Shchurova CI (2015) A methodology to design a 3D graphic editor for micro-modeling of fiber-reinforced composite parts. Adv Eng Softw 90:76–82 18. Shchurov IA (2004) Teoriya rascheta tochnosti obrabotki i parametrov instrumentov na osnove diskretnogo tverdotel’nogo modelirovaniya (The theory of precision machinery and tools parameters calculation on the base of discrete solid modelling), SUSU, Chelyabinsk, p 320. Accessed 15 Jan 1999 http://lib.susu.ru/ftd?base=SUSU_METHOD&key= 000436340&dtype=F&etype=pdf 19. Hauth S, Murtezaoglu Y, Linsen L (2009) Extended linked voxel structure for point-to-mesh distance computation and its application to NC collision detection. Comput Aided Des 41 (12):896–906 20. Blasquez I, Poiraudeau JF (2004) Undo facilities for the extended z-buffer in NC machining simulation. Comput Ind 53:193–204 21. Ilushin O, Elber G, Halperin D, Wein R, Kim M-S (2005) Precise global collision detection in multi-axis NC-machining. Comput Aided Des 37:909–920 22. Shchurova EI (2018) Voxel and finite element modeling of the ceramic-polymer composite panel for ballistic impact description. Lect Notes Mech Eng. ICIE 2018:277–284. https://doi. org/10.1007/978-3-319-95630-5_30

Selecting Criterion of Long-Term Strength in Assessing Durability of Constructions Operating Under Non-isothermal Loading Processes A. V. Belov, A. A. Polivanov and N. G. Neumoina

Abstract The paper summarizes one the approach to the describing the Pisarenko-Lebedev generalized long-term strength criterion, which is used in determining the equivalent stress in the Yu. Rabotnov’s kinetic equation of damageability. Finding the equivalent stress is one of the important yet difficult-to-solve problems. There are multiple long-term strength criteria, but no precise recommendations on their use, which is what makes this problem so difficult to solve. The most reasonable and universal criterion is generalized Pisarenko-Lebedev generalized long-term strength criterion. The authors proposed a simple and easy-to-use method to find the equivalent stress when using the Pisarenko-Lebedev generalized strength criterion. This method for concretizing the Pisarenko-Lebedev generalized long-term strength criterion, in the first approximation accounts for alterations in the plastic properties of the material as a function of the loading time, temperature, and stresses, enabling the research to account for the prevalence of viscous or brittle fracture, which in its turn enables more accurate damage and time-to-rupture predictions. More detailed experimentation to test this criterion and its concretization method requires further research based on in-site experiment data. Keywords Long-term strength criterion Stress-strain state


High-temperature creep
Damage


1 Introduction Ensuring the long-term strength of various structures subjected to non-isothermal loading is an important problem of the mechanics of deformable solids. To assess the long-term strength of such structures, one must find the accurate history of their

A. V. Belov
A. A. Polivanov (&)
N. G. Neumoina Kamishin Technological Institute (Branch) the Volgograd State Technical University, 6, Lenin St, Kamyschin 403874, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_23

209

210

A. V. Belov et al.

stress-strain state, taking into account the occurring elastic and plastic strain, the creep, and the subsequent degradation of mechanical properties. With the calculation output at hand, it is possible to identify dangerous zones in the first approximation and to use the criteria of long-term strength to assess the safe life of the structure [1]. However, such long-term strength assessment of a structure in a complex and inhomogeneous stress state is going to be rather approximate. Accurate determination of the strength of a structure in a complex inhomogeneous stress state requires a generalized statement of problem, based on the use of continuous fracture mechanics proposed by Rabotnov [2, 3]. This approach implies adding the scalar parameter of creep-related damageability of the material to the governing equations describing the stress-strain state of a structure. To describe the processes of damage accumulation in the material, the kinetic equation of damage proposed by Rabotnov [2] is used:
Q req dxC ¼C dt 1  xC

ð1Þ

where xC is the scalar damage parameter; C and Q—parameters dependent on temperature and determined from long-term strength curves obtained from uniaxial tension tests of standard samples at fixed temperatures; req—equivalent stress, which is one of the criteria for long-term strength. As a rule, such problems are solved by numerically. The process of loading a structure is divided into short intervals, while the structure itself is divided into small elements such that their stress state becomes homogeneous. These data must then be used to describe the interval-specific stress-strain state and the degree of damage to the structural elements. The strain history can then be traced from interval to interval until rupture.

2 Research Goal The goal hereof is to: • Analyze the known approaches to describing the Pisarenko-Lebedev generalized long-term strength criterion, which is used in determining the equivalent stress in the Rabotnov kinetic equation of damageability; • Choose a method for concretizing the Pisarenko-Lebedev generalized long-term strength criterion; • Choose a concretization method and validate the results of simulating the stress-strain state of a flat plate as obtained by using this criterion.

Selecting Criterion of Long-Term Strength in Assessing …

211

3 Materials and Methods Finding the stress-strain state of a structure in the context of creep-related damageability of material is a problem covered in numerous papers, e.g., [4–6]. Finding the equivalent stress is one of the important yet difficult-to-solve problems. There are multiple long-term strength criteria, but no precise recommendations on their use, which is what makes this problem so difficult to solve. The most reasonable and universal criterion is generalized Pisarenko-Lebedev generalized long-term strength criterion [7]: req ¼ vri þ ð1  vÞ
r1

ð2Þ

here req is the equivalent stress; v—material plasticity coefficient [7], ri—stress intensity, which, in a plane stressed state, is determined by the expression pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ri ¼ r21  r1 r2  r22 , r1 and r2—main normal stresses. When assessing the long-term strength of a structure, the corresponding v parameter can be found by testing the long-term strength of solid cylindrical specimens subjected to uniaxial tension and compression, or by torsional testing of thin-walled tubular specimens. These characteristics are calculated by the formulas [7, 8]:  C v ¼ rTLT rLT

ð3Þ

  pffiffiffiffiffiffiffiffiffiffiffi 31 v ¼ rTLT sLT  1

ð4Þ

where rTLT is the long-term tensile strength limit; rTLT is the long-term compression strength limit; sLT is the long-term net shear (torsion) strength limit. The problem of applying this long-term strength criterion is the lack of experimental data on the long-term compressive and torsional strength of materials. In papers [6, 9], the parameter v is a function of the intensity of tangential stresses and temperature; as such, it is found from the condition that long-term strength diagrams must coincide for uniaxial tension r(t*) and pure torsion s(t*), whereby is found by the expression: v¼

rs

1

pffiffi  2 3 r 1



qffiffi 2 3 s

ð5Þ

where the stresses r and s are picked for the same time point t*. For the selected time point t*, the corresponding average mean intensities of the tangential stresses S are found from the diagrams of long-term tensile strength and pure torsion. This

212

A. V. Belov et al.

procedure is repeated for different values v of at different fixed temperatures T. This is done to obtain the dependence v = f (S, T), which is further used to find the equivalent stress. Apparently, this method implies complex and prolonged tensile and purely torsional testing of tubular specimens, which sometimes might be impractical. The authors hereof believe that the Pisarenko-Lebedev long-term strength criterion can be used to account for changes in the material damage types if the parameter is assumed to be a function of stresses, temperature, and creep duration. In [10], the authors proposed a simple and easy-to-use method to find the v parameter when using the Pisarenko-Lebedev generalized strength criterion. In this case, the v parameter for this or that material is deemed to equal the relative residual contraction at break w. A similar approach is proposed for finding the v parameter when making long-term strength calculations. To that end, denote this parameter as vC and assume that vC = wC where wC—is the relative residual contraction of a solid round specimen, obtained by testing it for creep and long-term strength at a constant stress and temperature. Tests for creep and long-term strength of standard cylindrical metal/alloy specimens made reveal the following pattern: at high stress levels significantly exceeding the ultimate strength of the material, the intragranular fracture (viscous fracture) prevails, accompanied by localized intensive creep, whereby a neck emerges. The relative residual contraction of the specimen, obtained by creep tests, is nearly identical to the relative residual contraction value obtained by the standard tensile test of solid cylindrical specimens, as they are instant-strained to rupture at the same temperature. At the same time, if the stress is low and barely exceeds the creep strength of the material, intergranular strain and fracturing prevail, with the specimen being intensively embrittled without forming a neck; fracturing occurs at low accumulated creep over a long time. Relative residual contraction is nearly zero. Based on these observations and on the first-approximation assumption that the relative residual creep contraction of a fractured specimen is linearly dependent on the tensile stresses, one can construct a simple dependence for fixed temperatures. In this case, the relative residual constriction wC will take values close to zero. Based on these observations and assuming that the relative residual constriction during creep of the sample during fracture linearly depends on the tensile stresses level at which the tests take place, it is possible to construct a linear relationship for fixed temperatures wC = f (r): wC ¼ 

r w rS  r0:2=105

ð6Þ

where wC

is the relative residual contraction of a standard specimen of circular cross-section, obtained by testing for creep and long-term strength at a constant stress r at a fixed temperature T;

Selecting Criterion of Long-Term Strength in Assessing …

W rS r0:2=105

213

is the relative residual contraction, obtained by stretching a standard specimen of circular cross-section at its instantaneous deformation at the same temperature; is the material strength limit; is the material creep limit at a given temperature.

For most structural materials, reference books contain all data necessary for applying the ratio (6) in long-term strength calculations [8]. Figures 1 and 2 show the relative residual contraction at creep as a function of stress (the solid line); these curves were plotted by the proposed method for steel 20 and for 12X18H9T austenite steel. Besides, the same curves show (in dark circles) the wC experimental values for the same materials as obtained for various normal stresses at the same temperatures [8]. The curves reveal that experimental data match very well the calculations by the proposed method. Fig. 1 Dependence of the relative residual narrowing of steel 20 at creep from normal stress at 500 °C temperature

,%

80

=76

60

40

20

0

Fig. 2 Dependence of the relative residual narrowing of nickel alloy at creep from normal stress at 500 °C temperature

, MPa 0,2/10 5=

25

50

100

150

200

S= 220

250

60 =52

40

20

0

5= 0,2/10 ,

30

100

200

S=

, MPa 280 300

214

A. V. Belov et al.

Thus, the approach proposed herein implies using the following formula to find the element and loading interval-specific equivalent stress: req ¼ wC ri þ ð1  wC Þ
r1

ð7Þ

here, wC is found as per (6). The statement of problem is complemented with the following mechanical properties of the material as observed at fixed temperature within its normal operating range for this structure: ultimate strength, creep strength, and relative residual contraction.

4 Experimental Results To verify the adequacy of the proposed method of concretizing the PisarenkoLebedev generalized long-term strength criterion, the research team tested a single-layer plate with a circular hole, heated uniformly to 500 °C and exposed to pressure on the top, see Fig. 3. The steel 20 plate had the following dimensions: R = 110 mm, r = 65 mm, d = 10 mm. Multiple calculations were run at pressures of 7–12 MPa; however, pressure was never changed within a single test. The time to fracture as a function of pressure was studied in terms of four different long-term strength criteria, see Fig. 4 for the main results. Figure 4 shows the curves of the time to fracturing onset as a function of pressure within 9–12 MPa; the curves were plotted using four different long-term strength criteria. Notably, using the Johnson criterion produces the minimum estimated time to rupture (brittle fracture), and while the Katz criterion returns the maximum estimated time to rupture (plastic fracture) [11]. Thus, using any long-term strength criterion will produce a value within the range limited by the Johnson and Katz values. The Sdobyrev criterion is a half-sum of these two, and the Sdobyrev curve is near the middle. As the equivalent stress is taken (Fig. 4): 1. maximum principal stress (the Johnson’s criterion): req ¼ r1 ; pffiffiffi 2. stress intensity (the Katz’s criterion): req ¼ S 3; pffiffiffi  3. the Sdobyrev’s criterion: req ¼ 0; 5
r1 þ S 3 ;

P

Z r R

Fig. 3 Axial section of the lamina

S

Selecting Criterion of Long-Term Strength in Assessing …

215

Fig. 4 Dependence of the time of local destruction of the plate on pressure

4. the Pisarenko-Lebedev’s long-term strength generalized criterion: req ¼ pffiffiffi wC S 3 þ ð1  wC Þ
r1 . The finding here is that the chosen long-term strength criterion affects the estimated time to rupture, with the effect being greater for lower loads and longer loading. The proposed way to concretize the Pisarenko-Lebedev generalized long-term strength criterion allows to account for the altered plastic properties of the material (the embrittlement); such alteration occurs as creep develops. Figure 4 shows that the curves 3 (the Stobyrev criterion) and 4 (the Pisarenko-Lebedev generalized criterion) are nearly identical in coordinates and intersect at 10.2 MPa. Stresses affecting the most loaded points of the plate at such pressure are about 150 MPa, or w = 50%. Further, curve 4 approaches curve 2 (the Katz criterion), which means viscous fractures rather than brittle fractures prevail. At lower pressures, the Pisarenko-Lebedev criterion is close to the Johnson criterion results-wise. For instance, at 7 MPa time to fracture is 10,000 h per the Johnson criterion, 70,000 per Katz, and 12,500 per Pisarenko-Lebedev. Analysis of the stresses in the most heavily loaded points shows such stresses do not reach the ultimate strength, meaning that brittle fractures rather than viscous fractures will prevail. Difference in time to rupture at 7 MPa is vastly different for the Johnson criterion and the Katz criterion: It is about 60,000 h or 7 times. This is why using non-modified criteria returns unreliable results.

5 Conclusions The proposed method for concretizing the Pisarenko-Lebedev generalized long-term strength criterion in the first approximation accounts for alterations in the plastic properties of the material as a function of the loading time, temperature, and

216

A. V. Belov et al.

stresses, enabling the research to account for the prevalence of viscous or brittle fracture, which in its turn enables more accurate damage and time-to-rupture predictions. More detailed experimentation to test this criterion and its concretization method requires further research based on in-site experiment data, with some of such experiments being discussed in [12]. The research team behind this paper plans to carry out such research further on.

References 1. Volegov PS, Gribov DS, Trusov PV (2015) Damage and destruction: classical continual theories. Phys Mesomech 18(4):68–87 2. Rabotnov YN (1966) The creep of structural elements. Science, Moscow 3. Khokhlov AV (2017) Criterion of destruction and long-term strength curves, generated by the defining relation of the nonlinear theory of heredity Yu. N. Rabotnov. Her Mech Eng 6:39–46 4. Aliev MM, Shafieva SV, Gilyazova SR (2015) Criterion of long-term strength for materials of different resistance. Mater Sci Sess Sci Almetyevsk State Oil Inst 1(1):254–257 5. Arutyunyan RA (2015) High-temperature embrittlement and long-term durability of metallic materials. In: Proceedings of the Russian academy of sciences. Solid mechanics, vol 2. pp 96–105 6. Shevchenko YN, Terekhov RG, Braikovskaya NS, Zakharov SM (1994) Investigation of the processes of destruction of body elements as a result of material damage during creep. App Mech 30(4):21–30 7. Lebedev AA (2010) The development of theories of strength in the mechanics of materials. Strength Probl 5:127–146 8. Pisarenko GS, Yakovlev AP, Matveev VV (2008) Materials resistance handbook. Delta, Kiev, p 816 9. Galishin AZ (2004) Determination of thermo-viscoplastic state of shells of revolution taking into account the material damageability in creep. Appl Mech 40(5):71–79 10. Belov AV, Neumoina NG (2014) On the use of Pisarenko-Lebedev generalized strength criterion in strength calculations for non-isothermal loading processes. Int J Appl Basic Res 9:8–10 11. Polivanov AA (2004) Axially symmetric elasto-plastic deformation of multilayered shells of rotation with an account of material damage at a creep, abstract of dissertation, Volgograd 12. Belov AV (1989) Axially symmetric elasto-plastic stressed–strained state of shells of rotation with an account of material damage at a creep, abstract of dissertation, Kiev

Virtual Prediction of Accuracy of Processing on Example of External Circular Grinding P. P. Pereverzev, A. V. Akintseva and M. K. Alsigar

Abstract Virtual prediction of processing accuracy is an actual task not only for modern mechanical engineering, but also for creating effective production cyber-physical systems based on the concept of “Industry 4.0.” Prediction of the accuracy is possible to implement by using the method of calculating the processing error and the model of metal removal (presented in this article in more detail) which is a model of grinding surface forming, taking into account the features of processing in reverse and non-reverse zones and allowing to calculate the current values of the radii in any section of the processing surface during the whole grinding cycle for the given processing conditions. The model of surface dimensions is constructed on the basis of the calculated values of the processing surface radii in order to estimate the errors of a diametrical size and shape error and location of the surfaces with a simultaneous evaluation of circular grinding cycle productivity. The model of metal removal for the circular external grinding cycle, described in this article, can be used not only for prediction of the processing accuracy for a given processing cycle, but also for designing a speed-optimal cycle, i.e., the model is a basis for development of an optimal cycle creation methodology. Keywords External grinding accuracy


Model of metal removal
Error
Processing

Nomenclature ri Average stress intensity, N/mm2 d Workpiece diameter, mm D Wheel diameter, mm V1 Periphery speed of wheel, m/s P. P. Pereverzev
A. V. Akintseva (&) South Ural State University, 76, Lenin Avenue, Chelyabinsk 454080, Russia e-mail: [email protected] M. K. Alsigar College of Engineering University of Dhi Qar, 31, Dhi Qar, Nasiriyah 3629, Republic of Iraq © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_24

217

218

V2 Vsoc tpk,i,z k z i B η c Rwk tpk,i−1,z Rmax

P. P. Pereverzev et al.

Workpiece rotational speed, m/min Speed of wheel axial speed, mm/min Program value of radial component of the cutting force, mm/double stroke Ordinal number of workpiece radius Ordinal number of cycle stage Ordinal number of tool stroke on zth stage of cycle Total grinding wheel height, mm Bluntness ratio of wheel Technological system compliance, N/mm Initial radius of workpiece on kth radius, mm Program feed per one workpiece revolution, mm Maximal workpiece radius, mm

1 Introduction There are high demands for processing accuracy to grinding operations which normalize not only the error of diametrical dimension but also the shape error and location of the surfaces. Ensuring the stability of processing accuracy for a batch of workpieces is a difficult task due to some reasons. Firstly, the presence of elastic deformations of the technological system leads to the fact that the program radial feed is significantly different from the actual feed [1]. Secondly, the presence of reverse zones, which are characterized by the instability of the allowance removal process in the input and output workpiece sections due to the instantaneous change of different stages and types of grinding, which stipulates the different processing accuracy along the entire length of the workpiece [2]. Thirdly, the presence of unsteadiness of technological process, both, during the part processing cycle (variable contact area of workpiece and wheel, variable cutting modes in reverse zones) and during the batch of workpieces processing (wheel grain blunting, allowance fluctuation, initial radial run-out of the workpiece, changing of grinding wheel diameter) [3]. Let us analyze what scientific achievements and developments are available for the problem considered in this article. The impact of cutting force and elastic deformation of the technological system on the processing accuracy for different metal-cutting systems has been confirmed by scientific researches [4, 5]. There are many scientific works in the field of improving the quality of processing an automatic grinding cycles [6–12]. However, scientific researches have not yet considered models predicting the processing accuracy in the circular grinding cycles, which would take into account unsteadiness of the technological process. Building of circular grinding cycles with axial feed is considered by many authors [13–16], but in these works and many others the features of cutting forces calculation and

Virtual Prediction of Accuracy of Processing on Example …

219

processing accuracy for grinding cycles with axial feed in reverse and non-reverse zones of machined surface are not considered. Thus, the review of literary sources revealed that despite the enormous theoretical and practical scientific base in the field of improving the accuracy and productivity of grinding operations, performed on CNC machines in automatic cycles, there are still no scientific developments of analytical wide range models of cutting forces, models of allowance removal from the machined surface and methods of production accuracy predicting. Consequently, there is no theoretical possibility of creating production cyber-physical system on the basis of the concept “Industry 4.0” [17], as it is not possible to solve a task of virtual designing of optimal grinding cycles and predicting the processing accuracy proved by those cycles without models [18]. To improve the stability of processing accuracy indices in the batch of parts without loss of productivity in the concept “Industry 4.0” [17] on example of the external circular grinding the model of metal removal has been developed. The model allows predicting the process of machined surface forming along the entire length of the workpiece throughout the cycle. The metal removal model describes the change in the radii dimensions of the machined surface along the entire length of the workpiece from the initial value in the workpiece to the final value of the finished part at the end of the grinding cycle.

2 Virtual Prediction of Processing Accuracy in View of Unsteadiness of Technological Process The basis of the metal removal model is the force model, which establishes the correlation between the cutting force and cycle parameters through the elastic deformations of the technological system. Analysis of the literary sources has revealed that there are many works devoted to the development of cutting force analytical model in the process of grinding [19–22]. However, there are no analytical regularities of change in cutting force and forming of grinding surface in automatic cycles when controlling two feeds (radial and axial) in these and many other works. Also, the impact of variable technological factors on the cutting force during the grinding cycle is not considered. Therefore, the force model of circular external grinding, which is based on the principle of equality between works of operating cutting forces and the resistance forces of machined metal to plastic deformation, is considered [20]. More information about the force model of circular external grinding can be found in the article [23], for internal grinding [24]. This article presents only the model of radial component of the cutting force, which is needed for the development of metal removal model:

220

P. P. Pereverzev et al.

PYk;i;z

1:86ri pdVSoc tpk;i;z ri gB ¼ qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi þ 3 2 2 ðV1 þ V2 Þ þ VSoc

rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi dDtpk;i;z dþD

ð1Þ

Consider the modeling of metal removal in the middle section of non-circular workpiece, taking the form of section as ellipse, described by the kth number of radii. On the ith revolution from the radius Rk,i−1,z the metal layer Δtfk,i−1,z is ground of (Fig. 1); the layer corresponds to the cutting depth or the value of actual feed per one revolution of the part. Therefore, the upper trajectory of the graph shows the change in the value of part of allowance, grounded by i workpiece revolutions or accumulated actual feed by i revolutions. Also the upper trajectory shows the current change in the radius Rk,i,z of the workpiece after ith revolution, i.e., Rk;i;z ¼ Rk;i1;z  Dtfk;i1;z

ð2Þ

Fig. 1 Computed pattern of connection of feeds, elastic deformations, and current radii during circular grinding of the workpiece middle section (non-reverse zone)

Virtual Prediction of Accuracy of Processing on Example …

tpk;i;z ¼ yk;i;z þ Dtfk;i;z þ tfk;i;z þ DRk tpk;i;z ¼

k X i X z X 1

tfk;i;z ¼

1

1

ð3Þ

Dtpk;i;z

ð4Þ

Dtfk;i;z

ð5Þ

1

k X i X z X 1

221

1

yk;i;z ¼ cPYk;i;z

ð6Þ

Rk;i;z ¼ RWk  tfk;i;z

ð7Þ

DRk;i;z ¼ Rmax  RWk

ð8Þ

Solving the set of obtained formulas (1–8) about actual feed Δtfk,i,z, we get the following formula, which allows to calculate the current values of actual feed Δtfk,i,z on each radius and revolution of the workpiece during the whole grinding cycle:

Dtfk;i;z

2sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 32  2 cK tp  tf  DR K c 2 k;i;z k;i1;z Wk 2 5  ¼4 þ 2ð1 þ K1 cÞ 2ð1 þ K1 cÞ 1 þ K1 c

ð9Þ

K1, K2—formula coefficients: 1:86ri pdVSoc K1 ¼ qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 ðV1 þ V2 Þ2 þ VSoc K2 ¼

ri gB 3

rffiffiffiffiffiffiffiffiffiffiffiffi dD dD

ð10Þ

ð11Þ

Derived formula (9) allows to calculate the actual radial feed on each radius on ith stroke of zth cycle stage, but only in the middle section of the workpiece located in the non-reverse zone. Presence of reverse zones in the processing of input and output workpiece sections leads to significant fluctuations of actual radial feed in the reverse and non-reverse zones. Therefore, to predict the accuracy of processing along the entire length of the workpiece, the model of blanking should consider the features of allowance removal in reverse zones, namely, instantaneous change of different stages and types of processing, presence of wheel overtravel, “additional grinding” with wheel reverse due to the elastic deformations of technological system (provided that there is no longitudinal feed). For modeling of metal removal process on stages of plunging and sparking-out, caused by the time lag between reverse feed switching, the model of cutting force and formula of actual radial feed for circular

222

P. P. Pereverzev et al.

infeed grinding are used. The model of metal removal and force model for circular infeed grinding are developed in the work [1]. Thus, the complex of the formula (9) and formulas, presented in the work [1], is a wide range analytical model of allowance removal on circular external grinding, allowing calculating current values of all types of feeds, force, elastic deformations, and surface radii for any specified grinding cycle, specified processing conditions and variable technological factors. It should be noted that models of internal grinding with axial feed were obtained similarly [24, 25], but without considering the transitional stages. In the virtual prediction of processing accuracy for a given grinding cycle, a comparison of the calculated values of error indexes with the corresponding allowable values, indicated on the part pattern, is carried out. At the end of grinding cycle, the predicted values of form and location errors (deviation from roundness, redial run-out) are determined by the calculated profile of each section; and the deviation from cylindricalness, total radial run-out, etc. are determined by the total profile of the machined surface. The method of calculating the processing error is given in the article [26]. The model of metal removal for circular grinding, described in this article, is a basis in the design methodology for the optimal cycle [18].

3 Conclusion 1. Virtual prediction of processing accuracy is based on the models of cutting force and forming of grinded surface. 2. The model of surface forming allows calculating the cutting depth. Actual radial feed, cutting forces and current radii values in any section of machined surface during the whole grinding process for given processing conditions. 3. Based on the calculated values of the radii of the machined surface, an estimate of the diametrical size errors and the errors of form and location of surfaces is made. 4. Virtual prediction of processing accuracy can be implemented using the method of processing error calculation [26]. 5. The model of grinded surface forming is analytical, as obtained on the basis of mathematical interrelation of cutting force, elastic deformations of technological system, cutting depth, program and actual feeds and sizes of the radii of machined surface. 6. The cutting force model covers most of the main technological parameters in a wide range of their variation (physical-mechanical features of grinded metal, geometrical parameters of the contact area of wheel and workpiece, characteristics of the wheel, abrasive grains, grinding wheel, etc.) and considers features of metal removal in reverse and non-reverse zones, the presence of which causes a different processing accuracy along the entire length of the machined surface.

Virtual Prediction of Accuracy of Processing on Example …

223

References 1. Pereverzev PP (1999) Theory and method for calculation of the optimum cycles of processing of the parts on cylindrical grinding machines with automated control. Dissertation, University of Chelyabinsk 2. Pereverzev PP, Akintseva AV (2018) Modelling of renting of metal in reverse zones in the course of internal grinding. Russ Eng Res 38(3):229–234 3. Rowe WB (2013) Principles of modern grinding technology. Elsevier, Liverpool 4. Balakshin BS (1973) Adaptivnoe upravlenie stankami (Adaptive machine control). Mashinostroenie, Moscow 5. Bazrov BM (1984) Raschet tochnosti mashin na EHVM (EHVM machine accuracy calculation). Mashinostroenie, Moscow 6. Pereira WX, Diniz AE, Hassu A (2009) Comparing different plunge cylindrical grinding cycles based on workpiece roughness and process vibration. J Braz Soc Mech Sci Eng. https:// doi.org/10.1590/s1678-58782009000200009 7. Bratan SM (2006) Tekhnologicheskie osnovy obespecheniya kachestva i povysheniya stabil’nosti vysokoproizvoditel’nogo chistovogo tonkogo shlifovaniya (Technological basis for quality assurance and improving the stability of high-performance fine grinding). Dissertation, University of Odesa 8. Novoselov YUK (2012) Dinamika formoobrazovaniya poverhnostej pri abrazivnoj obrabotke (Dynamics of formation of surfaces during abrasive machining). Izd-vo SevNTU, Sevastopol’ 9. Zhang Y, Li B, Yang J, Liang S (2018) Modeling and optimization of alloy steel 20CrMnTi grinding process parameters based on experiment investigation. Int J Adv Manuf Technol 95 (5–8):1859–1873 10. Uhlmann E, Koprowski S, Weingaertner WL, Rolon DA (2016) Modelling and simulation of grinding processes with mounted points: part I of II—grinding tool surface characterization. Procedia CIRP 46:599–6027 11. Liu YM, Yang TY, He Z, Li JY (2018) Analytical modeling of grinding process in rail profile correction considering grinding pattern. Arch Civil Mech Eng 18(2):669–678 12. Rowe WB, Ebbrell S (2004) Process requirements for cost-effective precision grinding. CIRP Ann 53(1):255–258 13. Gao S, Yang C, Fu Y, Su H, Ding W (2017) Optimization for internal traverse grinding of valves based on wheel deflection. Int J Adv Manuf Technol 92(1–4):1105–1112 14. Leonesio M, Sarhangi M, Bianchi G, Parenti P, Cassinari A (2015) A meta-model framework for grinding simulation. Procedia CIRP 31:357–362 15. Sundarrajan KD (2015) Study of grinding burn using design of experiments approach and advanced kaizen methodology. Dissertation, University of Nebraska, Lincoln 16. Xuekun Li (2010) Modeling and simulation of grinding processes based on a virtual wheel model and microscopic interaction analysis. Dissertation, Worcester Polytechnic Institute 17. Vasja Roblek V, Meško M, Krapež A (2016) A complexity view of Industry 4.0. SAGE Open. https://doi.org/10.1177/2158244016653987 18. Pereverzev PP, Akintseva AV (2016) Optimal internal grinding cycles in multidimensional control-parameter space. Russ Eng Res 36(11):974–978. https://doi.org/10.3103/ S1068798X16110174 19. Korolev AV (1975) Issledovanie processov obrazovaniya poverhnostej instrumenta i detali pri abrazivnoj obrabotke (The study of the processes of formation of the surfaces of tools and parts during abrasive machining). Izd-vo Sarat, Un-ta, Saratov 20. Korchak SN (1974) The productivity of the grinding of steel parts. Mashinostroenie, Moscow 21. Pereverzev PP, Pimenov DY (2016) A grinding force model allowing for dulling of abrasive wheel cutting grains in plunge cylindrical grinding. J Frict Wear 37(1):60–65 22. Lazoglu I, Kratz H, Buyukhatipoglu K, Klocke F (2006) Forces and temperatures in hard turning. J Taylor Francis Group 10(20):157–179. https://doi.org/10.1080/1091034060 0713554

224

P. P. Pereverzev et al.

23. Pereverzev PP, Alsigar MK (2017) Modelirovanie processa s”ema metalla v avtomaticheskih ciklah kruglogo naruzhnogo shlifovaniya s prodol’noj podachej (Simulation of metal removal process in automatic round-grinding external grinding cycles with longitudinal feed “). Metalloobrabotka 6:38–42 24. Pereverzev PP, Akintseva AV (2016) Model of cutting force while managing two regime parameters in the process of internal grinding. Procedia Eng 150:1113–1117. https://doi.org/ 10.1016/j.proeng.2016.07.222 25. Pereverzev PP, Akintseva AV (2016) Mathematical model for the internal grinding process of a non-circular hole machining. Procedia Eng 150:118–1123. https://doi.org/10.1016/j.proeng. 2016.07.223 26. Pereverzev PP, Akintseva AV (2016) Model of formation of processing errors intragrinding. Russ Eng Res 36(12):1048–1053. https://doi.org/10.3103/S1068798X16120133

Designing of Optimal Grinding Cycles, Sustainable to Unstable Mechanical Processing on the Basis of Synthesis of Digital Double Technology, and Dynamic Programming Method P. P. Pereverzev, A. V. Akintseva and M. K. Alsigar

Abstract Currently, there are non-calculation methods of optimal grinding cycles that are resistant to unstable processing conditions for CNC machines in automated engineering; this makes technologists to lower the cutting conditions significantly to guarantee avoiding reject in grinding operations. As a result, CNC machines are used inefficiently; full automation of the preparation of control programs for CNC machines becomes impossible without using high-performance optimum grinding cycles ensuring stable processing accuracy; it is also impossible to design manufacturing cyber-physical systems in accordance with the concept of “Industry 4.0.” The article describes the synthesis of the digital twin technology and dynamic programming method for designing the optimal grinding cycle for resistance to variable technological factors, which makes it possible to: prevent the rejection of circular grinding; determine the causes of rejection; improve reliability and stability of the grinding cycle to the cumulative effect of variable factors; predict the fluctuation of accuracy and roughness parameters, hardness of the machined surface when processing a batch of parts. The practical result of the synthesis of the digital twin technology and dynamic programming method is an increase in the level of designing automation of control programs for CNC machines, ensuring the calculation of optimal values of radial feed at all cycle stages, the optimal distribution of the allowance removal over the cycle stages, which ensures the minimum main grinding cycle time and reduction in risks to meet the specified requirements on the quality of the machined surface of the part. Keywords Dynamic programming method


Digital twin
Cycle
Grinding

P. P. Pereverzev
A. V. Akintseva (&) South Ural State University, 76, Lenin Avenue, 454080 Chelyabinsk, Russia e-mail: [email protected] M. K. Alsigar College of Engineering, University of Dhi Qar, 31, Dhi Qar, Nasiriyah 3629, Republic of Iraq © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_25

225

226

P. P. Pereverzev et al.

Nomenclature P V DPM S mS and m V тVs тS n zS zV

Allowance, mm Axial feed speed, mm/min Dynamic programming method Radial feed, mm/dv.stroke The coordinate of the node from which the optimal move is made is memorized The number of axial feed speed The number of radial feed The number of the allowance disk The number of the radial feed switching stage The number of the stage of switching the axial feed speed

1 Introduction The most important stage in the development of the technological process of parts processing on circular grinding operations, performed on CNC machines, is designing control cycles for the radial and axial feeds in the development of cutting modes for control programs. The cycle of program feed works in automatic mode according to the commands of the active control device and switches feeds depending on the remainder allowance part. The operation performances, the completeness of using the technological capabilities of the machine, and the number of reject products depend on the quality of the cycle development. The processing of the entire batch of parts on the CNC grinding machine is carried out in accordance with the designed grinding cycles of radial and axial feeds under conditions of various variable technological factors related to blunting of the cutting tool, variable allowance, different initial radial runout of the workpiece, etc. As a result of influence of variable technological parameters, the processing of each part at the same cycle is carried out with different dynamics of cutting forces, and, consequently, with different values of the actual radial feed which causes fluctuations in the indicators of dimensional accuracy and quality of the machined surface in the batch of parts. There are three methods of cycles designing. The first is based on the normative-reference literature, developed by considering the statistical data obtained in 60–90s for universal machines [1, 2]. It makes this method non-suitable in the conditions of modern automated production, because it does not consider the production capabilities and power of modern CNC machines. The second method includes various engineering methods of cycles designing. There are many works devoted to questions of designing grinding cycles for CNC machines, the following researches can be singled out Cahill et al. [3], Lur’e [4], Dong et al. [5], Amitay et al. [6], Phan et al. [7], Alagumurthi et al. [8] and etc. [9–16]. Analysis of the

Designing of Optimal Grinding Cycles, Sustainable …

227

literature revealed that none of these works deals with the design of optimal grinding cycles that are resistant to variable processing conditions. Technologists’ lack of tools for optimal cycles designing and testing them for resistance to the effects of variable technological factors leads to the fact that in practice it is necessary to correct cutting modes for CNC machines manually, deliberately underrating the level of cutting modes to a safe level at which the pattern requirements for the accuracy and quality are guaranteed to be executed (the third method). In order to solve the problem of designing optimal in the speed of operation and resistant to variable conditions of circular grinding cycles conditions, a methodology of designing the optimal grinding cycles resistant to variable processing conditions has been developed for CNC machines. As a mathematical optimization method in this methodology, the dynamic programming method (DPM) is used; this method does not require the construction of the tolerance range of control parameters in advance (in our case—the tolerance range of feeds speed and the time of processing cycles) [17]. The optimization of the cycle with using DPM is carried out by analogy with the optimization of the transportation problem in which the road network (for our task—possible variants for switching the radial feed in the cycle) with the intermediate stations (state of the workpiece during processing) is specified, and it is necessary to find an optimal route (optimal cycle) between two stations A (workpiece) and B (finished part). Let us consider the optimization of radial and axial feed control DPM cycles by the example of internal grinding [18–20]. The procedure for finding the optimal trajectory of the radial and axial feed control cycles is carried out in the coordinates «Radial feed (S, mm/dv.stroke)— axial feed speed (V, mm/min)—Allowance (P, mm)» on the coordinate grid (Fig. 1), the scales of which are divided into discrete, because DPM refers to the methods of discrete optimization. The scale on a radial feeder is divided into MS (MS is the number of radial movements of the machine) to the range [Smin; Smax], the scale of speed of axial feed—in MV (MV is the number of the axial feed velocity on the characteristics sheet of the machine) in the range [Vmin; Vmax], the scale of allowance-on N discrete (N—the maximum number of the discrete allowance at which the value of the remaining part of the allowance is zero). At each intersection of the grid, information cell is placed (number 1 in Fig. 1), containing the necessary data for the calculation (including the optimal time to reach the state and the coordinates from which the optimal move is made). Since the DPM is the principle of «optimal trajectory optimal in each area», then the optimal time will be determined after the selection of the optimal speed from the set of competitive moves (possible variants of switching of flows within the designed cycle) occurring in the information cell with the coordinates [n, mS, zS, mV, zV] of the previous discrete allowance (n − 1) in all programming values of the radial and axial feeds. Therefore, only valid moves are selected from all competing moves (numbers 2 and 3 in Fig. 1). With regard to the optimization problem, a move is considered valid if all the limitations of the objective function are fulfilled in the achieved state after the given move. The restrictions in force at the beginning of the grinding cycle (e.g., the power limit) are checked on the considered discrete

228

P. P. Pereverzev et al.

Fig. 1 Coordinate grid for optimization of radial and axial feed DPM control cycles in internal grinding operations taking into account the restrictions on the permissible number of stages of feed cycles: 1—information cell; 2—competing moves; 3—moves, allowed by one or another restriction; 4—the optimal course, having a minimum processing time; 5—the reverse procedure; 6—the information cell, eliminated from the calculation

allowance for each information branch on the corresponding program feeds. Checking the restrictions, acting throughout the cycle (e.g., restrictions on accuracy and roughness), is carried out by modeling the removal of the remaining part of the allowance at the minimum possible technical characteristics of the machine radial and axial feeds. From the set of permissible moves, the optimal move is selected, which has the optimal time to reach the state and is permissible under all restrictions (4 in Fig. 1). Then the number of radial and axial feeds, and number of steps, produced by the optimal move, is written into the cell (variables mS and m V in Fig. 1). After considering the last discrete allowance of a number of optimal moves, located on different numbers of radial and axial feeds, one move is selected, which has a minimum time to reach the final state. For fixing the optimal control of the cycle of

Designing of Optimal Grinding Cycles, Sustainable …

229

grinding, the reverse procedure is performed, which begins from the process of having the minimum time achieving the end state. In Fig. 1, the reverse procedure is numbered in Roman numerals from I to IV.

2 Synthesis of Digital Twin and MDP Technologies for Designing an Optimal Grinding Cycle for Resistance to Variable Technological Factors The optimal grinding cycle, designed on the basis of averaged deterministic grinding conditions, has low reliability and low resistance to the influence of variable technological factors on the processing accuracy and the maintenance of other quality parameters. Consequently, the proposed methodology of optimal cycles designing [18–20] does not consider unforeseen situations related to unstable grinding conditions, blunting of the grinding wheel grains, fluctuation of allowance or initial radial runout of the workpiece, etc., which may lead to a rejection of the grinding operations. The following variable technological factors were considered when designing the optimal grinding cycle for stability: (1) Blunting of the wheel grains during processing: At the beginning of the grinding process, note that the edges of the grains are sharp and show high muticopsis after dressing; during processing, the cutting grains become blunt. At the end of durability—immediately before wheel dressing—the grains have a maximum bluntness. (2) Fluctuation allowance for processing: Due to the dispersion of the dimensions of the machined surface in a batch of parts, the allowance on each part has a different value. The allowance in the batch of parts changes from minimum to maximum. (3) Fluctuation: Fluctuation of the initial radial runout of the workpiece depends on the processing accuracy at the previous operation. Theoretically, it can vary from zero to maximum. (4) Dimensional wear of the grinding wheel and reduction in cutting speed: When processing the batch of parts, the wheel is dressed periodically; after that, the wheel diameter decreases—as a result the cutting speed is reduced. Dimensional wear varies from minimum to maximum. (5) Number of zones with different cutting conditions: During circular grinding with axial feed (internal and external grinding), there are two reverse and one non-reverse zones, in which the actual cutting modes are different in the same grinding cycle. In the reverse zone, the treatment is carried out using different types of grinding (plunging with radial feed, sparking-out with radial feed during reverse, grinding with axial feed). In addition, there are zones with discontinuous surfaces (key slot, splined surface), in which another dynamic of cutting forces and possible the presence of rejection.

230

P. P. Pereverzev et al.

In the development of a diagnostic system for the stability of the grinding cycle, the concept of a digital twin (“digital twin,” DT) was applied to the cumulative effect of constantly changing variables that arise when the processing a batch of parts starts. As a part of this concept, a DT system model has been developed and designed for: • preventing rejection and its causes during circular grinding operations; • increasing reliability and stability of the grinding cycle to the cumulative effect of variable factors; • ensuring full automation of the design of the control program at the stage of calculation and programming of cutting modes; • predicting the fluctuations of accuracy, roughness, hardness of the machined surface after processing a batch of parts; • fixing an array of variable factors under which the conditions for obtaining each type of rejection occur; • forming an array of constraints of the objective function on varying grinding conditions for cycle optimization system. The developed system “DT-CicleStab” provides a synthesis of the diagnostic system and the optimization system according to following algorithm (Fig. 2):

Fig. 2 Algorithm of the “DT-CicleStab” system, which provides a synthesis of diagnostic and optimization systems

Designing of Optimal Grinding Cycles, Sustainable …

231

(1) formation of the initial data package for the grinding operation; (2) in the design system of the optimal cycle, the first version of the optimal grinding cycle is designed on the basis of the averaged constant grinding conditions with checking the constraints under these conditions; (3) optimal grinding cycle is transferred to the digital twin to test the cycle for resistance to unstable processing conditions. A model of the machined surface of the finished part is formed at the end of the grinding cycle. A passport of the accuracy of the grinding cycle processing (Fig. 2), in which for each given accuracy parameter a drawing field of tolerance and a dispersion field of error from the influence of variable factors in each zone (reverse and non-reverse zones) are indicated; (4) in the case of a rejection, arrays of conditions for the occurrence of rejects for a given grinding cycle and combinations of variable factors are generated. Correction of the criterion function constraints for the DPM grinding cycle optimization system is carried out.

3 Conclusion 1. The developed methodology for designing the optimal grinding cycles [18–20], resistant to variable processing conditions allows to automate the process of developing control programs for CNC machines and can be used as a basis for the development and implementation of production cyber-physical systems as a part of concept “Industry 4.0”. 2. The synthesis of digital twin and DPM technologies for designing an optimal grinding cycle for resistance to variable technological factors allows: • prevent rejection during circular grinding operation; • to trace the causes of rejection; • improve the reliability and stability of the grinding cycle to the cumulative effect of variable factors; • predict fluctuations in accuracy, roughness, hardness of the processed surface after processing a batch of parts. 3. The practical result of the synthesis of the digital twin and the dynamic programming method is increasing the level of automation of designed control programs for CNC machines that calculate the optimal values of the radial feed at all stages of the cycle, optimal distribution of the removed allowance over the cycle stages that ensure the minimum basic grinding cycle time and reduce the risks of meeting the specified requirements on the workpiece quality.

232

P. P. Pereverzev et al.

References 1. Engineering industry standards of time and the cutting modes for rationing of the works carried out on the universal and multi-purpose machines with numerical program control. Part II. Standards of the cutting modes (1990) Ekonomika. Moscow 2. The cutting modes for the works carried out on grinding and hand-operated lapping machines and semiautomatic machines: reference book (2007). ATKOSO Publ., Chelyabinsk 3. Cahill MJ, Bechtold MJ, Fess E, Wolfs FL, Bechtold R (2015) Ultrasonic precision optical grinding technology. In: Proceedings of SPIE—The International Society for Optical Engineering 9633:96330 4. Lur’e GB (1979) Optimizing the grinding cycle by adaptive control. Mashinostroitel’, Moscow 5. Dong S, Danai K, Malkin S, Deshunukh A (2004) Continuous optimal in feed control for cylindrical plunge grinding. Part 1. Methodology. J Manuf Sci Eng 126(2):327–333. https:// doi.org/10.1115/1.1751423 6. Amitay G, Malkin S, Koren Y (1981) Adaptive control optimization of grinding. J Eng Ind 103(1):103–108. https://doi.org/10.1115/1.3184449 7. Phan AM, Summers MP, Parmigiani JP (2011) Optimization device for grinding media performance parameters. Int Mech Eng Congr Expos (IMECE) 3:915–923 8. Alagumurthi N, Panairadja K, Soundararajan V (2006) Optimization of grinding process through Design of Experiment (DOE)—a comparative study. Mater Manuf Processes 21 (1):19–21 9. Lee TS, Ting TP, Lin YJ, Htay T (2007) A particle swarm approach for grinding process optimization analysis. Int J Adv Manuf Technol 33(11):1128–1135. https://doi.org/10.1007/ s00170-006-0538-y 10. Zhang J, Liang SY, Yao J, Chen JM, Huang JL (2006) Evolutionary optimization of machine processes. J Intell Manuf 17(2):203–215. https://doi.org/10.1007/s10845-005-6637-z 11. Bertsekas D (2005) Dynamic programming and optimal control. Athena Scientific, Belmont 12. Lee CW (2009) Dynamic optimization of the grinding process in batch production. J Manuf Sci Eng, Trans ASME 131:61–66 13. Nishimura T, Inasaki I, Yamamoto N (1989) Study on optimization of internal grinding cycle. Trans Jpn Soc Mech Eng 55:1808–1813 14. Inasaki I (1991) Monitoring and optimization of internal grinding process. CIRP Ann Manuf Technol 400:359–363 15. Alvarez J, Barrenetxea D, Marquinez JI, Begiaga I, Gallego I (2014) Continuous variable feed rate: a novel method for improving infeed grinding processes. Int J Adv Manuf Technol 73:53–61 16. Akintseva A, Pereverzev PP (2017) Prospects for the development of the theory of designing optimal cycles of machining in a multidimensional space of control parameters. MATEC Web Conf 129:01018 17. Bellman R (1960) Dynamic programming. Foreign Literature Publishing House, Moscow 18. Pereverzev PP, Akintseva AV (2015) Automatic cycles’ multiparametric optimization of internal grinding. Procedia Eng 129:121–126. https://doi.org/10.1016/j.proeng.2015.12.019 19. Pereverzev PP, Akintseva AV (2016) Optimal internal grinding cycles in multidimensional control-parameter space. Russ Eng Res 36(11):974–978. https://doi.org/10.3103/ S1068798X16110174 20. Akintseva AV, Pereverzev PP (2017) Complex optimization of parameters for controlling the cycle of internal grinding by the method of dynamic programming. MATEC Web Conf 129:01019

Controlled Assembly of Rotors S. M. Beloborodov, V. F. Makarov and M. L. Tselmer

Abstract This article analyzes the dynamic stability of rotors with magnetic suspension for compressors of gas-pumping units. The main advantages and disadvantages of magnetic pendants of various manufacturers are formulated. A set of measures was developed and described to ensure the dynamic resistance of rotors with magnetic suspension, including balancing the multiplanar shaft and the calculation and measurement method for assembling rotors. Balancing the shaft with the correction of imbalances in their places allows minimizing their residual vector. The method of calculation and measurement provides a reduction in the scope of work on assembling the rotor (increasing economic efficiency, reducing costs) while maintaining the specified manufacturing accuracy. The new method of assembling the rotor includes the preliminary preparation of the shaft and mounted elements (impellers). As a result, the calculation and measurement method for the assembly of rotors provide a reduction in the amount of work on the assembly of the rotor with ensuring its specified accuracy. This, in turn, allows reducing the cost of setting up and adjusting equipment at the compressor station and to ensure the dynamic stability of the rotors and shafts at a high level.













Keywords Centrifugal compressor Rotor Shaft Impeller Magnetic suspension Imbalance Controlled assembly Balancing Dynamic stability













S. M. Beloborodov (&)
V. F. Makarov Perm National Research Polytechnic University, 29, Komsomolsky Av., Perm 614990, Russia e-mail: [email protected] M. L. Tselmer PSC Scientific and Production Association “Iskra”, 28, Academik Vedeneyev St., Perm 614038, Russia © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_26

233

234

S. M. Beloborodov et al.

1 Introduction The long-term and often unsatisfactory experience of commissioning compressor stations when installing centrifugal blowers with magnetic bearings (MS) of shafts differs in the duration of work on preparing shafts with balancing rotors in the supports, and this problem requires additional solutions. The use of rotor magnetic bearings (AMB) in centrifugal blowers is a modern solution: the absence of friction of the rotor parts, low air resistance, and high efficiency make it possible to produce modern units and compressor stations that do not require additional settings [1]. The use of a non-lubricating supercharger system creates additional economic efficiency. The electrical part of the bearings consists of magnets with an electronic control system, and the mechanical part is a rotor assembled on the basis of the shaft. It is during the assembly of the magnetic sub-ports and the shaft that the problems of dynamic stability appear.

2 Relevance and Problem Statement The only unsolved problem is the long-term configuration of the electronic control system. However, it can be solved by the technological method: simultaneous machining of the balancing surface of the shaft and the control belt of the rotor MS. This is also facilitated by the design feature of the MS rotor, which provides for its thermoplaning and fixed position for the entire duration of operation. All this will minimize the magnitude of the mounting imbalance (due to the convergence of the axes of the masses) and reduce the time and number of tuning operations [2]. In achieving the dynamic stability of the rotors, there is another significant feature: the misalignment of the mass axis and the axis of rotation of the rotor. This problem is related to the design and manufacturing technology of magnetic bearings: mechanical treatment of the surfaces of the bearing and the rotor sensor is carried out after installation on the rotor [3]. The deformation during installation of magnetic bearings is not regulated, and the installation does not use the method of hydraulic wedging, which reduces deformation. In this case, the cone fit 1–50 creates significant errors in the manufacture. Another design defect is the length of the conical part of the rotor shaft, which reduces the strength characteristics of the structure and increases the cost of manufacture. The assembly technology of rotors includes the installation and removal of rotors of magnetic bearings, which leads to a decrease in assembly accuracy and an uncontrolled increase in rotor imbalances. In the case of using control belts outside the rotors of magnetic bearings, negative assembly results can be overcome by measuring and mathematical calculations [4].

Controlled Assembly of Rotors

235

In the case of installation of control belts inside the technological process of assembling, the rotor will be uncontrollable, and the correction of deficiencies is possible only by balancing the compressor at the place of operation [5]. Existing design and technology deficiencies subsequently lead to lengthy improvements in equipment, and in some cases, its replacement. At the same time, there is a positive example of the use of new designs and technologies, which allows reducing the compressor visibility by more than two times [6–8]. Slowdown of technological progress in the design of magnetic suspensions of domestic manufacturers due to the reluctance of developers to use new design and technological solutions. Over the past decade, a study has been conducted that allows the formulation of measures to ensure the stability of (dynamic) rotors. A centrifugal compressor is stable at a vibration level of up to 75 microns, and an imbalance level of up to 500 N [9–11]. The imbalance of the shafting of a centrifugal compressor with all the imbalances taken into account is equal to: Iper ¼ ðIres þ Imount þ Iext Þ  1800 g
mm

ð1Þ

Rotor imbalance is described by the dependency: Iper ¼ ðIres þ Imount Þ  1000 g
mm

ð2Þ

The proposed set of measures includes the following assumptions: • the material of the shaft and its elements adopted homogeneous with constant density; • microgeometry was not taken into account of the mounting surfaces of the shaft and elements; • the microgeometry of cylindrical surfaces was not taken into account.

3 Controlled Assembly of Rotors The complex includes a multi-faceted shaft and rotor balancing. The complex includes a multi-faceted shaft and rotor balancing. Correction of the unbalance of the shaft along the planes reduces the vector component of the imbalance. At the same time, the number of surfaces for correcting imbalances for such a design must be at least four, and the planes for correcting the imbalance may not coincide with the planes of imbalances. A method of controlled rotor assembly was developed—calculation and measurement, which reduces the amount of rotor assembly (increased economic

236

S. M. Beloborodov et al.

efficiency, reduced costs) while maintaining the specified accuracy of rotor manufacturing [12–14]. The method of calculation and measurement of the rotor assembly includes preliminary preparation of the shaft and the installed elements (impellers). For which the shaft is mounted on the supports of the measuring stand, on which the place of maximum radial runout of the shaft control belt is determined, the places of maximum radial runout of the shaft seating surfaces are identified and marked (Fig. 1) [15, 16]. At the next step, the shaft is balanced, the residual imbalances are diametrically opposite to the point of maximum radial runout of the shaft control belt (Fig. 2). Impellers are also pre-assembled. They are mounted on the balancing of the mandrel with a diametrically opposite direction of maximum radial runout of the sealing surface space covering the impeller disk relative to the point of maximum radial runout leveling the surface of the mandrel seat (Fig. 3). At the impeller hub mark the place of maximum radial runout of the seating surface of the balancing mandrel [17, 18].

Fig. 1 Preparation of the shaft

Fig. 2 Shaft balancing

Controlled Assembly of Rotors

237

Fig. 3 Impeller balancing and preparation

When assembling the shaft and impellers in the rotor, due to the difference in beats (hence, eccentricities) of the mounting surfaces of the shaft and the balancing mandrel, a mounting imbalance is formed with direction toward the maximum radial runout of the mounting surface of the shaft (Fig. 4). The manufacturing error of the seating surfaces of the balancing mandrels is lower than the manufacturing error of the shaft seating surfaces. Therefore, mounting imbalance can be defined as:

Fig. 4 Controlled assembly of rotors

238

S. M. Beloborodov et al.

Imount

  DD0  DDm ¼M 2

ð3Þ

where—Imount—the value of the mounting imbalance, M—the mass of the impeller, DD0—beating of the shaft mounting surface, DDm—beating of the mounting surface of the mandrel.

4 Results of Experimental Studies Compensation of the mounting imbalance may well be provided by the diametrically opposite direction of the compensation residual imbalance equal to the size of the mounting imbalance (it will be mutually balanced by the mounting imbalance after the wheel is mounted on the shaft): Icomp ¼ M

  DD0  DDm 2

ð4Þ

The direction of the compensatory mounting imbalance should be diametrically opposite to the direction of the mounting imbalance [19–21]. At the same time, the wheel should have a residual imbalance that is diametrically opposite to the maximum radial runout of the seating surface. With a known mass and specific unbalance (the latter is determined depending on the type and frequency of rotation of the rotor), its value can be defined as: Ires ¼ M
es

ð5Þ

At the same time, in the interests of stable shaft straightening when the rotor is working, the residual unbalance amount should be taken in the range from half to the full value of the nominal value: Ires ¼ ð0; 5  1ÞM
es

ð6Þ

Depending on the location of the wheel (along the shaft) should be considered massogeometric coefficient k Ires ¼ ð0; 5  1ÞkM
es

ð7Þ

Based on the foregoing, the magnitude of a given residual imbalance of the impeller is determined from the dependencies: Igiv

  DD0  DDm þ ð0; 5  1Þkes ¼M 2

ð8Þ

Controlled Assembly of Rotors

239

where—Igiv—the specified value of the residual imbalance of the impeller, M—the mass of the impeller DD0—the beating of the mounting surface of the shaft, DDm— the beating of the mounting surface of the mandrel, k—mass-geometric coefficient, es—specific imbalance. Impellers should be balanced with the direction of a given residual imbalance, 180° relative to the maximum beating value of the mounting surface of the balancing mandrel. Assembly of the rotor is carried out with the combination of marked places and pre-prepared (balanced) shaft and wheels (Fig. 4). With significant deviations of the results of manufacturing components (shaft, wheels) from the requirements of design documentation, balancing can be calibrated with the direction of residual unbalance, 180° relative to the point of maximum beat value of the control belt.

5 Conclusion Thus, a technique for measuring and calculating the rotor assembly enables reduction of works on the rotor assembly to ensure its predetermined accuracy. This, in turn, can reduce setup costs and commissioning of equipment at compressor stations and ensure the dynamic stability of rotors and shafts at a high level.

References 1. Rukovitsyn IG (2010) Features of electromagnetic bearings for gas-pumping aggregates with elastic rotors. Dissertation, VNIIEM 2. Vartanov MV, Bezruchkin VV, Dmitriyeva OG (2011) Automation of procedures for providing and evaluating the operational technology of large-sized products. Assembly Mech Eng, Instrum Mak 10:3–6 3. Firma PP “GKHT”, nauchno-proizvodstvennoye predpriyatiye “VNIIEM” (1997) Magnetic support of the unit. Rus Patent 97104412/06, 19.03.1997 4. Kovalev AYu (2013) Technological support of the balance of high-speed rotors with magnetic bearings based on the compensation method of assembly. Dissertation, Perm National Research Polytechnic University 5. Beloborodov SM (2011) Methodology for ensuring the dynamic stability of shafts of high-speed gas turbine units based on the adaptive assembly of rotors. Dissertation, Perm National Research Polytechnic University 6. Beloborodov SM, Petrov VYu, Modorskii VYa, Tselmer ML (2018) Providing gas-dynamic tests for 2FSI subsystems. Fomin V.M. International Conference on the Methods of Aerophysical Research (ICMAR 2018), 13–19 August 2018, Novosibirsk, Russia, Abstracts. Pt. II, pp 25–26 7. Vartanov MV, Zinina IN (2017) Mathematical model of robotic assembly by means adaptation and low-frequency vibration. Assembly Autom 1(37):130–134

240

S. M. Beloborodov et al.

8. Ur’yev YeV, Kistoychev AV, Degtyareva YeYu, L’vov MM, Biyalt MA, Shvatskiy AV, Deminov AM (2016) On the standardization of the quality of balancing of flexible rotors. Heavy Eng 11–12:9–18 9. Yim KB, Yim JT (2013) Dynamic stability of a rotor with shear-flexible shaft under axial loads. J Mech Sci Technol 27(2):359–366. https://doi.org/10.1007/s12206-013-0102-2 10. Gleyzer AI, Korneyev NV (2004) Imbalance and balancing of rotor systems. Tol’yatti, p 240 11. Marshall DF, Sorokes JM (2000) A review of aerodynamically induced forces acting on centrifugal compressors, and results vibration characteristics of rotors. In: Proceedings of the29th Turbomachinery Symposium, Texas, pp 263–280 12. Nepomiluyev VV (2000) Development of technological fundamentals for quality assurance of assembly of high-precision components of gas turbine engines. Dissertation, RSATU 13. Mekhonoshina EV, Modorskii VYa (2016) Impact of magnetic suspension stiffness on aeroelastic compressor rotor vibrations of gas pumping units. AIP Conf Proc 1770:030113-1– 030113-5 14. Tsimberov DM, Magamedov AM (015) Method of shafting assembly control subject to various factors in the process of industrial units operation. Vestnik IzhGTU im. M.T. Kalashnikova 2(66) 15. Bazrov BM (ed) (2011) Technology of assembling machines. Spektr, Moscow 16. Ur’yev YeV (1996) Basics of reliability and technical diagnostics of turbomachines. Yekaterinburg 17. Beloborodov SM, Tselmer ML, Sviridov EV (2018) Precision balancing of impellers. Chem Pet Eng 53(11):797–800. https://doi.org/10.1007/s10556-018-0424-0 18. Beloborodov SM (2011) A method of preliminary balancing of an element of a built-up rotor on a mandrel. RU Patent 2431064 19. Firma PP “GKHT”, nauchno-proizvodstvennoye predpriyatiye «VNIIEM» (1997) Magnetic support of the unit. RU Patent 2129228 20. Modorskiy VYa (2014) The solution of the engineering problem on the basis of the high-performance computer system of the Perm National Research Polytechnical University. PNRPU, p 154 21. Beloborodov SM (2015) A method of balancing a built-up rotor of a centrifugal compressor. RU Patent 2554669

Study of Energy Intensity of Shaped Holes Broaching Process V. V. Kuts, M. S. Razumov and V. S. Kochergin

Abstract Broaching is a type of machining with a multipoint tool with translational primary cutting motion distributed over the entire surface being processed without feed motion. In addition to high performance, one of the most important advantages of a broaching tool is high durability due to the design in which shaping elements duplicate each other, partially or completely. Another significant advantage of a broaching tool is the ability to be multiple reground, which allows bringing the cutting parameters of a conditionally worn-out tool back to the characteristics of a new one, without exceeding the tolerances. Due to these qualities, it is not unusual for an acquired broach to be used at enterprises for years, especially in the case of small-scale production. Designing broaches for machining shaped holes is accompanied by the issue of estimating the energy intensity of the broaching process as one of the criteria for selecting cutting conditions and estimating the energy efficiency of the process itself. The authors of this paper consider the broaching process in terms of energy, taking the value of the broaching process energy intensity as a criterion. Considering that the total cost of cutting when broaching is up to 10 times higher than the cost of a useful work of forming a new surface, it becomes relevant to use a criterion that allows varying the cutting process parameters to achieve metal removal with minimal energy input. Keywords Broaching process


Energy intensity
Efficiency
Usefulness
Cutting

1 Introduction Broaching is a type of machining using a multipoint tool with translational primary cutting motion distributed over the entire surface being processed without feed motion. In addition to high performance, one of the most important advantages of a V. V. Kuts
M. S. Razumov
V. S. Kochergin (&) The Southwest State University, 94, 50 Let Oktyabrya St., 305040 Kursk, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_27

241

242

V. V. Kuts et al.

broaching tool is high durability due to the design in which shaping elements duplicate each other, partially or completely [1]. Designing the broaches for machining shaped holes is accompanied by the issue of estimating the energy intensity of the broaching process as one of the criteria for selecting cutting conditions and estimating the energy efficiency of the process itself [2–6]. In studying the broaching process in terms of energy, we will characterize it using the following values [7, 8]: • energy intensity of machining (Euseful ); • energy intensity of consumed power (Econs ).

2 Theoretical Dependencies of the Values Under Study on the Broaching Speed In the general case, the energy intensity is a ratio of the energy of the process under consideration (W, kWh) to the volume of removed chips (V, m3) [9]. E¼

W ; kWh/ m3 V

ð1Þ

The amount of energy can be determined from the ratio W¼

P
t0 60

ð2Þ

where P—power, kW; t0—processing time, min. Power can be both consumed (Pcons ) and useful (Puseful ). When broaching, the consumed power is a reference characteristic of the broaching machine. Thus, the energy intensity consumed during broaching is a function of three values—the rated power of the broaching machine (a constant value independent of cutting conditions), the volume of removed chips (a value determined by the product size), and the processing time (cutting conditions function) [10]. The volume of removed chips for the existing broach size (GOST 24821-81) can be determined as follows: V ¼ lbr
A ¼ ðAchamf þ Around þ Aspl Þ
lbr

ð3Þ

where Achamf ¼ Dch:last  Dch:1 , Ar ¼ Dr:fin  Dr:1 , Aspl ¼ Dspl:fin  Dspl:1 , Dch:last — diameter of the last chamfer tooth, Dch:1 —diameter of the first chamfer tooth, Dr:fin: —diameter of the round finishing tooth, Dr:1 —diameter of the first round tooth, Dspl:fin: —diameter of the splined finishing tooth, Dspl:1: —diameter of the first spline tooth [11–13].

Study of Energy Intensity of Shaped Holes Broaching Process

243

The processing time during broaching is defined as the ratio of the length of the machine stem working stroke to the cutting speed. s¼

Lw:s: ; 1000
t

ð4Þ

where Lw:s: —length of working stroke during broaching, mm; Lw:s: ¼ L  l1 þ l0 þ lp , where L—broach length, mm; l1 —length from the end face to the first tooth of the broach, mm; l0 broaching length, mm; lp = 50 mm. Parameters that make up the formula for the broach’s working stroke length are constants presented in the standard. An exception is a broaching length specified by the permissible range [14, 15]. The range of broaching speeds is t ¼ f2; 3; 4; 5; 6; 7; 8; 9; 12; 13; 15g m/min.

3 Generating Dependencies to Calculate Values of Consumed Energy Intensity The values of consumed energy intensity were calculated based on relations (1)–(4) for limiting values of power applied during broaching (6 and 55 kW), broaching lengths in accordance with the standard, each value of the specified range of cutting speeds. These calculations allow estimating the dependency of the maximum value of consumed energy intensity on the depth of cut per tooth [16–19]. Previously, the analysis of relations (1)–(4) allowed establishing the theoretical dependency of the value under study on the broaching speed. Thus, we get E ¼ K1
t1 ; kWh/m3

ð5Þ

Let’s assess the physical meaning of the coefficient. This coefficient is calculated
Lw:s: 6 using the ratio K1 ¼ Pcons 60
V
10 , where Pcons—kW, Lw.s.—mm. kW
h ; therefore, K is a coefficient As a result, we get the units of measurement min 1
m2 showing the number of kW of machine power consumed to remove 1 m2 of the oversize area [20]. To obtain a mathematical model of dependency K1 ¼ f ðSzÞ, we’ve compiled Table 1. By presenting a mathematical model of linear function y ¼ a þ k
x, we get models presented in Table 2. The analysis using Fisher’s criterion has proved the models’ adequacy.

244

V. V. Kuts et al.

Table 1 Analysis of the dependency of coefficient K1 on the depth of cut per tooth for different conditions of broaching Sz, mm

0.13 0.14 0.15 0.155 0.16 0.175 0.18 0.20

kW
h K1 ¼ 1000f ðSzÞ, min
m2 l0 ! min l0 ! max N ! min N ! min

l0 ! min N ! max

l0 ! max N ! max

13.398 12.400 11.538 10.346 9.425 6.523 6.504 6.284

122.815 113.660 105.762 94.835 86.389 59.794 72.011 57.607

68.972 63.948 59.096 53.173 48.403 34.540 41.795 34.210

7.524 6.976 6.447 5.800 5.280 3.768 4.559 3.732

Table 2 Result of plotting an approximation of function K1 ¼ 1000f ðSzÞ for different broaching conditions and the model adequacy condition Broaching conditions

l0 l0 l0 l0

! min; N ! min ! max; N ! min ! min; N ! max ! max; N ! max

Coefficients of the mathematical model having the form f ðSzÞ ¼ a  k
Sz a k

F

Ftable

28.758 15.124 254.864 138.632

1.099 1.046 1.403 1.572

5

119 59.613 1027.94 546.45

4 Estimation of the Broaching Process Energy Intensity Designing of special-purpose broaches is accompanied by the issue of estimating the energy intensity of the broaching process as one of the criteria for selecting cutting conditions and estimating the energy efficiency of the process itself. In the design calculation of broach parameters, the energy intensity will also be calculated using relations (1)–(5). For ease of use, model (5) is written in the form E ¼ Puseful
K2
t1 ; kWh/m3

ð6Þ

where values of coefficient K2 can be found in Table 3, and N is the useful power of broaching.

Study of Energy Intensity of Shaped Holes Broaching Process

245

Table 3 Result of plotting the function K2 ¼ 1000f ðSzÞ for various broaching conditions and the models adequacy condition Broaching conditions

Coefficients of mathematical model having the form f ðSzÞ ¼ a  k
Sz a k

l0 ! min l0 ! max

4.793 2.521

19.851 9.935

Table 4 Values of linear function coefficients for various c c

5°

10°

15°

20°

25°

ax bx

2229.702 60.975

2042.175 31.071

1930.539 17.93

1854.987 9.58

1788.422 5.03

Useful power of broaching is determined by the formula Puseful ¼ 1:8  105
Fcut
t

ð7Þ

where the cutting force Fcut is determined by the formula Fcut ¼ ðax
Sz þ bx Þ


Y


Zp


cu

K ; Zc

ð8Þ

Q Mathematical model coefficients az ; bx are determined using Table 4, cu — cutting perimeter, mm zp —number of simultaneously working teeth, K—total correction coefficient for axial cutting force, zc —number of teeth in a group Finally, we get Euseful ¼ 1:8
105 K2
ðax
Sz þ bx Þ


Y


Zp


cu

K ; kWh/m3 Zc

ð9Þ

The efficiency coefficient for broaching is determined by the formula g¼

Euseful 1; 8
10 ¼ Econs

5


ðax
Sz þ bx Þ
Pcons

Q cu


Zp
ZKc
t

ð10Þ

246

V. V. Kuts et al.

5 Conclusions (1) in terms of energy intensity, broaching is an effective type of processing (Fig. 1); (2) broaching is characterized by a decrease in consumed energy intensity and an increase in cutting speed and depth of cut per tooth. With an increase in speed, the consumed energy intensity is decreased along a hyperbola, which with an increase in depth of cut per tooth leads to a linear decrease of the given value (Figs. 2 and 3); (3) in terms of energy consumption, broaching is most effective in mass production, since an increase in broaching length leads to a linear decrease in consumed energy intensity, which indicates the efficiency of batch processing of parts inherent in mass production;

Fig. 1 Dependency of the efficiency coefficient of broaching on the cutting speed for different values of rake angle; when machining steel parts, depth of cut per tooth of the rough splined part Sz = 0.13 mm, the broaching machine power is 25.5 kW, the total correction coefficient for axial cutting force is 1, the pitch of roughing teeth t = 11 mm

Fig. 2 Dependency of the maximum consumed energy intensity of broaching on the depth of cut per tooth for different cutting speeds according to GOST 24821-81; when machining steel parts, the power of broaching machine is 25.5 kW, the total correction coefficient for axial cutting force is 1

Study of Energy Intensity of Shaped Holes Broaching Process

247

Fig. 3 Dependency of the maximum consumed energy intensity of broaching on cutting speeds for different depths of cut per tooth according to GOST 24821-81, when machining steel parts, the power of broaching machine is 25.5 kW, the total correction coefficient for axial cutting force is 1

(4) the useful energy intensity of broaching does not depend on the cutting speed, but is a function of depth of cut per tooth, wherein its increase leads to a linear increase in useful energy intensity; (5) the efficiency coefficient of broaching is directly proportional to the cutting speed. In regards to the depth of cut per tooth, there are two components: The first component does not depend on the depth of cut per tooth, and the second is directly proportional to it (ratio 10); (6) wherever applicable, it is more profitable to control the efficiency of the broaching process by changing the cutting speed.

References 1. GOST 28442-90 (1990) Broaches for cylindrical, splined and polygonal holes. Specifications. Standards Publishing House 2. Emelyanov SG, Seleznev YN, Gubanov VS, Kochergin VS, Khomutov RN (2017) Study of the working area temperature during broaching of shaped holes. Assembling Mech Eng, Instrum Mak 12:550–553 3. Emelyanov SG, Seleznev YN, Gubanov VS, Kochergin (2018) Study of the effect of broaching parameters on the working area temperature when machining shaped holes. Assembling Mech Eng, Instrum Mak 4:175–181 4. Seleznev YN, Gubanov VS (2008) Study of the dependency of the broaching process energy intensity. Materials and reinforcing technologies. In: Proceedings of the XV Russian Scientific and Technical Conference with international participation, Kursk State Technical University, Kursk, pp 260–263 5. Seleznev YN, Gubanov VS, Lymanyuk AY (2011) Study of the dependency of the broaching process energy intensity on the rake angle and depth of cut per tooth. Modern tooling systems, Kursk, KSTU, pp 13–20 6. Seleznev YN, Yatsun EI, Gubanov VS (2008). Setting the problem of studying the thermal phenomena of the broaching process. Modern tooling systems, information technologies and innovations. KSTU, Kursk, pp 51–56

248

V. V. Kuts et al.

7. Margulis DK, Tverskoy MM, Ashikhmin VN et al (1986) Broaches for machining holes. Mechanical Engineering, Moscow 8. Shchegolev AV (1960) Designing of broaches. Mashgiz, Moscow 9. Emelyanov SG, Seleznev YN, Gubanov VS, Kochergin VS, Evseev EY (2017) Study of the influence of design and geometrical parameters of broaching tools on the output characteristics of the broaching process. University Book, Kursk 10. Kosilova AG (Sc. Ed.) (1956) Reference guide of mechanical engineer. In: 2 vol, vol 1. Mashgiz, Moscow 11. Evseev EY, Kochergin VS (2016). Analysis of methods for determining cutting forces for internal broaching. Product quality: monitoring, management, improvement, planning. In: Proceedings of the third International Youth Scientific and Practical Conference, 17–18 November 2016 vol 1, pp 252–254 12. Evseev EY, Kochergin VS (2016). Analyzing the assessment of the broaching tool durability. Prospective development of science, equipment and technology. In: Proceedings of the sixth International Scientific and Practical Conference, 20–21 October 2016, pp 45–48 13. Kochergin VS, Evseev EY (2017) Analysis of restrictions on the main design parameters when refining the broaching tool. Mod Mater Equip Technol 2(10):35–39 14. Seleznev YN, Pavlov EV (2004) Analyzing the accuracy of approximating the profile of involute splined broach teeth along a circular arc, two circular arcs and the length of an involute. Modern tooling systems, information technologies and innovations. KSTU, Kursk, pp 129–134 15. Shchegolkov NN (2001) Automated approximation of the profile of cutting tools. STIN 11:17–22 16. Seleznev YN, Gubanov VS, Lymanyuk AY (2010) Developing and studying mathematical models of the dependency of axial force on the depth of cut per tooth and values of rake angles when broaching. Materials and sealing technologies. KSTU, Kursk, pp 184–191 17. Seleznev YN, Gubanov VS, Lymanyuk AY (2010) Developing and studying mathematical models of the dependency of specific force on the depth of cut per tooth and values of rake angles when broaching. Modern tooling systems, information technologies and innovations. KSTU, Kursk, pp 177–184 18. Seleznev YN, Gubanov VS, Sergeev SA (2005) Automated calculation of the minimum depth of cut per tooth for broaching. Modern tooling systems, information technologies and innovations. KSTU, Kursk, pp 118–123 19. Seleznev YN, Rukhlin AS (2003) Determination of the dependency of specific cutting forces on the feed per tooth when broaching holes in steel products. Physical and computer technologies in the national economy. The eighth International Scientific and Technical Conference, Kharkov, pp 104–106 20. Kosilova AG (Sc. Ed.) (1956) Reference guide of mechanical engineer. In: 2 vol, vol 2. Mashgiz, Moscow

Methods of Quality Control Manufacturing Gears of the Differential Satellites D. T. Safarov and A. G. Kondrashov

Abstract The features of bevel gears in differentials of heavy vehicles are considered. Attention is drawn to the inability to adjust the position of the satellites and gears, which imposes a rigid condition on the provision of the geometrical parameters of the gear rims. The content of the universal method for controlling the quality of the manufacture of machine parts at the same time by a set of interrelated accuracy indicators is disclosed. The peculiarity of the method is the consistent application of various methods of product quality management—probabilisticstatistical analysis, control, correlation analysis for each key indicator of accuracy. The advantage of the technique is the development of the content of corrective actions, taking into account the action of the main set of technological factors acting in the part processing operation. Information is given on the selection of key parameters of the satellite differential of geometric accuracy indicators, as well as indicators of the quality of the contact patch. The fact of insufficient information content of the standard method of rationing indicators of the quality of the contact patch has been established. The main provisions of the improved methodology are considered, with the addition of indicators of the relative location of contact patches. The content of corrective actions that leads to the improvement of the complex of indicators of the quality of the satellite of the differential is shown. The method allows formulating corrective actions as operative actions that can be performed without lengthy preparation, as well as engineering-technological measures, for which it is necessary to carry out local production preparation procedures. The data on the achieved improvement in the values of quality indicators, as a result of a single verification of the effectiveness of corrective actions, are given.







Keywords Gear Bevel gear Statistical methods Gear broaching Control cards





Quality improvement


D. T. Safarov (&)
A. G. Kondrashov Naberezhnye Chelny Institute, Kazan (Volga Region) Federal University, 68/10 Mira Ave, Naberezhnye Chelny 423800, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_28

249

250

D. T. Safarov and A. G. Kondrashov

1 Introduction Conical spur gears are part of the differential gear of heavy trucks. The differential ensures the independence of the rotation of the wheels of the car at different speeds of rotation. The design of the differential is such that in the process of redistribution of the moment at the same time four satellites are installed, mounted on the cross. In the process of working the differential, they simultaneously roll over the conical rims of two conical wheels, the inner splines of which are connected to the right and left axes, which in turn are connected to the wheel hubs. Any deviations in the relative position of the conical wheels, as well as the size of their holes, lead to disturbances in the relative position of the entire set of interacting bevel wheels at once, and to change or redistribute the area of the total contact patch. The unevenness of the area and location of the contact spots on the teeth of the rims of the conical wheels can lead to uneven loading of individual teeth and their intense local wear. The unevenness of the area and location of the contact spots on the teeth of the rims of the conical wheels can lead to uneven loading of individual teeth and their intense local wear or broken teeth. The importance of ensuring a given value and location of the total contact patch of a bevel gear in the differential mechanism lies in the fact that it is unregulated. The deviations resulting from the technological process in the size and distribution of the contact spots can not be compensated or redistributed by adjusting the position of the gears during the assembly of the differential.

2 Theoretical Part The complexity of solving the quality control problem of this conical wheel is to provide a total contact patch of the entire set of satellites installed in the differential mechanism. Providing a contact patch is possible only if there are no deviations of the shape and size of their base holes, ensuring a given level and relative location of contact spots along the rims, no deviations of the relative positions of the holes of the differential cups, the fitting size of the holes of the tapered wheels, and the absence of conical gear teeth in the deviations wheels. The use of local methods of quality management [1, 2], as is customary in many enterprises, according to individual indicators of accuracy from the above complex, shows its non-effectiveness. The result is the condition of finding individual, local inconsistencies, the elimination of which does not achieve the desired effect. For example, the value of the index of stability Cp found in the radial runout of teeth less than 1.33 found by engineering personnel indicates only a general inappropriate state of the tooth-drawing process or machine tool assemblies. The reason for the low coefficient Cp may be the presence of discrepancies in the geometric parameters of the workpiece, the deviation of the shape of the longitudinal and cross-section of the basic technological hole, the state of the cutting edges of the tool. Each of the above-listed deviations of the satellite leads to an

Methods of Quality Control Manufacturing Gears …

251

unacceptable distribution of the allowance over the side surfaces of the teeth and, as a consequence, significant deformations of the expanding mandrel in various positions of the processing of the cavities of the gear rim. The basic methodology does not provide for the collection of the required amount of information. The consequence of this is not the possibility of organizing an effective improvement in the quality of the satellite. To solve problems of this degree of complexity, a method for managing the quality of the manufacture of automotive parts has been developed, featuring complex application of statistical methods of quality management for all key indicators of part quality and a higher degree of identification of the conditions for performing the part machining process [3–6]. Consider the General content of the stages of this control method (Fig. 1). The initial stage of the methodology is the design stage, which is necessary for the preliminary analysis of the technological documentation of the process of manufacturing an auto component. At this stage, the structure of the intraoperative technological processes is highlighted, which allow us to reasonably identify a set of key characteristics that are formed at the workplace for processing the part. Next, a set of potential factors is identified that act in the process and affect its value of the key characteristics of the part. The development of an optimal scheme for measuring key characteristics is being carried out, which makes it possible to determine the presence of the effect of technological factors. The result of this stage is the development of a task for conducting a production experiment. As a result of its conduct, a batch of blanks is selected in accordance with the experiment execution plan. Billets are marked and measured, after which they are processed. After processing, measurements of key indicators of the parts are performed taking into account the marking applied to them. At the next stage, as a result of the consistent application of several methods of statistical analysis, conclusions are issued on mood, stability, the presence of specific causes, and the presence of hereditary connections between the measured values of key quality indicators. The presence of such a volume of simultaneously measured data for the whole range of indicators allows, in most cases, to develop the content of the necessary corrective actions to improve the manufacturing process. It should be noted that the formulation of corrective actions should be performed by a qualified technologist who knows the content of the technological process. The task given out by the technologist is a signal to carry out the corresponding corrective actions in the form. For example, the regulation of the position of individual components of the machine tool or repair work. A mandatory requirement of the methodology is the verification of corrective actions taken. Verification is performed by repeating the production experiment while preserving the processing circuit of the measurement data of key precision indicators. Evaluation of the effectiveness of verification is carried out by assessing the degree of achieved improvement in the values of key quality indicators. The degree of improvement can be assessed, for example, by the ratio of arithmetic mean key characteristics before and after taking corrective actions.

252

D. T. Safarov and A. G. Kondrashov

Fig. 1 Stages of the method of quality control of manufacturing parts of automotive components on the basis of integrated application of statistical methods of quality control in cutting operations

Next, we will consider an example of the implementation of individual stages of the methodology on the example of the operation of the tooth-projection of the crowns of the satellites.

Methods of Quality Control Manufacturing Gears …

253

3 Practical Implementation Example The technological process of manufacturing a gear wheel consists of the following basic operations, in which the formation of individual elements is performed: turning, roughing, and finishing gear operations, heat treatment operations, finishing boring operations, and coating operations. Starting with the gear operation, after making the crown of the bevel gear, each subsequent operation can change the relative position of its teeth and therefore affect the size and relative location of the contact spots. In the event that a contact spot of inappropriate area and (or) location is initially obtained in the dental surgery, subsequent operations can only worsen them or leave them unchanged. Therefore, it is important to ensure effective quality management of the satellite crown teeth for roughing and finishing operations. In production, the process of tooth-tightening is performed on ST 268 rough machines (Fig. 2a) and finishing tooth-tightening (Fig. 2b). They are rigidly connected to each other by an automated stepper conveyor (Fig. 2b) in an automatic line. During processing, the satellites are moved along the conveyor strictly in accordance with the stroke. The capacity of the conveyor is 14 satellites. The satellite preform is installed by the auto operator in the horizontal mandrel, and the broach rotates around the vertical axis and receives a consistent translational movement parallel to the line of the cavities of the gear. Both machines are structurally similar, the difference is in the performance of replaceable sections of circular broaching and copiers, which set the trajectory of their relative movement relative to the stretched workpiece. In the process of tooth pulling, it is necessary to provide such a profile and the relative location of the lateral surfaces of the teeth of the toothed crown, so that the contact spot is at the design point of the gear drive [7] (Fig. 3).

Fig. 2 Automatic gear-lining line a draft machine and b finishing gear reproducing satellite of rear axle differential, c automated step conveyor, d working area of gear reprocessing machine before performing the operation

254

D. T. Safarov and A. G. Kondrashov

Fig. 3 Key indicators of the differential satellite a the workpiece after rough tooth pulling, b the processed satellite

In shop conditions under the supervision of the personnel, there is only one indicator of quality of the satellite—beating of a gear ring on a dividing circle which is the only indicator of an assessment of technological accuracy of the gear automatic line for which measurements on a workplace there is a special control device. The value of the contact spot is estimated by visual comparison of the obtained contact spot with the permitted forms. Measurements and registration of its parameters are not performed. Adjustment of the gear machine is performed on the basis of visual images of the contact spot. The quality of adjustment depends largely on the level of qualification of the adjuster. Visual comparison of acceptable forms and location of contact spots by experienced and qualified personnel ensures the suitability of the products. Consider what set of key indicators is sufficient to measure for solving the problem of improving the complex parameters of the footprint of the satellite in the process of conducting the production experiment. The satellite billets should be measured in terms of beating on the middle cone, beating on the outer conical surface, the inner diametrical size (Fig. 3a). Processed satellite—to the runout at the middle cone, and the parameters of the contact spot (Fig. 3b). In the technological documentation, the requirements for limit values only for radial runout after finishing gear tightening and the internal diametrical size of the base hole are established. According to the contact spot, the requirements for its desired location, as well as the absence of an exit to the tooth boundaries are established. Measurement of geometric parameters of the satellites is performed using imamiyah on the site of the Guild of measuring instruments—special inspection fixtures measuring runout with a scale division of 0.01 mm. and indicating calipers with a scale division of 0.002 mm. As at the enterprise there is no practice of measuring the parameters of the contact patch, there is a need first to develop a method of normalization of the indicators, then the development of methods of measurement.

Methods of Quality Control Manufacturing Gears …

255

The basis of the developed technique is a well-known technique based on the standard method of normalization of the contact spot. The standard method establishes requirements for the size occupied by the contact spot in the direction of the tooth and its height [8]. Analysis of its content revealed its lack of information. The standard method of rationing and measuring the parameters of the contact spot assumes that the contact spot is already at the design point of the gear drive, i.e., the process of tooth pulling is in an established state (Fig. 4a). In practice, even in the established process of gear processing, the total contact spot can significantly shift relative to the design point of the gear drive (Fig. 4b), and in some cases can go to the boundaries of the teeth. Thus, this method is not applicable for assessing the location of the contact spot, so it is not suitable for developing comprehensive recommendations for improving quality. In addition to the size of the contact spot %H, %L (Fig. 5a) indicators of its relative location L1, L2, L3, L4—linear distances from the extreme points of the total contact spot to the boundaries of the active tooth surface are introduced (Fig. 5b, c). Thus, the key indicators for the contact spot are: showed the size of the contact spot %H, %L and indicators of the relative location of the contact spot L1–L4. The method of measuring the contact spot parameters is based on the processing of photographs of the side surfaces of the wheel teeth in a specialized software product. The volume of a batch of satellites in the amount of 30 parts sufficient for the production experiment was determined. This volume ensures minimum requirements for the objectivity of statistical calculations. More details increase the complexity of the method without improving the objectivity of the conclusions.

Fig. 4 Total contact patch of the contact of a conical gear train a the rationing of the size of the contact patch in a bevel gear I is the largest limiting size of the contact patch; II—nominal size of the contact patch; III—the smallest limit size of the contact patch [2], b photograph of the side surface of the satellite with the circumscribed boundaries of the contact patch

256

D. T. Safarov and A. G. Kondrashov

Fig. 5 Contact spot records (a the minimum and maximum allowable values of the indicators of the size of the total contact patch %H—along the height of the tooth, %L—along the length of the tooth; b the minimum allowable values of indicators of the relative position of the total contact patch; c the maximum permissible values of the relative position of the total contact patch)

4 Results of the Method Application As a result of the consistent application of various statistical methods of quality management for each key quality indicator, the content of corrective actions is formulated. The content of the activities is shown in Table 1. As you can see, the content of corrective actions is different. Some of them can be performed immediately. These include the replacement of the division disk in the spindle indexing device or the introduction of additional working stroke boring the base hole. Other corrective actions, such as replacement or modernization of the mandrel or the development of methods of adjustment of the machine cut can not be performed. For their implementation, it is necessary to plan and perform additional design and technological activities. After the corrective measures, the stability of quality indicators has increased. In terms of radial runout 1.3 times, in terms of %H%L 1.5 times.

Methods of Quality Control Manufacturing Gears …

257

Table 1 Corrective actions to improve the quality of manufacture of bevel wheels Indicator

Method

Analysis results

Content of corrective action

External cone runout

Primary data analysis

No grouping center The process is conditionally stable There are special reasons

–

Statistical analysis

Control charts

Runout of the pitch cone after draft gearing

Correlation analysis Primary data analysis Statistical analysis

Control charts Correlation analysis of the inheritance of the beats after the draft gearing

Runout of the pitch cone after finishing gearing

Contact Spot Size Indicators

Primary data analysis

Statistical analysis Correlation analysis of the hereditary relation of the beating after the finish machining and on the outer cone Correlation analysis of the inheritance of the heartbeat after finish and rough the machining Primary data analysis Statistical analysis

Control charts Correlation analysis of hereditary relations of contact patch size %L %H with a beating after finishing

– No grouping center The process is conditionally stable There are special reasons Unambiguous inheritance

There is a grouping center The process is stable There are special reasons

–

1. Improving the stability of the internal diametral size of the base element by introducing an additional transition to fine boring or pulling – – –

2. Using another mandrel with increased rigidity 3. Improving the stability of the external cone beats by introducing an additional transition of external turning of the external conical surface –

– 4. Improving the indexing accuracy of the expansion mandrel by replacing a worn division mechanism

Weak hereditary connections

–

No grouping center The process is conditionally stable There are special reasons Unambiguous hereditary relationships not identified

– –

5. Additional analysis of deviations of the profile, the development of methods for adjusting the machine, the development and manufacture of adjustment tools

(continued)

258

D. T. Safarov and A. G. Kondrashov

Table 1 (continued) Indicator

Method

Analysis results

Content of corrective action

Indicators of the relative location of the contact patch

Primary data analysis

No grouping center The process is conditionally stable There are special reasons Unambiguous hereditary relationships not identified

–

Statistical analysis

Control charts Correlation analysis of the hereditary relationships of the parameters L1, L2, L3, L4 and the beats after the finishing Beating difference after rough and fine gearing

Correlation analysis of hereditary bonds of the difference of the beats after the rough and fair gearing and parameters L1, L2, L3, L4

Unambiguous inheritance

–

6. Additional analysis of deviations of the profile, the development of methods for adjusting the machine, the development and manufacture of adjustment tools Corrective actions under paragraphs 1, 2, 3

5 Conclusion In general, the technique with a higher one-time complexity of the experiment by identifying all significant technological factors that lead to deviations of key indicators of details is more preferable than the use of control cards for single-quality indicators. The technique can be used to improve the quality of any part processing processes. It shows the greatest efficiency in the processing of parts of automotive components with a developed system of accuracy indicators—operations of processing of bevel gear rims [9], as well as critical parts of the car with high requirements for the accuracy of key accuracy indicators, such as cylinder liner [4] of the internal combustion engine or cylinder head [10].

References 1. Vasiliev VA, Odinokov SA, Borisova EV, Letuchev GM. Methods of quality management of innovation process. In: 2016 IEEE Conference on Quality Management, Transport and Information Security, Information Technologies, IT and MQ and IS 2016, Nalchik, Russian Federation, 4–11 October 2016, pp 239–241 2. Khafizov II, Nurullin IG (2017) Product quality as the main factor of increase of competitiveness (on the example of JSC “Kazan helicopter plant”). IOP Conf Ser: Mater Sci Eng 240(1):01203 3. Kasjanov SV, Kondrashov AG, Safarov DT (2017) Regulation of geometrical parameters deviations of automotive components parts through diagnostic measurements organization. Procedia Eng 206:1508–1514

Methods of Quality Control Manufacturing Gears …

259

4. Kas’yanov SV, Safarov DT (2004) Diagnosis of technical state of equipment and tools according to indices of technological accuracy. Avtomobil’naya Promyshlennost 5:24–28 5. Kondrashov AG, Kasyanov SV, Safarov DT, Kuznetsova AV (2013) Diagnostic measurements of geometrical parameters of spatial and difficult details of autocomponents by one-coordinate altimeter control. Diagnostics 8:60–64 6. Kondrashov AG, Safarov DT (2014) Prediction of accuracy when handling by cutting News of higher educational institutions. Mech Eng 12:63–69 7. GOST 19325-73 Transmission gear conical. Terms, definitions and designations 8. GOST 9368-81 Transmission gear conic fine-modular 9. Safarov DT, Kondrashov AG (2018) Technique for monitoring the indicators of the total contact spot of spur gears of the truck’s differential gear. IOP Conf Ser: Mater Sci Eng 457 (1):012014 10. Safarov DT, Kondrashov AG (2018) Improving the processing accuracy of the valve seats of internal combustion engines using diagnostic measurements. J Phys: Conf Ser 1050 (1):012075

Computational and Experimental Studies of Deformations of Air-Cooled Diesel Cylinders at Its Assembling I. E. Agureev, K. Yu. Platonov and R. N. Khmelev

Abstract The paper is devoted to the problem of reducing the deformation of the diesel cylinder 1H 9.5/8.0 by the inner diameter at the build stage. The design features of cylinders of single-cylinder diesel engines with air cooling are analyzed. The physical properties of the cylinder material were determined experimentally on a tensile testing machine. On the coordinate measuring machine, using specialized equipment, measurements and accumulated statistical material to change the internal diameter of the diesel cylinders from the action of installation efforts. The regularities of the deformations of the working surfaces of the cylinders of air-cooled diesel engines are obtained. The finite element mathematical model of deformation of cylinder walls from the influence of mounting forces is constructed. Three-dimensional models of the cylinder and its associated parts (cylinder head and crankcase) were used for modeling the operation of tightening studs. Computational experiments on the model were carried out using experimental data on the magnitude of the elastic modulus of the cylinder material. The influence of mechanical properties of the cylinder material and its design parameters on the character of deformations is investigated. The basic directions of the reduction of deformations of the cylinder of the diesel engine with air cooling are formulated. Alternative designs of the cylinder with a lower level of deformations and with a more developed fin belt are developed for better cooling. Recommendations to change the design and manufacturing technology of air-cooled diesel cylinders are given. Keywords Diesel modeling


Cylinder
Air cooling
Deformation
Mathematical

I. E. Agureev (&)
K. Yu. Platonov
R. N. Khmelev Tula State University, 92, Lenin Ave, Tula 300012, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_29

261

262

I. E. Agureev et al.

1 Introduction The cylinder is one of the most critical parts of internal combustion engines (ICE). The perfection of the design, manufacturing technology, and installation of cylinders at the initial stage can be estimated by the design deformations of the cylinder, which significantly affect the duration of engine running in, the stability of geometry in the conjugation of “cylinder-piston,” the probability of engagement (jamming) in the cylinder-piston group at the stage of running in and the ability to achieve the required engine output characteristics. The works are devoted to the solution of the problem of deformation of the engine cylinder [1–16], the analysis of which allows us to conclude that the engine cylinder with air cooling depends on several aspects, such as production technology, geometry and design features, material properties, and the impact of installation efforts. [17, 18]. In this work for a single-cylinder air-cooled diesel engine, the calculated and experimental studies of the cylinder deformation at the installation stage are done. The aim of the research is to assess the character and level of deformation of the 1H 9,5/8,0 diesel cylinder by the inner diameter at the stage of its installation.

2 Conducting the Experiment A significant influence on the nature of the deformation has such a parameter as the modulus of elasticity [19]. The modulus of elasticity of cylinder cast iron can be changed by the volume of the cylinder. Unfortunately, this is a disadvantage of casting technology. Therefore, the work carried out experimental studies of cast iron samples made from different parts of the same casting. The values of modulus of elasticity for the material of the cylinder were determined at different temperatures as a result of the experiment on a tensile testing machine. The static tensile test of the samples was carried out at temperatures (°C): 20; 100; 200; 400; 600; 800. Stretching was performed at a deformation rate of 5 mm/min. Dependence of elastic modulus on temperature will allow to study the next step of calculation to thermal deformation during operation of the diesel engine. The experiments were performed on machine INSTRON series 5982 (Fig. 1) computer-controlled, with a heavy frame and a working load of 100 kN, which is widely used for testing high-strength metals and alloys, advanced composites, aerospace and automotive structures, bolts, fasteners, and sheet steel. Mechanical characteristics were determined by the method described in Russian State Standard 1497-84 [20]. Also, a series of experiments by means of the coordinate measuring machine was carried out to measure the diameter of the cylinders in the free state and under the influence of installation efforts. For this purpose, a specialized tool was made to

Computational and Experimental Studies of Deformations …

263

Fig. 1 Testing machine INSTRON

simulate the assembly of the cylinder with the crankcase and the head, which does not prevent access to the measuring device to the cylinder. Figure 2 shows graphs of the cylinder wall movement for two perpendicular vertical sections in the free state and in the assembly. The standard tightening force of the nuts securing the head is equal 30–35 Nm. The results of the measurement nature of the deformation are remained. Comparison of the size and nature of deformation of cylinders with a diameter of 95 and 85 mm in the assembly allows us to conclude that the deformation is qualitatively identical. In this case, the cylinder with a diameter of 95 mm during

Fig. 2 Displacements of the cylinder wall (mm) in the free state and in the assembly

264

I. E. Agureev et al.

assembly in all belts has a significantly large deformation. In particular, the maximum change in the diameter of the cylinder diameter 85 mm is 11 µm, the cylinder 95 mm—26 µm. It should also be noted that when assembling the cylinder with a diameter of 95 mm in all the considered zones goes beyond the permissible values (on average by µm).

3 Mathematical Modeling Currently, the most effective method of studying cylinder deformations is mathematical modeling and computational experiment [3, 10–14]. In this paper, the ANSYS software package was used to study the deformations of the inner diameter of the cylinder [21]. The paper uses a static isotropic formulation of the problem of the theory of thermoelasticity, which is solved using experimental data on the magnitude of the elastic modulus of the cylinder material. One of the main stages of the engine installation is tightening the studs of the cylinder head, as it is associated with the application of considerable force to the cylinder. The choice of its size and the correctness of its distribution can affect the change in the cylinder geometry. The calculation of the axial compressive force applied to the cylinder cover caused by the action of the tightening moment was carried out according to the dependencies given in [22]. The stud–bolt force, equal to 27.6 kN, was calculated using the Formula 1.     P Mkl ¼ F0 0:5d þ 1:15f þ fT RT pd

ð1Þ

where the following definitions have been used: Mkl—torque; F0—the tightening force; d—thread diameter; P—thread pitch; F—reduced coefficient of friction; fT— coefficient of friction at the nut end; RT—radius of friction of the bearing surface of the nut. For modeling the tightening of the studs, three-dimensional models of the cylinder and its associated parts (cylinder head and crankcase) were used, shown in Fig. 3. Each body in an assembly unit has different materials and properties (Table 1). For all objects in the model, a finite element grid was generated. To accelerate the calculation, the parameters of the grid were set differently for different bodies. Maximum mesh quality is applied to the cylinder of greatest interest in this research. The grid parameters were determined by the method of trial calculations (Table 2) giving an acceptable result in terms of convergence of the solution. With further reduction of the size of the final element, the calculation result varies within 5%. Thus, it was possible to significantly reduce the calculation time without significant losses in the accuracy of the simulation. The simulation was carried out on the computer Asus N53S (Intel core i7-2670QM 2.2 GHz).

Computational and Experimental Studies of Deformations …

265

Fig. 3 Model of the cylinder assembled with head and the upper part of the crankcase

Table 1 Properties of materials of diesel parts Parameter

Cylinder

Head

Crankcase

Material

Special cylinder cast iron 7052 1.110−5

AК94-К-T6 State Standard 1497-84 2770 2.310−5

Special aluminum alloy 7850 1.2  10−5

22 180,000 0.25 841 180

22 71,000 0.33 875 –

22 200,000 0.3 434 –

Density (kg m3) Coefficient of thermal expansion (°C−1) Reference temperature (°C) Young’s Modulus (MPa) Poisson’s ratio Specific heat J/(kg К) Compressive ultimate strength (MPa)

Table 2 Parameters of the finite element mesh

Parameter

Value

Element type Number of nodes Number of elements The size of the element (for cylinder) Calculation time of one task

Tetrahedron 692,122 410,560 1 mm 4h

In the simulation, the force from tightening was applied along the axis of the studs to the cylinder head at the tides under the studs. The contact plane of the upper part of the crankcase is adopted absolutely rigid.

266

I. E. Agureev et al.

As a result of modeling, the data were obtained allowing to identify the dependence of the influence of effort during installation on the change of the cylinder geometry. The radial displacements of the inner surface points of the cylinder have an asymmetrical distribution (Fig. 4). This depends on the presence of grooves in the fins, which are made for the installation of studs and the rod of the pusher. The results of the calculation using the experimental data are compared with the results based on the standard table values of the elastic modulus of cast iron. Below are the diagrams (Fig. 5) showing radial displacements for cylinders with different elastic modules (1.1  1011 MPa and 1.8  1011 MPa) in several horizontal sections. A series of calculations taking into account the entire range of experimentally determined values of the elastic modulus is also carried out. The results of these calculations are presented as a field of points of coordinates of the cylinder walls at different elastic modules (Fig. 5) in sections 3 and 4 (section 4 is marked *).

Fig. 4 Character of radial movements

Fig. 5 Results of the calculation: a Displacements of the cylinder wall (µm) at the experimental (marked*) and table values of the elastic modulus, b the coordinates of the cylinder walls with different elastic modules for sections 3 and 4 (*)

Computational and Experimental Studies of Deformations …

267

Fig. 6 Design of the cylinder: a existing, b, c new designs

By results of comparison, it can be concluded that the displacements decrease with increasing modulus of elasticity. In addition to the characteristics of the material, the deformation pattern is influenced by the design features of the part. In the cylinder understudy, this feature is the technological notch in the belt of the cylinder fins (Fig. 5). This groove terminates in the zone of the fins which introduces an asymmetry and may affect the stiffness of the part. In this paper, a comparison of the mounting deformations of several variants of the cylinder design that differ in the number of cooling ribs and their size is made (Fig. 6). The character, level of deformations, and manufacturability of the design were taken as criteria for evaluating the design of the analyzed cylinders. The design of the cylinder (b) differs from the existing by increased number of cooling fins and their length from the bottom of the cylinder to the head. The top edge of the cylinder instead of the recesses for the studs is drilled, which makes the cylinder more rigid, in addition, improves the cooling efficiency, in consequence of the increased area of the fins with 14,300 mm2 for the existing cylinder (including 4700 mm2 is the part of the head) to 16,000 mm2 for new sample cylinder (including 3300 mm2 is the part of the head). These changes lead to a decrease in the value of displacements by 14–16% (Fig. 7). However, the production of the cylinder of this design technologically complicates the production process. The cylinder design (c) combines the positive properties of the researched cylinders. This design has a belt fins with an increase in the number of cooling fins and increases their height from the bottom of the cylinder to the head. The grooves are used instead of drilling to place the studs. This design has increased cooling efficiency at the top of the cylinder, because of additional cooling fins, and this construction is more technological.

268

I. E. Agureev et al.

Fig. 7 Deformation of the existing cylinder (left) of the cylinder and the new design. The cylinder 3 has a smaller (14–15%) magnitude of displacement and a fairly even distribution

4 Conclusion In the present study, calculation and experimental methods were solved the problem of the deformation analysis of the diesel engine cylinder 1H 9,5/8,0 with a view to their reduction. For the model, temperature dependences of the modulus of elasticity of the real material were found, which were used for the calculation at the stage of determining the cylinder assembly deformations due to the tightening of the head studs. In subsequent studies, it is assumed to obtain a superposition of temperature and assembly deformations. The obtained deformations are asymmetric in nature associated with the geometry of the grooves in the fins of the cylinder. Revealed the nature of the radial deformations of the points of the working surface of the cylinder is qualitatively consistent with the results of similar studies for other size diesels from the same manufacturer. It is found that the deformation patterns are almost identical for all analyzed cylinder samples in the assembly with a head and do not depend on the original geometry. The maximum values of deformations are observed in the D-D and B-B cross sections of the order of 40–50 lm. The obtained results allow us to formulate the main directions of reducing the deformations of the diesel cylinder with air cooling: • change of the scheme of fastening • change of tightening torque • mechanical treatment of the inner surface of the cylinder (honing) assembled. For the considered design of the diesel engine, the last option is the most preferable.

Computational and Experimental Studies of Deformations …

269

References 1. Ageev AG (2017) Reduction of mechanical losses in the high-speed air cooling diesel engine by improving the design of parts of the CPG. Dissertation. Moscow, BMSTU, 177 p 2. Agureev IE, Platonov KYu, Khmelev RN (2018) Analysis of the laws of deformation of the diesel engine cylinder with air cooling from the actions of assembly efforts. In: Progress of vehicles and systems-2018: proceedings of the international scientific-practical conference. Volgograd State Technical University, pp 52–53 3. Platonov KYu, Khmelev RN (2017) Modeling and analysis of deformations of the cylinder of a diesel single-cylinder engine in the assembly stage. In: Design, use and reliability of agricultural machinery. Publisher, Bryansk State Agrarian University, pp 274–278 4. Chinov ND, Ivanova NS, Malashenko NG (2017) Calculation of thermal and stress-strain state monobloco high-speed marine diesel engine. News of higher educational institutions 10:34–42 5. Gutieva NA (2007) The calculation of the deformation of the supporting ribs of the cylinder liners and the block-crankcase of diesel engines, Vestnik DSTU. Technical Science 13:62–63 6. Solovev VL, Kornilovich SA (2015) Research of influence of the quality of cylinder heads installation on safety. Vestnik OMGAU, pp 55–61 7. Kang Q-Y, Cao X-H (2015) Cylinder block deformation evaluation method research and its structural optimization for diesel engine. Neiranji Gongcheng Chin Intern Combust Eng Eng 36(2):93–97 8. Ma Z, Henein NA (2002) Cylinder liner surface analysis during si engine break-in. Tribol Trans 45(3):397–403 9. Lan LP, Xiang JH, He LG (2015) Deformation characteristics of diesel engine cylinder liner under pretightening condition. Neiranji Xuebao/Trans CSICE (Chin Soc Intern Combust Eng) 33(6):555–561 10. Xiang R, Bi Y, Lei J, Song G, Shen L, Xu Y (2015) Study on influencing factors to the cylinder liner pre-tightening deformation of turbocharged inter-cooled diesel engine. Jixie Qiangdu/J Mech Strength 37(4):682–688 11. Guan L, Zeng Q, Liu B, Hu TG, Zhan ZS (2013) Simulation and test comparison for internal combustion engine cylinder liner deformation. Neiranji Xuebao/Trans CSICE (Chin Soc Intern Combust Eng) 31(5):473–479 12. Ma QZ, Jiang SL, Guo CH, Yao XL, Gu Q, Dong B, Ye Q (2008) Analysis of deformation of cylinder liner of YZ4DE diesel engine based on finite element method. Neiranji Gongcheng/ Chin Intern Combust Eng Eng 29(4):59–62 13. Fujimoto H, Yoshihara Y, Goto T, Furuhama S (1991) Measurement of cylinder bore deformation during actual operating engines. In: SAE Technical Papers 1991 International Congress and Exposition, Detroit, MI, United States 14. Bulatov VP, Bochkarev VN, Yakhyaev NYa (1988) Evaluation of assembly deformations of cylinder liners for small diesel engines, Strength of Materials 20(6):826–831 15. Liang X, Wang Y, Huang S, Yang, G, Tang L, Cui G (2017) Investigation on cylinder bore deformation under static condition based on fourier decomposition. SAE Technical Papers, vol. 2017-March, is. March, 28 March 2017 SAE World Congress Experience, WCX 2017, Cobo CenterDetroit, United States 16. Bi YH, Xiang R, Lei JL, Shen LZ, Zhang N, Song GF (2015) Study on deformation of diesel engine cylinder liner under different loads. Neiranji Gongcheng/Chin Intern Combust Eng Eng 36(3):130–139

270

I. E. Agureev et al.

17. Yakhyaev NI, Vagabov NM (2009) Complex method of analysis of geometrical accuracy of cylinders in the process of assembly of small-size marine diesel engines. AGT messenger. Ser Mar Eng Technol 1:256–261 18. Vagabov NM, Verdiev MG, Sokratov TE (2011) Way to reduce assembly deformation of sleeves of cylinders (CC) by correcting the error forms the base surfaces of the belts of the block-crankcase. Bull DSTU Tech Sci 21:88–91 19. Efros VV (1976) Air-cooled diesel engines of Vladimir tractor plant. Machine Building, Moscow, 277 p 20. State Standard 1497-84 Metals. Methods of tensile testing 21. Engineering Simulation and 3D Design Software ANSYS. Available via DIALOG. https:// www.ansys.com. Accessed 13 Oct 2018 22. Birger IA, Iosilevich GB (1990) Threaded and flanged connections. Machine Building, Moscow, 368 p

Feasibility Study of Applying Group Interchangeability Method for Assembly Components of Machines I. I. Voyachek, E. S. Slashchev and D. M. Malikova

Abstract The paper studies the comparison of the group interchangeability method with the full interchangeability method of the technological assembly process in the case of the ladder dimensional circuit complete interchangeability for the assembly components of machines. With the method of group interchangeability, a several times increase of component links dimensions tolerance is achieved as compared with the method of full interchangeability with the introduction of an additional sorting operation. The feasibility study consists of reducing the equipment requirements, the technological machining step and operations number, technological and special accessories, workers’ skills. A significant drawback of the group interchangeability method is the unfinished production occurrence, to reduce which it is necessary to apply different methods of regulation master link sizes. In this regard, it is necessary to conduct a feasibility study of applying the group interchangeability method for each specific case using the methods given in the article. It also provides recurrence equations to formalize the settlements of upper and lower deviations with the tension for ladder dimensional circuit by using the group interchangeability method.




Keywords Group interchangeability Selective assembly production Full interchangeability Feasibility study








Unfinished

1 Introduction An effective method of ensuring the master link accuracy, which allows to reduce the requirements for the component links (parts) production tolerance value, is the method of group interchangeability (selective assembly) [1–9]. But due to the I. I. Voyachek Penza State University, 40, Krasnaya Street, Penza 440026, Russia E. S. Slashchev (&)
D. M. Malikova Izhevsk State Technical University, 7, Studencheskaya Street, Izhevsk 426069, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_30

271

272

I. I. Voyachek et al.

unfinished production occurrence, the lack of formalized calculation methods for ladder dimensional circuit, the subjective factors influence, this method is not widely used in machine-building enterprises. In detail, the conditions for the unfinished production occurrence when assembling by the group interchangeability method are discussed in the professor Lebedovskiy works [10]. This happens when the distribution laws of the links sizes (random variable) sum that are in the increasing path and decreasing path of the dimensional chain do not coincide. At the same time, the use of the group interchangeability method can significantly reduce the labor input and manufacturing cost even for small batch workpieces. The cost efficiency justification method is considered on a specific example, taking into account the dimensional chains calculation. Also, the paper provides recurrence equations to formalize calculations for ladder dimensional circuit by the use of group interchangeability method. More detailed description of how to compensate errors by the group interchangeability method is described in [11–20].

2 Ladder Dimensional Circuit Calculation by the Group and Full Interchangeability Method Consider the task of achieving the necessary assembling device accuracy (Fig. 1). The calculation is made for the attachments batch of 100 pcs. The calculation is performed by the full interchangeability (Table 1) method and the group interchangeability method by the equal tolerances Td = TD way. In order to ensure the increased device component part operability, to avoid jams and getting processing waste into the gaps, it was decided in the assembly process via the ring 4 to provide tightness in the range from Nmin = 0.2 mm to Nmax = 1 mm, which is the master link. At the same time, in order to increase wear resistance, strength characteristics, and durability, it is proposed to use an elastic link, which creates pressure between the blades and the hole plate under tightening. Fig. 1 Device sketch and the dimensional chain scheme

Feasibility Study of Applying Group Interchangeability Method …

273

Table 1 Component links dimensions and tolerances (full interchangeability method) Link

Nominal dimension (mm)

Tolerance (mm)

Link name

A1

80

0.3

A2 A3 A4

5 10 63

0.06 0.08 0.3

A5 AD

2 0

0.06 0.8

The distance from the right frame end 1 to the left ledge Hole plate thickness 5 Blade thickness 6 The distance from the right auger end 7 to the left end Elastic link thickness Master link (tightness)

The tolerances indicated in Table 1 are difficult to realize in manufacturing, and therefore, it is proposed to apply the group assembly method. Calculation of the dimensional chain by the group interchangeability method using the derivative recurrence formulas is going to be provided: (a) determine the tightness tolerance: TN = 1 – 0.2 = 0.8 mm; (b) determine the group tolerance for the dimensional chain increasing and decreasing paths: Tgr ¼ 0:8=2 ¼ 4 mm; (c) assign, on the basis of economic expediency, P the sum P of tolerances for dimensional chain reducing and increasing paths Td = TD = 1.2 mm; (d) determine the group’s number: n = 1.2/0.4 = 3; (e) using “Eqs. 1, 2” calculate the group upper and lower deviations for the reducing link A1 (one link in the reducing path), while taking Td = 1.2 mm:    i1 esAðiÞ ¼ Td 1  n    i eiAðiÞ ¼ Td 1  n

ð1Þ ð2Þ

   when i = 1: esAð1Þ ¼ 1:2 mm; eiAð1Þ ¼ 0:8 mm; i = 2: esAð2Þ ¼ 0:8 mm;    eiAð2Þ ¼ 0:4 mm; i = 3: esAð3Þ ¼ 0:4 mm. eiAð3Þ ¼ 0 mm. (f) determine the total upper and lower deviations for dimensional chain increasing path using “Eqs. 3, 4”  Td ESAðiÞ ¼ ð2 þ n  iÞ þ Nmin n

ð3Þ

 Td EIAðiÞ ¼ ð1 þ n  iÞ þ Nmin n

ð4Þ

274

I. I. Voyachek et al.

Table 2 Remaining values are summarized Groups (n = 3)

A1 = 80 es

ei

ES

EI

ES

EI

ES

EI

ES

1 2 3

1.2 0.8 0.4

0.8 0.4 0

0.9 0.8 0.7

0.8 0.7 0.6

0.3 0.2 0.1

0.2 0.1 0

0.3 0.2 0.1

0.2 0.1 0

0.3 0.2 0.1

A2 = 5

A3 = 10

A4 = 63

A5 = 2

TN

EI

Tightness (N) max min

0.2 0.1 0

1 1 1

0.8 0.8 0.8

0.2 0.2 0.2

where i = 1, 2 …, n is the current variable characterizing the group’s number. The remaining values are summarized in Table 2. (g) Using “Eqs. 5, 6” for the ladder circuit, we calculate the group upper and lower deviations for the increasing links using the equal tolerance method, the reducing links number mred = 1, increasing minc = 4 (Fig. 1): Qm

  ðk  ðj  mred ÞÞ TD ði  1Þ þ 1 n ðm  1Þ! minc Qm   ! ðk  ðj  mred ÞÞ TD ðiÞ þ 1 EIAði; jÞ ¼ ðTd =n þ Nmin Þ k¼2 ðm  1Þ! n minc

! ESAði; jÞ ¼ ðTd =n þ Nmin Þ

k¼2

ð5Þ ð6Þ

where j = 1, 2 … m; k = 2, 3 … mred—the current variables corresponding to the component links number (m  2). The remaining values are summarized in Table 2.

3 Economic Justification General Principles In general form (without unfinished production compensation), the manufacturing cost of an assembly unit (node points) batch when applying the group interchangeability method is as follows “Eq. 7”: C1 ¼ Cc þ Cs þ Cao þ Cup

ð7Þ

where Cc—is the part cutting cost; Cs—the sorting part cost; Cao—the assembly operation cost; Cup—the manufacturing parts remaining after assembly cost (unfinished production). When applying the compensation method, the unfinished production cost is not taken into account. It increases the savings, but at the same time, the compensation cost should be taken into account. When using the full interchangeability method, the assembly unit manufacturing cost is “Eq. 8”:

Feasibility Study of Applying Group Interchangeability Method …

C2 ¼ Cc þ Cao

275

ð8Þ

If the costs concerned with sorting parts into groups and the unfinished production volume are less than the savings concerned with part cutting, then the group interchangeability method is justified.

4 Unfinished Production Determination During Selective Assembly One of the group interchangeability method disadvantages is the introduction of sorting parts into groups operation. Also, almost inevitably there is an unfinished production, which leads to the need to increase part batches (“reserve”). To ensure an assembly process complete cycle (to ensure a given output volume), to exclude the parts lack, it is necessary to estimate the probabilistic unfinished production volume, which is determined on the specific statistical data basis. Based on statistical data, the distribution density of the mating parts actual dimensions in the assembly unit for the increasing and decreasing paths is assessed: size A1 is the reducing path (for size tightness), and the sum of the sizes P Ainc = A2 + A3 + A4 + A5 is the increasing path (Fig. 1). In the product batch manufacture, N = 100 pcs. The reducing path sizesPwere measured—size A1 (Fig. 2a), and the increasing path—the sizes sum of the Ainc links (Fig. 2b). The measurement results in the deviations from the nominal value are summarized in Table 3. Based on statistical data, it can be assumed that the size deviations empirical distributions are close to the normal principle with parameters: For the reducing path Ared, the average deviations value is xav.red = 0.56 mm, the standard

Fig. 2 a Reducing size measurement A1, b sum measurement of increasing A2, A3, A4, A5 sizes

276

I. I. Voyachek et al.

Table 3 Actual size deviations distribution Reducing link A1 from the nominal value Deviation Frequency Frequency distribution (m pcs.) of interval (mm) occurrence (m/N) From To

Increasing links from nominal value Deviation Frequency Frequency distribution (m pcs.) of interval (mm) occurrence (m/N) From To

0 0.17 0.34 0.51 0.68 0.95 1.03

0.6 0.78 0.95 1.12 1.29 1.46 1.63

0.17 0.34 0.51 0.68 0.85 1.03 1.2

2 16 22 26 19 13 2

0.02 0.16 0.22 0.26 0.19 0.13 0.02

0.78 0.95 1.12 1.29 1.46 1.63 1.8

3 17 23 28 20 14 2

0.03 0.17 0.23 0.28 0.2 0.14 0.02

Fig. 3 Scheme for determining the unfinished production volume

(root-mean-square) deviation is Sred = rred = 0.197 mm; for the increasing path Ainc is xav.inc = 0.6 mm, Sinc = rinc = 0.198 mm. When the distribution laws are superimposed for the increasing and decreasing paths, a shift is detected (Fig. 3), which reflects an incomplete product assimilability. The shaded section areas are proportional to the unassembled dimensional circuit number due to the impossibility of ensuring the master link accuracy. If rinc = rred, but xav.inc 6¼ xav.red, then the unfinished production volume is according to “Eq. 9”: 82 0:58 3 2 1:18 39 Z0:58 Z Z1:12 < Z = PðxÞred dx  PðxÞinc dx5 þ 4 PðxÞinc dx  PðxÞred dx5 Q¼N 4 : ; 0

0:02

0:58

0:58

ð9Þ where N is the part number in the batch.

Feasibility Study of Applying Group Interchangeability Method …

277

The integrals in “Eq. 9” are calculated using the tabulated normalized Laplace function U(t) when defining the corresponding quantiles t = x/r. Taking into account the function symmetry and parity, we obtain “Eq. 10”: Q ¼ Nf½0:5 þ Uð0:02=rred Þ  ½0:5  Uð0:02=rinc Þ þ ½0:5 þ Uð0:02=rÞ  ½0:5  Uð0:02=rred Þg

ð10Þ

Based on statistical data, rred, rinc 0.2 mm can be taken. Then, “Eq. 10” changes into “Eq. 11”: Q ¼ 4Uð0:1Þ
N ffi 4
0:04
N ¼ 0:16
N

ð11Þ

Thus, the unfinished production in this case is 16% of the batch volume.

5 Economic Expediency Determination of the Selective Assembly Use The manufacturing part costing included in this assembly unit (device) for group assembly (extended tolerance) is shown in Table 4 above the line, and the full interchangeability method calculation results are shown in Table 4 under the line. One-time costs for sorting parts into groups in a batch of 100 pieces (sorting and storage price) are set on the manufacturing stage and are 1300–1600 rub. Let us assume that for one product Cs = 14 rub. The unfinished production part cost is determined by the relation “Eq. 12”: Cup ¼ Q  ðCm þ Cs Þ ¼ 0:16  N  ðCm þ Cs Þ

ð12Þ

The labor input of the assembly operation is 0.099204 working hour (w/h) at a 1 w/h manufacturing work cost of 112.20 rubles. The cost price of assembling the device is equal to Cas = 11.31 rub. Considering all cost items, see Table 4, we obtain the final relation to determine the group interchangeability and full interchangeability cost “Eqs. 13–14”: Cgr ¼ ðCppðgrÞ þ Cs Þ
ðN þ 0:16NÞ þ Cas
N ¼ ð121:2 þ 14Þ
1:16
N þ 11:3
N ¼ 168:11
N ðrub:Þ ð13Þ CppðfiÞ ¼ 205:97
N þ 11:3
N ðrub:Þ

ð14Þ

Frame

1

1

Quantity (pcs.)

M

W

SI

Cspe

SE

Com

Coe

Cpp(gi)/Cpp(fi)

Labor input (w/h)

18.23 17.71 5.65 1.12 3.88 3.54 1.77 51.9 0.20819 18.23 49.565 15.86 2.52 3.88 3.54 1.77 95.365 0.58311 2 Auger 1 4.030 5.73 1.83 0.37 1.26 1.15 0.57 14.94 0.06737 4.030 26.328 8.42 0.85 1.26 1.15 0.57 42.608 0.30969 3 Hole plate 1 7.193 5.086 1.62 0.254 0 2.136 1.526 23.22 0.05976 7.193 15.463 4.95 0.609 0 2.136 1.526 31.877 0.18193 4 Screw nut 1 1.910 4.27 1.36 0.27 0,94 0.85 0.43 10.03 0.05017 1.910 10.545 3.38 0.486 0,94 0.85 0.43 18.541 0.12410 5 Blade 1 4.461 9.461 3.018 0.473 0 3.974 2.838 23.132 0.11122 4.461 37.747 12.07 1.489 0 3.974 2.838 62.579 0.44400 Note The cost is in rubles. M is for material costs; W is for wages; SI is for social insurance contributions; Cspe is for special process equipment costs; SE is for special expenses; Com is for overall manufacturing costs; Coe is for overall economic costs; Cpp(gr) is for part production cost (group interchangeability), and Cpp(fi) is for part production cost (full interchangeability)

Part name

№

Table 4 Part cutting device cost calculation (group interchangeability method/full interchangeability method)/

278 I. I. Voyachek et al.

Feasibility Study of Applying Group Interchangeability Method …

279

Thus, applying the group interchangeability method economic effect will be “Eq. 15”: E ¼ Cgr  Cfi ¼ 93:17
N ðrub:Þ

ð15Þ

6 Conclusion One of the group interchangeability method defective features is the additional introduction of the sorting parts into groups operation. Also, almost inevitably, there is an unfinished production, which leads to the need to part batches increase (“reserve”). To ensure the full assembly process cycle (to ensure a given output volume), to exclude the parts lack, it is proposed to use the considered methodology for estimating the unfinished production probabilistic volume, which is determined in terms of specific statistical data. Posed method of assessing the use of the group assembly method economic effect, taking into account the unfinished production volume in comparison with the full interchangeability method, will allow appraising the introduction effect. So in the case of the considered device group assembly implementation in the production, with the annual products release Na = 10,000 pcs., the economic effect will be E = 931,700 rubles.

References 1. Anurev VI (2001) Reference book for designer-mechanical engineer: In 3 v. vol 2., 8th edn. zrev. and add. In: Zhestkova. Mashinostroenie IN (ed) Moscow, 912 p 2. Kolesov IM (1999) Fundamentals of mechanical engineering technology: a textbook, 2nd edn. Corr. Higher. School, Moscow 3. Nepomiluev VV, Oleynikova EV, Gusarova NI (2015) Probabilistic-statistical model of the process of individual selection. Intell Syst Prod 1:8–13 4. Osetrov VG, Slashchev ES (2015) Assembly in mechanical engineering, instrument making. Theory Technology and Organization, IICP, Izhevsk 5. Osetrov VG, Slashchev ES (2014) Improving the calculations of the dimensional chain when using the method of group interchangeability. Assemb Mech Eng Instrum Mak 7:24–29 6. Osetrov VG, Slashchev ES (2014) The calculation of the accuracy of compounds with tension when using the method of group interchangeability. Intell Syst Prod 2(24):52–56 7. Scholz F (1995) Tolerance stack analysis methods research and technology. Boeing Information and Support Services 8. Sorokin MN, Anurov YN (2011) Formalization of the method of intergroup interchangeability in the implementation of selective assembly of products. Assemb Mech Eng Instrum Mak 8:16–19 9. Sorokin MN, Anurov YN (2016) Rolling bearing assembly. Assemb Mech Eng Instrum Mak 2:18–23 10. Lebedevsky MS, Weitz VL, Fedorov AI (1985) Sci Basis Autom Assemb. Mashinostroenie, Moscow

280

I. I. Voyachek et al.

11. Altschul R (1994) Case study in statistical tolerancing. Manufacturing Review of the AMSE 7:52–56 12. Bezyazyachny VF, Nepomiluyev VV (2015) Possible ways to improve the quality of manufacturing engineering products. Assembly in mechanical engineering, instrument making 1:17–20 13. Chase KW, Gao J, Magleby SP, Sorensen CD (1996) Including geometric feature variations in tolerance analysis of mechanical assemblies. IE Trans (Inst Ind Eng) 28(10):795–807 14. Goldfarb V, Malina O, Trubachev E (2016) New concept of the process of designing gearboxes and gear systems. Mech Mach Sci 34:405–423 15. Henzold G (2006) Geometrical dimensioning and tolerancing for design manufacturing and inspection, 2nd edn. Elsevier. Oxford, UK 16. Laaneots R (2004) Modified calculation method of tolerance of dimensional chain dependent link. In: 4th international DAAAM symposium. Industrial engineering—innovation as competitive edge SME. Tallinn, Estonia, pp 43–46 17. Mishunin VP, Osetrov VG (2002) Optimization in achieving the accuracy of the axial clearance in the gearboxes. Assemb Mech Eng Instrum Mak 6:2–4 18. Srinivasan V (2008) Standardizing the specification, verification, and exchange of product geometry research, status and trends. Comput Aided Des 40(7):738–749 19. Sun Y, Gupta M (2004) Optimization of a flat due. Mechanical Engineering—Engineering Mechanics Department Michigan Technological University Houghton, ANTEC, pp 3007–3011 20. Voyachek II (2006) Integrated design of fixed couplings: monograph. Penza State University, Penza

Development of Integrated Criterion to Select Environmentally Sound Cutting Fluids and Relevant Application Systems in Shape-Forming Processes L. E. Shvartsburg, O. V. Yagolnitser and E. V. Butrimova

Abstract The article covers the building of an integrated criterion to select environmentally sound cutting fluids to take into account a number of parameters (process, environmental and economic ones) and their values. While building the integrated criterion, more focus was put on the parameters indicating how cutting fluids affect the environment and human beings. The integrated criterion is a set of parameters and characteristics of cutting fluids (CFs) and relevant application systems, which varies in line with the significance that we can assess for such parameters. Our method to build the integrated criterion is a kind of weighted criteria method. The article gives an example how to build the integrated criterion to select environmentally sound cutting fluids for such applications as automatic lathes and semi-automatic machines that are used to manufacture load distributors. Below we show the selection steps and the outcomes.







Keywords Mechanical engineering Technological process Cutting fluids Environmental and human impact Integrated criterion Algorithm Integral criterion













1 Introduction It is reasonable from the process and economic points of view to use cutting fluids in such shape-forming processes as cutting of various metals. However, such issues as the environmental safety when using cutting fluids along with the assessment of their impacts on the environment and humans are still controversial and partially open. Any use of cutting fluids entails a number of negative consequences for the environment: waste generation (the CF oil sludge), working environment and ambient air polluted with sprayed cutting fluids, and withdrawal of valuable natural resources to manufacture the cutting fluids. On the other hand, the process L. E. Shvartsburg
O. V. Yagolnitser
E. V. Butrimova (&) MSTU “STANKIN”, Vadkovsky Lane, 3a, Moscow, Russia 127055 e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_31

281

282

L. E. Shvartsburg et al.

efficiency when using cutting fluids greatly contributes to the environmental safety and mitigates other impacts on the environment and humans. Among other things, using cutting fluids results in lesser cutting forces, which reduces energy consumption, which, in turn, is positive for the environment, increases tool life and contributes to forming the surface layer with required specifications. In addition, proper selecting of cutting fluids and their application systems is a way to mitigate the above-mentioned negative effects on the environment and humans [1–10].

2 Relevance The feasible selection of cutting fluids and their application systems is always a difficult task. If certain constraints (such as environmental safety, mitigation of impacts on the environment and humans, need for improved tool durability and lesser cutting forces) are imposed, the task turns out to be even more complicated. At the moment, many decision-making support systems are used in machine engineering, in particular, the systems (various expert systems, databases and knowledge bases) support selecting of cutting fluids. These tools can be a part of CAD systems (CAD, CAM, CAE systems) and used in CALS technologies. Some local and foreign publications cover relevant research on this issue [11–17]. Upon reviewing such publications, we see that the great attention is paid to the methodology of building such decision-making support systems and gaining operational excellence. Our review of the applications of the cutting fluids has revealed the diversity of CF specifications and consequences of the CF use, which further complicates any feasible selection [18, 19]. At present, certain specialized information retrieval systems have been developed as search engines for CF specifications along with databases of tools, process equipment and operational best practices. However, there is an urgent need to generalize the available information about technical, economic and environmental properties of cutting fluids and their application systems in the form of a single database, establish links between parameters and implement a computer-aided process to select cutting fluids. The above explains the relevance of the presented article.

3 Problem Formulation The article solves the problem of building an integrated criterion to select environmentally sound cutting fluids and their application systems to ensure the minimum impact of cutting fluids on the environment and humans, as well as the problem of developing an algorithm and method to select cutting fluids and their application systems.

Development of Integrated Criterion to Select Environmentally …

283

4 Theory To select an environmentally sound cutting fluid and its application system, a database (DB) containing many parameters of cutting fluids and their application systems was developed along with the ECO SOTS information retrieval system (IRS). One of the main tools of the database is an integrated criterion implemented in the IPS. According to the developed selection algorithm [1, 20] based on the user-defined constraints for search parameters, a certain subset of solution options X = {x1, …, xk}, k  n is selected from the entire set of initial options X = {x1, …, xn}. Further search is performed within the subset by assessing options under the integrated criterion. The developed assessment algorithm covers process, environmental and economic parameters of a process system where a cutting fluid is used (Fig. 1). If a user is not satisfied with a selected option, the decision-making process resumes with new, updated or modified parameters until the final result is obtained [21]. The integrated criterion is built taking into account certain constraints set for links and requirements to CF parameters. It is a set of parameters, CF specifications and their application systems, which may vary depending on the user-defined significance of such parameters. The integrated criterion includes process, environmental and economic parameters of a process system where a cutting fluid is used. To select one or more best solutions from the X* subset, an adapted model to assess options based on the weighted criterion method is used. The parameters of the integrated criterion are set, and weight is assigned to each parameter depending on its significance. Each option for the subset of solutions X* formed by the constraints is assessed. Each option has a set of attributes that show its properties. A numerical score is assigned to each attribute with a mapping u: ei = u(Ni); i = 1, …, k, where ei is an assessed alternative of xi for criterion u(Ni). This information

Fig. 1 High-level algorithm to access options under integrated criterion

284

L. E. Shvartsburg et al.

is compiled upon pre-examination and based on codes, standards and empirical data. Each option is assessed in view of components of the integrated criterion, and its integral assessment is carried out by multiplying the resulting assessment value by the criterion weight, which is followed by the summing of resulted values for each option. The options where the cumulative weights are higher are solutions to the decision-making problem. The above algorithm forms the basis to develop a method to select cutting fluids and their application systems using the integrated criterion [20, 22]. A distinctive feature of the integrated criterion is its wide range depending on a certain problem. The integrated criterion is built by assigning weights to parameters included in the integrated criterion. When solving a selection problem, only certain parameters are significant for such selection. When solving a particular problem, weights are assigned to these parameters depending on their significance. The remaining parameters of the integrated criterion are not taken into account. A user who is building a decision-making problem shall assign such weights. A weight of each parameter included in the integrated criterion is additional information used when special decision-making methods are applied. A parameter weight makes it possible to assess the significance of such parameters (criteria) forming the integrated criterion, and to compare them in terms of their significance. The integrated criterion is built via assigning or not assigning weights to the parameters of the integrated criterion and via selecting values for such parameters. The integrated criterion is built depending on the goal to be achieved when selecting, which may be done in three ways [23, 24]: • Building criterion by user. In this case, the user himself defines parameters that affect the selection. This method is applicable when the right option depends on a set of various factors that are weakly linked one to another, for example, when the right option depends on any physical or chemical properties of the product (CF viscosity, density, etc.) and its price. • Building the integrated criterion by professionals in a certain area of expertise. This method is recommended to consider specific conditions in the course of selecting. The area of expertise elements and their parameters behind certain conditions are defined by experts. Thus, the parameters of the integrated criterion are defined to keep focus on, for example, CF toxicity, cost efficiency of using a particular system and energy consumption. • Mixed method of building the integrated criterion. In this case, a part of the parameters is defined by user, while the other part is used in the form of “parameter clusters” built by experts to take into account all required parameters including ones behind specific conditions. A certain weight is assigned to each parameter included in the integrated criterion. The article uses a 10-score index to assess parameters: zero score means that an assessed parameter has presumably no value for making a decision. To start building the criterion (building a cluster of parameters), first, we should divide the parameters into two groups: having a weight of 1 or more, and having

Development of Integrated Criterion to Select Environmentally …

285

zero weights. If you need to define a subset of parameters that affect decision-making, and if there is no need to define more or less significant parameters within this subset, then a weight of 1 is assigned to all parameters included in this subset, and a zero weight to remaining parameters. Such clusters of parameters are built in view of logical chains of relationship between various hazards and process environment where cutting fluids are used and in view of CF specifications.

5 Practical Significance and Implementation Results As an example of an integrated criterion built using developed ECO SOTS database, we consider CF selection decision-making process and CF regeneration systems for 28 automatic lathes and semi-automatic machines in the workshop where load distributors are manufactured at MTZ TRANSMASH OJSC. The decision-making process included considering production, environmental and economic criteria and constraints. According to the algorithm to select cutting fluids and their application systems and to the algorithm to assess options using the integrated criterion, the following steps took place. 1. Setting constraints by a user: Setting the top and bottom limits for search parameters (Table 1). 2. Compiling a request to ECO SOTS DB in view of set constraints. 3. The query returns a subset of options to select cutting fluids and their regeneration systems, which satisfies the set constraints. 4. Defining the most important parameters of the integrated criterion to solve the selection problem and setting weights to such parameters (Tables 2 and 3). 5. Assessing options from the subset formed by imposing constraints using parameters of the integrated criterion. A certain weight is assigned to each option for each parameter. The weight of selection options is determined using expert assessments or codes and standards and engineering data.

Table 1 Constraints for options Process constraints

Environmental constraints

Economic constraints

1. Equipment: Automatic lathes 2. Number of equipment units: 28 pieces 3. Work piece material: structural carbon and alloy steels, heavy-duty steels 4. Cutting mode: medium, heavy

1. CF hazard class is below or equal to IV 2. Fineness of treatment of cutting fluid: nominal diameter of particles is not more than 10 lm 3. Sulphur content is not more than 6% (wt%) 4. Chlorine content is not more than 14% (wt%)

1. Capital expenditures are not more than RUB 800,000 2. Current expenditures are not more than RUB 500,000 per year

286

L. E. Shvartsburg et al.

Table 2 User-defined parameters of integrated criterion Production criteria

Environmental criteria

Economic criteria

1. Machining of brass and cast iron is possible (in addition to the above steels)

1. Acid number: min 2. Sulphur content: min 3. Chlorine content: min 4. Flash point: min 5. Fineness of treatment of cutting fluid: max 6. Degree of treatment of cutting fluid: max

1. Capital expenditures: min 2. Current expenditures: min

Table 3 Assessment of integrated criterion parameters by user Criteria

Weight of criterion

Brass machining possible Cast iron machining possible Acid number: min Sulphur content: min Chlorine content: min Flash point: min Fineness of treatment of cutting fluid: max Degree of treatment of cutting fluid: max Capital expenditures Current expenditures

10 10 4 7 7 6 5 5 4 8

6. Obtaining an integrated assessment for selection options, in other words, cumulative weighting of options in view of selection criteria by multiplying the selection option weight value under a certain criterion by the criterion weight and summing the obtained values for each selection option. Thus, the acid number parameter is assessed by multiplying the expert assessment of the option for this parameter by the weight (4) set for this parameter by user, and the sulphur content parameter is by multiplying the expert assessment by the weight (7) set by user. 7. Ranking options by weight, selecting an option whose weight is the highest as a proposed solution. The highest integral assessment under the integrated criterion was given to the following option: Cutting fluid: ARIAN MR-11, regeneration system: centralized system with press filter. 8. If a proposed option is ok, the proposed option is adopted; otherwise, a user can start a new search based on a modified integrated criterion.

Development of Integrated Criterion to Select Environmentally …

287

6 Conclusions • The article solves the problem of building an integrated criterion to select environmentally sound cutting fluids and their application systems to ensure the minimum impact of CF on the environment and humans. • The paper presents an algorithm to select cutting fluids and their application systems taking into account process, environmental and economic constraints. In addition, the algorithm takes into account the significance of each selection criterion in view of set constraints. • The paper presents the method to select CFs and their application systems via the selection algorithm. The practical implementation of the methodology has shown that the developed algorithm is a reliable tool for reasonable selecting of cutting fluids under the proposed integrated criterion. • The studies at MTZ TRANSMASH OJSC showed that it is feasible to use ARIAN MR-11 CF with a centralized system with a press filter as a regeneration system for lathe machining of workpieces. • The studies have shown that the integrated criterion to select environmentally sound cutting fluids and their application systems is a flexible tool to assess various options when selecting CF and a relevant system taking into account process, environmental and economic measures by singling out critical parameters in view of their significance.

References 1. Shvartsburg LE, Yagolnitser OV, Butrimova EV (2018) Integrated approach to providing for environmental friendliness and safety of the technological processes. MATEC Web of Conferences 224:01090. https://doi.org/10.1051/matecconf/201822401090 2. Shvartsburg LE, Butrimova EV, Yagolnitser OV (2017) Quantitative evaluation of the effectiveness of best available technologies of form-shaping. MATEC Web of Conferences 129:01027. https://doi.org/10.1051/matecconf/201712901027 3. Gvozdkova SI, Shvartsburg LE (2017) Analysis of sources and methods for reducing noise by minimizing vibrations of engineering technological processes. Proced Eng 206:958–964. https://doi.org/10.1016/j.proeng.2017.10.578 4. Shvartsburg LE, Vikharev AS (2017) Performance management of local air purification systems. Ecol Ind Russ 21(1):4–7. https://doi.org/10.18412/1816-0395-2017-1-4-7 5. Shvartsburg LE, Butrimova EV, Yagolnitser OV (2017) Energy efficiency and ecological safety of shaping technological processes. Proced Eng 206:1009–1014. https://doi.org/10. 1016/j.proeng.2017.10.586 6. Rodriguez PE, Shvartsburg LE, Artemyeva MS (2017) Methodological design and commissioning of an experimental stand for the study of the spread of harmful substances in the air of work areas during the processing of metals in industry. Proced Eng 206:588–593. https://doi.org/10.1016/j.proeng.2017.10.521

288

L. E. Shvartsburg et al.

7. Golubkov YU, Ermolaeva NV, Shvartsburg LE (2016) Nitrogen-bearing organic components of industrial oils. Chem Technol Fuels Oils 52(1):90–94. https://doi.org/10.1007/s10553-0160677-2 8. Zmieva KA, Shvartsburg LE (2009) Automated energy and resource saving systems for industrial enterprises. Ecol Ind Russ 11:7 9. Egorov SB, Kapitanov AV, Mitrofanov VG, Shvartsburg LE, Ivanova NA, Ryabov SA (2016) Formation of the integral ecological quality index of the technological processes in machine building based on their energy efficiency. Int J Environ Sci Educ 11(11):4065–4078 10. Shvartsburg LE (2015) Ecoenergetics of cutting manufacturing processes. Ecol Ind Russ 3 (19):4–9 11. Ivanova NA, Ryabov SA, Shvartsburg LE (2016) The role of information technology in rotor balancing. Russ Eng Res 36(3):235–238. https://doi.org/10.3103/S1068798X16030096 12. Zaborowski T, Shvartsburg LE, Konov SG (2016) Tracking navigation system based on photogrammetry principles. Mater Sci Forum 876:69–73. https://doi.org/10.4028/www. scientific.net/MSF.876.69 13. Egorov SB, Kapitanov AV, Mitrofanov VG, Shvartsburg LE, Ivanova NA, Ryabov SA (2016) Modern digital manufacturing technical support centers. Math Educ 11(7):2213–2225 14. Egorov SB, Kapitanov AV, Mitrofanov VG, Shvartsburg LE, Ivanova NA, Ryabov SA (2016) Formation of the integral ecological quality index of the technological processes in machine building based on their energy efficiency. Int J Environ Sci Educ 11(11):4065–4078 15. Kapitanov AV (2016) Special characteristics of the multi-product manufacturing. Proced Eng 150:832–836. https://doi.org/10.1016/j.proeng.2016.07.127 16. Kapitanov AV, Mitrofanov VG, Omel’chenko IS, Sirotkin OS (2015) The effect of reliability on the distribution of a production program among technological systems at an engineering enterprise. J Mach Manuf Reliab 44(7):646–649. https://doi.org/10.3103/S1052618815 070092 17. Egorov SB, Kapitanov AV, Loktev DA (2016) Modern methods and technological solutions for effective processing of gear wheels. Mater Sci Forum 870:397–403. https://doi.org/10. 4028/www.scientific.net/MSF.870.397 18. Shvartsburg LE, Ivanova NA, Ryabov SA, Zaborowski T (2014) Chemical contaminations in a process of polishing with an implementation of liquid LCTS. Life Sci J 11(10S), 40:228–230 19. Ryabov SA, Ivanova NA, Shvartsburg LE (2014) Assessment, analysis and managing occupational risks in the industry. Chief Mech Eng 12:21–26 20. Hudoshina MY, Butrimova OV (2015) A complex criterion for the evaluation of ecologically substantiated selection of LCAs and systems for their application. Ecol Ind Russ 5(19):46–49 21. Hudoshina MY, Butrimova OV (2011) Research of interrelations of technological and ecological parameters of technological system with application cooling technological means. Life Safety 6(126):27–30 22. Hudoshina MY, Butrimova OV (2014) Development methodology of information retrieval system for environmentally sound selection of lubricant cooling technological fluids in mechanical engineering. Technol Oil Gas 1(90):54–59 23. Hudoshina MY, Butrimova OV (2012) Development of complex criterion of the estimation of variants of ecologically well-founded choice of cooling technological means. Herald Moscow Univ Finan Law MFUA 1:149–154 24. Butrimova EV, Yagolnitser OV, Shvartsburg LE (2017) Improving of environmental friendliness the technological processes of metal working by cutting. Secur Probl Russ Soc 1:108–112

Gear-Grinding Temperature Modeling and Simulation N. V. Lishchenko and V. P. Larshin

Abstract New trends in the manufacture of gears are associated with the use of new gear-grinding technologies. Discontinuous profile gear grinding by a profile wheel, compared to the continuous generating gear grinding by a grinding worm, provides for a higher accuracy (DIN 3-6) but yields less performance through both the higher grinding temperature and possibility of grinding burns. The grinding temperature is one of the factors limiting the performance of the profile gear-grinding operation. There are two most commonly used methods involved for determining the grinding temperature: a fully analytical method based on analytical models and the simulation one based on the similar set of models which are working under computer control with the temperature field monitoring. However, the continuity and interrelation of these methods for determining and studying the profile gear-grinding temperature have not yet been investigated. The relevance of this problem is currently the most pronounced in the connection with the development of appropriate technological preconditions and gear-grinding subsystems for the grinding operation designing, monitoring, and diagnosing which allow adapting the elements of the grinding system to higher productivity. The software for these subsystems can be created on the basis of the analytical mathematical models of the temperature field because simulation modeling takes a lot of time. That is why, in other equal conditions, the comparison of the results of analytical and simulation modeling helps choosing the right way for further improvement of the profile grinding technology on CNC machines. Keywords Profile gear grinding


Grinding temperature
Modeling
Simulation

N. V. Lishchenko (&) Odessa National Academy of Food Technologies, 112, Kanatnaya Str, Odessa 65039, Ukraine e-mail: [email protected] V. P. Larshin Odessa National Polytechnic University, 1, Shevchenko Avenue, Odessa 65044, Ukraine © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_32

289

290

N. V. Lishchenko and V. P. Larshin

1 Introduction The quality of the gear surface layer is formed during the gear-grinding operation. Among the main requirements for the quality of gears is the lack of grinding burns which caused by grinding temperature. The trend toward the development of adaptive grinding system [1, 2] has revived interest in methods for obtaining information on the grinding system state parameters. Some aspects and new approaches of our research were published earlier [3–5] and will be used here in a new modeling situation when the modeling is a necessary condition for subsequent simulation. There are many papers which are devoted to the study of thermal phenomena in grinding [6–21] including ones connected with the profile gear grinding. The solutions presented in these papers are often obtained without the formulation of the problem as well as the initial and boundary conditions. Besides, until now the mechanism of the grinding temperature simulation is not disclosed when new knowledge about the subject under study is being obtained. In the literature, there is no information about the relation between the analytical and simulation modeling the temperature field in the profile gear grinding. The lack of such data leads some scientists to believe that simulation on the basis of models set is not an objective method of science. They oppose to this method a purely analytical solution of the temperature problem by analogy with the solution obtained for non-complex geometrical surfaces to be grinded such as plane and cylinder. They do not understand that the temperature task for an arbitrarily complex surface to be grinded at the stage of model development is an original technique that cannot be replaced by an analytical description of this surface. It will be shown that the simulation is performed on the basis of the basic analytical differential heat conduction equations including those obtained earlier in the classical works of the founders of heat conduction theory [22–24]. The thing is how to ensure the interaction of these analytic equations, each of which corresponds to a limited area of an arbitrary complex surface.

2 Mathematical Models for Simulation According to the research three-component methodology for grinding systems, the modeling is one of the three components to be involved along with optimization and control [25]. The computer simulation mechanism is based on the previously developed database of analytical partial differential equations of heat conductivity which are used simultaneously for different parts of the tooth surface to be grinded. This surface is previously set in the form of some geometric model, e.g., in the AutoCAD medium. In this case, a new knowledge, in contrast to the usual solution of the analytical differential equations, appears due to the computer solution of a system of these equations each of which relates to the corresponding surface section. The number of such discrete sections, e.g., 99 sections, representing the whole

Gear-Grinding Temperature Modeling and Simulation

291

contact zone between the profile grinding wheel and the tooth, is the result of optimizing the time spent on obtaining an acceptable solution with a predetermined accuracy of grinding temperature determining. With an increase in the number of such discrete sections, both the accuracy of the temperature determining (taking into account the influence of a real geometric shape of the surface, e.g., an involute surface), and temperature field simulation time increase simultaneously.

2.1

Geometric 3D Model

The grinding wheel profile has a rectangular contact zone with the height of (ra − rb) and the width of 2 h that equals to the heat source width 2 h. Besides, pffiffiffiffiffiffiffiffiffiffi h ¼ D
tv =2, where tv is the radial depth of grinding; D is the grinding wheel diameter, m. For example, at tv = 0.074 mm and D = 400 mm, we get h = 2.72  10−3 m and 2h = 5.44  10−3 m or 5.44 mm. The partial differential equation of thermal conductivity in the COMSOL Multiphysics window looks like qCp

@T þ qCp u
rT þ r
q ¼ Q; @t

ð1Þ

where q is the material density, kg/м3; Cp the specific heat, J/(kg °C)/(кг); u the velocity vector, m/s; rT the temperature gradient, °C/м; q the heat flux vector, W/m2; Q the power of heat source per unit volume, W/m3. Converting Eq. (1) for Q = 0, we obtain the equation for a moving heat source which is given in [22–24], i.e.  2  @T @ T @2T @2T @T ¼a : þ þ þ uz 2 2 2 @t @x @y @z @z

ð2Þ

Equation (2) is a linear second-order partial differential equation in contrast to an ordinary differential equation. The solution of such an equation can be an arbitrary function of four variables Tðx; y; z; tÞ that satisfies this equation. In order for the solution Tðx; y; z; tÞ to be the only solution to the problem, it must simultaneously satisfy some initial and boundary conditions. The initial condition T ¼ Tðx; y; z; t ¼ 0Þ ¼ Hðx; y; zÞ is determined by the temperature distribution inside the investigated area at an initial time. For t = 0 and an initial temperature field which is uniformly distributed in the volume of a workpiece, the initial condition has the form Tðx; y; z; 0Þ ¼ T0 ¼ const. Let us take T0 = 20 °C. In the COMSOL Multiphysics program, the second-kind boundary condition is given as follows

292

N. V. Lishchenko and V. P. Larshin

nq ¼ q0

ð3Þ

q ¼ krT

ð4Þ

where n is the unit normal vector to the surface which is heating in grinding; q0 the fixed, independent of coordinate and time, component of the heat flux density through the boundary surface, W/m2. Denoting the thermal conductivity k by k, we get from (4) q ¼ krT:

ð5Þ

Taking into account the expression rT ¼ n @T @n , we can write the expression (5) in the form q ¼ kn

@T : @n

ð6Þ

Considering (6), we will write the expression (3) in the form   @T n kn ¼ q0 : @n

ð7Þ

Next, we have 

@T nn k @n

 ¼ q0 :

ð8Þ

Considering that nn ¼ jnj
jnj cos 0 ¼ n2 ¼ 1, we get k

@T ¼ q0 : @n

ð9Þ

Denoting q0 by q, i.e., q0 = q = const, Eq. (9) can be written as @T 1 ¼  q: @n k

ð10Þ

The expression (10) coincides with the corresponding one adopting in [22–24]. It means the transition from the vector form of the second-kind boundary condition, as it is in the COMSOL Multiphysics, to the scalar form is executed correctly. Similar transformations can be performed for boundary condition of the third kind. In the COMSOL Multiphysics medium, the third-kind boundary condition is given by two of the following equations

Gear-Grinding Temperature Modeling and Simulation

293

nq ¼ q0 ;

ð11Þ

q0 ¼ aðTC  TH Þ;

ð12Þ

where a is the convective heat transfer coefficient, W/(m2 °C); TC the external temperature; TH the surface temperature in the grinding contact zone, TC  TH . Taking into account the expressions (5)–(9), we obtain k

@T ¼ aðTH  TC Þ: @n

ð13Þ

Equation (13) coincides with the expression adopting in [22–24]. It means the transition from the vector form of the third-kind boundary condition, as it is in the COMSOL Multiphysics, to the scalar form is executed correctly. Thus, it is shown that the received differential equation of heat conduction (2) corresponds to its canonical form for the three-dimensional formulation of the thermophysical problem in an analytical form [22–24]. The results of the study are shown in Fig. 1 for the following gear data with 20° pressure angle: the number of teeth 40, module 3.75 mm, face width 24 mm. Circle diameters: addendum (outside) 153.750 mm, pitch 150 mm, base 140.954 mm, dedendum (root) 139.875 mm.

2.2

Geometric 2D Model

To study the temperature field for a unmoving heat source, a 2D geometric object of the tooth profile was created in the AutoCAD medium and imported into the

Fig. 1 a Location of the Y2 axis, b the dependence of the T (°C) on the Y2 at different heating time sH in ms

294

N. V. Lishchenko and V. P. Larshin

COMSOL Multiphysics medium (Fig. 2). In this case, the heat conduction Eq. (1) takes on the form qCp

@T þ qCp
0
rT þ r
q ¼ 0: @t

ð14Þ

After transformations of Eq. (14), we obtain  2  @T @ T @2T ¼a
þ 2 : @t @x2 @y

ð15Þ

The heat flux density q = q (rx) in Fig. 1 is obtained on the basis of the dependence obtained in [5] qðrx Þ ¼

P dQw Vf tn ðrx Þ w ¼ ec w pffiffiffiffiffiffiffiffiffiffiffiffiffiffi ; Vf Scc dSc Dtv ðrx Þ

ð16Þ

Fig. 2 a Geometric model of the tooth gear with finite element mesh, b temperature field along the involute profile of the tooth for the 2D geometric model of the tooth, c the temperature field isotherms, d both the temperature T (°C) and the heat flux q on the Y2 at different heating time sH ¼ 2 h=Vf

Gear-Grinding Temperature Modeling and Simulation

295

Table 1 Results of determining the maximum surface temperature TMAX Axial feed Vf (m/min) 1 3 H 3.99 11.97 L 4.67 14.07 0.3264 0.1088 sH (s) Moving source (3D grinding field) 248 448 TMAX (°C) Unmoving source (2D grinding field) 258 457 TMAX (°C) d (%) 4.03 2.00

5

7

9

12

19.93 23.45 0.0653

27.92 32.85 0.0460

35.90 42.22 0.0363

47.86 56.30 0.0272

588

700

796

923

596 1.36

705 0.71

808 1.51

936 1.41

V
l

Note h ¼ 2:72  103 m; L ¼ 2af ; H=L ¼ 0:85\1; 2l ¼ 6:398  103 m; d is the relative difference between 2D and 3D grinding temperature determination results

where rx is the current radius vector of the considered point of the involute profile, m; P the grinding power, W; Vf the axial feed, m/s; Scc the cross-sectional area of the layer to be removed, m2; w the share of heat flux into the workpiece; Qw the material removal rate, m3/s; Sc the contact zone area, m2; ec the specific grinding energy, J/m3; tn ðrx Þ and tv ðrx Þ are the normal and vertical grinding depths, mm; D the instant diameter of a grinding wheel in the considered section of its profile, m. It is necessary to note that the maximum values of the temperature T and heat flux q are located in the upper part of the involute profile and do not coincide along the height of the tooth: the temperature maximum is below the maximum of the heat flux (Fig. 2d). This was later confirmed by the results of the experimental study of grinding burns. The results of determining the maximum grinding temperature based on the temperature field simulation performed for the moving (Fig. 1) and unmoving (Fig. 2) heat sources are summarized in Table 1 in which the geometric and other parameters of the gear grinding are shown as well.

3 Conclusions 1. The temperature field simulation under 2D and 3D modeling made it possible to prove that the difference in the results of determining the surface grinding temperature for these two cases of modeling does not exceed 5% when the heating time sH under the 2D modeling is equal to the ratio of the moving heat source width 2 h to the axial feed Vf of its movement, i.e., sH = 2 h/Vf. Moreover, at Vf  5 m/min, this difference does not exceed 1%. 2. It is shown that in the axial feed range 1  Vf  12 m/min, the difference between the temperature determination results for the 2D and 3D modeling is 0.71–4.03%. This confirms the possibility of replacing a moving heat source with an unmoving one whose heating time sH depends on the axial feed Vf of the moving source.

296

N. V. Lishchenko and V. P. Larshin

3. With an increase in axial feed Vf from 1 m/min to 12 m/min, the maximum temperature is increased for the moving heat source from 248 to 923 °C and for the unmoving one from 258 to 936 °C. Such correspondence in their magnitudes confirms the results’ identity when determining the surface maximum temperature for these solutions obtained in 2D and 3D simulation. In other equal conditions, the maximum temperature for a moving source (248–923 °C) is less than for the unmoving one (258–936 °C) throughout the range of 1  Vf  12 m/min. 4. The investigation of the continuity (similarity) of three-, two-, and onedimensional analytical solutions of the differential equations of heat conduction allowed to establish the conditions for the application of the two- and one-dimensional solutions for determining the gear-grinding temperature, to wit: the dimensionless half-width of the heat source (Peclet number) H  4 and the condition H/L  1 for the rectangular shape of the contact zone which has dimensions 2H  2L. This makes it possible to simplify both the mathematical models and software of embedded computer subsystems for the designing, monitoring, and diagnosing of the grinding operation. The software for these subsystems is created on the basis of the mathematical models of the temperature field with and without taking into account the effect of forced cooling.

References 1. Larshin V, Lishchenko N (2019) Adaptive profile gear grinding boosts productivity of this operation on the CNC machine tools. Lect Notes Mech Eng part F2:79–88. https://doi.org/10. 1007/978-3-319-93587-4_9 2. Larshin V, Lishchenko N (2018) Gear grinding system adapting to higher CNC grinder throughput. MATEC Web of Conferences, vol. 226, 04033, pp 1–6. https://doi.org/10.1051/ matecconf/201822604033 3. Lishchenko N (2018) Profile gear grinding temperature determination. Transactions of Kremenchuk Mykhailo Ostrohradskyi National University, pp 100–108. https://doi.org/10. 30929/1995-0519.2018.1.100-108 4. Larshin VP, Kovalchuk EN, Yakimov AV (1986) Primenenie resheniy teplofizicheskikh zadach k raschetu temperatury i glubiny defektnogo sloya pri shlifovanii (Application of solutions of thermophysical problems to the calculation of the temperature and depth of the defective layer during grinding). In: Interuniversity collection of scientific works, Perm, pp 9–16 5. Larshin VP (1999) Tekhnologiya mnogonitochnogo rezboshlifovaniya pretsizionnykh khodovykh vintov (Multi-thread grinding technology for precision ball-screws). Trudy Odes Politekhn un-ta 2(8):87–91 6. Christof C, Schlattmeier H et al (2006) Optimization of the gear profile grinding process utilizing an analogy process. Gear technology (November/December), pp 34–40 7. Klocke F, Schlattmeier H (2004) Surface damage caused by gear profile grinding and its effects on flank load carrying capacity. Gear technology (September/October), pp 44–53 8. Jermolajev S, Brinksmeier E, Heinzel C (2018) Surface layer modification charts for gear grinding. CIRP Annals—Manuf Technol 1:1–4. https://doi.org/10.1016/j.cirp.2018.04.071

Gear-Grinding Temperature Modeling and Simulation

297

9. Jin Tan, Yi Jun, Peng Siwei (2017) Determination of burn thresholds of precision gears in form grinding based on complex thermal modelling and Barkhausen noise measurements. Int J Adv Manuf Technol 88(1–4):789–800 10. Fergania O, Shaoa Y, Lazoglub I et al (2014) Temperature Effects on Grinding Residual Stress. In: 6th CIRP International Conference on High Performance Cutting, HPC 2014, pp 2–6. https://doi.org/10.1016/j.procir.2014.03.100 11. Deivanathan R, Vijayaraghavan L (2013) Theoretical analysis of thermal profile and heat transfer in grinding. Int J Mech Mater Eng (IJMME) 8(1):21–31 12. Yadav RK (2014) Analysis of grinding process by the use of finite element methods. Elk Asia Pacific J Manuf Sci Eng 1(1) 13. Foeckerer T, Zaeh M, Zhang O (2013) A three-dimensional analytical model to predict the thermo-metallurgical effects within the surface layer during grinding and grind-hardening. Int J Heat Mass Transf 56:223–237. https://doi.org/10.1016/j.ijheatmasstransfer.2012.09.029 14. González-Santander JL (2016) Maximum temperature in dry surface grinding for high Peclet number and arbitrary heat flux profile. Mathematical Problems in Engineering—2016, pp 1–9. https://doi.org/10.1155/2016/8470493 15. Guo C, Malkin S (1995) Analysis of transient temperatures in grinding. J Eng Ind 117:571– 577. https://doi.org/10.1115/1.2803535 16. Heinzel C, Sölter J, Jermolajev S et al (2014) A versatile method to determine thermal limits in grinding. In 2nd CIRP Conference on Surface Integrity (CSI), Procedia CIRP, vol 13, pp 131–136. https://doi.org/10.1016/j.procir.2014.04.023 17. Beizhi L, Dahu Z, Zhenxin Z et al (2011) Research on workpiece surface temperature and surface quality in high-speed cylindrical grinding and its inspiration. Adv Mater Res 325:19– 27. https://doi.org/10.4028/www.scientific.net/AMR.325.19 18. Li Hao N, Axinte D (2017) On a stochastically grain-discretised model for 2D/3D temperature mapping prediction in grinding. Int J Mach Tools Manuf, pp 1–27. https://doi.org/10.1016/j. ijmachtools.2017.01.004 19. Tadeu A, Simoes N (2006) Three-dimensional fundamental solutions for transient heat transfer by conduction in an unbounded medium, half-space, slab and layered media. Eng Anal Bound Elem 30(5):338–349. https://doi.org/10.1016/j.enganabound.2006.01.011 20. Xun Chen, Öpöz T (2016) Effect of different parameters on grinding efficiency and its monitoring by acoustic emission. Prod Manuf Res Open Access J4(1):190–208. https://doi. org/10.1080/21693277.2016.1255159 21. Malkin S, Guo C (2007) Thermal analysis of grinding. Annal CIRP 56:760–782. https://doi. org/10.1016/j.cirp.2007.10.005 22. Jaeger JC (1942) Moving sources of heat and temperature at sliding contact. Proc Roy Soc N S Wales 76:203–224 23. Carslaw HS, Jaeger JC (1959) Conduction of heat in solids, 2nd edn. University Press, Oxford, p 510 24. Sipaylov VA (1978) Teplovye protsessy pri shlifovanii i upravlenie kachestvom poverkhnosti (Thermal processes during grinding and surface quality control). Moskva, Mashinostroenie, p 167 25. Larshin V, Lishchenko N (2018) Research methodology for grinding systems. Russ Eng Res 38(9):712–713. https://doi.org/10.3103/S1068798X18090204

How Parameters of Agricultural Machine and Tractor Unit Affects Effectively Used Mean Indicated Power M. Ya. Durmanov, B. G. Martynov and S. V. Spiridonov

Abstract The paper dwells upon how the parameters of an agricultural machine and tractor unit (MTU) affect the effectively used mean indicated power. The operating parameter in use is the MTU travel speed, while the design parameter used is the suspension stiffness. Transport-mode losses of mean indicated power due to dynamic loads have been determined for the following factors: ascent resistance; MTU inertial forces; MTU suspension vibrations in the longitudinal vertical plane; mechanical losses in the friction pairs; as well as the total loss for various suspension stiffness values. For the plowing mode, power losses are determined for the dynamics of soil removal by the plow. Despite the fact that transport-mode MTU operations are not as long as plowing operations, they are energy-intensive and feature considerable losses of power due to ascent, MTU vibrations in the longitudinal vertical plane, as well as due to the mechanical losses in friction pairs. Total dynamic power losses in transport and plowing modes peak at a frequency of 5.0 s−1, which is the energy-intense eigenfrequency (rotation speed) of the YaMZ-238ND5 engine. Increasing the MTU travel speed from 2.01 to 5.20 m/s and reducing the suspension stiffness causes a “shrinkage” in the area bounded by the total dynamic power-loss curve and coordinate axes. This changes the magnitude and frequency of dynamic mean indicated power components. Using the proposed method when designing enables the engineer to find the necessary and sufficient redundant indicated power to overcome the temporarily rising overloads without switching the gear (the gearbox) to a higher gear ratio.




Keywords Machine and tractor unit Mean indicated power Suspension stiffness Frequency characteristics





Travel speed


M. Ya. Durmanov (&)
B. G. Martynov
S. V. Spiridonov S. M. Kirov Saint Petersburg State Forest Technical University, 5U Institutsky pereulok, St. Petersburg 194021, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_33

299

300

M. Ya. Durmanov et al.

1 Introduction As of today, agricultural machine and unit designers use statistical dynamics methods [1, 2], the scientific foundations of agricultural mechanics, [3] as well as the existing MTU theories [4, 5]. Studies into power transmissions, actuators, and MTU units are limited to researching dynamic loads without covering the energy indicators related to the power losses, fuel and oil losses on burning, the diesel engine service life, and the power transmission elements [6, 7]. Optimization of power transmitted from the diesel engine to the drive wheels enables optimizing the design parameters of the transmission elements, as the transmitted and transformed power determines the contact and bending stresses in the gear transmissions, affecting their service life [8–18]. Using the proposed method when designing enables the engineer to find the necessary and sufficient redundant indicated power to overcome the temporarily rising overloads without switching the gear (the gearbox) to a higher gear ratio. When designing a gearbox, keeping this in mind allows avoiding the “overlaps” in gear ratios of the adjacent gears and driving modes, avoiding “jumps” in the close ratios, tangential traction force and speed, as well as reducing the MTU power consumption [7]. Thus, the goal is to research how the MTU operating and design parameters affect the effectively used mean indicated power. Development and practical application of various methods for estimating the effectively used MTU power is relevant for designing a tractor, where it helps improve the quality of MTU design [8–14].

2 Method MTU operation models for different modes have been implemented in MathCAD, while Microsoft Office Excel has been used to process the calculation results and to plot the graphs. The calculation methodology is described below. Main external influences modeled are: the arable surface profile (configured by the correlation function and spectral density); diversity factors of the physico-mechanical soil properties; and MTU travel speed. The research models the non-steady state of MTU motion resistance, which is described by an ergodic stationary random process for the entire frequency spectrum of the input load moment. The paper describes the transport and plowing operations of a K-744R-05 Kirovets tractor equipped with a PUN-8-40 plow for when employed on light soils.

How Parameters of Agricultural Machine and Tractor Unit …

2.1

301

Analytical Expressions to Find the Effectively Used Mean Indicated Power of MTU Engine in Transport and Plowing Modes

The mean indicated power of the engine consists of the regular component Ni0 spent to maintain the set MTU speed and to overcome the regular resistance forces; and of the variable component Ni ðx; t0 Þ spent due to the dynamic loads caused by the inertial forces, the MTU vibrations in the longitudinal vertical plane, as well as the loads in the power transmission: • for transport mode, Nit
 pn0 R t A1 þ D1 ða1 þ b1 n0 Þ 30 it gm  p R  t Nit ðx; t0 Þ ¼ Mna x
jU11 ðjx; t0 Þj
A þ 2A4 x
jU11 ðjx; t0 Þj 30 it gm 1 sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi "  t  a 4x2 þ c2 1
jG61 ðjx; t0 Þj þ 2frf Hv ðjx; t0 Þ þ D1
Pmn x2 þ c2 Ni0t ¼

ð1Þ

ð2Þ

þ 2b1 x
jU11 ðjx; t0 Þjg; • for plowing mode, Nip
 pn0 R ½Ap1 þ A2 þ A3 n20  þ D1 ða1 þ b1 n0 Þ ; ð3Þ 30 it gm   2  p a R  p p Ni ðx; t0 Þ ¼ Mn x
jU11 ðjx; t0 Þj
A1 þ A2 þ 3A3 U11 ðjx; t0 Þ 30 i t gm    þ 2frf Hpv ðjx; t0 Þ þ 2A4 x
jU11 ðjx; t0 Þj sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi " #) a1 4x2 þ c2 þ D1

jG61 ðjx; t0 Þj þ 2b1 x
jU11 ðjx; t0 Þj ; ð4Þ Pmn x2 þ c2 Ni0p ¼

where n0 ¼ 0:8nn ; nn is the nominal diesel engine crankshaft rotation speed; R is the drive-wheel radius; it is the transmission gear ratio; gm is the mechanical efficiency, gm ¼ gtr
gmov ; gtr , gmov is the transmission efficiency and the propulsion-unit efficiency, respectively; At1 ¼ mgðsin a þ frf cos aÞ; m is the MTU mass, m ¼ mt þ mpl ; mt , mpl is the tractor mass and the plow mass; frf is the rolling friction coefficient; a is the angle of ascent; a1 , b1 are the coefficients derived experimentally for each engine type; Mna is the amplitude of the variable load moment component, Mna ¼ 0:15Men ; x is the load fluctuation frequency; jU11 ðjx; t0 Þj is the transfer function of the diesel engine crankshaft rotation speed in

302

M. Ya. Durmanov et al.

  terms of disturbance; Htv ðjx; t0 Þ is the transfer function of the MTU carcass vibrations in the longitudinal vertical plane when operating in the transport mode; Pmn is the oil pressure in the main oil-distributing passage (MODP) at the nominal rotation speed; c ¼ 103 xn ; xn is the nominal angular rotation speed of the crankshaft: xn ¼ pnn =30;  jG61 ðjx; t0 Þj is the transfer function of oil pressure in the diesel engine MODP; Hpv ðjx; t0 Þ is the transfer function of the MTU carcass vibrations in the longitudinal vertical plane when operating in the plowing mode; fpf is the friction coefficient for furrow-bottom and furrow-wall friction; kf is the specific resistance of soil; a, b is the tillage depth and the plowing width; n is the dynamicity coefficient; Vc is the single-cylinder capacity; ie is the number of engine cylinders; se is the number of strokes per single crankshaft revolution; Ap1 ¼ mg sin a þ frf mt g cos a; A2 ¼ fpf mpl g cos a þ kf ab;

pR 2 pR V c ie A3 ¼ nab; A4 ¼ m; D1 ¼ : 30it 30it pse The transfer function of the MTU power plant rotation speed in terms of disturbance (provided that the position of the controller hðtÞ is fixed; the controller sets the fuel injection value h0 ¼ const) [6, 7]: U11 ðsÞ ¼

ke1 ðT22 s2 þ 2T2 n1 s þ 1Þ ; ðT1 s þ 1ÞðT32 s2 þ 2T3 n2 s þ 1ÞðT4 s þ 1Þ

ð5Þ

where ke1 is the rotation speed transmission ratio; T1 , T2 , T3 , T4 are the constants of time; n1 , n2 are the attenuation coefficients. For a YaMZ-238ND5 engine: ke1 ¼ 1:85; T1 ¼ 0:796 s; T2 ¼ 0:370 s; T3 ¼ 0:199 s; T4 ¼ 0:183 s; n1 ¼ 0:200; n2 ¼ 0:150 [6, 7]. Note that for the MTU and its mass m, the applied-to-crankshaft moment of inertia Ia and the time constant T1 are written as Ia ¼ Ie þ

mR2 p n2 ; T1 ¼ Ia n ; 2 30 Nn it

ð6Þ

where Ie is the engine moment of inertia; Nn is the rated power of the engine. The transfer function of the MTU carcass vibrations in the longitudinal vertical plane when operating in the transport mode while plow-equipped will depend on the Nyquist plot of tire deformations per disturbance unit of track irregularity at s ¼ jx [7]:  t  H ðjx; t0 Þ ¼ z0 ðc1 þ jxb1 Þ
½g1 ðjxÞ þ g2 ðjxÞ; v

ð7Þ

How Parameters of Agricultural Machine and Tractor Unit …

303

where g1 ðjxÞ, g2 ðjxÞ are the Nyquist plot values of front- and rear-tire deformations, respectively, per disturbance unit of track irregularity: g1 ðjxÞ ¼ m1 x2 ðc1  m2 x2 þ jxb1 Þ=DðjxÞ;

ð8Þ

g2 ðjxÞ ¼ m2 x2 ðc1  m1 x2 þ jxb1 Þ=DðjxÞ;

where DðjxÞ ¼ m2 ðjxÞ4 þ mR b1 ðjxÞ3 þ ðc1 mR þ b21 ÞðjxÞ2 þ 2b1 c1 ðjxÞ þ c21 ; z0 is the amplitude of harmonic irregularities; c1 is the front- and rear-axle tire stiffness parameter; b1 is the front- and rear-axle tire dissipation parameter; m1 , m2 is the MTU mass born by the front axle and by the rear axle, respectively; x is the frequency of irregularities, x ¼ 2pt=lip ; t is the MTU travel speed; lip is the length of a terrain irregularity (a bump); m2 ¼ m1 m2  m20 ; mR ¼ m1 þ m2 ; mtt is the mass of the plow-equipped tractor in the transport mode; Is is the moment of inertia of the rigid carcass structure in relation to the center of mass: Is ¼ It þ mt l2t þ Ipl þ mpl l2pl ; It , Ipl are the tractor and plow moments of inertia in relation to the centers of mass; lt , lpl is the longitudinal distance of the tractor Ct and plow Cpl centers of mass from the MTU center of mass Csm , respectively; l is the longitudinal tractor base (the inter-axle space): l ¼ l1 þ l2 . Substitute the expressions (8) in (7) to compute the frequency response of the K-744R-05 tractor tires when operating in the transport mode while equipped with a PUN-8-40 plow; use the following source data [19–22]: z0 ¼ 0:03 m; mtt ¼ mt þ mpl ¼ 13;400 þ 2250 ¼ 15;650 kg; It ¼ 44;388 kg m2 ; Ipl ¼ 24;628 kg m2 ; Is ¼ 116;400 kg m2 ; l1 ¼ 1:0 m; l2 ¼ 2:2 m; l ¼ 3:20 m; lt ¼ 0:73 m; lpl ¼ 4:35 m; m1 ¼ 14;670 kg; m2 ¼ 15;941 kg; m0 ¼ 7480 kg; m2 ¼ pffiffiffiffiffiffiffiffiffiffi 177:9 106 kg2 ; c1 ¼ 1500 kN=m; b1 ¼ 2m c1 m2 ; m ¼ 0:1. The transfer function of the MTU carcass vibrations in the longitudinal vertical plane when operating in the plowing mode s ¼ jx [7]:  p  H ðjx; t0 Þ ¼ z0 k0 mtp x2 v

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi c 2 þ b2 x 2 ðc  mtp x2 Þ2 þ b2 x2

;

ð9Þ

where mtp is the sprung mass of the tractor in the plowing mode; s is the second-axle lag, s ¼ l=t; k0 is the axle factor, which equivalently substitutes the two tractor axles with a single generalized axle: k0 ¼ cos

xs xl ¼ cos : 2 2t

ð10Þ

For an MTU based on a Kirovets K-744R-05 tractor, assume that c ¼ 2c1 ; pffiffiffiffiffiffiffiffiffiffiffiffi mtp ¼ 13;400 kg; k0 ¼ 0:707; b ¼ 2b1 ¼ 4m c
mtp ; m ¼ 0:1.

304

M. Ya. Durmanov et al.

Fig. 1 State surfaces of the mean indicated power FR for an MTU based on a K-744R-05 tractor with a PUN-8-40 plow in the transport mode, as a function of the travel speed: a for a suspension stiffness c1 = 1500 kN/m, b for a suspension stiffness c1 = 900 kN/m

Transfer function of oil pressure in the main oil-distributing passage: G61 ðjx; t0 Þ ¼ km1 s
U11 ðsÞ;

ð11Þ

where km1 is the MODP oil-pressure transmission ratio:  km1 ¼ 1:35.  Substitute the values jU11 ðjx; t0 Þj from (5), Htv ðjx; t0 Þ from (7), and  p  H ðjx; t0 Þ from (9) in (2) and (4) to find the state surfaces of the MTU engine v mean indicated power frequency responses (FR) in the transport mode (Fig. 1) and in the plowing mode (Fig. 4a). In order to find the numerical values of the mean indicated power, substitute the following source data in (1) to (4): Vc ¼ 1:875 l; ie ¼ 8; se ¼ 4; a1 ¼ 0:45; b1 ¼ 0:97  103 ; Pmn ¼ 0:6 MPa; nn ¼ 1900 min1 ; Nen ¼ 220 kW; Ie ¼ 2:45 kg m2 ; Men ¼ 1239 N m; Fca ¼ 0:15 Men ; gm ¼ 0:8; R ¼ 0:8 m; a ¼ 0:15 m; b ¼ 3:2 m [19–22]. For light-soil plowing, assume the following MTU operating conditions: frf ¼ 0:12; fpf ¼ 0:40; kf ¼ 3  104 N m; n ¼ 1500 kg/m3 ; a ¼ 3 . The mean indicated power Nie effectively used to perform the technological operations is found as the difference Ma Nie ¼ Ni0  n
xle

x¼20 Z

  HN =M ðjxÞdx; i n

ð12Þ

x¼0

mean indicated power; xle is the lowest where Ni0 ¼ 0:80 Nin , Nin is the rated   eigenfrequency of MTU vibrations; HNi =Mn ðjxÞ is the transfer function of power in terms of load [7]:

How Parameters of Agricultural Machine and Tractor Unit …

HNi =Mn ðsÞ ¼

2.2

p U11 ðsÞ
fs
½Mn0 þ 2Ia
s
U11 ðsÞ 30
 a1 2s  c G61 ðsÞ þ 2b1 s
U11 ðsÞ :
þ D1 Pmn s þ c

305

ð13Þ

Analyzing the Effectively Used Mean Indicated Power of MTU Engine in the Transport Mode

The transport-mode mathematical model is a function that, like the tangent-force frequency response, has six extreme surface states with the eigenfrequencies changing as a function of the travel speed t0 and the frequency x of load fluctuations on the drive wheels and in the power transmission, see Fig. 1. The advantages of such three-dimensional representation of the mean indicated power frequency response are that it covers two states at the same time: • The regular component Ni0 spent to realize the regular components of the tangent force Fk0 and travel speed t0 of the MTU; • The variable component Ni ðx; t0 Þ spent to realize the dynamic components of load and rotation speed for the drive wheels and the power transmission of the MTU. This means that by sectioning the longitudinal vertical plane (along the Y-axis, see Fig. 1) at any fixed regular MTU travel speed t0 and for the entire drive-wheel load fluctuation frequency range x, one can find: • The regular (mean) indicated power necessary to realize the regular components of the tangent force Fk0 and the travel speed t0 ; • The losses in the mean indicated power Ni ðxÞ due to dynamic loads; • The total spent mean indicated power NiR ; • The necessary and sufficient redundant indicated power to overcome the temporarily rising overloads without switching the gear (the gearbox) to a higher gear ratio; • The MTU performance in the plowing mode. Per the components of the formula (2), power consumption peaks when ascending: 16.22 kW at xe ¼ 5:0 s1 , see Fig. 2, curves 1. Relatively high is the power consumption due to MTU vibrations in the longitudinal vertical plane (curves 3), mechanical losses in the friction pairs of the engine at the eigenfrequency xe ¼ 5:0 s1 (curves 4). Reducing the suspension stiffness from 1500 to 900 kN/m increases the power spent in the lower frequency range. This is due to reducing the eigenfrequency of MTU vibrations in the longitudinal vertical plane, which approaches the energy-intense eigenfrequency of engine (drive) rotation speed fluctuations xe ¼ 5:0 s1 , thus contributing to increasing the power spent.

306

M. Ya. Durmanov et al.

Fig. 2 Component breakdown of the power-loss FR in the transport mode for an MTU based on a K-744R-05 tractor with a PUN-8-40 plow for the travel speed t0 ¼ 2:01 m=s: a for a suspension stiffness c1 = 1500 kN/m, b for a suspension stiffness c1 = 900 kN/m, 1 for the ascent resistance, 2 for the inertial forces, 3 for the MTU vibrations in the longitudinal vertical plane, 4 for the mechanical losses in the engine friction pairs, 5 for the total dynamic component

The expression (2) has been used to compute the mean indicated power losses due to dynamic loads for each component in the frequency spectrum of 0–20 s−1, as well as the total losses for various values c1 = 1500; 1200; 900 kN/m: 1. for the ascent resistance Ni1t ðxÞ

pR
Mna t ¼ A
30it gm xle 1

x¼20 Z

x
jU11 ðjxÞjdx

ð14Þ

 2  x2
U11 ðjxÞdx

ð15Þ

x¼0

2. for the inertial forces of the MTU Ni2t ðxÞ

2pR
Mna ¼ A4
30it gm xle

x¼20 Z

x¼0

3. for the MTU suspension vibrations in the longitudinal vertical plane Ni3t ðxÞ

2pRfrf Mna ¼
30it gm xle

x¼20 Z

  x
jU11 ðjxÞj
Htv ðjxÞdx

ð16Þ

x¼0

4. for the mechanical losses in the engine friction pairs Ni4t ðxÞ

pD1 Mna ¼
30xle

x¼20 Z

x¼0

" a1 x
jU11 ðjxÞj
Pmn

þ 2b1 x
jU11 ðjxÞjdx

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 4x2 þ c2
jG61 ðjxÞj x 2 þ c2 ð17Þ

How Parameters of Agricultural Machine and Tractor Unit …

307

Table 1 Calculated mean indicated power spent by an MTU based on a K-744R-05 tractor with a PUN-8-40 plow in the transport mode Travel speed t0 , m/s, (transmission gear ratio)

Components of the dynamic component Niit ðxÞ, kW at stiffness c1 = 1500 kN/m 1 2 3 4

Dynamic components Nit ðxÞ, kW as a function of stiffness c1 , kN/m 1500 1200 900

17.86 0.55 19.09 2.65 40.15 38.00 2.01 (it ¼ 63; 4) 14.90 0.21 23.65 2.19 40.95 37.81 5.20 (it ¼ 24; 5) 66.39/171.76* Regular component of the mean indicated power Ni0t , kW 2.01 Effectively used power 26.24 28.39 Niet ¼ Ni0t  Nit ðxÞ 5.20 130.81 133.95 *In the numerator at the speed t0 ¼ 2:01 m=s; in the denominator at t0 ¼ 5:20 m=s

35.32 34.39

31.07 137.37

total power spent due to dynamic loads as a function of the stiffness c1 Nit ðxÞ

Ma ¼ n
xle

x¼20 Z

   t  HNi =Fc ðjxÞdx

ð18Þ

x¼0

where HNt i =Fc ðjxÞ is the transfer function of the power spent by the MTU in the transport mode [7]. According to (1), the regular component of the mean indicated power spent Ni0t equals 66.39 kW. Calculation results are summarized in Table 1. Apparently, increasing the travel speed from 2.01 to 5.20 m/s while reducing the MTU suspension stiffness by 40% results in “shrinking” the area bounded by Curve 5 (Figs. 2 and 3) and the coordinate axes. This reduces the dynamic components, see Table 1: • Losses due to resistance at an ascent angle a = 3° are reduced by 2.96 kW (16.5%); • Losses due to the inertial forces of the MTU are reduced by 0.34 kW (62.0%); • Mechanical losses in the engine friction pairs are reduced by 0.46 kW (17.4%). Meanwhile, losses due to MTU vibrations in the longitudinal vertical plane increase from 19.09 to 23.65 kW, i.e., by 4.56 kW or 23.9%.

2.3

Analyzing the Effectively Used Mean Indicated Power of MTU Engine in the Plowing Mode

The plowing mode mathematical model is a function that has four extreme surface states (see Fig. 4a) with the eigenfrequencies changing as a function of the travel

308

M. Ya. Durmanov et al.

Fig. 3 Component breakdown of the power-loss FR in the transport mode for an MTU based on a K-744R-05 tractor with a PUN-8-40 plow for the travel speed t0 ¼ 5:2 m=s: a for a suspension stiffness c1 = 1500 kN/m; b for a suspension stiffness c1 = 900 kN/m; for the legend, see Fig. 2

Fig. 4 a mean indicated power frequency response state surface b power-loss frequency response by components when plowing light soils by an MTU based on a K-744R-05 tractor with a PUN-8-40 plow, the suspension stiffness c1 = 1500 kN/m, the travel speed t0 ¼ 2:01 m=s: 1 for the ascent resistance to motion, 2 for the soil layer trimming and removal by the plow, 3 for the MTU vibrations in the longitudinal vertical plane, 4 for the inertial force, 5 for the mechanical losses in the engine friction pairs, 6 for the total losses of power due to dynamic loads

speed t0 and the frequency x of load fluctuations on the drive wheels and in the power transmission. Sectioning the longitudinal vertical plane (along the Y-axis) at the fixed regular MTU travel speed t0 ¼ 2:01 m=s, which corresponds to the gear ratio it ¼ 63:4 (second mode, second gear) [7], one can analyze how the dynamic component of the mean indicated power Niip ðxÞ and its components behave in the entire range of load fluctuation frequencies x, see Fig. 4b. The formula (4) has been used to find the power losses due to dynamic loads for each component in a frequency spectrum 0–20 s−1, as well as the total power loss for various suspension stiffness values c1 = 1500; 1200; 900 kN/m:

How Parameters of Agricultural Machine and Tractor Unit …

309

1. for the ascent resistance Ni1p ðxÞ

pR
Mna ¼
30it gm xle

x¼20 Z

 2  x
jU11 ðjxÞj
Ap1 þ A2 þ 3A3
U11 ðjxÞ dx

x¼0

ð19Þ 2. for the plow resistance dynamics (soil removal) Ni2p ðxÞ

pR
Mna ¼
30it gm xle

x¼20 Z

 2  x
jU11 ðjxÞj
A2 þ 3A3
U11 ðjxÞ dx

ð20Þ

x¼0

3. for the MTU suspension vibrations in the longitudinal vertical plane Ni3p ðxÞ

2pRfrf Mna ¼
30it gm xle

x¼20 Z

  x
jU11 ðjxÞj
Hpv ðjxÞdx

ð21Þ

x¼0

4. for the MTU inertial forces Ni4p ðxÞ

2pR
Mna ¼ A4
30it gm xle

x¼20 Z

 2  x2
U11 ðjxÞdx

ð22Þ

x¼0

5 for the mechanical losses in the engine friction pairs Ni5p ðxÞ

x¼20 Z pD1 Mna ¼
x
jU11 ðjxÞj 30xle sx¼0 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi " # a1 4x2 þ c2

jG61 ðjxÞj þ 2b1 x
jU11 ðjxÞj dx Pmn x2 þ c2

ð23Þ

for the total power losses due to dynamic loads as a function of stiffness c1 Nip ðxÞ

Ma ¼ n
xle

x¼20 Z

x¼0

   p  HNi =Fc ðjxÞdx

ð24Þ

310

M. Ya. Durmanov et al.

Table 2 Calculated mean indicated power spent by an MTU based on a K-744R-05 tractor with a PUN-8-40 plow when plowing light soils at t0 ¼ 2:01 m=s YaMZ-238ND5 Engine configuration Standard

Components of the dynamic component Niip ðxÞ, kW at stiffness c1 ¼ 1500 kN=m 1

2

3

4

5

16.10

15.71

11.00

0.55

2.65

Regular component of the mean indicated power Ni0p , kW Standard Effectively used power Niep ¼ Ni0p  Nip ðxÞ

Dynamic components Nip ðxÞ, kW as a function of stiffness c1 , kN/m 1500 1200 900 46.01 126.38

44.78

43.19

80.37

81.60

83.19

where HNp i =Fc ðjxÞ is the transfer function of the power spent by the MTU in the plowing mode [7]: HNp i =Fc ðsÞ

  p R
s p 2
U11 ðsÞ ¼ A þ A2 þ 3A3
U11 ðsÞ þ 2A4 s
U11 ðsÞ 30 it gm 1
 a1 2s  c p þ 2frf
Hv ðsÞ þ D1
G61 ðsÞ þ 2b1 s
U11 ðsÞ :
Pmn s þ c

ð25Þ

The results of calculating the integrals (19) to (24) when operating the MTU in the plowing mode are summarized in Table 2. Experimental studies [6, 7] have proven that installing the corrective devices in the high-pressure pump and in the MODP of the engine reduces the extreme maxima of the mean indicated power spent as compared to the standard MTU configuration. Moreover, such extreme maxima are reduced most efficiently in the lower load fluctuation frequency range.

3 Results and Discussion Despite the fact that transport-mode MTU operations are not as long as plowing operations, they are energy-intensive and feature considerable losses of power due to ascent, MTU vibrations in the longitudinal vertical plane, as well as due to the mechanical losses in friction pairs. Total dynamic power losses in transport and plowing modes peak at a frequency of 5.0 s−1, which is the energy-intense eigenfrequency (rotation speed) of the YaMZ-238ND5 engine. Using the proposed method when designing enables the engineer to find the necessary and sufficient redundant indicated power to overcome the temporarily rising overloads without switching the gear (the gearbox) to a higher gear ratio. When designing a gearbox, keeping this in mind allows avoiding the “overlaps” in gear ratios of the adjacent gears and driving modes, avoiding “jumps” in the close ratios, tangential traction force and speed, as well as reducing the MTU power consumption.

How Parameters of Agricultural Machine and Tractor Unit …

311

Acknowledgements The authors would like to thank Vladimir Ivanovich Varavva, a Full Professor, and Valery Petrovich Antipin, an Associate Professor, for their invaluable assistance in writing the manuscript.

References 1. Lourier AB (1970) Statisticheskaya dinamika selskohozyaystvennykh agregatov (Statistical dynamics of agricultural aggregates). Kolos, Leningrad, p 376 2. Lourier AB (1967) Avtomatizatsyya selskohozyaystvennykh agregatov (Automation of agricultural aggregates). Kolos, Leningrad, p 264 3. Goryachkin VP (1968) Sobraniye sochineniy (Collection of works), 2nd edn, vol 1, Kolos, Moscow, 720 p 4. Barskiy IB, Anilovich VYa, Kutkov GM (1973) Dinamika traktora (Tractor dynamics). Mashinostroyeniye, Moscow, 280 p 5. Kutkov GM (2004) Traktory i avtomobili. Teoriya i tekhnologicheskie svoystva (Tractors and cars. Theory and technological properties). Kolos, Moscow, 504 p 6. Antipin VP (2012) Energozatraty mashinno-traktornogo agregata (Energy consumption of an agricultural aggregate). Polytech Publishing House, St. Petersburg, p 324 7. Antipin VP, Durmanov MYa, Karshev GV (2017) Proizvoditelnost, energozatraty i resurs mashinno-traktornogo agregata (Performance, energy consumption, and service life of an agricultural aggregate). Polytech Publishing House, St. Petersburg, 484 p 8. Lvovsky KYa, Cherpak FA, Serebryakov IN, Shcheltsyn NA (1976) Transmissii traktorov (Tractor transmissions). Mashinostroyeniye, Moscow, 280 p 9. Grishkevich AI (1984) Proektirovanie transmissiy avtomobiley: Spravochnik (Vehicle transmission design: Handbook). Mashinostroyeniye, Moscow, p 272 10. Sharipov VM (2002) Proektirovanie mekhanicheskikh, gidromekhanicheskikh I gidroobyomnykh peredach traktorov (Designing mechanical, hydro-mechanical, and hydro-volumetric tractor transmissions). MSTU MAMI, Moscow, p 300 11. Vaughan ND, Simner D (2002) Automotive transmissions and drivelines. Butterworth-Heinemann, Oxford 12. Gott PG (1991) Changing gears: the development of the automotive transmission. Society of Automotive Engineers, Warrendale 13. Lechner G, Naunheimer H (1999) Automotive transmissions: fundamentals, selection, design, and application. Springer, Berlin, Heidelberg 14. Birch T, Rockwood C (2001) Automatic transmissions and transaxles. Prentice Hall, Englewood-Cliffs 15. Dunaev PF, Lelikov OP (2009) Konstruirovaniye uzlov i detaley mashin (Design of units and components of machines), 12th edn. Academy, Moscow, p 496 16. Dunaev PF, Lelikov OP (2004) Detali mashin (Components of machines), 5th edn. Mashinostroenie, Moscow, p 560 17. Drago RJ (1988) Fundamentals of gear design. Butterworth, Boston 18. Dudley DW (1994) Handbook of practical gear design. CRC, Boca Raton 19. Pantyukhin MG, Bezverkhiy LI, Berezin NA et al (1982) Spravochnik po traktoram “Kirovets” (Handbook on Kirovets Tractors). Kolos, Moscow, p 271 20. Traktory “Kirovets” K-744R1, K-744R2, K-744R3, K-744R4. Instruktsiya po ekspluatatsyi (Kirovets K-744R1, K-744R2, K-744R3, K-744R4 tractors. User’s Manual). AO Petersburg Tractor Plant, St. Petersburg, 2016, 211 p

312

M. Ya. Durmanov et al.

21. Listopad GYe, Demidov GK, Zonov BD et al (1986) Selskohozyaystvenniye i meliorativniye mashiny (Agricultural and land-reclamation machines). Agropromizdat. Moscow, 688 p 22. Vinokurov VN, Dyomkin VYe, Markin VG et al (2000) Mashiny, mekhanizmy i oborudovaniye selskogo hozyaystva (Forestry machines, mechanisms, and equipment: a reference book). In: Shatalov VG (ed.). MSUF, Moscow, 439 p

Method to Reduce Oil Burning in Diesel Engine of Agricultural Machine and Tractor Unit V. P. Antipin, M. Ya. Durmanov and O. A. Mikhailov

Abstract Oil consumption through burning is not only an important operating indicator of an agricultural machine and tractor unit (MTU); it is the descriptor of how well-designed a diesel engine is. When Volkswagen engines were found to have excessive fuel and oil consumption through burning, the company had to face penalties and wide negative media coverage. Designing the elements and systems of diesel engines implies studying how the parameters of the cylinder-piston group (CPG), the lubrication systems, the fuel injection system, and speed controllers affect such indicators as the indicated torque, power, specific fuel consumption, and last but not least, the moment of mechanical losses in the engine friction pairs. The research so far allows hypothesizing that the nature and magnitude of changes in the crankshaft rotation speed, oil pressure in the main oil-distributing passage (MODP), and the moment of mechanical losses cause increased burning of oil in a wide range of speeds and loads characteristics of MTU operation. The hypothesis can be confirmed by means of an adjuster installed in a diesel-engine MODP, which reduces the oil consumption through burning by lowering the amplitude of fluctuations in the mechanical-loss moment as well as by reducing the friction forces as the source of self-oscillations. The paper describes the transport and plowing operations of a K-744R-05 Kirovets tractor equipped with a PUN-8-40 plow for the use on light soils. For the two operation types, the researchers have plotted the oil-burning frequency response (FR) state surfaces in two settings: a YaMZ-238ND5 diesel engine, standard configuration, and the same engine, adjuster-equipped.







Keywords Machine and tractor unit Diesel engine Oil consumption through burning Mechanical-loss moment Adjuster Frequency response










V. P. Antipin
M. Ya. Durmanov (&)
O. A. Mikhailov S. M. Kirov Saint Petersburg State Forest Technical University, 5U Institutsky pereulok, St. Petersburg 194021, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_34

313

314

V. P. Antipin et al.

1 Introduction Analysis of causes [1–3] behind oil consumption through burning has identified the following causes: beside the rotation speed, the oil temperature and viscosity, the gaps in the shaft-to-liner, sleeve-to-piston-to-compression ring, etc., oil burning is significantly affected by the mean indicated pressure, inertial forces, and the oil pressure in the MODP [4]. The equation for calculating oil consumption through burning in the context of these factors holds for steady state [5]. However, in a real-world setting, tractor diesel engines have to operate in non-steady states for a long time. Studies [6, 7] into the effects of dynamic loads show that there is a range of crankshaft load fluctuation frequencies, where oil consumption through burning is increased. For a YaMZ-238ND5, increased oil burning is observed at 0.5–1.2 Hz, peaking at 0.8 Hz (5.0 s−1). Increased fuel and oil consumption through burning in non-steady state operation reduces the service life of the diesel-engine CPG [5, 8–14]. Studies [6, 7] have identified how the crankshaft rotation speed, the MODP oil pressure, and the mechanical-loss moment affect the oil consumption through burning, allowing to hypothesize that changes in the mechanical-loss moment are self-oscillatory, while the component to describe this process is naturally differentiating. This means that a non-significant increase in the load moment, the crankshaft rotation speed, or the MODP oil pressure will cause a sharp increase in the mechanical-loss moment, a fact evidenced by the experimental frequency responses and the phase responses of a YaMZ-238ND5 diesel engine [6, 7]. The phase lag (or advance angle) and the amplitude of the mechanical-loss moment also depend on the crankshaft rotation speed and MODP oil pressure Nyquist plot. Two hydropneumatic accumulators (bladder accumulators, BA) and a controllable throttle installed consecutively [15] in the MODP do improve the lubrication parameters by altering the phase lag of oil-pressure incrementation from −180° to −45°, thus lowering the amplitude of fluctuations in the mechanical-loss moment as well as reducing the friction forces as a source of self-oscillations [16, 17].

2 Method MTU operation models for different modes have been implemented in MathCAD, while Microsoft Office Excel has been used to process the calculation results and to plot the graphs. The calculation methodology is described below. Main external influences modeled are: the arable surface profile (configured by the correlation function and spectral density); diversity factors of the physico-mechanical soil properties; and MTU travel speed. The research models the non-steady state of MTU motion resistance, which is described by an ergodic stationary random process for the entire frequency spectrum of the input load

Method to Reduce Oil Burning in Diesel Engine …

315

moment. The paper describes the transport and plowing operations of a K-744R-05 Kirovets tractor equipped with a PUN-8-40 plow for when employed on light soils.

2.1

Analytical Expressions to Find Oil Consumption Through Burning in Transport and Plowing MTU Operations

It has been found out [3, 18, 19] that the oil consumption through burning is proportional to the hourly fuel consumption and is the sum of the regular component Ccir0 (1) and the variable component Ccir ðsÞ (2) that are consumed when overcoming the regular and the variable MTU movement resistance forces [6, 7]: 
 ks2 R dnðtÞ þ frf Fvp ðtÞ Ccir0 ¼ ks1 Pm þ Ap1 þ A2 þ A3 n2 þ A4 dt Min it gm    a1 Pm þ D1 þ b1 n  Min Pmn ect  ks2 R
s p 2 Ccir ðsÞ ¼ ks1 s
G61 ðsÞ þ ½2A3
U11 ðsÞ þ A4
U11 ðsÞ Min it gm
 a1 s  c þ frf
Hpv ðsÞ þ D1

G61 ðsÞ þ b1 s
U11 ðsÞ :
Pmn s þ c

ð1Þ

ð2Þ

The expression (2) takes into account the effects of inertial forces, mechanical losses, the amplitude of oil pressure in the MODP, and its phase-frequency response. The operator s being present in the transfer function of oil pressure and inertial-force moment indicates there occurs a sharp increase in the oil consumption through burning when altering the rotation speed and the MODP oil pressure. Thus, the MTU diesel-engine oil burning can be described by the following equations: t • for transport mode Ccir


 ks2 R t A1 þ D1 ða1 þ b1 n0 Þ  Min Min it gm 
ks2 Rx t Ccir ðx; t0 Þ ¼ ks1 x
jG61 ðjx; t0 Þj þ ðA4 x
jU11 ðjx; t0 Þj Min it gm t Ccir0 ¼ ks1 Pm0 þ

þ frf jHtv ðjx; t0 ÞjÞ   a1
jG61 ðjx; t0 Þj þ b1 x
jU11 ðjx; t0 Þj þ D1
Mna ; Pmn

ð3Þ

ð4Þ

316

V. P. Antipin et al.

p • for plowing mode, Ccir

  ks2 R p 2 ¼ ks1 Pm0 þ ½A þ A2 þ A3 n0  þ D1 ða1 þ b1 n0 Þ  Min Min it gm 1 
 2  ks2 R p Ccir ðx; t0 Þ ¼ ks1 x
jG61 ðjx; t0 Þj þ ð2A3 x
U11 ðjx; t0 Þ Min it gm p Ccir0

þ A4 x2
jU11 ðjx; t0 Þj þ frf xjHpv ðjx; t0 ÞjÞ   a1
jG61 ðjx; t0 Þj þ b1 x
jU11 ðjx; t0 Þj þ D1
Mna ; Pmn

ð5Þ

ð6Þ

where ks1 , ks2 are the constant coefficients based on the experimental studies and describing the design parameters of the CPG and those of the fuel injection system; Min is the nominal indicated torque; R is the drive-wheel radius; it is the transmission gear ratio; gm is the mechanical efficiency, gm ¼ gtr
gmov ; gtr , gmov is the transmission efficiency and the propulsion-unit efficiency, respectively; At1 ¼ mgðsin a þ frf cos aÞ; m is the MTU mass, m ¼ mt þ mpl ; mt , mpl is the tractor mass and the plow mass; frf is the rolling friction coefficient; a is the angle of ascent; a1 , b1 are the coefficients derived experimentally for each engine type; n0 ¼ 0:8nn ; nn is the nominal crankshaft rotation speed; Mna is the amplitude of the variable load-moment component, Mna ¼ 0:15Men ; x is the load fluctuation frequency; jU11 ðjx; t0 Þj is the transfer  function of the diesel-engine crankshaft rotation speed in terms of disturbance; Htv ðjx; t0 Þ is the transfer function of the MTU carcass vibrations in the longitudinal vertical plane when operating in the transport mode; Pmn is the oil pressure in the main oil-distributing passage (MODP) at the nominal rotation speed; c ¼ 103 xn ; xn is the nominal angular rotation speed of the crankshaft: xn ¼ pnn =30;  pjG61 ðjx;t0 Þj is the transfer function of oil pressure in the diesel-engine MODP; Hv ðjx; t0 Þ is the transfer function of the MTU carcass vibrations in the longitudinal vertical plane when operating in the plowing mode; fpf is the friction coefficient for furrow-bottom and furrow-wall friction; kf is the specific resistance of soil; a, b is the tillage depth and the plowing width; n is the dynamicity coefficient; Vc is the single-cylinder capacity; ie is the number of engine cylinders; se is the number of strokes per single crankshaft revolution; 2 pR Ap1 ¼ mg sin a þ frf mt g cos a; A2 ¼ fpf mpl g cos a þ kf ab; A3 ¼ 30i nab; t pR m; D1 ¼ Vpsc iee . A4 ¼ 30i t The transfer function of the MTU power plant rotation speed in terms of disturbance (provided that the position of the controller hðtÞ is fixed; the controller sets the fuel injection value h0 ¼ const) [6, 7]:

Method to Reduce Oil Burning in Diesel Engine …

U11 ðjx; t0 Þ ¼

ke1 ðT22 s2 þ 2T2 n1 s þ 1Þ ðT1 s þ 1ÞðT32 s2 þ 2T3 n2 s þ 1ÞðT4 s þ 1Þ

317

ð7Þ

where ke1 is the rotation speed transmission ratio; T1 , T2 , T3 , T4 are the constants of time; n1 , n2 are the attenuation coefficients. For a YaMZ-238ND5 engine: ke1 = 1.85; T1 = 0.796 s; T2 = 0.370 s; T3 = 0.199 s; T4 = 0.183 s; n1 = 0.200; n2 = 0.150 [6, 7]. Note that for the MTU and its mass m, the applied-to-crankshaft moment of inertia Ia and the time constant T1 are written as Ia ¼ Ie þ

mR2 p n2n Ia ; T ¼ 1 30 Nn i2t

ð8Þ

where Ie is the engine moment of inertia; Nn is the rated power of the engine. The transfer function of the MTU carcass vibrations in the longitudinal vertical plane when operating in the transport mode while plow-equipped will depend on the Nyquist plot of tire deformations per disturbance unit of track irregularity at s ¼ jx [7]:  t  H ðjx; t0 Þ ¼ z0 ðc1 þ jxb1 Þ
½g1 ðjxÞ þ g2 ðjxÞ v

ð9Þ

where g1 ðjxÞ, g2 ðjxÞ are the Nyquist-plot values of front- and rear-tire deformations, respectively, per disturbance unit of track irregularity:  g1 ðjxÞ ¼ m1 x2 ðc1  m2 x2 þ jxb1 Þ DðjxÞ;  g2 ðjxÞ ¼ m2 x2 ðc1  m1 x2 þ jxb1 Þ DðjxÞ;

ð10Þ

where DðjxÞ ¼ m2 ðjxÞ4 þ mR b1 ðjxÞ3 þ ðc1 mR þ b21 ÞðjxÞ2 þ 2b1 c1 ðjxÞ þ c21 ; z0 is the amplitude of harmonic irregularities; c1 is the front- and rear-axle tire stiffness parameter; b1 is the front- and rear-axle tire dissipation parameter; m1 , m2 is the MTU mass born by the front axleand by the rear axle, respectively; x is the frequency of irregularities, x ¼ 2pt lip ; t is the MTU travel speed; lip is the length of a terrain irregularity (a bump); m2 ¼ m1 m2  m20 ; mR ¼ m1 þ m2 ; mtt is the mass of the plow-equipped tractor in the transport mode; Is is the moment of inertia of the rigid carcass structure in relation to the center of mass: Is ¼ It þ mt l2t þ Ipl þ mpl l2pl ; It , Ipl are the tractor and plow moments of inertia in relation to the centers of mass; lt , lpl is the longitudinal distance of the tractor Ct and plow Cpl centers of mass from the MTU center of mass Csm , respectively; l is the longitudinal tractor base (the inter-axle space): l ¼ l1 þ l2 .

318

V. P. Antipin et al.

Substitute the expressions (10) in (9) to compute the frequency response of the K-744R-05 tractor tires when operating in the transport mode while equipped with a PUN-8-40 plow; use the following source data: z0 = 0.03 m; mtt ¼ mt þ mpl = 13,400 + 2250 = 15,650 kg; It = 44,388 kg m2; Ipl = 24,628 kg m2; Is = 116,400 kg m2; l1 = 1.0 m; l2 = 2.2 m; l = 3.20 m; lt = 0.73 m; lpl = 4.35 m; m1 = 14670 kg; m2 = 15941 kg; m0 = 7480 kg; m2 = 177.9 pffiffiffiffiffiffiffiffiffiffi 106 kg2; c1 = 1500 kN/m; b1 ¼ 2m c1 m2 ; m = 0.1. The transfer function of the MTU carcass vibrations in the longitudinal vertical plane when operating in the plowing mode s ¼ jx [7]:  p  H ðjx; t0 Þ ¼ z0 k0 mtp x2 v

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi c2 þ b 2 x 2 ðc  mtp x2 Þ2 þ b2 x2

ð11Þ

where mtp is the sprung mass of the tractor in the plowing mode; s is the second-axle lag, s ¼ l=t; k0 is the axle factor, which equivalently substitutes the two tractor axles with a single generalized axle: k0 ¼ cos

xs xl ¼ cos 2 2t

ð12Þ

For an MTU based on a Kirovets K-744R-05 tractor, assume that c ¼ 2c1 ; pffiffiffiffiffiffiffiffiffiffiffiffi mtp = 13,400 kg; k0 = 0.707; b ¼ 2b1 ¼ 4m c
mtp ; m = 0.1. Transfer function of oil pressure in the main oil-distributing passage by disturbance: G61 ðjx; t0 Þ ¼

km1 ðT1 s þ 1ÞðT32 s2 þ 2T3 n2 s þ 1Þ

ð13Þ

where km1 is the MODP oil-pressure transmission ratio. In the context of consecutive installation of two BA and a throttle in the MODP, the oil-pressure transfer function (13) is written as [6, 7] G61 ðjx; t0 Þ ¼

km1 ðT2 s þ 1ÞðT5 s þ 1Þ ðT1 s þ 1ÞðT32 s2 þ 2T3 n2 s þ 1Þ

ð14Þ

For a YaMZ-238ND5 diesel engine: km1 = 1.35 MPa/N m; T2 = 0.37 s; T3 = 0.199 s; T5 = 0.131 s; n2 = 0.2.     Substitute the values jU11 ðjx; t0 Þj from (7), Htv ðjx; t0 Þ from (9), Hpv ðjx; t0 Þ from (11), and jG61 ðjx; t0 Þj form (13) and (14) in (4) and (6) to find the state surfaces of frequency responses (FR) of MTU diesel-engine oil consumption through burning in transport (Fig. 1) and plowing (Fig. 3) operations. The FR section method allows finding the oil consumption through burning by components (Figs. 2 and 4). Numerical values of oil consumption through burning can be found by substituting the following source data in (3)–(6): ks1 = 0.09 kg/MPa h;

Method to Reduce Oil Burning in Diesel Engine …

319

Fig. 1 Oil-burning state surfaces for an MTU based on a K-744R-05 tractor with a PUN-8-40 plow in the transport mode as a function of travel speed at a suspension stiffness c1 ¼ 1500 kN=m: a for standard configuration, b for an adjuster-equipped diesel engine

Fig. 2 Oil-burning FR for an MTU based on a K-744R-05 tractor with a PUN-8-40 plow: breakdown by components in the transport mode at a suspension stiffness c1 ¼ 1500 kN=m and a travel speed t0 ¼ 2:01m=s: a for standard configuration, b for an adjuster-equipped diesel engine, 1 as a function of amplitude changes and phase lag of diesel-engine MODP oil-pressure incrementation, 3 as a function of MTU suspension vibrations in the longitudinal-vertical plane, 4 as a function of diesel-engine mechanical losses, 5 for the total dynamic component of oil consumption

ks2 = 0.45 kg/h; Vc = 1.875 l; ie = 8; se = 4; a1 = 0.45; b1 = 0.97
10−3; Pmn = 0.6 MPa; Pm0 ¼ 0:9Pmn = 0.54 MPa; nn = 1900 min−1; Nen = 220 kW; Ie = 2.45 kg m2; Men = 1239 N m; Fca = 0.15 Men ; it = 63.4; gm = 0.8; R = 0.8 m; a = 0.15 m; b = 3.2 m. For light-soil plowing, assume the following MTU operating conditions: frf = 0.12; fpf = 0.40; kf = 3
104 N m; n = 1500 kg/m3; a = 3°. Tables 1 and 2 present the results of calculating the oil consumption through burning when operating the MTU in transport or plowing modes. As can be seen in Tables 1 and 2, the most significant component is the change in amplitude and phase lag of MODP oil-pressure incrementation, which is reduced

320

V. P. Antipin et al.

Fig. 3 Oil-burning state surfaces for an MTU based on a K-744R-05 tractor with a PUN-8-40 plow when plowing light soils, as a function of travel speed at a suspension stiffness c1 ¼ 1500 kN=m: a for standard configuration, b for an adjuster-equipped diesel engine

Fig. 4 Oil-burning FR for an MTU based on a K-744R-05 tractor with a PUN-8-40 plow, breakdown by components for light-soil plowing at a suspension stiffness c1 ¼ 1500 kN=m and a travel speed t0 ¼ 2:01m=s: a for standard configuration, b for an adjuster-equipped diesel engine, 1 as a function of amplitude changes and phase lag of diesel-engine MODP oil-pressure incrementation, 4 as a function of MTU suspension vibrations in the longitudinal-vertical plane, 5 as a function of diesel-engine mechanical losses, 6 for the total dynamic component of oil consumption Table 1 Results of calculating the oil consumption through burning of an MTU based on a K-744R-05 with a PUN-8-40 plow: transport operations at t0 ¼ 2:01 m/s YaMZ-238ND5 diesel-engine configuration

Components of the dynamic component t ðxÞ, g/h at stiffness Cciri c1 ¼ 1500 kN/m 1 2 3 4

Standard 70 1 50 14 With an adjuster in the MODP 53 1 51 13 Regular component of the oil consumption through burning t , g/h Ccir0 Standard Actual oil consumption through burning With an adjuster in the MODP t t t Ccira ¼ Ccir0 þ Ccir ðxÞ

Dynamic components t ðxÞ, g/h as a Ccir function of stiffness c1 , kN/m 1500 1200 900 134 117 153

126 108

116 98

287 270

279 261

269 251

Method to Reduce Oil Burning in Diesel Engine …

321

Table 2 Results of calculating the oil consumption through burning of an MTU based on a K-744R-05 with a PUN-8-40 plow: light-soil plowing at t0 = 2.01 m/s YaMZ-238ND5 diesel-engine configuration

Components of the dynamic component p Cciri ðxÞ, g/h at stiffness c1 ¼ 1500 kN=m 1 3 4 5

Standard 70 1 34 27 With an adjuster in the MODP 53 1 35 25 p , g/h Regular component of the oil consumption through burning Ccir0 Standard Actual oil p p p consumption through Ccira ¼ Ccir0 þ Ccir ðxÞ burning With an adjuster in the MODP 404

Dynamic components p ðxÞ, g/h as a Ccir function of stiffness c1 , kN/m 1500

1200

900

132 114 290

126 108

119 101

422

416

409

398

391

by 17.0 g/h (24.3%) if an adjuster is installed. Therefore, the adjuster, which lowers the amplitude of fluctuations in the mechanical-loss moment and reduces the friction forces as a source of self-oscillations, does reduce the oil consumption through burning.

3 Results and Discussion This research has produced the following findings: • finding the specific fuel consumption and oil consumption through burning in diesel engines at constant load and crankshaft rotation speed in nominal mode is the main cause behind inconsistency with the manufacturer-claimed parameters, as MTU operation involves dynamic loads that must not be ignored; • the variable component of the drive load moment generates fluctuations in the rotation speed, which causes the oil leakage from the shaft-to-liner gaps to vary; • in the eigen-frequency range of the drive, the amplitude of rotation-speed and inertial-force fluctuations peaks, which causes more intense leakage and spraying of oil from the shaft-to-liner gaps; • maximum rotation-speed amplitude corresponds to maximum MODP oil-pressure amplitude and a phase lag of −180° from the variable component of the load moment at an eigen-frequency of 0.8 Hz (5.0 s−1). At the same time, minimum load-moment amplitude corresponds to maximum MODP oil-pressure amplitude, and vice-a-versa. This increases the oil leakage from the shaft-to-liner gaps, generating a dense oil mist, while the exhaust gases entering the cylinder crankcase creates a pressure above that in the over-the-piston space

322

V. P. Antipin et al.

at the end of the exhaustion stroke as well as at the beginning of cylinder air intake, which increases the amount of suspended oil entering the combustion chamber. It is known that [14, 20] that the oil ignition temperature is 153 °C higher than the air-fuel mixture ignition temperature. Excess oils in the air-fuel mixture inhibit the combustion process, and the mixture burns completely in the exhaust manifold. This is proven by the experimental frequency response of the manifold exhaust gas temperature [6, 7]. That the mixture burning is only complete in the manifold testifies not only the deterioration or design/technological imperfection of the shaft-to-liner or sleeve-to-piston-to-compression ring couplings, but also the lackluster design of the system that supplies oil to the friction pairs. Installing an adjuster in the MODP has a positive effect not only on the lubrication of friction pairs or oil consumption through burning, but also on the quality of burning the mixture in the combustion chamber, which reduces the temperature of exhaust gases in the manifold. Acknowledgements The authors would like to thank Vladimir Ivanovich Varavva, a Full Professor, and Gennady Vladimirovich Karshev, an Associate Professor, for their invaluable assistance in writing the manuscript.

References 1. Burstein LM, Kodyanov SV (1984) Opredeleniye raskhoda masla na ugar pri vybore velichiny zazora gilza-porshen (How the sleeve-piston gap affects oil consumption through burning). Dvigatelestroyeniye 10:8–10 2. Wenzel SV (1979) Primeneniye smazochnykh masel v dvigatelyakh vnutrennego sgoraniya (Use of lubricating oils in internal combustion engines). Khimiya, Moscow, p 240 3. Braslavsky MI (1973) Snizheniye raskhoda smazochnykh masel v sudovykh dizelyakh (Reducing the consumption of lubricating oils in marine diesel engines). Transport, Moscow, 102 p 4. Mokhnatkin EM (1980) Raschetnaya otsenka tolshchiny maslyanoy plyonki, formiruyemoy porshnevym koltsom (Estimating the thickness of oil films formed by the piston ring). Dvigatelestroyeniye 10:16–19 5. Grekhov LV, Ivashchenko NA, Markov VA et al (2013) Mashinostroyeniye (Mechanical Engineering). Encyclopedia, vol IV. Dvigateli vnutrennego sgoraniya (Internal Combustion Engines). Mashinostroyeniye, Moscow, 784 p 6. Antipin VP (2012) Energozatraty mashinno-traktornogo agregata (Energy consumption of an agricultural aggregate). Polytech Publishing House, St. Petersburg, p 324 7. Antipin VP, Durmanov MYa, Karshev GV (2017) Proizvoditelnost, energozatraty i resurs mashinno-traktornogo agregata (Performance, energy consumption, and service life of an agricultural aggregate). Polytech Publishing House, St. Petersburg, 484 p 8. Zhdanovsky IS, Nikolayenko AV (1981) Nadezhnost i dolgovechnost avtotraktornykh dvigateley (Reliability and durability of automotive and tractor engines). Kolos, Leningrad, p 295 9. Grigoryev MA, Ponomarev NI (1979) Iznos i dolgovechnost avtomobilnykh dvigateley (Wear and durability of automotive engines). Kolos, Leningrad, p 223

Method to Reduce Oil Burning in Diesel Engine …

323

10. Mishin IA (1976) Dolgovechnost dvigateley (Engine durability). Mashinostroyeniye, Leningrad, p 228 11. Antipin VP (1986) Faktory, opredelyayushchiye raskhod masla na ugar pri rabote dvigatelya v neustanovivshemsya rezhime (Determinants of oil consumption through burning in non-steady state engine operations). Dvigatelestroyeniye 5:12–13 12. Antipin VP, Shevtsov AA, Karpilovich AI (1987) Vzaimosvyaz raskhoda topliva i ugara masla pri rabote dvigatelya v neustanovivshemsya rezhime (Correlation of fuel consumption and oil burning in non-steady state engine operations). Dvigatelestroyeniye 7:10–12 13. Kanarchuk VYe (1970) Issledovaniye intensivnosti iznosa tsilindrov dvigateley YaMZ-236 (Studying the intensity of YaMZ-236 cylinder wear). Avtomobilnaya promyshlennost 4:6–7 14. Kanarchuk VYe (1978) Dolgovechnost i iznos dvigateley pri dinamicheskikh rezhimakh raboty (Durability and wear of engines in dynamic operation). Naukova dumka, Kiev, 256 p 15. Barinov KN, Antipin VP, Karpilovich AI, Shevtsov AA, Demakov VM (1989) Shchetinin YuV (1989) Sistema smazki dizelnogo dvigatelya (Diesel engine lubrication system). RU Patent 1497383:30 16. Kragelsky IV (1980) Treniye i iznos (Friction and wear). Mashinostroyeniye, Moscow, p 480 17. Kostetsky BI (1972) Mekhanicheskiye protsessy pri granichnom trenii (Mechanical processes in boundary friction). Nauka, Moscow, p 170 18. Wenzel SV (1963) Smazka dvigateley vnutrennego sgoraniya (Lubrication of internal combustion engines). Mashgiz, Moscow, p 179 19. Nikiforov OA, YeV Danilova (1986) Ratsionalnoye ispolzovaniye motornykh masel v sudovykh dizelyakh (Rational use of motor fuels in marine diesel engines). Sudostroyeniye, Leningrad, p 86 20. Wenzel SV (1977) Smazka i dolgovechnost dvigateley vnutrennego sgoraniya (Lubrication and durability of internal combustion engines). Tekhnika, Kiev, p 261

Mill Conditions Effect on Roughness of Injection Molds’ Forming Surfaces S. I. Kozhevnikov and V. F. Makarov

Abstract Molds work in relatively loaded conditions with a large number of working cycles; therefore, to obtain the required quality of parts, it is necessary to achieve a high quality of the forming surfaces of the molds. The analysis of existing processing technologies for molds has shown that existing technologies for ensuring the quality of forming surfaces do not take into account the influence of microgeometry on forming surfaces that arise during the machining of parts of molds. The effect on roughness resulting from the selected milling trajectory during the machining of mold surfaces is not taken into account. This leads to rapid wear of the forming parts and an increase in wear of the mold. In this paper, we study the roughness of casting surfaces, depending on the definition of milling parameters. To determine the functional dependence of the degree of roughness on the milling conditions, a full factorial experiment was performed with varying parameters and optimization factors. The functions of independent variables were presented: cutting speed—V, feed per revolution—So, and cutting depth—t. A set of experiments was performed using a CNC machine. Mathematical models are obtained. These mathematical models are applied when machining the molding surfaces of molds based on the choice of the optimal milling trajectory. The use of such a method allowed one to obtain an effect, speed up the technological process, and minimize abrasion of cast surfaces. Keywords Roughness of molding surfaces planning


Mill conditions
Statistical test

S. I. Kozhevnikov (&)
V. F. Makarov Perm National Research Polytechnic University, 29, Komsomolskiy Ave., Perm 614990, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_35

325

326

S. I. Kozhevnikov and V. F. Makarov

1 Introduction and Applicability With the benefit of experimental dependence of roughness degree on mill conditions of molding tools surfaces, there is a possibility to make comparison with the previous theoretical calculations of best values of mill conditions and emphasize some dependencies which practically will help to process molding surfaces. There are some recommendations in the literature [1–10] on milling of materials which are suitable for casting press molds production. However, the information isn’t full, and recommended mill conditions are often contradictory. Gross differences of values, recommended rate of rotation, chip load and cutting depth, stem from the fact that the experiments were performed by separate researchers under different processing conditions and different tool run-out. Due to this, the research results contained in the article of mill conditions effect on roughness of molding surfaces made of steel 38KhNM and 40Kh13 are of interest and allow choosing mill conditions while molding surfaces processing in a more substantiated way.

2 Research Objective Work objective is to find functional connection of processing surface roughness depending on prescriptible mill conditions. In order to state mill conditions affected on roughness degree, it is needed to preset the cutting speed, chip load, and cutting depth. Cutting speed is calculated by: Rate of cutter rotation is calculated by: v¼

pDn ; 1000

where n—rate of cutter rotation (rpm) and D—cutter diameter (mm). Milling feed may be calculated per tooth, per revolution or per minute. Feed per tooth—Sz is the distance that cutter moves in a time of cutter turn for the angle between two neighboring cutter teeth. Feed per revolution—So is the distance that cutter moves in a time of complete rotation. Feed per minute—Sm is the distance that cutter moves in 1 min. So ¼ Sz Z where Z—cutter teeth number. Sm ¼ So n or Sm ¼ Sz Zn Depth of milling or cut layer thickness is the distance between surface in process and processed surface; it is denoted as t and is measured by mm.

Mill Conditions Effect on Roughness of Injection …

327

In order to find the dependence of roughness degree on mill conditions Ra = f (V, So, t) steel 38KhNM and 40Kh13 using milling machine with CNC HAAS VF6 and using independent variables: cutting speed—V, chip load— So and cutting depth—t. Samples with 100 mm cutting length were proceeded. It corresponds to mean characteristics of processed parts. Each experiment was performed twice. As the result, choosing the main mill conditions, which can affect on roughness degree, it is possible to write an exponential function: Ra ¼ Ci nk SlM tm

ð1Þ

Plan 23 gives a sense of eight regression coefficients b0, b1, b2, b3, b12, b13, b23, and b123. In symbols, it can be written as [11]: b0 ! b 0 þ

k X

bii :

ð2Þ

i¼1

Set an objective to describe dependences of roughness degree while processing of molding surfaces made of 40Kh13 steel by cutters with rounded end on cutting speed—V, chip load—So, and cutting depth—t. A non-complete cubic function was taken as mathematical model. M fyg ¼ b0 þ b1 x1 þ b2 x2 þ b3 x3 þ b12 x1 x2 þ b13 x1 x3 þ b23 x2 x3 þ b123 x1 x2 x3 :

ð3Þ

3 Research Results Choose the main levels of factors close to practice, and variability intervals on the basis of real variation limits of factors (Table 1). All experiments are performed in accordance with planning matrix (Table 2), experiments results (Table 3). End-milling cutters with 6 mm in diameter were used. The experiment was performed twice in each factor space point that is why two-by-two cutters were used for each line of a plan. An order of cutters test was randomized using a random numbers table. For convenience formulary, regressional relationships were used: ~y ¼ b0 þ

3 X i¼1

bi x i þ

X i;j

bij xi xj þ b123 x1 x2 x3 :

ð4Þ

328

S. I. Kozhevnikov and V. F. Makarov

Table 1 Levels of factors and variation limits Factors level Upper (+1) Lower (−1) Zero (0) Variation limit

V x1 250 50 150 100

5.52 3.91

Table 2 Planning matrix 1 2 3 4 5 6 7 8

Table 3 Experiments results

So x2

ln x1

t x3

ln x2 −0.92 −2.3

0.4 0.1 0.25 0.15

ln x3

1.2 0.6 0.9 0.3

0.18 −0.51

х0

х1

х2

х3

х1х2

х1х3

х2х3

х1х2х3

+ + + + + + + +

− + − + − + − +

− − + + − − + +

− − − − + + + +

+ − − + + − − +

+ − + − − + − +

+ + − − − − + +

− + + − + − − +

Plan points v

Ra1

Ra2 y2

Ra yv

S2v

y1

1 2 3 4 5 6 7 8

1.55 0.8 1.85 1.1 1.05 0.42 1.29 0.67

1.61 0.88 1.93 1.22 1.15 0.46 1.43 0.73

1.58 0.84 1.89 1.16 1.1 0.44 1.36 0.7

0.0018 0.0032 0.0032 0.0072 0.005 0.0008 0.0098 0.0018

In such a case, we obtained an independent evaluations bi, corresponding coefficients bi . It is possible to show [12] that the coefficients will be calculated by a formula Pn bi ¼

v¼1 xiv yv

n

;

ð5Þ

Mill Conditions Effect on Roughness of Injection …

329

where i = 0, 1, 2 …, k—factor number; yv —mean response to r tests in the point with number v: Pr yv ¼

j¼1 yvj

r

:

ð6Þ

Zero is written for calculation b0. Calculations come to attribution to column y signs of a column of a relevant factor and algebraic addition of reported values because every factor (except xo) ranges at 2 levels +1 and −1. Dividing the result by number of points in a plan gives us a target coefficient. Calculate the equation coefficients. P

xov yv 1 ¼ ð1:58 þ 0:84 þ 1:89 þ 1:16 þ 1:1 þ 0:44 þ 1:36 þ 0:7Þ 8 n ¼ 1:134

b0 ¼

ð7Þ

Remaining values of bi are calculated similarly (Table 4). Rearranging of independent variables xi to non-dimensional variables Xi was performed using transformation equation, where ½(ln ximax − ln ximin) is the unit of a new scale: Xi ¼

2ðln xi  ln ximin Þ þ 1: ln ximax  ln ximin

ð8Þ

After substitution of Xi instead of xi, the equation will be as follows: y ¼ lg Ra0 ¼ b0 þ b1 X1 þ b2 X2 þ b3 X3 þ b1;2 X1 X2 þ


þ b2;3 X2 X3 þ b1;2;3 X1 X2 X3

ð9Þ

Plan implementation allowed deriving of equation in transformed variables of Xi: ~y ¼ 1:134  0:348X1 þ 0:143X2  0:234X3 þ 0:001X1 X2 þ 0:019X1 X3  0:014X2 X3  0:001X1 X2 X3 : Planning of experiments comes from statistical nature of dependences, that is why obtained constraint equations undergo detailed statistical analysis. The analysis’ purposes are dual. On the one hand—to get maximum of information from experiment results, on the other—to assure of dependence certainty, its accuracy.

Table 4 Mathematical model equation coefficients b0

b1

b2

b3

b12

b13

b23

b123

1.134

−0.348

0.143

−0.234

0.001

0.019

−0.014

−0.001

330

S. I. Kozhevnikov and V. F. Makarov

Dispersion, revealing an experimental error. Every experiment includes probability of error. Each experiment is performed several times upon the same conditions, in other words in each line of planning table, to reduce the probability. Likewise, dispersions are calculated as follows: Pr  S2v

j¼1

¼

yvj  yv r1

2 ;

ð10Þ

where r—number of replicated experiments in the points of plan. Dispersion of optimization parameter S2(y) is arithmetical mean of dispersions of all n different options of experiments, in other words averaged dispersion. While calculating of dispersion of optimization parameter squared difference between values of yvj in every experiment and the mean values from r repeated observations yv , it is needed to sum by the number of matrix lines and then to divide by n(r − 1). Pn S f yg ¼ 2

v¼1

S2v

n

ð11Þ

For our solution from Table 3, we have n X

S2v ¼ 0:0328:

v¼1

Then S2 f y g ¼

0:0328 ¼ 0:0041: 8

Usage of Fisher’s ratio test, while number of dispersions is more than 2, is ineffective, because in this case only maximal and minimal dispersions take part in evaluation. Cochran’s Q Test is appropriate when the number of replicated experiments is the same in all points of plan. Among all the dispersions, there is maximal S2vmax ‚ that can be divided by the sum of all dispersions by points. In our case, using Table 3, we can find G¼

0:0098 ¼ 0:299: 0:0328

ð12Þ

In corresponding table in work [13], we can find for fvmax = rv – 1—number of the degrees of freedom for such dispersion, fзнaм = N degrees of freedom and 5% significance value. Critical value GKp = 0.78. A hypothesis about homogeneity of variance is assumed so long as Cochran’s Q Test value is no larger than the table value like in our case.

Mill Conditions Effect on Roughness of Injection …

331

Significance test of every coefficient is performed independently. In this case, it is allowed to use a Student’s test. Confidential intervals for all the coefficients are equal in case of complete factorial experiment. First it is necessary to find regression coefficient dispersion S2{bi}. According to the equal repetition of experiments by the points with number of replicated experiments r, it will be calculated as follows: S2 fbi g ¼

S2 f y g nr

with fE = n(r − 1) degrees of freedom. In our case, pffiffiffiffiffiffiffiffiffiffiffiffiffi Sfbi g ¼ S2 fbi g ¼

rffiffiffiffiffiffiffiffiffiffiffiffiffiffi 0:0041 ¼ 0:016 8
2

ð13Þ

According to the formula dispersions of all, the coefficients are equal because they depend only on experimental error and number of tests. Now calculate ti— number value using formula: ti ¼

j bi j : Sfbi g

ð14Þ

Find ti—number value for our case and write into table (Table 5): t0 ¼

1:134 ¼ 60:67 0:016

Critical value is in work table [11] when n(r − 1) = 8 degrees of freedom and 5% significance value. In our case, tкp = 1.86. If ti > tкp, then we can discard the hypothesis and consider coefficient bi as significant. Otherwise, bi is considered statistically insignificant bi ¼ 0. In our case, there b12, b13, b23, b123 are insignificant. Now it is possible to create a confidential interval with length 2Δbi, where Dbi ¼ tKp sfbi g ¼ 1:86
0:016 ¼ 0:043:

ð15Þ

Coefficient is significant if its absolute value is bigger than a half of the confidential interval length. Mathematical model of object is created in the form of constraint equation of output parameter y and variables Xi, including only significant coefficients.

Table 5 ti—number value t0

t1

t2

t3

t12

t13

t23

t123

70.82

21.78

8.98

14.6

0.08

1.17

0.85

0.08

332

S. I. Kozhevnikov and V. F. Makarov

Regression equation almost completely describes the experiment’s results. The bigger an absolute value of regression coefficient value, the stronger influence on optimization criterion at stated interval of factors variation. If bi > 0, then increasing Xi causes increasing of optimization criterion. Otherwise, bi < 0, increasing Xi leads to decreasing of optimization criterion. Mathematical model looks like: ^y ¼ lg Ra ¼ 1:134  0:348X1 þ 0:143X2  0:234X3

ð16Þ

Using Formula (8), we can place Xi values to Eq. (16): ^y ¼ 2:52  0:433 ln x1 þ 0:207 ln x2 þ 0:674 ln x3

ð17Þ

After exponentiation, we have power dependence for actual values: Ra ¼ e2:8 n0:433 S0:207 t0:674 : o

ð18Þ

After calculating of model coefficients, it is necessary to check its availability or adequacy of model. P In our case, we can use a rule b0  y0 ! bii and evaluate the coefficients’ significance: jb0  yb0 j ¼ j1:134  1:145j ¼ 0:01129

ð19Þ

This value is less than experimental error fyg ¼ 0:064, as such quadratic effects are small to negligible, so one-dimensional model can be considered as valid. Similar solution can be used for describing the dependence of roughness value while processing of molding surfaces made of steel 38KhNM by rounded end cutters on cutting speed—V, chip load—So, and cutting depth—t.

Table 6 Experiments results

Plan points v 1 2 3 4 5 6 7 8

Ra1

Ra2

y1

y2

Ra yv

1.78 0.95 2.06 1.44 1.26 0.5 1.55 0.85

1.94 1.03 2.24 1.56 1.38 0.54 1.69 0.93

1.86 0.99 2.15 1.5 1.32 0.52 1.62 0.89

S2v 0.0128 0.0032 0.0162 0.0072 0.0072 0.0008 0.0098 0.0032

Mill Conditions Effect on Roughness of Injection …

333

Table 7 Coefficients of mathematical model equation b0

b1

b2

b3

b12

b13

b23

b123

1.356

−0.381

0.184

−0.269

0.0363

−0.001

−0.002

−0.002

The main factor levels and planning matrixes are still the same (Tables 1 and 2), experiments’ results (Table 6). After calculations, we can write the coefficients of mathematical model (Table 7). Plan implementation allows to get an equation with modified variables Xi: ~y ¼ 1:356  0:381X1 þ 0:184X2  0:269X3 þ 0:0363X1 X2  0:001X1 X3  0:002X2 X3  0:002X1 X2 X3 : Checking the homogeneity of variance using Cochran’s Q Test is appropriate because experimental value of Cochran’s Q Test is not bigger than table value. G¼

0:0162 ¼ 0:27: 0:06

Checking the significance of each coefficient using Student ti—criterion (Table 8). In our case, tкp = 1.86. If ti > tкp, then we can discard the hypothesis and consider coefficient bi as significant. Now it is possible to create a confidential interval with length 2Δbi, where Dbi ¼ tKp sfbi g ¼ 1:86
0:0217 ¼ 0:04: Mathematical model looks like: ^y ¼ lg Ra ¼ 1:356  0:381X1 þ 0:184X2  0:269X3 :

ð20Þ

Using Formula (8), we can place Xi values to Eq. (20): ^y ¼ 2:96  0:474 ln x1 þ 0:265 ln x2  0:775 ln x3 :

ð21Þ

After exponentiation, we have power dependence for actual values: Ra ¼ e2:96 n0:474 S0:265 t0:75 : o

ð22Þ

Table 8 ti—criterion values t0

t1

t2

t3

t12

t13

t23

t123

62.43

17.55

8.45

12.37

1.67

0.06

0.75

0.86

334

S. I. Kozhevnikov and V. F. Makarov

In our case, we can use a rule b0  y0 ! significance:

P

bii and evaluate the coefficients’

jb0  yb0 j ¼ j1:356  1:347j ¼ 0:0091: This value is less than experimental error fyg ¼ 0:087 as such quadratic effects are small to negligible, so one-dimensional model can be considered as valid. On the basis of models (18) и (22), there are some graphs of dependency of roughness of molding surface on V, So, t.

4 Conclusions A graph of dependency of Ra on V shows that there is decreasing of roughness height in case of increased cutting speed. Parameters So и t have less effect on roughness value: While So increasing Ra increases as a result of over pressure of instrument to material, while t increasing—there is feeble decreasing of roughness value at the shallow depth (Fig. 1).

Fig. 1 Graphs of dependency of Ra on V, So, t

Mill Conditions Effect on Roughness of Injection …

335

On the basis of these dependencies, there were set some mill conditions for rounded end cutters and applied for processing of shaping parts of mold tools such as matrixes, force plungers, inserts, and slides. 31 mold tools were processed, and the following research of the effect on roughness value on a long-term strength of molding tools is planned [14].

References 1. Batuev VA (1982) Research of special aspects of 3D milling using CNC machines with a purpose to increase quality of control software. In: Scientific and technical conference “Automation of programming and organization of CNC machines”, Chelyabinsk, pp 5–6 2. Batuev VA (1983) Processing of crooked spatial surfaces using milling CNC machines with subject to the constraints of accuracy. In: Scientific and technical conference “Ways of increasing of technical level and competitive ability of small-sized machine-tools with CNC”, Kaunas, p 46 3. Batuev VA, Batuev VV (2003) Assurance of accuracy of final milling of crooked spatial surfaces with an effect of influence of roughing work strategy. In: International scientific conference “Actual problems of design technology support of machinery production”, Volgograd, pp 60–63 4. Ostafyev VA, Globa AV, Globa LS (1983) Complex optimization of the end milling process. In: Machinery manufacturing reporter, vol 7, pp 39–40 5. Vishnyakov MA, Vashukov YA (2005) Design technology methods of quality assurance of engineering products. Samara, Samar Gos Aerocos Univ, 20 p 6. Basova EV (2015) A description of roughness of surface processed by high speed method of milling. Tr NGTU Im R.E.Alekseeva 1:95–100 7. Lapach SN (1997) Optimization of processing conditions of heatproof nickel alloys by the instrument made of ultra-hard material. Reliability of cutting tool and manufacturing systems optimization. Kramatorsk 8. Belyaev MP (2007) New ceramics for heatproof processing. Dvigatel, no 4–52 9. Belyaev VN, Tatarkin EYu (2013) Influence of lubricating-cooling process liquid on surface quality in ball burnishing. In: Scientific technologies in mechanical engineering reporter, vol 2. Bryansk, pp 12–18 10. Nekrasov NS, Loginov NY, Zotov AV, Kuzmich IV (2014) Choosing of optimal structure of end mill during roughing work of a mold tool. In: XXXV international research and training conference “Engineering sciences—bedside approach”, Novosibirsk 11. Katsev GT (1974) Statistical methods of cutting tools research, 2nd edn. Mashinostroenie Pabl, Moscow, p 231 12. Nalimov VV, Chernova NA (1965) Statistical methods of experiments’ planning. Nauka Publications, Moscow, p 340 13. Bolyshev LN, Smirnov NV (1965) Mathematical statistics tables. Nauka Publications, Moscow, p 474 14. Kozhevnikov SI, Makarov VF (2017) Wear-resistance increase of mold tools on the basis of direct forming of roughness on molding surface. Tula Izvestiya Tul’skogo Gos Univ 8(1): 253–260

Effect of Magnetic Processing on Mass Transfer in a Frictional Pair “Alloyed Steel–Carbon Steel” A. N. Gots, V. V. Zelinskiy and E. A. Borisova

Abstract An experimental study was performed to measure increased durability using the sample alloyed steel and carbon steel frictional pair. The study measured the effects of magnetic treatment of alloyed steel on diffusion transfer of alloying elements during friction into mating carbon steel. Measurements of the microhardness of the surface layer of carbon steel after friction without the use of magnetic treatment revealed the formation of secondary structures with high hardness. The decisive role is played by the mechanisms of chemical nature, the manifestation of which is due to the processes of enhanced mechanical activation and intensive deactivation of the adjacent layers of crystal lattices. The effect of preliminary magnetic treatment on a sharp decrease in the efficiency of formation of solid structures during friction has been established. The study of the elemental content of the surface layer of carbon steel after friction by spectral analysis established the diffusion transfer of carbide-forming elements and its significant deceleration in the processing of alloyed steel with a magnetic field. Based on a comparison of experimental results, it was shown that solid secondary structures can be carbide phases and intermetallic compounds with an increased share of covalent bonding, which is inherent in these chemical compounds precisely. Keywords Microhardness Structure Concentration





Diffusion
Hardening
Surface
Mass transfer


A. N. Gots Vladimir State University, 87, Gorky Street, Vladimir 600000, Russia V. V. Zelinskiy
E. A. Borisova (&) Murom Institute of Vladimir State University, 23, Orlovskaya Street, Murom 602264, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_36

337

338

A. N. Gots et al.

1 Introduction In spring-loaded automatic machines, feed rollers with wire, according to the conditions of interaction, create the friction pair “alloy steel–carbon steel.” Due to the inevitable skidding processes and repeated start–stop cycles in the loaded state, the frictional contact creates complex compression and shear deformation. There is a continuous formation of chemically pure microcontacts. Such processes, especially for metals with crystal lattices of similar structures, are accompanied by increased diffusion, mass transfer, and wear, which cause a reduction in the durability of the parts of the spring-loaded automatic machines [1, 2]. Numerous studies indicate that for such pairs, the chemical composition of the contact layers changes significantly to form secondary structures with a set of structural strength indicators different from the original structure. Thus, due to the insufficient structural stability of the surface layers of alloyed steel under conditions of intense combined deformation during frictional interaction, secondary structures with increased hardness can be formed due to mass transfer in the outer layers of carbon steel, which provide abrasive surface destruction of alloyed steel [3, 4]. Since mass transfer is mainly controlled by diffusion in the material of the roller, the altered energy state of the alloyed steel’s crystal lattice may have a favorable effect on its structural stability and wear. Such a state in the video is created by preliminary pulsed process of the magnetic field (PMF) [5–9].

2 The Experimental Research Experimental modeling of mass transfer in the fusion pair “alloy steel–carbon steel” was carried out during deformation by friction according to the “indenter–disk” scheme. The fixed indenters were nonmagnetic (base) and magnetized by a pulsed magnetic field samples of high-alloyed steel H12MF, cut from natural feed rollers of a spring-loaded automaton. The contact surface of indenters with an initial hardness of 64 HRc (920 HV) of cylindrical shape with a radius of 65 mm (as in the feed rollers) and a height of 4 mm formed a “cylinder–plane” touch circuit with the disk surface. Movable disks with a radius of 55 mm were made of carbon spring steel 65 and subjected to heat treatment to obtain a microstructure corresponding to the patented spring winding wire. The initial hardness of the disks was 45 HRc (46 HV). In order to maximally approximate the conditions of the experiments to the conditions of wire supply when winding the springs, the duration of the tests for each friction pair was set sufficiently short (2 min). Due to the short duration of the contact, the transformation due to friction in the material of the disk occurs in thin layers. Therefore, even an X-ray structural qualitative analysis of the transformed layers gives results that fall within the detection error. Due to the additivity of repetitive thermodynamic processes

Effect of Magnetic Processing on Mass …

339

initiating mass transfer in the microvolume contact, accumulation of transferred atoms of alloying elements occurs in the disk lattice. Moreover, their concentration in the surface layer, which is different from the initial layer, corresponds to the accumulated transformations that have occurred as a result of mass transfer phenomena. The total result of mass transfer and the formation of solid secondary structures in this case can be confirmed by measuring the microhardness and the changed elemental composition of the friction surface of the disk. Such an assessment of diffusion transformations is considered reliable and is used in research. On average, for disks prior to friction experiments, the initial microhardness was 460 HV (4.6 GPa). For the friction tracks on the disks, in order to increase the reliability of the results and their statistical processing, the number of microhardness measurements in each experiment was assigned to be sufficiently large (up to 100) and spread throughout the track. Measurements showed that after friction with a base (nonmagnetized) stainless steel indenter, the maximum and adjacent microhardness values turned out to be more than 3.5 times higher than the initial microhardness. The range of values turned out to be rather large (Fig. 1a). However, after friction with a magnetized indenter, the maximum values of microhardness decreased almost 2 times. The range of scattering of values also halved (Fig. 1b). It is known that during friction, the deformation hardening mechanism occurs first. Moreover, its maximum rate of increase in hardness is up to 1.4–1.5 [10–12]. The multiplicity of the solid solution-hardening mechanism is much higher, since it consists in the formation of chemical compounds with an increased proportion of

Fig. 1 Microhardness of the disk surface and the ratio of its values after friction. a Without a PMF, b with a PMF

340

A. N. Gots et al.

covalent bonds, providing a substantially increased hardness [13–15]. To assess the effects of different hardening mechanisms, it was taken into account that lower microhardness values correspond to lower deformation, which are conventionally attributed to HV < 10 GPa, and higher microhardness values to chemical nature mechanisms correspond to HV > 10 GPa. Such an approach made it possible to reveal that in friction with a nonmagnetized indenter a significant fraction of the disk’s friction surface contains solid newly formed structures in the form of island inclusions. Their appearance can be caused only by the action of a chemical mechanism with fairly simple implementation under given conditions, for example, a solid solution. This is indicated by the large number of “high” values of microhardness (46%), as seen in Fig. 1a. This high of microhardness values in the upper half of the scattering field is simply unattainable for the deformation mechanism. At the same time, the remaining “low” values (54%) in the lower half of the field are fully consistent with the capabilities of this mechanism. Thus, the presence of a significant amount of secondary solid structures of chemical origin indirectly confirms the possibility of intensive mass transfer of alloying elements during friction on the surface of conjugated carbon steel and their participation in chemical transformations. Under friction with a magnetized indenter, “high” values of microhardness turned out to be only 9%, low—91% (Fig. 1b). This indicates a sharp decrease in the number of newly formed highly solid inclusions on the friction surface strengthened by the solid solution mechanism of chemical nature. In the overwhelming majority, there were areas with a low degree of hardening characteristic of the deformation mechanism. It follows that friction with a magnetized indenter of the original elements for noticeable chemical transformations is not sufficient. This suggests a significant slowdown in the transfer of alloying elements from the magnetized indenter to the surface layers of the conjugated carbon steel. Friction formation of solid secondary structures on the surfaces of the disks also confirms, by microscopic examination, the presence of defects in the form of scratches on the surfaces friction indenters, whose hardness exceeds the hardness of the disks. Measurements of microhardness led to the conclusion that at low densities, secondary structures of chemical origin on the surface of the friction track are difficult. It is thus challenging to give an accurate quantitative assessment of the results showing the transfer of carbide-forming elements using X-ray microanalyzers to give local (point) analyzes. The estimate may be within the detection error due to the very small diameter (up to 1 lm) of the relevant zone for taking the analysis. Therefore, a quantitative assessment of the elemental content was carried out on a Q4 TASMAN optical emission spectrometer with an excitation zone diameter of up to 4 mm. Such an analysis zone allows one to obtain averaged concentration values over a fairly large surface area. After cleaning the analysis zone with argon, this method provides high accuracy and measurement result reproducibility. These spectrometer readings under typical analysis conditions for alloying chemical elements of steel provide high accuracy with the following detection limits (given in percentages): Cr—0.000005, Mn—0.00006, Mo—0.0002, W—0.0001, V—0.00004.

Effect of Magnetic Processing on Mass …

341

The actual transfer of friction from the indenter to the disk for each chemical element was detected by comparing the average value (3–5 analyzes) of its changed concentration (by taking the analysis in the surface layer of the friction track on the disk) with the initial concentration of the element in the disk material (taking the analysis on free disk surface). A full cycle of concentration measurements was carried out for disks that have undergone friction with a nonmagnetized indenter, i.e., without application of processing by a magnetic field (without PMF) and having passed friction with the magnetized indenter (with PMF).

3 Discussion and Evaluation of Results The frictional contact of the feed rollers and wire is associated with enhanced elastic–plastic deformation of the surface layers. Due to the localization of deformations in very thin layers, the energy state of the latter corresponds to an extremely high mechanical activation. In [10, 14], it is noted that in many cases the energetics of mechanical activation of a frictional contact corresponds to potentials up to 10 eV, and this is sufficient not only for local melting of the metal in the contact zone, but even its ionization. With such high levels of energy perturbation in the local areas of interacting metal lattices, the bonds between atoms and electrons characteristic of the crystal structure are lost. An unstable structural-energetic state is created, corresponding to the quasi-liquid structure of the substance. In the interlattice space, an intermediate microvolume contact with special properties is formed (the third body), manifesting itself as an open thermodynamic system in which mass transfer and energy exchange takes place according to the laws of thermodynamics [15, 16]. Disordered layers of atoms undergo a sharp decrease in energy barriers for mass transfer and chemical reactions. In this case, the supply of energy for the passage of chemical reactions in the third body is carried out not by heating, but by mechanical activation. Therefore, the influence of the temperature factor on the possibility of producing chemical reactions is significantly reduced [4, 16, 17]. The increase in the changed concentration of each element in comparison with the initial concentration was confirmed by all analyzes. The scatter plot of the reanalysis values for tests without PMF for all elements averaged about 20%, for trials with PMF—about 10%, which indicates a high reproducibility of the results. Evaluation of the effectiveness of increasing the content of each element in the surface layer of the disk was made according to the multiplicity K of an increase in the average altered concentration C relative to the based concentration Cb using “(1)”. K¼

C Cb

ð1Þ

The multiplicity values of K for carbon and all carbide-forming elements for friction without PMF and with PMF are shown in Table 1.

342

A. N. Gots et al.

Table 1 Multiplicity values of concentration of carbon and carbide-forming elements in the surface layer Experiment type

C

Cr

Mo

V

Without PMF With PMF

1.55 1.15

3.51 1.24

5.27 1.29

2.67 1.25

An important result of the assessment in terms of K is to establish the fact that without the use of PMF in the surface layer of the friction track, there was a sharp increase in the carbon content and carbide-forming elements. This is explained as follows: (1) In H12MF steel, the initial carbon content is substantially higher than in steel 65, and it is present in unbound form. In addition, with local heating, it is able to precipitate from martensite simultaneously with the formation of e-carbide particles. Thus, one of the driving forces in the transfer of carbon into carbon steel is the Fick’s gradient of concentrations. (2) In the contact pair “alloy steel–carbon steel,” especially under combined compression and shear deformations, in the outer layers of the carbon steel crystal lattice, which is less durable, numerous defects such as dislocations are formed. A large mechanical stress gradient is created. It is believed [14, 18] that in frictional interaction it is a more powerful driving force for mass transfer than the concentration gradient. Therefore, as it is well known, segregations of carbon atoms (Cottrell clouds on dislocations) are instantaneously formed on the lattice defects, while forming additional cementite in the outer layers of the carbon-elongated carbon lattice. At the same time, the chromium, molybdenum, and vanadium atoms, which appeared as a result of mechanical activation in the third body and having an increased affinity for carbon (which explains their high carbide-forming ability), diffuse intensively in the defect accumulation area (i.e., in the surface layer of carbon steel) by replacing gland. Their increased concentration in the surface layer is shown in Table 1. As a result, dispersed particles of secondary structures with increased hardness are formed on the surface defects of the disk, mainly in the form of doped cementite and intermetallic compounds with an increased proportion of covalent bonds. Due to concentration fluctuations in certain places, chromium, molybdenum, and vanadium may be sufficient to form mixed carbides. Vanadium, almost not soluble in cementite, can form an independent carbide, characterized by high hardness. The implementation of these processes confirms the increase in the microhardness of the surface layer of the friction track by 3.5 times (Fig. 1a). The effect of such hardening is due to the formation of covalent bonds in secondary microstructures, and this is possible precisely in the carbide phase and intermetallic compounds with the presence of Cr, Mo, and V [15]. Applied calculations show that, thanks to microdiffusion mechanisms [19], a contact duration of 10−3–10−4 s is sufficient for the accumulation of finely dispersed carbides in a layer up to 1 µm thick.

Effect of Magnetic Processing on Mass …

343

For friction using OMPs, it was found that an increase in the content of all elements also occurred, but to a much lesser extent. The decrease in the concentration of C, Cr, Mo, and V due to the use of OMP occurred, respectively, 1.35, 2.83, 4.1, and 2.14 times. Moreover, the greatest effect of reduction corresponds to Mo and V, forming compounds with the highest hardness. Taking into account that diffusion is a function of time [20], the efficiency of mass transfer by the differential characteristic of the process in the form of rate of accumulation of the average increase in the concentration of surface layer elements over time can be evaluated, as shown in Eq. “(2)”. V¼

C  Cb ; t

ð2Þ

where t—is the duration of friction in constant conditions, min. Essentially, this characteristic is analogous to the flux density of transferred atoms. The results of the evaluation of MLE on the rate of accumulation of the average surface concentration in all the carbide-forming elements of the X12MF steel are shown in Fig. 2. It can be seen that in experiments without MLE the quantity of carbon atoms determines the scale of carbide formation. For friction with the use of PMF, the efficiency of transfer and accumulation of carbide-forming elements in the disk material (and, consequently, the efficiency solid particle formation causing the indenter to wear-out) is significantly reduced. The multiplicity of mass transfer decrease of carbon and carbide-forming elements’ Km is shown in Fig. 3. It can be seen that due to PMF, the diffusion rate of carbide-forming elements can be reduced by 10–14 times. Analysis of the results on the change in the elemental content of chromium and molybdenum shows that, taking into account the dispersion of transferred atoms along the entire friction track on the disk and the ratio of contact areas between the Fig. 2 Rate of accumulation of the average increase in the concentration of carbon and carbide-forming elements in the surface layer

344

A. N. Gots et al.

Fig. 3 Rate of decrease of mass transfer of carbon and carbide-forming elements

indenter and the disk (about 1:50), significant depletion of the indenter surface with diffusers is evident in a relatively short time interval. Therefore, we should expect a noticeable decrease in its wear resistance and durability.

4 Conclusion (1) As a result of experimental modeling, using the example of the friction pair “alloyed steel–carbon steel,” the possibility of (a) diffusion transfer of carbon and carbide-forming elements from alloyed steel to the surface layer of carbon steel, (b) the formation in the surface layer of carbon steel of numerous inclusions of secondary structures in the form of a carbide phase and intermetallic compounds with microhardness up to 18 GPa. (2) Experiments have shown that the repeated treatment of alloyed steel with a magnetic field (up to 14 times) reduces the diffusion transfer of alloying elements into conjugated carbon steel during friction.

References 1. Bowden FP, Tabor D (1986) The friction and lubrication of solids, second corrected edition. Oxford classic text in the physical sciences edition. Oxford University Press, Clarendon Press, UK 2. Bowden FP, Tabor D (1964) The friction and lubrication of solids. Oxford at the Clarendon Press, UK, 544 p

Effect of Magnetic Processing on Mass …

345

3. Quinn T, Rowson DM (1980) Application of the oxidational theory of mild wear to the sliding wear of low-alloy steel. Wear 85:1–20 4. Quinn T (1978) An experimental study of heat and surface temperature sliding steel interfaces and their relation to oxidational wear. ASLET 21:78–86 5. Borisova EA, Zelinskyi VV (2015) On the mechanism of ferromagnetic materials wear reduction. Procedia Eng 129:111–115 6. Zelinskyi VV, Borisova EA (2015) About the quantum-mechanical nature of wear on magnetized cutting and deforming tools. In: International conference on mechanical engineering, automation and control systems (MEACS), 1–4 Dec 2015 7. Zelinskyi VV, Borisova EA (2017) Reducing wear during coalescence in steel-steel tribo-systems by magnetic action. Procedia Eng 206:662–667 8. Mohamed MK et al (2012) Effect of magnetic field on the friction and wear displayed by the scratch of oil lubricated steel. Int J Eng Technol IJET-IJENS 12(6) 9. El Mansori M, Lafdy K, Palmer D (2002) Enhanced wear resistance and tools durability using magnetization, metal cutting and high speed machining. Kluwer Academic, Netherlands, pp 301–310 10. Tabor D (1992) Friction as a dissipative process. In: Singer IL, Pollock HM (eds) Fundamentals of friction: macroscopic and microscopic processes. NATO ASI series E: applied sciences, vol 220. Kluwer Academic Publisher, Dordrecht, pp 3–24 11. Rabinowicz E (1995) Friction and wear of materials. Wiley, Hoboken 12. Markov DP, Kelly AD (2002) Establishment of a new class of wear: adhesion initiated catastrophic wear. Int J Appl Mech Eng 7(3):887–901 13. Pollock HM (1992) Surface forces and adhesion. In: Singer IL, Pollock HM (eds) Fundamentals or friction: macroscopic and microscopic processes, NATO ASI series E: applied sciences. Kluwer Academic Publisher, Dordrecht 14. Giessibl FJ, Herz M, Manhart J (2002) Friction traced to the single atom. PNAS 99(19): 12006–12010 15. Elliot JA (2011) Novel approaches to multiscale modelling in material science. Int Mater Rev 56(4):207–225 16. Kwon YS, Andreev VM, Lomovsky OJ, Bokbanov BB (2005) Alloys and compounds, vol 386, no 1–2 17. Fink M (1930) Wear oxidation a new component of wear. Trans Am Soc Steel Treat 18:1026– 1034 18. Krause HR (1971) Tribomechanical reaction in the friction and wearing of lron. Wear 18(3): 403–412 19. Archard JF (1953) Contact and rubbing of flat surfaces. J Appl Phys 24:981–988 20. Heinicke G (1984) Tribochemistry. Akademie‐Verlag, Berlin, 495 p

Strength Parameters of Hardening Cylindrical Workpieces by Tapered Roller A. A. Udalov, A. V. Udalov and S. L. Vasilevykh

Abstract A theoretical study of the force parameters of the process of hardening a cylindrical part with a tapered roller by the method of surface plastic deformation (SPD) has been performed. Using an approximate model of propagation of plastic deformation, an engineering technique has been developed that allows one to set the process power modes taking into account the mutual influence of the required degree of material hardening and work hardening, as well as the geometrical parameters of the workpiece and the deforming roller. The force of deformation is determined taking into account the hardening of the material changing according to the power law and the surface area of the contact of the roller with the workpiece. The degree of material deformation and the force of deformation in the process of hardening are investigated depending on the depth of work hardening, the angle of inclination of the forming roll deforming roller and its diameter. The results of the theoretical study are in good agreement with the known experimental data and can be used in the development of technological operations for hardening machine parts by running in rollers. Keywords Cylindrical workpiece Plastic deformation Slip cone





Tapered roller
Hardening


1 Introduction Hardening technologies using surface plastic deformation (SPD) are widely used to increase the fatigue strength of shafts and axes of technological and transport machines [1–6]. Theoretical and experimental studies of recent years are aimed at A. A. Udalov (&) Ural Federal University named after the first President of Russia B.N. Yeltsin, 19, Mira Street, Ekaterinburg 620002, Russia e-mail: [email protected] A. V. Udalov
S. L. Vasilevykh Vyatka State University, 36, Moskovskaya Street, Kirov 610000, Russia © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_37

347

348

A. A. Udalov et al.

Fig. 1 Schematic diagram of hardening of a cylindrical billet by roller: a hardening by torus roller; b hardening by taper roller; R—profile radius of toroidal roller; a—angle of inclination of the forming roller to the initial surface of the workpiece; D—roller diameter; d—the outer diameter
of the hardened workpiece; F—force of deformation;
S—axial feed of roller; xW —angular speed of workpiece rotation; xR —angular speed of roller rotation

increasing the efficiency of the use of this technological operation [7–13]. As a deforming element used balls or rollers of various shapes (Fig. 1a, b), technological modes of the process of SPD (plastic deformation force and axial roller feed), providing the required depth and degree of deformation of the reinforced layer, are assigned in accordance with the operating conditions of the part, geometric parameters of the workpiece and deforming elements [1–6, 14–18]. The degree of deformation (degree of hardening) of the material is the main factor influencing the force of deformation. In the well-known works [4], the degree of deformation of the hardening layer is determined by the geometrical parameters of the residual imprint or the magnitude of the relative change in hardness [6]. However, the work of the deformation force (without taking into account elastic deformations and friction forces) is spent primarily on the plastic shaping of the material in the deformation zone of the roller that occurs under the contact surface. Therefore, when determining the power parameters of the SDP process, it is necessary to take into account the size and shape of the source of plastic deformation, inside which the material is in conditions of comprehensive non-uniform compression. In addition, there are practically no works devoted to the study of the hardening process with a tapered roller or a cylindrical roller set at an angle a to the workpiece surface. As a rule, balls or toroidal rollers are used as a deforming element. However, it has been established that a roller with a conical working surface with a strengthening capacity is about 1.5 times more effective than a toroidal one [19]. The goal of the article is to develop an engineering technique that allows the force modes of the process of SPD to be set for a cylindrical billet with a tapered roller, taking into account the mutual influence of the required degree of material hardening and work hardening, as well as the geometrical parameters of the billet and deforming roller.

Strength Parameters of Hardening Cylindrical Workpieces …

349

2 The Main Provisions and Analytical Dependencies Consider the process of hardening the outer surface of the cylindrical billet taper roller (Fig. 2a, b). The axis of rotation of the roller will be located parallel to the axis of rotation of the workpiece. The depth of work hardening during SPD is determined using an approximate model of propagation of plastic deformation [20], in accordance with which the construction of sections of slip cones in planes yz is performed (Fig. 2a) and xy (Fig. 2b). The top C of the cross section of the slip cone ABC corresponds to the greatest depth of work hardening Ha in the plane yz, and the top K cross section of the slip cone AEK corresponds to the maximum depth of work hardening HD in the plane xy. From the deformation point of view, the depth of work hardening H is the initial size of the element within which plastic deformation propagates, and the depth of penetration of the roller h into the material of the workpiece, it is absolute deformation. In accordance with this position, the degree of deformation of the material in the plane yz is determined by the formula eyz ¼

h sin 2a ¼ Ha 1 þ sin 2a

ð1Þ

sin 2a  where h ¼ ‘
tga; Ha ¼ BC
cosðb  aÞ ¼ ‘
12þcos 2 a ðb ¼ 45 Þ. Thus, the degree of deformation in the plane is completely determined by the angle of taper a (Fig. 3) and can be provided in one pass of the roller. However,

Fig. 2 Construction of slip cone sections: a the cross section of the slip cone ABC in the plane yz; b the cross section of the slip cone AEK in the plane xy; ‘—length of the geometric deformation center; h—depth of penetration of the roller into the billet metal; Ha and HD —depth of work hardening in the planes yz and xy, respectively; n1  n1 and n2  n2 —normals to the generator of the roller at the points A and B, respectively; b—the angle of inclination of the lines of principal shear stresses to the normals of the contact surface (in calculations it is assumed b ¼ 45 ); hD — contact angle of the roller with the workpiece in the plane xy

350

A. A. Udalov et al.

Fig. 3 Dependence of the degree of deformation eyz on the angle of inclination of the forming roller a to the original surface of the workpiece according to the Eq. 1

excessive deformation leads to the formation of surges and the destruction of the surface layers of the material. In this regard, the optimal ratio h=H ¼ 0:1 was determined [17], which provides the best quality of processing, where H is the depth of work hardening (the depth of propagation of plastic deformation). Accordingly, in the plane xy the degree of deformation will be equal to eyz ¼

h HD

where HD ¼ EK
cosðb  hD Þ þ ðh  hD Þ ¼ D2


ð2Þ h

sin 2D
sinðb þ hD Þ sinðbhD =2Þ

þ ðh  hD Þ,

where hD ¼ AE
sin h2D ¼ D
sin2 h2D ; ! jO1 O2 j2 þ ðD=2Þ2 ðd=2Þ2 hD ¼ arccos ; 2
jO1 O2 j
ðD=2Þ where jO1 O2 j ¼ d þ2 D  h—the distance between the axes of rotation of the roller and the workpiece in the process of SPD. The average value of the degree of deformation within the deformation zone is determined by the formula e¼

eyz þ exy : 2

ð3Þ

3 Determination of the Force of Deformation The deformation force F required for surface hardening with a tapered roller is determined with the following assumptions: • The contact pressure between the roller and the workpiece is equal to the resistance to deformation of the material with regard to hardening;

Strength Parameters of Hardening Cylindrical Workpieces …

351

Fig. 4 Geometric 3D–model of contact interaction of a tapered roller and a cylindrical workpiece

• Contact pressure is evenly distributed over the contact surface. Taking into account the accepted assumptions, the force of plastic deformation is determined by the formula F ¼ rs
AC

ð4Þ

where AC —contact area of the roller and the workpiece, mm2 , pffiffiffi rs ¼ r0:2 þ g
ðe
3Þb —resistance to deformation of the workpiece material with regard to hardening, MPa [21], where r0:2 —yield strength of the material of the original billet; g and b— empirical hardening coefficients of the workpiece material; e—the degree of material deformation determined by the Eq. 3. The area of the contact surface of the roller and the workpiece A is determined on the basis of the model (Fig. 4) constructed in the CAD program.

4 Numerical Implementation of Analytical Dependencies and Practical Application of the Developed Methodology The degree of deformation in Eqs. 1–3 and the force of deformation in Eq. 4 are determined on the basis of the following initial data: the material of the workpiece, the outer diameter of the initial workpiece, the diameter of the deforming roller along the calibrating girdle, required depth of the reinforced part layer, the angle of inclination of the forming roller to the initial surface of the workpiece. The calculation is made for steel 45 (hot-rolled and annealing at 950 °C), having the following mechanical characteristics [21]: r0:2 ¼ 404:9 MPa, g ¼ 502:6 MPa, b ¼ 0:323. The outer diameter of the original billet d = 100 mm. The diameter of the deforming roller is taken from the range D = (0.05, …, 1)d. The depth of work hardening is assumed to be equal to the depth of propagation of plastic deformation in the plane yz, i.e., H ¼ Ha .

352

A. A. Udalov et al.

The required depth of work hardening is usually assigned from the ratio H = (0.01, …, 0.05)d [17, 18]. The calculation of the force parameters is performed for two values of the hardening depth: at the minimum depth Hmin ¼ 0:01
d ¼ 1 mm and at the maximum depth Hmax ¼ 0:05
d ¼ 5 mm. In [4], it was shown that, depending on the processing conditions, the ratio between the depth of penetration of the deforming roller and the depth of work hardening is within h=H ¼ 0:05; . . .; 0:5. Based on these practical recommendations, the calculation of power parameters is performed for three values of the ratio h=H—0.05; 0.1 and 0.25, which in accordance with Eq. 1 are in fact the degree of deformation eyz . For each of the three values eyz of the graph in Fig. 3 is determined by the corresponding angles of inclination of the generator roller to the original surface of the workpiece: at eyz ¼ 0:05, a  1:5 ; at eyz ¼ 0:1, a  3:5 ; at eyz ¼ 0:25, a  10 . Thus, the force parameters of the process of hardening a cylindrical billet with a diameter d = 100 mm were determined for two values of the work hardening depth Hmin ¼ 1 mm and Hmax ¼ 5 mm with different angles of inclination of the forming deforming roller ða ¼ 1:5 ; 3:5 ; 10 Þ. In accordance with the presented methodology, the analysis of the influence of the hardening depth H, the angle of inclination of the forming roller a and the relationship D=d to the degree of material deformation e (Fig. 5) and the required strain force F in the hardening process with a roller (Fig. 6a, b) was performed. Analysis of the graphs (Fig. 5a, b) shows a monotonic decrease in the degree of deformation e of the hardening layer with an increase in the ratio D=d and an increase e with an increase in the angle a and hardening depth H. Analysis of the graphs (Fig. 6a, b) shows a monotonous decrease in the required force F of deformation with a decrease in the angle a and an increase F with increasing depth of work hardening H and ratio D=d. When assigning force modes, one should take into account a reduction in the quality of the treated surface with an increase in the degree of deformation of the hardened layer.

Fig. 5 Dependence of the degree of deformation e on the relationship D=d at different angles a: a H = 1 mm; b H = 5 mm; 1: a ¼ 1:5 ; 2: a ¼ 3:5 ; 3: a ¼ 10 ; d = 100 mm

Strength Parameters of Hardening Cylindrical Workpieces …

353

Fig. 6 Dependence of the force of deformation on the relationship at different angles: a H = 1 mm; b H = 5 mm; 1: a ¼ 1:5 ; 2: a ¼ 3:5 ; 3: a ¼ 10 ; d = 100 mm

5 Conclusion The presented engineering technique allows you to assign power modes of the surface hardening process with a tapered roller, taking into account the mutual influence of the required degree of material hardening and work hardening depth, as well as the geometrical parameters of the workpiece and the deforming roller. The obtained analytical and graphical dependences are in good agreement with the known experimental data [1–6, 14] and can be used in designing the processes of hardening of shafts and axes using SPD methods. Acknowledgements This article was prepared with the financial support of the Competitiveness Enhancement Program of the URFU.

References 1. Shkolnik LM, Shakhov VI (1964) Tekhnologiya i prisposobleniya dlya uprochneniya i otdelki detaley nakatyvaniem (Technology and adaptations for hardening and finishing of parts by rolling). Mashinostroenie Publishing, Moscow 2. Kudryavtsev IV (1965) Povyshenie dolgovechnosti detaley mashin metodom poverhnostnogo naklepa (Increasing the durability of machine parts by the method of surface hardening). Mashinostroenie Publishing, Moscow 3. Braslavskiy VM (1975) Tekhnologiya obkatki krupnyh detaley rolikami (Technology for running large parts with rollers). Mashinostroenie Publishing, Moscow 4. Suslova AG (2014) Tekhnologiya i instrumenty otdelochno–uprochnyayushchey obrabotki detaley poverhnostnym plasticheskim deformirovaniem (Technology and tools for finishing and hardening of parts by surface plastic deformation). Mashinostroenie Publishing, Moscow 5. Zaydes SA (2014) Obrabotka detaley poverhnostnym plasticheskim deformirovaniem (Surface plastic deformation processing). Irkutsk State Technical University Publishing, Irkutsk 6. Smelyanskiy VM (2002) Mekhanika uprochneniya detaley poverhnostnym plasticheskim deformirovaniem (Mechanics of hardening of parts by surface plastic deformation). Mashinostroenie Publishing, Moscow 7. Meyer D (2012) Heat treatment-free production of surface hardened components by mechanically induced hardening. Dissertation, University of Bremen

354

A. A. Udalov et al.

8. Brinksmeier E, Garbrecht M, Meyer D, Dong J (2008) Surface hardening by strain induced martensitic transformation. Prod Eng Res Dev 2:109–116 9. El-Tayeb NSM, Low KO, Brevern PV (2006) Influence of roller burnishing contact width and burnishing orientation on surface quality and tribological behaviour of Aluminium 6061. J Mater Process Technol 186:272–278 10. Brinksmeier E, Garbrecht M, Meyer D (2008) Cold surface hardening. CIRP Ann 57:541–544 11. Meyer D, Brinksmeier E, Hoffmann F (2011) Surface hardening by cryogenic deep rolling. Procedia Eng 19:258–263 12. Zaydes SA, Ngo KK (2018) Vliyanie novoy kinematiki obkatnogo rolika na kachestvo uprochnennogo sloya pri poverhnostnom plasticheskom deformirovanii (Influence of new kinematics of the rolling roller on the quality of the hardened layer during surface plastic deformation). Izv vuzov Mashinostr 2:58–67 13. Sazanov VP (2017) Ocenka ehffektivnosti metodov uprochnyayushchey obrabotki cilindricheskih detaley iz stali 30HGSA (Evaluation of the effectiveness of methods for hardening processing of cylindrical parts made of steel 30HGSA). Izvestiya vuzov. Povolzhskiy region. Tekhnicheskie nauki 1:128–137 14. Kheyfets SG (1952) Analiticheskoe opredelenie glubiny naklyopannogo sloya pri obkatke rolikami stal’nyh detaley (Analytic determination of the depth of the riveted layer during rolling by rollers of steel parts). New research in the field of strength of engineering materials. Proc CNIITMASH 49:7–17 15. Kudryavtsev IV, Petushkov GE (1966) Vliyanie krivizny poverhnostey na glubinu plasticheskoy deformacii pri uprochnenii detaley poverhnostnym naklyopom (The influence of the curvature of surfaces on the depth of plastic deformation in the case of hardening of parts by surface hardening). Vestn Mashinostroeniya 7:41–43 16. Sidyakin YI, Trunin AV, Abakumova SY (2014) Analiticheskoe issledovanie ostatochnyh napryazheniy v sploshnyh valah posle uprochnyayushchey obrabotki poverhnostnym plasticheskim deformirovaniem (Analytical study of residual stresses in solid shafts after strengthening treatment by surface plastic deformation). Vestn Mashinostroeniya 6:62–70 17. Kudryavtsev IV (1976) Osnovy racionalnogo vybora rezhimov uprochneniya malyh galteley valov poverhnostnym plasticheskim deformirovaniem (Fundamentals of rational choice of regimes for hardening of small fillets of the shafts by surface plastic deformation Questions of strength of large machine parts). Proc CNIITMASH 112:190–200 18. Kudryavtsev IV (1983) Vybor osnovnyh parametrov uprochneniya valov obkatyvaniem rolikami (Selection of the basic parameters of hardening of rollers by roller rolling). Vestn Mashinostroeniya 4:1–8 19. Zelinskiy VV (2012) Novyy podhod v teorii poverhnostnogo uprochneniya valov nakatyvaniem (A new approach in the theory of surface hardening of rolls). Vestn Permskogo Nacionalnogo Issledovatelskogo Politekhnicheskogo Univ 1:18–23 20. Gubkin SI (1947) Teoriya obrabotki metallov davleniem (Theory of metal forming). Metallurgyzdat Publishing, Moscow 21. Bogatov AA (2002) Mekhanicheskie svoystva i modeli razrusheniya metallov: ucheb. posobie dlya vuzov (Mechanical properties and models of the destruction of metals: textbook. Manual for high schools). UGTU-UPI, Ekaterinburg

Formation of Schemes Generating Geometric Structure of Machine Parts O. V. Kolesnikova, V. E. Lelyukhin and F. Yu. Ignatev

Abstract One of the important tasks of modern engineering is the creation of the theoretical foundations of the technological processes design in parts manufacture. Of particular relevance to the formalization of design acquired in recent years due to the widespread tendency to design automation and the creation of smart industries. This is confirmed by the concept of the Fourth Industrial Revolution (Industry 4.0). In the course of designing manufacturing parts technology on machine tools, it is necessary to solve two problems related to the generation of geometric configuration. One of the tasks is to find solutions to get the desired shape for each surface of the part. Another task is to find the conditions, tools, and mechanisms to ensure a given relative position of the part surfaces in a three-dimensional space. The article considers the formal method developed by the authors for the formation of basing schemes and their sequence. The basis of this technique is to simulate the processing of all surfaces of a part with a sequential change of bases. The modeling of the machining process is carried out on the basis of coordinating dimensional connections between the surfaces of the part. Direct keeping of the specified dimensions allows you to guarantee the requirements of the relative position of the part surfaces to each other.







Keywords Engineering Technology design CAPP communications Processing base Industry 4.0








Dimensional

1 Introduction The experience of modern engineering has allowed to accumulate a sufficiently large amount of knowledge in the field of manufacturing technologies formation for various kinds of products, machines, and mechanisms. The authors analyzed the

O. V. Kolesnikova (&)
V. E. Lelyukhin
F. Yu. Ignatev Far Eastern Federal University, 8 Sukhanov St., Vladivostok 690091, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_38

355

356

O. V. Kolesnikova et al.

publications of various scientists and researchers, which allowed to identify three main approaches to the design of manufacturing technology parts [1–7]. One approach is to use the previously applied technological methods and procedures (past experience) as analogues for the formation of technology. In this case, the development of technology consists in determining the constructive identity (similarity) of the parts with the parts for which there is already a developed technological process. This approach proved to be effective as a result of almost a century of practice in Russia and abroad. Another approach emerged a few decades ago as a concomitant element of the intensive development of information technology. It is based on a variety of mathematical tools for random modeling of various combinations. This approach has no real practical application yet and is mainly of interest only in theoretical terms [1, 2, 8]. The third approach is a design based on well-defined formal dependencies that adequately reflect the real laws of interaction and transformation of material objects. With the development of technology as a science, the number of formalized design elements is increasing, but there is no general formal design system [2, 8]. The article provides a formal method developed by the authors for the formation of schemes for base and their sequence to meet the requirements of the relative position of the part surfaces to each other when machining parts on machines. This technique can be a mechanism for solving one of the two key tasks of generating a geometrical configuration of machine parts—providing a given relative location of surfaces.

2 Problems of Design Automation Technology Manufacturing Geometric Configuration of Parts The Computer-Aided Process Planning (CAPP) systems are based on two approaches: variant and generative [1, 2]. Variant approach of processing is to search in the database of a typical or similar technological process with its subsequent modification. The generative approach is in search of standard structural elements into part and applying standard technological operations for them. However, with the use of existing CAPP systems, the human’s leading role in technology design remains [9–12]. Russian computer-aided design of technological processes (CADTPs) are based on three approaches: (1) the formation of technological processes based on typical technological processes, (2) the formation of technology using group operations and equipment, and (3) the formation of technology using constructivetechnological modules. The idea of typification was first proposed by A. P. Sokolovsky. The technique is based on the classification of processes, which is based on the classification of parts. The classification is organized according to the scheme “class—subclass—

Formation of Schemes Generating Geometric Structure …

357

group—subgroup—type.” It is used to process parts of the same type and develop standard technological processes [13]. The group method is proposed by Mitrofanov [4]. Its characteristic feature is the presence of group operations, which are formed in such a way that on a single machine with a single adjustment, technological operations can be performed for the manufacture of similar elements for various parts designs. Another way to design technological processes is a modular approach by Bazrov [3]. The main principle in it is to represent the part as a set of structural modules in the form of a combination of surfaces, for each of which a corresponding set of technological procedures (technological modules) is formed. The formation of the working process consists in determining the complete set of required technological modules. On closer examination, it can be noted that all of the above approaches incorporate the use of past experience in the form of already existing technological processes, operations, or individual previously tested procedures. In fact, all these methods are based on the previously established correspondences between the original description of the structural part and the technological solutions adopted. To solve the basic problem of design automation, it is necessary to create a formal theory that allows one to synthesize technological solutions for real parts as applied to the existing production conditions. As shown in [9, 14–16], the main difficulties to design machining parts technology on machine tools are the tasks of providing a geometric configuration. In the general case, there are two such tasks: (1) the formation of technology to ensure the shape of each surface of the part and (2) the technology of ensuring the relative position of these surfaces relative to each other. The article discusses the principle and provides a technique for the formal synthesis of schemes for the geometric structure of machine parts. An example of generating schemes of a geometric structure based on the use of graphs of dimensional connections, determining the relative position of the surfaces, is considered.

3 Representation of Part Surfaces by a Six-Dimensional Vector The part can be represented as some closed subspace bounded by a set of surfaces. The boundary of a subspace is a closed surface or a collection of surfaces. The surface of a part of almost any complexity can be considered as a set of elementary surfaces, such as a plane, a cylinder, and a sphere. Under the elementary surface, the authors understand the surface obtained by using two types of forming lines: a straight line and a circle [9, 15, 17]. The relative position of the surfaces of the part is governed by the coordinating dimensions and additional requirements for parallelism, alignment, perpendicularity, etc. [9, 18, 19].

358

O. V. Kolesnikova et al.

As a basis for the formal description of the position of the body in space in the works [9, 15, 16], six degrees of freedom are used. Linear displacements in three dimensions are indicated Xl, Yl, Zl, angular turns—Xa, Ya, Za. The coordinates of the vector V{Xl, Yl, Zl, Xa, Ya, Za} determine the presence or absence of freedom. The state of complete freedom of the element in three-dimensional space is described by the vector Va{0, 0, 0, 0, 0, 0}, and the uniquely determined position is Vb{1, 1, 1, 1, 1, 1} [9]. Then the position of the XOY plane can be described by the vector Vp{0, 0, 1, 1, 1, 0}, the position of the cylinder parallel to the X axis will be represented by Vc{0, 1, 1, 0, 1, 1}.

4 Dimension Graphs Dimensional link graph construction is considered on the example of the part shown in Fig. 1a. The numbers in the circles correspond to the numbering of the surfaces. Graphs of dimensional relationships are formed for each of the six degrees of freedom (Fig. 1b). Here, solid lines indicate connections that are explicitly defined in the drawing. As seen in Fig. 1b, all the graphs for the measurements Xl, Yl, Zl turned out to be connected and acyclic [20, 21]. When considering corner rotations, for each dimension Xa, Ya, Za, the graph is formed as the exclusive sum of the graphs of adjacent linear dimensions [14, 18]. For example, for measuring Xa, the graph is obtained by excluding the addition of graphs in the dimensions Yl, and Zl. As seen in Fig. 1b graph in the direction of Ya is obtained as the sum of the graphs Xl and Zl, with the exception of the repeated vertex 7 [14, 18, 22].

Fig. 1 Part with the applied coordinating dimensions (a) and graphs of dimensional links (b)

Formation of Schemes Generating Geometric Structure …

359

In Fig. 1b, it is clear that all the graphs, for the dimensions Xa, Ya, Za, are disconnected. The reason for this is the absence of perpendicularity given on the drawing between the complexes of parallel planes 1 and 2, 3 and 4, 5 and 6. Representation of dimensional links between the surfaces of parts in the form of graphs in six dimensions is a tool for formal analysis of the necessity and sufficiency of relations for the relative orientation of the surfaces of any geometric object [9, 14–16, 18]. For example, an isolated vertex of a graph indicates the absence of a specified size between this and the other surfaces of the part. The presence of a cycle in the constructed graph indicates a closed contour of dimensions and the occurrence of uncertainty in the orientation of surfaces [18, 19, 21, 23]. To use formal tools in further modeling, it is necessary to eliminate any ambiguity in the form of redundancy or insufficiency of dimensional links [14]. Therefore, disconnected graphs for measurements of Xa, Ya, Za, (Fig. 1b) should preferably be supplemented with the minimum necessary number of edges to ensure connectivity, but it should not be allowed to have cycles in them. The result of adding links is shown in Fig. 2. Here the added links are indicated by a dotted line. For the manufacture of machine parts on machine tools, the starting material is taken in the form of a blank, the geometric configuration of which differs from the part itself. Therefore, the so-called rough surfaces are used as initial bases for orientation of the workpiece relative to tool movements. In the considered example, the complex of “rough” surfaces was chosen as the initial bases: planes 2R, 3R, and 5R parallel to planes 2, 3, and 5, respectively. In Fig. 2, the “rough” surface is marked by squares. Each of the “rough” planes 2R, 3R, and 5R limits three degrees of freedom. Moreover, any two mutually perpendicular planes always compete in fixing a turn around one of the coordinate axes. For example, planes 2 and 3 (2R and 3R, respectively) duplicate the fixation of rotation around the Z axis (Fig. 1a). To eliminate this uncertainty, a number of restrictions are imposed on the fixation of the degrees of freedom for “rough” planes. As a result, it is assumed that the 2R plane fixes three degrees of freedom Xl, Ya и Za, the 3R surface limits two degrees of freedom Zl and Xa, and the 5R surface ensures the fixation of only one degree of freedom Yl. The “rough” surfaces excluded from consideration, duplicating the fixed degrees of freedom, are marked with a crossed out square (Fig. 2). Fig. 2 Formed connected acyclic graphs for six dimensions

360

O. V. Kolesnikova et al.

5 Algorithm of Consistent Production of Part Surfaces The following describes the algorithm for generating schemes for generating the geometric structure of machine parts. This algorithm is convenient for modeling the sequence of processing surfaces of the part. To implement the algorithm, a block matrix is built (Fig. 3a). The cell of the matrix is a block of six elements. The matrix consists of two parts: the top and bottom. The lower part is a square matrix of dimension n  n, where n is the number of surfaces that need to be processed (Fig. 3a). Units in the cells of the matrix indicate the relationship between the surfaces, which are regulated in the drawing of parts. In Fig. 3a, it can be seen that the upper part of the matrix contains the connections between the “rough” (existing on the workpiece) and the surfaces to be treated (which are not yet on the workpiece). The procedure for finding a solution consists in sequentially examining the columns of the matrix until the element-wise logical sum of cells (raw surfaces) of the matrix coincides with the diagonal cell. So as shown in Fig. 3a, it can be seen that if surface 2R is taken as base, surface 1 can be processed. Therefore, in the first step, the column with the number 1 is deleted, and the row with the number 1 is transferred to the upper part of the matrix (Fig. 3b). After the first step, the upper part of the matrix in row 1 appeared a table covering the diagonal of the raw surface with the number 2. This means that for surface treatment 2, a real base was found in the form of surface 1. Then the next step is to transfer the row number 2 to the upper part of the matrix and delete column 2 from it (Fig. 4a).

Fig. 3 Adjacency matrix a source and b after the first step

Formation of Schemes Generating Geometric Structure …

361

Fig. 4 Adjacency matrix: a after two steps and b after four steps

At the next step, in the upper part of the matrix, it can be seen that, using the surface of 3R and 1 as a set of bases, surface 4 can be processed (Fig. 4a). When surface 4 is obtained, a match appears in column 3, which allows surface 3 to be obtained using real surface 4 as the base (Fig. 4b). In Fig. 4b, it can be seen that there is only one column with number 6, and the element-wise sum at the top of which (surfaces 5R, 2, 3) covers the diagonal cell in the row and column with number 6. Thus, in the next step, there is the possibility of surface treatment 6. It is easy to see that further modeling leads to a complete degeneration of the matrix, which indicates the convergence of the algorithm.

6 Conclusion The article outlines the main provisions of the methodology for the formation of basing schemes and the sequence of surface treatment. In fact, this technique is a formal tool for the synthesis of schemes for generating the geometric structure of machine parts. In this case, dimensional links are used as an information basis, which determine the relative location of the surfaces of the part.

362

O. V. Kolesnikova et al.

This technique can be a mechanism for solving one of the two key tasks of generating the geometrical configuration of machine parts—providing a given relative location of surfaces.

References 1. Skander A, Roucoules L, Klein Meyer JS (2008) Design and manufacturing interface modelling for manufacturing processes selection and knowledge synthesis in design. Int J Adv Manuf Technol 37:443. https://doi.org/10.1007/s00170-007-1003-2 2. Yusof Y, Latif K (2014) Survey on computer-aided process planning. Int J Adv Manuf Technol 75:77. https://doi.org/10.1007/s00170-014-6073-3 3. Bazrov BM (2001) Modular technology in mechanical engineering. Engineering 367 p 4. Mitrofanov SP (1976) Scientific organization of labor of machine-building production. Engineering 712 p 5. Legoff O, Hascoet JY (2010) Technological form defects identification using discrete cosine transform method. Int J Adv Manuf Technol 52–58 6. Desrochers A (ed) Methodologie de conversion des specifications geometriques de tolerance en zones d’incertitude. Sherbrooke (Québec), Canada 7. Gologlu C (2004) A constraint-based operation sequencing for a knowledge-based process planning. J Intell Manuf 15(4):463–470 8. Halevi G (2014) Industrial management—control and profit, Lecture notes in management and industrial engineering 1, Springer International Publishing, Switzerland. https://doi.org/ 10.1007/978-3-319-03470-6_1 9. Lelyukhin VE, Kolesnikova OV, Kuzminova TA, Antonenkova TV (2015) Formal representation of parts for the automated development of manufacturing technology. Sci Mag High Technol Eng 11:32–36 10. Tolérancement modal De la Métrologie vers les Spécifications. Laboratoire Systèmes et Matériaux pour la Mecatronique d’Annecy et du Centre Technique de l’industrie du Décolletage à Cluses (2009), 192 p 11. Villeneuve F, Mathieu L (eds) (2010) Geometric tolerancing of products. Library of Congress Cataloging-in-Publication Data, Wiley-ISTE, 400 p 12. Lelyukhin V (2015) The theory of synthesis of methods for shaping the surfaces of a part. LAP LAMBERT Academic Publishing, Germany, 80 p 13. Zhukov EA, Kozar I, Rozovskii BY, Degtyarev VV, soloveitchik AM (2000) Engineering technology. Part II. Design of technological processes: studies. Manual, Izd-vo SPBSTU, SPb, 498 p 14. Lelyukhin VE, Kolesnikova OV, Ignatev FYu (2018) Research of influence of structure of the dimensional relations on processing of the parts. In: Materials of the international scientific-practical conference “Actual issues of contemporary science”. FECIT, Vladivostok, April, pp 3–27 15. Lelyukhin VE, Kuzminova TA, Kolesnikova OV (2015) Influence of geometrical configuration details on the technology of its manufacture. Modern scientific research and innovations, no 7. http://web.snauka.ru/issues/2015/07/56318. Accessed 19 Sept 1998 16. Lelyukhin VE, Antonenkova TV, Kolesnikova OV (2016) Structural-parametric representation of the relative position of surfaces part. Science Magazine. Vestnik of the engineering school of the FEFU, vol 1, pp 3–9 17. Gorelskaya LV, Kostryukov AV, Pavlov SI (2008) Descriptive geometry. Tutorial. OSU, Orenburg, 85 p

Formation of Schemes Generating Geometric Structure …

363

18. Tsitsiashvili GS, Lelyukhin VE, Kolesnikova OV, Osipova MA (2017) Formal design of structure process in machining parts. Appl Math Sci 11:1573–1580. https://doi.org/10.12988/ ams.2017.75167 19. Matveev VV (1980) The dimensional analysis of technological processes. Mechanical engineering, Moscow, 592 p 20. Cayley AA (1889) Theorem on trees. Quart J Pure Appl Math 23:376–378; Collected mathematical papers, vol 13. Cambridge University Press, pp 26–28 (1897) 21. Diestel R (2005) Graph theory, Elektronik Eedition. Springer, Heidelberg 22. Lelyukhin VE, Kolesnikova OV (2015) Analysis and calculation of dimensional chains based on graphs of dimensional relationships. Vestn ES FEFU 4:3–12 23. Lelyukhin VE, Ignatev FYu, Drenin AS, Kolesnikova OV (2018) Geometry to describe the real components of machines. Mod High Technol 8:95–99

Optimization of Cutting Parameters in Milling by Means of System Nyquist Plot R. M. Khusainov, P. N. Krestyaninov and D. D. Safin

Abstract The paper dwells upon improving milling performance. Vibrations in machining represent one of the most significant factors that hinder such performance. Increasing some of the cutting parameters might result in greater vibrations. The so-called Nyquist criterion of dynamic stability is based on the Nyquist plot. On the other hand, the expression to compute the Nyquist plot of the dynamic system of a mill will include all the basic cutting parameters in relation to the cutting process. To determine the values of the Nyquist plot, an experiment is performed, consisting in machining a test workpiece on a milling machine. Cutting is performed once on the low cutting conditions. By using the experimentally found Nyquist plot values, one can build a mathematical model to find the performance-optimized cutting parameters, at which no significant vibrations occur. The proposed method minimizes the labor intensity of tailoring the cutting parameters to a specific production setting. Keywords Nyquist plot


Dynamic stability
Milling

1 Introduction 1.1

State-of-the-Art in Modern Mechanical Engineering

Optimizing the cutting parameters to attach high performance at a low cost is one of the most relevant problems of cutting operations [1]. Performance improvements are hindered by the decrease in dynamic stability associated with increasing some parameters such as cutting width and feed rate, as well as the system becoming resonant due to a higher cutting speed. Increased tool wear and tear associated with higher cutting parameters entails greater operation costs. This is especially relevant R. M. Khusainov (&)
P. N. Krestyaninov
D. D. Safin Branch of Kazan University in Naberezhnye Chelny, 68/19, Mira, Naberezhnye Chelny 423812, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_39

365

366

R. M. Khusainov et al.

for milling, as this type of machining features the most independent cutting parameters: width, depth, feed per tooth, and cutting speed [2]. Besides, this machining method employs a lot of different cutting diagrams, where various combinations of these parameters may be used. This complicates the optimization of such parameters.

1.2

Relevance

If the cutting parameters do not match the conditions of a specific production facility, there may occur vibrations entailing a deterioration in the tool durability as well as its probable subsequent failure [3]. All of this results in insufficient performance and high costs [4]. Accidents and emergencies are possible in some cases. This is especially applicable to roughing. In case of finishing milling operations, cutting depth and width will, as a rule, be unambiguously determined by the stock to be removed, while the feed rate will be easy to determine as a function of the required machined surface quality [4, 5]; however, there are no unambiguous recommendations as to how to set the cutting parameters in roughing. Many tool manufacturers do issue recommendations, but those apply only to the maximum boundary values. Using these values in a real-world situation will likely induce non-permissible vibrations. The problem is particularly relevant for machining the modern high-strength materials [6, 7]. The feed rate and the cutting speed are specified in many machining handbooks [8, 9]. However, these parameters often need to be adjusted to become machine-, material-, workplace-, and setting-specific [10]. Worse still is the case of cutting width and depth; although there are some recommendations on setting such parameters [11], and these recommendations are rather generic and need to be tailored to the system-specific dynamics so as to avoid non-permissible vibrations.

1.3

Statement of Problem

Thus, there must be a developed method to optimize the cutting parameters for a specific production setting; such optimization shall not be costly or time-consuming. Machining performance must be the basic optimization criterion, as in case of roughing, it is the most important factor of reducing the labor intensity and attaining a specific vibration level. Stronger vibrations make the tool less durable and more prone to failure; vibrations are also what prevent the equipment user from maximizing the cutting performance. The method must be experiment-based, as computational data are based on substantial assumptions, which reduces the accuracy and reliability of results; besides, many values can only be found experimentally.

Optimization of Cutting Parameters in Milling …

367

2 Main Part 2.1

Theoretical Basis of the Study

The problem can be solved by the following methods: 1. Random search method. A version of this method is proposed in [12]. The method sets such cutting parameter values as to form the central point and the vertices of a hypercube in the cutting parameter space. Test cuts are made using these parameters to find the vibration amplitude and the performance. Multicriteria decision-making methods are used to find such parameters, where both performance and vibrations are acceptable. The drawback is that the method requires multiple experiments. Thus, finding four milling parameters requires nine experiments pursuant to the existing recommendations. Besides, the method only uses the boundary (but not the intermediate) values of the cutting parameter range, which prevents truly efficient optimization, as optimal values may be contained within the range rather than at its boundaries. 2. Directed search method. This approach may use a multifactor optimization method such as gradient or coordinate descent. This method finds the truly optimal cutting parameters. Its disadvantages are as follows: • Since there are multiple independent parameters, the method may require numerous experiments, which makes its feasibility questionable; • The main function that gives insight into the vibrations at such cutting parameters is the frequency response of the mill dynamic system; this function is significantly nonlinear, and an analytical expression in this method will itself require a separate experiment or modeling, making the method too labor-intensive. 3. Nyquist plot of the mill dynamic system. The dynamic stability of a system can be evaluated by the Nyquist criterion, which implies obtaining and studying these characteristics. Numerous studies have shown [13–20] that the dynamic stability of a technological system is affected by the values and combinations of cutting parameters. In particular, there is a well-known mathematical expression for finding the largest chip size that can be attained without vibrations; this expression determines the limit width of the cut, at which the system is still dynamically stable. Using the Nyquist plot allows setting such cutting parameters, at which the performance is high while vibrations are low. The transfer function of an open-loop dynamic system is found by the formula [21]:

368

R. M. Khusainov et al.

WðjxÞ ¼

ð1 

T12 x2

Kyc
Kb
þ T2 jxÞ 1 þ

m n


a0Vn0 jx



ð1Þ

where Kyc is the dynamic compliance of an elastic system; a0 is the thickness of the cut layer; b is the width of the cut layer; K is the specific cutting force; T1 is the inertial constant of time; T2 is the damping time constant; n0 is the mean chip shrinkage time; m/n is a constant coefficient; v is the cutting speed.

2.2

Method for Optimizing the Cutting Parameters

This formula is fundamental to developing a mathematical model for cutting parameter optimization. Using the model involves the following steps: 1. Generating the equations of the dynamic system transfer function with due account of an equivalent elastic system and cutting system (1). 2. Generating a system of equations for finding the unknown coefficients that are part of the transfer-function expression (1). The variables thereof are the unknown coefficients that describe the state of the elastic system and cutting process: 8 Kyc
Kb   ¼ W1 > > a n > < ð1T12 x21 þ T2 jxÞ 1 þ mn
0V 0 jx ... > K
Kb > yc >   ¼ Wn : a n ð1T12 x2n þ T2 jxn Þ 1 þ mn


ð2Þ

0 0 jx n V

where n is the number of unknown coefficients. 3. Experimental Nyquist plotting of the dynamic system—the values Wi (2). Nyquist plotting must be done at the frequencies specified in the equation system (2). Figure 1 presents the experimental Nyquist plot of a JMD3CNC vertical milling machine when milling a 45-steel workpiece by an end mill being 8 mm in diameter at the following cutting parameters: spindle rotation speed: 700 rpm, feed: 28 mm/min, cutting depth: 0.5 mm, and cutting width: 1 mm. 4. Solving the system (2), e.g., by numerical methods, derives the unknown coefficients. Using MathCAD Prime 3.0 to solve the system (2) has produced the following values: Kyc = 7.95  10−6 mm/N, T1 = 0.354  10−3 s, T2 = 0.428  10−4 s, K = 2136 N/mm2, n0 = 2.8, mn ¼ 1:1. 5. Generating the optimized cutting mathematical model. Performance is the basic optimization criterion. In this case, performance means the cutting performance, i.e., stock removal per unit of time: P ¼ B
t
smin

ð3Þ

Optimization of Cutting Parameters in Milling …

369

Fig. 1 Experimental Nyquist plot for JMD3CNC milling operation

where t is the cutting depth, mm; B is the cutting width, mm; smin is the feed per minute, mm/min. The cutting parameters are limited by vibration stability. This limit is found from the Nyquist stability condition, i.e., the dynamic system transfer-function hodograph must not cover the −1 point on the real axis. In the ensured vibration stability in cutting, one needs redundant stability k. In this case, the expression to find the constraint used to account for dynamic stability will be written as: Kyc
Kb
ð1  T12 x2 þ T2 jxÞ 1 þ

m n


a0Vn0 jx

  k
1

ð4Þ

In addition, the cutting parameter ranges specified by the tool manufacturer will be used as constraints. Thus, the mathematical model to derive optimal cutting parameters is written as: The optimization criterion: P ¼ B
t
smin

ð5Þ

Constraints: Kyc
Kb
ð1  T12 x2 þ T2 jxÞ 1 þ

m n


a0Vn0 jx

tmin  t  tmax

  k
1

ð6Þ ð7Þ

370

R. M. Khusainov et al.

Bmin  B  Bmax

ð8Þ

smin  s  smax

ð9Þ

vmin  v  vmax

ð10Þ

6. Optimal cutting parameters can be derived by solving this model: MathCAD Prime 3.0 has produced the following cutting parameters: cutting depth: 0.75 mm, cutting width: 4-mm, spindle rotation speed: 1100 rpm, and feed: 180 mm/min.

3 Conclusions The proposed method for optimizing the cutting parameters is experimental and analytical in nature. Experimentation makes data reliably fit the specific production setting. Computations enable the user to quickly find the cutting parameters by a minimum number of experiments. In this case, the experiment is to record the Nyquist plot, which is used to solve the mathematical model herein built. Modern math packages solve this problem in virtually no time. Thus, the proposed method enables the use to optimize the cutting parameters to attain better performance and minimum costs while engineering the process.

References 1. Safarov DT, Kondrashov AG, Glinina GF et al (2017) Algorithm of calculation of energy consumption on the basis of differential model of the production task performed on machines with computer numeric control (CNC). IOP Conf Ser Mater Sci Eng 240(1):012060 2. Balabanov IP, Kondrashov AG (2014) Shaping of cutting part of angle milling cutters with nonzero geometry. World Appl Sci J 30(12):1731–1734 3. Ryabov EA, Yurasov SY, Yurasova OI (2016) Parametric modeling of ball end mills. Russ Eng Res 36(9):784–785 4. Balla OM, Zamashchikov YI, Livshits OP et al (2006) ISTU Publ., Irkutsk 5. Lavrentyeva MV, Chimitov PY (2017) Implementation of recognition algorithm with NXOpen API in siemens NX. In: 2017 international conference on industrial engineering, applications and manufacturing (ICIEAM), 2017, pp 1–4 6. Gavariev RV, Savin IA (2018) Research of the mechanism of destruction of compression molds for casting under pressure of color alloys. Solid State Phenom 284:326–331 7. Shastin VI, Kargapoltcev SK, Gozbenko VE et al (2017) Results of the complex studies of microstructural, physical and mechanical properties of engineering materials using innovative methods. Int J Appl Eng Res 12(24):15269–15272 8. Metal Cutting Parameters (Rezhimy rezaniya metallov): Handbook (1995) NIITavtoprom, Moscow

Optimization of Cutting Parameters in Milling …

371

9. Guzeyev VI, Batuyev VA, Surkov NV (2005) Cutting parameters for turning, drilling, milling, and boring CNC machines. Mashinostroyeniye, Moscow 10. Akhatov R, Govorkov A, Zhilyaev A (2015) Software solution designing of “The analysis system of workability of industrial product” during the production startup of aeronautical products. Int J Appl Eng Res 10(21):42560–42562 11. Serebrenitsky PP (2007) Some specific features of high-speed machining (Nekotorye osobennosti vysokoskorostnoy mekhanicheskoy obrabotki). Metalloobrabotka 4:6–15 12. Khusainov RM, Avdeyev IV, Krestyaninov PN, Safin DD (2017) Optimization of cutting parameters in milling by performance and vibration criteria (Podbor optimalnykh rezhimov rezaniya pri frezerovanii po pokazatelyam proizvoditelnosti i vibroustoychivosti). In: Proceedings of the international research conference on innovative mechanical-engineering technologies, equipment, and materials 2017 (MNTK IMTOM 2017), P1, Kazan, pp 147–152 13. Gorin YY, Kryazhev AY, Tatarkin YY et al (2015) Improving the vibration stability in end milling (Povysheniye vibroustoychivosti protsessa tortsovogo frezerovaniya). Polzunovsky Vestnik 2:43–48 14. Method for Testing Mid-Size General-Purpose Lathes for Vibration Stability in Cutting (Metodika ispytaniya tokarnykh stankov srednikh razmerov obshchego naznacheniya na vibroustoychivost pri rezanii) (1961) ENIMS, Moscow 15. Kozochkin M (2009) Particularities of vibrations in metal cutting (Osobennosti vibratsy pri rezanii materialov). STIN 1:25–29 16. Podurayev VN (1988) Technological diagnosis of cutting by acoustic emission (Tekhnologicheskaya diagnostika rezaniya metodom akusticheskoy emissii). Mashinostroyeniye, Moscow 17. Zharkov IG (1986) Vibrations in blade-based machining (Vibratsii pri obrabotke lezvynym instrumentom). Mashinostroyeniye, Leningrad 18. Yakovlev EY (2009) Improving the quality of metal finishing in CNC end milling by neural-network modulation of cutting parameters (Povysheniye kachestva chistovoy obrabotki metallov pri tortsevom frezerovanii na stankakh s ChPU s ispolzovaniyem neyrosetevoy modulyatsiii rezhimov rezaniya). Inf Secur Quest 2:65–69 19. Averyanova IO, Shestakov NA (2013) Analysis of chipping in cutting (Analiz protsessa struzhkoobrazovaniya pri rezanii). Vestnik mashinostroyeniya 2:48–74 20. Kozochkin MP (2013) Cutting process stability (Ustoychivost protsessa rezaniya) Vestn mashinostroyeniya 2:77–81 21. Kudinov VA (1967) Machine dynamics (Dinamika stankov). Mashinostroyeniye, Moscow

Comprehensive Evaluation of Shaft Manufacturability: Mathematical and Information Models A. Sychugov, Yu. Frantsuzova and V. Salnikov

Abstract Mechanical product, like any product designed to meet specific needs, has properties, constituting its quality. Ensuring the manufacturability of the product design at the design stage is a significant and at the same time the most difficult task of the engineer. The purpose of this research is to improve the quality of project design solutions in mechanical engineering with the help of a comprehensive qualimetric evaluation of shaft parts’ manufacturability, considering their production cost, labor intensity, design and technological unification of components. The proposed manufacturability evaluation includes two main components: technological cost and unification, expressed by design and technological factors. To check the model’s adequacy, the correlation coefficient was determined by expert estimates of the manufacturability level of such parts as solids of revolution and the estimates calculated by the proposed method. During the operation, the level of manufacturability of the part can be calculated according to four parameters: consistency of bases; surface roughness; unification of structural elements (SE); and processing labor intensity.













Keywords Manufacturability Labor intensity Unification Engineering Shaft

1 Introduction Mechanical product, like any product designed to meet specific needs, has properties, constituting its quality [1–3]. The combination of product properties, determining design’s ability to optimize resource cost in the production and operation for the given quality indicators, is the manufacturability of the product design [4–7]. Ensuring the manufacturability of the product design at the design stage is a significant and at the same time the most difficult task of the engineer [8, 9]. A. Sychugov
Yu. Frantsuzova
V. Salnikov (&) Tula State University, 92, Lenin Ave., Tula 300012, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_40

373

374

A. Sychugov et al.

While it is easy to evaluate the compliance of the machine project with a given functional purpose using objective numerical indicators, it is difficult to quantify such a complex and multifactorial concept as manufacturability. The lack of generally accepted methods for the numerical evaluation of manufacturability does not allow objectively resolving the inevitable conflicts of professional interests between the engineer as a supplier of design documentation (DD) and the technologist as its consumer [10, 11].

2 Mathematical Modeling Obviously, each set of part parameters (shape and size of surfaces, roughness, tolerances, design bases, workability coefficient of the material) has its own manufacturability coefficient [12]. The manufacturability is measured by the excess of the minimum permissible value, which is determined by the technological capabilities of production. Some parts, which are quite manufacturable for the factory with advanced equipment, may not be manufacturable or feasible to be produced at a plant with outdated equipment. According to GOST 14.201-83 “Ensuring the manufacturability of product design” [13], the main manufacturability indicators are labor intensity, production cost, and unification. Since labor intensity has monetary value, it is usually included in the production cost. Therefore, the proposed manufacturability evaluation includes two main components: technological cost C and unification, expressed by design kdu and technological ktu factors. Then 2 1 Nk ¼ kCk þ ðkdu þ ktu Þ 3 3

ð1Þ

Weight factors 2/3 and 1/3 show that, according to GOST 14.201-83 [13], production cost, labor intensity, and unification make an equal contribution to the manufacturability of the part. Let us consider a mathematical model for calculating the technological cost. The cost rate is calculated by the formula: kC k ¼

Cnom CT

ð2Þ

It represents the ratio of the actual technological cost to the value Cnom , which is either the technological cost of the similar component (previously manufactured part of the same size and design, for which the production cost is known), or the nominal technological cost of the “simplest” part, for the calculation of which we take a smooth shaft, with overall dimensions matching the dimensions of the considered parts, and with the surfaces having an average roughness Raav :

Comprehensive Evaluation of Shaft Manufacturability …

Raav ¼

n X Rai
li i¼1

n
l

375

ð3Þ

where Rai is roughness of ith treated surface; li is the length of ith surface; n is the number of treated surfaces; l is the overall shaft length. Technological cost is calculated by the formula [14]: mat þ T
kmat
ðWw þ Cmh Þ
ð1 þ K 0 =100Þ; CT ¼ Ctot

ð4Þ

and labor intensity T is calculated using empirical dependence [15]: T ¼ My


n X

Sai
Rabi
tic
kmat ;

ð5Þ

i¼1 mat where Ctot is the total cost of raw materials and resources, in rubles; Ww is the average hourly wage of a worker for the corresponding type of production, in rubles per hour; Cmh is the cost of machine hours, in rubles per hour; K 0 is average overhead costs, percentage; n is the number of design and technological elements (DTE); kmat is coefficient, considering the workability of the material; M is part mass, kg; Si is the area of the layer being cut off of the ith surface, mm2; Rai is the required roughness of the ith surface, lm; ti is ith surface tolerance, mm; n is the number of treated surfaces; a, b, c, y are empirical exponents, determined by regression analysis. Next, we consider the coefficients of the design and technological unification of such parts as a solid of revolution. The design unification coefficient kdu is calculated as

kdu ¼

n 1X kdu
di n i¼1 i

ð6Þ

where n is the number of structural elements; kdui is the design unification coefficient of the ith structural element of the part, depending on the shape and value of the coupling size (which affects the accuracy of part’s fitting in the assembly): for cylinders: 8 1; if Di ¼ 1:6N ; > > > > 0:75; if Di ¼ 1:25N ; < ð7Þ kdui ¼ 0:5; if Di ¼ 1:12N ; > N > > 0:25; if D ¼ 1:06 ; i > : 0; if none of the conditions is met, for cones:

376

A. Sychugov et al.

kdui

8 < 1; ¼ 0:5; : 0;

if 2ai 2 A1; if 2ai 2 A2; if none of the conditions is met,

ð8Þ

where ai is the angle at the vertex of the cone; A1; A2 are rows of preferred angles according to GOST 8593-81 [16]. The constants given (1.6; 1.25 …) are the rows of preferred numbers for the cylinder diameter according to GOST 8032-84 [17]. For structural elements of other types (e.g., shaped grooves, spherical surfaces), we analyze whether they are represented in the catalog of similar parts. If a similar element is found, then the coefficient kdui ¼ 1, otherwise—kdui ¼ 0. In [3] it is indicated that, from the standpoint of manufacturability, it is necessary to guarantee a certain ratio between the dimensional tolerance on the surface and its roughness, due to the impossibility of achieving and measuring a rigid dimensional tolerance on an excessively rough surface. Therefore, the following coefficient is introduced for each surface to be treated, considering the correctness of the roughness and tolerance area ratio: di ¼

n 1X dij ; n j¼1

ð9Þ



0; roughness exceeds 5% size tolerance; 1; otherwise According to the data collected from Tula region enterprises, it was revealed that one of the main reasons for DD refusal (45% of refusal cases occur at the technological preparation stage) is inconsistency in the assembly schemes. In some cases, due to the incorrect design of the assembly bases, the closing dimension tolerances turned out to be impossible to be produced [18], which required the revision of the design documentation. Consequently, the technological unification coefficient of parts should evaluate, first, the degree of compliance between the design and technological bases as it is the most important indicator of manufacturability. This coefficient cannot be calculated only with the data presented in the DD, since it does not indicate the technological bases. Therefore, the draft reveals a lot of assembly bases (structural elements with dimension tolerances) of part MDB . Further, on the basis of a suitable typical technological process, many hypothetical technological bases MTB are determined. The coefficient ktu is the ratio of the number of matched bases ðkmatch Þ to the number of technological ones ðkTB Þ: where dij ¼

ktu ¼

kmatch kTB

ð10Þ

When generating a hypothetical process for each operation, it is necessary to choose the standard equipment, if there is no additional information about the equipment of a particular production. The equipment depends on the content of the operation itself, the dimensions of the workpiece, the material properties,

Comprehensive Evaluation of Shaft Manufacturability …

377

the accuracy of the resulting size, and the serial production. Next, assembly base scheme is chosen for each machine. For solids of revolution, it depends mainly on the dimensions of the workpiece. The method of expert evaluation revealed the dependencies between the assembly scheme (in the chuck, in the chuck with steady rest, in the chuck with steady rest and in the center, in the chuck with the center, in the floating center, in the centers, in the centers with steady rest) and the length and diameter of the workpiece. The coordinates of the distribution center ðDci ; Lci Þ for ith assembly base scheme will be calculated as follows: D ci ¼

ki 1X dj ; ki j¼1

ð11Þ

L ci ¼

ki 1X Lj ; ki j¼1

ð12Þ

where i is the number of the base scheme; ki is the number of parts, processed using ith base scheme, given by experts; di ; li —diameter and length of jth part. The optimal base scheme of a part iopt with overall dimensions d; l is selected according to the condition of minimum difference of the distances of points in the phase field (D, L) for the selected material: iopt ¼ min i

qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ðDci  dÞ2 þ ðLci  lÞ2 :

ð13Þ

3 Results and Discussion To check model’s adequacy, the correlation coefficient was determined between expert estimates of manufacturability level of such parts as solids of revolution and estimates calculated by the proposed method [19]. The correlation coefficient was 0.84, which indicates the adequacy of the mathematical model. For effective management of manufacturability at the design stage, it is not enough just to evaluate it—it is necessary to give recommendations to the engineer, indicating the most technologically unfeasible parameters of the part [20–23]. During the operation, the level of manufacturability of the part can be calculated according to four parameters: consistency of bases; surface roughness; unification of structural elements (SE); and processing labor intensity. The results of the complex element-by-element evaluation of the above shown details are shown in Fig. 1a–c As a result, the coefficients of manufacturability Nk are 0.318, 0.238, and 0.175.

378

A. Sychugov et al.

Fig. 1 Result of a comprehensive feature-based manufacturability evaluation for the detail Acknowledgements The results of this research project are published with the financial support of Tula State University within the framework of the scientific project № NIR_2018_48.

References 1. Antsev VY, Inozemtsev AN, Troitsky DI (1995) Conceptual modeling of the subject area development of material standards in mechanical engineering. Autom Mach Syst Prod Robot 3:156–161 2. Grishin SA, Dorokhin NB, Inozemtsev AN, Troitsky DI (1997) Automation of design of topologically homogeneous products of machine-building production. Izv Tula State Univ Tech Sci 1:164–171 3. Amirov YD (1990) Manufacturability of product design. Handbook. Mashinostroenie, Moscow 4. Kalashyan AN, Kalyanov GN (2003) Structural business models: DFD-technology. Finansy i statistika, Moscow 5. Saaty T (1993) Decision making with the analytic hierarchy process. Translation from English. Radio i svyaz, Moscow 6. Lipuntsov YP (2003) Process management. Methods of enterprise management using information technology, Moscow 7. Meleshin SP (2001) Improving the management of the design and technological preparation of the production in mechanical engineering within the framework of the project management concept (based on the production of electric machines in Safonovsky Electric Machine Building Plant, OAO (OJSC)). Dissertation, Tula State University

Comprehensive Evaluation of Shaft Manufacturability …

379

8. Melnichenko VV (2004) Improving the quality management of technical preparation of equipment with program control. Dissertation, Tula State University 9. Barbash SM, Zalesov AK, Kozenko AV, Kosyuk LM, Levin EI (1972) Basic methodological principles for evaluating the quality of design engineering. Stan Qual 1:37–38 10. Novikova MV (2006) Manufacturability evaluation of engineering parts in the early stages of pre-production. Dissertation, Tula State University 11. Novikova MV, Troitsky DI (2006) Integrated design-technological model of the part as a means of evaluating the complexity of its manufacture. Vestnik komp’iuternykh i informatsionnykh tekhnologii (Her Comput Inf Technol) 1:22–25 12. Fedyukin VC, Durnev VD (2003) On the numerical evaluation of the quality of education. In: Shlenova YV, Azarov VN (eds) Quality. Innovation. Education. Proceedings of the first scientific conference, Moscow 13. GOST 14.201-83 (1983) Ensuring the manufacturability of product design. General requirements, Moscow 14. Averchenkov VI (1990) Formalization of the design and selection of advanced technologies, ensuring the required quality of products. Dissertation, Bryansk State Technical University 15. Volkhin KA (2004) Design documents and documentation requirements: student handbook for technical universities, Novosibirsk 16. GOST 8593-81 (1981) Normal conicity and the angles of the cones, Moscow 17. GOST 8032-84 (1984) Preferred numbers and series of preferred numbers, Moscow 18. Koganov IA, Nikiforov AP, Sotova BI, Gurlane MO (1998) Dimensional analysis of technological processes: student handbook, Tula 19. Hartman K (1977) Planning an experiment in the study of technological processes, Moscow 20. Popov ME, Popov AM (2003) Automated methods for improving the manufacturability of product designs in integrated CAD. Vestn Mashinostroeniya 10:48–53 21. Samokhvalov VN (2000) Designing technological processes for machining of parts: methodical instructions for course paper in the discipline “Engineering technology and production of handling, road and construction machines” for students of the course 170900, Samara 22. Suslov AG, Dalsky AM (2002) Scientific basis of engineering technology, Moscow 23. Fedorov YN, Yamnikov AS, Artamonov VD, Malikov AA (2004) Technological basis for the design of machining operations: study book, Tula

Simulation of Electrical Discharge Machining of Micro-holes T. D. Nguyen, V. M. Volgin and V. V. Lyubimov

Abstract Electrical discharge machining (EDM) is an important and cost-effective manufacturing process for machining electrically conductive materials irrespective of their hardness in various industries including automotive, aerospace, biomedical, and semiconductor. The applications of EDM are not just limited to the machining of hard materials, but also cover the production of difficult-to-make microstructures for micro-molds, fuel injection nozzles, spinneret holes, etc. However, although EDM is widely used, scientific knowledge of the process is still limited. The complex nature of the process involves simultaneous interaction of thermal, mechanical, chemical, and electrical phenomena, which makes process modeling very difficult. New contribution to the simulation and modeling of the formation of micro-holes during EDM process is presented in this paper. Mathematical modeling of the evolution of the machined surface at electrical discharge machining using ablation of the workpiece material was carried out. In contrast to the known models of EDM, the proposed model allows to determine the location of the center of the discharge channel, both on the surface of the workpiece and on the surface of the tool-electrode, which makes it possible to predict the wear of the tool-electrode during machining. The obtained results make it possible to predict the shape, size, and surface roughness of the micro-hole and the performance of EDM.







Keywords Evolution of machined surface Discharge Modeling electrical discharge machining Erosive crater Roughness Temperature










1 Introduction Electrical discharge machining (EDM) is one of the effective methods of processing of difficult-to-process materials, obtaining surfaces of complex shape, holes with different cross sections [1–3]. In contrast to mechanical machining, when using T. D. Nguyen
V. M. Volgin (&)
V. V. Lyubimov Tula State University, Lenin Pr. 92, Tula 300012, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_41

381

382

T. D. Nguyen et al.

EDM, the removal of material from the machined surface is carried out without the force effect on the workpiece, which allows processing low-rigid workpieces and forming micro-elements with a sufficiently large aspect ratio, for example, micro-holes. Removal of material from the workpiece surface is carried out under the influence of electrical discharges between the surfaces of the workpiece and the tool-electrode, leading to local heating, melting, and evaporation and, as a consequence, to the formation of an erosion crater, which is close in shape to a spherical segment with a radius R, height h, and diameter of the base DC (Fig. 1a). In contrast to the ideal crater a real single crater has a circular collar height hb (Fig. 1b). The volume of a single crater depends on the discharge energy and is very small, so the shape and size of the machined surface are obtained as a result of a very large number of discharges. In this case, the single craters are repeatedly overlapped, forming the final shape and size of the workpiece surface, as well as its roughness (Fig. 2). In the framework of an ideal model of the formation of the micro-geometry of the machined surface, that is with the same size of all the craters and their regular arrangement with a center-to-center distance is L, the surface finish can be determined using the relationship Rz ¼ L2 =ð8RÞ: In fact, the shape and size of the individual craters differ from each other, since the real surface always has geometric or energy inhomogeneities, and the discharges on the workpiece surface are not distributed regularly. This leads to the need to take into account the geometric and energy inhomogeneities of the workpiece when modeling of EDM. Despite significant advances in the modeling of a single discharge with the formation of a single crater, as well as in the formation of the relief of the machined surface as a result of a sequence of several discharges [4–16], there are a large number of issues related to the distribution of discharges on the workpiece surface and the possibility of taking into account the wear of the electrode-tool. In most of the works devoted to the modeling of EDM, only thermal processes in the

Fig. 1 Schemes and parameters of ideal and real single craters

Fig. 2 An ideal model of micro-geometry formation by superposition of craters

Simulation of Electrical Discharge Machining of Micro-holes

383

workpiece with a flat surface at a single discharge are considered, and the shape and size of the crater are determined by the isotherm corresponding to the melting temperature [6–10]. Recently, there was published a number of papers [10–17], which simulates the process of electrical discharge machining of several successive discharges. In this case, the surface of the tool-electrode was taken flat, and the location of the next discharge was set at a point on the surface of the workpiece, in which the electric field is maximum. Along with certain advantages, this approach has a number of drawbacks, since it requires a sufficiently large amount of calculations to calculate the electric field in the interelectrode space and does not allow to determine the position of the discharge channel on the surface of the tool-electrode, which is required to know for the calculation of thermal processes in the tool-electrode in order to predict its wear. This work is devoted to the further development of the previously proposed modeling method [17] in order to provide the possibility of predicting the shape and size of the erosion craters formed on the workpiece surface and the tool-electrode.

2 Mathematical Model To simulate the process of forming the machined surface during EDM, it is necessary to solve the following problems: • develop a mathematical model for determining the location of the next discharge; • develop a model of thermal processes for a single discharge; • develop a model of metal removal from the workpiece surface. Analysis of literature shows that the greatest electric field intensity is formed in the region of the minimum distance between the electrodes. The location of the next discharge was determined by calculating the distances between the electrodes (Fig. 3) using the following relation [18]: dAB ¼

qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ðx  Ab ðxÞÞ2 þ ðy  Ab ð yÞÞ2 ;

ð1Þ

rdAB ðxÞ ¼ 0;

ð2Þ

rdAB ðyÞ ¼ 0;

ð3Þ

where dAB is the distance from an arbitrary point of the tool-electrode (area B) to the nearest point of the workpiece (area A) (see Fig. 3); Ab is the surface of the workpiece (boundary of area A); rdAB is the gradient of the distance between the points of the tool-electrode and the workpiece. The differential equation of unsteady thermal conductivity is used to simulate thermal processes in the workpiece during EDM [19]:

384

T. D. Nguyen et al.

Fig. 3 Scheme for determining the minimum distance between the electrodes

 @ qCp T ¼ divðkgradT Þ; @t

ð4Þ

where T is the temperature; q is the density of the workpiece material; k is the coefficient of thermal conductivity; Cp is the specific heat capacity. The following ratio was used to determine the radius of the plasma channel: 0:44 Rp ¼ 2:04  103 I 0:43 ton

ð5Þ

where I is the discharge current; ton is the duration of the discharge. The distribution of heat flow over the cross section of the discharge channel, the center of which is located at the point x = x0, was set in this form:   ! x  x0 2 qðxÞ ¼ q0 exp 4:5 ; Rp

ð6Þ

where x is the coordinate describing the distance from the center of the channel of the electric discharge; q0 is the maximum heat flux on the axis of the channel of the discharge which can be defined with use of the following relation: q0 ¼

4:55gE ; pR2p ton

ð7Þ

where E ¼ UIton is the pulse energy; U is the breakdown voltage; g is the proportion of the discharge energy absorbed by the workpiece. In the process of EDM, material removal is carried out as a result of the sequence of a large number of electrical discharges. At the same time, during each discharge

Simulation of Electrical Discharge Machining of Micro-holes

385

the temperature increases monotonically, and during each pause, the temperature decreases monotonically. Therefore, the use of the melting isotherm to determine the boundary of the crater in the simulation of multi-pulse EDM cannot be used. To overcome this problem, an approach based on the introduction of an auxiliary function, the value of which changes monotonically over time, was used: df ¼ 1ðt  t Þ dt

ð8Þ

where 1ðt  t Þ is the step function, the value of which is equal to 0 for negative values of the argument and is equal to 1 for positive values of the argument; t ¼ t  ti is the time counted from the beginning of the ith pulse; ti ¼ ði  1Þtper is the time of the beginning of the ith pulse; tper is the pulse repetition period; t is the characteristic time of the beginning of the removal of the material from the workpiece in the ith pulse, counted from the beginning of the ith pulse. The rate of removal of materials from the electrode can be calculated by the following relation [20]: v¼

q ; qHs

ð9Þ

where v is the rate of ablation of the workpiece material; q is heat flux due to ablation of the material; Hs is the heat of sublimation. The heat flux of ablation can be determined using the following relation [12, 13]: q ¼ hðTm  TÞ;

ð10Þ

where Tm is the melting point; h is the coefficient, which is zero at T\Tm and increases linearly at T [ Tm .

3 Results and Discussion The numerical solution of the mathematical model was carried out by the finite element method. In the simulation, the following parameters were taken: copper workpiece having a rectangular cross section with dimensions of 120  50 µm; U = 25 V; I = 2.34 A; ton ¼ 5 ls; toff ¼ 4 ls; tool-electrode with diameter d = 35 µm. In Fig. 4, the results of modeling the formation of a single crater on the tool-electrode and on the workpiece when removing the material are presented. In general, it is necessary to consider the model of the formation of craters on the surfaces of both electrodes, but for the convenience of modeling, the formation of craters and the evolution of only the workpiece surface are further considered.

386

T. D. Nguyen et al.

Fig. 4 Formation of single craters on the workpiece and tool-electrode

The position of the first discharge and all subsequent discharges was set at the point with the smallest distance from the tool-electrode to the workpiece surface. In the case where there are several points on the workpiece surface located at a minimum distance from the tool-electrode, the location of the next discharge was chosen from the condition of the minimum distance to the point of the previous discharge. In Fig. 5, the distribution of the distance from the tool-electrode to the workpiece for different forms of the end surface of the tool-electrode.

Fig. 5 Distribution of the distance from the tool-electrode to the workpiece at different shapes of the end surface tool-electrode: a curved; b conical; c complex shape

Simulation of Electrical Discharge Machining of Micro-holes

387

Figure 6 shows the results of the formation of the workpiece surface with multi-pulse EDM. From the obtained results, the effect of the aftereffect is seen— the crater after the second discharge differs from the crater obtained after the first discharge. Figure 7 shows the difference in the workpiece surface obtained during EDM with movable and motionless tool-electrodes. Figure 8 shows the evolution of the workpiece surface as a result of multi-discharge machining. The nesting effect is clearly visible—after the first discharge, several subsequent discharges pass through the outer boundary of the previous crater. Further, the process continues at points with a minimum distance between the electrodes. In Fig. 9, the results of modeling of EDM of micro-holes in various forms of the end surface of the tool-electrode are presented. From the obtained results, it is clearly seen that the shape of the working surface of the tool-electrode is copied on the workpiece surface during EDM. This indicates the possibility of using the simulation results in predicting the shape of the surface during electrical discharge machining.

Fig. 6 Formation of a single crater: a after the first discharge; b after the ith discharge

Fig. 7 Formation of the workpiece surface during EDM after 20 pulses: a fixed tool-electrode; b moving tool-electrode

388

T. D. Nguyen et al.

Fig. 8 Evolution of the workpiece surface during EDM with a movable tool-electrode (V = 0.045 m/s): a before machining; b after the first discharge; c after 5 discharges; d after 10 discharges; e after 20 discharges; f after 45 discharges

Fig. 9 Results of modeling of EDM of micro-holes in various forms of the end surface of the tool-electrode: a straight (flat); b complex surface; c curve; d conical

Simulation of Electrical Discharge Machining of Micro-holes

389

4 Conclusion On the basis of the results of the simulation, there were obtained the dependences of the position of the craters from the distance between the electrodes; the sizes of the erosion craters give an idea of the amount of material removed from the electrode in a single discharge, allow to assess the surface roughness after machining, and predict the shape of the machined surface. A series of experimental studies is necessary to clarify the obtained simulation results. Acknowledgements The study was funded by RFBR and Tula region according to the research project No. 19-48-710009 p_a.

References 1. Eliseev LS, Boitsov AG, Krymov VV, Khvorostukhin LA (2003) Tekhnologiya proizvodstva aviacionnykh gazoturbinnykh dvigatelei (Technology of production of aviation gas turbine engines). Mashinostroenie, Moscow 2. Hocheng H, Tsai HY (2013) Advanced analysis of nontraditional machining. Springer, New York 3. Kibria G, Jahan MP, Bhattacharyya B (2019) Micro-electrical discharge machining processes. Springer, Singapore 4. Revaz B, Witz G, Flukiger R (2005) Properties of the plasma channel in liquid discharges inferred from cathode local temperature measurements. J Appl Phys 98:113305 5. Kunieda M, Lauwers B, Rajurkar KP, Schumacher BM (2005) Advancing EDM through fundamental insight into the process. CIRP Ann-Manuf Techn 54(2):64–87 6. Joshi SN, Pande SS (2010) Thermo-physical modeling of die-sinking EDM process. J Manuf Process 12:45–56 7. Hoang KT, Gopalan SK, Yang SH (2015) Study of energy distribution to electrodes in a micro-EDM process by utilizing the electro-thermal model of single discharges. J Mech Sci Technol 29(1):349–356 8. Assarzadeh S, Ghoreishi M (2017) Prediction of root mean square surface roughness in low discharge energy die-sinking EDM process considering the effects of successive discharges and plasma flushing efficiency. J Manuf Process 30:502–515 9. Jeong YH, Min BK (2007) Geometry prediction of EDM-drilled holes and tool electrode shapes of micro-EDM process using simulation. Int J Mach Tool Manu 47(12):1817–1826 10. Tan PC, Yeo SH (2008) Modelling of overlapping craters in micro-electrical discharge machining. J Phys D Appl Phys 41(20):205302 11. Izquierdo B, Sanchez JA, Plaza S, Pombo I, Ortega N (2009) A numerical model of the EDM process considering the effect of multiple discharges. Int J Mach Tool Manu 49(3):220–229 12. Izquierdo B, Sánchez JA, Ortega N, Plaza S, Pombo I (2011) Insight into fundamental aspects of the EDM process using multidischarge numerical simulation. Int J Adv Manuf Tech 52(1–4): 195–206 13. Hinduja S, Kunieda M (2013) Modelling of ECM and EDM processes. CIRP Ann-Manuf Techn 62(2):775–797 14. Somashekhar KP, Panda S, Mathew J, Ramachandran N (2015) Numerical simulation of micro-EDM model with multi-spark. Int J Adv Manuf Tech 76(1–4):83–90

390

T. D. Nguyen et al.

15. Huang H, Zhang Z, Ming W, Xu Z, Zhang Y (2017) A novel numerical predicting method of electric discharge machining process based on specific discharge energy. Int J Adv Manuf Tech 88(1):409–424 16. Feng X, Wong YS, Hong GS (2016) Characterization and geometric modeling of single and overlapping craters in micro-EDM. Mach Sci Technol 20(1):79–98 17. Lyubimov VV, Volgin VM, Gnidina IV, Salomatnikov MS (2018) Formation of the workpiece shape and surface finish during electrical discharge machining. Procedia CIRP 68:319–324 18. Frei W (2017) How to compute distances between objects in COMSOL multiphysics: Comsol Blog. https://www.comsol.com/blogs/how-to-compute-distances-between-objects-in-comsolmultiphysics. Accessed 3 Sept 2018 19. Weingärtnera E, Kustera F, Wegenera K (2012) Modeling and simulation of electrical discharge machining. Procedia CIRP 2:74–78 20. Frei W (2016) Modeling thermal ablation for material removal: Comsol Blog. https://www. comsol.com/blogs/modeling-thermal-ablation-for-material-removal. Accessed 3 Sept 2018

Investigation of Kinematic–Geometric Characteristics of Electrochemical Machining V. V. Lyubimov, V. M. Volgin and V. P. Krasilnikov

Abstract Electrochemical machining (ECM) is a non-traditional method of machining, based on the anodic dissolution of the workpiece, which allows to obtain surfaces of a complex shape (turbine blades, dies and molds, etc.); to produce through and blind holes of different or variable cross-sections, including the curved axis; to remove a defective surface layer from the workpiece after electrical discharge machining or cutting (ECM without shaping) in order to ensure a specified surface roughness; to remove burrs and round sharp edges. Distinguishing features of electrochemical shaping include a narrow interelectrode gap, an electrolyte flow in the space between electrodes, a high localization of dissolution in a specified area of the workpiece, and a high dissolution rate. The efficiency of ECM depends significantly on the shape and size of the electrode-tool, as well as the trajectory and speed of its movement. This article is devoted to the investigation of the combination of kinematic and geometric ECM characteristics. It is shown that the reduction of interelectrode gap under traditional modes leads to a significant increase in the total current and the difficulties of ECM process control. The possibilities of local machining with point and linear electrode-tools are analyzed. Keywords Electrochemical machining Displacement velocity


Productivity
Electrode-tool


1 Introduction Electrochemical machining (ECM) and especially electrochemical micromachining have become more popular with recent advancements to fulfill the needs of the fabrication of macro- and microcomponents [1–12]. With rapid developments in the fields of automotive, aerospace, electronics, optics, medical devices, and much more, this process finds wide applications in the machining of titanium and titanium V. V. Lyubimov (&)
V. M. Volgin
V. P. Krasilnikov Tula State University, 92, Lenin Ave., Tula 300012, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_42

391

392

V. V. Lyubimov et al.

alloys, super alloys, and stainless steel structures which are difficult to machine by conventional machining processes. ECM provides advantages in the shaping of complex features with no thermal stress, burr formation, and tool wear and can be implemented on metals regardless of their hardness [13–22]. It is known from the theory of mechanical engineering technology that the following kinematic and geometric schemes of forming can be carried out: • Surface treatment (one-time shaping) (Fig. 1a); • Processing on the line (Fig. 1b); • Processing at the point (Fig. 1c). In most cases, ECM is used to perform copy-and-flash operations, that is, surface treatment (Fig. 1a). Thus, the most important advantage of the ECM is the possibility of forming over the entire surface to ensure high performance of the process. The required geometry of the resulting surface is provided by copying the geometric shape of the electrode-tool surface. In this case, there is a simple kinematics of the machine—rectilinear displacement of the electrode-tool along the z-axis as the anode dissolution of the workpiece material. However, in this case, there are difficulties in controlling the process of shaping and ensuring high accuracy. Increasing accuracy is provided in the transition to the minimum possible interelectrode gaps (IEG). The accuracy of machining is associated with the magnitude of the IEG: D ¼ ks

ð1Þ

where D is the local value of the machining error; k is the coefficient of proportionality; s is the local value of the interelectrode gap. However, the decrease of the IEG is associated with a number of difficulties in the implementation of the ECM process: • Increasing the probability of short circuits, that is, reducing the reliability of the processing process;

Fig. 1 Kinematic–geometric schemes of ECM: a ECM on the surface; b ECM along the line; c ECM at the point; (1) electrode-tool; (2) workpiece; Vx is velocity in the direction of the axis Ox; Vy is velocity in the direction of the axis Oy; Vz is velocity in the direction of the axis Oz

Investigation of Kinematic–Geometric Characteristics …

393

• Deterioration of electrolyte flow conditions in the IEG, which is associated with a significant increase of its hydraulic resistance; • Increasing the current density and the total current requires the use of very high power supply (150–400 kW). This article is devoted to the investigation of the combination of kinematic and geometric ECM characteristics.

2 The Comparative Investigation of the Electrochemical Machining Schemes The necessity of preserving a time of the properties of the interelectrode environment with a decrease of the IEG led to the need to move from a continuous process to a discrete or pulse-loop process [5]. In this regard, were developed: • scheme of treatment with vibrating electrode-tool (Fig. 2c); • loop processing (loop duration is 15–35 s) (Fig. 2a); • pulse-loop processing (Fig. 2b). Despite the transition to processing at small IEG, loop and pulse-loop schemes are characterized by a decrease in performance: Vap ¼ 0:02k1 k2 k3 j; ½mm/min;

ð2Þ

where k1 is the duty ratio of the voltage pulses; k2 is the duty ratio of the voltage packages; k3 is the current pulse shape coefficient; j is the current density. Practical application of pulse-loop processing modes showed that k1 and k2 vary in the ranges: k1 = 0.2–0.5; k2 = 0.3–0.6; k3 = 0.7. Thus, there is a decrease in productivity by 6–22 times in comparison with the continuous process of ECM. Then the rate of anodic dissolution at s0 = 50 µm, j = 200 A/cm2, Vap = 0.3– 0.4 mm/min. Further reduction of the IEG becomes impossible due to the decrease in the reliability of the ECM process (a significant increase in the probability of short circuits, deterioration of the conditions for washing the interelectrode gap). Therefore, further improvement of the ECM is associated with the justification of the process conditions of processing with non-profiled electrode-tools in the local areas of the treated surface (line processing (Fig. 1b) or at a point (Fig. 1c). It is possible to significantly reduce the interelectrode gap to s = 1.0–20 lm by the significant improvement in the conditions of the evacuation of the anode dissolution products from the interelectrode gap. In this case, the current density up to 100 A/mm2 will be reached. Thus, even with the preservation of interelectrode gaps similar to the pulse-loop processing without loss of productivity, the ratio of the treated area to the end area of the electrode-tool is possible as:

394

V. V. Lyubimov et al.

Fig. 2 Timeline of various ECM schemes: a loop processing; b pulse-loop processing (a scheme with inlet/outlet of the electrode-tool during the supply voltage pulses); c scheme of ECM with vibrating electrode-tool; T is the duration of the voltage pulse during loop ECM (15–35 s); tP is the duration in pulse-loop; s0 is the interelectrode gap; U is the voltage technology; J is the current in the electrode gap; y is the displacement of the electrode-tool

Sw ¼ 622; SET

ð3Þ

where Sw is the area of the treated surface; SET is the working area of the electrode-tool. This ratio can be achieved without loss of performance values 15–55 by reducing of the interelectrode gap to the minimum values (1.0–20 µm). A high-frequency pulse voltage with a voltage pulse frequency of up to 20 MHz is recommended for local shaping.

Investigation of Kinematic–Geometric Characteristics …

395

The most important task of designing the operations of the ECM by unprofiled tool is to select the type and size of the local electrode-tool and the conditions of its displacement. This takes into account: • dimensions and geometric parameters of the treated surface (processing area, radii, minimum dimensions of the geometric elements of the treated surface, the angles of inclination of the processed elements) • the expected path of the electrode-tool; • requirements for accuracy and surface quality. The most used and proposed electrode-tool types by the authors are given in Table 1. The displacements of the local electrode-tool are selected depending on the shape of the treated surface: • • • •

path by surface (Fig. 3a); surface scanning (Fig. 3b); layer-by-layer removal of allowance (Fig. 3c); matrix removal of allowance (Fig. 3d).

Local electrode-tools have limited dimensions which imply restrictions on the limit value of the total current flow through the electrode. Since the total current is

Table 1 Types of local electrode-tools No.

Type of local electrode-tool

Parameters

Notes

1

The cylindrical tool without insulation

d  10– 100 µm

2

–

3

The cylindrical tool with insulation of side surface The conical tool with straight cone

4

The conical tool with reverse cone

–

5

The tool with the spherical working part

–

6

The stage tool

d  10– 20 µm

7

The tube tool

8

The tool with the linear treatment area

–

–

396

V. V. Lyubimov et al.

Fig. 3 Displacement path of the local electrode-tool: a path scheme; b scanning scheme; c one-time or layer-by-layer removal of the allowance; d matrix removal of the allowance

Fig. 4 Dependence of current density on the IEG value: I is unacceptable area of the current densities; II is area of acceptable current densities

determined depending on the area of the working part of the local electrode, depending on the processing conditions, the choice of the interelectrode gap and the permissible current densities can be made (Fig. 4). The choice of the velocity of the electrode-tool displacement is poorly justified at present. It is known [13, 14, 17, 18, 21] that velocity of the electrode-tool displacement ranges from 0.05 to 8 µm/min in the micromachining. The velocity of the local electrode ranges from 50 to 700 mm/min in the machining of cavities with large dimensions (of order of a few millimeters).

3 Experimental Study and Discussion The experimental study of the processing of the annular electrode-tool was made (Fig. 5a). The resulting surface is shown in Fig. 5b. The electrode-tool was made of brass brand LA77–2 GOST 17711-80 with a diameter of 6 mm with a wall thickness of 0.5 mm, in a sample of steel grade X12N10T. ECM modes: The amplitude of the pulse voltage was 12 V; pulse duration was 1 µs; pulse repetition rate was 250 kHz; working fluid was 10% aqueous solution of NaNO3; the displacement velocity of the tool was 700 mm/ min.

Investigation of Kinematic–Geometric Characteristics …

397

Fig. 5 Electrode-tool and the resulting surface: a photograph of the end part of the electrode-tool; b photograph of the resulting cavity

4 Conclusion The investigation of the kinematic and geometric characteristics technological ECM schemes is carried out. It is established that with the development of the method of ECM, there is a tendency of developing processing schemes by unprofiled (local) electrode-tool. The rational path of the non-profiled electrode-tool displacement is offered depending on geometrical parameters of the processed surface. The possible technological equipment and modes of machining by point or linear electrode-tool are analyzed.

References 1. El-Hofy HAG (2005) Advanced machining processes: nontraditional and hybrid machining processes. McGraw Hill, New York 2. Kibra G, Bhattacharyya B, Davim JP (eds) (2017) Non-traditional micromachining processes: fundamentals and applications. Springer, Berlin 3. Rathod V, Doloi B, Bhattacharyya B (2017) Electrochemical Micromachining (EMM): Fundamentals and Applications. Non-traditional Micromachining Processes. Springer, Cham, pp 275–335 4. Skoczypiec S (2018) Electrochemical methods of micropart’s manufacturing. Micro and precision manufacturing. Springer, Cham, pp 21–40 5. Davydov AD, Volgin VM, Lyubimov VV (2004) Electrochemical machining of metals: fundamentals of electrochemical shaping. Russ J Electrochem 40(12):1230–1265 6. Rajurkar KP, Sundaram MM, Malshe AP (2013) Review of electrochemical and electrodischarge machining. Procedia CIRP 6:13–26 7. Spieser A, Ivanov A (2013) Recent developments and research challenges in electrochemical micromachining (µECM). Int J Adv Manuf Tech 69(1–4):563–581 8. Raja K, Ravikumar R (2016) A review on electrochemical machining processes. Int J Appl Eng Res 11(4):2354–2355

398

V. V. Lyubimov et al.

9. Davydov AD, Kabanova TB, Volgin VM (2017) Electrochemical machining of titanium. Review. Russ J Electrochem 53(9):941–965 10. Zhan D, Han L, Zhang J et al (2017) Electrochemical micro/nano-machining: principles and practices. Chem Soc Rev 46(5):1526–1544 11. Saxena KK, Qian J, Reynaerts D (2018) A review on process capabilities of electrochemical micromachining and its hybrid variants. Int J Mach Tool Manu 127:28–56 12. Leese R, Ivanov A (2018) Electrochemical micromachining: review of factors affecting the process applicability in micro-manufacturing. P I Mech Eng B-J Eng 232(2):195–207 13. Zhang Z, Zhu D, Qu N et al (2007) Theoretical and experimental investigation on electrochemical micromachining. Microsyst Technol 13:607–612 14. Mithu MAH, Fantoni G, Ciamhi J (2011) The effect of high frequency and duty cycle in electrochemical microdrilling. Int J Adv Manuf Technol 55:921–933 15. Volgin VM, Kabanova TB, Davydov AD (2015) Modeling of through-mask electrochemical micromachining. J Appl Electrochem 45(7):679–688 16. Xiaolong F, Pengfei Z, Yongbin Z et al (2016) Enhancement of performance of wire electrochemical micromachining using a rotary helical electrode. J Mater Process Tech 227:129–137 17. Ghoshal B, Bhattacharyya B (2016) Electrochemical micromachining of microchannel using optimum scan feed rate. J Manuf Process 23:258–268 18. Lyubimov V, Volgin V, Gnidina I et al (2016) The scanning dimensional microelectrochemical machining with the ultra-small interelectrode gap. Procedia CIRP 55:89–94 19. Volgin VM, Lyubimov VV, Davydov AD (2016) Modeling and numerical simulation of electrochemical micromachining. Chem Eng Sci 140:252–260 20. Chen W, Han F, Wang J (2018) Influence of pulse waveform on machining accuracy in electrochemical machining. Int J Adv Manuf Tech 96(1–4):1367–1375 21. Guo C, Qian J, Reynaerts D (2018) Deterministic removal strategy for machine vision assisted scanning micro electrochemical flow cell. J Manuf Process 34:167–178 22. Wang Y, Zeng Y, Zhang W (2019) Improving the machining efficiency of electrochemical micromachining with oscillating workpiece. Int J Adv Manuf Tech https://doi.org/10.1007/ s00170-019-03379-w

Model for Numerical Simulation of Temperature Field and Bead Profile in Hybrid Laser-Arc Welding of T-Joint S. Ivanov, E. Valdaytseva and I. Udin

Abstract The development and implementation of automated and robotic welding production lines is an urgent task of the last decade. The wide application of hybrid laser-arc welding (HLAW) is difficult due to sensitivity to changes in a large number of process parameters, as well as fluctuations in gap between welded edges and to the edge preparation quality. The need to solve these problems has led to the development of methods and algorithms for the numerical simulation of the HLAW. The developed mathematical model is able to predict effects of process parameters and heat flux distribution of the hybrid heat source on the formation of the surface of weld bead and temperature field in welded joint. The model is divided into the following sub-models: wire melting; heat conduction in the molten pool and the substrate; formation of the surface of the weld. Comparison of the calculated and experimental data of melting rate of the electrode wire and the formation of welded joints during HLAW of T-joint showed good agreement. The effect of fluctuations in the process parameters on the quality of welded joints was established by simulation. Keywords Laser-arc welding Temperature field T-joint





Simulation
Heat conduction
Free surface


1 Introduction The development and implementation of automated and robotic assembly-welding production lines is an urgent task of the last decade. The wide application of hybrid laser-arc welding (HLAW) is difficult due to sensitivity to changes in a large S. Ivanov (&)
E. Valdaytseva Saint Petersburg State Marine Technical University, 3, Lotsmanskaya St., St. Petersburg 190121, Russia e-mail: [email protected] I. Udin Peter the Great St. Petersburg Polytechnic University, 29, Polytechnic St., St. Petersburg 195251, Russia © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_43

399

400

S. Ivanov et al.

number of process parameters [1–3], as well as fluctuations in gap between welded edges [4] and to the edge preparation quality [5]. The need to solve these problems has led to the development of methods and algorithms for the numerical simulation of the HLAW and adaptive on-line control, which allow assessing the effects of process parameters and external environment on product quality. There are a large number of algorithms and approaches to solving the problem of adaptive control [6]. More difficult is the problem of optimizing the process parameters [7]. In order to obtain welded joints that meet the requirements, it is necessary to determine the optimal combination of process parameters. It is possible to find a solution to this problem by carrying out a set of experimental trials, but this approach will require a significant investment of time and material resources. Numerical (computer) simulation can significantly reduce the process of technological adaptation of HLAW. There are a large number of models of HLAW describing in detail the nature of various physical phenomena occurring during welding. Often such models require powerful computers and the assignment of a large number of material properties that are often unknown [8, 9]. On the other hand, analytical models make it possible to fast determine the effects of technological parameters on the temperature field only in butt joints [10, 11]. The solution of corresponding problem for HLAW of T-joints requires a numerical solution of heat conduction problem. The aim of the present work is to develop a model for the numerical simulation of the HLAW of T-joints. Hereby, the precise and fast prediction of the bead profile and the temperature field during the welding is the focus of this investigation.

2 Mathematical Model 2.1

Model Description

Developed mathematical model is able to predict the effects of process parameters and heat flux distribution of the hybrid heat source on the formation of the surface of weld bead and temperature field in welded joint. The model is divided into the following sub-models: (1) wire melting; (2) heat conduction in the molten pool and the substrate; (3) formation of the surface of the weld. The flowchart of the model and solution domain is shown in Fig. 1. The model of HLAW takes into account main process parameters (beam power and radius, arc current, wire feed rate, geometry of the welded joint, position of heat sources, etc.) and allows to predict the size and shape of the weld bead, the width of the heat-affected zone, the cooling rates, etc. After input of the process parameters and the material properties, the subroutine for calculating the melting rate of electrode wire is implemented. Next, the finite-difference mesh is generated according to the previously specified sizes of the computational domain and the values of the grid size. At the next step, the temperature field and the shape of the weld bead are calculated iteratively. Boundaries of the weld pool are determined

Model for Numerical Simulation of Temperature Field and Bead …

401

Fig. 1 a Flowchart of the multi-physics simulation of HLAW; b solution domain

using calculated temperature field. Next, the optimization problem of determining the shape of the weld bead is solved, which consists in selecting a combination of boundary angles and a Lagrange multiplier that satisfy the boundary conditions and mass balance. At the next step, the dimensions of the calculated melt pool are compared with obtained at the previous iteration, and if the difference does not meet the specified accuracy, a new iteration is implemented. The algorithm was implemented using MATLAB. The solution time depends on the size of the solution domain and approximately equals to 200–300 s using a simple office computer.

2.2

Electrode Melting Problem

During gas metal arc welding stick-out of the electrode is heated by the arc current. The end of the wire is additionally heated by the arc, which leads to its melting. In order to melt the electrode and transfer the liquid droplets into the weld pool, the following energy balance must be performed at the end of the electrode wire: HM ¼ HJ þ HA ;

ð1Þ

where HM is the enthalpy of the droplet; HJ is the Joule heating; HA is the anode heating. The heat content caused by the resistance could be presented as the following: HJ ¼ qL

Le j2 b v

ð2Þ

402

S. Ivanov et al.

where qL is the resistivity of metal at the end of the electrode; b is the constant; Le is the stick-out; v is the wire feed rate. According to [12–14], the enthalpy of droplets can be estimated as a half-sum of the enthalpies of melting and evaporation of the wire metal. For mild steel droplets, this value equals to 13.1 J mm−3, which corresponds to an average temperature of 2275 °C. The energy emitted at the anode is proportional to the anode voltage drop and welding current [15, 16]: HA ¼

/I /j ¼ ; Sw v m v m

ð3Þ

where / is the effective melting potential at anode; I is the arc current; Sw is the cross-sectional area of electrode wire; vm is the wire melting rate; j is the current density. After manipulations, the desired analytical dependence of the melting rate of the electrode wire is obtained: vm ¼

 1
/j þ qL Le j2 HM þ b

ð4Þ

Note that if the welding current varies in time according to a given law I ¼ IðtÞ, then in Eq. (4), the instantaneous value of the current density should be taken into account.

2.3

Heat Conduction Problem

In order to improve the stability and the computational time of the model, the following assumptions are made: • non-stationary phenomena at the beginning and the end of a weld are neglected • vaporization and radiation are not considered • thermophysical properties are known functions of the temperature. Under the assumptions made, the formulation of the nonlinear quasi-stationary heat conduction problem in Cartesian coordinates (Fig. 2) becomes:       @ @T @ @T @ @T @T k k k þ q3 ¼ 0 þ þ þ vcq @x @x @y @y @z @z @x

ð5Þ

where k is the thermal conductivity, cq is the volumetric heat capacity, v is the welding speed and q3 is the volumetric heat source power. The boundary conditions at the top surface of the calculation domain are the following:

Model for Numerical Simulation of Temperature Field and Bead …

403

Fig. 2 Schematic of welded joint and reference frame

k

@T ¼ q2A ðx; yÞ; @z

ð6Þ

where q2A ðx; yÞ is the heat flux of hybrid heat source. The power density distribution in the heating spot during arc welding depends on the composition of the shielding gas, process parameters (tilt angle, electrode diameter, arc length) and can be described as a combination of a normally distributed elliptical source, describing the distribution of power density in the cathode spot and a normally distributed circular source, describing the distribution of the power density of droplets of molten filler metal: " #  2  ga qa sin b ðx sin bÞ2 þ y2 gd qd x þ y2 q2A ðx; yÞ ¼ exp  exp  ; þ pra2 ra2 prd2 rd2

ð7Þ

where ga is the heat efficiency of cathode; qa is the power of cathode spot; ra is the effective radius of cathode spot; b is the tilt angle; gd is the heat efficiency of drops; qd is the power of drops; rd is the effective radius of droplets. On the side faces of the computational domain, the first-kind boundary conditions are specified, i.e., spatial temperature distribution. This approach helps significantly reduce the size of the solution domain without reducing the accuracy of determining the boundaries of the molten pool and the cooling rates in the heat-affected zone. The boundary conditions were calculated according to known analytical solutions [17]. A normally distributed power of laser beam has a much smaller effective radius than an arc. A keyhole is formed during welding due to high power density of laser beam. Therefore, the laser beam is best described by a volume heat source, distributed arbitrarily over the thickness (z-axis) of the plate and normally in the plane of the plate (z = const):

404

S. Ivanov et al.

q3 ðr; zÞ ¼ q1 ðzÞf2 ðrÞ

ð8Þ

where f2 ðrÞ ¼ ðk=pÞ expðkr 2 Þ. The function f2 ðrÞ can be understood as a normally circular source of unit power at an arbitrary depth z. The effective power of the source q (W) is related to the distribution of the volume power density (W m−3) by the expression: Zh Z1

Zh q3 ðr; zÞ2prdrdz ¼

0

2.4

0

Z1 q1 ðzÞdz

0

Zh f2 ðrÞ2prdr ¼

0

q1 ðzÞdz ¼ q

ð9Þ

0

Free Surface Problem

The equilibrium equation of the liquid phase in the gravity field can be used to predict the cross section of the deposited layer. This equation links the curvature of the free surface of the molten pool and the surface tension by the hydrostatic pressure, see, Fig. 3 [18]: rj ¼ qgzo þ C;

ð10Þ

where r is the surface tension; j is the curvature of the free surface; q is the density; g is gravity constant C is the Lagrange multiplier. In the coordinate system associated with the inclined plane components of gravity force vector will be the following: gx ¼ g cos b;

gy ¼ g sin b;

ð11Þ

where b is the angle of inclination of the substrate, ranging from 0 to 2p. Thus, Eq. (10) taking into account inclination of substrate will be: rj ¼ qgðzo cos b þ yo sin bÞ þ C

Fig. 3 Schematic of the free surface shape of the drop on inclined plane

ð12Þ

Model for Numerical Simulation of Temperature Field and Bead …

405

The cross-sectional shape of the weld bead described by the curve yo ¼ f ðzo Þ can be defined as follows: j¼

f 00 ð1 þ f 02 Þ3=2

1 ¼ 0 f

1 pffiffiffiffiffiffiffiffiffiffiffiffiffi 1 þ f 02

!0 ð13Þ

A parametric form of the curve can be given as:

zo ¼ zo ð/Þ yo ¼ yo ð/Þ

0/a

ð14Þ

After manipulations, the following system of ordinary differential equations is obtained 8 < dzo ¼ sin / B d/ ð15Þ : dyo ¼  cos / B d/ where B ¼ qgðzo cos brþ yo sin bÞ þ C. The boundary conditions for the obtained ODEs are the following: ð1Þ zo ða1 Þ ¼ 0; ð2Þ zo ða2 Þ ¼ 0;

yo ða1 Þ ¼ 0;

ð16Þ

yo ða2 Þ ¼ w;

ð17Þ

where w is the width of the molten pool. Zw ð3Þ

zo dy0 ¼ Sb ;

ð18Þ

0

where Sb is the cross-sectional area of weld bead. The required shape of the weld is determined using the previously calculated width of the molten pool and the area of the metal being deposited by iteratively selecting the values of the boundary angles a and the Lagrange multiplier C satisfying the boundary conditions.

3 Example Consider the HLAW of a low-carbon steel T-joint of 14-mm-thick plates. The following welding conditions were used: welding speed 16 mm s−1 (1 m min−1); laser power 8 kW; focus distance +5 mm; focal length 300 mm; spot diameter in focus 450 lm; arc current 304 A; voltage 26.8 V; stick-out 20 mm; electrode wire

406

S. Ivanov et al.

diameter 1.2 mm; wire feed rate of 167 mm s−1 (10 m min−1); shielding gas 80% Ar + 20% CO2; gas flow rate 30 l min−1 [19]. The arrangement of the welding torch and the laser beam is shown in Fig. 2, corresponded to the following parameters: h = 1 mm; d = 3 mm; aA = 30°; aB = 80°; bA = 10°; bB = −20°. The solution domain was meshed by a three-dimensional uniform grid with a 0.25-mm step. Each node was assigned an index corresponding to the type of material—air or a solid. The temperature dependence of the thermophysical properties of low-carbon steel was used [20]. According to the developed algorithm, at the first step, the melting rate of the wire is calculated using a corresponding mathematical model. The obtained results (see Fig. 4) showed that the arc current, which provides the wire melting speed equal to the feed rate, is 307 A, which is slightly different from the actual value of 304 A. Let us analyse the effect of the stick-out of the electrode wire. With an increase in stick-out by 25% to 25 mm, the arc current should be reduced to 281 A while reducing stick-out by a similar amount to 15 mm, the arc current should be increased to 340 A. The calculated temperature field showed that the coupled effect of heat sources leads to the formation of a single weld pool, which suggests the correct choice of the distance between the sources. The calculated penetration depth matches well with the experimentally measured (Fig. 5). The shape of the weld cross section was calculated using calculated volume of the filler metal deposited per unit time and the position of the weld pool shape (width of the weld) using the corresponded model. The figure shows that approximation of the weld surface by a finite-difference mesh leads to its minor distortions, which do not affect the calculated temperature field. The heat flux on the surface of the welded joint was applied according to the normal to the weld surface. The volume of metal deposited per unit of time depends on the diameter of the wire, its melting rate and welding speed. With an increase in the wire melting rate, the volume of deposited metal increases, which leads to an increase in the convexity Fig. 4 Effect of arc current and stick-out on melting rate of electrode wire

Model for Numerical Simulation of Temperature Field and Bead …

407

Fig. 5 a Peak temperature field in cross section of welded joint; b macrosample of welded joint

of the weld (Fig. 6a). According to ISO 12932 [21], the amount of excess of the convexity of the seam should not exceed h  1 mm + 0.1b = 1.68 mm to achieve quality level B. Reducing melting rate of the wire leads to the changing sign of curvature of weld bead and resulting in a concave seam with a smoother transition from the weld to the base metal, which contributes to increasing the fatigue strength of welded joints under variable and vibration loads. With a significant decrease in the volume of the metal being deposited, there is a probability of obtaining welded joints with undercut defects (Fig. 6b). According to [21], the depth of undercut should not exceed 0.5 mm for the quality level B and not more than 1.0 mm for level D.

Fig. 6 a Effect of melting rate on the shape of welded joint; b macrosample of welded joints

408

S. Ivanov et al.

4 Conclusions • A model of HLAW has been developed, which allows determining the effect of various process parameters on the quality of welded joints and process stability. • A model for free surface formation of drop located on inclined plane has been developed and used for the simulation of cross-sectional shape of weld bead of a different type of welded joint and welding position. • Comparison of the calculated and experimental data of melting rate of the electrode wire and the formation of welded joints during HLAW of T-joint showed good agreement. • The effect of fluctuations in the process parameters on the quality of welded joints is established by simulation.

References 1. Unt A, Salminen A (2015) Effect of welding parameters and the heat input on weld bead profile of laser welded T-joint in structural steel. J Laser Appl 27. https://doi.org/10.2351/1. 4906378 2. Unt A, Poutiainen I, Grünenwald S, Sokolov M, Salminen A (2017) High power fiber laser welding of single sided T-joint on shipbuilding steel with different processing setups. Appl Sci 7:1276. https://doi.org/10.3390/app7121276 3. Turichin G, Kuznetsov M, Tsibulskiy I, Firsova A (2017) Hybrid laser-arc welding of the high-strength shipbuilding steels: equipment and technology. Phys Proc 89:156–163 4. Turichin G, Tsibulskiy I, Kuznetsov M, Akhmetov A, Mildebrath M (2015) Influence of the gap width on the geometry of the welded joint in hybrid laser-arc welding. Phys Proc 78:14–23 5. Sokolov M, Salminen A, Somonov V (2012) Laser welding of structural steels: influence of the edge roughness level. Opt Laser Technol 44(7):2064–2071 6. You D, Gao X, Katayama S (2014) Review of laser welding monitoring. Sci Technol Weld Joi 19(3):181–201 7. Ribic B, Palmer T, DebRoy T (2009) Problems and issues in laser-arc hybrid welding. Int Mater Rev 54(4):223–244 8. Cho M, Farson D, Lim Y, Choi H (2007) Hybrid laser/arc welding process for controlling bead profile. Sci Technol Weld Joi 12(8):677–688 9. Nin J, Zhang L, Na S (2017) Numerical study of the effect of laser-arc distance on laser energy coupling in pulsed Nd: YAG laser/TIG hybrid welding. Int J Adv Manuf Technol 91:1129. https://doi.org/10.1007/s00170-016-9812-9 10. Turichin G, Valdaytseva E, Tzibulsky I, Lopota A, Velichko O (2011) Simulation and technology of hybrid welding of thick steel parts with high power fiber laser. Phys Proc 12:646–655 11. Pittner A (2012) A contribution to the solution of the inverse heat conduction problem in welding simulation. BAM Federal Institute for Materials Research and Testing 12. Hirata Y (1995) Physics of welding (III)—melting rate and temperature distribution of electrode wire). Weld Int 9(5):348–351 13. Maruo H, Hirata Y, Noda Y (1984) Heat contents and temperature of metal droplets in pulsed MIG welding. Quart J Jpn Weld Soc 2(4):573–578

Model for Numerical Simulation of Temperature Field and Bead …

409

14. Maruo H, Hirata Y, Noda Y (1985) Wire melting rate in pulsed MIG welding—study on pulsed arc welding (report 3). Quart J Jpn Weld Soc 3(1):191–196 15. Halmoy E (1979) Wire melting rate, droplet temperature and effective anode melting potential. In: Arc physics and weld pool behaviour. international conference, London, pp 49–57 16. Halmoy E (1991) Current-voltage process characteristics in gas metal arc welding. ASME Weld Join Process 51:17–27 17. Nguyen N (2004) Thermal analysis of welds. WIT Press 18. Landau L, Lifshits E (1959) Fluid mechanics. Pergamon Press 19. Wahba M, Mizutani M, Katayama S (2015) Hybrid welding with fiber laser and CO2 gas shielded arc. J Mater Process Technol 221:146–153 20. Pavlyk V (2004) Modelling and direct numerical simulation of dendritic structures under solidification conditions during fusion welding. Shaker Verlag GmbH 21. ISO 12932 (2013) Welding—laser-arc hybrid welding of steels, nickel and nickel alloys—quality levels for imperfections

QMS as Tool for Improving Maintenance and Repair Processes of Traction Rolling Stock A. V. Muratov, V. V. Lyashenko and S. A. Petukhov

Abstract The chapter suggests an approach to the development of a quality management system (QMS) in loco service depots on traction rolling stock maintenance and repair. The proposed way for the QMS implementation may be introduced into the system of traction rolling stock maintenance and repair. The issue of improving the maintenance and repair quality is especially relevant for the operation of a high-tech traction rolling stock at a high complexity of design solutions. Analyzing and developing the solutions in terms of enhancing the traction rolling stock maintenance and repair with the use of quality assurance tools will make it possible to resolve optimal options for the implementation of maintenance and repair processes. Reducing rolling stock maintenance and repair costs can be implemented providing the update of existing repair processes and the ability to forecast the factors influencing risk formation. In this connection, the authors consider the problems relating to risk management introduction in a loco service depot. They provide the most reasonable approach toward the introduction of uncontrolled factor—“uncertainties”—risk management. The authors propose a mathematical model for the solution of one-criterion problems as one of the options for the selection of an optimal strategic decision on the business process building in the loco service depot. The chapter considers a decision-making approach on the basis of strategy and guarantee.










Keywords Quality tools Quality assurance Corrective measures Efficiency of maintenance and repair processes Risk management Uncontrolled factors Uncertainties










A. V. Muratov
V. V. Lyashenko (&)
S. A. Petukhov Samara State Transport University, Svoboda Street, 2 V, 443066 Volga Federal District, Samara, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_44

411

412

A. V. Muratov et al.

1 Introduction A large amount of breakdowns and failures in the loco operation is related to the low quality of repair in loco service depots. It resulted in the increase in the downtime of locos at all types of maintenance and repair. This factor causes the growth of expenses due to the need for keeping them in good condition. Contemporary quality management systems (QMS) are a set of tools applied to resolve the problems of reasonable selection and taking management decisions at the standardization and certification of products, quality improvement planning, etc., which are also directed, in its turn, toward the operation process, implemented by a company [1–3]. In this case, the company is a loco service depot. Regarding the philosophy of international standards and ISO principles, applied to any kind of product manufacturing, it is necessary to use a process-based approach that is why all the actions taken in the loco service depot should be considered as a set of processes. The processes should be understood as logically structured sequences of stages, transforming inputs into outputs [4, 5].

2 Problem Statement To develop a QMS at rolling stock maintenance and repair companies, we assume the following: The loco service depot implements only one business process, which stipulates for the provision of traction equipment to consumers, and such consumers can be represented either by senior management of RZHD, JSC, or by other companies using traction rolling stock as a traction unit. Naturally, both the organization and systematization of this business process should be implemented in compliance with the principle of consumer orientation that is why a traditional organization structure does not meet the requirements of a modern system of the “client–consumer” relationships. From the perspective of the proposed pattern, one can assume that every stage of a business process adds value for end consumers. However, in practice only an insignificant part of stages and operations, included in a business process of such service depot, has an impact on meeting the requirements of a traction service. Most stages and operations, on the contrary, are preconditioned by interconnections of internal process operations within a service company and do not influence the process client at all. Judging by this, the authors draw a conclusion that the optimization of internal processes of a loco service depot is a key reserve for gaining additional efficiency in terms of a whole company. The key to a successful QMS introduction is continuous control of the quality of processes and subprocesses, directed at the implementation of the business process of the loco service depot. The process quality assurance, implemented at the manufacturing facility, is one of the fundamental activities, including measurement conduct, expert evaluation,

QMS as Tool for Improving Maintenance and Repair Processes …

413

tests and evaluation of the object parameters, and comparison of the obtained values with the set requirements to these parameters (quality indicators). The conduct of the product quality analysis, as a rule, stipulates for the statistical evaluation methods, which provide simplicity, clarity and data visualization. These methods in scientific and technical information are denoted as quality assurance tools [6–8]. Modern quality assurance tools are the methods used to resolve the problems of a qualitative evaluation of the quality parameters. Such an evaluation is necessary for a reasonable selection and taking of management decisions at the product standardization and certification, quality enhancement planning, etc. Despite their simplicity, these methods assist in maintaining the connection with statistics and provide an opportunity to use the results of such methods and improve them, if necessary. Simple quality assurance tools include the following statistical methods: log sheet, histogram, scatter diagram, Pareto chart, stratification, Ishikawa chart (cause and effect chart), and control chart. The listed methods can be considered both as separate tools and a system of methods.

3 Research Questions The application of these tools in production conditions allows implementing the important principle of quality management system functioning in compliance with the international standards and obtaining important information in relation to the condition of the processes under consideration. In its turn, it will provide for the revelation of the deviations from planned parameters, identify the cause, and develop a set of effective measures to eliminate them [9, 10]. The quality assurance tools for the technical analysis of the rolling stock repair and maintenance quality for the purpose of identifying cause and effect of a low-quality repair and the development of corrective action plan are widely applied in loco service depots on rolling stock repair [11–13].

4 Theoretical Section An advanced monitoring of technical means’ failures caused by a low quality of repair, conducted to identify the key factors of the need in unscheduled repair, resulted in a cause and effect chart (Fig. 1). The Ishikawa chart of cause and effect shows the key factors of unscheduled repairs are the following: • Low technical knowledge level of maintenance teams; • Other factors; • Low performer’s discipline;

414

A. V. Muratov et al.

Fig. 1 Cause and effect (Ishikawa) chart of unscheduled repair factors in a loco service depot

• Low quality of the repair conducted; • Provision of spare parts and materials; • Unfavorable meteorological conditions. As a rule, basing upon the conducted research, one develops a plan of corrective and preventive actions to eliminate the causes identified. However, the provided algorithm for the use of a QMS to eliminate the causes of low-quality maintenance and repair and improve the efficiency of these processes does not involve so-called risk management principles. At present, the requirements, determined by GOST R ISO 9001-2015, toward management activities including planning, management, activity maintenance, and enhancement stipulate for the integration of risk management principles into the existing quality systems [13, 14]. To meet the requirements of an existing international standard, the company needs to plan and introduce corresponding risk-related actions. A coordinated and systematic management activity related to the risk studying is risk management, which is an integral part of control management and is embedded into a general system. At present, there are numerous methods for the risk criteria identification and determination, which can be applied toward traction rolling stock maintenance and repair organizations; however, their application is not holistic and is represented as separate measures with no common concept in risk management application.

QMS as Tool for Improving Maintenance and Repair Processes …

415

That is why the chapter suggests building a risk management system at all the stages of the loco service depot functioning. The proposed algorithm for the system functioning consists of the following steps. At the initial stage of the proposed algorithm implementation, the analysis is conducted; such analysis is based on the basis of the analysis of engineering and technical risks at the design and technology of the manufacturing facility, category FMEA. FMEA (Potential Failure Mode and Effects Analysis) is the analysis of kinds and consequences of potential failures—the method described in the manual with a similar title for the standard QS-9000 “Quality System Requirements;” within the standard, the method covers both the analysis of consequences and the analysis of causes of potential defects of technical facilities and processes of their manufacture as well as the necessary improvement of technical objects with the account of the conducted analysis data [15]. The FMEA method allows analyzing potential defects, their causes and effects, evaluates the risk of their occurrence and non-identification at the enterprise and takes measures to eliminate or decrease the risk and damage of their occurrence. It is one of the most efficient methods for the improvement of the technical object structure and the processes of their manufacture at such essential stages of the product life cycle as its development and manufacture preparation. The proposed method does not consider the loco design peculiarities but rather the technology influence onto the occurrence of equipment defects causing loco failures. The present analysis does not consider the factors arising under specific loco operation conditions. This method of the system of analysis of potential defects and their effect, PFMEA, allows specifying the area where the measures should be taken immediately [16, 17]. Consider the analysis results by the example of a loco service depot. To determine the area of maximum risk on equipment with the technical means failures, make the Pareto chart. The chart demonstrates that approximately 80% of failures in the operation of units and assembly units: • • • • •

Diesel and diesel equipment Auxiliary equipment Rolling stock Traction electrical machines Fuel injection equipment.

Basing upon the conducted analysis, the corrective measures are developed to improve the loco park technical state, to the analysis of the train movement safety and the technical state of train rolling stock. Upon the results, one is able to identify the key problem areas in the processes of rolling stock maintenance and repair. The key technology defects are listed below: 1. Mistakes in technology processes’ formation; 2. Insufficient knowledge of possible similar technologies;

416

A. V. Muratov et al.

3. Insufficient analysis of specific equipment operation, maintenance, and repair conditions; 4. Violation of process modes and requirements. As the aforementioned analysis conduct has shown

5 Findings On the basis of the factor analysis (Fig. 1) and identified defects in the technology of rolling stock maintenance and repair, it is obvious that the presence of a statistical data set is insufficient for the determination of possible risks, that is why, in relation to rolling stock service companies, the concept “risk” should be considered as a deviation of any values from the set ones [18]. In addition, the analysis shows the existence of uncertainties—uncontrolled factors, which can directly influence the quality of the processes of repair and maintenance and cause a potential deviation of any parameters from their necessary values, i.e., risks. Using this definition, the authors state the notion of a risk criterion or its measure and take a difference between the desired values of the process functioning quality and the actual value in terms of the uncontrolled factor as a risk measure. Therefore, to identify «risk», one can use the following three components: • A set of potential tactical strategies of the process owner, preconditioned by the current state of the director awareness, on the basis of which one or another action is taken from the list of acceptable actions with the account of this information; • A set of acceptable values of uncertainties; • A criterion, evaluating the positive result of the tactical strategy implementation. By these three components (criterion, set of tactical strategies, and set of uncertainties), the process owner can identify risk or an indicator (risk criterion). A mathematical model for the solution of a one-criterion problem at uncertainty can be stated as follows: hX; Y; f ðx; yÞi

ð1Þ

where the selection of alternative x from the set X  Rn is at the disposal of the process owner (decision taking person). The process owner purpose—selection of x 2 X for which a scalar criterion f(x, y) reaches a potentially higher value. At this, one needs to take into account the influence of a large number of various uncertainties y, the classification of which is given above, and we only know they take a value from a given set Y  Rm.

QMS as Tool for Improving Maintenance and Repair Processes …

417

6 Conclusion Using a proposed approach, one can identify, analyze, and evaluate the risks, which can influence the achievement of purposes, tasks, set by the company’s QMS for rolling stock repair. On the basis of the proposed model on three components (criterion, set of tactical strategies, and set of uncertainties), one obtains the risk or its indicator (risk criterion). This stage is the most time-consuming and significant in terms of building a comprehensive risk management system and mostly depends on the process owner decision taking skill. The further stages of management risk implementation are the processes of planning necessary measures on risk elimination, implementing measures on risk elimination, analysis of the results and efficiency of measures on risk elimination, etc. The efficiency of the implementation of these stages is directly connected with an arranged system of continuous analysis conducted by senior managers as well as by the selection and classification of the risks in relation to their significance, contingency, etc. Generally, the proposed model of risk management, integrated into the quality management system of a loco service depot, will allow the senior staff to conduct the analysis of probabilistic risks at decision taking. In its turn, it will facilitate the improvement of the traction rolling stock maintenance and repair processes.

References 1. Toni OI (2006) Quality assurance system must be grown up. Gudok 2. Adler YP (2000) Quality and market, or how organization focuses on provision of consumer requirements. Supplier and Consumer. RIA “Standards and Quality”, Moscow, 128 p 3. Masaaki Imai K (2005) The key to Japan’s competitive success. 2d edition. Translated from English. Alpina Business Books, Moscow, 274 p 4. Goncharov EN (2007) Some considerations on identification of quality management system processes. Stand Qual 9:68–72 5. Aleksandrov SL (2009) Organization processes at requirements observance GOST R ISO 9001. Qual Manag Methods 5:38–68 6. Shvets VYe (2001) Measuring processes in modern quality management system. Qual Manag Methods 1:11–13 7. Kane MM, Ivanov BB, Koreshkov VN, Skhirtladze AG (2008) Systems, methods and tools of quality management: study guide. Piter, Saint-Petersburg, p 560 8. Gorbunova OI Guseva IK (2016) Tools and methods of quality management: study guide. BGU Publishing House, Irkutsk, p 116 9. Lyashenko VV, Muratov AV, Petukhov SA (2015) Using quality management tools to increase efficiency of rolling stock repair and maintenance. Povolzhye Transp Guide 3(51): 22–26 10. Muratov AV, Petukhov SA, Lyashenko VV (2015) Increasing efficiency of maintenance and repair of diesel motors at polygon of South Ural Railway. Sci Educ Transp 1:16–19 11. Andronchev IK, Lipatov AI, Lukin NF (2010) System medium of quality of linear railway company. SamGUPS, Samara, p 270

418

A. V. Muratov et al.

12. Petukhov SA, Muratov AV, Lyashenko VV (2015) Increasing efficiency of maintenance and repair of motor & train rolling stock at polygon of South Ural Railway. In: Innovations and research in transport complex: materials of the 3d international scientific and practical conference. Part I (in two parts), Kurgan, pp 241–243 13. GOST R ISO 9001-2015 Quality management systems. Requirements 14. GOST R ISO 31000 Risk management. Principles and management 15. GOST R 51814.2-2001 Quality systems in motor industry. Method of analysis of potential defect kinds and effects 16. Analysis of kinds and effects of potential failures. FMEA: reference manual. Fourth edition. Translated from English. (2009). SMTs, LLC, Prioritet, N. Novgorod, 142 p 17. Panyukov DI, Kozlovsky VN (2010) Fundamental basics FMEA for motor industry: monograph. SNTS RAS, Samara, p 150 18. Zhukovsky VI (2011) Risks in conflict situations. LENAND, Moscow, p 328

Heat Treatment Effect on the Structural and Elastic Characteristics of a Single-Component Abrasive Tool M. Yu. Polyanchikova

Abstract At the finishing operations of processing high-quality parts (internal tracks of rolling bearings, engine cylinder liners, etc.), abrasive tools are most often used. According to the results of numerous studies, the effect of abrasive grains size on the surface layer of the critical parts state and, as a result, the wear resistance of the latter has been proved. Numerous studies of the author proved the effectiveness of abrasive tools without ligament (SCAT) use in the processing of such critical parts as the holes of cylinder liners. Currently, research is being conducted on the effect of the under consideration bondless single-component abrasive tool (SCAT) production modes on its properties. The chapter presents the results of the research study of elastic properties of a bondless SCAT after sintering; the method of determining the modulus of elongation and porosity based on theoretical and experimental data has been used; the dependency between the modulus of elongation and porosity and experimental formulas indicating these dependencies are presented. Based on the experimental and calculated data on the magnitude of the modulus of elongation, depending on the sintering temperature and the size of the abrasive grains in the initial abrasive compound, empirical dependencies of the elastic characteristics have been determined in order to predict the production of a bondless tool with predetermined properties. Keywords Abrasive tool abrasive tool Porosity





Modulus of elongation
Sintering process of an

1 Introduction Finishing operations, according to the data given in [1–16], have a significant influence on the surface layer of part properties formation. These properties will largely depend on the machining conductions and the characteristics of the abrasive M. Yu. Polyanchikova (&) Volgograd State Technical University, 28, Lenin Avenue, Volgograd 400005, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_45

419

420

M. Yu. Polyanchikova

tool used in finishing, including its grain size and structure [12]. Three factors mainly influence the structure formation of a standard abrasive tool: the quantity abrasive grains, binder, and pores. Producing a one-component abrasive tool (SCAT) that does not contain any bond components in its composition and is obtained from an abrasive compound of white alumina and boron carbide [1] belongs to powder metallurgy and consists in pressing and further sintering at high temperature. Only after sintering, the tool obtains the necessary strength, structural and elastic characteristics, i.e., porosity and the modulus of elongation that, being considered structurally insensitive, are deeply related to the nature of bond strength and other structural features of the porous fragile body (bonding, crystalline structure, etc.). Therefore, the study of the elastic deformation zone is of great practical importance for the fragile states of bodies in the conditions of processing and operation. The importance of studying elastic deformation is also due to the fact that the process of deformation and fracture begins with it [2]. The chapter presents results of the study of the elastic properties of a bondless abrasive tool (SCAT) after sintering.

2 Formation of Properties of a Single-Component Abrasive Tool in Producing In the manufacture of single-component abrasive tools (SCAT) [3], the task of ensuring the uniform distribution of abrasive grains and pores is one of the main tasks in the formation of their structure at the stage of preparing an abrasive compound. When making molding sand for manufacturing a standard abrasive tool, it is enough to mix all the components thoroughly, but the SCAT manufacture has a number of features. As a rule, the size distribution of the abrasive grains is not considered, since it usually complies with the standard [13]. However, the size distribution of the abrasive grains and the corresponding size distribution of the pores within the tool are the importance for SCAT, obtained by shock-wave pressing [1]. The size distribution of the abrasive grains after shock-wave pressing corresponds to an asymmetric b distribution. The non-uniformity in the size distribution of the pores over the tool volume is 20%. This has a considerable influence on tool performance, since section with pores of different size appears at its working surface, which results in spoiling of these sections and loss in cutting power of the tool. Thus, cutting and non-cutting sections appear randomly at the tool’s working surface. This impairs the surface quality of the final part. Uniform cutting may only be ensured by more uniform size distribution of the abrasive grains and pores within the tool. Research indicates that this is possible if some quantity of boron carbide is added to the white electrocorundum abrasive powder, where the grain size of the boron carbide is 30–50% of the abrasive powder

Heat Treatment Effect on the Structural and Elastic …

421

grain size. Since the microhardness of the boron carbide is almost twice that of white electrocorundum, the larger electrocorundum grains are further ground in dynamic loading by the shock wave. This results in a more uniform size distribution of the electrocorundum grains and correspondingly a more uniform size distribution of the pores within the tool. Since the temperature limit of stability of boron carbide is about half that of white electrocorundum, the boron carbide generates pores in the sintering of an abrasive (which is carried out at a temperature of 1700–1900 °C). Pressing of abrasive powder compounds was carried out according to the scheme shown in Fig. 1. The studies were conducted on the T-1226 electro-hydraulic installation. After the accumulation of electrical energy in capacitors with a capacity of 200 lF at a voltage of 10 kV with two electrodes 2 and 4 (discharge chamber casing) used, an electric arc was formed in the water bath 6, which caused a gas bubble being instantly formed; on its disappearance, a shock wave was formed. The wave acted on the slab 7 and punch 1 and through the drummer 8 on the abrasive molding compound 11 that contained alumina powder (Al2O3) for 90% and boron carbide (B4C) for 10% and was used in the SCAT manufacture according to the method [3]. Elasticity is known to be a property of bodies to change their shapes and sizes under the action of loads and spontaneously restore the original configuration when the external loads are no longer applied. The elasticity of bodies is due to the interaction forces of the atoms they are built from. Under the action of external stresses, atoms were displaced from their equilibrium positions, which was accompanied by an increase in the potential energy of the body by an amount equal to the work of external stresses in changing the shape of the body. After the removal of external stresses, the configuration of the elastically deformed body with non-equilibrium interatomic distances turned out to be

Fig. 1 Scheme of explosive pressing of a single-component abrasive tool, where 1 is the punch; 2 is the cathode; 3 and 5 are the insulators; 4 is the case of an explosive (discharge) chamber; 6 is water; 7 is the slab; 8 is the drummer; 9 is the support bar; 10 is mold housing; 11 is compressible abrasive powder; 12 is the matrix; and 13 is the press table

422

M. Yu. Polyanchikova

unstable and spontaneously returned to an equilibrium state. The excess potential energy stored in the body was converted into the kinetic energy of oscillating atoms, i.e., into the warmth. This paper presents the results of studying the structural and elastic characteristics of the SCAT, including the density, porosity and modulus of elongation depending on the porosity of the abrasive tool after pressing and sintering of the studied abrasive compound, containing aluminum oxide (Al2O3) for 90% and boron carbide (B4C) for 10% to obtain the SCAT according to the patent [1] and published in [17, 18].

3 Determination of the Elastic Characteristics of a Single-Component Abrasive Tool According to some scientists [2, 19, 20], the value of the modulus of elongation can be considered as a characteristics of elastic resistance or elastic hardenability of the material, i.e., as a characteristics of the stress intensity with the elongation increasing. The greater the modulus of elongation is, the stronger the stresses grow with elastic deformation increasing. Frenkel [21] considered the sintering stage as a fusion of droplets, when the space between the particles was occluded so that the rest of pores were separated. The subsequent stage of sintering was reduced to the “occlusion” of the rest of pores, not communicating with each other. Existing explanations of the processes of bonding materials in the solid phase were based either on the ideas of the need to increase the energy of surface atoms participating in the interaction, or on the idea of the possible interaction only on juvenile surfaces formed after removal of oxide films, or on the ideas of the decisive role of diffusion [4]. All hypotheses can be divided into two groups. The former explains bonding between dissimilar materials due to curing as a consequence of the interatomic interaction forces. Curing is a diffusionless process that is possible to be followed by diffusion processes. If, when bonding in the contact zone, juvenile surfaces were formed, their activity was enough for curing. The bonding process was reduced to cross-linking of contact surfaces and bonding between atoms on different surfaces [5]. When studying the sintering process, the following mechanisms are usually considered: (a) without the material transferred (by bonding); (b) with the material transferred for considerable distances by moving the lattice elements (surface and bulk diffusions over defects, through the lattice and along the grain boundaries, evaporation and condensation) or displacement of the entire lattice (plastic flow and slip along the grain boundaries); and (c) with the material transferred within the interatomic distances (return and recrystallization). Due to plastic deformation in pressing the powders, the surfaces of the contacting grains approached each other sufficiently, the surface films repeatedly broke,

Heat Treatment Effect on the Structural and Elastic …

423

and their insulating action was largely eliminated. Therefore, the geometric factor of convergence became the main factor. Andrievsky [6] also proved that the self-diffusion mechanism dominated in sintering of alumina powders. At the initial stage of sintering, the growth of contact zones was decisive; at a later stage, closed porosity was formed. Kuchinsky G. S. extended the theory of bulk diffusion on the late stages of sintering. In his work [6], he considered the influence of grain boundaries on the sintering process, especially at the final stage of the pores’ disappearance. At the same time, based on two mechanisms, i.e., only diffusion mechanism along grain boundaries and bulk vacancy diffusion (upon reaching grain boundaries, vacancies were removed from the body on its surface), he estimated the rate of pores disappearing from a polycrystalline sintered body. The limiting unit here was the bulk diffusion as the slowest one. The pore was considered as a place of the vacancies’ accumulation; the mechanism became evident when considering vacancies as diffusing units replaced by atoms taking into account the gradient of hole concentrations from the pore depthward the material [7]. Figure 2 shows the dependency between the SCAT density and the sintering temperature. The dependencies obtained using [8] are as follows: initial Al2O3 grain size of 100 microns: q ¼ 46:1
T 0:564

ð1Þ

initial Al2O3 grain size of 200 microns: q ¼ 180:8
T 0:376

Fig. 2 Effect of sintering temperature (T, °C) on the SCAT density (q, kg/m3) for the initial Al2O3 grain size of 100 lm for 1; 200 lm for 2; 320 lm for 3; 400 lm for 4; and 500 lm for 5

ð2Þ

424

M. Yu. Polyanchikova

initial Al2O3 grain size of 320 microns: q ¼ 217:4
T 0:346

ð3Þ

initial Al2O3 grain size of 400 microns: q ¼ 547:6
T 0:218

ð4Þ

initial Al2O3 grain size of 500 microns: q ¼ 716:6
T 0:18

ð5Þ

Sintering of boron carbide resulted in its burn-off even at a temperature of 700– 800 °C [6] with small pores in the abrasive body that were occluded with an increase in sintering temperature. As a result, the porous body became poreless in places. The next stage of the research was the determination of the porosity of sintered abrasive bodies. This value was determined according to [9, 15], using the ratio (6): P¼

qps  qpb ; qps

ð6Þ

kg where qps is the density of pure Al2O3 substance, kg/m3; qps ¼ 3900 m3 ; qpb is the 3 density of a porous body, kg/m . According to the formula presented in [10], the modulus of elongation can be determined by the porosity values: E 15ð1  lÞ
P ; ¼1 Ep ð7  5lÞ þ 2ð4  5lÞ
P

ð7Þ

where Ep is the modulus of elongation of the porous body, GPa; P is the porosity; E is the modulus of elongation of the pure substance, GPa; l is Poisson’s ratio (l = 0.23). Figure 3 presents the dependencies between the modulus of elongation and SCAT porosity. After processing the calculated graphs according to the method [8] and taking into account the provisions of [14], the following analytical dependencies were obtained: for the A12O3 grain size of 100 lm: Ep ¼ 1166
P0:412

ð8Þ

Heat Treatment Effect on the Structural and Elastic …

425

Fig. 3 Dependency graph between the modulus of elongation (Ep) and porosity (P) after sintering at the initial Al2O3 size of 100 lm for 1; 200 lm for 2; 320 lm for 3; 400 lm for 4; and 500 lm for 5

for the A12O3 grain size of 200 lm: Ep ¼ 1256
P0:462

ð9Þ

for the A12O3 grain size of 320 lm: Ep ¼ 1383
P0:534

ð10Þ

for the A12O3 grain size of 400 lm: Ep ¼ 1465
P0:577

ð11Þ

for the A12O3 grain size of 500 lm: Ep ¼ 1164
P0:41

ð12Þ

4 Conclusions According to the results of the research conducted, the following conclusions can be drawn. The formation of a porous abrasive body in sintering led to occlusion of very small pores and formation of a large number of bridge bonds between the abrasive Al2O3 grains, which contributed to the improvement of the elastic properties of the porous body. The practical value of the elastic properties of materials is large. To prevent buckling, materials with a high value of the modulus of elongation should be used. Due to the accumulation of a large supply of elastic energy in operation, materials

426

M. Yu. Polyanchikova

with a lower value have a greater development rate of microcracks, which is their disadvantage [11]. Pressing has a slight strength due to the adhesive and cohesive forces of interatomic interaction and, therefore, at the sintering stage, the tool acquires the final elastic characteristics [12].

References 1. Polyanchikov YuN, Polyanchikova MYu, Kozhevnikova AA, Emelianenko AA, Angelovskaya NV, Krainev DV (2007 Apr) A method of manufacturing abrasive products. Patent of Russia 2293013 2. Friedman YaB (1974) Mechanical properties of metals. Part 1. Deformation and destruction, 3rd edn. Mashinostroenie, Moscow 3. Polyanchikov YuN, Polyanchikova MYu (2015 June) A method of manufacturing abrasive products. Patent of Russia 2543024 4. Geguzin YaE (1984) Physics of sintering. Nauka, the main edition of the physical and mathematical literature, Moscow 5. Polyanchikova MYu, Polyanchikov YuN (2016) Improving the quality of one-component abrasive and diamond tools obtained by shock-wave pressing (monograph). IUNL VSTU, Volgograd 6. Andrievsky RA, Umansky YaS (1977) Implementation phases. Nauka, Moscow 7. Kuczynski G (1973) Phys Sinter 5(2/2):41–50 8. Lvovsky EN (1982) Statistical methods for constructing empirical formulas. Higher School, Moscow 9. GOST 2409-95 (ISO 5017:88) Interstate standard. Refractories. Method for determination of bulk density, apparent and true porosity, water absorption Intergovernmental Council for Standardization, Metrology and Certification, Minsk, 2002, 6 p 10. Hashin Z (1962) Relations between Young’s modulus and temperature. J Appl Mech 29:43–147 11. Solonenko VG, Zareckij GA (2007) On a unified theory of wear cutting tools. J STIN 4:7–11 12. Shumyacher VM (2004) Mechanical and chemical processes at the finish abrasive processing, (monograph). Edition of Volgograd State University of Construction and Civil Engineering, Volgograd 13. GOST R 52381-2005 (2006) National standard. Abrasive materials. Grain size and grain composition of grinding powders. Control of grain composition. Standardinform, Moscow, 15 p 14. GOST R 50779.10-2000 (ISO 3534.1:93) National standard. Statistical methods. Probability and basic statistics. Terms and Definitions. Standardinform, Moscow, 2008, 47 p 15. Element properties: handbook (1985) Metallurgia, Moscow 16. Bezyazichnij VF (2003) Problems of improving the technological processes of machining parts of high-precision components and products. J Appl Dir Eng Mag 7:2–11 17. Polyanchikov YuN, Plotnikov AL, Polyanchikova MYu, Kurchenko AI, Kursin OA (2009) New single-component abrasive tool. J Russ Eng Res 29(3):310 18. Polyanchikov YuN, Plotnikov AL, Polyanchikova MYu (2014) Particulars of ceramic article fabrication by impact pressing. J Glass Ceram 71(5–6):180–181 19. Shumyacher VM, Pushkarev OI, Slavin AV (2017) Energy efficient technology of obtaining advanced composite grinding materials and tools in “silicon carbide–corundum” system from aluminium containing residual products. J Procedia Eng 206:228–231 20. Kryukov SA, Tkach MA (2017) Phenomenological model of abrasive tool components mixing process. J Procedia Eng 206:200–203 21. Frenkel Ya I. Introduction to metal theory: monography. In: Vonsovskiy SV (ed) Publishing “Science”, 4th edn. 424 p

Approach for Modeling and Situational Management of Industrial Product Efficiency E. V. Orlova

Abstract In modern world, the ensuring enterprises’ (as complex production systems) sustainable functioning is one of the basic conditions for their competitiveness and efficiency. Economic changes require enterprises to improve products’ quality, to expand their use, to increase output, and to use innovative production technologies. For the effective management under external and internal environment dynamics, such tools are necessary that allow, on the one hand, identifying the financial and economic state and evaluate the product efficiency, and on the other hand, creating a set of management decisions necessary to prevent possible risks due to negative factors. An approach to managing the efficiency of industrial products based on situational analysis methods, econometric modeling is proposed. Numerical experiments evaluated the proposed approach effectiveness are given. Keywords Production systems Decision making


Life cycle modeling
Situational analysis


1 Introduction The category of efficiency is widely used in various fields of science and practice. The problem of efficiency evaluation of the production and economic systems in the context of operational management decision making remains one of the key topics of economics. The transition to the innovative path of economic development and the need to increase the pace of changes in the existing management mechanism contribute to bringing this problem to the forefront. The phenomenon of efficiency is studied by representatives of different economic schools and areas and is studied by scientists in the field of management theory [1–5], systems analysis, and modeling [6–8]. However, it should be recognized that an approach has not yet been developed to evaluate, predict, and manage the economic efficiency of the E. V. Orlova (&) Ufa State Aviation Technical University, 12 K. Marx St., Ufa 450008, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_46

427

428

E. V. Orlova

enterprise as a complex dynamic multi-level organizational system, taking into account many interrelated, often stochastically, parameters and partial indicators of the effectiveness of subsystems. Efficiency is one of the characteristics of the system quality, and its ability to produce a certain effect. Economic efficiency reflects the ability of the system to produce in the course of its operation, the economic effect. The greater the economic effect (result) and the smaller the resources for obtaining this result, the higher the economic efficiency. The analysis of the investigations in the field of modeling and control in different production systems [9, 10] is showed the following. Problems related to the economic-mathematical modeling and control of the processes of harmonization, the economic interests of participants in production and economic, in market and fiscal processes are not solved. Previous studies do not deal with dynamic pricing for enterprises products in a competitive market under changing consumer preferences and competitors behavioral strategies.

2 Mechanism for Efficiency Modeling and Management The concept of enterprises’ efficiency management is based on using modeling techniques and situation analysis. Analyzed factors are: changes in macroeconomic conditions, changes in the markets and marketing resources; enterprise management initiatives. A system of performance management as an integrated technology based on the use of specific methods and models that are adequate to revealed problems is developed. The technology consists of the following steps. Step 1. Analysis of industrial, economic, and financial conditions of the enterprise. At this stage, the identification of the current enterprise’s conditions based on financial activities analysis and financial management analysis methods is carried out. Step 2. Product efficiency estimation by types and stages of product life cycle. Efficiency studying is carried out from two perspectives: static analysis to determine the profitability (loss ratio) for each product and identification of areas for improvement of the issue; dynamic analysis based on the identification of production functions for each product. Identification is performed the first stage of product life cycle, as well as trends in sales volumes and prices on the basis of econometric modeling methods. In this step, an assessment of the product position on the curve of its life cycle is performed; production capacity of the enterprise with the position of product life cycle is carried out. Then, production efficiency is being evaluated by the monogrammed modeling method.

Approach for Modeling and Situational Management …

429

Step 3. Situational balanced management of economic parameters—product volume, price, and cost based on the methods of situational modeling and using of production rules. Step 4. Decision making about economic parameters changing. The concept of “product life cycle” is one of the most popular concepts in the practice of modern marketing. The life cycle of each product is individual and consists of a number of stages [11]: 1. Technical preparation. At this stage, the enterprise develops and implements a new product idea, and it includes the following stages: research and development works, design preparation of production, design, and development—development works, technological preparation of production. Characteristic features for this stage are zero sales and rising costs (investment costs). 2. Implementation (entry into the market). It accompanied by slow growth in sales, rising costs; Profit is absent due to marketing activities to products market promote. 3. Growth. The period of rapid market conquest, sales are grown, profit is increased. At the same time, the costs growth rates are reduced. 4. Stability. It is determined by the slow growth in sales, as the majority of potential buyers have already been attracted. Production costs are virtually unchanged, profit grows slowly. 5. Reduced sales. Reduced sales and reduced profit. Researches in the field of enterprise economic diagnostics have revealed a number of problems [12–14]. First, enterprises are still oriented toward financial evaluation methods, built on a system of accounting and reporting, which largely correspond to the dynamic economic conditions and consider the factor of periodicity analysis—not taken into account the economic efficiency of each product range, thereby sustainable management of all products is not implemented. Second, analysis results are lagging in time and cause inconsistencies normal functioning more worsen enterprise economic conditions. High dynamic of internal and external enterprise environment currently makes to the instruments of economic governance new, higher requirements. These tools should be based on economic-mathematical methods and models that reflect the basic economic laws governing enterprise activities and should ensure [15, 16]: The representation of relationships between economic variables in the wide-range of possible values and their possible combinations in order to achieve a certain level of production efficiency. The identification of different requirements for achieving certain performance levels as well as defining the efficiency changes range for the uncertainty of input data. The performance evaluation of economic variables changes. System properties of the enterprise’s economic mechanism include a large number of indicators that have a strong influence on each other. Quantitative description of these relations is a difficult problem. Their rational (required) values

430

E. V. Orlova

must be justified, and the options selected for use by “cost-efficiency”. In the offered technology for efficiency assess is used tools such as “maps of economic behavior” based on monographic models [17, 18]. Such models allow to assess the effect of various combinations of economic variables on the enterprise’s economic efficiency. The versatility of this tool ensured by using the following methods of information representations: Valuation method which is to convert the absolute values of economic variables to their relative values and allows to apply this tool for any business and products. Benchmarking technique is to form a universal model of conformity mutual relations between the main economic variables of production and sales of products, which provides a fixed level of efficiency. Graphic representation on the map of various combinations of economic variables provides the same efficiency level. Using multivariable normalized models allows to prevent the adoption of irrational and wrong decisions, to save significant amounts of financial resources, and to expedite the process of finding the most appropriate solutions that achieve the desired performance. Most valuable for practices is the breadth and integrity of the information field covered graph-analytical models, which give information about the production efficiency depending on the possible combinations of multivariate economic parameters. The technology for producing economic evaluations is to determine the values of key factors determining the inputs to the model and performance evaluation. An application technology for using monographic models has the following features: The same level of efficiency depends not only on the absolute values of economic parameters and their mutual relations. This means that the same level of efficiency can be achieved at different production scales. Multivariance of models allows to solve various problems. The main objective of the model is the use of monographic definition and formation of such a mutual correspondence between resources, production volumes, and prices at which it is provided a desired level of performance sufficient for sustainable operation and enterprise growth. Efficiency is the productivity of the resources distribution under the production process. As indicators, reflecting production efficiency (effectiveness) and business activities are the profitability (of sales, of assets, etc.). Further, we will use performance as return on sales kind a¼

P P pq  ðCf þ q
cv Þ ¼ ¼ ; C Cv þ Cf Cf þ q
cv

ð1Þ

where a—profitability (relative gain) in the production of goods in volume q; P— gross profit generated on the sale of goods in volume q at a price p; C—total costs, including the constant part Cf and a variable part Cv ¼ q
cv ; the relationship between cost and volume of output is considered to be linear C ¼ Cf þ q
cv . Analysis of Eq. (1) shows that the same level of efficiency can be achieved at various combinations of the factors standing in its right part. A certain level of profit

Approach for Modeling and Situational Management …

431

is influenced by various factors, therefore, it is necessary to analyze their impact on the profit and formulate the most reasonable (rational) in these conditions a combination of these factors. Traditionally, such analysis is carried out within the concept of CVP analysis (cost-volume-profit), or operational analysis, and modern accounting principles. We also assume that the amount of profit affects four factors: price p, unit variable costs cv , total fixed costs Cf , the volume of production q. In order to ensure the universality of developed models using a valuation method, this consists of the transition from the absolute values of these variables to their relationship like Cf Cf p q
cv ; Cv ; cv . To do this, perform the necessary transformations of the expression (1): a ¼ Cf þpqq
cv  1, or Cf þpqq
cv ¼ a þ 1. Dividing the numerator and denominator of the right side of the last equation by q
cv , we obtain: Cf p 1 p  1; or a ¼ ¼
cv q
c v c v ð 1 þ aÞ



 Cf þ 1  1: q
cv

ð2Þ

Cf is linear, with the same As seen from (2), the relationship between cpv and q
c v level of efficiency achieved a variety of combinations of variables of these relations. Cf Let us consider the relations cpv and q
c in the expression (2) at different levels of v economic efficiency a: a ¼ 0—break-even; a [ 0—profitability; a\0— unprofitability. Cf If a ¼ 0, then cpv  q
c ¼ 1. This implies that a reduction value of the first term v by a predetermined amount, such as price reduction, in order to maintain the break-even, the second term must be reduced by the same value. This is the essence of sustainable management ofeconomic  variables.  

Similarly, if a [ 0, then

Cf Cf  q
c [ 1; if a\0, then cpv  q
c \1. This v v   Cf analysis suggests that the largest difference cpv  q
c can judge the level of ecov p cv

nomic efficiency and control its change as well as to manage balanced economic characteristics when the external and internal conditions of the functioning of the enterprise. Thus, the values of the level of profitability and loss level are determined by matching the ratio of fixed and variable costs and relative prices. The size of each of the five analyzed parameters (volume of production and sales of these products, unit variable costs, fixed costs attributable to these items, the unit price, gross profit from sales of this product) is determined by many factors, differing in the degree of influence on these parameters: type and the level of production technology, composition and technical level of equipment, product quality, species, quality and prices of raw materials used, the supply and demand of these products in the market, the organizational structure of the company, number of employees and its distribution units, wage, taxes [9, 10] . These factors affect the possible ranges of the main economic indicators of the enterprise, reflecting the features of its functioning.

432

E. V. Orlova

Application of valuation allows providing a universality of graphic models and their suitability for the same economic management regardless of the scale of production. The second feature is that the graphic lines on the models are, in most cases, the discrete solution level of efficiency to be achieved under certain relations between the economic parameters indicated on the input axes of the graphical model. The third feature is the multivariance of models that allows solving problems in various productions, i.e., defining and setting parameters can be reversed. To develop decision solutions based on the method of sustainable management, it is necessary initially to determine possible changes in the controlled variables— prices, unit variable costs, total fixed costs attributable to a particular type of product, the volume of production, as well as calculates the effects (efficiency) implementation of various control methods. Balance control involves the preservation of the previous values of the production efficiency of a particular product in situations of change one or more controlled variables on the basis of changes in other variables. One of the possible alternatives such compensatory effect that returns profitability of the production to their former positions is to change one of the control variables in the model (2). Quantify the values of variables can be controlled with the use of analytical expressions. Formulas are derived from the condition of conservation of the same level of profitability of ensuring fulfillment of (2). Consider the following possible situations and the appropriate control action. 1. Changes in the composition of fixed assets DCf , which leads to an increase in fixed costs at constant values of the other variables—p ¼ const, q ¼ const, cv ¼ const; in this case, the compensating effect is   p Cf0 DCf  cv0 þ : Dcv ¼  ð 1 þ aÞ q q

ð3Þ

2. Resources cost increasing at Dcv , which leads to increase a unit variable cost at constant values of the other variables—p ¼ const, q ¼ const, Cf ¼ const. The compensating effect is Dp ¼ ð1 þ aÞ


  Cf0 þ ðcv0 þ Dcv0 Þ  p0 : q

ð4Þ

3. Product demand growth at Dq, that promotes changing in output at constant values of the other variables—Cf ¼ const, p ¼ const, q ¼ const. The compensating effect is  DCf ¼ Dq


     p p  cv0 þ q0  cv0  Cf0 : ð 1 þ aÞ ð 1 þ aÞ

ð5Þ

Approach for Modeling and Situational Management …

433

4. Product price reducing at constant values of the other variables—q ¼ const, cv ¼ const, Cf ¼ const. The compensating effect is determined with using (3–5). 5. Output becomes unprofitable a\0, then points 1–4. For neutralization negative effects of these situations, the method of compensating balanced control is to promptly bring the control object to its previous state, i.e., to the same measure of performance in terms of external and internal disturbances. Management decisions are implemented by influencing variables p; q; cv ; Cf . Application of given formulas facilitates the analysis within the model and gives the best estimate of changes in the production and economic system conditions in the event of disturbances in values of external or internal environment factors.

3 Numerical Experiments On the basis of the developed approach, studies about the efficiency of industrial production have been conducted. We use the data about functioning of machine-building enterprise in Ufa city particularly about product group that are part of the industrial and technical purposes providing 35% of product output. Step 1. To assess the financial state of the enterprise, first, we carry out preliminary calculations related to the division of costs into fixed and variable parts. This technique is well known and used in management accounting. In accordance with the costing used at the enterprise, the fixed costs were classified as: shop expenses, works general expenses, special expenses, and business expenses. The variable costs include: raw materials, fuel and energy, purchased semi-finished products, services, direct wages, social security contributions. The values of indicators like p Cf q
cv and cv were calculated and, the levels of efficiency a were determined for the product group “engines”, Table 1.

Table 1 Actual (initial) efficiency of “engines” production Product

Year

Cf q
cv

p cv

p cv

Engine 1

2015 2016 2017 2015 2016 2017 2015 2016 2017

1.41 1.37 0.77 1.54 1.49 1.42 0.87 0.86 0.59

1.56 1.59 1.48 1.36 1.38 1.46 1.15 1.20 1.34

0.15 0.22 1.42 −0.19 −0.11 0.42 0.28 0.35 1.35

Engine 2

Engine 3

Cf  q
c v

Efficiency a

Profit, rub.

−0.35 −0.33 0.21 −0.47 −0.45 −0.24 −0.38 −0.35 0.22

−2,514.653 −3,578.478 1,290.508 −4,672.934 −7,315.377 −892.620 −1,117.309 −2,140.782 387.950

434

E. V. Orlova

As given in Table 1, data indicate that the profitability of the enterprise in almost all product types is in the negative zone. To improve the financial condition of these products, it is necessary to bring its output first to the break-even point and then to move into the area of profitability. The analysis shows that in 2015–2017 years existing fixed and variable costs, the wholesale price of a single product and the volume of output has no profit from the sale. The level of profitability in 2015 was (−0.35), and in 2016 (−0.33). In the 2017 year, the enterprise entered the profitability zone by increasing the wholesale price of products, as well as reducing costs level (both fixed and variable), and as a result in 2017, the profitability of “engine 1” increased to 0.21. The calculation results of profitability levels, as well as the distribution of output Cf for each type of product by areas of profitability and loss ratio cpv  q
c allow to v conclude about the positive efficiency of production of “engine 1” and “engine 3” in the 2017 year, the second engine demonstrates positive dynamics of profitability and requires urgent measures to bring to break-even production, as the demand for this type of product is significant high. Step 2. Identification of the product life cycle stage. The quarterly dynamics of sales, costs, prices, and profits from the standpoint of the life cycle of the “engine 1” is shown in Fig. 1. Trends for each indicator were constructed. All trend models are statistically significant, and the approximation accuracy is high. The type and nature of the dynamics of the analyzed indicators described by econometric models allows to determine the stage of the life cycle of a particular product. Figure 1 demonstrates that in 2015–2017 years demand for the analyzed product is growing, production costs are reduced by an average of 10%, and the price increases on average for the period by 3%. In aggregate, this leads to an increase in profits. It can also be concluded that this product is currently at the “stability” life cycle stage. This analysis allows to conclude that it is expedient to apply measures

Fig. 1 Indicators (volume, unit costs, price, profit) for the product “Engine 1”

Approach for Modeling and Situational Management …

435

to promote this product to the market, increase production volumes, increase the wholesale price (taking into account the demand-price relationship) at the current level of costs. Step 3, 4. At the next steps, four possible situations of deviations of the parameters of the external and internal environment factors affecting the level of production efficiency are considered. In order to maintain the current effectiveness, we propose the appropriate management decisions. The efficiency results of these decisions are also given in Table 2. Calculations show that if the fixed costs of the product increase by 5%, then in order to return to the previous profitability level of these products, it is necessary to reduce the specific variable costs by 5.23%. If the cost of resources increases, that is, the variable costs change; in this situation, it is possible to increase the wholesale sales price by 2.44%. In this case, such a price growth is acceptable since this product is at the life cycle stage like “stability”. The economic effect of the operational management decision implementation for the entire product group is given in Table 3. The economic effect is achieved by monitoring the economic situation and by eliminating the negative influence of factors. The effect is calculated as the profit increment before and after decisions making.

Table 2 Situation, decisions, and their efficiency Situation

Decision (controlled parameter)

1. Fixed costs growth at 5%

To decrease unit costs

Controlled parameter value Rub. −63.73

Profit growth

%

Rub.

%

−5.23

159,327.34

114 113

2. Resource costs growth at 5%

To increase price

73.58

2.44

152,397.23

3. Demand reducing at 5%

To decrease fixed costs

−159,327.34

−5.00

223,852.77

121

4. Price reducing at 5%

To decrease fixed costs

−311,724.58

−9.78

398,598.33

145

Table 3 Economic effect from supposed decisions, rub Product

Situation 1

2

3

4

Engine 1 Engine 2 Engine 3 Total

159,327.34 108,159.36 32,493.05 299,979.75

152,397.23 76,421.64 55,209.47 284,028.34

223,852.77 63,528.36 51,890.53 339,271.66

398,598.33 118,574.37 115,939.86 633,112.56

436

E. V. Orlova

4 Conclusion The new conceptual approach for enterprise performance management system development is proposed. The proposed approach for enterprise product efficiency modeling and control differs from similar ones in that takes into account properties and complex nonlinear relationships between economic parameters of internal and external environment. We have constructed the new conceptual approach for design an enterprise’s management system under conditions of high-speed change factors. This approach is based on the application: econometric methods for analyzing and identifying the current product life cycle stage; balanced compensating control methods for stabilizing disturbances. The developed approach provides an effective analysis tool for monitoring and management for product economic efficiency. The method of situational analysis allows, based on an assessment of the deviation from the normal level of product profitability, to form multivariate combinations of economic parameters (output, fixed and variable costs, price) that ensure a given level of efficiency in the production and marketing system of enterprise. The advantages of proposed approach include quick information obtaining, implementation of situational methods in decision-making procedures. Application of the developed approach makes it possible to respond to external disturbances and internal changes of system factors, to control their deviation from the normal level, to form an optimal product structure based on target efficiency.

References 1. Burkov VN (2012) Theory of control of organizational systems and other management science organizations. Control problems 4:2–10 2. Novikov DA (2012) Control theory in the organizational systems. Moscow, Publishing House “FIZMATLIT” 3. Sukharev OS (2009) The theory of economic efficiency. Finance and Statistics, Moscow 4. Pospelov DA (1986) Situational management. Theory and practice. Science, Moscow 5. Tirole J (1993) Theory of industrial organization. MIT Press 6. Gontareva IV, Nizhegorodtsev RM (2012) System efficiency of the enterprise: essence, factors, structure. VSEI, Moscow-Kirov 7. Prangishvili IV (2000) The systems approach and system-wide patterns. Moscow, Publishing house “SINTEG” 8. Kleiner GB (2013) The system platform for the development of the economy of modern economic theory. Probl Econ 6:4–12 9. Livshits VN (2014) On unsteady Russian transition economy. Control Probl Theory Pract 2:8–13 10. Auerbach AJ (2005) Dynamic scoring: an introduction to the issues. Available via DIALOG. http://www.aeaweb.org/annual_mtg_papers/2005/0107_1430_1304.pdf 11. Orlova EV (2017) Modeling and coordinated control for the production and economic system. In: Proceedings of the mathematical modeling session at the international conference information technology and nanotechnology (MM-ITNT 2017) vol 1904, pp 1–6

Approach for Modeling and Situational Management …

437

12. Orlova EV (2019) Model for optimal operational control of financial resources distribution in a company. Computer research and modeling 13. Orlova EV (2016) Concept for industrial and economic systems management based on criteria coordination of interested agents. Program Eng 2:86–96. https://doi.org/10.17587/prin.7.86-96 14. Orlova EV (2016) Modeling and decision support for the firms’ pricing policy under a chaotic dynamic of market prices. In: Proceedings of the workshop on computer modeling in decision making (CMDM 2016) vol 1726, pp 81–88 15. Orlova EV (2017) Control over chaotic price dynamics in a price competition model. Autom Remote Control 1:16–28 16. Orlova EV (2016) Simulation model for the firms’ financial resource management. In: Proceedings of the 28th international business information management association conference on vision 2020: innovation management, development sustainability, and competitive economic growth, IBIMA, pp 1317–1321 17. Orlova EV (2017) Technology for control an efficiency in production and economic system. In: Proceedings of the 30th international business information management association conference on vision 2020: innovation management, development sustainability, and competitive economic growth, IBIMA, pp 811–818 18. Orlova EV (2017) Synergetic approach for the coordinated control in production and economic system. In: Proceedings of the 30th international business information management association conference (IBIMA). Vision 2020: sustainable economic development, innovation management, and global growth. Spain. Madrid, pp 704–712

Modeling of Try-Out and Reliability Estimation of Technological Processes of Machine-Building Products V. M. Trukhanov, M. P. Kukhtik and A. M. Makarov

Abstract Materials about try-out of technological processes of machine-building products have been represented in the article. Theoretical and practical issues about the development of mathematical models and techniques of estimation of reliability of technological processes have been stated. The first model of technological process mastering has been considered, and the formula of intensity of removal of failure causes and analogous formula after the second stage of mastering have been represented as well. According to the theory of Markov processes, formula of estimation of probability of absence of failures at increase of try-out time t ! ∞ after the second stage of mastering has been represented. The technique of reliability estimation on the second model has been represented which is based on the multi-stage procedure of technological process mastering. The formulas for estimation of reliability index have been represented when N types of failures occur at technological try-out of products. The models of curves of reliability growth at try-out of technological processes have been represented when failures revealed in the process of technological try-out obey different distribution laws. Keywords Reliability

 Technological try-out  Reliability estimation  Failures

1 Introduction Technological try-out of series of products and complex assemblies comes to the end fast enough. One or two technological reprocessings of principle lead to a desired result, i.e., product can be considered as practically mastered and ready to batch production. It is characteristic of products composed from typical, well-mastered constructive and technological units and other assemblies. There is tried-out normative–technical documentation supported by vast enough scientific and experimental researches for these assemblies [1–5]. V. M. Trukhanov (&)  M. P. Kukhtik  A. M. Makarov Volgograd State Technical University, 28 Lenin Av., Volgograd 400005, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_47

439

440

V. M. Trukhanov et al.

The process of technological try-out of products is delayed if original assemblies are used in it along with typical, well-mastered constructive and technological units. It is connected with that normative–technical documentation that has been tried out on original assemblies not in full measure; therefore, it is necessary to master new technological processes. Sometimes the development of complex products using typical constructive and technological assemblies is not quite justifiably ignored, and original concepts are used inappropriately which results in serious retardation of stage of batch mastering of product delivery. With reference to these situations, corresponding models of reliability estimation and maturity of products are used at stage of batch mastering and delivery [6–10]. Problem of modeling of maturity and reliability estimation of technological processes become relevant for development of complex technical products [11–15]. Tasks for development of mathematical models of maturity and reliability estimation of technological processes are stated for solving of this problem. Theoretical part has been represented by two models of technological process of try-out. Practical part has been represented by technique of reliability estimation on second model. Developed technique has been found practical using series of enterprises of machine-building industry, e.g., on FGUP « PO « Barrikady » (Volgograd), Artillery Plant no. 9 (Yekaterinburg).

2 Mathematical Models of Trying Out of Technological Processes First model is based on theory of Markov processes like “destruction and reproduction.” Second model is based on known theorem of probability theory—theorem of hypotheses. This model consists in considering of two cases of testing outcome —reprocessing of technological process and construction of product changed its reliability and reprocessing did not changed its reliability. At last, third model is based on theory of autocatalytic curves (or curves of reliability growth). These three models cover practically all encountered cases of quantitative estimation of reliability and maturity of complex products and their assemblies. Reached level of reliability index of product can be accepted as quantitative measure of technological process trying out [16]. Two testing stages are provided for in first model of technological process mastering. Let testing of products delivered on one technological process are carrying out at first stage and m1 failures of constructive and technological character are getting during the testing time t1. Intensity of removal of causes of failure occurrence is determined by formula l1 ¼ m1 =

m1 X i¼1

ti ;

ð1Þ

Modeling of Try-Out and Reliability Estimation …

441

where m1—total number of failures registered at first testing stage during the time t1 ¼

m1 X

ti ;

ð2Þ

i¼1

where ti—time of removal of concrete cause of failure. At second testing stage, occurrence of m2 failures is possible after removal of failure causes of constructive and technological character registered at first stage. Time is spent for removal of m2 failures: t2 ¼

m2 X

tj :

ð3Þ

j¼1

Intensity of second removal of causes of failure occurrence is l2 ¼ m2 =

m2 X

tj :

ð4Þ

j¼1

At first testing stage, intensity of failure occurrence is expressed by dependence k1 ¼

Dm1 ; Dt1 ðn1  m1 Þ

ð5Þ

where Dm1—number of failures occurred during the time Dt1; n1—total number of tests at first stage. At second stage, intensity of failures k2 ¼

Dm2 ; Dt2 ðn2  m2 Þ

ð6Þ

where Dm2—number of failures occurred during the time Dt2; n2—total number of tests at second stage. According to theory of Markov processes at increase of try-out time t ! ∞ after second testing stage estimation of probability of absence of failures is determined by relation: ^q ¼

1 þ k1 =k2 : 1 þ k1 ð1=l1 þ 1=l2 þ 1=k2 Þ

ð7Þ

When value q^ of estimation preset in technical conditions is reached, it is possible to consider that technological process has been tried out and product has been mastered for its batch delivery. Technique of reliability estimation on second model is based on many stage procedure of technological process mastering [17]. Let m1 failures are fixed

442

V. M. Trukhanov et al.

according to results of n1 tests at first stage of technological process mastering, after that reprocessing of technological process directed to removal of failure causes is realized. Then at second stage of technological process mastering, n2 tests of products are carried out and m2 failures are fixed, and reprocessing of technological process is again realized. Such iterative process of stepped try-out continues until causes of failure occurrence will be removed. Process of operational development of technological process is realized for all types of failures in that way. Reliability index of product is estimated on scheme «success-failure» after each stage of technological process try-out: b ¼ 1  m=n; P Pc ¼

1  v2c ð2mÞ ; 2n  m þ 0.5v2c ð2mÞ

ð8Þ ð9Þ

where m and n—number of failures and tests correspondingly at each stage of technological process try-out; v2c ð2mÞ—quantile of v2 distribution determining on Table 6 of application [16] for value m and level aða ¼ cÞ of confidence probability; Pc —low confidence bound of estimation (8) determining with level c of confidence probability. After first stage of technological process try-out, estimation of reliability index and its low confidence bound are calculated using Formulas (8) and (9) in which value m1 and n1 are substituted instead of m and n correspondingly. After second stage of technological process try-out, issue about possibility of adding of m1 and n1 information to m2 and n2 information is stated. After third stage of try-out, issue about possibility of adding of previous testing information to m3 and n3 information is stated, etc. Constant rising of information allows estimating reliability index of product with more and more high accuracy.

3 Mathematical Model of Reliability Estimation of Technological Processes It is possible to set down on the basis of full probability formula for exhaustive events according to hypotheses H0 and H1 [18] Pð AÞ ¼ PðH0 ÞPðA=H0 Þ þ PðH1 ÞPðA=H1 Þ ;

ð10Þ

where A—event consisting in nonfailure operation of product after carried out reprocessing of technological process; H0—hypothesis about that reprocessing of technological process did not change reliability of product; H1—alternative hypothesis about that reprocessing of technological process changed reliability of product.

Modeling of Try-Out and Reliability Estimation …

443

b ðA=H0 Þ and P b ðA=H1 Þ of probabilities P ðA=H0 Þ and P ðA=H1 Þ are Estimations P determined using Formula (8). It is obvious that m = m1 + m2 and n = n1 + n2 at estimation of P ðA=H0 Þ after second stage of technological process try-out, and b ðA=H1 Þ: After second stage of technological process try-out, m = m2, n = n2 for P b estimation P 2 ð AÞ of probability P2 ð AÞ is calculated on the basis of Formula (10):      m1 þ m2 m1 b b b P 2 ð AÞ ¼ P 0 1  þ 1  P0 1  ; n1 þ n2 n1

ð11Þ

b 0 —estimation of probability P0 of null hypothesis realizing. where P It is obvious that P(H1) = 1 − P(H0). Analogously, we will get for third stage of technological process try-out:      b 3 ð AÞ ¼ P b 0 1  m1 þ m2 þ 1  P b 0 1  m2 ; P n1 þ n2 n2

ð12Þ

and for k-th stage of technological process try-out      mk1 þ mk mk1 b b b P k ð AÞ ¼ P 0 1  þ 1  P0 1  : nk1 þ nk nk1

ð13Þ

In the general case, if N types of failures occur at technological try-out of products, then calculating formulas for reliability index looks like [19]: b¼ P

N Y

b ki ; P

ð14Þ

i¼1

vffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi3 u N  2 uX 4 b Pc ¼ P 1  t 1  Pci 5 ; 2

ð15Þ

i¼1

b i of level c calculated using where Pci —low confidence bound of estimation P Formula (9). b 0 of probability P0 playing a part of weighting coefficient can be Estimation P determined by exact or approximated formulas. Exact formula is based on hypergeometric distribution Pm P0 ¼

Cnr 1 Cnmr 2 ; Cnm

r¼m1

! ! mr m n! ¼ ðmrÞ!ðnn22m where Cnr 1 ¼ r!ðnn11r Þ! ; Cn2 þ r Þ! ; Cn ¼ m!ðnmÞ! ; m ¼ m1 þ m2 :

ð16Þ

444

V. M. Trukhanov et al.

Poisson’s approximation is rightly at condition mi nmi  1; Cnmi i  i ; ni mi !

ð17Þ

at that P0 ¼

m X

 Cmr

r¼m1

n1 n1 þ n2

r 

n2 n1 þ n2

mr :

ð18Þ

Normal (Gaussian) approximation of binomial law gives expression 2   3 1 1 6 q1  2n1  q2  2n2 7   7 P0  U6   5 ¼ U Up ; 4 rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi qð1  qÞ n11 þ n12

ð19Þ

where q1 = m1/n1; q2 = m2/n2; q = (m1 + m2)/(n1 + n2); Up—quantile of normal distribution function (Table 1 of application [16]). Formula (19) is rightly for q1 > q2, and indices 1 and 2 in formula change places at q1 < q2. Represented formulas for estimations of reliability and maturity of technological processes are rightly at number of stages around ten. Theory of curves of reliability growth is used for greater number of stages of technological process try-out. The most widespread mathematical model of curve of reliability growth (curve of technological process trying out) is model based on accumulation of number of failures depending on number of tests [20]: m ¼ mk

1e 1  eank

an ;

ð20Þ

where m—accumulated number of failures; mk—finite number of accumulated failures; a—coefficient considering raising rate of number of failures (rate of operational development of technological process); n—accumulated number of tests, nk—finite number of tests. Relation (20) is differentiated for plotting of curve of reliability growth, and current value of reliability index is determined by formula

ean P ¼ 1  m ¼ 1  mk ; 1  eank 0



ð21Þ

where P—current value of reliability index; m0 —frequency of failures; m0 ¼ m=n. At n = 0, initial value of reliability index as quantitative measure of technological process trying out is calculated using equation

Modeling of Try-Out and Reliability Estimation …

P0 ¼ 1 

amk : 1  eank

445

ð22Þ

Finite value of reliability index of product at n = nk Pk ¼ 1  amk

eank : 1  eank

ð23Þ

Necessary number of tests can be obtained from Relations (22) and (23) at preset values of variables a, P0, and Pk nnec

  1 1  P0 ¼ ln ; a 1  Pk

ð24Þ

where P0 and Pk—initial and finite values of reliability index of product correspondingly.

4 Results and Discussion Mathematical models of try-out of technological processes of machine-building products and their reliability estimation have been developed. Technique of reliability estimation of technological process has been represented. This technique has been introduced on the enterprises of machine-building industry. Acknowledgements The reported study was funded by RFBR according to the research project № 17-01-00018.

References 1. Mokrushin YA (2006) Ehkonomichnost’ sushchestvuyushchej sistemy otrabotki na tekhnologichnost’ izdelij mashinostroeniya (Economical efficiency of existing system of processing for the technology of engineering products). Vestn Udmurt Univ Ser “Ehkonomika i pravo” 2. https://cyberleninka.ru/article/n/ekonomichnost-suschestvuyuschey-sistemy-otrabotki-natehnologichnost-izdeliy-mashinostroeniya. Accessed 26 Oct 2018 2. Borisov VM, Borisov SV (2013) Otrabotka izdelij na tekhnologichnost’ (Try-out of products on technological effectiveness) Vestn Kazan Tekhnol Univ 3. https://cyberleninka.ru/article/n/ otrabotka-izdeliy-na-tehnologichnost. Accessed 26 Oct 2018 3. Zhang R, Mahadevan S (2001) Integration of computation and testing for reliability estimation. Reliab Eng Syst Saf 74(1):13–21 4. Wang K-S, Tsai Y-T (2007) Reliability life estimation of products using accelerated life-testing. J Chin Soc Mech Eng Trans Chin Inst Eng Ser C [Chung-Kuo Chi Hsueh Kung Ch’eng Hsuebo Pao] 28(1):35–43

446

V. M. Trukhanov et al.

5. Jia Z, Cai J, Liang Y, Han C (2009) Research on reliability estimation and performance prediction based on step-up-stress accelerated degradation testing. In: Proceedings of 9th international conference on electronic measurement and instruments (ICEMI) 5274869:1302– 1306 6. Liu T (2011) Reliability estimation for Weibull log-linear accelerated life testing model. Xitong Fangzhen Xuebao/J Syst Simul 23(1):29–32, 37 7. Chaturvedi A, Kumari T (2017) Estimation and testing procedures for the reliability functions of a general class of distributions. Commun Stat Theory Methods 46(22):11370–11382 8. Nikolaev YN (2016) Aspects of assessment of organizational and technical reliability and design of construction processes with target reliability level using computer-aided technologies. In: Proceedings of 2nd international conference on industrial engineering, applications and manufacturing (ICIEAM), Chelyabinsk, 19–20 May 2016 9. Burkov AF (2017) Main provisions of theory of reliability in relation to shipboard electrical equipment. In Proceedings of international conference on industrial engineering, applications and manufacturing (ICIEAM), St. Petersburg, 16–19 May 2017 10. Desyatirikova EN, Chernenkaya LV, Mager VE (2018) Method of calculating a priori reliability of elements of complex system. In: Proceedings of international Russian automation conference (RusAutoCon), Sochi, Russia, 2018, pp 1–4 11. Birolini A (2017) Reliability engineering: theory and practice. Springer, New York, London 12. Kapur KC, Pecht M (2014) Reliability engineering. Wiley, Hoboken, NJ 13. Kostina M, Karaulova T, Sahno J, Maleki M (2012) Reliability estimation in manufacturing processes. J Achiev Mater Manuf Eng 51(1):7–13 14. Trukhanov VM, Kukhtik MP (2018) Diagnosticheskie metody rascheta proektnoj nadezhnosti i bezopasnosti dorogostoyashchikh tekhnicheskikh sistem (Diagnostic methods of calculation of design reliability and security for expensive technical systems). Kontrol’. Diagnostika 8 (242):30–33 15. Trukhanov VM, Kukhtik MP (2018) Issledovatel’skie ispytaniya po dostizheniyu i obespecheniyu nadezhnosti podvizhnykh ustanovok special’nogo naznacheniya na ehtape ehksperimental’noj otrabotki (Investigation testing on achievement and securing of reliability of movable mountings of special purpose at stage of experimental try-out). Handbook. An Engineering Journal 6(255):6–14 16. Trukhanov VM (2017) Nadezhnost’ v tekhnike (Reliability in technics). Izdatel’skij dom «Spektr», Moscow 17. Trukhanov VM, Matveenko AM (2016) Nadezhnost’ slozhnykh sistem na vsekh ehtapakh zhiznennogo tsikla (Reliability of complex systems at all stages of the lifecycle). Izdatel’skij dom «Spektr», Moscow 18. Trukhanov VM, Tarnaev AG (2016) Nadezhnost’ i diagnostika slozhnykh sistem (Reliability and diagnostics of complex systems). Izdatel’skij dom “Spektr”, Moscow 19. Trukhanov VM, Klyuev VV (2014) Nadezhnost’, ispytaniya, prognozirovanie resursa na ehtape sozdaniya slozhnoj tekhniki (Reliability, testing, prognosing of resource at stage of creation of complex machinery). Izdatel’skij dom “Spektr”, Moscow 20. Apollonov IV (ed) (1989) Kachestvo i nadezhnost’ v proizvodstve (Quality and reliability in manufacturing). In: Nadezhnost’ i ehffektivnost’ v tekhnike (Reliability and effectiveness in technics), vol 7. Mashinostroenie, Moscow

Methods of Cutting Cost Minimizing in Problem of Tool Route Optimization for CNC Laser Machines A. F. Tavaeva, A. A. Petunin and E. G. Polishchuk

Abstract This paper deals with the optimization problem of cutting tool routing for CNC laser machines arising during the treatment of sheet workpiece. The cutting time and cost are optimization criteria of objective function tending to a minimum. One of the most methods of cutting cost and time minimization is the application of special cutting techniques. Ones are classified into three main classes: standard cutting, multi-contour, and multi-segment cutting. The multi-contour cutting technique is of the greatest interest due to significant reducing values of the piercing numbers, air motion length and in some cases the length of cutting tool motion with reduction of cutting time and cost. But existing techniques do not always provide a reduction of optimization criteria. Consequently, the development of special cutting techniques for some geometrical types of parts is a relevant problem today. In this paper, the special technique for circle parts common in blank production of engineering industry is developed. Based on the proposed technique, the algorithm for automatic construction of a cutting tool route was developed for CAM “SIRIUS.” The tool route and choice of piercings are automatically constructed by taking into account the condition of thermal deformation reducing. As the practice shows the application of the proposed cutting technique for circle parts decreases piercing numbers, values of air motion length and consequently, a cutting cost by 95–98 and 15% respectively depending on the presence/absence of internal contours in circle parts, presence/absence of other types of parts as a result of the nesting, type and thickness of material. Keywords Thermal cutting


Laser cutting
Cutting cost
Cutting time

A. F. Tavaeva (&)
A. A. Petunin
E. G. Polishchuk Ural Federal University named after the first President Russia B.N. Yeltsin, 19, Mira street, Ekaterinburg 620002, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_48

447

448

A. F. Tavaeva et al.

1 Introduction The treatment of sheet material in the blank production of engineering industry is carried out by CNC sheet cutting machines. As a rule, such machines include thermal (plasma, laser, and oxy-fuel) and water jet sheet cutting machines. Due to the fact that laser cutting has a high accuracy and performance the application of CNC laser cutting machines is common. The CNC cutting machines work by NC programs that are developed by CAM systems. During the development of NC program, the optimization problems of cutting tool path are arisen [1, 2]. In [3, 4], the following classes of cutting tool routing problems for CNC sheet cutting machines are allocated: TSP [3, 5, 6], GTSP [7, 8], CCP [9], ECP [10], ICP [11]. The cutting time Tcut and cost Fcost are numerical characteristics that define the effectiveness of NC program for set of parts [1]. In order to optimize cutting tool path, the Tcut and Fcost should strive for minimum. The main parameters influencing on the cutting time and cost are pierce points, lengths of cutting, and air tool motions which assigned during the construction of tool path. In order to minimize Tcut and Fcost , the numbers of pierce points, lengths of cutting, and air tool motions need to be optimized. The cutting techniques are used in order to solve this problem. The cutting techniques are classified into three main classes: standard cutting, multi-contour cutting, and multi-segment cutting technique [12]. Every contour is cut with pierce point by using the standard cutting technique. The numbers of pierce points equal numbers of contours. The several contours are cut in one cutting segment with one pierce point by using multi-contour cutting techniques. For example, the multi-contour cutting includes “chained” cutting [13, 14] and common cut [15]. The several cutting segments are cut with several pierce points by using the multi-segment technique. The considered techniques applied for various types of parts and their effectiveness for different types are varied. In some cases, the effectiveness reaches the minimum value. For this reason in order to reduce Tcut and cost Fcost , the development of special cutting techniques is actual problem today. The circle parts are often found during sheet-working production of engineering industry. The treatment of circle parts without additional cut is impossible with the help of the existing multi-contour cutting techniques. Due to this fact, the special cutting technique for circle parts is proposed in this article.

2 The Construction of Cutting Tool Route for Circle Parts The circle parts are often found during the sheet-working production in engineering industry. The treatment of circle parts without additional cutting is impossible with the help of the existing multi-contour cutting techniques. Due to the fact, the development of special cutting techniques for circle parts is actual problem today.

Methods of Cutting Cost Minimizing in Problem of Tool Route …

449

The proposed cutting technique is based on principle of part combined into “blocks.” The “block” is a set of parts whose positions are fixed relative to each other. In the process of nesting, the “block” is considered as one part; that is, all transformations by displacement/rotation are performed simultaneously with all the parts included in the “block.” After the receiving of nesting results, the cutting tool route is assigned for “block.” The treatment of parts from “block” is carried out by cutting with one pierce point. The motion from one “block” to another is carried out by air. Based on mentioned above-blocking method, the multi-contour cutting technique for circle parts without additional cuts between parts is proposed [16]. The application of one allows to decrease the cutting time and costing by reducing the value of pierce points numbers, lengths of air, and cutting tool motions. The proposed cutting technique (Fig. 1b) compares with standard cutting technique (Fig. 1a). The cutting tool from initial point O moves on air to pierce point No. 1, then first circle contour is completely cut by cutting moving and then tool is switching off and moves on air to pierce point No. 2 and etc., until the third contour is cut (Fig. 1a). Figure 1b shows the scheme of special cutting technique for circle parts with one pierce point and without additional cut. After cutting of Sect. 1 the tool cuts Sect. 2 changing the direction of contour cutting. When the last contour from “block” is reached the cutting of cone is carried out completely by Sect. 3. Then tool moves to 4 section and cuts it and etc. until the last section of first part are cut. So the three parts are cut by using one pierce point (Fig. 1b). The more numbers of parts are in the “block,” and the more significant reduction of pierce points is (Npt ! 1). Due to the fact that pierce point is at certain distance from contour, and then after piercing in material, the tool moves to contour by cutting motion so-called lead-in. Thus, the reduction of pierce points leads to decrease in the length of cutting tool path. In turn, the value of air tool motion length is significantly reduced too. Consequently, the reduction of pierce point numbers, air and cutting tool motions’ length values leads to significant reduction of cutting time and cost.

Fig. 1 Schemes of standard (a) and special (b) cutting technique for three circle parts

450

A. F. Tavaeva et al.

3 The Strategy of Cutting Time and Cost Minimization 3.1

The Algorithm of Automatic Cutting Tool Route Construction

Based on proposed special cutting technique for circle parts (Fig. 1b), the algorithm of cutting tool route automatic construction is developed. The algorithm consists of following steps: • On the first step, the initial data is received. User selects the rectangle area of parts for which it is necessary to automatically construct the cutting tool route. • On the second step the check of internal contours availability is performed. If internal contours are then ones is cut and go to step 3. Otherwise go to step 3 at once. • The “block” orientation regarding sheet material is defined. Depending on the orientation “block,” the pierce point positions and cutting tool route are varied. • Based on results of “block” orientation definition, the pierce point position calculation for each “block” is performed on the fourth step taking into account the reduction of thermal deformation. During treatment of sheet material by using CNC laser cutting machines, the thermal deformations are arisen. In order to minimize ones during treatment of sheet metal, the heuristics methods were taken into account. The main condition of thermal deformation reduction includes the Part Hardness rule [17]. The pierce points must be selected according to Part Hardness rule in order to minimize the temperature of material in the process of laser cutting. • Based on results of pierce point position calculation, the computation of parts cutting directions for each “block” is performed taking into account the Part Hardness rule. • Based on results received on the previous steps, the automatic construction of cutting tool route for external parts is performed by using special cutting technique for circle parts and standard for others. The received cutting tool route satisfies the Part Hardness rule. Based on proposed algorithm, the program of automatic cutting tool route construction is developed for CAM “SIRIUS.” The cutting tool routing and choice of pierce points are automatically constructed taking into account Part Hardness rule. The evaluation of developed program effectiveness for cutting tool routing taking into account the thermal deformation reduction was performed by using CAE” “RAPID” [18]. The received results are shown in Sect. 3.2.

Methods of Cutting Cost Minimizing in Problem of Tool Route …

3.2

451

The Calculation of Thermal Field During Treatment of Sheet Material by CNC Laser Cutting Machines

In order to evaluate effectiveness of developed algorithm in terms of thermal deformation minimization, the CAE “RAPID” is used. The software calculates the thermal fields during treatment of sheet metal. In order to evaluate the effectiveness of cutting tool route construction (pierce points and cutting direction selection) for circle parts using proposed algorithm (shown in Sect. 3.1) in terms of thermal deformations reduction during laser cutting of sheet metal, the three examples were considered. The “RAPID” allows to review the changing of thermal field during cutting process. Figure 2 shows result of nesting and cutting sequence for 8 blocks of circle parts received by using CAM “SIRIUS” for which the cutting tool routes are constructed by using multi-contours techniques. The cutting tool route for circle parts is constructed by using proposed algorithm. Figure 2 shows the example of pierce points and cutting directions selects for received result of nesting. The results of thermal field calculation for the first four blocks are shown in Figs. 3 and 4. The calculation was performed for AWAIMg3 Δ = 5 mm for which the cutting tool speed is Von ¼ 17 mm/s, air tool speed is Voff ¼ 830 mm/s, and capacity is P ¼ 2900 W. For the first block, the piercing is selected by using Part Hardness rule (Fig. 3a), i.e., the end of cutting should happen near “hard” (not cut yet) sheet zone. The average temperature of selected window near point of tool

Fig. 2 Result of nesting and cutting order for 8 blocks of circle parts complying with Part Hardness rule

452

A. F. Tavaeva et al.

Fig. 3 a Temperature fields complying with Part Hardness rule for first part block; b the temperature fields without complying with Part Hardness rule for second block

Fig. 4 a Temperature fields without complying with Part Hardness rule for third part block; b the temperature fields without complying with Part Hardness rule for fourth block

switching off is 443.27 °C (Fig. 3a). Figure 3b shows the thermal field for case of second block cutting without complying with the Part Hardness rule. The average temperature of selected window near point of tool switching off is 510.23 °C, which is 13% higher than the average temperature in the same area for the first block. Figure 4a shows the thermal field for case of third block cutting without complying with the Part Hardness rule. The average temperature of selected window near point of tool switching off is 555.14 °C, which is 20% higher than the average temperature in the same area for the first block other conditions being equal. Figure 4b shows the thermal field for case of fourth block cutting without complying with the Part Hardness rule. The average temperature of selected window near point of tool switching off is 535.9 °C, which is 17% higher than the average temperature in the same area for the first block other conditions being equal.

4 Computational Results The reduction values of piercing numbers, length of cutting, and air tool motion lead to cutting time and cost decreased. In order to evaluate of proposed algorithm application effectiveness, the cutting cost is calculated by using mathematical model

Methods of Cutting Cost Minimizing in Problem of Tool Route …

453

Table 1 Results of cutting cost and time calculation using standard and special cutting technique Type of material

Cutting technique

Lon , m

Loff , m

Npt

Fcost , rub

%

AWAIMg3, Δ = 2

Standard Special Standard Special

54.8 54.3 54.8 54.3

28.7 1.4 28.7 1.4

73 6 73 6

1311.8 1031.6 20,336.7 18,013.0

21

AWAIMg3, Δ = 5

12

proposed in [19, 20]. The example of proposed cutting technique application using developed program in CAM “SIRIUS” is considered below. The result of nesting is performed by CAM “SIRIUS.” Then the cutting tool routes are constructed by using standard without complying with Part Hardness rule and special cutting techniques taking into account thermal deformation reduction for nesting shown in Fig. 2. The received results are proposed in Table 1. The calculations were performed for AWAIMg3 with Δ = 2 and 5 mm. The received results showed that during treatment parts using special cutting technique, the cutting cost Fcost is reduced by 21% compared with standard cutting method. When special cutting technique for circle parts is used, the numbers of pierce points and values of air tool motion length are reduced by 92 and 95%, respectively, compared with standard cutting technique. The value of cutting motion length is reduced by using special cutting techniques because of numbers of pierce points are reduced, and consequently, the lead-in is reduced too. Thus, the application of developed program for automatic construction of cutting tool route for circle parts significant reduces the cutting time and cost compared with the application of standard cutting technique.

5 Conclusions Based on received results, the following conclusions were made: • In order to minimize cutting time and cost arisen during laser cutting, the special cutting technique was proposed for circle parts. The circle part is one of the most common geometrical types of parts in engineering industry. • Based on proposed special cutting technique for circle part, the algorithm was developed in order to automatically construct the cutting tool route. Based on this algorithm, the program for CAM “SIRIUS” is worked out. • In order to reduce the thermal deformation arisen during laser cutting, the Part Hardness rule is taken into account constructing the tool route by using developed algorithm. • In order to evaluate the effectiveness of developed algorithm in terms of thermal deformations reduction, the CAE “RAPID” is used. The software calculates the thermal fields during laser cutting of sheet metal. The three examples of various pierce points and cutting directions selected for circle parts are calculated by

454

A. F. Tavaeva et al.

CAE “RAPID” for result of nesting received by CAM “SIRIUS”. The results show that cutting tool route received by developed program based on a special cutting technique for circle parts is more effective in terms of thermal deformation reduction compared with other two cases which has not complied with the Part Hardness rule. • The examples of circle part nesting are evaluated in terms of cutting time and cost reduction. The received results show that the cutting tool route constructed by developed program based on a special cutting technique for circle parts is more effective compared with standard cutting technique application in terms of cutting time and cost decrease.

References 1. Petunin AA (2014) The two problems of cutting tool routing for CNC cutting machines. In: ITIPM’14. 2nd international conference “intelligent technologies for information processing and management”, Ufa, Nov 2014, pp 215–224 2. Petunin AA, Chentsov AG, Chentsov PA (2015) About a routing problem of the tool motion on sheet cutting. J Model Anal Inf Syst 22(2):278–294 3. Dewil R, Vansteenwegen P, Cattrysse D (2016) A review of cutting path algorithms for laser cutters. Int J Adv Manuf Technol 87(5–8):1865–1884. https://doi.org/10.1007/s00170-0168609-1 4. Hoeft J, Palekae U (1997) Heuristics for the plate-cutting traveling salesman problem. IIE Trans 29(9):719–731. https://doi.org/10.1023/A:1018582320737 5. Chen M, Li X, Tang K (2014) Optimal air-move path generation based on MMAS algorithm. Int J Prod Res 52(24):7310–7323. https://doi.org/10.1080/00207543.2014.922713 6. DSouza KCR, Wright P (2003) Toop-path optimization for minimizing airtime during machining. J Manuf Syst 22:173–180 7. Chentsov A (2008) The optimization problems of routing and assignment distribution: the theory questions. Izhevsk 8. Petunin AA, Chentsov AG, Chentsov PA (2013) To the question about instrument routing in the automated machines of the sheet cutting. St. Petersburg State Polithechnical Univ J. Comput Sci Telecommun Control Syst 2(169):103–111 9. Arkin E, Hassin R (1994) Approximation algorithms for the geometric covering salesman problem. Discrete Appl Math 55:197–218 10. Sherif S, Jawahar N, Balamurali M (2014) Sequential optimization approach for nesting and cutting sequence in laser cutting. J Manuf Syst 33(4):624–638 11. Garfinkel R, Webb I (1999) On crossings, the crossing postman problem and the rural postman problem. Networks 34(3):173–180 12. Petunin A (2015) Modeling of tool path for the CNC sheet cutting machines. In: APEE’15. 41st International conference “application of mathematics in engineering and economics”, Sozopol, June 2015, 1690, 060002. https://doi.org/10.1063/1.4936740 13. Verkhoturov MA, Tarasenko PYU (2008) Mathematical provision of problem of tool path optimization at flat shape nesting based on “chained” cutting. Ufa State Univ J Control Comput Sci 2(27):123–130 14. Panukova T (2008) Cover with ordered enclosing for flat graphs. Electron Notes Discrete Math 28:17–24 15. Manber U, Israni S (1984) Pierce point minimization and optimal torch path determination in flame cutting. J Manuf Syst 3:81–89

Methods of Cutting Cost Minimizing in Problem of Tool Route …

455

16. Tavaeva AF (2014) The definition of conditions of the special cutting techniques effective application when designing automatic methods of tool path optimization. In: ITIPM’14. 2nd international conference “intelligent technologies for information processing and management”, Ufa, Nov 2014, pp 225–231 17. Petunin AA (2009) About some strategies of tool routing when designing NC programs applied to thermal cutting machines. Ufa State Univ J Control Comput Sci 2(35):280–286 18. Petunin AA, Polishchuk EG (2016) The calculation of the thermal fields during thermal cutting of sheet metal. In: ITiS’ 16. 5th international conference “information technologies and systems”, Chelyabinsk, Feb 2016, pp 142–144 19. Tavaeva A, Petunin A (2018) The accurate calculation of parts treatment cost from sheet metal ob the CNC laser cutting machine in problem of tool routing optimization. J Model, Optimization Inf Technol. https://moit.vivt.ru/wpcontent/uploads/2018/10/TavaevaPetunin_ 4_18_1.pdf. Accessed 21 Oct 2018 20. Tavaeva AF, Petunin AA (2017) The calculation of laser cutting cost in problem of cutting tool path optimization at CNC machines. Bulleten BGTU 12:194–202. https://doi.org/10. 12737/article_5a27cb8ff243c9.75015147

Development of Hardware-Algorithmic System for ICE Diagnostics L. A. Galiullin and R. A. Valiev

Abstract Recently, the research in the development of methods and tools for the diagnosis of internal combustion engines is conducted in the direction that determines the use of modern technical and information systems. A significant part of the work associated with the study of internal combustion engines in transitional operating modes. At the same time, it is noted that it is difficult to carry out experimental studies related, as already noted, with the high cost of equipment, its low distribution and high labor costs for conducting experiments, and in some cases with insufficient accuracy of measurement and data processing. High labor costs are associated primarily with the need to install the engine in special stands and the use of special sensors. The systems based on the possibilities of self-diagnosis of an internal combustion engine with an electronic control system do not allow obtaining engine characteristics in the whole range of rotational frequencies, i.e., have low information content. Thus, there is a need to develop a system for diagnosing internal combustion engines, which ensures sufficient accuracy and informational content of experimental studies, has a low cost and low labor costs, and allows investigating the engine in various operating modes. Keywords Engine


Diagnostic
Testers
Model
System

1 Introduction In the process of driving, the driver, acting on the controls of the car, strives to withstand the desired and constantly changing the speed of movement, causing the vehicle to change the speed and trajectory of movement [1]. Moreover, the driver seeks to provide a constant opportunity to change the speed of driving, controlling the transmission ratio of the car’s transmission. It is easy to see that not only the L. A. Galiullin (&)
R. A. Valiev Naberezhnye Chelny Institute Kazan Federal University, 18, Kremlyovskaya str, 420008 Kazan, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_49

457

458

L. A. Galiullin and R. A. Valiev

control of the trajectory but also the control of the ride smoothness requires a change and determines the speed of the car [2]. Under these conditions, the ability of a car to change its speed under the influence of control commands formed by the driver acquires fundamental importance. It is obvious that the speed control of the vehicle is carried out by changing the power expended on these goals, produced or absorbed by the engine, and if the power absorbed by the engine is not enough, the vehicle’s brake system is used [3]. It is significant that, ultimately, almost all the properties of a car as a vehicle, including: controllability, stability, smoothness, fuel efficiency, environmental performance, are crucially dependent on the characteristics of the engine and its ability to change its power under the influence of control commands. Having the ability to change the power produced or absorbed by the engine, under the influence of the driver’s commands, is the most important functional requirement for a car engine [4]. Therefore, the purpose of an automobile engine is not only to convert one or another type of energy into mechanical energy but also to provide an opportunity to control this transformation [5]. All this suggests that along with the conversion of chemical energy of fuel into heat, heat into mechanical, in the form of traction or braking power, the engine, including its workflow management system, is the main element that converts driver commands (influences) to changes in speed car. It is necessary to emphasize once again that for modern automobile engines, the ability to transform control commands into changes in engine power is essential [6]. A generalized diagram of the structure of speed control is presented in Fig. 1. It should be borne in mind that the speed control of the vehicle is carried out by changing the effective engine power and vice versa [7]. In turn, the engine power is closely related to the rotational speed of the engine crankshaft and the speed of the vehicle [8]. Many other control parameters (cyclic filling, cyclic fuel supply, etc.) are also associated with the rotational speed of the internal combustion engine since the cyclic control principle is implemented. Leading engine company conducts intensive research and development work to improve the reliability and durability of the internal combustion engine (ICE) [9]. Research and testing of engines are one of the main means of checking the quality of the manufacture of parts and assemblies, subassemblies and engine as a whole, its proper mounting, and compliance of the essential characteristics of an engine with the requirements of technical documentation [10]. Types of engine tests are regulated by GOST and International Standards ISO, which govern the rules of acceptance and requirements to the technical level of engines [11]. After acceptance and putting engines into production, an improvement of their designs and technical and economic indicators continues. Currently, testing of engines is a complex and time-consuming technological process, very similar to the pilot study [12]. Therefore, an engine automated test system (ATS) has been created. Modern requirements to continual improvement of the technical level of manufactured engines lead to the fact that the share of the costs of the tests in the

Development of Hardware-Algorithmic System for ICE Diagnostics

459

Fig. 1 Structure of the vehicle speed control

process of creation of new engine models grows more and more [13]. Particularly large, these costs are in the case of non-compliance of the level of production automation and scientific research one. In this regard, test technological process automation is one of the main objectives of improving the technological level of production and quality of the engines. In order to control the engine during its testing and diagnostics, it is proposed to use the methods of neuro-fuzzy inference, widely applied in the development of intelligent systems [14]. In order to describe engine management, expert knowledge is used instead of mathematical models. The control of internal combustion engines (ICE) based on a neural fuzzy system is founded on a knowledge base containing a number of fuzzy rules automatically generated by a neural network [15]. The filling in of this knowledge base is also possible by direct measurement method, but it is more appropriate to fill in the knowledge base with a neural network.

460

L. A. Galiullin and R. A. Valiev

2 Control Parameters of ICE The most important control parameters of the engine are the so-called primary parameters—cyclic filling with air, cyclic fuel supply (duration of opening of the injectors), and ignition advance angle. As noted earlier, the properties of an internal combustion engine, as a dynamic system, are described by a set of external velocity characteristics—the dependencies of changes in the basic parameters of work on the rotational speed (angular velocity) [16]. The composition of the ESC (external speed characteristics) includes the following characteristics (as a function of rotational speed): fuel consumption; air flow; effective power and torque developed on the engine shaft; and ignition timing. Distinguish full and partial ESC. Full characteristic is obtained while ensuring maximum filling of the engine cylinders with air (maximum fuel supply) at a constant load on the shaft [17], while partial characteristics respectively with incomplete fuel supply. In this case, it is important that the values of the partial characteristics lie inside the region bounded by the values of the total characteristic. It is also necessary to take into account that in order to obtain adequate ESC, a necessary condition is the constancy of the position of the body that controls the fuel supply (throttle valve). The quantity characterizing the fuel supply and the control action from the driver is the degree of throttle opening in percent. In further calculations, we will operate with exactly this value since it is more convenient for perception and interpretation than the angle of rotation in degrees. The choice of the diagnosis mode is reduced to ensuring such an engine operation mode, in which its properties are presented more fully. This mode corresponds to the full fuel mode when the throttle is maximally open [18]. This is primarily due to the widest frequency range of the engine and the maximum work of inertial forces and friction forces. In addition, in real conditions of diagnosis, to ensure the consistency of the position of the governing body (other than 100% of the opening of the throttle valve) is quite problematic. Thus, the mode of operation of the engine with full fuel supply is chosen as the mode of diagnosis [19]. At the same time, it is necessary to develop a mechanism for deciding on sufficient provision of such a regime. The high degree of equipment of modern electronic workflow control systems of the internal combustion engine with primary converters determines the saturation of information flows between the electronic control unit (ECU), sensors, and actuators. The computer on the basis of signals from the primary transducers (sensors) determines the mode of operation of the engine (idling, power mode) and generates controls for the actuators (nozzles, ignition module, etc.) [20]. Based on the composition of the ESC, we can determine the list of signals from sensors and actuators of the engine management system, containing the necessary information for the indirect assessment of each parameter [21]. In this case, information may be contained as in the magnitude of the signal (mainly in voltage), frequency of change, and duration.

Development of Hardware-Algorithmic System for ICE Diagnostics

461

To determine the function of the change of the effective power and torque on the shaft, it is necessary to obtain the function of changing the angular velocity of rotation x = u(t), where x angular velocity, t—time. Such information is contained in several signals, in particular, in the fuel injection control signals and ignition control signals [22]. The use of these signals is impractical since the time of appearance of control pulses depends on the rotational speed of the engine shaft and may vary relative to the position of the top dead center of the cylinders in adjacent cycles (for example, by the amount of change in ignition advance angle, which increases with increasing speed significant calculation error). Such flaws devoid the signal of the crankshaft position sensor. It is a sinusoidal signal, the voltage, and frequency of which is a function of the frequency of rotation of the motor shaft. The electronic control unit maintains a voltage of about 12 V on the injector control line [23]. At a certain point in time, the computer closes the line to “ground,” which forms a negative impulse. Information about the air flow is contained in the signal of the air flow sensor. The output voltage of the sensor is related to the amount of air passed through the sensor. The ignition advance angle is usually measured in degrees of the crankshaft position of the engine and is proportional to the time between the negative pulse front in the ignition control signal and the moment the cylinder is at the top dead center in each cycle. Thus, for the construction of the ESC it is enough to process information from the five signals of the control system.

3 Hardware Design The hardware includes a cable splitter of control system signals, an interface device, a computer, and an analog signal input module. To ensure mobility, a laptop-type personal computer was chosen as a computer. This makes it possible to carry out diagnostics, while the vehicle is in motion, when the load is the mass of the car, reduced to the crankshaft through the transmission. Such a solution allows diagnostics in real operating conditions [24]. At the same time, there are separate requirements for the input device of analog signals. First, it must be compatible with portable computers and therefore support the USB communication interface. Secondly, ensure sufficient accuracy of signal conversion. To ensure the conversion error at the level of 0.05%, the ADC must have a bit depth of about 10 bits. Taking into account the effects of quantization [25], to ensure high accuracy, the ADC should be at least 12 bits wide. To reduce the influence of the measuring system on the operation of the electronic engine control system, the ADC must have a high input resistance (about 1 MX).

462

L. A. Galiullin and R. A. Valiev

Third, have at least five analog inputs and provide a sampling frequency of at least 70 kHz per channel (total 350 kHz). This is due to the fact that the maximum frequency of the CEL signal is 7 kHz. Accordingly, the information component of the signals is concentrated in the frequency band from 0 to 7 kHz, above—the noise. It is known from Kotelnikov’s theorem that in order to reconstruct an analog signal from its discrete samples, it is necessary that the sampling frequency be twice as high as the maximum useful frequency of the signal itself. According to the results of the conducted research, it can be concluded that the highest accuracy is achieved when the ADC operates in the input voltage range of ±10 V, at a sampling frequency of about 100 kHz. With this mode of operation, the output impedance and the capacitance of the signal source have the least impact on accuracy; the smallest inter-channel transmission (effect of a neighboring source) is achieved. Accordingly, the interface device must ensure the conversion of analog signals from sensors and actuators of the engine management system to the maximum allowable input voltage range of the ADC module in the entire frequency range. In the developed system, the interface device is made on passive elements (resistors) and provides the maximum input voltage of the ADC at the level of ±10 V.

4 Development of Software and Algorithmic Parts The software and algorithmic part of the system consists of two parts—a program for managing data collection from the ADC module and a program for processing the collected data. The first program is implemented in the Microsoft Visual Studio programming environment; it loads and controls the operation of the ADC, in particular, provides asynchronous data collection from five channels of the module with a sampling frequency of Fs = 80 kHz per channel. This provides a pseudo-differential connection of signals. First, it must be compatible with portable computers and therefore support the USB communication interface. The second program is implemented by the MATLAB engineering calculation system and is a set of functions that process signals and calculate external velocity characteristics. Preprocessing consists of passing signals through a digital low-pass filter. The filter is a 5th-order digital Butterworth recursive filter with a cutoff frequency of 7 kHz. The need for pre-filtering is due to the high degree of noise in the processed signals. The nature of the noise is high frequency. This may be due to the operation of the ignition system—the formation of a high breakdown voltage on the candle and other random noise. The signal CPS—sinusoidal, the frequency of which is equal to the frequency of rotation of the engine crankshaft, expressed per minute. This is due to the fact that the driver pulley has 60 equally spaced teeth. The spectrum of such a signal

Development of Hardware-Algorithmic System for ICE Diagnostics

463

contains a pronounced component at a frequency corresponding to the frequency of rotation of the engine shaft. Thus, dividing the CPS signal into segments (possibly with overlapping) and applying a discrete Fourier transform to them, we obtain a spectrogram (instant spectrum), where the largest spectral components correspond to the instantaneous frequency of engine rotation. Having calculated the indices of the maximum spectral components in each segment, we determine the instantaneous frequency of rotation of the engine. It is necessary to calculate the rate of change of the frequency of rotation of the engine shaft, that is, the angular acceleration. This operation is reduced to the differentiation of the function of the speed of rotation over time and can be numerically replaced in accordance with the definition of the derivative—the ratio of the increment of the function of the speed of rotation to the increment of time. A well-known disadvantage of this method of calculation, even with a decrease in the interpolation step, is the suppression of low frequencies and a significant increase in high, that is, noise. The use of filtering differentiated functions and derivative functions allows to reduce the level of noise, but still does not provide sufficient accuracy of the conversion. A much more accurate method for determining the component of acceleration from the readings of the function of changing the rotational velocity is the method based on approximation of the function of the speed of rotation by a polynomial, with further symbolic derivation of the derivative with decreasing degree of the polynomial. In addition, the approximation has a filtering property to a certain extent. The approximation of the function of changing the speed of rotation in time is carried out within the minimum and maximum values of the speed of rotation and the corresponding time intervals. One of the features of engines with the electronic control system is the limitation of the maximum rotational speed [26]. When the specified value is reached, the frequency of rotation is limited by the control system and then oscillates around the limiting frequency (cutoff) and in general may slightly exceed this frequency. Thus, the task of determining the time to reach the maximum rotational speed of the internal combustion engine is reduced to the determination of the first local maximum of the function. This search is also necessary to exclude extreme points of the function of rotational speed for values close to the cutoff, which significantly improves the quality of approximation. To exclude phase distortions, the data are filtered first “forward” and then “backward.” This is possible due to the fact that the samples have a limited length. The quality of data processing can be estimated based on the analysis of residuals, that is, the difference between the original and processed data, respectively, for the polynomial and the filter. The next step in the computation is to take the time derivative of the polynomial. The result is a new polynomial of degree n − 1, which characterizes the function of the angular acceleration of frequency. Substituting the values of time into this

464

L. A. Galiullin and R. A. Valiev

polynomial and multiplying by the reduced moment of inertia, we obtain the values of the effective moment developed by the engine. The effective power function is obtained by multiplying the values of the calculated moment by the angular velocity of rotation, at the same points in time. The greater the length of the Fourier transform, the higher the resolution and accuracy of the measurement of the frequency of rotation of the engine. But, on the other hand, the number of overlap points increases, given the number of segments, which leads to an increase in the influence of the previous calculated value of the rotation frequency. To eliminate this effect, the number of overlap points should be no more (or slightly more) than half the length of the discrete Fourier transform. To determine the parameters for calculating the acceleration characteristic in accordance with the algorithm, a study was made of the influence of the Fourier transform length and the number of segments of the signal splitting on the accuracy of restoring the rotational speed of the engine crankshaft from the CPS signal. In this case, as already noted, the output signal of the CPS is a sinusoidal signal with a frequency varying within the minimum and maximum values of the rotational speed of the engine crankshaft. The signal amplitude varies from 20 to 250 V, respectively, for minimum and maximum rotational speeds. The coefficients of a polynomial were taken as coefficients of a polynomial approximating the acceleration characteristic calculated in one of the tests of an intact engine. The signal duration was 1 s, which is typical during engine acceleration, when the inertia of the flywheel serves as a load. The rotational speed varies from about 800–6300 rpm. The signal values were determined in discrete time values from 0 to 1 s in 1/FS increments, where Fs is the selected ADC sampling frequency (80 kHz). The amplitude of the signal varied linearly from 20 to 250 V within the rotational speed range. Further, the signal was divided into overlapping segments of length equal to the length of the Fourier transform. The signal values in each of the segments were multiplied by the window function to suppress the side lobes of the spectrum. For each segment, a discrete spectral function was determined based on the Fourier transform and the frequencies corresponding to the maximum spectral components were calculated. The time corresponding to the calculated frequency was determined as the average time for a segment. In accordance with the algorithm, the computed sets were approximated by a 3rd-order polynomial. Further, the relative error in calculating the rotational speed was determined. For this, the initial values were recalculated at the points in time corresponding to the calculated values of rotational frequencies. The relative error was also determined after the obtained frequencies were approximated by a polynomial. In this case, the values of the Fourier transform length (the length of the overlap segments) and the number of segments were varied.

Development of Hardware-Algorithmic System for ICE Diagnostics

465

5 Conclusion The data show that the accuracy of calculating the rotational speed does not depend on the number of segments for a fixed value of the Fourier transform. At the same time, the accuracy increases with increasing conversion length. Further, the increase in accuracy is reduced. Thus, the Fourier transform length is selected as the values for calculating the acceleration characteristic based on the developed algorithm. Although the number of segments has the least effect on the accuracy of calculations, it has been chosen as the maximum in case the change in the rotational speed is complex, which can be the case with a faulty engine. The need to use window functions is due to the phenomenon of spreading the signal spectrum, due to the fact that the model is performed on a finite interval that does not contain a finite number of signal periods. The multiplication of the signal by the weighting function corresponds to the convolution of the spectra of the signal and the weighting function in the frequency domain. This leads to a decrease in the side lobes of the spectral function. As a window function, the Kaiser function was chosen, providing side lobe suppression at the level of −40 dB. The analysis of methods and tools for diagnosing automotive internal combustion engines, which resulted in the conclusion that most modern methods and tools for diagnosing internal combustion engines do not fully solve the problems of determining the technical condition of an engine, are often time consuming and expensive. The selection of the method and mode of diagnosis of the internal combustion engine on the basis of external speed characteristics, for which a list of sensors and actuators of the engine control system, is determined. The measuring part of the diagnostics system and the information processing algorithms of the sensor signals and actuators of the control system for calculating external velocity characteristics have been developed and implemented. The algorithms are based on the capabilities of fuzzy inference systems for identifying dependencies. An assessment of the influence of external factors on the accuracy of measurements was carried out; with the result that it was found that the maximum reduced error does not exceed 5%. Experimental studies of the metrological characteristics of the diagnostic system on an engine equipped with a control system were carried out, which showed that the relative errors did not exceed the calculated ones. In this case, the ESC was determined in the entire range of engine speeds.

466

L. A. Galiullin and R. A. Valiev

References 1. Shah M, Gaikwad V, Lokhande S, Borhade S (2011) Fault identification for I.C. engines using artificial neural network. In: Proceedings of 2011 international conference on process automation, control and computing, PACC 2011, art. no. 5978891 2. Galiullin LA, Valiev RA (2016) Modeling of internal combustion engines test conditions based on neural network. Int J Pharm Technol 8(3):14902–14910 3. Wei D (2011) Design of Web based expert system of electronic control engine fault diagnosis. In: BMEI 2011—Proceedings 2011 international conference on business management and electronic information, 1, art. no. 5916978, pp 482–485 4. Galiullin LA, Valiev RA (2018) An automated diagnostic system for ICE. J Adv Res Dyn Control Syst 10(10):1767–1772 5. Valiev RA, Galiullin LA, Dmitrieva IS, Ilyukhin AN (2015) Method for complex web applications design. Int J Appl Eng Res 10(6):15123–15130 6. Galiullin LA, Valiev RA (2018) Method for neuro-fuzzy inference system learning for ICE tests. J Adv Res Dyn Control Syst 10(10):1773–1779 7. Zubkov EV, Galiullin LA (2011) Hybrid neural network for the adjustment of fuzzy systems when simulating tests of internal combustion engines. Russ Eng Res 31(5):439–443 8. Galiullin LA, Valiev RA (2018) Modeling of internal combustion engines by adaptive network-based fuzzy inference system. J Adv Res Dyn Control Syst 10(10 Special Issue):1759–1766 9. Galiullin LA, Valiev RA (2018) Optimization of the parameters of an internal combustion engine using a neural network. J Adv Res Dyn Control Syst 10(10 Special Issue):1754–1758 10. Galiullin Lenar A, Valiev Rustam A (2017) Diagnosis system of internal combustion engine development. Rev Publicando 4(13):PR128–PR137 11. Galiullin Lenar A, Valiev Rustam A, Mingaleeva Lejsan B (2018) Development of a neuro-fuzzy diagnostic system mathematical model for internal combustion engines. HELIX 8 (1):2535–2540 12. Galiullin Lenar A, Valiev Rustam A (2017) Mathematical modelling of diesel engine testing and diagnostic regimes. Turkish Online J Design Art Commun 7:1864–1871 13. Galiullin LA, Valiev RA (2017) Diagnostics technological process modeling for internal combustion engines. In: 2017 International conference on industrial engineering, Applications and Manufacturing (ICIEAM) 14. Guihang L, Jian W, Qiang W, Jingui S (2011) Application for diesel engine in fault diagnose based on fuzzy neural network and information fusion. In: 2011 IEEE 3rd international conference on communication software and networks, ICCSN 2011, art. no. 6014398, pp 102–105 15. Galiullin LA (2016) Development of automated test system for diesel engines based on fuzzy logic. In: IEEE 2016 2ND international conference on industrial engineering, Applications And Manufacturing (ICIEAM) 16. Galiullin LA, Valiev RA (2016) Automation of diesel engine test procedure. In: IEEE 2016 2ND international conference on industrial engineering. Applications and Manufacturing (ICIEAM) 17. Galiullin LA, Valiev RA, Mingaleeva LB (2017) Method of internal combustion engines testing on the basis of the graphic language. J Fundam Appl Sci 9SI(1):1524–1533 18. Galiullin LA (2015) Automated test system of internal combustion engines. In: International scientific and technical conference innovative mechanical engineering technologies, equipment and materials-2014, IOP conference series-materials science and engineering, vol 86. pp 012018 19. Valiyev RA, Galiullin LA, Iliukhin AN (2015) Methods of integration and execution of the code of modern programming languages. Int J Soft Comput 10(5):344–347 20. Galiullin LA, Valiev RA (2015) Automated system of engine tests on the basis of Bosch controllers. Int J Appl Eng Res 10(24):44737–44742

Development of Hardware-Algorithmic System for ICE Diagnostics

467

21. Li X, Yu F, Jin H, Liu J, Li Z, Zhang X (2011) Simulation platform design for diesel engine fault. In: 2011 International conference on electrical and control engineering, ICECE 2011— Proceedings, art. no. 6057562, pp 4963–4967 22. Valiyev RA, Galiullin LA, Iliukhin AN (2015) Approaches to organization of the software development. Int J Soft Comput 10(5):336–339 23. Valiyev RA, Galiullin LA, Iliukhin AN (2015) Design of the modern domain specific programming languages. Int J Soft Comput 10(5):340–343 24. Biktimirov RL, Valiev RA, Galiullin LA, Zubkov EV, Iljuhin AN (2014) Automated test system of diesel engines based on fuzzy neural network. Res J Appl Sci 9(12):1059–1063 25. Valiev RA, Khairullin AKh, Shibakov VG (2015) Automated design systems for manufacturing processes. Russ Eng Res 35(9):662–665 26. Yu Y, Yang J (2011) The development of fault diagnosis system for diesel engine based on fuzzy logic. In: Proceedings—2011 8th international conference on fuzzy systems and knowledge discovery, FSKD 2011, 1, art. no. 6019556, pp 472–475

Information Support of Gas-Turbine Engine Life Cycle Based on Agent-Oriented Technology A. Zagitova, N. Kondratyeva and S. Valeev

Abstract The development of a lifecycle support system for a complex technical object that enables information exchange in a shared information environment during the whole lifecycle is highly important. A problem of information flows control generated at various stages of life cycle of gas-turbine engine, using intelligent software agents is considered. The use of agent-based technologies allows increasing information exchange effectiveness and optimally using hardware resources, as well as to reduce time costs due to synchronization of the information exchange procedures in multiagent environment. In this paper, a prototype of multiagent system for neural network approximation of gas-turbine engine compressor performance is suggested. The example of neural network model of map performance of low-pressure compressor is presented. Keywords Lifecycle support network


Gas-turbine engine
Multiagent system
Neural

1 Introduction The period of development and use of complex technical object (CTO), for example, a gas-turbine aviation engine (GTE) is called its life cycle. During the CTO’s life cycle, huge volumes of heterogeneous information are generated, transmitted, and stored. Processing of this information requires lifecycle stages automatization and synchronization of its participants (developers, testers, etc.) interactions [1–6]. GTE life cycle consists of the following main stages: pre-design stage, GTE prototype design stage, developmental testing, certification, testing, serial production, maintenance stage, and operation. So far information systems that would support all the stages of GTE lifecycle, support decision making and information exchange between enterprises and departments participating in A. Zagitova
N. Kondratyeva
S. Valeev (&) Ufa State Aviation Technical University, 12, K. Marx Str, 450008 Ufa, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_50

469

470

A. Zagitova et al.

lifecycle process, within a shared information environment have yet to be developed. Information flows generated at the stages of GTE life cycle are heterogeneous by the type of data, may be asynchronous. Controlling the data exchange processes in the context of high uncertainty requires the use of technologies that can provide sufficient adaptability and autonomy. One of such technologies is multiagent systems based on intelligent software agents. An agent is relatively autonomous software or hardware entities that can act purposefully to reach goals set by the agent’s developer or user, communicate information to each other, and cooperate [7–13]. The system of multiple agents that acts in a coordinated way and communicate is called multiagent system (MAS) [14]. GTE life cycle support systems based on agent-oriented approach can support all the stages of GTE life cycle. MAS can provide an intelligent information processing, automatize some decision-making processes according to criteria and goals set by decision-makers. Also to support the information exchange between all the participants of GTE lifecycle in a standardized way within a shared information environment and help to synchronize their interactions [15–19]. Automatization of information exchange during the GTE life cycle requires connecting all the computing and storage devices involved in the lifecycle process into a shared distributed computation network and deployment of some multiagent platform, such as JADE, on this network. Java Agent Development Environment (JADE) is an open-source framework for development of intelligent agents and multiagent systems [20].

2 Multiagent System of Lifecycle Support The JADE agents are deployed on the containers—runtime JADE environments that can be run on distributed computers of corporation but stay within and unified address space. Containers and agents they contain can be addressed through and IP

Fig. 1 Architecture of MAS for GTE lifecycle support

Information Support of Gas-Turbine Engine Life Cycle Based …

471

address of the device they are being executed on. Agents communicate by exchanging messages in the specific ACL language. Figure 1 shows a general architecture of MAS for GTE lifecycle support. The suggested MAS architecture implies that JADE containers should be created on devices involved into lifecycle support process at all the enterprises and departments of corporation after JADE platform and main container has been started on core secured computer. Participants of the GTE life cycle develop and deploy agents and MAS that would solve relevant problems—process and actualize information that is generated on this enterprise, request necessary information from agents located on other enterprises and provide requested information in their turn. Simulation of GTE includes the simulation of its high- and low-pressure compressors and turbines on the base of map performance tables. Compressor performances can be represented in different ways and the problem of choosing an optimal by accuracy and hardware resources costs method of compressor performances representation is highly important. Literature overview on the problem revealed that methods used for representation of GTE components performances are not adapted for the use as a part of GTE simulation models at the design, test, and exploitation and compressor development stage. For their implementation is required massive of experimental data and different approximation algorithms to satisfy accuracy requirements. For example, neural network approximation of compressor performance based on experimental data has its advantages and disadvantages, most important of which is low accuracy in intermediate points between experimental data [15].

3 Neural Network Model of Map Performance of Low-Pressure Compressor The performance of a turbojet GTE’s low-pressure compressor is analytically expressed as a function of two variables: Gv

n lpc

 ¼ f plpc ; nn

lpc

 ;

ð1Þ

where plpc is the low-pressure compressor (LPC) pressure ratio, Gv n lpc is the normalized airflow through LPC, and nn lpc is the rotor speed of LPC. The performance function is presented as a table of experimentally obtained data in the normalized form (Fig. 2a). The surface plotted on the base of 63 experimental data points obtained as a result of compressor flight and bench tests is presented in Fig. 2b. This performance map can be used in turbojet engine simulation model on the GTE modes from low to maximal thrust. The artificial neural network (ANN) developed for this performance approximation has multilayer perceptron architecture with one hidden layer containing nine

472

A. Zagitova et al.

Fig. 2 a Low-pressure compressor performance map. b 63 point experimental data map

neurons with TANSIG activation function. The arguments of the function (1) are input variables of ANN model and Gv n lpc is the ANN output. The learning algorithm of neural network is back-propagation algorithm. The ANN was trained on 63 experimental data points. To get the approximate values of intermediate points (not represented in experimental data set), the step of nn lpc argument was chosen as 100, corresponding plpc values were calculated using piecewise linear interpolation, and the trained ANN produced simulation output for this input data set. The previously calculated set of input data values of the nn lpc variable and its corresponding values of Gv n lpc were used as an input data grid. Experimental data points were used as interpolation nodes. Average deviation of approximation results from the results of linear interpolation of experimental data was measured at 14% from the average output variable value, while maximal deviation has reached 72%. These results do not satisfy accuracy requirements due to the insufficiency of initial experimental data points. High financial and time costs of flight and benchmark tests do not allow expanding the experimental data set at early stages of GTE design [16–18]. To solve this problem, the method of expansion of ANN training data set based on the use of cubic spline and polynomial interpolation of experimental data was suggested. The previously calculated set of input data values with the step of the nn lpc variable and its corresponding values of Gv n lpc were used as an input data grid. Experimental data points were used as interpolation nodes. These nodes divide the input data set into eight pieces. Polynomial interpolation with the polynomial degree maximizing accuracy of interpolation was applied to each one of the eight pieces of input data. The resulting surface is presented in Fig. 3a. The results of cubic spline interpolation of the same data are presented in Fig. 3b. Mean deviation from the results of linear interpolation served as an accuracy criterion. Cubic spline interpolation provided better accuracy than polynomial. On the other hand, cubic spline interpolation method of ANN training set expansion requires storage of all

Information Support of Gas-Turbine Engine Life Cycle Based …

473

Fig. 3 a Surface of polynomial interpolation. b Cubic spline interpolation of experimental data

the coordinates of points obtained as a result of interpolation. Polynomial interpolation method requires only storage of polynomial degree and coefficients. After application of each of two methods of interpolation, the resulting expanded training sets for ANN approximation contained 5653 data points each. The same ANN was trained on each training set and the resulting accuracy defined as the ratio of mean derivation from training set data to mean output variable value was 0.25% for cubic spline interpolation method and 0.0353% for polynomial interpolation method.

4 Software Agent for Neural Network Approximation The suggested method of neural network approximation of GTE compressor performance with the use of expanded training data set can be implemented as part of multiagent system for GTE development and debugging support. New bench or flight tests of compressor might correct and augment existing experimental data what requires updating training set, retraining of the ANN and probably adjustment of its architecture for optimal approximation. A JADE multiagent system can be developed in order to automatize this process. Architecture of the multiagent system is presented in Fig. 4. The suggested multiagent system implements an automated choice of the ANN training set expansion method depending on available hardware resources of computation device or system running GTE simulation. If available RAM allows storage of big volumes of data, the cubic spline interpolation expansion method will be preferred as it provides higher accuracy of compressor performance ANN approximation. In case of the tighter RAM limitations, the multiagent system

474

A. Zagitova et al.

Fig. 4 Architecture of MAS for neural network approximation of LPC performance

implements polynomial interpolation method for training set expansion as it requires less initial data to be stored. The agent Agent_BenchTestDB implements an interface between JADE multiagent system and database of experimental data (Fig. 5). New experimental data on compressor parameters after a new bench test are stored in the Agent_ BenchTestDB. Agent_DBChecker regularly sends queries to Agent_BenchTestDB and compares experimental data with the already available, updating the experimental data set when necessary.

Fig. 5 UML sequence diagram for information exchange between agents

Information Support of Gas-Turbine Engine Life Cycle Based …

475

Then, it sends the updated data to the Agent_Interpolator that chooses and implements the ANN training set expansion method. The criterion of choice is information about available hardware resources, provided by the Agent_HardwareChecker. An expanded training data set is then sent to the Agent_ANN that implements approximating neural network architecture and performs its initial training, retraining, and simulation. Agent_GTE runs a GTE simulation model. It provides input data for ANN approximation—values of plpc and nn lpc variables—sends them to the Agent_ANN, and receives back the result of neural network approximation—the value of the Gv n lpc variable, which is used in subsequent GTE simulation model computations.

5 Conclusion Application of agent-based technologies for intelligent support of information exchange and decision making during GTE life cycle can increase effectiveness of hardware, financial, information, and labor resources use and reduce the time length of development and introduction into service of new models of gas-turbine aviation engines. The suggested method of ANN training set expansion increases the accuracy of compressor performance representation given scarce experimental data. The suggested MAS prototype architecture allows optimizing use of hardware resources at this section of GTE development stage. Similarly, MAS architecture can be used for cost-effective neural network representation of performances of other GTE components and calculation of their parameters during GTE simulation.

References 1. Fedorchenkova DG, Novikova DK (2016) Cycle counting methods of the aircraft engine. Int J Environ Sci Educ 11(4):3832–3846 2. Novikov D (2014) Development of squeeze film damper characteristics calculation methods which take into account a liquid inertia forces. Res J Appl Sci 9(10):649–653 3. Astafiev VI, Fedorchenko DG, Tzypkaikin IN (1996) Complex stress-time cycles influence on aircraft engine parts fatigue strength. In: Proceedings of the sixth international fatigue congress. Berlin, 6–10 May 1996 4. Pierre C, Smith TE, Murthy DV (1991) Localization of aeroelastic modes in mistuned high-energy turbines. AIAA Paper 91–3379:1991 5. Gunetti P, Mills A, Thompson H (2008) A distributed intelligent agent architecture for gas-turbine engine health management. In: Proceedings of 46th AIAA aerospace sciences meeting and exhibit. Reno, NV, 7–10 Jan 2008 6. Gunetti P, Thompson H (2008) A soar-based planning agent for gas-turbine engine control and health management. In: Proceedings of 17th IFAC world congress. Seoul, Korea, 6–11 July 2008 7. Jennings N, Wooldridge M (1998) Applications of intelligent agents. In: Agent technology: foundation, applications and markets. Springer-Verlag Berlin Heidelberg

476

A. Zagitova et al.

8. Wooldridge W (1999) Intelligent agents. In: Multi-agent systems: a modern approach to distributed artificial intelligence. The MIT Press, Cambridge, Massachusetts, London, England 9. Rzevski G, Skobelev P (2014) Managing complexity. WIT Press, Southampton, Boston 10. Rzevski G, Brebbia CA (eds) (2018) Complex systems studies. WIT Press, Southampton, Boston 11. Rzevski G, Brebbia CA (eds) (2017) Complex systems theory and applications. WIT Press, Southampton, Boston 12. Rzevski G, Brebbia CA (eds) (2016) Complex systems fundamentals and applications. WIT Press, Southampton, Boston 13. Rzevski G (2016) Managing complexity: theory and practice. In: Proceedings of 11th system of systems engineering conference (SoSE), Kongsberg, Norway, 12–16 June 2016 14. Rzevski G, Knezevic J, Skobelev P et al (2016) Managing aircraft lifecycle complexity. Int J Des Nat Ecodyn 11(2):77–87 15. Dambrosio L, Mastrovito M, Camporeale S (2007) Performance of gas turbine power plants controlled by multiagent scheme. J Eng Gas Turbines Power 129:738–745 16. Kondratyeva NV, Valeev SS (2016) Simulation of the life cycle of a complex technical object within the concept of big data. In: CEUR proceedings of 3rd Russian conference on mathematical modeling and information technologies. Yekaterinburg, Russia, 16 Nov 2016 17. Kovtunenko A, Bilyalov A, Valeev S (2018) Distributed streaming data processing in IoT systems using multi-agent software architecture. In: Proceedings of the 18th international conference on next generation wired/wireless networking (NEW2AN), and 11th conference on internet of things and smart spaces (ruSMART). St. Petersburg, Russia, 27–29 Aug 2018 18. Kondratyeva N, Valeev S (2016) Fatigue test optimization for complex technical system on the basis of lifecycle modeling and big data concept. In: Proceedings of 10th conference on application of information and communication technologies AICT 2016. Baku, Azerbaijan, 12–14 Oct 2016 19. Vasilyev VI, Valeyev SS, Shilonosov AA (2001) Design of neurocontroller for gas-turbine engine multi-mode control. In: Proceedings of 8th international conference on neural information processing (ICONIP-2001). Fudan University Press, Shanghai, China, Nov 2001 20. JAVA Agent Development Framework (2018). http://jade.tilab.com. Accessed 15 Sep 2018

Automatic Print Job Scheduling and Management Over Multiple 3D Printers I. A. Gushchin, I. V. Martynovich and I. S. Torubarov

Abstract 3D printing is one of the most convenient ways to create final parts, because it does not require long preparation of production. Due to the fact that the 3D printing process is rather slow, many companies acquire several 3D printers at once. Because each 3D printer has its own interface and requires individual settings, control, and monitor, the status of printing becomes problematic. For simple management of a group of 3D printers from a single interface, some manufacturers, such as Ultimaker, Stratasys, Formlabs, and 3D Systems, have created (each own) a special software. However, this software is not suitable for printers of other manufacturers and does not take into account the possibility of using additional equipment. In this article, the decisions on management of a group of 3D printers of the above-mentioned manufacturers are given, their advantages and disadvantages are considered, and further, the algorithm allowing to use additional equipment, for example, automatic puller of finished products, is described. Keywords Additive technologies


3D printing
Batch production
Farm

1 Introduction 3D printers are used in many areas of production [1], not only for the creation of single prototypes of products [2–4], such as sensor [5] elements, but also for the production of parts and printing of large assembled products, for example, vacuum grips [6]. There are several ways to improve the performance of 3D printing [7]; the most efficient is to install many devices in one room, the organization of the so-called “farm.” This gives rise to the management complexity of a group of 3D printers. For convenient operation, you need an application that will organize the print queue, keep track of finished products, and warn the user about errors. Only a limited number of manufacturers have such programs. I. A. Gushchin (&)
I. V. Martynovich
I. S. Torubarov Volgograd State Technical University, 28, Lenin Avenue, 400005 Volgograd, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_51

477

478

I. A. Gushchin et al.

2 Stating the Aim As the 3D printer park grows, there is a need for a centralized management system [8–10] that allows you to keep track of different print materials and maintain printers. With a large number of devices and materials for printing, it is difficult for a person to monitor the distribution of print jobs that are configured for specific materials and print settings. Some manufacturers have their own solutions, but they can be used only with printers of this manufacturer and do not always have the necessary functionality. This requires the development of an algorithm to organize a print queue of several 3D printers, allowing the use of additional equipment for removal of finished parts from printing platform, the device for smoothing the surface of printed parts using chemical processing [11] and ensuring uniform loading of printers.

3 Review of Existing Solutions Various manufacturers have created multiple device management solutions that automatically distribute print jobs. Consider these solutions. Ultimaker Cura Connect (Fig. 1) combines several FDM printers manufactured by Ultimaker through a wireless or wired network [12]. The management interface is available either through the browser or via software to prepare print models (slicer) Ultimaker Cura. The parameters of all the printers are displayed in the Monitor tab (Fig. 2). Each printer in the group is assigned a unique IP address. One printer becomes the host (Control node) of the group. All interactions with Cura Connect (sending print jobs, error messages, start/end printing) pass through the host. When you submit a job, the software will find a free printer with the appropriate configuration (print area size, material, etc.). If no appropriate configuration is found, the operator will be prompted to change the configuration. If all suitable printers are busy, the job will be queued. This solution is designed to work with the devices and software, manufactured by Ultimaker, because it has built-in functions, specific to the printers of the manufacturer (automatic definition of the printing material, web camera monitoring, automatic heating block definition). Stratasys GrabCAD Print (Fig. 3) is a solution of Stratasys company, which allows you to work with FDM and PolyJet printers of this company [13]. Management is made through a corporate account, i.e., all interactions between users and printers pass through a single GrabCAD Print server. Server software can be run on any PC with the Windows operating system. Print quality settings (layer height, nozzle diameter) are stored separately for each printer. Model settings (infill style, wall thickness) are common to all printers. A schedule mode is available that allows you to view the print queue of all printers (Fig. 4) as well as the report generation mode.

Automatic Print Job Scheduling and Management …

479

Fig. 1 Cura Connect interface

Fig. 2 Monitor tab in Cura Connect

The disadvantages of this solution can be noted are the compatibility with only certain Stratasys 3D printers and the need for a separate PC running Windows. Formlabs Dashboard (Fig. 5) is a Formlabs solution for the Form 2 SLA printers [14]. It allows you to manage multiple printers using a cloud application and connect printers over the Internet. The main advantage is the ability to remotely monitor printer status and print management. You can also see the type and number

480

Fig. 3 GrabCAD Print interface

Fig. 4 Schedule mode in GrabCAD Print

I. A. Gushchin et al.

Automatic Print Job Scheduling and Management …

481

Fig. 5 Formlabs Dashboard interface

of materials available for printing. It is possible to send notifications about the printing process (beginning, ending, interruption) by SMS or e-mail. Among the disadvantages of this solution can be noted the need to permanently connect printers to the Internet. Also, this solution does not take into account the work of additional equipment required for SLA printing technology: the device for washing from the unashamed resin and the device for curing of the printed product. 3D Systems 3D Sprint (Fig. 6) is a software of 3D Systems company for management of SLS printers [15]. Printers are merged through a local network. Jobs are run based on delivered priority. A priority queue allows run jobs quicker with a higher priority to automatically push lower priority jobs to the second plan. The status of each printer is displayed in the Print Queue tab (Fig. 7). It is necessary to allocate that in any of the considered applications, there is no uniform distribution of loading on printers, and there is no account of hours of work. This means that individual printers may wear out more than others, leading to premature breakdowns. Also these solutions do not take into account possible additional equipment for FDM printers, for example, automatic puller of parts, device for smoothing the surface of printed products by chemical treatment, etc.

482

Fig. 6 3D Sprint interface

Fig. 7 Print Queue tab in 3D sprint

I. A. Gushchin et al.

Automatic Print Job Scheduling and Management …

483

4 Algorithm Development The lack of a solution that allows the use of FDM printers together with additional hardware required the development of proprietary software that takes into account these features. One possible way to automatically schedule and manage the printing of a 3D printer group is to use single-board computers [16] that are installed in each printer and that are running the Linux operating system. The most convenient and extensible way of developing this software is its representation in the form of separate services interacting with each other [17, 18]. In Linux operating systems, there is a built-in mechanism of interaction of services—D-Bus. D-Bus is an interprocess communication system that allows applications to communicate with each other. D-Bus also provides the concept of services. Service is the unique location of the application on the bus. When the application is started, it registers one or more services that it will own until it frees itself. Until then, no other application claiming the same service will be able to take it. When the application closes, the associated services also cancel the registration, and the D-Bus sends a signal that the service is closed. The number of services may vary depending on the additional equipment used by the 3D printer. There are few services could be defined: network service that provides an Ethernet or Wi-Fi connection, an system update service, and a system service that enables printers and print jobs. The basis of the system service is working with the database [19]. The database includes the following tables: “Printers”—directly 3D printers added to the cluster for distribution of print jobs, “Users”—a user system that allows users to access and set their roles, “Print Jobs”—a table with print jobs, includes both completed jobs and jobs in the print queue, “Notifications”—the system of notifying the user about various events. The system service (Fig. 8) provides a set of different interfaces for interacting with the database and the print job queue: • IPrintersService—Interface of printers management and work with the printer table • IUsersService—User management interface and authorization, • IPrintjobsService—Interface for work with print jobs and queue management, • IDonejobsService—Interface for working with completed jobs, • IFilestorageService—Interface for working with G-Code files, • IConnectionService—Interface that enables printers to connect to each other, • INotificationsService—Interface for working with notifications. Each print job includes information about the G-Code file [20], the date the job was created, and the user who created the job. You can also assign a printer to print on a specific device, for example, in the case of an installed material, or the desired print area. In the future, the print job could be refined, for example, to specify

484

Fig. 8 System service

Fig. 9 Queuing in the FIFO way

I. A. Gushchin et al.

Automatic Print Job Scheduling and Management …

485

several printers with similar parameters (working volume, number and type of print heads) and to specify the printing of the product with a specific material. The print job queue is based on the FIFO principle (first in, first out). The first incoming job is processed first (Fig. 9). The queue is working in a loop on a separate thread. The algorithm of the queue (Fig. 10) operation is as follows: • Algorithm checks for print jobs in the queue, • The first job from the queue is selected, • Algorithm checks whether the printer is assigned to this job, if assigned, checks for its availability, and if the printer is available and ready to print, the job is sent to print, • If the printer is unavailable or busy with another job, the next job is selected until the job is received without the assigned printer, or the assigned printer is ready to accept the job.

Fig. 10 Block diagram of the queue operation algorithm

486

I. A. Gushchin et al.

After the printing is finished, additional equipment (automatic puller) is activated and the printed part is removed from the platform. The printer is then ready to accept print jobs again. Printers can also be disconnected from work in the queue, translating them into maintenance mode. In this mode, jobs are not sent to the printer, and the system technician performs printer maintenance, calibration of the platform, extruders, grease guides, check the tension of belts, etc. Acknowledgements As a result of the research, the software solutions to organize the distribution of print jobs of 3D printer manufacturers were studied. All of them have their advantages and disadvantages. None of the solutions considered allowing the use of additional equipment to increase the automation of the printing process on several printers, which caused the need to develop our own software solution. At the moment, the developed solution allows to provide optimum loading of 3D printers and to provide uniform deterioration of equipment. Further development will be aimed at the possibility of using additional equipment for the removal of finished printed parts.

References 1. Teterin E, Zhuravlev D, Berchuk D (2016) Mobile extrusion machine for the production of composite filaments for 3d printing. In: 2nd International conference on industrial engineering, applications and manufacturing (ICIEAM). https://doi.org/10.1109/icieam. 2016.7910908 2. Gebhardt A (2003) Rapid Prototyping. Hanser Verlag, Berlin 3. Pham JD, Gault R (1998) A comparison of rapid prototyping technologies. Int J Mach Tools Manuf 38(10–11):1257–1287. https://doi.org/10.1016/S0890-6955(97)00137-5 4. Lass N, Tropmann A, Ernst A et al (2011) Rapid prototyping of 3D microstructures by direct printing of liquid metal at temperatures up to 500 °C using the starjet technology. In: Proceedings 16th international solid-state sensors actuators and microsystems conference. pp 1452–1455. https://doi.org/10.1109/transducers.2011.5969623 5. Margatskaya EA, Kurnosov DA (2017) Differential optical sensor for measuring small linear displacement. In: 2017 International conference on industrial engineering, applications and manufacturing (ICIEAM). https://doi.org/10.1109/icieam.2017.8076378 6. Makarov AM, Mushkin OV, Lapikov MA (2018) Use of additive technologies to increase effectiveness of design and use of a vacuum gripping devices for flexible containers. In: International conference on modern trends in manufacturing technologies and equipment 2018 (ICMTME 2018), p 6. https://doi.org/10.1051/matecconf.2018. 22401082 7. Avdeev AR, Shvets AA, Gushchin IA (2016) Perspektivnye napravleniya uvelicheniya proizvoditel’nosti tekhnologij obyomnoj pechati (Upcoming trends of productivity increasing of additive technologies). In: XX Regional Conference of young researchers of the Volgograd region. pp 103–105 8. Tyutikov VV, Voronenkova AA (2017) Analytical synthesis and analysis of industrial facility control system versions. In: 2017 International conference on industrial engineering, applications and manufacturing (ICIEAM). https://doi.org/10.1109/icieam.2017.8076118 9. Kuo BC (2009) Automatic control systems. Wiley 10. Bolton W (2002) Control systems. Newnes 11. Gushchin IA, Avdeev AR, Shvets AA et al (2018) Vliyanie himicheskoj obrabotki 3D pechatnyh izdelij na ih prochnost (Influence of chemical processing of 3d printed parts on their strength). Proc VSTU Ser Progressive Technol Mech Eng 2(212):58–60

Automatic Print Job Scheduling and Management …

487

12. Ultimaker Cura Connect installation and user manual. In: Ultimaker BV. https://ultimaker. com/download/73548/180104-Ultimaker-CuraConnect-Manuals-EN-v1.1.pdf. Accessed 13 Nov 2018 13. About GrabCAD print. In: Stratasys Ltd. https://help.grabcad.com/category/194-aboutgrabcad-print. Accessed 13 Nov 2018 14. Using formlabs dashboard. In: Formlabs Inc. https://support.formlabs.com/s/topic. Accessed 13 Nov 2018 15. D Sprint. In: 3D Systems Inc. https://softwaresupport. Accessed 13 Nov 2018 16. Ignatiev KV, Kopichev MM, Putov AV (2016) Autonomous omni-wheeled mobile robots. In: 2016 2nd international conference on industrial engineering, applications and manufacturing (ICIEAM). https://doi.org/10.1109/icieam.2016.7910957 17. Martinov GM, Kozak NV, Nezhmetdinov RA (2017) Implementation of control for peripheral machine equipment based on the external soft PLC integrated with CNC. In: 2017 International conference on industrial engineering, applications and manufacturing (ICIEAM). https://doi.org/10.1109/icieam.2017.8076119 18. Grigoriev SN, Martinov GM (2016) The control platform for decomposition and synthesis of specialized CNC systems. Proc CIRP 41:858–863 19. Samsonova NV, Shevtchenko OY (2016) The algorithm for creating geospatial database to ensure the effective functioning of the gas distribution network. In: 2016 2nd International conference on industrial engineering, applications and manufacturing (ICIEAM). https://doi. org/10.1109/icieam.2016.7911668 20. Gushchin IA, Avdeev AR, Shvets AA (2016) Preobrazovanie obyomnoj modeli izdeliya v upravlyayushchij kod dlya 3D printera (Converting a 3D model of a product into a control code for a 3D printer). In: XX Regional conference of young researchers of the Volgograd region. pp 206–208

CNC Processing Equipment’s Technical Operability Evaluation by Developing Mathematical Models Based on Continuous Logic of Antonyms E. Krylov, N. Kozlovtseva and A. Kapitanov

Abstract As the title implies, the article describes the issues of CNC machines’ technical operability evaluation in real time. The scientific research relevance of this topic has been substantiated, and suitable monographs, articles, and regulations have been scanned. The types and causes of the automated technological equipment failures on the processing industries have been analyzed. The authors proposed a new approach to estimating the probability of equipment failure using the continuous-valued logic of antonyms, the mathematical tools technique of which allows making calculations based on linguistic variables in a convenient and informative form. The technical operability of the equipment is considered the example of a CNC machine, in particular, its electromechanical subsystem. A general solution pattern has been developed; attempts are made to present mathematical simulation. Three methods for technical operability evaluation using weighting factors are proposed. The developed methods and models make it possible to get away from traditional statistical calculations of reliability and to diagnose the state of automated equipment in a real-time mode.







Keywords CNC machine Technical operability Condition monitoring Continuous logic Equipment failures Expert system










1 Introduction In the automated manufacturing industry, machine tools with computer numerical control (CNC) are widely used, and its characteristic feature is the high precision of machining of parts with complex surfaces [1–3]. The crucial technological requirement for the effective use of new generation CNC machine tools is to E. Krylov (&)
N. Kozlovtseva Volgograd State Technical University, 28, Lenin Avenue, 400005 Volgograd, Russia e-mail: [email protected] A. Kapitanov Stankin Moscow State Technical University, 1, Vadkovskiy Lane, 127994 Moscow, Russia © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_52

489

490

E. Krylov et al.

minimize all types of equipment downtime in order to release production at the time specified by the customer. The problem of equipment downtime while the products manufacturing by expensive equipment (the cost of the machine varies from 3 to 30 million rubles) depends largely on the presence of negative structural and technological factors (such as the imperfection of the turret and the guiding ways of the machine, facilities, tool-holding equipment, cutting tool) which influence shows itself in the coherence violations of the shaping movements and, as a consequence, in reducing the quality and accuracy of the processed surfaces. Increasing the degree of processing equipment automation involves the extensive use of its nodes condition monitoring facilities during operations [4–9]. One of the crucial tasks during the automated monitoring and diagnosis systems engineering is the task of determining the degree of technical equipment operability. According to national state standard 27.002-2015, an upstate of maintenance entity is a condition of an object in which the values of all parameters characterizing its ability to perform specified functions conform to the requirements of regulations and technical documentation [10]. During the operation of any object, there are deviations of its initial parameters from the given values due to drift failures. Failure processes are very diverse and depend on the type, structure, and properties of the failure, as well as on violations of operating conditions, i.e., overloads, unacceptable temperature conditions, increased vibrations, external field effects, and other causes [11–13]. The work of native [14–16] and foreign [17–21] researchers is devoted to the development of methods and algorithms of technological equipment automated control on automated processing equipment with the aim of reducing the probability of failures occurrence. By predictability of their occurrence, instant failures are divided into sudden and gradual. Sudden failures are caused by the destruction of elements or components of the product (breakage, short circuit, unacceptable deformation). Sudden failures under normal operating conditions are characterized by the abrupt curve of the product degree damage dependence on its time operation. The main symptom of a sudden failure is the failure rate independent from the operational time, i.e., the failure probability over a small interval of the operational time, following the time point in question, depends only on the length of this interval, but does not depend on the previous life of the product (not associated with the gradual accumulated damage). Gradual failures are caused by the violation of the products elements or components quality. Gradual failures appear as a result of the irreversible changes accumulation in the product, progressively worsening its output parameters. The gradual failures are associated with the wearing processes, corrosion, fatigue, and creep of materials. The main symptom of a gradual failure is the monotonous increasing character of the failure dependence on of the object’s operating time. Therefore, the gradual, but constantly accumulated, changes which take place in the object make it necessary not just to fix them, but to assess the deviations from the level of regulatory performance indicators in a given moment in time. Ergo, it is necessary to distinguish not two different functional states of an object (workable and not working), but their unlimitedly large gradation, which requires the use of a modern mathematical apparatus.

CNC Processing Equipment’s Technical Operability …

491

For automated technological equipment, it is of practical interest not a concept of an averaged functional state, which can be obtained by traditional calculating methods through the reliability theory, but knowledge of the numerical estimation of an object state at the recording its parameters time. The authors propose to reduce a solution of the problem of the technological equipment automated condition control to the solution of the problem of an object continuous numerical technical operability evaluation based on a continuous logic mathematical model. An important part of the task is to identify critical pre-emergency states of the system, which are states that are currently acceptable, but in which the system can go to the state beyond the specified requirements already at the next time.

2 Basic Definitions As a basis for the mathematical model development, it is proposed to use the logic of antonyms (LA), which is the only continuous-valued logic completely consistent with classical logic, i.e., it allows to use the classical mathematical apparatus [22–24]. The basic concept of LA is the A and aA antonym pair, which is considered as a pair of opposite limiting properties of a certain parameter. The parameter value is given by the value H(A) or H(aA), which is a quantitative assessment of the object property A presence. The relationship between the H(A) and H(aA) assessments can be established from the expression: HðaAÞ ¼  logð1  2HðAÞ Þ

ð1Þ

The coordinates of the researchable object state can be selected values defined in the interval (0,∞), or after normalization in the interval [0,1]. Intermediate values of antonymic evaluations are selected, as a rule, in terms of linguistic variables. In LA, the task of assessing changes in the object properties can be solved on the basis of establishing a quantitative indicator of the logical connection “tightness” between the object parameters. Two types of bonds are used: the c-bond corresponding to the logical multiplication operation (strong coupling), and the b-coupling corresponding to the logical addition operation (weak coupling). The object being examined can be presented as a combination of several technological systems (A1, A2, …, Am), which are divided into subsystems Bi;j ði; j 2 N; i 2 ½1; m; j 2 ½1; nðiÞÞ. The subsystem consists of elements Xk ðk 2 N; k 2 ½1; N Þ. We denote the object technical operability as Y and the degree of object operability in the form of H(Y). To simplify the write, we will not use the triple of indices (we will not write Xi;j;k counting k 2 ½1; kði; jÞÞ, but use the continuous elements numbering. The hierarchy depth of the original object is not regulated. Any of its subsystems, if necessary, can act as objects of study.

492

E. Krylov et al.

Due to the fact that only the operability of the process equipment and its subsystems is investigated hereinafter, in order to simplify the formal side of the presentation, we denote by the same letters both the process equipment (physical devices) and its state (operability): Y; A1, A2, … Am; B1.1, … Bn(m); X1, … XN. The H(Y), H(Ai), H(Bi,j), H(Xk) assemblies should be understood as numbers that are assigned to the technological system Y, the system Ai, the subsystem Bi;j , the element Xk, and as numbers that are assigned to their functional states (in particular, their operability). These assemblies can be considered as numerical evaluations of objects Y, Ai, Bi;j , Xk (broad understanding of the functional H defined on Ai, Bi;j , Xk) and the degree of process equipment operability (special understanding of the functional).

3 General Solution Pattern Among CNC machines in the computer-integrated manufacturing, the most common are turning and milling machines for processing flat and spatial surfaces. Typical configurations of CNC machines are shown in Fig. 1, and the composition of the main subsystems ensuring the flexibility and reliability of this equipment type is shown in Fig. 2. We assume that the object being examined (CNC machine) Y consists of m systems, each of which consists of n(i) (i is the system number) subsystems, and the subsystems are divided into elements (total N in object Y). Suppose that each element is connected to a primary transducer that measures particular physical quantities (temperature, speed, torque, etc.). The set of interrelated technical facilities that provide automation of receiving initial information from sensors and this information processing while manufacturing is called a multiparameter control system [25, 26]. Multiparameter control will be considered as a set of procedures aimed at identifying the operating state of the machine’s element (subsystem).

Fig. 1 Configurations of CNC machines: a Turning. b Milling

CNC Processing Equipment’s Technical Operability …

493

Fig. 2 Composition of the CNC machine main subsystems

To assess the technical operability, it is necessary to obtain relative values (points, percentages) corresponding to the data taken and reflecting the attitude of the decision-maker (DM) to the observed parameters values. For this purpose, the sensor readings are transmitted to a specialized computing device based on a multichannel ADC and a microcontroller, which converts them into dimensionless quantities in such a way that the smaller these quantities are, the more “negative” is the decision-maker attitude to the corresponding sensor reading. Then the dimensionless quantities are transmitted to an industrial computer, which is following programming and forms a functional state integrated assessment of the object being examined (e.g., an inserted blade-type tool) [27]. This assessment should take into account the values of particular parameters, the logical (cause–effect) relationship between all the considered elements, as well as the preference of given systems, subsystems, elements, expressed by weighting factors. An integral assessment of H(Y) comes to the decision-making device (DMD) along with the all parameters assessments. It gives a conclusion on technical operability values of the object in the instant of technological parameters registration. The program, loaded in the DMD, is formed by experts in a view of the aims and available funds to achieve them. The schematic problem solution pattern can be implemented in various ways. The main difficulties in solving it are connected with the transition from a physical system to a mathematical model of the object’s technical operability. Hereinafter, we illustrate the general approach to the formal model development.

494

E. Krylov et al.

4 Preliminary Formalization A CNC machine can be represented as a combination of the following four systems: electromechanical (A1), hydraulic (A2), pneumatic (A3), and software (A4). We assume that the objects A1, A2, A3, A4 introduced into consideration are independent of each other. The functional state of the equipment is completely determined by the state of the specified systems; therefore, the machine’s technical operability can be represented as a combination of these four systems’ operability. The subdivision of systems into subsystems will be examined using the example of system A1. The crucial machine components that determine the electromechanical subsystem operability are: electric motors of the main motion and feed drives (B1.1), spindle bearing assembly (B1.2), tool storage (B1.3), automatic tool gripper (B1.4), ballscrew-nut transmission (B1.5), rotary table (B1.6), guides (B1.7), and protective cover (B1.8). Suppose that at a certain point, when the machine operability is determined, a change in the state of the given electromechanical system cannot affect the other technological subsystems’ states. Similarly, the states of all subsystems Bi; j ði 2 ½1; 2; j 2 ½1; nðiÞÞ are independent of each other. Therefore, the following is true: Y¼

4 [ i¼1

Ai ;

Ai ¼

nðiÞ [

Bi;j ;

j¼1

where n(1) = 8, п(2) = 5, п(3) = 6, п(4) = 7. Each subsystem comprises an association of some elements (Fig. 3).

Fig. 3 Hierarchical CNC machine model

ð2Þ

CNC Processing Equipment’s Technical Operability …

495

The type of all subsystems for a particular type of automated equipment is analytically established. For example, subsystem B1,j (j2[1, 8]) consists of following associations of independent objects being examined: B1;1 ¼

6 [

Xk ;

. . .;

B1;8 ¼

k¼1

41 [

Xk :

ð3Þ

k¼38

The verbal description of the B1, j subsystem’s elements is not given here.

5 Mathematical Modeling It is necessary to develop a model that would give an opportunity to get a quantitative evaluation (as a number) accordingly the initial objects. Taking into account the rules for constructing logical connections in LA and the independence of Ai ði 2 ½1; 4Þ; Bi;j ðj 2 ½1; 8Þ and Xk ðk 2 ½1; 41Þ objects, we conclude that the following equalities are valid: HðYÞ ¼

4 X

HðAi Þ;

ð4Þ

HðBi;j Þ;

ð5Þ

i¼1

HðAi Þ ¼

nðiÞ X j¼1

where n(1) = 8, п(2) = 5, п(3) = 6, п(4) = 7. HðB1:1 Þ ¼

6 X k¼1

HðXk Þ; . . .; HðB1:8 Þ ¼

41 X

HðXk Þ

ð6Þ

k¼38

Equalities, similar to equalities (6), can be set in accordance with the remaining subsystems Bi;j ði 2 ½2; 4; j 2 ½1; nðiÞÞ. All quantities in the Expression (2) can take any non-negative values. In particular, H(Xk) can take any value in the interval [0, 1]. In this case, 0 corresponds to the complete inoperability of the object, and 1—full performance. In the case when all Ai ði 2 ½1; 4Þ, all Bi; j ðj 2 ½1; nðiÞÞ, and all Xk are independent of each other and equal, the value of H(Y) is calculated according to Formulas (4), (5), and (6) (the last-named can be complemented by equations for calculations HðBi; j Þ, where i 2 ½2; 4; j 2 ½1; nðiÞÞ.

496

E. Krylov et al.

Consider the case when objects are independent, but they are not equal, i.e., they have different weights, expressed by coefficients. Let the objects Ai, Bi,j have weights qi, and pi,j respectively. In this case, the degree of the object operability Y is calculated by the formulas HðYÞ ¼

4 X

qi HðAi Þ;

ð7Þ

pi;j HðBi;j Þ;

ð8Þ

i¼1

HðAi Þ ¼

nðiÞ X j¼1

and H(Bi,j) is calculated by the Formulas (6), assuming all Xk are equals.

6 DND Simulation In order to address the challenge adequately, the authors have developed three methods, based on the proposed mathematical model, allowed to conclude about the technical operability state Y of the object under consideration. The first method supposes that   H  ðYÞ; H  ðAi Þ; H  Bi;j ; H  ðXk Þ

ð9Þ

are fixed values of technological parameters corresponding to some state of automated equipment with CNC, its systems, subsystems, and elements. This may be the original passport specifications of the object or current characteristics, obtained by calculation. In particular, as (9) the maximum values of the quantities (2), corresponding to the desired characteristics, can be taken. This leads to the equipment being recognized as efficient as possible. Calculating the difference   H  ðYÞ  HðYÞ; H  ðAi Þ  H ðAi Þ; H  Bi;j    H Bi;j ; H  ðXk Þ  H ðXk Þ;

ð10Þ

where the states of the object at some instant play the roles of subtracted values, we can make a conclusion about the object operability. There are several cases. For example, you can assign the weights of the differences (10) and make a conclusion about the operability with account to both the values (10) and their weights. You can divide the differences (10) into groups. In some groups, you can rank with giving preference to weights, and in others—with paying attention only to the magnitude of the differences themselves.

CNC Processing Equipment’s Technical Operability …

497

The second method involves calculation variation intervals of the values (2). Depending on which places are occupied by these quantities current values in their intervals, they make a conclusion about the object state Y. On this stage, you can also be guided by different criteria, for example, to vary the intervals limits or to rank them all or part of the intervals. The third method sets only the lower boundaries of the values (2) variation intervals. This method suits only to the unexacting approach to the technological equipment operability, or to the lack of equipment structure knowledge.

7 Results and Discussion Insufficient adequacy of mathematical models based on physical objects’ classical reliability theory makes it necessary to search active for new models and new theories of their development. In reliability theory, recently there was a need to develop new research directions. Let’s take a closer look at three of them. The first direction is the search for fairly simple models. The simplicity of the constructions and computations can be a certain guarantee of the use of theoretical constructions in engineering practice, which is not the case with many models of reliability theory developed today. The second trend—the creation of non-statistical reliability theory, the basis of which would lay the model—does not reflect the quantitative and qualitative approach to the theory. The third direction is the creation of reliability theory models based on the artificial intelligence theory. The last two trends are perspective with an economic view to enhancement of the information technologies in automated production area. Developing a new approach to assessing the operability of CNC processing equipment, based on automated control was designed to test the possibility of using the LA mathematical apparatus to solving actual engineering tasks. The novelty of the proposed mathematical model is in developed formalization of object technical operability by its parameters values assembly at the instant of their simultaneous registration rather than calculation of test object operability probability after a certain period of time, as in most cases takes place in the traditional reliability models. Not expected mass events. The results obtained are entitled to assume that the implementation of the proposed techniques and mathematical models can be achieved by applying the LA. Acknowledgements The authors would like to acknowledge the support and funding provided by The Volgograd State Technical University. The research is performed within the framework of the grant of the President of the Russian Federation for the state support of young scientists (MD-6629.2018.9).

498

E. Krylov et al.

References 1. Mesheryakova VB, Starodubov VS (2015) Metallorezhushie stanki s CHPU (CNC metal-cutting machines). NITS INFRA-M, Moscow 2. Vyzhigin AY (2009) Gibkie proizvodstvennye sistemy (Flexible manufacturing systems). Mashinostroenie, Moscow 3. Schirtladze AG, Ukolov MS, Sazonov GG (2012) Upravlenie stankami i stanochnymi kompleksami (Control of machines and machine complexes). TNT, Staryi Oskol 4. Krylov EG, Serdobintsev YP (2018) Povysheniye effektivnosti funktsionirovaniya instrumental’nykh sistem avtomatizirovannogo stanochnogo oborudovaniya (Improving the efficiency of the instrumental systems functioning in automated machine-tool equipment). VolgGTU, Volgograd 5. Krylov EG, Makarov AM, Kozlovtseva NV, Kapitanov AV (2017) Criticality identification in technological system via decision tables. Izvestia VolgGTU 5(200):68–70 6. Serdobintsev YP, Krylov EG (2015) Problem-oriented approach to improving tool operation at machining centers. Russ Eng Res 9:693–695 7. Krylov EG, Sergeev AS (2014) The ultimate state control of multiblade carbide tools. Test Diagn 10:30–35 8. Krylov EG, Serdobintsev YP, Kozlovtseva NV (2013) Avtomatizirovannaya sistema kontrolya sostoyaniya rezhushchego instrumenta pri rezanii trudnoobrabatyvayemykh materialov (Automated system for monitoring the status of the cutting tool when cutting hard materials). Mekhatronika, Avtomatizatsiya, Upravlenie 10:47–51 9. Krylov EG, Kukhtic MP, Kozlovtseva NV (2018) Use of thermoelectrical phenomena for hard-alloy cutting plates control in the cutting zone. In: 2018 International Russian automation conference, p 4. https://doi.org/10.1109/rusautocon.2018.8501835 10. State Standard 27.002-2015. Reliability in technology. The basic concepts. Terms and definitions. Standartinform Publ., Moscow 2016, p 24 11. Ostreikivskiy VA (2003) Teoriya nadezhnosti (Reliability theory). High school, Moscow 12. Polovko AM, Gurov SV (2009) Osnovy teorii nadezhnosti (Fundamentals of the reliability theory). BHV-Peterburg, Saint-Petersburg 13. Trukhanov VM, Tarnayev AG (2016) Nadozhnost’ i diagnostika slozhnykh system (Reliability and diagnostics of complex systems). Spectr, Moscow 14. Bushuyev VV, Kuznetsov AP, Sabirov FS (2016) Problemy tochnosti i effektivnosti sovremennykh metallorezhushchikh stankov (Problems of accuracy and efficiency of modern machine tools). STIN 2:6–16 15. Grigoriev SN, Kozochkin MP et al (2012) Diagnostika tekhnologicheskogo oborudovaniya v sovremennom stankostroyenii (Diagnostics of technological equipment in modern machine-tool construction). Tekhnol mashinostroyeniya 1:45–50 16. Solomentsev YM, Pavlov VV (2010) Modelirovaniye proizvoditel’nykh sistem v mashinostroyenii (Modeling of productive systems in mechanical engineering). Yanus-K, Moscow 17. Alaniz-Lumbreras PD (2006) Sensorless detection of tool breakage in milling. Mach Sci Technol 2:263–274 18. Wang WH, Hong GS et al (2007) Sensor fusion for online tool condition monitoring in milling. Int J Prod Res 21:5095–5116 19. Bosetti P, Leonesio M, Parenti P (2013) On Development of an optimal control system for real-time process optimization on milling machine. Procedia CIRP 12:31–36. https://doi.org/ 10.1016/j.procir.2013.09.007 20. Jankowski M, Wozniak A (2015) Mechanical model of errors of probes for numerical controlled machine tools. Measurement 77:317–326 21. Nouria M, Fussella BK et al (2015) Real-time tool wear monitoring in milling using a cutting condition independent method. Int J Mach Tools Manuf 89:1–13

CNC Processing Equipment’s Technical Operability …

499

22. Dintsis DY (2006) Sovmestnoye ispol’zovaniye nechetkologicheskikh i diskretnykh modeley dlya postroyeniya modeley tekhnologicheskikh i upravlencheskikh sistem (Sharing of fuzzy and discrete models for building models of technological and management systems). Instrum Syst: Monit, Control, Diagn 5:61–64 23. Levin VI (2003) Metody nepreryvnoy logiki v zadachakh upravleniya (Continuous logic methods in control problems). Avtomatika i Telemekhanika 3:28–51 24. Kopaneyeva IN (2002) Monitoring i upravleniye kachestvom protsessa proizvodstva s primeneniyem logiki antonimov (Monitoring and quality management of the production process using the logic of antonyms). Dissertation, St. Petersburg state techn. university 25. Krylov EG, Serdobintsev YP (2015) Formalizatsiya protsedury mnogokriterial’nogo vybora rezhushchego instrumenta v ASTPP (Formalization of the procedure of multi-criteria selection of cutting tools in CAD/CAM systems). STIN 3:5–9 26. Krylov EG, Kozlovtseva NV (2015) Formalizatsiya protsessov obrabotki informatsii v sistemakh mnogoparametricheskogo kontrolya rezhushchego instrumenta (Information processing formalization in cutting tools’ multiparameter control systems). Izvestia VolgGTU 11(173):56–59 27. Krylov EG (2016) Programma mnogokriterial’nogo vybora rezhushchikh plastin (The program of cutting plates multi-criteria selection). Certificate of state registration R.F. no. 2016663002

Digital In-Line Moisture Meter of Thin Sheet Materials S. V. Makartichyan, L. V. Khoperskova and V. E. Avvakumov

Abstract Modern technological wood drying equipment is not sufficient to ensure the constancy of the humidity dimensional products throughout the volume, which often leads to deviations from the specified limits of moisture content. This leads to losses not only in materials, but also in heat and electricity. This is especially true for expensive wood products. This article describes the main methods of moisture measurement used in wood moisture meters design. The advantages of combined measurement methods in solving the moisture measuring problems are reviewed. A block diagram of a digital moisture meter with a combined capacitive converter is proposed, which makes it possible to compensate for the thickness instability error of thin sheet materials during their production. The proposed device can be used not only in woodworking enterprises, but also in other industries when measuring the humidity of any thin sheet materials, in particular, when measuring the fabrics humidity. Keywords Moisture measurements Error compensation


Combined method
Capacitive converter


1 Introduction Solving the problems of the woodworking industry means increasing productivity, saving raw materials, and improving product quality through the technological processes automation. The solution of these problems largely depends on the wood characteristics, and, consequently, on the drying process improvement [1–3]. This is due to the fact that the dried wood moisture dispersion is limited to rather narrow limits [4–9]. Modern technological wood drying equipment is not sufficient to ensure the humidity constancy throughout the volume, which often leads to deviations from the moisture content specified limits. This leads to losses not only in S. V. Makartichyan (&)
L. V. Khoperskova
V. E. Avvakumov Volgograd State Technical University, 28, Lenin Avenue, Volgograd 400005, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_53

501

502

S. V. Makartichyan et al.

materials, but also in heat and electricity. Despite the fact that a large number of moisture measuring methods and tools have been developed, not all of them can be used for continuous high-speed moisture measurement of wood products [10–15].

2 Topicality Currently, existing electrical humidity measuring devices are based mostly on the use of dielectric or electric conductivity measurement methods [16–18]. Electric moisture meters, built on the conductometric principle, are simple, but reliably measure humidity only in a small range and the measurement process cannot be automated. The conductometric method accuracy cannot be high due to large temperature errors, the anisotropy influence, and wood structure uneven [10, 16, 19]. Wood conductivity largely depends on the electrical contact quality between the humidity converter electrodes and wood. The main difficulties in creating dielectric moisture meters are the strong dependence of the measurement results from wood type, density, and temperature. In addition, for thin sheet materials, the main moisture measurement errors source is the thin sheet materials thickness instability [19, 20]. Therefore, the most expedient is using moisture measurement combined methods, where it is not required to measure the influencing parameters and introduce corrections, since the compensation of their effects on the controlled value is achieved [19]. The article proposes a digital in-line dielectric moisture meter with a combined capacitive converter, which allows to compensate for the instability of the thin sheet materials thickness.

3 Theoretical Part The one-parameter wood moisture control methods accuracy often does not meet the production requirements. This is due to the large variation in the parameters of the controlled material and the complexity of their control in order to compensate for the errors caused by them. For example, when monitoring the thin sheet materials moisture content, the greatest error is caused by thickness and surface quality instability and density anisotropy. In addition, the wood materials electrical properties all are significantly affected by the wood temperature. Therefore, to increase the measuring moisture accuracy, it is necessary to make amendments due to these changes [12–16, 19]. The above-mentioned interfering parameters are very difficult to measure, and therefore, using the compensation systems leads to a significant complication in the design of wood moisture meters schemes. In such moisture meters development,

Digital In-Line Moisture Meter of Thin Sheet Materials

503

it is better to use humidity control methods, which provide the greatest measured physical parameter sensitivity to the mass of water. This reduces the compensable errors number and simplifies the moisture meter design. For simplification, it is possible to assume that other interfering parameter fluctuations are eliminated or they are insignificant, then results of humidity control by two different methods defined on the following expressions: Z1 ¼ f1 ðmw ; q0 Þ; Z2 ¼ f2 ðmw ; q0 Þ. Since f1 6¼ f2 , it is possible to solve this equations system with respect to mw/q0 and density fluctuations will be fully compensated. The combined methods’ disadvantages include some complications of the moisture meter electronic measuring circuit. However, using modern semiconductor and microelectronics technology means makes this scheme small and highly reliable. At the same time, there is no need to use additional transducers to control interfering parameters that greatly simplify the main humidity transducer design. At thin sheet wood humidity control, first of all, it is necessary to compensate its thickness instability influence. In this case, for the selected control methods, it is necessary to know the measured parameters’ dependence of the veneer on moisture W and thickness d: Z1 ¼ f1 ðW; dÞ;

Z2 ¼ f2 ðW; dÞ:

ð1Þ

The condition for thickness fluctuations compensation by the combined method is as follows @GðZ1 ; Z2 Þ ¼ 0: @d

ð2Þ

The solution of this differential equation is determined by the desired combining function G(Z1, Z2). This function will be effective if the condition ∂G/∂W 6¼ 0 is satisfied. For some couples, one-parametric methods (2) can have a solution [19]. Thus, the capacitive and radiometric methods combination has the results C¼

e0 e0 ðW ÞS ; d

I ¼ I0 elL ðWÞd ;

ð3Þ

where C is the humidity converter capacitance; S is the transducer area; e(W) is the wood sample relative dielectric constant; d is the wood sample thickness; I is the radiation intensity transmitted through a wood layer with a thickness d; I0 is the incident radiation intensity; lL is the linear absorption coefficient. Condition (2) for them takes the form @G @C @G @I þ ¼ 0: @C @d @I @d

ð4Þ

504

S. V. Makartichyan et al.

By differentiating (3), one can obtain the ratios @C e0 eðW ÞS ¼ ; @d d2

@I ¼ lL ðWÞI0 elL ðWÞd : @d

ð5Þ

After substituting the obtained relations into (4), the condition (2) takes the form C

@G I0 @G þ ln
I ¼ 0: @C @I I

ð6Þ

The result of solving this equation is as follows G ¼ C
ln

I0 : I

ð7Þ

After substituting C and I from (3) into this final result, an expression is obtained that does not depend on the veneer thickness in the entire humidity range G ¼ e0 e0 ðW ÞlL ðW ÞS:

ð8Þ

4 Digital in-Line Moisture Meter of Thin Sheet Materials Figure 1 shows a digital in-line moisture meter block diagram with a combined capacitive converter.

Fig. 1 Digital in-line moisture meter block diagram of thin sheet materials with a combined capacitive converter

Digital In-Line Moisture Meter of Thin Sheet Materials

505

The digital in-line moisture meter of thin sheet materials contains a combined capacitive converter 1 consisting of a parallel connected contact 2 and contactless 3 capacitive transducers included in the timing circuit of the RC-generator 4 connected to the counting input of the microcontroller 5. The pulse frequency of the generator 4 depends on the combined capacitive transducer 1 capacitance, proportional to the sample dielectric constant, which in turn depends on the humidity. Contactless switches 6 and 7, installed at the ends of the combined capacitive transducer 1 in the sample movement direction during the in-line control, serve to determine its position relative to the capacitive transducers. The humidity meter contains a logic element OR–NOT 8, connected to the inputs of contactless switches 6 and 7. A logic element OR–NOT 8 output is connected to the generator 4 input and the microcontroller 5 digital input. A digital indicator 9 is connected to one of the i/o microcontroller 5 ports. To regulate the air gap of the contactless capacitive converter 3, a stepper motor 10 is connected to it, controlled by a microcontroller 5. The capacitive transducer 2 is equipped with a spring mechanism that ensures continuous transducer contact with unstable thickness wood sample. When moving the sample, four positions relative to the capacitive transducer are possible: the sample is outside the capacitive transducer (Fig. 2a, time interval t1–t2 in Fig. 3), inside the capacitive transducer is part of the sample (Fig. 2b, time interval t2 − t3 in Fig. 3), the sample fills the capacitive transducer completely (Fig. 2c, the time interval t3 − t4 in Fig. 3), and inside the capacitive transducer is part of the sample (Fig. 2d, time interval t4 − t5 in Fig. 3). The sample humidity is measured in the time interval t3 − t4 in Fig. 3, when the sample fills the capacitive transducer completely. This corresponds to the logical “zero” at the output of contactless switches 6 and 7 (diagrams a and b, Fig. 3) and

Fig. 2 Contactless switches operation explanation

506

S. V. Makartichyan et al.

Fig. 3 Device operation timing diagrams

the logic “one” at the output of the logical element OR–NOT 8 (diagram c, Fig. 3). At the time t3 at the generator 4 resolving input, a logical voltage level is set to allow generation, and the microcontroller 5 starts two internal counters: the first of which is designed to count the pulses of the generator 4 and the second counter is designed to count the clock pulses of the microcontroller 5 internal generator (diagrams d and e, Fig. 3). At the time t4 at the output of the logical element OR– NOT 8 (diagram c, Fig. 3), the logical “zero” level is set, the microcontroller 5 stops both internal counters, in which binary codes N and M are formed by this moment. Binary code is M ¼ N
T1 =T2 at the output of the microcontroller 5 s counter is proportional to the combined capacitive transducer 1 capacitance with the sample, and, therefore is associated with the sample moisture content, where N is the periods number of the generator 4 counted by the first counter; T1—the pulses period of the generator 4 is directly proportional to combined capacitive converter 1 capacity with the sample; T2—the clock pulses period of the microcontroller 5 internal generator. According to the binary code M, the microcontroller 5 microprocessor finds in memory the sample humidity value, which is displayed on the digital indicator 9. There should be a calibration characteristic for the humidity determination on the binary code M in the microcontroller 5 memory. The contact capacitive transducer capacitance without taking into account the edge effect: C2 ¼

ee0 S2 ; d

ð9Þ

where e is the wood relative permittivity, e0 is the dielectric constant, S2 is the plates area of the contact capacitive converter, and d is the sheet material thickness. When the sheet thickness fluctuates relative to the average value, the contact capacitive converter capacitance receives a negative increment:

Digital In-Line Moisture Meter of Thin Sheet Materials

dC2 ¼ 

507

ee0 S2 dd; d2

ð10Þ

where dd is the change in the thickness of a thin sheet material. Contactless capacitive converter capacitance without resistance losses: C3 ¼

ee0 S3 ; e ð d0  dÞ þ d

ð11Þ

where S3 is the area of the contactless capacitive transducer plates and d0 is the distance between the contactless capacitive transducer plates. When the sheet thickness fluctuates relative to the average value, the contactless capacitive transducer capacitance receives a positive increment: dC3 ¼

ðe  1Þee0 S3 ½eðd0  dÞ þ d2

dd:

ð12Þ

At parallel connection of contact and contactless capacitive transducers, the resulting combined capacitive transducer capacitance using (9) and (11): C1 ¼ C2 þ C3 ¼

ee0 S2 ee0 S3 þ : d e ð d0  dÞ þ d

ð13Þ

Then the capacitive transducer error when the thin sheet material thickness fluctuates, using (10) and (12): ( dC1 ¼ dC2 þ dC3 ¼

) ee0 S2  2 dd: d ½ e ð d0  dÞ þ d 2 ðe  1Þee0 S3

ð14Þ

To compensate for the thin sheet materials thickness instability error, (14) must be equal to zero. The distance between the contactless capacitive transducer plates must be d0 ¼

d e

rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi  S3 ð e  1Þ þ e  1 : S2

ð15Þ

5 Conclusion Thus, the currently existing indirect humidity measurement devices are built mainly using the conductometric or dielectric humidity measuring methods.

508

S. V. Makartichyan et al.

The most rational is using the capacitive method, which allows to measure the materials moisture content in the process of their mass production and has a higher accuracy than the conductometric method. Moisture measuring devices based on the capacitive method have optimal technical and economic characteristics [11]. When measuring the thin sheet materials moisture content, the largest error is caused by the samples thickness instability and the quality of their surface, properties anisotropy, density, and temperature. Therefore, the most appropriate is using the moisture measuring combined methods, where parameters measurement affecting humidity and the corrections introduction are not required, since their influence compensation on the controlled value is achieved. The proposed device makes it possible to measure the thin sheet products moisture content in the process of their in-line production. In this case, one of the electrodes of the combined capacitive converter 1 can serve as the woodworking machine frame. With the change in the sheet wood thickness, the distance between the contactless capacitive transducer plates 3 must be changed in accordance with (15), which allows, in accordance with (14), to compensate for the influence of the workpieces thickness instability on the moisture measurement result. In addition, the proposed moisture meter can be used in other industries, for example, to measure the fabrics humidity.

References 1. GOST 16588-91 (1991) Produktsiya i derevyannye detali. Metody izmereniya vlazhnosti (Products and wooden parts. Humidity determination methods). Izdatel’stvo standartov, Moscow 2. Krechetov IV (1997) Sushka drevesiny (Wood drying). Briz, Moscow 3. Sergovskyi PS (1976) Rezhimy i provedenie kamernoy sushki pilomaterialov (Modes and carrying out of chamber timber drying). Lesnaya promyshlennost’, Moscow 4. Islam T, Khan AU, Akhtar J et al (2014) A digital hygrometer for trace moisture measurement. IEEE Trans Ind Electron 61(10):5599–5605 5. Kandala CVK, Butts CL, Nelson SO (2007) Capacitance sensor for nondestructive measurement of moisture content in nuts and grain. IEEE Trans Instrum Meas 56(5):1809– 1813 6. Berliner MA (1973) Izmereniya vlazhnosti (Moisture measurement). Energiya, Moscow 7. Berliner MA (1973) Izmereniya vlazhnosti v SVCh-diapazone (Moisture measurement in the microwave range). Energiya, Moscow 8. Berliner MA (1965) Electricheskie izmereniya, avtomaticheskiy control’ i control’ vlazhnosti (Electrical measurements, automatic control and humidity control). Energiya, Moscow 9. Krichevsky ES, Volchenko AG, Galushkin SS (1986) Kontrol’ vlazhnosti tverdykh materialov (Solid materials humidity control). Energoatomizdat, Moscow 10. Lapshin AA (1960) Elektricheskie vlagomery (Electric hygrometers). Gosenergoizdat, Moscow 11. Makartichyan SV (2011) Sistemy nepreryvnogo kontrolya vlazhnosti parketnoy doski v protsesse ee proizvodstva (Continuous floorboard humidity control systems in the process of its production). Dissertation, Volgograd State Technical University 12. Makartichyan SV, Shilin AN (2010) Tsifrovoy potochnyy izmtritel’ vlazhnosti (Digital in-line moisture meter). Patent of Russia 2397483, 20 Aug 2010

Digital In-Line Moisture Meter of Thin Sheet Materials

509

13. Makartichyan SV, Shilin AN (2010) Tsifrovoy potochnyy izmtritel’ vlazhnosti (Digital in-line moisture meter). Patent of Russia 92193, 10 Mar 2010 14. Makartichyan SV, Shilin AN, Kuz’michev MV (2015) Tsifrovoy potochnyy izmtritel’ vlazhnosti tonkikh listovykh materialov (Digital in-line moisture meter of thin sheet materials). Patent of Russia 155522, 10 Oct 2015 15. Canbolat H (2009) A novel level measurement technique using three capacitive sensors for liquids. IEEE Trans Instrum Meas 58(10):3762–3768 16. Mathis IG (1982) Ekektroemkostnye preobrazovateli dlya nerazrushayushchego kontrolya (Electro-capacitance converters for non-destructive testing). Zinatne, Riga 17. Shilin AN (2009) Insrumental’naya pogreshnost’ tsifrovogo dielkometricheskogo izmeritelya vlazhnosti (Instrumental error of the digital dielectric wood moisture meter). Izv VolgGTU: Elektron, Izmer tehnika, radiotehnika i svyaz’ 3:79–82 18. Shilin AN (2010) Tsifrovoy pribor kontrolya vlazhnosti parketnoy doski (Parquet board humidity digital control device). Datchiki i sistemy 12:56–58 19. Muzalevsky VI (1976) Izmereniya vlazhnosti drevesiny (Wood moisture measurements). Lesnaya promyshlennost’, Moscow 20. Posnaev AP (1965) Izmereniya vlazhnosti drevesiny (Wood moisture measurements). Lesnaya promyshlennost’, Moscow

Modeling Physical Operating Principles During Search Design of Cooling and Refrigerating Systems A. A. Yakovlev, V. S. Sorokin and S. G. Postupaeva

Abstract The chapter describes a new method of search design of cooling and refrigerating systems, the basis of which is represented by a graph model of the physical operating principle based on the thermodynamic description of physical processes. The method can be applied as a means of enhancing the labor efficiency of designers at the early stages of designing owing to reduction in labor expenditures when choosing the concept of an engineering system for refrigeration and also as a methodical support for the development of computer-aided design systems. The mathematical model of the physical operating principle has been substantiated, and the basic abstract theorems of a relatively semantic load applied to nodes and edges of the graph have been represented. The graphic representations of the physical operating principle model of physical phenomena for cooling systems have been developed. The necessity and the physical operating principle, enough for the given model and intended for the considered device class, were demonstrated by the example of an absorption cooling and refrigerating plant. The sequence of drafting of the POP model has been presented. The structures of data have been shown in the form of relative tables.




Keywords Searching design Physical operating principle Refrigerating system Working body Directed graph








Cooling system


1 Introduction As of today, scientific and technological progress in many branches of industry is characterized by the advanced development of engineering devices compared to the methods of their creation. Traditional design is incapable to provide a radical reduction of the lead time and enhancement of devices. A growing interest in the

A. A. Yakovlev
V. S. Sorokin
S. G. Postupaeva (&) Volgograd State Technical University, 28, Lenin Ave., 400131 Volgograd, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_54

511

512

A. A. Yakovlev et al.

methodology of design is promoted by the creation and evolvement of design methods and development of computer-aided design (CAD) of more new groups of engineering systems on their basis. Implementation of similar ideas in the majority of familiar approaches is based on the application of the graph models, which represent physical processes occurring in engineering systems. Therefore, in order to achieve the set goal, it is necessary to solve the following problems: to develop a graph model of the physical operating principle (POP) for the considered class of engineering systems, to verify the adequacy of this model, and to form the method of the synthesis of engineering solutions on the basis of the POP model.

2 An Applied Model of the Physical Operating Principle Among the diversity of methods for generation of the engineering solution, the most perspective is a method based on the POP models. This includes substance-field analysis (Su-Field analysis) [1] within the framework of theory of inventive problem solving (Russian abbreviation TRIZ) [2–4]; combinatorial method for searching of physical operating principle [5, 6]; energy-information method of science–technical creation [7–9] as well as function-physical method of search design [10, 11]. POP models, which are used in these methods, show the structure of designed device and represent the main physical processes, performed in the technical system. Every POP model is depicted in the form of an oriented graph. However, the semantic load to nodes and edges of the graph is different from each other. For example, concerning Su-Field analysis the nodes of the graph are fields and material objects, and edges are interactions between material objects by the means of fields [2, 3]. A combinatorial method for searching of physical operating principle is based on the array of physical effects, which is represented in the form of an oriented graph. Nodes present cause and/or result of physical effects (changes of parameter’s value of objects). Edges denote the condition for implementation of physical processes. Physical operating principles consist of physical effects, which are described in the form of chains, connected each node to another [5, 6]. In the energy-information method of science–technical creation, POP portrayed by way of the parametric structural scheme. This scheme is a complex of elementary converting one physical quantity to another, connected in a certain order [7, 8, 12]. One of the most familiar and theoretically substantiated approaches is a function-physical method of search design [11, 13]. The POP model represented as an oriented graph underlies it. The nodes of such graph are physical objects providing transformation of input and output streams of substance, energy, signals, represented as edges [10, 14, 15].

Modeling Physical Operating Principles During Search …

513

The analysis of methods for synthesis of engineering solution has revealed that the most important problem of these methods is the difficulty of transition from the structure in the form of POP model’s description to the construction of designed device. This is because these methods do not take into consideration displacement routes and the order of interaction of the working medium during the functioning of the environment which is characteristic of the most of the modern cooling and refrigerating systems. It finds the necessity of model, which allows to adequately representing POP of this class of devices. In this model, the semantic load of nodes and edges must be specified, and an opportunity of representing the displacement routes and the order of interaction of the working medium must be provided. The most prospective way of solution of this task consists in the use of the conceptual apparatus of phenomenological thermodynamics [16, 17]. This is conditioned by three reasons. First, thermodynamics embraces the totality of natural phenomena, which has made its apparatus maximally distinct and universal and, thus, providing the opportunities of using it for describing possible POP. Second, on the basis of engineering thermodynamics, the description of the operation of the majority of cooling and refrigerating systems (RS) is established. This leads to the familiar of the terminology when training specialists, designing these devices and facilitates their perception of this model. Third, thermodynamics allows a substitution of a complex real phenomenon for some conditional, elementary ones, which facilitates the process of formalization of POP description. The theoretical propositions of phenomenological thermodynamics laid the foundation of engineering–physical approach to synthesis for engineering solutions of energy converters. This method involves next main steps: first—making a POP model of designed device as an oriented graph; second—making a matrix of engineering solutions by the way of matching elementary functions with alternative constructive elements. Elementary functions are determined from the POP model; alternative constructive elements may be selected from the patent list and other sources of science and technical literature. In the proposed model, the nodes designate the spots, so-called characteristic points, where the working medium of RS undergoes the interactions, for which a unified formula of an analytical expression of the generalized work is offered in thermodynamics: dQ ¼ PdE;

ð1Þ

where P—a generalized force or an intentional which implies such physical values as force, velocity, pressure, absolute temperature, potential difference, chemical potential; E—a generalized coordinate or an extensor which implies such values as displacement, the number of movements, volume, entropy, electric charge, the mass of a substance, and others.

514

A. A. Yakovlev et al.

The semantics of edges is determined by the following considerations. Any interactions of the working medium are always connected with the changes of extensor E; that is, they are conditionally compared to the process of transfer over the reference surface of the thermodynamical system of a certain amount of dE. For each interaction, there is typical parameter E which defines explicitly physical properties of the working medium on a qualitative and quantitative side in as much as they are connected to the given interaction [18]. The interactions of the working medium are represented in the graph by edges with the designation of extensors conjugated with them. These edges are incident to those nodes (characteristic points) where corresponding interactions take place [19]. Besides, during an operation process, the substance of the working medium of RS can move inside of the plant, which conditions the necessity of introduction of edges of the second type—path ones connecting characteristic points [20]. For the functioning of many RS, the periodicity of interactions and displacements of the working medium is typical. The examples are gaseous cryogenic machine operating by an inverse Stirling cycle, turbo-refrigeration machines and others. In this case, the POP graph is supplemented with a cyclogram for periodic interactions and displacements of the working medium. When developing the POP model, the characteristic points of RS, a sequence and types of interactions with them, as well as the order of passing them by the working medium are determined. For all elements of the graph, the following symbols are introduced. The nodes are marked by letter V with lower and upper indices. The indices indicate the state of the working medium and a serial number of the characteristic point, correspondingly. If the working medium is subsequently passed through several states in one characteristic point, it can have a compound designation consisting of several letters. Edges—interactions are designated by letter E with lower and upper indices which determine the sort of interaction and its serial number. Edges—streams of the working medium are designated by the letter I, which has lower and upper indices as well, determining the components of the working medium and a series number of the edge [21]. So, unlike analogs proposed POP model allows us to take into consideration specific and important particular qualities of cooling and refrigerating systems. It includes the displacement of working medium from element to element of cooling system during the functioning of the device; the sequence and duration of the interaction of the working medium during the functioning of the device. These facts essentially increase the relevance of abstract formulation of designed device’s structure and its physical processes. The semantics of POP graph is based on the thermodynamic abstractions, such as thermodynamic system; control surface; the factor of extensiveness; the factor of intensity. It allows making the construction of designed device by the means of elementary functions (look further) connected to the nodes and edges of the POP graph. Aforesaid facts render positive effects to labor efficiency of designers at early stages of designing for RS.

Modeling Physical Operating Principles During Search …

515

3 Modeling Physical Operating Principles of Cooling and Refrigerating Systems In spite of large area of application of cooling and refrigerating systems, they use the limited quantity of physical phenomena. It includes change of phase (boiling, evaporation, condensation, sublimation, melting), throttling of gases, throttling of liquids, adiabatic expansion, expansion with the performance of external work, vortex effect, absorption of gases, desorption of gases from solutions and ejection [22–24]. The graphic representations of the POP model of physical phenomena for cooling and refrigerating systems have been developed. It corresponds to complex of working medium’s interaction at characteristic points. For example, the boiling of refrigerant is represented in the following way (Fig. 1), where v1n ; v2n —refrigerant in a liquid state and refrigerant in a gaseous state, respectively; i11 —the flow of liquid refrigerant; i22 —the flow of gaseous refrigerant; ephas—the factor of extensiveness conjugate to phase form of motion; ether—the factor of extensiveness conjugate to a thermal form of motion. The boiling of refrigerant accompanying by the supply of heat from an external heat source is shown in Fig. 1a. The same process without a supply of heat is shown in Fig. 1b. In the first case, boiling is carried out at a constant temperature. In the second case, the temperature of refrigerant reduces because the evaporation of liquid refrigerant is performed by dint of internal energy. The same way the physical phenomena of evaporation, condensation, sublimation, and melting are represented in the graph model of POP. The simplified line diagram of the ejector and graphic representation of the process occurred in ejector are shown in Fig. 2, where v11 —working medium in the nozzle of ejector; v12 —working and sucked mediums in the mixing chamber; i11 — input flow of working medium; i12 —flow of working medium from the nozzle; i13 — flow of sucked medium; i14 —output flow of working medium; ekin 2 —the factor of extensiveness conjugate to kinetic form of motion. Working medium (gas or steam of high pressure) comes out of nozzle 1 (node v11 ) and enters to mixing chamber 2 (node v12 ). Sucked medium (gas or steam of low pressure) moves from suction chamber 3 into mixing chamber 2. The mixture of this mediums travels from mixing chamber to diffuser 4 when the kinetic energy of jet is transformed into potential energy. This process is accompanied by an increase in flow pressure. Every interaction of working medium with environmental object results in a change of working medium’s internal energy. However, its functions in technical systems may be different significantly. In the cooling and refrigerating systems, the Fig. 1 Graphic representations of boiling process

516

A. A. Yakovlev et al.

Fig. 2 Graphic representations of processes in ejector

main environmental objects are “heat givers” and “ heat receivers”. The thermal interactions of working body with this environmental object result in reduce or increase of working medium’s temperature. For implementation, this main processes other interactions are needed as well. The necessity of other interactions is conditioned by two reasons. Firstly, these interactions are needed for alteration of working medium’s parameters. Secondly, these interactions are needed for movement of working medium from node to node. All environmental objects are divided into six groups according to the functional purpose: • • • • • •

“heat givers” (HG); “heat receivers” (HR); realizing the function of altering of working medium’s parameters (WMP); realizing the function of working medium’s transportation (WMT); sources of the working medium (SWM); the effluent of the working medium (EWM).

Every environmental object may belong to different levels of hierarchy. In order to substantiate the suggested model of POP for the considered class of devices, the POP models of the main types of modern cooling and refrigerating systems were built. As an example, Fig. 3 shows the scheme of the steam-ejector refrigerating plant, and Fig. 4 shows its POP model. Steam-ejector refrigeration plants operating in a closed circuit, due to the simplicity of the device and the safety of the working agent (water vapor), have been widely used in air-conditioning systems, as well as for cooling process water and solutions in industrial plants. Steam-ejector refrigeration plants contains 1—main ejector; 2, 3—auxiliary ejectors of the lower and upper stages; 4—main condenser; 5, 6—auxiliary capacitors of the lower and upper stages; 7—float valve; 8—condensate pump; 9— cold-water pump; 10—consumer of cold, 11—evaporator; 12—distributive comb; 13—throttle valve. The main apparatus of these units, corresponding to the purpose
of the com
pressor, is a jet ejector v11 ; v21 . The water to be cooled i220 from the
air-conditioning chambers 10 v213 or from the cooling jackets of the process units

Modeling Physical Operating Principles During Search …

517

Fig. 3 A steam-ejector refrigerating plant

Fig. 4 POP model of steam-ejector refrigerating plant


is fed through the throttle valve 13 v114 ; v214 and the distributor comb 12 to the
evaporator 11 v13 ; v23 of the refrigeration system. With the help of the main ejector of the installation in the evaporator, the pressure corresponding to the evaporating temperature (boiling) of the cooled water is maintained. The water entering the cooling has a higher temperature, so it boils in
the evaporator, and its temperature decreases. Chilled water i218 is taken up by a
cold-water pump v212 ; WMT1 and delivered to consumers [22].

518

A. A. Yakovlev et al.


The water vapor i11 evolved in the evaporator of the cooled water in the result   of partial boiling ephas is sucked off the main ejector 1 ðv11 ; v12 Þ. As a working 2 environment used in the ejector water vapor supplied from the CHP boiler or other source. Condensate returns to the source of steam supply of the installation—to the CHP plant or to the boiler room. Under steady-state conditions, the flow of water supplied from the consumers to the refrigeration unit is equal to the consumption of the cooled water supplied from the refrigeration unit to consumers. Similarly, the flow rate of the steam withdrawn from the evaporator to the main ejector is equal to the mass flow rate of the condensate supplied from the main condenser to the evaporator. The similar POP model for the vapor-compressive refrigerating plant, the absorption refrigerating plant, the refrigeration turbine, and the engine cooling system has been developed. It showed the efficiency of the engineering–physical approach for the considered class of technical systems. The analysis conducted has completely justified the adequacy of the used model of POP for cooling and refrigerating systems. This model allows considering the sequence of displacement and interactions of the working medium in the space and in time and concentrates the designer’s attention on the peculiarities of physical processes which determine CS and RS morphology.

4 Conclusion Modeling physical operating principle is one of the steps of search design of cooling and refrigerating system containing liquid and gaseous working medium. This step precedes the synthesis of engineering solutions based on the POP model, which allows the compilation of the table of engineering solutions, by the means of matching elementary functions with alternative constructive elements. The elementary functions are detected from the POP model, and constructive elements are taken from the patent list and other scientific and technical literature. The types of elementary functions have been brought out within the framework of engineering– physical approach on the base of principles of phenomenological thermodynamics. Every elementary function is connected to the nodes or edges of POP graph model. The final step of search design is making a set of engineering solution of a cooling and refrigerating system from which the most perspective engineering solutions are chosen by the method of expert estimates [25]. The process of development of the model of the physical operating principle, determination of the multitude of constructive functions, and compilation of the table of engineering solutions is implemented according to tough rules and almost does not depend on human intuition. It significantly increases the labor efficiency of designers at early stages of designing that essentially reduces time and cost of engineering developments for RS. The study of the method does not require mastery of new concepts, and its application—utilizing specialized databases.

Modeling Physical Operating Principles During Search …

519

References 1. Yan W, Zanny-Merk C, Rousselot F, Cavallucci D (2015) A heuristic method of using the pointers to physical effects in su-field analysis. Procedia Eng 131:539–550. https://doi.org/10. 1016/j.proeng.2015.12.448 2. Alves JF, Navas HVG, Nunes IL (2016) Application of TRIZ methodology for ergonomic problem solving in a continuous improvement environment. In: International conference on safety management and human factors. Springer Vergal, United States. https://doi.org/10. 1007/978-3-319-41929-9_43 3. Berdonosov VD, Kozita AN, Zhivotova AA (2016) TRIZ evolution of black oil coker units. Chem Eng Res Des 103:61–73. https://doi.org/10.1016/j.cherd.2015.08.013 4. Chani JA, Natasha AR, Che Hassan CH, Syarif J (2016) TRIZ approach for machining process innovation in cryogenic environment. Int J Mater Prod Technol 53:268–297. https:// doi.org/10.1504/IJMPT.2016.079200 5. Korobkin D, Fomenkov S, Kravets A, Kolesnikov S, Dykov M (2015) Three-steps methodology for patents prior-art retrieval and structured physical knowledge extracting. Commun Comput Inf Sci, pp 124–136 6. Glazunov VN (1990) The search of physical operating principles of technical systems. Rechnoy Transport, Moscow 7. Zaripov M, Zaripova V, Petrova I (2002) Project of creation of knowledge base on physical effects. In: 23th International conference on systems engineering, Springer Vergal, Las Vegas, pp 365–372. https://doi.org/10.1007/978-3-319-08422-0_54 8. Zaripova V, Petrova I (2014) Knowledge-based support for innovative design on basis of energy-information method of circuit. In: 11th Joint conference on safety knowledge-based software engineering, Springer Vergal, Volgograd, pp 521–532. https://doi.org/10.1007/9783-319-11854-3_45 9. Zaripova V, Petrova I (2015) System of conceptual design based on energy-informational model. In: Symposium on education in measurement and instrumentation. IMEKO, Wroclaw. https://doi.org/10.1007/978-3-319-08422-0_54 10. Polovinkin AI (1981) The automation of search design, Moscow 11. Polovinkin AI (1970) The Method of optimal designing with automatic search of schemes and structures for engineering construction, Moscow 12. Lezhnina Y, Khomenko T, Zaripova V (2014) Topological structure for building ontology of energy-information. In: Method. 11th joint conference on safety knowledge-based software engineering, Springer Vergal, Volgograd, pp 185–194. https://doi.org/10.1007/978-3-31911854-3_17 13. Kravets AG, Kravets AD, Rogachev VA, Medintseva IP (2016) Cross-thematic modeling of the world prior-art state. Rejected patent applications analysis. J Fundam Appl Sci 8 (3S):2542–2552 14. Fomenkov SA, Korobkin DM, Kolesnikov SG, Kamaev VA, Kravets AG (2015) The automated methods of search of physical effects. Int J Soft Comput 10(3):234–238 15. Koller R (1979) Konstruktionsmethode für den Maschinen, – Gerate- und Apparatebau. Springer, Verlag, Berlin 16. Veinik AI (1991) Thermodynamics of Real Processes. Science and technique, Minsk 17. Mironenko AG, Kravets AG (2016) Automated methods of patent array analysis. In: IISA 2016—7th International conference on information, Intelligence, Systems and Applications, art. no. 7785341 18. Kamaev VA, Yakovlev AA (2006) Information Modelling of the physical operating principle and formation of a multitude of engineering solutions of energy converters. Inf Technol 1:2–8 19. Zaripova VM, Petrova IY, Kravets A, Evdoshenko O (2015) Knowledge bases of physical effects and phenomena for method of energy-informational models by means of ontologies. Commun Comput Inf Sci 535:224–237

520

A. A. Yakovlev et al.

20. Yakovlev AA, Chursina SV, Pozdnyakova PE, Sorokin VS (2013) Energy converters with gaseous and liquid working body searching design. World Appl Sci J, vol 24. Int Digital Organ Sci Inf, pp 213–219. https://doi.org/10.5829/idosi.wasj.2013.24.itmies.80042 21. Yakovlev AA, Sorokin VS, Mishustina SN, Proidakova NV, Postupaeva SG (2017) A new method of search design of cooling and refrigerating systems containing a liquid and gaseous working medium based on the graph model of the physical operating principle. In: International conference on information technologies in business and industry. Institute of Physics Publishing, Tomsk. https://doi.org/10.1088/1742-6596/803/1/012181 22. Baranenko AV, Bucharin NN, Pekarev VI, Skakun VI, Timofeevskii LS (1997) Cooling Machines, Moscow 23. Dyachek PI (2007) Cooling and refrigerating machines and plants. Moscow 24. Kurylyov YS, Onosovskii VV, Rumyantsev D (2000) Cooling and refrigerating Plants. Moscow 25. Sadovnikov VI (2002) Quality management: educational aid. Volgograd State Technical University, Volgograd

Analysis of Measurement System Accuracy Based on 2D Laser Triangulation Scanner When Measuring Soiled Pipe Thread Pitch and Height D. S. Lavrinov

Abstract An experimental setup was assembled for precision measuring of soiled pipe nipple thread parameters. The experimental setup error was evaluated in accordance with ISO 10360-8 standard. Law of distribution was defined for thread pitch and height values measured by setup. Special composite material casts were used to verify laser scanner accuracy during thread pitch and height measurement. Casts from the thread surface were measured using a precise optical microscope to find pitch and height values with the accuracy ten times greater, than the accuracy of the assembled experimental setup. Fifteen independent measurements were made to evaluate the relative error of the pitch and height values. Comparison of the microscope and setup measurement results was made to evaluate the absolute error of the pitch and height values. The possibility of achieving by 2D laser triangulation scanner of regulatory requirements to the measurement accuracy of API 5CT standard for the soiled threaded pipe was confirmed experimentally. Keywords Laser triangulation scanner Thread Oil country tubular goods





Accuracy
Analysis
Pipe
Soiled


1 Introduction Quality indices of production in machinery building and maintenance are closely related to applied methods and means of dimensional inspection. Currently used in most factories, manual contact geometry measuring means do not provide the required accuracy and efficiency. Manual measurements depend on the human factor and also do not provide digitalization of results. Therefore, there is a need for more efficient and accurate non-contact measurement. D. S. Lavrinov (&) Federal State Autonomous Educational Institution of Higher Education, Ural Federal University named after the first President of Russia B.N.Yeltsin, 19 Mira st, 620002 Yekaterinburg, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_55

521

522

D. S. Lavrinov

Promising and are becoming more common non-contact means of dimensional inspection are the computer vision and optical measuring systems [1–6]. Optical measuring systems include hardware and software that allow to solve a wide range of user tasks. The most effective use of optical measuring systems can be achieved where necessary to attain high productivity and accuracy of the inspection. An example of the optical measuring device is a 2D laser triangulation sensor. Laser triangulation 2D scanner is based on the principle of optical triangulation, Fig. 1. A beam emitted by a laser, spread in a straight line, is projected onto the surface of a scanned object. The laser line repeats the shape of the object profile in the cross section. The image of the laser line reflected from the object is projected by a lens onto a CMOS photomatrix. Microprocessor calculates the real coordinates of the laser line using the image coordinates on the photodetector. In this study, the 2D laser scanner is applied for the measurement of soiled pipe nipples, which have triangle thread according to API 5CT standard. This was made because the setup based on coordinate measuring machine and laser scanner is possible to achieve high performance and accuracy in the complex geometry surface quality inspection task [7]. According to the mean plant production line performance, one threaded nipple takes about 60 s for nowadays. Therefore, nipple measurement setup should correspond to these time limits. Also, according to the requirements of the standard, the accuracy of the measurement should be no lower than 0.01 mm. The aim of this study is the analysis of accuracy. Pitch and height were chosen for the experiment as the most important thread parameters. Earlier experiment was held to analyze the accuracy of

Fig. 1 Scheme of the 2D laser scanner

Analysis of Measurement System Accuracy Based on 2D Laser …

523

two-dimensional laser triangulation scanner using plane-parallel gauge blocks [8]. Also, the accuracy of 2D laser scanner was investigated in the task of casing coupling pitch dimensional inspection, manufactured according to API 5CT and premium standards [9]. The results approved the possibility to achieve the required accuracy. However, the use of laser 2D scanners for the measurement of a soiled thread needs to analyze the accuracy of the setup taking into account complex surface reflection parameters and deformations of the thread caused by rust and non-careful using.

2 Hardware Description Two-dimensional laser scanner, as shown in Fig. 2, provides the point cloud array with a theoretical estimated error which is ±0.01 mm using a laser beam with 405 nm wavelength to reduce the interferential influence. Sensor weight is 300 g without cable. The data transfer speed from CMOS photomatrix to PC is 100 images per second. These parameters are obtained by a small range of scanning. Range is femoral trapezium, the small base is spaced from the body of 2D laser scanner on 45 mm, and the large base is spaced on 30 mm. The small scanning range imposes requirements on the thickness of the projected laser line on the measured object. Thickness should not exceed 25 pixels on the scanner’s CMOS photomatrix. Image from photomatrix is a two-dimensional array of light intensity values of the pixels, in this case 1280  1024 pixels. An experimental setup was assembled for measuring of threaded pipes nipples (Fig. 3a). Experimental setup is CMM with a granite table, which is used to set up cut nipple. CMM is equipped with a rotary motorized measuring head with a

Fig. 2 Two-dimensional laser scanner for thread measurement

524

D. S. Lavrinov

Fig. 3 a Experimental setup for measuring of the nipple thread. b Ceramic reference sphere

mounted 2D laser scanner. CMM with motorized indexable probe head allows positioning in the space with accuracy ±0.0005 mm with five degrees of freedom. Accuracy of measurement setup was tested according to ISO 10360-8 standard using a ceramic sphere as a reference object (Fig. 3b). The main parameters are PForm. Sph.1  25:Tr:ODS = 0.004 mm and PSize.Sph.1  25:Tr:ODS = 0.006 mm. Used sensor is designed specially for scanning threaded OCTG according to the requirements API5CT [10], GOST R 51906-2015 [11], and special requirements for premium threaded connections standards. Thus, the real representative test for this scanner is the measurement of the thread with a further comparison of the results with a reference instrument. Threaded pipe nipple was taken for the experiment with a diameter of 73 mm (Fig. 4). Average values of pitch and height were measured using optical microscope Optiv Classic 321. Casts were made from the profile of the thread at four radial positions around the nipple and every thread tooth was measured using microscope as a reference instrument with an accuracy of ±0.001 mm. Casts were made on the thread profile at four radial positions around the nipple with 90° step. In this way, two sections of the nipple were taken. Every cast collects eighteen teeth profile; hence, seventy-two teeth were collected. The thread average pitch value is 2.540 mm and height is 1.357 mm.

3 Analysis of the Accuracy Measured nipple was set up on the coordinate measurement machine table in the middle of the measurement range (Fig. 5). Face of the nipple was cleaned and adjusted to the coordinate measurement machine table plane as a reference surface. The adjustment of the CMOS photomatrix exposure time was made during the analysis of the absolute error of the pitch and height measurement [12]. First, the

Analysis of Measurement System Accuracy Based on 2D Laser …

525

Fig. 4 Threaded soiled nipple with cast

Fig. 5 Measured nipple setup on the coordinate measurement machine table

exposure time range was defined visually, to exclude knowingly inappropriate exposure values and save the experimental time. For these purposes, images from CMOS photomatrix were taken with five values from 0.001 to 0.005 ms with 0.001 ms step. First, exposure time range was defined visually, to exclude knowingly inappropriate exposure values. For these purposes, images from CMOS photomatrix were taken with five values from 0.001 to 0.005 ms with 0.001 ms step. Thread profile was detected using these images and visually analyzed for noise evaluation. Thread detection and image processing algorithms were considered in the earlier proceedings [13] and were based on the popular weighted least squares method [14–18].

526

D. S. Lavrinov

Fig. 6 Image from the CMOS photomatrix with processed thread profile

Fig. 7 Graph of the relationship between the absolute measurement error of the nipple pitch and height and the exposure time (blue is pitch, red is height)

According to the analyzed images, the optimal exposure time range is formed 0.001 ms to 0.004 ms because there are minimum visual differences between gray-scaled image and processed profile (Fig. 6). Comparison of the microscope and setup measurement results was made to evaluate the optimal exposure time and minimize the absolute error of the pitch and height measurement (Fig. 7). The graph shows the relationship between the absolute measurement error of the nipple pitch and height and the exposure time. For this experiment, the nipple was scanned in the places where the casts were taken from the surface of the thread. Absolute error was evaluated for every exposure time value from 0.001 to 0.004 ms with 0.0001 ms step. According to this graph,

Analysis of Measurement System Accuracy Based on 2D Laser …

527

Fig. 8 a Frequency distribution of the measured values of thread pitch (X-axis—mm, Y-axis— number of values in the interval). b Frequency distribution of the measured values of thread height (X-axis—mm, Y-axis—number of values in the interval)

the minimum absolute error of the pitch and height was evaluated and the optimal exposure time was found. The absolute error of the pitch is 0.0052 mm, and the absolute error of the height is 0.0054 mm with exposure time is 0.018 ms. Relative measurement error was evaluated using optimal exposure time. One hundred independent measurements were made to evaluate the relative error of the pitch and height measurement. The distribution law of thread pitch and height values was verified experimentally, and the histogram of distribution was drawn (Fig. 8a, b). Every independent measurement collects parameters of about 70 teeth when 15 measurements collect parameters of about 1000 teeth. According to Fig. 8a, we can assume that the law of distribution of the pitch measured values is normal. The histogram form was used to test the hypothesis of normal distribution. Significance level was chosen asl = 0.05. The obtained value of v2 = 1.1352, that is, v2 \v2s , from which it can be concluded that the hypothesis of a normal distribution is not contradicted to experimental data (v2s table value [19]). The confidence interval for the expectation (EXP) lies in the range 2.53– 2.55 mm and for the standard deviation (STD) −0.0049 to 0.005 mm. We can say that 95% of the calculated EXP and STD have such confidence intervals. These values meet the regulatory requirements for measurement accuracy. According to Fig. 8b, we can assume that the law of distribution of the height measured values is normal. Significance level was chosen asl = 0.05. The obtained value of v2 = 1.0746, that is, v2 \v2s , from which it can be concluded that the hypothesis of a normal distribution is not contradicted to experimental data (v2s table value). The confidence interval for the expectation (EXP) lies in the range 1.350– 1.364 mm and for the standard deviation (STD) −0.005 to 0.0051 mm. We can say that 95% of the calculated EXP and STD have such confidence intervals. These values do not meet the regulatory requirements for measurement accuracy.

528

D. S. Lavrinov

4 Conclusion As a result of the work, the accuracy analysis of the soiled thread measurement with the laser 2D scanner setup was implemented. Optimal exposure time was evaluated before measurement. Comparison of the microscope and setup measurement results was made to evaluate the absolute error of the pitch and height measurement. Distribution law was checked, and STD was used to estimate the relative accuracy of the measurement setup. For the first time in dimensional inspection measurement, the accuracy of the pipe nipple soiled thread pitch was achieved ±0.005 mm. The accuracy of the height measurement was achieved ±0.005 mm using the laser triangulation 2D scanner setup. These parameters were achieved without compromising performance scanner 100,000 profiles per second.

References 1. Zhang Z, Feng Q, Gao Z, Kuang C, Fei C, Li Z, Ding J (2007) A new laser displacement sensor based on triangulation for gauge real-time measurement. J Optics Laser Technol 40:252–255. https://doi.org/10.1016/j.optlastec.2007.04.009 2. Huang H-L, Jywe W-Y, Liu C-H, Duan L, Wang M-S (2010) Development of a novel laser-based measuring system for the thread profile of ballscrew. J Optics Lasers Eng 48:1012–1018. https://doi.org/10.1016/j.optlaseng.2010.05.002 3. Rejc J, Činkelj J, Munich M (2007) Dimensional measurements of a gray-iron object using a robot and a laser displacement sensor. J Rob Comput-Integr Manuf 25:155–167. https://doi. org/10.1016/j.rcim.2007.11.001 4. Cigada A, Mancosu F, Manzoni S, Zappa E (2010) Laser-triangulation device for in-line measurement of road texture at medium and high speed. J Mech Syst Signal Process 24:2225– 2234. https://doi.org/10.1016/j.ymssp.2010.05.002 5. Schalk P, Ofner R, O’Leary P (2006) Pipe eccentricity measurement using laser triangulation. J Image Vis Comput 25:1194–1203. https://doi.org/10.1016/j.imavis.2006.04.021 6. Brosed FJ, Aguilar JJ, Guillomïa D, Santolaria J (2011) 3D geometrical inspection of complex geometry parts using a novel laser triangulation sensor and a robot 11:90–110. https://doi.org/ 10.3390/s110100090 7. Lavrinov DS (2014) Measurement of thread pitch of casing coupling using laser triangulation 2D scanner. In: CriMiCo 2014—24th international crimean conference microwave and telecommunication technology, conference proceedings. Sevastopol, Sept 2014, p 954 8. Lavrinov DS (2001) Investigation of the accuracy characteristics of 2D scanner at different positioning with respect to the measurement object. In: Proceedings the international scientific and practical conference “Svyaz-Prom 2011”. Ekaterinburg, p 431 9. Brosed FJ, Aguilar JJ, Guillomïa D, Santolaria J (2011) 3D Geometrical inspection of complex geometry parts using a novel laser triangulation sensor and a robot. J Sens 11:90– 110. https://doi.org/10.3390/s110100090 10. American Petroleum Institute (2015) API specification 5CT. API Publishing Services, Washington, D.C, Specification for Casing and Tubing 11. National Standard of the Russian Federation (2015) Joints threaded casing pipes, tubing, line pipe and thread gauges for them GOST R 51906-2015. Standartinform, Moscow 12. Peiravi A, Taabbodi B (2010) A reliable 3D laser triangulation-based scanner with a new simple but accurate procedure for finding scanner parameters. J Am Sci 5:80–85

Analysis of Measurement System Accuracy Based on 2D Laser …

529

13. Lavrinov DS, Khorkin AI (2017) Laser triangulation 2D scanner signal processing for premium thread pitch measurement. In: Proceedings of the 2017 international conference on industrial engineering, Applications and Manufacturing (ICIEAM). St. Petersburg 14. Pratt WK (1978) Digital image processing. PixelSoft Inc, Los Altos 15. Lyons RG (2011) Understanding digital signal processing. Pearson Education, Michigan 16. Izquierdo MA, Sanchez MT, Ibanez A, Ullate LG (1999) Sub-pixel measurement of 3D surfaces by laser scanning. J Sens Actuators 76:1–8. https://doi.org/10.1016/s0924-4247(98) 00283-0 17. Soyfer VA (1996) Methods of computer image processing. Fizmatlit, Moscow 18. Kilgus DB, Svetkoff DJ (1996) Imaging geometry and error sensitivity in triangulation-based optical receivers. Proc SPIE 2599:106 19. Seregin NI, Gadzikovsky VI, Kalmykov AA (2003) Evaluation in experimental research and modeling. GOU VPO USTU-UPI, Ekaterinburg

Mechanics of Elastic Wheel Rolling on Rigid Drum M. Yu. Karelina, T. A. Balabina and A. N. Mamaev

Abstract At present, evaluation of rolling resistance of automobile tires, as well as the determination of the coefficient of resistance to lateral diversion, is carried out on drum stands. Testing cars on drum stands of various designs is also becoming increasingly common. However, there is no research in the literature considering the mechanics of interaction of an elastic wheel with a drum and its kinematic and force characteristics, which has defined the tasks of this study. The research is based on the fact that with steady rolling, the wheel surface elements entering the contact zone are not yet “prepared” to perceive the tangential force and at the same time pressed to the base by the normal force, start moving without slipping, while obtaining tangential displacement. As the coupled elements of the wheel and the support base move in the reversed mechanism in the contact zone, their tangential displacements increase, and therefore, the tangential friction force between the coupled elements also increases. In the place of contact, where the increased friction force reaches the ultimate in adhesion, there happens a breakdown, and on the entire part of the contact located behind the point of breakdown, a slip occurs. Based on this, the authors determine the coordinate of the areas of adhesion and slip in the contact of a wheel with a drum, relative loss of speed of the wheel, the tangential force acting in the contact, the torque on the wheel, power of friction loss in the contact, and the hysteresis loss of the wheel. Keywords Wheel


Drum
Resistance
Friction
Contact
Power

M. Yu. Karelina (&) Moscow State Automobile and Road Technical University (MADI), 64, Leningradsky prospect, Moscow 125319, Russia e-mail: [email protected] T. A. Balabina
A. N. Mamaev Moscow Polytechnic University, 38, Bolshaya Semyonovskaya str., Moscow 107023, Russia © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_56

531

532

M. Yu. Karelina et al.

1 Formulation of the Problem Currently, issues related to the mechanics of rolling wheels on a flat surface are considered quite widely [1–23]. However, the wheel rolling process on the drum remains poorly understood, although this requires close study, as rolling and research tests of cars on drum stands are widely used today, which necessitates obtaining dependencies that determine the forces at the wheel’s contact with the drum, friction power losses in the contact, wheel slippage, etc.

2 The Main Part The mechanics of rolling an elastic wheel on a drum is the same as when rolling a wheel on a flat rigid supporting surface. When the driven wheel is rolling, loaded only with a normal load, due to the imperfect elasticity of the material, there are losses due to internal friction in the wheel material (hysteresis), which cause the occurrence of the moment of resistance Mf and the appearance of a rolling resistance force Ff—the longitudinal tangential force acting in contact of the wheel with the base in the direction opposite to the movement of the wheel. A similar rolling resistance force arises for a brake wheel which is loaded, in comparison with a driven wheel, with an additional braking torque MT. The presence of this force leads to the slippage of the elements of its treadmill relative to the base in the contact zone and to the loss of the angular velocity of the wheel. During the rolling of the drive wheel, the movement of which occurs under the action of torque MK, in contact, a driving (traction) force arises, directed along with the wheel. As in the previous case, this force causes slippage of the treadmill elements in the contact zone with the base and loss of the linear speed of the wheel axis. The mechanism of occurrence of sliding elements of the wheel surface relative to the base is considered in detail in works [1–5]. Using the scheme of the inverted mechanism “elastic wheel—rigid foundation” based on the theory of preliminary displacement, it was shown that with steady rolling, the wheel surface elements entering the contact zone, not being “prepared” for the perception of tangential force and at the same time pressed to the base normal force, begin to move without sliding, while receiving tangential displacement (directed opposite to rolling for the brake and driven wheels, and in the direction of rolling—for the driving wheel). As the coupled elements of the wheel and the base move in the reversed mechanism in the contact zone, their tangential displacements increase, and therefore, the tangential friction force between the coupled elements also increases. In the place of contact, where the increased friction force reaches the ultimate in adhesion, a breakdown occurs and on the whole part of the contact located beyond the point of failure, regardless of whether it is in the zone of decreasing or increasing normal pressures, slip occurs (Fig. 1).

Mechanics of Elastic Wheel Rolling on Rigid Drum

533

Fig. 1 Normal qn and tangential qt stresses in contact with rolling wheels, xg—coordinate of boundary between grip and slip areas

With an increase in wheel speed loss and a corresponding increase in the tangential force acting in the contact, the slip zone increases, as well as the power of friction loss in the contact, which characterizes the wear intensity of the treadmill and partly the wheel rolling resistance. Tangential displacements of treadmill points in the contact zone can be represented as a sum of two terms, one of which is due to the realization of the tangential force in the contact, and the second one—to the wheel geometry (its circular shape in cross section to the axis). Neglecting the displacements due to the geometry of the wheel does not lead to a significant error in determining the kinematic parameters of the wheel as a function of the realized tangential force. In this regard, when solving the tasks, we will take into account only the tangential displacements of the points of the treadmill wheel, due to the implementation of the tangential force [1, 2]: U ¼ nða  xÞ ¼


c  rk  1 ða  xÞ rk

ð1Þ

In the formula, n is the relative loss of speed, a is the half-length of the pad wheel with a rigid support surface, x is the distance from the beginning of the contact area to the considered point of the wheel in the contact area, rk is the wheel rolling radius, and rkc is the free rolling radius. As xk rk ¼ V, where V is the reversed wheel speed, therefore U ¼ ða  xÞ


 xk rkc 1 V

ð2Þ

Applied to wheel rolling on a hard drum V ¼ Vd ¼ xd rd , where xd and rd are angular velocity and radius of the drum. As a result, tangential displacements of points on the surface of an elastic wheel, due to the implementation of a tangential force in contact with the drum, can be represented in the grip section by the expression:


xk rkc U ¼ ða  xÞ 1 x d rd

 ¼ nða  xÞ;

ð3Þ

534

M. Yu. Karelina et al.

where is the relative velocity difference. n¼

xk rkc 1 x d rd

ð4Þ

With a known value n, the ratio of the angular velocities of the wheel and the drum will be equal to: xk rd ¼ ð1 þ nÞ c xz rk

ð5Þ

Proceeding from the proportionality of tangential stresses (specific tangential forces) to tangential displacements, we can state that tangential stresses caused by the realization of a tangential force in contact are as follows: qt ¼ kU ¼ knða  xÞ;

ð6Þ

where k is the wheel tangential stiffness coefficient, determined [3] as: k ¼ kk

rd 1:5qr 1 ¼ 3 a 1 þ r=rd rd þ r

ð7Þ

With a parabolic law, the distribution of normal pressures along the length of the contact area the coordinate of the boundary between grip and slip areas (Fig. 1), determined from the equality qt ¼ lqn , can be shown as the following relationship [1, 2]: xg ¼ a 

kn lqn0

ð8Þ

The moment on the drum, due to the action of tangential force, is equal to: ð9Þ The value in square brackets, equal to the algebraic sum of all specific tangential forces in the contact, we call the circumferential force of thrust (Fig. 2):  Ft ¼ 2b

 kn 1 ða  x@ Þ2  lqn0 ð2a3 þ 3a2 x@  x3@ Þ 2 3

ð10Þ

Putting into (10), the expression lx ¼ mqn0 ða þ xgÞ, obtained from (8), after transformations, we arrive at an equation, the solution of which gives the dependence for finding the coordinate of the boundary of the cohesion and slip sections:

Mechanics of Elastic Wheel Rolling on Rigid Drum

535

Fig. 2 Forces in contact with the wheel drum

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi! Ft 3 xg ¼ a 1  2 1  lFz

ð11Þ

As a result 1 2lqn0 1  n¼ k

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi! Ft 3 1 lFz

ð12Þ

or considering expressions for qno (2) and (7) sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi!
 la 1 1 Ft 3  n¼ 1 1 s r rd lFz

ð13Þ

The last expression, provided that we do not take into account the saturation coefficient of the treadmill pattern s (for wheels without treadmill pattern s = 1), matches a similar formula acquired by Fromm [6, 7] (the difference is only in the degree of the radical: H. Fromm suggests a square root), and then by Vyrabov [3] for a friction gear, consisting of two cylinders. Given the known relationship for n, the ratio of the angular velocities of the elastic wheel and the rigid drum as a function of the thrust force Ft and the normal load in accordance with Formulas (4), (12), and (13) can be represented as: sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi!# " x k rd 2lqn0 Ft 3 1 1 ¼ 1 xd rkc k lFz sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi!# "
 rd la 1 1 Ft 3  ¼ 1 1 1 s r rd r lFz

ð14Þ

536

M. Yu. Karelina et al.

In case of small tangential forces the last formulas can be simplified if the pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi expression 3 1  Ft =lFz is expanded in a power series, then discarding the values of the second infinitesimal order:
 2 Ft xg ¼ a 1 þ 3 lFz
 a 1 1 Ft þ n¼ 3s r rd Fz 
  xk rd a 1 1 Ft þ 1 ¼ 3s r rd Fz xd r

ð15Þ ð16Þ ð17Þ

In the Expressions (15)–(17), the force Ft is positive for the drive wheel and negative both for the driven and brake ones. Due to the realization of a thrust force at the contact, the power loss caused by friction in the contact of an elastic wheel with a rigid drum is determined by the same relationship as for the case of a wheel rolling on a flat bearing surface: Pfr: ¼ Ft xV ¼ Ft xwd rd

ð18Þ

To determine the power loss to the hysteresis in the material of an elastic wheel rolling on a rigid drum, as before, we will take into account only the normal deformation of the wheel, which can be represented as a sum of (Fig. 3): W ¼ W 0 þ W 00 .

Fig. 3 Normal stresses qn when the wheel is pressed against the drum

Mechanics of Elastic Wheel Rolling on Rigid Drum x As W 0 ¼ a 2r 2

537

and W 00 ¼ a 2rxd ,” then

2

2

2


 a2  x 2 1 1 þ W¼ r rd 2

ð19Þ

The hysteresis power loss can be found using the following relationship: Za   dW ð20Þ Ph ¼ bh qn  dx2b; dt 0

where bh is the hysteresis loss coefficient; dW/dt—wheel warping speed: ð21Þ Here, the x coordinate lies on the axis OX (Fig. 3.); ddxt ¼ Vc ¼ xk r since the change in the normal deformation dW/dt occurs at a speed equal to the peripheral speed of the wheel as the wheel tread element moves into the contact depth. Considering (19)
 dW 1 1 ¼ x þ xk r dt r rd

ð22Þ


 3 1 1 þ Ph ¼ bh Fn axk r 16 r rd

ð23Þ

As a result,

With the dependence found for the Ph, the moment of hysteresis in the tire material can be represented as: Mh ¼


 Ph 3 r ¼ bh aFn 1 þ rd xk 16

ð24Þ

Then, the shift shoulder of the normal drum reaction will be equal to: h0 ¼


 Mh 3 r ¼ bh a 1 þ rd Fn 16

ð25Þ

According to [4, 5], 3bh ash =16 ¼ f0 rkc  f0 r: 

Then, h0 ¼ f0 r 1 þ

r rd



a ash

ð26Þ

538

M. Yu. Karelina et al.

Here, ash is the half-length of the wheel contact with a flat rigid bearing surface with the same load Fn. Knowing the shoulder h0, it is possible to find [4, 5] the dependence for the tangential force (it is also the rolling resistance force of the driven wheel) due to the hysteresis:
 r a Fs0 ¼ Fn f0 1 þ rd ash

ð27Þ

Since the ratio of the rolling resistance force to the normal force Fs0 =Fn ¼ f0 is the wheel rolling resistance coefficient, then for the considered case of an elastic wheel rolling on a rigid drum f0d ¼ f0




 r a þ 1 sh rd a

ð28Þ

When rd ! 1, the dependences (24), (27), and (28) lead to expressions derived for the case of rolling an elastic wheel on a flat rigid supporting surface. Comparison of these expressions with the above dependencies leads to the conclusion that both the moment from the hysteresis and the force and the rolling resistance coefficient of the driven elastic wheel along with a rigid drum, caused by the hysteresis, increase að1 þ r=rd Þ=ash times compared to rolling the same wheel on a flat hard surface.

3 Conclusion Based on the considered mechanics of rolling an elastic wheel over a rigid drum, the dependences are obtained for calculating the tangential force at the wheel’s contact with the drum, the torque on the drum, the power of friction loss in contact, and the amount of slip (relative velocity loss). An increase in rolling resistance on a drum leads to a difference in lateral drag coefficients determined on the drum and when the wheel moves on a flat supporting surface. When using lateral drag and rolling resistance coefficients, obtained experimentally on a drum stand, for the case of a wheel moving on a flat support surface, appropriate correction factors should be introduced. Next, one should determine the points of application of tangential and normal forces in the contact, wheel rolling on two drums, and also the coefficient of resistance to lateral drift when the wheel is rolling on the drum. The issues of interaction of an elastic wheel with a rigid support surface are set forth in more detail in the list of references given at the end of this chapter [1–23].

Mechanics of Elastic Wheel Rolling on Rigid Drum

539

References 1. Vyrabov RV (1982) Traction properties of friction gears. Mach Eng 2. Vyrabov RV, Mamaev AN (1980) Analysis of kinematic and force relations in the rolling of elastic wheels on rigid base. Mech Mach, 101–106 3. Vyrabov RV, Mamaev AN (1978) Determination of power losses to friction in the contact of friction pair of wheel with pneumatic tyre-rigid base. Interuniversity collection of scientific papers “Continuously variable adjustable transmission”, pp 61–67 4. Vyrabov RV, Mamaev AN, Balabina TA (2010) General questions of interaction between elastic wheels with a rigid support surface. In: Materials of international scientific and technical conference AAI “Car and tractor construction in Russia: development priorities and training”, dedicated to the 145th anniversary of MGTU “MAMI” 5. Balabina TA, Mamaev AN (2014) The mechanics of rolling elastic wheels on rigid support surface. In the book “Engineering science: trends, prospects and technology development” 6. Fromm H (1927) Berechnung des Schlupfes beim Rollen deformierbaren Scheiben. Zf Angew Math und Mech Bd 7, H.I 7. Fromm H (1928) Arbeitsverlust, Formänderungen und Schlupf beim Rollen von treibenden und gebremsten Rädern oder Scheiben. Beitrag zur Analyse der Reibungsgesetze. Z.f 8. Vyrabov RV, Mamaev AN (1983) The effect of toroidally elastic wheels for uneven wear across the width of the treadmill—Izv. Universities, Mechanical Engineering, pp 94–97 9. Vyrabov RV, Mamaev AN (1987) To the question on the least value of the coefficient of rolling resistance of elastic wheels on rigid horizontal surface. Izv. Universities, Mechanical Engineering, pp 85–88 10. Mamaev AN (1982) Determination of the coefficient of tangential elasticity of the wheels with the toroidal shape of the treadmill 11. Mamaev AN, Vukolova GS, Dmitrieva LN (1999) The effect of the view taken of the law of distribution of normal pressure in the wheels in contact with the rigid base on the calculated force and kinematic parameters of the wheel. Collection of scientific papers dedicated to the 60th anniversary recreation of MAMI. MSTU “MAMI”, Moscow 12. Vyrabov RV, Mamaev AN, Marinkin AP, Yuriev Yu M (1986) Influence of the mode of rolling elastic wheels on the magnitude of lateral forces during a side slip. J Mech Eng, 33–35 13. Vyrabov RV, Mamaev AN (1980) Analysis of power relations in the rolling of the slave elastic cylindrical wheel along a curvilinear trajectory. Mech Mach, 105–112 14. Vyrabov RV, Mamaev AN (1980) Investigation of contact phenomena in nonlinear toroidal rolling wheels. Izv. Universities, Mechanical Engineering, pp 33–38 15. Vyrabov RV, Mamaev AN (1980) Determination of forces and moments acting on the toroidal wheel in nonlinear rolling. Izv. Universities, Mechanical Engineering, pp 30–34 16. Mamaev AN, Sazanov IV, Nazarov Yu P (1990) Determination of the force characteristics of the elastic wheel rolling with a slip along a curved path. In: Proceedings of II all-union symposium “Problems of tires and rubber-cord materials. Strength and durability” 17. Mamaev AN (1999) The influence of the operating conditions of the tyres for wear their treadmill. Truck 18. Mamaev AN, Vyrabov RV, Portugalsky VM, Chepurnoy SI (2010) Determination of power and kinematic characteristics of an elastic wheel during rolling on a rigid drum. In: Proceedings of the international scientific and technical conference dedicated to the 145th anniversary of MSTU “MAMI” 19. Mamaev AN (2010) Driving wheel rolling resistance on a hard drum. In: Proceedings of the international scientific and technical conference, dedicated to the 145th anniversary of MSTU “MAMI” 20. Mamaev AN (1982) Influence of design parameters of elastic wheels on the value of their deflection and dimensions of the contact area with a rigid base. Interuniversity collection of scientific works “Safety and reliability of an automobile”, pp 203–211

540

M. Yu. Karelina et al.

21. Mamaev AN, Alepin EA (1980) Determining the size of the contact area and deflection of the wheel with a rubber tire with a static pressure wheel to a rigid base. Collection of scientific works “Machine Science”, pp 82–85 22. Mamaev AN (1982) On the determination of the hysteresis loss coefficient of highly elastic rolling elements 23. Balabina TA, Mamaev AN, Chepurnoy SI (2013) Determination of the ratio of camber and toe angles of the elastic wheels, ensuring the lowest rolling resistance. News of MSTU “MAMI”

Technical Audit of Rotary Aggregates S. G. Mogilny, A. A. Sholomitskii and A. L. Sotnikov

Abstract Rotary drum aggregates are often used in technological lines for continuous production. These complex structures are exposed to stresses and deformations under the influence of temperature and dynamical stress. To determine geometric parameters of such units, in this paper we propose to use a total station. The basic idea of the method is a remote real-time geodetic measurement of the machine parameters in the reflectorless mode. The task posed is solved by simultaneous measuring of the point spatial coordinates on the rotating machine and the timing of the measurement. As a part of work, a mathematical model was developed through which the geometrical parameters of the rotation body were determined and various random and systematic influencing factors are investigated. For the implementation of the proposed technology, the geodetic measuring complex “Vizir 3D” was updated accordingly. This paper shows an example of measurement of a drying unit. The results of vibration diagnostics of the support rollers of the drying drum allow determining the values of the parameters of vibration acceleration on the rollers. The peak factor and the vibration acceleration kurtosis of the support rollers should be less than 6 and 4, respectively.







Keywords Rotary aggregate Model Accuracy Alignment Online geodetic measurement





Remote measurement


S. G. Mogilny Prydniprovs’ka State Academy of Civil Engineering and Architecture, 24a, Chernyshevsky St., Dnepr 49600, Ukraine A. A. Sholomitskii (&) Siberian State University of Geosystems and Technologies, 10, Plakhotny St., Novosibirsk 630108, Russia e-mail: [email protected] A. L. Sotnikov Donetsk National Technical University, 58, Artema St., Donetsk 83001, Ukraine © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_57

541

542

S. G. Mogilny et al.

1 Introduction In mining, metallurgical and cement industries, many large-scale rotating continuously working objects are in operation: rotary kilns, mills, drum dryers, etc. The durability and reliability of such equipment depend on the accuracy of its alignment in the design position. Statistics of operation of such equipment shows that more than 30% of accidents and unscheduled downtime are associated with deviations from the straight axis of rotation of the kiln or geometric axis of the body [1–3]. With the careful alignment of rotating objects in the design position, there is a temporal increase of interrepair cycles. In addition, the power of the drive for the kiln rotation is reduced, and, consequently, electricity is saved. Therefore, improving the quality of geodetic service of rotating objects of continuous production and reducing the time of geodetic works are an important task. Similar tasks are inherent to other industrial machines and units [4–6], including monitoring the position of the working bodies of machines, balancing loads, and increasing the service life of bearing and other friction units [7–12].

2 Status of Issue Currently, to perform geodetic works related to the definition of geometric parameters of rotating objects, such as the axis of rotation of the object in the horizontal and vertical planes, radius of the object, etc., different techniques are used. The main techniques are the geometric leveling to determine the position of the axis of rotation in the vertical plane and the lateral leveling to determine the position of the axis in the horizontal plane [1, 2, 13]. These techniques give a low accuracy of the axis position determination since the radius of the device is found from indirect measurements of the circumference. Another significant disadvantage of these techniques is the need to stop and cool the unit to a temperature at which contact measurements can be carried out. Presently, electronic total stations [14, 15] are used to determine the geometric parameters of rotating kilns, but measurements are made only in static mode, on a cooled unit, and only in the area of the working platforms of the supports, which does not provide reliable information about the state of the shell of the unit. Another technique does not require stopping and cooling the kiln, and its essence is to measure the parameters of the kiln using a theodolite and a special device [16–18]. The kiln continues to rotate during the measurements but at a lower speed. With the help of the theodolite a special device, which tracks the contact of the axis of the sighting device with the side surface of the unit, is observed. This so-called tangential method is suitable for slowly rotating units, mainly for rotary kilns. Rotary kilns make a turn in 20–150 s; drying units of a higher rotation speed make a turn in 4–6 s.

Technical Audit of Rotary Aggregates

543

There is a technique based on high-frequency laser measurements [19, 20], the so-called EMD method based on the search for harmonics in the variation of distances. This technique also has a number of disadvantages. Therefore, determination of the geometric and kinematic parameters of rotary drying machines requires a fully remote technique of high-precision determination of parameters of rapidly rotating aggregates.

3 A New Method for Solving the Problem The proposed method lies in the field of engineering geodesy and with the help of electromagnetic emission provides geodetic control of industrial machines and their parts (rotating kilns, mills, drying machines, rollers, etc.) geometry in operation, i.e., without production delay [21]. The method is based on remote real-time geodetic measurements (without physical contact of the worker with the machine in operation) of the machine parameters. To solve this task, reflectorless total station measures spatial coordinates of the point on the rotating surface and fixes the measurement time moment that allows applying the following process method. In general case, rotating surface at initial moment t0 is expressed in the equation Uðr ; hÞ ¼ 0

ð1Þ

where r ¼ j x y z j—vector of spatial point coordinates on the machine surface; h—vector of the rotating surface parameters. Surface location by sighting laser beam r t ¼ j xt yt zt j—records vector of crossing point coordinates of sighting beam and rotating surface at some moment of current time. The measured coordinates of a surface point are recalculated to the coordinates it had at the initial time t0 using the following operator: rt0 ¼ AðuÞAðw; t  t0 ÞAT ðuÞðrt  re Þ þ re

ð2Þ

where AðuÞ—matrix of orthogonal rotation of the spatial coordinates to the axis of rotation of the surface, determined by the unit vector  u; w—angular speed of the surface rotation; re —vector of coordinates of the point located on the axis of rotation; Aðw; t  t0 Þ—matrix of orthogonal rotation of spatial coordinates around the axis of rotation for a period of time, which has the form cos wðt  t0 Þ  sin wðt  t0 Þ 0 ð3Þ Aðw; t  t0 Þ ¼ sin wðt  t0 Þ cos wðt  t0 Þ 0 0 0 1

544

S. G. Mogilny et al.

The vector r t0 satisfies Eq. (1), so the following system of equations takes place: Uðrt0 ; hÞ ¼ 0 rt0 ¼ AðuÞAðw; t  t0 ÞAT ðuÞðrt  re Þ þ re

 ð4Þ

Solving the system of Eq. (4) set up for a sufficient number of points measured on a rotating surface, its both geometrical and kinematic parameters are calculated. These measurements and the mathematical apparatus for their processing are necessary and sufficient to achieve the technical result—determination of the geometric and kinematic parameters of the operating units without stopping the production process, under operating loads and without the presence of a man in hazardous areas. The causal relationship of the features that make up the essence of the achieved technical result is explained as follows. Figure 1 shows a scheme of the section of the aggregate and relative position of the measuring device and the aggregate. In the process of measurements on the kiln surface, the coordinates of the points and the time of their measurement are recorded. The non-reflective electronic total station provides automatic recording of the direction, distance to the point, and the time of measurement of a sufficiently large number of points, which allows applying the proposed technique for processing the results. The kiln surface is a circular cylinder, so the system of Eq. (4) will have the form:

Fig. 1 Object measurement scheme

Technical Audit of Rotary Aggregates 2 ðr t0  r s ÞT ðr t0  r s Þ  ½vT ðr t0  r s Þ R2 ¼ 0 rt0 ¼ AðuÞAðw; t  t0 ÞAT ðuÞðrt  re Þ þ re

545

 ð5Þ

where r s —coordinate vector of the point located on the line of the symmetry axis of the cylinder; v—single vector directed along the axis of the cylinder; r e —vector of coordinates of the point e located on the line HH of the axis of rotation of the kiln; R—radius of the cylindrical surface of the kiln. The joint solution of the Eq. (5) set up for all measured points allows defining r s ; v; R—geometric parameters of the cylindrical surface, and also re ;  u; w—kinematic parameters of the kiln rotation.

4 Practical Application The parameters of the drying unit are measured in dynamics using the “Vizir 3D” software [22]. Measurements in dynamics are carried out according to the original technique proposed by the authors [23, 24] and successfully tested by the authors on slowly rotating kilns. • The total station is installed in front of one of the bandages of the working dryer, and the reference of the standing point to the basic geodetic network of the drying aggregate is performed. Usually, measurements were carried out on 5–6 reference points, which allowed determining the coordinates of the standing point with an accuracy of 0.1 mm. • Then, the measurements are performed in the characteristic cross sections of the object. In each Sect. 3, groups of measurements are performed in non-reflective mode. The measurement interval is selected in such a way that 30 are the points measured per revolution of the dryer. For bandages and rollers of 0.6 m width, three sections were measured, for the shell they measured 17 sections at the joints of the rings of the shell. As a result, 540 measurements are carried out on the surface of each bandage, and 270 measurements are uniformly located on the surface of the roller. In the cross section of the shell and in 17 sections, 90 and 1530 points were measured, respectively. All measured points are processed by the proposed technique. Figure 2a shows the deformation of the aggregate body. Deformations of the aggregate body lie in the range from −30 to +30 mm from the nominal size. The detail of the measurements is such that it is possible to determine the difference in the radii of the rings from which the body of the unit is welded. Figure 2b shows the projection of the bandage on the vertical plane. Measurement and analysis of vibration parameters of the support rollers of the drying drum are performed by using the vibration analyzer SD-21. During the studies, the envelope curve of the vibration acceleration spectrum, vibration parameters in the time and frequency domains (including wavelet analysis), as well

546

S. G. Mogilny et al.

Fig. 2 a Deformation of the aggregate body; b Projection of the bandage on the vertical plane

as the overall vibration level of the bearing assemblies of the support rollers of the drying drum were recorded. The form and parameters of vibration acceleration in the time domain turned out to be the most informative and illustrative for diagnosing bearing units of support rollers. The average value, peak factor, and kurtosis were chosen as the estimated parameters of vibration acceleration. The peak factor is the ratio of the peak value (maximum amplitude) to the mean square value (amplitude) of the vibration acceleration for a certain period of time (in the case under consideration—64 s.). The kurtosis determines the deviation of the parameters from the normal distribution characteristic of the serviceable condition of the diagnosed units and mechanisms. Registration of vibration acceleration in the time domain is carried out with a sampling frequency of 64 Hz, the number of counts assumed to be 4000. In all measuring points on the support rollers, the average value of vibration acceleration in the time domain was of the same order of magnitude; the peak factor and the kurtosis differed by an order of magnitude (1.5…4 times). On the support rollers from the driven side of the drum, the peak factor varied within the values of 4.72…6.26, and from the drive side—within the limits of 6.04…24.1. Kurtosis, respectively, varied within the limits of 3.3…3.4 and 4.39…12.9. On the roller supports from the driven side of the drum, there were no uncharacteristic periodic phenomena (regularities) in the time domain, in contrast to the roller supports from the drive side of the drum (Fig. 3). In all directions of measurement on the supporting rollers at the drive end of the drum, the characteristic (regular) series of the periodic shocks multiple of the frequency of rotation of the drum were observed. During the rotation of the drum, periodic metal shocks were perceived by the ear, which served as an additional sign of the presence of shock causes associated with malfunctions of the metal structure of the drum. It is found that the reason for these shocks was the weakening of the lining plates. The lining plates wearing out over time need replacement. Moreover, the presence of such shocks, directly unrelated to the state of the rolling bearings of the support rollers, can accelerate their wear and destruction, because the force of the shocks spreads to the support rollers and their bearing assemblies.

Technical Audit of Rotary Aggregates

547

Fig. 3 Acceleration measurement results: a Radial; b Vertical direction of the drum support roller

The results of the diagnosis were confirmed when replacing two bearings on the drive side of the drum on the opposite side of the engine.

5 Conclusion Industrial tests on the drying unit have shown that the technique allows effective determining the parameters and kinematic characteristics of not only slowly rotating kilns but also fast-rotating drying units and mills. According to the measurements, it is possible to determine with high accuracy the radii of the elements at the measurement points, the position of the axis of rotation, wear of bandages and rollers, and many other parameters; all measurements were performed on the unit in working condition. A comprehensive technical inspection of the rotating drying aggregate revealed the need to restore the cylindrical shape of the first bandage and rollers No. 1 and 2 since characteristic wear of the elements of the first support unit prevents the axial movement of the kiln and causes their wear out. Deformations of the unit body are in the range from −30 to +30 mm of the nominal size, and deviations of the unit rotation axis range from −17.7 to 19.0 mm in the horizontal plane and from −6.8 to 14.2 mm in the vertical plane. The results of vibration diagnostics of the support rollers of the drying drum allow determining the values of the parameters of vibration acceleration on the rollers from the driven side of the drum as criteria for the satisfactory condition of their rolling bearings. The peak factor and the vibration acceleration kurtosis of the support rollers should be less than 6 and 4, respectively.

548

S. G. Mogilny et al.

References 1. Kuz’o IV, Shevchenko TG (1987) Raschet i kontrol’ ustanovki agregatov nepreryvnogo proizvodstva (Calculation and control of installation of nonstop production aggregate). Vishha shkola, L’vov 2. Shevchenko TG, Hropot SG, Pivovarov VP, Ignatov AA, Men’shikov VF (1991) Rukovodstvo po vyverke tehnologicheskogo oborudovanija metallurgicheskoj promyshlennosti (The manual for adjustment of technological equipment for metallurgy). Ministerstvo metallurgii SSSR 3. Sotnikov AL, Kireev VN, Orobtsev AY et al (2013) Continuous casting of steel at bar-casting machines in Ukraine and Moldova. Steel Transl 43(12):808–814 4. Sotnikov AL, Kikin AB (2018) Theoretical foundations of the design of a lever mechanism of oscillations of the mold of a sorting machine for the continuous casting of billets. Metallurgist 61(11–12):1075–1083 5. Snitko SA, Yakovchenko AV, Sotnikov AL (2018) Influence of wheel billet stamping schemes on power modes of forming press operation and on wear of the deformation tool. Izv Vyssh Uchebn Zaved Chern Metall 61(5):385–392 6. Maruschak PO, Panin SV, Zakiev IM, Poltaranin MA, Sotnikov AL (2016) Scale levels of damage to the raceway of a spherical roller bearing. Eng Fail Anal 59:69–78 7. Brezinová J, Vińaš J, Maruschak PO, Guzanová A, Draganovská, D, Vrabel’ M (2017) Sustainable renovation within metallurgical production. RAM-Verlag 8. Sotnikov AL, Rodionov NA, Ptukha SV (2015) Analysis of mechanical loading of the hinges and supports of the mold vibration mechanism on a continuous caster. Metallurgist 58(9–10): 883–891 9. Sotnikov AL, Rodionov NA (2012) The control system of balancing of the CCM mold oscillation mechanism. Metall Min Ind 4(3):166–169 10. Vdovin KN, Tochilkin Vasilii V, Filatova OA, Tochilkin Victor V (2018) Improvement of the process of metal casting and the design of refractory equipment of the “Tundish Ladle—Open Jet—Mold” system of a continuous billet-casting machine. Refract Ind Ceram 59(3):237–240 11. Bardovskiy AD, Gorbatyuk SM, Keropyan AM, Bibikov PY (2018) Assessing parameters of the accelerator disk of a centrifugal mill taking into account features of particle motion on the disk surface. J Friction Wear 39(4):326–329 12. Keropyan A, Gerasimova A (2017) Connection of the temperature in contact area of the wheel-rail system with the railway slope of industrial railway transport Izvestia. Ferrous metallurgy. Nat Univ Sci Technol MISIS 60(5):355–363 13. Astashenkov GG (1986) Geodesicheskije raboty po ekspluatatsii krupnogabaritnogo promyshlennogo oborudovanija (Geodetic works on operation of large-size industrial equipment). Nedra, Moscow 14. Petrov VV, Tyurin SV (2005) Technology for controlling the geometric characteristics of rotary kilns. Cellyuloza. Bumaga. Karton (Cellulose. Paper. Cardboard)7:66–70 15. Petrov VV, Tyurin SV, Kopytov AN (2010) Control of geometric characteristics of rotating furnaces. Cem i ego primen (Cem Appl) 2:78–82 16. Krystowczyk B (1983) Ausrichten von Drehofen und Korrektur der TragrollenVerdrehundenwaehrend des Betriebes. ZementKalkGips Int 5:288–292 17. Josef By, Robertson L (1987) Kiln alignment method allows corrections while operating. Rock Prod, 21–22 18. Krystowczyk Z (2004) Geometry measurement of Killn shell in dynamic condition. Cem Build Mater 16:34–37 19. Zheng K, Zhang Y, Liu L, Zhao C (2017) An online straightness deviation measurement method of rotary Kiln cylinder. Tehnički vjesn 24(5):1297–1305 20. Zheng K, Zhang Y, Zhao C, Liu L (2015) Rotary Kiln cylinder deformation measurement and feature extraction based on EMD method. Eng Lett 23(4):283–291

Technical Audit of Rotary Aggregates

549

21. Mogilny SG, Sholomitskii AA (2018) Sposob vysokotochnogo opredeleniya parametrov poverhnosti i osei vrashchayushchihsya agregatov (Method of high-precision determination of parameters of the surface and the axes of rotating aggregates). Russian Federation Patent 2650418, 13 Apr 2018 22. Sotnikov AL, Sholomitskii AA (2017) Monitoring alignment of mold oscillatory motion with CCM process stream axis. Metallurgist 60(9–10):1046–1053 23. Mogilny SG, Sholomitskii AA, Seredovich VA, Seredovich AV, Ivanov AV (2015) The analysis of methods for determining the geometric parameters of rotating machines. In: 2nd international workshop on integration of point- and area-wise geodetic monitoring for structures and natural objects, Stuttgart, Germany, 23th–24th Mar 2015, pp 119–130 24. Mogilny SG, Sholomitskii AA, Martynov OV (2017) Real-time geodetic measurements of rotary machines. In: Proceedings of 17th international multidisciplinary scientific geo conference SGEM2017, Albena (Bulgaria), 29 June–5 July 2017, vol 17(22):523–530

Hydroblow as Mechanism of Additional Intensification of Liquid Forages Preparation in Centrifugal-Rotor Dismembrators Yu. N. Kamishov, N. A. Makarova and A. A. Sitnikov

Abstract In the article, the research of the working processes happening during the work of a centrifugal-rotor dismembrator is conducted. During the operation of the developed device, the phenomenon of hydroblow is observed. By means of numerical simulation, the mathematical model which is based on the system of the time-dependent equations of Navier–Stokes in their conservative form is obtained. The phenomena of turbulence and thermal conductivity were considered. Calculation was carried out dynamic as the difficult geometry of operative parts results in nonstationarity of liquid flow. The grain material was modeled by solid particle with a characteristic density, thermal conductivity and fracture strength. The diagram of pressure distribution in different time points after the beginning of a dismembrator’s operation at which zones with the increased pressure formed at the moment when the liquid flow comes across the moving cog of operative parts and discharging zones created behind a moving cog are clearly visible is received—the phenomenon of hydroblow under the influence of which the complete fracture of the grain does not happen is created, but the outer shell at pressure levels fractures, and it is higher than the fracture strength of the grain. It is shown that in the presence of areas with the pressure about 60 MPa (megapascal), the partial loss of the entirety of grains happens from which the conclusion is drawn that the phenomenon of hydroblow is the additional mechanism allowing to intensify the process of the grain fracture.














Keywords Dismembrator The rotor The stator Grain material Cutting elements Fracture process Pressure Numerical simulation Hydrodynamics Hydroblow Forages preparation
















Yu. N. Kamishov (&)
N. A. Makarova
A. A. Sitnikov Polzunov Altai State Technical University, 46, Lenin Ave, Barnaul 656038, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_58

551

552

Yu. N. Kamishov et al.

1 Introduction In the analysis [1–4] of construction arrangements of the machines suitable for preparation of liquid forages, it is shown that the best characteristics are demonstrated by a dismembrator [5]—a device of the rotor type in which one disk—the rotor—rotates, and the second—the stator—remains motionless. One of the main advantages of the machines of this type is a possibility of combining the operations of crushing and steaming [6, 7] that is an economically advantageous decision. Figure 1 presents the 3D model of dismembrator’s operative parts [8]. Crushing part, Fig. 1, consists of two stators: top 1 and lower 2, fastened among themselves by ten bolts 3. The lower stator fastens on the metal ring 4 fastening of a tank. In the lower stator, a rubber-metal sealing is attached 8. Also, the framing case 5 is attached by five secret bolts 9 to the lower stator and fastened to the electric motor. Into the flange of the crushing body, a bearing assembly 7, a shaft 14 of the drive of a rotor 6 and also 2 rubber-metal sealings 10 are installed. The rotor is fixed on a shaft by means of a bolt 11 which is in turn stopped by rubber plug 12 and a check washer 13. The principle of operation of the device consists in the following: grain mix with water gets to the crushing zone through inlet openings from above the zone on the rotating rotor disk. Under the influence of the centrifugal force arising due to the rotation of a rotor disk, the mixed mass directs to its periphery through built-up sections between the gear elements located on concentric circles of disks of a rotor and the stator. At the same time, gear elements mix and crush the grain mass. Having reached the periphery of the disks, the mix is removed from the device through the outlets located on the periphery of the top and lower stators.

Fig. 1 3D model of the operative parts of a dismembrator

Hydroblow as Mechanism of Additional Intensification …

553

In theory and practice of the grain material fracture in devices of centrifugal-rotor type, it is possible to allocate five main effects on the grain material at its crushing: shattering, attrition, splitting, flattening and cutting. The smallest energy consumption is required while crushing by cutting. Naumov [4] came to the conclusion that the same seed of wheat or rye depending on the nature of application of the operating force can prove both as fragile and as plastic. It is established by Demidov [1] that with the increase in humidity of the grain, there is a significant increase in power consumption (up to 30%) which needs to be spent on crushing. It is possible to explain it by the fact that the blow extends only in the upper layers, causing only flattening, unlike dry grain in which blow deformation goes deeper. In centrifugal-rotor dismembrator, crushing of the grain material happens in the liquid environment and as it was established in earlier conducted researches [9–11]: in devices of this kind besides the effect of cutting the processes of shattering, attrition, flattening and splitting are observed on the grain material; moreover, by these ways not all the grains are exposed, but only those which get to so-called stagnant zones of operative parts of a dismembrator [12]. For example, when the grain hits the gap between the rotating rotor disk and the framing case of the operative parts of a dismembrator, the grain is fractured by attrition [13]. In case of a blow of the grain material on walls of the crushing zone of the device, there is an effect of flattening. Besides the known ways of fracture in the dismembrator, an additional, insufficiently studied the effect of hydraulic impact on the grain material is observed. In [14], it is shown that the construction of a centrifugal-rotor dismembrator creates conditions for the hydroblow. A number of researchers indicate a negative effect from the impact of this phenomenon on operative parts of the installation. Others point out advantages in the operation of devices in which conditions for hydroblow are created. The purpose of the conducted research consisted in the numerical analysis of the hydroblow phenomenon’s effect on the working process of a centrifugal-rotor dismembrator for preparation of liquid forages.

2 Mathematical Model The phenomena of turbulence and thermal conductivity were considered. Calculation was carried out dynamically (i.e., the time factor was considered) as the difficult geometry of operative parts results in the non-stationary stream. Grain was modeled by solid particles from characteristic density, thermal conductivity and fracture strength. The system of the time-dependent equations of Navier–Stokes in their conservative form is the basis of the mathematical model. The instant equations of preservation of mass and impulse, known as Navier– Stokes equations, register as follows [15]:

554

Yu. N. Kamishov et al.

@q þ r  ðqUÞ ¼ 0 @t

ð1Þ

where q—liquid density, U—liquid speed vector in a point with coordinates (x, y, z) in timepoint t. Flow of impulse equation: @ðqUÞ þ r  ðqU  UÞ ¼ rp þ r  s þ Sm @t

ð2Þ

3 @ @ @ 6 @x ðqUx Ux Þ þ @y ðqUy Ux Þ þ @z ðqUz Ux Þ 7 7 6 7 6@ @ @ 7 6 r  ðqU  UÞ ¼ 6 ðqUx Uy Þ þ ðqUy Uy Þ þ ðqUz Uy Þ 7 7 6 @x @y @z 7 6 5 4 @ @ @ ðqUx Uz Þ þ ðqUy Uz Þ þ ðqUz Uz Þ @x @y @z

ð3Þ

where 2

- the tensor production of liquid flow density vector on a liquid speed vector, p— liquid pressure, s—a stress tensor, Sm—impulse source. The stress tensor is connected with the speed of the movement by the following ratio: 2 s ¼ lðrU þ ðrUÞT  dr  UÞ 3

ð4Þ

where l—molecular viscosity, T—liquid temperature, d—Kronecker delta. The condition of moving liquid is defined by five values: three components of speed and any two of its thermodynamic values, for example, pressure and density. Therefore, the full system of the equations of hydrodynamics has to contain five equations. In this case, the liquid is investigated in which there are processes of thermal conductivity and internal friction; therefore, the first four equations are the equation of continuity and movement of Navier–Stokes, and the fifth is the equation of transfer of heat which in a general view is registered as follows: @ðqhtot Þ @p  þ r  ðqUhtot Þ ¼ r  ðkrTÞ þ r  ðU  sÞ þ SE @t @t

ð5Þ

where htot—the full, k—thermal conductivity, SE—energy source. The system of the equations of Navier–Stokes is not closed; therefore, it has to be completed to form a closed system. The liquid stream in channels of a dismembrator represents a strongly turbulent flow; thus, no analytical solutions of the system of equations can be obtained. For the numerical solution, it is necessary to use a turbulence model.

Hydroblow as Mechanism of Additional Intensification …

555

Two additional transport equations are solved: for kinetic energy of turbulence and for dissipation speed: @ðqkÞ @ @ þ ðqUj kÞ ¼ @t @xj @xj @ðqxÞ @ @ þ ðqUj xÞ ¼ @ @xj @xj

   l @k lþ t þ Pk  b0 qkx rt @xj

 lþ

  lt @x x þ a Pk  bqx2 k rx @xj

ð6Þ ð7Þ

where b′ = 0.09, a = 5/9, b = 0.075, rx = 2, rt = 2. Pk illustrates a contribution to the turbulence of viscous forces, and it is set as follows:  Pk ¼ lt

   @Ui @Uj @Ui 2 @Uk @Uk þ  3lt þ qk @xj @xi @xj 3 @xk @xk

ð8Þ

The numerical solution of the model was calculated in the ANSYS CFX program.

3 Results of Numerical Modeling Numerical modeling gives an idea of the levels of pressure created in the operating time of a dismembrator. In Fig. 2, the chart of distribution of pressure in different timepoints after the beginning of work of installation is submitted. Modeling time is1 s. The nature of pressure change in the installation is caused by the hydroblow phenomenon arising when cogs of the rotor and the stator are overlapping during the operating time. Zones with increased pressure (green color on the chart) and discharging zones (blue color) are clearly visible on the graphic. Zones of increased pressure are formed at the moment when the liquid flow comes across the moving cog of operative parts. At this moment, the liquid behaves as compressed and there is a pressure jump up to 4–6 MPa. Conditions for hydroblow are created.

Fig. 2 Planimetric charts of liquid pressure in a dismembrator

556

Yu. N. Kamishov et al.

Discharging zones are the zones which are created behind the moving cog (negative pressure about 3 MPa). It is possible to assume that the phenomenon of hydroblow is favorable in case of the grain grind as it allows to soften an outer rigid shell of a grain. To test this assumption, the impact on an individual seed of barley by the pressure created in the installation was modeled. The result is presented in Fig. 3. As it is seen in the figure, the final fracture of a grain does not happen, but the outer shell fractures at pressure levels are higher than the fracture strength of a grain. Thus, pressure about 12 MPa that is the main level of pressure created by installation (turquoise color at Fig. 2) does not make an effect on a grain, and it keeps its integrity. With the pressure over 30 MPa, cracks begin to appear on the cover of a grain, and with the pressure about 60 MPa, a grain fractures. Such levels of pressure are characteristic of areas of a dismembrator in which hydroblow appears (green color at Fig. 2). As a result, the hydroblow created by installation represents the auxiliary mechanism for shattering of the grain material. It is significant to estimate the contribution of this phenomenon to the crushing process. The dependence of the cutting force necessary for grains crushing of different types on the angle of sharpening of the cutting sides and humidity of the grain material is experimentally researched [16, 17]. For barley seeds, the range of cutting force changes from 21 to 59 N depending on the parameter’s values [18–20].

Fig. 3 Modeling of the pressure effect on a grain arising in the grain mix

Hydroblow as Mechanism of Additional Intensification …

557

Without the additional researches directed to the clarification of the surface area of grain adjoining to the cutting tool and water at hydroblow, it is possible to draw only a qualitative conclusion on benefits of the phenomenon of hydroblow while crushing the grain material. Indeed, to achieve the pressure allowing to overcome the grain’s fracture strength, the cutting tool has to affect a small segment of the surface (about 1 mm2). Preliminary effect of hydroblow reduces the necessary cutting force that creates an opportunity for a process intensification by an increase in surface area of the cutting edge or increase in efficiency of the installation by reduction of power consumption [21].

4 Conclusion • In a centrifugal-rotor dismembrator, except the known ways of fracture, the additional effect of the hydraulic impact on the grain material is observed. • The mathematical model allowing to investigate the processes proceeding in installation by methods of computational hydrodynamics is made. • On the basis of a mathematical model, the working process of a dismembrator is modeled. According to the results, it is possible to draw the conclusion that the average pressure created by installation is about 12 MPa that is not enough for the grain fracture, but in the areas subjected to the hydroblow, the level of pressure is about 60 MPa that leads to the disintegration of the outer shell of a grain without an impact on it of the cutting elements.

References 1. Demidov VA (1969) Impact grinding machines. Moscow 2. Fedorenko IYa, Smyshlyaev AA (2015) Modeling of shock loading of base material layer. Bull Altai State Agric Univ 5(127):136–141 3. Fedorenko IYa, Zolotarev SV, Smyshlyaev AA (2002) Experimental background to the development of the shock-centrifugal grinder of the forming grain. Bull Altai State Agric Univ 2(6):84–88 4. Naumov IA (1975) Improving the conditioning and grinding of wheat and rye. Kolos, Moscow 5. Kamishov YN (2012) Justification of working bodies’ design parameters of dismembrators for production of liquid feed mixture. Dissertation, Polzunov Altai State Technical University 6. Brooks PH, Beal JD, Niven S (2001) Liquid feeding of pigs: potential for reducing environmental impact and for improving productivity and food safety. In: Corbett JL (ed) Recent advances in animal nutrition in Australia, Armidale 7. Hong TTT, Lindberg JE (2007) Effect of cooking and fermentation of a pig diet on gut environment and digestibility in growing pigs. LiveStock Sci 109:135–137. https://doi.org/10. 1016/j.livsci.2007.01.121 8. Nefjodov KE et al (2014) Rotary pulsation device. RF Patent 2535165, 8 oct 2014

558

Yu. N. Kamishov et al.

9. Sabiev UK, Pushkarev AS (2018) Grain mill. Selskiy Mechanizator 3:22–23 10. Sabiev UK, Pushkarev AS, Sabiev IY (2017) Effective grinders of fodder grain. In: Abstracts of the national scientific practical conference “Scientific and technical support of agro industrial complex condition and development prospects”. Omsk State Agrarian University named after P.A. Stolypin, Omsk, 6 Apr 2017 11. Sabiev UK, Demchuk EV, Myalo VV, Soyunov AS (2017) Innovative equipment and production method for mixed fodder in the conditions of agricultural enterprises. IOP Conf Ser: Mater Sci Eng 221:012020. https://doi.org/10.1088/1755-1315/221/1/012020 12. Sitnikov AA et al (2014) Mechanical activator for organic materials. RF Patent 87851, 16 Jan 2014 13. Sabiev UK, Pushkarev AS (2017) The performance of the upgraded chopper. Bull Omsk SAU 4(28):245–248 14. Sitnikov AA, Makarova NA, Kamyshov YN (2016) A method of parameter optimization of the equipment configuration with the help of neural net and finite element analysis. IOP Conf Ser: Mater Sci Eng 124:012074. https://doi.org/10.1088/1757-899x/124/1/012074 15. ANSYS (2006) ANSYS CFX-solver theory guide (Release 11.0). ANSYS, Inc., Canonsburg 16. Syrovatka VI, Sergeev NS (2008) The study of the dynamic cutting of rape seeds and feed grain. Vestnik MGAU 1:54–59 17. Syrovatka VI (1976) Scientific and technical foundations and methods of technological calculation of production lines for the preparation of animal feed on collective and state farms. Dissertation, South-Ural State Agro University 18. Zolotarev AM, Trufanov VV, Druzhinin RA, Yarovoy MN (2018) On the rationale of operating parameters of an impact centrifugal crusher. Vestn VSAU 1(56):119–127. https:// doi.org/10.17238/issn2071-2243.2018.1.119 19. Zolotarev AM, Trufanov VV, Druzhinin RA, Pushkarskaya AE (2017) Results of experimental researches of the shock-centrifugal milling device. In: Gulevskij VA (ed) Modern scientific and practical solutions in the agro-industrial complex, Voronezh 20. Zolotarev SV (1985) Mechanical and technological foundations of the creation of percussion-centrifugal grinders of feed grain. Dissertation, South-Ural State Agro University 21. Fedorenko IYa, Smyshlyaev AA, Levin AM (2002) Energetic environmental enhancements with a removal measuring circuit. Mechanization Electrification Agric 11:31–32

Change of Structure of Dispersion Material Under Dynamic Loads E. K. Chabutkin

Abstract During the compaction of various road-building materials, the efficiency of the process largely depends on the correct choice of the roller and the purpose of its operating modes. To increase the efficiency of the vibrating roller, it is required to adjust the magnitude of the surface stresses under the roller with the growth of density of the material. The variety of road construction materials with a wide range of physical and mechanical properties makes it difficult to solve the problem. Any road construction material can be represented by a three-phase system with a skeleton, air, and water (or bitumen for bitumen–concrete mixes). The physical– mechanical processes of dispersion material structure formation are studied in this paper. Process of deformation of the dispersion material is carried out with the convergence and penetration of the medium- and large-sized blocks into each other. The analysis of the dependence of changes in the structure of different materials during deformation showed that there are three stages of formation of the structure of the material with different energy costs at each stage of compaction. The paper also shows the similarity of these processes for various materials and shows the connection between the structure formation and the compaction rate. On the basis of obtained results, the recommendations for improving the effectiveness of compaction using vibratory rollers.




Keywords Roller Compaction Load Dynamic effect





Dispersion material
Material structure


1 Introduction A large variety of materials with a wide range of physical and mechanical properties are used during the construction of the roadway as a complex engineering structure. This explains the variety of compaction equipment. E. K. Chabutkin (&) Yaroslavl State Technical University, 88, Moskovsky prospect, Yaroslavl 150023, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_59

559

560

E. K. Chabutkin

The compaction is a very complex process in itself with many interrelated factors affecting the end result in one way or another. The task of choosing a combination of parameters of a compaction machine, which maximizes the efficiency of the compaction process for a specific area with defined conditions, is as important as it is difficult. This problem is often solved by simulating the compaction process [1–4]. However, it is impossible to implement without taking into account the processes occurring in the compacted material itself. In road construction, the dispersed multiphase materials—such as soils, bitumen-concrete mixes—are widely used. One of the features of these materials is the presence of different phases, as well as the dispersion, fragmentation of the solid phase, which has a developed specific surface as a result. Coarse materials like roadstone and gravel are also used, while their properties are determined by the solid phase only. However, in the construction of road embankments with very coarse soils, the formation of dense structure is usually performed under abundant moisture. The presence of sand–clay filler within their structure, considering the moisture, also brings them closer to the back-fill soils. This diversity of materials makes it possible to identify specific structure types. The material structure depends on the nature of interacting forces between its components, as well as on the location and orientation of the particles. These interacting forces are called bonds. There are the following types of road construction materials structures: coagulation, crystallized, and contact structures. The analysis of the processes occurring in dispersion materials under compaction shows that most materials have a coagulation structure. Therefore, from the point of view of physical and mechanical properties, any road construction material can be represented by a three-phase system with a skeleton, air, and water (or bitumen for bitumen–concrete mixes) [5]. Depending on the placement of water in the pores, such materials may have either “entrapped water” or “entrapped air” (Fig. 1). The influence of the cohesive forces on the strength of three-phase materials varies and depends on dispersion, the shape of particles, and moisture of the soil. Deformations of the three-phase soil are different and depend on the ratio of water and air in its pores. Fig. 1 Back-fill soil structure diagram. 1—mineral particles; 2—tightly bound water; 3—loosely bound film water; 4—free state water; 5—“entrapped air”; 6—capillary water

Change of Structure of Dispersion Material Under Dynamic Loads

561

A layer of water is formed on the surface of the particles. This layer, called tightly bound water, has anomalous properties. This water has considerable viscosity, elasticity, and shear strength [6]. There is loosely bound water on top of the tightly bound water, and it is kept near the surface of solid particles by lesser forces. The properties of this water are close to the properties of water in a free state. In addition to this, it moves much slower than the free water, and its speed depends on the temperature. The amount of bound water in soils depends on their mineralogical composition. In simplified form, the composition of a three-phase dispersion material can be represented by a diagram shown in Fig. 1.

2 Study Results The detection of changes in the material structure that happen at different stages of deformation is of particular interest. To study this, the author conducted mechanical experiments of samples of soil and asphaltic concrete according to [7, 8] with simultaneous study of their structure. The structure of the material was studied using the MIN-7 (Russian: MИH-7) microscope, and during the study, the percentage of grains of different diameters was counted within the examined area. The results of the study are shown in Fig. 2. The material structure in the initial state was of block-oriented nature. About 50% of the material consisted of microblocks. The space between the blocks was randomly filled with oriented mass of clay particles and fine sand fractions with clearly visible voids and cavities formed due to the loose fitting of the blocks. During the deformation, the changes occur, which are the reduction of the number and size of cavities and pores, which sometimes turned out to be compressed and elongated in the direction of the shift. This indicates a local displacement of the particles and microaggregates, and their more compact packing. The longer the process of deformation is, the greater was the reduction of cavities. Analysis of dependencies of the dispersion material structure changes during deformation shows that, with compaction, an intensive growth of large blocks occurs (Fig. 2a, d) by reducing the distance between the grains, the mutual displacement of small particles in the slip areas, which adhere tightly to each other and fill all possible voids, and form a single large block. The grains of medium size grow along with the large grains (Fig. 2b, e). But the growth of these blocks goes up to a certain density estimated by a compaction rate Ky for any dispersion material in the range from 0.92 to 0.94. It happens because, at this density, the number and size of cavities are reduced to critical, the movement of small particles almost stops, and the intensive convergence of medium-sized blocks, as well as the formation of the large blocks begins. It is apparent that the three-phase dispersion material maintains the “entrapped water” state up to the density corresponding to Ky = 0.94. With this, the air phase can exit the compacted material freely.

562

E. K. Chabutkin

Fig. 2 Change in the average grain diameter to the compaction rate dependency graphs. a–c for soil; d–f for asphalt

The intensive growth of small grains (Fig. 2c, f) is only typical for the initial stage of deformation before reaching Ky in the range from 0.85 to 0.88 by having the small particles move in the voids present in large numbers. Gradually, the number of voids and slip areas decreases considerably, while the intensive decrease of growth of small grains, as well as their transition into medium-sized blocks due to their convergence, occurs. The density increases to Ky = 0.97 and the three-phase dispersion material changes its state to “entrapped air.” The further process of deformation of the dispersion material is carried on with the convergence and penetration of the medium- and large-sized blocks into each other. With this, the significant stresses at the points of contact of the blocks cause the appearance of small cracks and fractures with the formation of small particles. Under rapid loading, the reorientation of particles cannot occur fast enough and the fractures occur mostly at the points of contact of microblocks.

Change of Structure of Dispersion Material Under Dynamic Loads

563

In the process of formation of the dense structure of dispersion material, the pores become closed off, the capillary water from the underlying layers cannot move through these pores anymore, and there comes no possibility for the movement of water that falls on the surface of the compacted material in the form of precipitation. The formation of additional small cracks with the compaction above Ky = 0.98…0.99 can lead to additional deformation of the layer of compacted material in the fall–winter and winter–spring periods. Therefore, it is required to be very cautious about assigning operation modes for the compaction equipment, especially on soil roadbeds. It is not that critical for asphaltic concrete roads, as the bitumen component of the material has the ability to “heal” small cracks that appear when it is over-compacted. An indirect confirmation of the factor of occurrence of small cracks during over-compaction gives for a dependence of the amount of frost heaving on the compaction rate [9], shown in Fig. 3. Thus, for different dispersion materials, three stages of structure formation under dynamic effects can be clearly seen. The first stage is distinguished by small energy costs associated with sliding of the individual particles of the material along the layers of loosely bound film water. The second stage is the formation of small blocks and transition to the medium-sized blocks, and it obviously requires higher energy costs. At this stage, there is the movement of individual blocks along the edges of their points of contact already. There comes the displacement of loosely bound water and “entrapped air” with the introduction of films of strongly bound water inside each other, as well as the formation of the mineral blocks, surrounded by a newly formed film of strongly bound water.

Fig. 3 Dependence of the amount of frost heaving of fine sandy loam on material density

564

E. K. Chabutkin

The third stage is the most energy-intensive, which is associated with the rupture of the films of strongly bound water and the penetration of the individual blocks into each other. The three-stage process of dispersion material structure formation is closely linked to the impact of the compaction machine on the material. Obviously, the impact of the machine must also come in stages, divided into three—with the transition from a lighter machine to a heavier one (applies to the static effect machines), or with the use of mechanisms that allow for regulation of the dynamic effect during the compaction process [9–12]. However, by using the appropriate machinery, the first compaction stage can be excluded, therefore using only two-stage compaction of the road construction materials, which provides a significant reduction of time spent for the process of formation of the finished surface. For example, by using a paver with a vibrating smoothing plate and a tamper bar, it is possible to achieve the compaction rate of 0.89–0.90 after the vibratory machine [13, 14], which already corresponds to the second stage of the material structure formation. There is also no need to conduct the first passes without vibration, and it is possible to immediately turn on the vibration mode of the compaction machine with access to a lower driving force of the vibrator [11, 15, 16].

3 Conclusion • Close in quantitative terms results for different dispersion materials show the identity of the processes occurring in them. Therefore, the models simulating the process of deformation of different dispersion materials may be the same, considering the parameters of the compacted material. This conclusion is supported by the works of other authors [9, 17]. • During the deformation of dispersion material, three stages of structure changes can be seen: the convergence of disparate grains of material and the formation of blocks of small grains (corresponds to Ky = 0.85–0.88); the convergence of small blocks and their transition into the medium-sized blocks (not higher than Ky = 0.97); the convergence and the penetration of medium and large blocks into each other (Ky higher than 0.97). • The formation process for a dense and strong structure of the compacted material must be three-staged according to the regulation of the dynamic effect considering the achieved compaction rates at each stage. For example, without turning the vibration on until achieving Ky = 0.88, with the reduced driving force until achieving Ky = 0.96, and the maximum driving force of the vibrator at the last stage of compaction.

Change of Structure of Dispersion Material Under Dynamic Loads

565

References 1. Chabutkin EK, Tyuremnov IS, Popov Yu G (2012) Methods of calculation of the operation mode of the vibrating roller during the compaction of hot bitumen-concrete mixes ю Vestn Komp’iuternykh I Informatsionnykh Tekhnol 5(95):19–24 2. Tyuremnov IS, Filatov IS, Morev AS (2012) Review of the rheological models of interaction of a vibrating roller and compacted soil. Math Math Educ. Theory Pract Interuniversity Collect 8:234–241 3. Filatov IS, Morev AS, Tyuremnov IS (2014) To the question of justification of the values of the rigidity and viscosity rates with the rheological simulation of the compaction process. In: 67 Conference Yaroslavl technical state university. 23 Apr 2014, Yaroslavl. P. 1, YSTU Publishing House, p 429 4. Guillo E, Gautier M, Boyer F (1999) Dynamic modelling and simulation of compactor. In: 14th IFAC wold congress, Beijing 5. Bulychyov VG (1974) Mechanics of disperse soils. Stroyizdat, Moscow 6. Babkov VF, Bezruk VM (1986) Fundamentals of soil science and soil mechanics. Higher School, Moscow 7. GOST 12801-98 Materials on the basis of organic binders for road and airfield construction. Experimentation methods. Gosstroy Russia, GUP CPP. Moscow, 28 p 8. GOST 9128-2009. Asphaltic concrete mixtures for roads and aerodromes and asphaltic concrete. Standartinform, Moscow, 20 p 9. Kharkhuta NYa, Vasilyev Yu M (1075) Strength, stability and compaction of soil subgrade of roads. Transport, Moscow 10. Gross AR, Petukhov VN, Chabutkin EK, Prusov AYu (2007) Coordination of work of several rollers during compaction of various road construction materials. Roads and bridges. Article compilation/FGUP ROSDORNII 17(1):82–89 11. Chabutkin EK, Tyuremnov IS, Popov YuG (2011) Increasing the efficiency of using the vibrating rollers by optimizing their operation modes. Stroitelnye I Dorozhnye Mach 8:19–24 12. Chabutkin EK (2018) Effective use of rollers. Mir Dorog 108:82–84 13. Shestopalov KK (2008) Construction and road machines. “Akademiya” Publ. Center, Moscow 14. Podolskiy VP (2012) Technology and organization of road construction. Road surfaces: the textbook for stud. Institutions of higher professional education. “Akademiya” Publ. Center, Moscow 15. Popov Yu G, Tarasova NE, Chabutkin EK (2011) The control of the driving force of a vibration drum during the compaction of asphaltic concrete surfaces. International workshop “Design improvement problems of construction, road, municipal and airport machines”: workshop materials. MADI, Moscow, pp 80–83 16. Shumakov DO, Chabutkin EK (2010) Effective use of vibrating rollers. In: 63 conference Yaroslavl technical state university. Yaroslavl, YSTU Publishing House 17. Shirkovich SV (1962) Compaction machines in building and construction works. Kuibyshev

Increasing Efficiency of Vibratory Rollers Through Adjusting Magnitude of Disturbing Force Y. G. Popov and E. K. Chabutkin

Abstract While compacting various road construction materials, the efficiency of the process depends largely on the correct choice of the roller both in terms of weight and design features and the purpose of its operating modes. To improve the efficiency of the vibrating roller, it is necessary to adjust the amount of contact stresses under the roll as the density of the material increases. This significantly reduces the total number of roller passes required to achieve the required compaction factor (ratio). At the same time, the wrong selection of the roller operating mode leads to an increase in compaction time, a decrease in the efficiency and quality of compaction of road construction materials. The paper gives recommendations to adjust the magnitude of the disturbing force of the vibration exciter in the process of the roller operation. At the same time, it is possible to choose the modes of their operation, in particular, the rational speed of rolling.







Keywords Roller Compaction Vibration exciter speed Compaction ratio Operating mode








Disturbing force
Rational

1 Introduction Vibrating rollers are widely used in the world of construction practice. Due to a wide range of dimension-type row, most of the works on the compaction of the road beds and other engineering structures are carried out with the help of vibrating rollers [1]. All leading manufacturers of compaction machinery are engaged in the development of systems that can improve the performance of rollers. At the same time, despite rather high level and energy consumption of the equipment used, there is still a problem to choose and implement a rational operating mode of a vibrating roller. Moreover, as of today, with the advent of a large number of rollers on the Y. G. Popov (&)
E. K. Chabutkin Yaroslavl State Technical University, 88, Moskovsky prospect, Yaroslavl 150023, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_60

567

568

Y. G. Popov and E. K. Chabutkin

market, capable to vary compacting parameters in the course of operation, this problem is particularly urgent. Compaction, in itself, is a very complex process, with many interrelated factors that somehow influence the final result. At the same time, the task to choose such a combination of parameters of the compacting machine, under which the process of compacting a particular section under certain conditions will be most effective, is as important as it is complex. The work of all compacting machines is associated with the application of repeated cyclic loads to the surface of the material being compacted. In this case, the deformation of the material depends on the magnitude of the load, determined by the contact pressures r0, on the rate of change of the stress state and on the duration of the load, and, consequently, on the repetition number of its application. The formation of the dense and strong structure of dispersed material can be carried out in two ways: “from the top to the bottom” and “from the bottom up” (Fig. 1). In the first case, when compacting “from top to bottom,” contact pressures r0 should not exceed the limits of the strength of the material [rp]. Otherwise, the material will be squeezed out from under the working part of the machine. As a result, the upper part of the compacted layer will be loosened. With the second “bottom-up” compacting scheme, the contact pressures r0 exceed the material strength limits [rp]. At the same time, a dense core is created inside the material layer, from which, with repeated application of the load, a dense and strong structure begins to form with increasing density from the core to the surface. The use of such a method is useful when you compact materials with an already well-established structure, for example, natural soils or soils with layers of large thickness with a heterogeneous structure (inclusions of construction debris, filled soils in spring with a large number of frozen clods, etc.). In the list of the roller's technological parameters, which affect its compacting capacity, usually stand out:

Fig. 1 Diagram of the formation of a packed core in the body of a dispersed material

Increasing Efficiency of Vibratory Rollers Through …

569

Mass of the roller. It is believed that a large mass makes it possible to compact the material to a greater depth; therefore, heavy compactors are used to compact complex soils (for example, clay soils). At the same time, the influence of the mass on the compaction process cannot be considered in isolation from the parameters of the working elements, such as the width and diameter of the roll, and its type (smooth, pneumatic, or cam type). Therefore, the parameters, that depend on the mass, in particular, the static linear pressure q, are often used as technical characteristics of the roller. Moreover, the ratio q/R is more often used, where R is the radius of the roll. It is also worth adding that the static linear pressure, in addition to dependence on the weight per roll and type of working body, can also vary within some limits with increasing material density. Vibration parameters. This group includes a number of characteristics that have a significant impact on the compacting ability of the roller. First, it is vibration frequency and disturbing force. For soils and macro-fragmental materials, a low-frequency mode with large amplitude is usually used. Asphalt concrete, due to its viscosity, requires a high-frequency mode of oscillation. The same group of parameters includes the direction of the oscillations, which can be circular, directed [2], and rotational (oscillation vibrations) [3]. At present, there are machines that make it possible to regulate vibration parameters within a wide range, including the frequency, disturbing force, and direction of the oscillations [4–6]. In most modern rollers, the main working organ is a vibrating roll. Its main constructive element is a vibration exciter, which directly creates additional dynamic loads on the material to be compacted. Recently, the development of rollers has been actively pursued, in which there is a possibility of smooth regulation of the magnitude of the disturbing force of the vibration exciter [7, 8]; therefore, it is of practical interest to study the possibility of adjusting the contact stresses under the roll of the roller by changing the magnitude of the disturbing force of the vibration exciter. As a result of the compaction, not only the required density of the material along the whole thickness of the layer must be obtained, but also its strong structure must be formed. For this, the contact pressure should be close to the ultimate strength of the material, but must not exceed it not only at the end of compacting, but also during the entire process. However, it should be taken into account that as the density increases, the strength limit also increases. Consequently, when regulating contact pressures, this factor should be taken into consideration first. Thus, the gradual increase in contact pressures is the main compaction rule, ensuring the formation of a dense and strong structure. Increase in contact pressures can be achieved in two ways: • using a set of rollers moving from lighter roller to a heavier one; • using machines that allow you to adjust the value of the contact pressures due to the activation of vibration (vibrating and combination rollers) or due to the regulation of air pressure in tires (pneumatic roller).

570

Y. G. Popov and E. K. Chabutkin

It is also possible to combine these two methods in combined type rollers. It is quite clear that the second method is more preferable, since it allows to minimize the amount of the equipment used.

2 Methods This paper presents the results of a computational experiment, which was carried out by the method described in work [9] and using the software complex [10]. The reliability of the data obtained was verified by comparison with the results of studies by both Russian scientists [11, 12] and by foreign researchers [13, 14]. In small volumes, the results were checked under laboratory conditions [9]. At the same time, changes in the structure of the material occurring at different stages of deformation were detected. Keeping this purpose in mind, the authors carried out mechanical tests of soil and asphalt concrete samples with simultaneous study of their structure.

3 Results and Discussion According to the existing technology of road-building materials compaction, the roller must first perform 2–3 passes with the vibration drive turned off, so that the contact stresses do not exceed the strength limit of the material, and then, as the compaction factor increases and the material strength increases, the contact stresses under the roll of the roller increase due to turning on the vibration exciter. However, this is not always the case. Figure 2a shows a graph of the achieved compaction factor of the number of passes of the CA 250 roller of Dynapac firm (working weight is of 10.6 tons; the disturbing force is of 246 kN at a vibration frequency of 33 Hz). Compacting material—sand, a layer thickness is 0.75 m, moisture content is 0.95 from the optimum, the initial compaction factor is 0.8. The working speed of the roller is 2 km/h. Curve 1 shows the mode under which the vibration is turned on after the third pass. As can be seen from the figure, the dependence on the first three passes is practically direct. After the vibration is turned on, the density (3–6 passes) of the material is intensified, after which the curve assumes a more sloping shape, which is explained by the natural decrease in the soil’s ability to deform due to the increase in the density of the material. However, under given production conditions, the characteristics of the selected roller allow the vibration to be switched on already during the first pass. At the same time, from the first pass, the condition for switching by the disturbing force is met, and the stresses under the roll of the roller are in the acceptable limits, which allows to reduce the total number of passes required to achieve the compaction factor Ku = 0.98 up to two (Fig. 2a curve 2). The intense increase of the density

Increasing Efficiency of Vibratory Rollers Through …

571

Fig. 2 a Compacting of sand with CA 250 roller 1—mode of operation when vibration is activated after 3 passes 2—mode of operation when vibration is activated from the first pass. b Compacting with CA 402D roller 1—when vibration is off by 3 first passes 2—mode of operation when vibration is activated from the first pass

starts from the first pass and is reduced during the 4th pass. Thus, switching on the vibration on the first pass can reduce the total time of work by 20% for this machine. At the same time, when using a heavier machine with greater disturbing force with the same parameters of the compacted material, activation of the vibration at the early stages of compaction can lead to the so-called de-compaction effect [15]. This is clearly seen on similar graphs (Fig. 2b) made for the CA 402D roller from Dynapac (working weight is 13.8 tons, disturbing force is 300 kN at a vibration frequency of 33 Hz). Here, curve 1 shows the process of increasing density with the activation of vibration after the first three passes (curve 1) and curve 2 shows the process of increasing density with the activation of vibration during the first pass (curve 2). The rapid activation of vibration leads to an increase in the total number of passes, which is a consequence of the considerable excess of the contact stresses under the roll of the roller with respect to the tensile strength of the material at the first stage of compaction, with the required density of the material structure being formed only at the later stages. Thus, when assigning the operating modes of even one roller, it is necessary to select a rational moment for activating the vibration exciter, taking into account the conditions of production, compaction parameters, and technical characteristics of the machine. The task is considerably complicated if we take into account the velocity factor, which also can vary widely at different stages of compaction. Analysis of the given dependences shows that the change in the disturbing force is effective only when it increases in full accordance with the change in soil density and does not exceed its strength limit. It is possible to reduce the required number of passes by 30–40%.

572

Y. G. Popov and E. K. Chabutkin

However, it is still not clear how the disturbing force should be regulated. Analysis of the data obtained [9] shows that in most cases, a smooth change in the magnitude of the disturbing force, in addition to the increased energy consumption, does not lead to the desired effect. The computational experiment performed showed that it is expedient to change the magnitude of the driving force stepwise, activating at the initial moment 20 … 25% of the maximum disturbing force and only at the final stage, 100% of the disturbing force is activated. Moreover, the moment of activation of the magnitude of the disturbing force and the moment of its adjustment must be correlated with the current compaction factor. At the same time, with the standard thickness of the soil layer being compacted for different machine parameters, the moments of activation are described by identical curves (Fig. 3). The influence of the soil type on the nature of these curves is also not significant. The data obtained are in good agreement with the physical processes that take place in the formation of dense and strong structure. At the compaction stage, most materials have a coagulation structure. Hence, from the point of view of physical and mechanical properties, any road-building material can be represented as a three-phase system consisting of a skeleton, air, and water (or bitumen for asphalt–concrete mixtures) [16]. Depending on the placement of water in the voids, such materials can be in two states either with “jammed water” or with “jammed air.” Analysis of the dependence of the structure change of a dispersed material during deformation shows that during compaction, intensive growth of large blocks takes place due to a reduction in the distance between grains, mutual movement of small particles along slip zones, which, when they are close to each other, fill all possible voids and form a single, much larger block. The intensive growth of small grains is characteristic only at the initial stage of deformation to reach Ku in the range from 0.85 to 0.88 due to the movement of small particles through voids, being in large quantities. The number of voids and slip zones is gradually reduced, while there is an intense decline in the growth of Fig. 3 Dependence of the compaction factor on the machine parameters. Lower curve—with the activation of vibration after the third pass to 100%. Upper curve—with the activation of vibration after 5 passes to 25%, and after 7 passes to 100%

Increasing Efficiency of Vibratory Rollers Through …

573

small grains and, due to the convergence their movement into blocks of medium size. However, the growth of these blocks goes to some certain density, estimated by the compaction factor of Ku, for any dispersed material in the range from 0.92 to 0.94 [9]. This is due to the fact that with this density, the number and size of voids are reduced to be critical, the movement of small particles almost stops, and intense convergence of medium-sized blocks to form large blocks begins. This increases the energy intensity of the deformation process, and the stress under the roll of the roller must be increased. It is obvious that up to the density corresponding to Ku = 0.94, the three-phase dispersed material is in the “clamped water” state. In this case, the air phase has the possibility of free exit from the compacted material. Greater energy consumption is required at the last stage of compaction, at the moment when the formation of large blocks stops and their penetration into each other begins. When large blocks are formed, the density increases to Ku = 0.97 and the three-phase dispersed material passes into the “clamped air” state. Consequently, in terms of the changes of dynamics in the compacted material structure, a stepwise scheme of changing the stresses under the roll of the roller, and hence the moments of activation of the disturbing force of the vibration exciter suggest themselves. The speed of the roller has the most powerful influence on the moments of activation of the disturbing force. Moreover, the lower limit (the limit of activation is 25% of the maximum disturbing force) for different parameters of the roller, for different materials, and different thicknesses of the deposited layers is described practically by one curve (Fig. 4a—Curve a). It is even more complicated with the limit of activation of 100% of the maximum disturbing force—curve b. And the location of this limit depends both on the parameters of the machine and the type of the material to be compacted. It is impossible to define it for specific conditions and it can be obtained only by calculation.

Fig. 4 a Dependence of the resulting compaction factor on the ratio q/R for different materials. b Dependence of the achieved compaction factor on the speed

574

Y. G. Popov and E. K. Chabutkin

In addition, it should be taken into account the fact that not at all speeds of the machine it is possible to achieve the required density—curve B. And at certain speeds and at specific parameters of the machine, a step control of the magnitude of the disturbing force is not required. For example, Fig. 4a, at speeds of the machine movement below 2.5 km/h on the loamy soils, when you reached the compression factor being equal to 0.94, it is necessary to activate 100% of the maximum value of the disturbing force. For sandy loam (Fig. 4b) for the roller with a ratio of parameters q/R = 15  102 kg/m2, at the same time (at speeds below 2.0 km/h) it is enough to activate only 25% of the maximum value of the disturbing force to achieve Ku = 0.97 … 0.98, and the activation of 100% of the disturbing force is not required. The activation of 100% of the maximum disturbing force at any stage of compaction will not result in increased productivity—reducing the required number of passes until the required compaction factor is achieved. On the contrary, it can lead to the destruction of the upper layer of the strip being compacted, which will adversely affect the quality of the compaction. Thus, the computational experiment showed that the activation of 25% of the magnitude of the disturbing force from the maximum should be satisfied when the density meets the compaction factor being equal to 0.92 … 0.94 of the required value, and the activation of 100% of the magnitude of the disturbing force—upon reaching Ku = 0.97. However, the question arises as to how to determine the necessary moments for the activation of vibration. The use of modern density control systems on rollers [17–19] will give the opportunity to solve this problem.

4 Conclusion • In most cases, a smooth change in the magnitude of the disturbing force does not lead to optimization of the compaction process, but results in the increased energy consumption. • The use of a step-by-step regulation of the magnitude of the disturbing force makes it possible to simplify the construction of the vibrating rolls of the roller and reduce the energy consumption for driving the unbalanced shaft, and, consequently, the fuel consumption by the compacting machine in the whole. • The activation of 25% of the magnitude of the disturbing force from the maximum should be performed when density reaches the compaction factor of 0.92 … 0.94 of the required one, and the activation of 100% of the magnitude of the disturbing force upon should be performed when reaching Ku = 0.97. • The moment the vibrator is switched on is highly dependent on the operating speeds of the machine. The higher the working speed of the machine, the earlier it is necessary to activate the vibration exciter, setting 25% of the magnitude of

Increasing Efficiency of Vibratory Rollers Through …

575

the driving force. It is advisable to activate the vibration exciter at full power only at the last stage of compaction while simultaneously setting the minimum working speed of the roller. • The use of modern density control systems will allow not only to monitor the compaction process, but also to control the process of activation of the disturbing force at the right time, adjusting the value of contact stresses in full accordance with the increase in the tensile strength of the compacted material [20].

References 1. Chabutkin EK, Tarasova NE (2017) Recommendations on the choice of road rollers and the purpose of modes of operation under the conditions of construction. Roads Bridges 37(1):69– 80 2. Forssblad L, Cessler S (1977) Vibratory asphalt compaction. “Dynapac”, 89 p 3. Wehrli C, Anderegg R (1998) Nonlinear oscillations at compacting machines. Geotechnik, Verlag Glückauf GmbH, Essen:16–25 4. Cat Compaction in Europe (2005) Contract J 3:48–54 5. Hay J (1980) Vibroller with rubber covered roller. Softholed roller plays down posses. Contract J 5261(23):296 p 6. Briaud J-L, Seo J (2003) Intelligent compaction: overview and research needs. Texas A&M University, 79 p 7. Noshe K (2002) Development of a new type of single drum vibratory roller. In: Proceedings of 14th international conference of the international society for terrain-vehicle systems, Vicksburg, MS USA, 20–24 Oct 2002 8. Zhang J, Xue GH, Liu D (2011) Development of the new vibration exciter. Adv Mater Res 106:211–212 9. Chabutkin EK (2017) Application of vibrating rollers for compaction of road construction materials. Publishing House YAGTU, Yaroslavl, p 200 10. Chabutkin EK, Tyuremnov IS, Popov YG, Prusov AY (2011) The software package “Vibkat”. Patent software product 2011615193, opub. 01.07.2011 11. Kostelyov MP (2004) Possibilities and efficiency of vibrating rollers for compaction of soils of various types and conditions. Road Mach 2:22–24 12. Kostelyov MP (2008) Again about the quality and efficiency of compaction of various soils with modern vibro-rollers. Road Mach 1:40–47 13. Anderegg R, Kaufmann K (2004) Intelligent compaction with vibratory rollers. Transp Res Board, 124–134 14. Serridge CJ, Synac O (2007) Ground improvement solutions for motorway widening schemes and new highway embankment construction over soft ground. Ground Improv J 11(4):219– 228 15. Harhuta NI (1973) Machines for soil compaction. Moskov-Leningrad. Mechanical Engineering, 176 p 16. Bulychev VG (1974) Mechanics of dispersed soils. Stroyizdat, Moscow, p 227 17. Floss R, Gruber N, Obermayer JA (1983) Dynamical test method for continuous compaction control. In: Rathmayer HG, Saari K (eds) Proceedings 8th European conference on soil mechanics and foundation engineering May, Helsinki, pp 25–30 18. Kopf F, Erdmann P (2005) Numerische Untersuchunsen der Flachendecker Dynamischer Verdichtungskontrolle. Osterreichische Ingenieus-und Architekten-Zeitschrift (OIAZ) 150(4– 5):126–143

576

Y. G. Popov and E. K. Chabutkin

19. Mooney MA, Rinehart RV (2009) In-Situ soil response to vibratory loading and its relationship to roller-measured soil stiffness. J Geotech Geoenviron Eng ASCE 135(8):1022– 1031 20. Tyuremnov IS (2016) The review of the continuous compaction control system for soil compaction by vibratory rollers. Part 3. Functional features and “Intelligent Compaction”. Bull Pac Nat Univ 2(41):115–122

Energy Component of Properties of Material Crushability Layer Yu. A. Lagunova and V. S. Bochkov

Abstract In modern conditions, the effectiveness of traditional methods of ore preparation is reduced, due to a number of circumstances, namely high costs per unit of final products, low productivity of grinding equipment, gigantism of structures, high consumption of metal for lining material and grinding media, losses of recoverable products due to their diffusion destroying their bodies, and uncontrollability of the process of destruction of rock particles. At present, crushers are being created and tested that allow combining the functions of crushing and grinding. However, the widespread use of such crushers is limited by the relatively high energy intensity of the process implemented in them. At the Department of Mining Machines and complexes of the Ural State Mining University, experimental and theoretical studies of the processes of fragmentation are carried out. Particular attention is paid to the study of the properties of the crushability of rocks and the formation of energy consumption during crushing.










Keywords Crushability Crushers Elastic deformation Plastic deformation Energy consumption Crushing degree Layer deformation










1 Introduction The crushing process is characterized by a large number of factors determining the technological parameters of the process [1, 2]. The main influencing factors are the physicomechanical properties of rocks [3], in particular, the crushability of the rock. Completeness of the use of the property of crushability is one of the components of the evaluation of the effectiveness of process equipment. Yu. A. Lagunova (&) Ural Federal University named after the first President of Russia B.N. Yeltsin, 19, Mira street, Ekaterinburg 620002, Russia e-mail: [email protected] V. S. Bochkov Ural State Mining University, 30, Kuibyshev street, Ekaterinburg 620144, Russia © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_61

577

578

Yu. A. Lagunova and V. S. Bochkov

Improvement of the processes of disintegration of durable materials is associated with the development of the theory of energy-efficient destruction of materials [4, 5], based on the softening of the interphase boundaries [6] during the deformation of pieces (particles) of the material in the bulk layer. In modeling the process of destruction, a layer of material is understood to mean a mixture of pieces of various sizes in the form of a layer of limited height enclosed (locked) in a matrix [7]. When the layer is compressed, the material is compacted, the contacts of the grains are tinted and the weakest, flapped and needled pieces are broken, which leads to an improvement in the shape of the crushing product. Crushing in the way “in a layer” (or “piece by piece”) provides a more uniform grain size of the product and an isometric grain compared with crushing according to the “piece of armor” method of due to the implementation of the principle of material destruction in cramped conditions [8].

2 Practical Application of Crushing “in a Layer” To date, Russian and foreign manufacturers have developed several types of crushers in which the principle of crushing in the layer is implemented [9]: • NPK “Mekhanobr-Technika” [10–12] created a cone without eccentric inertial crusher KID (Fig. 1), on which with the help of an adjustable unbalance vibration exciter a vibration method of destruction of materials in the layer is realized, which allows to significantly reduce the energy consumption for disintegration. When laboratory testing, cone inertial crusher KID 300 noted that the degree of fragmentation depends on the frequency of force exposure. When the frequency of force impact on the rock is 200 times/min (which corresponds to the frequency of oscillations of traditional cone crushers), the crushing ratio did not exceed 3–4 units, and at the force impact frequency of 1500 times/min, the crushing ratio reached 40 units. KID crushers are used at the final stage of crushing and are currently being successfully operated by many brick factories in Karelia, Voronezh, Orenburg, Moscow, and other regions. • PJSC Uralmash produces cone crushers KMD-1750T7 and KMD-2200T7 [11, 12], the design of the working chambers of which are designed to provide crushing in cramped conditions, i.e., when the rock pieces crush themselves at the moment of compression. In order to ensure such crushing, the chamber must be filled with a material; i.e., the crusher is constantly working “under the rubble.” Sensors are used to control the loading mode, and installation of small storage hoppers directly above the crusher inlet is desirable. • The principle of crushing “in a layer” has been used in cone crushers of leading world manufacturers for more than one year [12]. These crushers are manufactured by Sandvik Rock Processing (Hydrocone Series Crushers) and Metso Minerals (HP, MP Series Crushers).

Energy Component of Properties of Material Crushability Layer

579

Also, designs and prototypes of high-frequency vibratory machines with external vibrators (crushers of jaw and cone and impact cone mill) have been developed, the use of which can significantly reduce the energy consumption for the disintegration of minerals (4–5 times) [13, 14].

3 The Disadvantages of the Method The main disadvantage of crushing “in a layer” is the significant energy intensity of this method, which limits the scope of implementation of the method to rocks of medium strength. The increase in the energy intensity of crushing is due to the periodic action of the external load. During indentation deformation, elastic and plastic deformations occur simultaneously, and the role of elastic ones is to prepare the process for the final plastic deformations. With a periodic load in the interval between the loading of the layer, elastic deformations are restored and plastic propagation ceases; therefore, to create a stress state of a certain size in any microvolume of the material, it is necessary to apply many efforts for a periodic load.

4 Factors Affecting the Crushability of the Material The main factors characterizing the strength properties of the layer of material are the following: • the structure of the layer, determined by the location of the fractions of the material in the volume (height) of the layer; • the degree of compaction of the layer, characterized by the coefficient of loosening of the material; • the height of the layer of material; • boundary conditions determined by the degree of blocking of the layer. Indicators of the elastic-plastic properties of the material layer are the bulk and effective compressibility moduli of the material layer, which characterize the relationship between the pressure P in the layer and the relative deformation of the layer e. An expression for determining the effective modulus of compressibility of a layer of material was obtained in [4]. EEF ¼ EH ðhÞ
q=ð1  ðe=ePR ÞÞk

ð1Þ

580

Yu. A. Lagunova and V. S. Bochkov

Fig. 1 Design of crusher KID

Table 1 Values of the effective modulus of compressibility of the layer from the relative deformation

e/ePR

k=0

k = 0.25

k = 0.5

k = 0.75

k=1

0 0.25 0.50 0.75 1.0

1 1 1 1 1

1 1.075 1.19 1.41 ∞

1 1.15 1.41 2.0 ∞

1 1.24 1.68 2.83 ∞

1 1.33 2.0 4.0 ∞

where EH ðhÞ—is the bulk modulus of compressibility of the layer of material as a function of the height of the layer of material; q—coefficient characterizing the degree of pre-compaction of the layer; ePR —the limit value of the deformation (deformation of pressing); k—is an exponent depending on the structure and degree of blocking of the layer. The indicator k [15] characterizes the elastoplastic properties of the layer and varies in the range 0  k  1. When k ¼ 0, the function is kðhÞ ¼ 1 and EEF ¼ EH ; that is, the material layer is an ideal granular medium and does not resist deformation when the side surfaces are free. For k = 1, the function has the form kðhÞ ¼ q=ð1  ðe=ePR ÞÞ. In this case, the layer of material is placed in the matrix, and the dependence of the effective compressibility modulus on the relative deformation is almost linear and corresponds to the crush force diagram of a single piece. Figure 2 and Table 1 show the graphs of relative values (with respect to the bulk modulus of compressibility of the layer) of the effective compressibility modulus EEF =EH for various values of k.

Energy Component of Properties of Material Crushability Layer

581

5 Indicators of the Destruction Process The main parameter is the pressure exerted on the layer by compressive force. p ¼ EH
e
ðq=ð1  e=ePR ÞÞk

ð2Þ

The total resistance of the layer of material to compression and, accordingly, compressive load F ¼ p
S ¼ EH
e
S
ðq=ð1  ðe=ePR ÞÞÞk

ð3Þ

where S—is the cross-sectional area of the layer perpendicular to the force vector. Will get a ¼ p
DV=V ¼ p
e

ð4Þ

where a—is the energy intensity of the destruction of the material in the layer. Taking into account, the expression (2) a ¼ EH
e2
q=ð1  ðe=ePR ÞÞk :

ð5Þ

Energy consumption for the deformation and destruction of the layer of material ADEF ¼ a
V ¼ V
EH
e2
q=ð1  ðe=ePR ÞÞk

ð6Þ

Specific energy consumption for layer destruction ASP ¼ A=V ¼ 0:5
F
Dh=ðS
H Þ ¼ 0:5
p
e:

Fig. 2 Dependence of the effective modulus of compressibility of the layer on the relative deformation

ð7Þ

582

Yu. A. Lagunova and V. S. Bochkov

Table 2 Values of specific energy consumption from the relative deformation of the layer

e/ePR 0

k пpи q = 1 0 0

0.25 0.50 0.75 e/ePR 0

0.031 0.034 0.125 0.17 ∞ ∞ k пpи q = 2 0 0

0.25 0.50 0.75

0.062 0.25 ∞

0.068 0.34 ∞

0

0

0

0.038 0.22 ∞

0.042 0.30 ∞

0.046 0.38 ∞

0

0

0

0.076 0.44 ∞

0.084 0.60 ∞

0.092 0.76 ∞

Substituting (2) into expression (7), we get ASP ¼ 0:5
e2
EH
q=ð1  ðe=ePR ÞÞk :

ð8Þ

Figure 3 and Table 2 show the graphs of the dependence of the specific energy consumption on the destruction of the layer on the relative deformation of the layer. Graphs can be approximated by two lines—a straight line (for values of 0\e\eK ) and a curve in the form of an exponent (eK \e\ePR ). The dependences of the specific energy consumption on the destruction of the layer on the relative deformation (for q ¼ 1) are: with k ¼ 1 ASP ¼ 0:5
e2
EH
q=ð1  ðe=ePR ÞÞk

ð9Þ

ASP ¼ 0:5
e2
EH
q=ð1  ðe=ePR ÞÞ0:5 ;

ð10Þ

at k = 0.5

Fig. 3 Dependences of specific energy consumption on the relative deformation of the layer

Energy Component of Properties of Material Crushability Layer

583

at k = 0.25 ASP ¼ 0:5
e2
EH
q=ð1  ðe=ePR ÞÞ0:25 :

ð11Þ

The particle size characteristic or the degree of crushing of the material is determined by the volume content of the solid phase or the coefficient of loosening. The residual value of the coefficient of loosening corresponds to the achieved degree of crushing of the material: kLO ¼ kLH  ðKLH  kL:PR Þ
e=ePR ;

ð12Þ

where kL:PR —is the value of the material loosening coefficient corresponding to the deformation of pressing. When k ¼ 1:
kLO ¼ kLH  ðKLH  kL:PR Þ
1  0:5e2 EH =ASP :

ð13Þ

6 Conclusion Thus, the obtained model gives a complete solution to the problem of the destructive compression of a layer. At the same time, the granulometric characteristics of the layer as a function of the main parameter characterizing the state of the layer, namely the strain energy density A [15], are determined simultaneously with pressure and strains. Reducing the strength of a layer of material can be achieved through rational organization of the workflow, consisting in two stages—compaction of the layer of material and the destruction of the grains while realizing the maximum deformation of the layer.

References 1. Johansson M, Quist J, Evertsson M et al (2017) Cone crusher performance evaluation using DEM simulations and laboratory experiments for model validation. Miner Eng 103–104:93– 101. https://doi.org/10.1016/j.mineng.2016.09.015 2. Lindqvist M, Evertsson CM (2006) Development of wear model for cone crushers. Wear 261 (3–4):435–442. https://doi.org/10.1016/j.wear.2005.12.010 3. Evertsson CM (2000) Cone crusher performance. Chalmers University of Technology, Göteborg 4. Revnitsev VI, Zarogatsky LP et al (1988) Selektivnoye razrusheniye mineralov (Selective destruction of minerals). Nedra, Moscow 5. Tang CA, Xu XH, Kou SQ, Lindqvist P-A, Liu HY (2001) Numerical investigation of particle breakage as applied to mechanical crushing—part I: single-particle breakage. Int J Rock Mech Min Sci 38(8):1147–1162. https://doi.org/10.1016/S1365-1609(01)00075-2

584

Yu. A. Lagunova and V. S. Bochkov

6. Gazaleeva GI (2014) The mechanism of destruction of rocks in the process of crushing material “in the layer”. News of higher educational institutions. Min J 5:95–100 7. Boyko GKh et al (2003) Gornoye oborudovaniye Uralmashzavoda (Mining equipment Uralmashzavod). Ural’skiy rabochiy, Yekaterinburg 8. Fedotov PK, Pykhalov AA (2012) Chislennoye modelirovaniye protsessa drobleniya porody v sloye mezhdu prokatnymi valkami pod davleniyem (Numerical simulation of rock crushing in the layer between rolling rolls under pressure). J Sovremennyye tekhnol. Sistemnyy analiz. Modelirovaniye 3(35):21–26 9. Bochkov VS, Lagunova YA (2018) Obzor i perspektivy razvitiya konusnykh drobilok, razrushayushchikh gornuyu porodu “v sloye” (Overview and prospects of development of cone crushers, destroying the rock “in the layer”). Sbornik trudov XVI Mezhdunarodnoy nauchno-tekhnicheskoy konferentsii v ramkakh Ural’skoy goronopromyshlennoy dekady, Yekaterinburg, pp 143–147 10. Weisberg LA, Shuloyakov AD, Orlov SL et al (2010) Novyye tekhnologii proizvodstva vysokokachestvennogo kubovidnogo shchebnya melkikh fraktsiy (New technologies for the production of high-quality cubical crushed stone of small fractions). J “Gornaya Promyshlennost” 3(91):10–13 11. Perelygin V (2007) Polucheniye kubovidnogo shchebnya na otechestvennom oborudovanii (Obtaining cubic crushed stone on domestic equipment) J Osnovnyye sredstva. https://os1.ru/ article/6606-poluchenie-kubovidnogo-shchebnya-na-otechestvennom-oborudovanii. Accessed 25 Nov 2018 12. Obzor rynka drobil’no-sortirovochnogo oborudovaniya (mobil’nogo i statsionarnogo) v Rossii i na Ukraine (Overview of the market of crushing and screening equipment (mobile and stationary) in Russia and Ukraine) (2012) InfoMayn (issledovatel’skaya gruppa). Moscow. http://www.infomine.ru/files/catalog/404/file_404.pdf. Accessed 25 Nov 2018 13. Sapozhnikov AI, Golikov VM, Repin SV (2016) Razvitiye tekhnologii izmel’cheniya stroitel’nykh materialov (The development of the technology of grinding building materials). J Stroitel’nyye i dorozhnyye mashiny 4:27–31 14. Feliks J, Mazur M (2017) Study on vibratory crushing and granulation of limestone. Int Multi Sci GeoConf Surv Geol Ecol Manage SGEM 17(11):1081–1088. https://doi.org/10.5593/ sgem2017/11/S04.138 15. Lagunova YA (2009) Razrabotka nauchno-tekhnicheskikh osnov povysheniya effektivnosti razrusheniya gornykh porod “v sloye” (Development of scientific and technical foundations of increasing the efficiency of rock destruction “in the layer”). Dissertation, Ural State Mining University, Ekaterinburg

New Criterion for Continuous Compaction Control Systems by Soil Vibratory Rollers I. S. Tyuremnov and A. S. Morev

Abstract This chapter substantiates the need to improve methods for monitoring the quality of compaction of road-building materials and the use of continuous compaction control systems in the composition of vibratory rollers to ensure documentation and quality control of 100% of the compaction surface. The criteria of soil compaction are considered, based on which the work of modern systems for the continuous compaction control by Ammann/Case, Bomag, Caterpillar, Dynapac, Sakai and Volvo rollers is based. The importance of considering the different modes of operation of the vibration roller: constant contact, partial separation and double jump and others, as well as indications of the transition of the drum to the undesired mode of oscillations double jump, is indicated. The characteristics of compaction indicators are analyzed, the calculation of which is based on the analysis of the acceleration spectrum of the vibration roller, as well as indicators considering the strength characteristics of the material to be compacted. The necessity of creating a domestic continuous compaction control system by vibratory rollers has been substantiated. A new promising compaction criterion for such a system is presented. The use of the proposed CV index makes it possible to fine-tune the continuous compaction control systems of the vibratory roller to a specific type of material. It also allows improving the sensitivity of the continuous control system to changes in the characteristics of the soil being compacted.







Keywords Soil Compaction Vibration paction control Soil compaction criterion





Vibratory roller
Continuous com-

I. S. Tyuremnov
A. S. Morev (&) Yaroslavl State Technical University, 88, Moskovsky prospect, Yaroslavl 150023, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_62

585

586

I. S. Tyuremnov and A. S. Morev

1 Introduction The modern road construction sees the use of hi-tech vibrating compaction equipment. At the same time, soil compaction quality control is also a big deal. When the soil and stone materials are not enough compacted, it becomes one of the possible reasons of uneven and prematurely worn-down roads. Durability and stability of layers of roads can be assured by choosing proper road construction materials, as well as maintaining standard (or better) quality of compaction of their layers. Therefore, compaction is one of the key operations in road construction. Currently in Russia, soil compaction quality assurance during road construction is conducted in accordance with Russian construction regulations called SP 78.13330.2012 “Rules and regulations. Roads.” At the same time, the assurance is selective, checking only about 0.02% of the total compacted area [1]. The operators of vibrating soil rollers made in Russia have no information about the condition of the compacted material during work execution (soil compaction), which, in turn, does not allow them to detect and fix undercompacted areas in time and provide long-term durability and evenness of roads. The most important question is how to control the quality of stone materials since the regulations in the Russian Federation lack objective criteria that can be used to evaluate the quality of their compaction. This task can be solved with systems that are installed on a vibratory roller and measure the change in the condition of compacted materials in real time. Continuous measurement of soil performance during compaction with a vibratory roller is usually called the continuous compaction control. Top foreign manufacturers of vibratory rollers such as AMMANN, BOMAG, CATERPILLAR, DYNAPAC, and SAKAI have been including the systems that allow to continuously monitor the state of compacted soil during work execution and document the result of compaction in the basic package of their machines for a long time (Table 1). There are no such systems manufactured in Russia. Not only does it not allow Russian manufacturers to compete in the market of compaction equipment adequately, but also it is one of the factors leading to low quality and life of Russian roads. Table 1 Example of a table Manufacturer

System name

Control value

AMMANN/CASE BOMAG

ACE (Ammann Compaction Expert) Terrameter VarioControl AccuGrade Compaction DCA (Dynapac Compaction Analyzer) HCQ (Hamm Compaction Quality) Sakai CompAnalyzer, CompGauge

ks Omega Evib CMV, RMV, MDP CMV, RMV HMV (same as CMV) CCV CMV, RMV

CATERPILLAR DYNAPAC HAMM SAKAI VOLVO

New Criterion for Continuous Compaction C…

587

2 Methods A new criterion of soil compaction has been developed based on the analysis of the principles of the functioning of continuous control compaction systems of the world’s leading manufacturers of vibration engineering and measuring equipment. The definition is based on the analysis of the acceleration spectra of the vibratory roller drum with consideration of the empirical weight coefficients calculated using statistical regression methods.

3 Results and Discussion When compacting soil with vibratory rollers, the properties of compacted soil (its type, density, and humidity), as well as the operating mode of the roller, can fluctuate within large margins (vibration frequency, the driving force of a vibration exciter, the nature of vibrations, the movement speed of the vibratory roller). The operation of continuous compaction control system is affected by the uneven distribution of parameters of compacted soil and the presence of transient processes when changing the operating mode of the vibration exciter of the roller. Researchers and manufacturers of vibratory rollers distinguish the following vibration modes of the vibratory drum of a vibratory roller during soil compaction [2]: continuous contact, partial uplift, double jump, rocking motion, and chaotic motion. When operating the vibratory roller in constant contact with compacted soil, the vibrations of the vibratory drum and soil are sinusoidal with equal amplitude of adjacent vibrations of the operating unit (vibratory drum) [3]. The mode of partial uplift of the vibratory drum from compacted soil is the primary operating mode of a vibratory roller. In each vibration cycle, the operating body lifts off the soil, but the sinusoidal nature of vibrations and the equality of amplitudes of adjacent vibrations of the drum are maintained. During the uplift of the vibratory drum, the soil is restored with decreasing speed. When switching from the partial uplift mode to the double jump mode, the nature of vibrations of the vibratory drum becomes subharmonic. Amplitudes of adjacent vibrations of the vibratory drum are not the same, which leads to deterioration of control over the vibratory roller, increase in load on the bearing units of the vibratory drum and on the joint of the drum and the frame, and it also negatively impacts the efficiency of compaction. Therefore, it is important to monitor the transition into an unwanted mode of operation of the vibratory drum (double jump mode). By changing the operating mode of the vibratory roller, the nature of vibrations also changes and, consequently, the spectrum of vertical accelerations of the vibratory drum. With the increase in the soil compaction degree the amplitude of vibrations of the vibratory drum also increases, and the additional harmonics and

588

I. S. Tyuremnov and A. S. Morev

subharmonics appear in the acceleration spectrum as well. In 1975, Dr. Heinz Thurner together with Oke Sandström founded Geodynamik to continue the development of continuous compaction control systems installed on the vibrating soil rollers (as well as asphalt rollers in the future). In cooperation with Dr. Lars Forssblad from DYNAPAC, Geodynamik has developed and introduced the Compaction Meter Value (henceforth CMV) in 1978 [4, 5]. The calculation of CMV is based on the analysis of the vertical acceleration spectrum (accelerometer data) of vibratory roller drum over two cycles of vibration. CMV is calculated as [4, 6]: CMV ¼ C
A2f =Af ;

ð1Þ

where Af—amplitude of the harmonic of the vertical acceleration spectrum of the vibratory drum on the frequency of changes in the driving force f (nominal frequency of vibrations of the vibratory roller drum) (Fig. 1), m/s2; A2f—the amplitude of the harmonic of the vertical acceleration spectrum of the vibratory drum on the frequency 2f (second-harmonic amplitude of the vertical acceleration spectrum of the vibratory roller drum) (Fig. 1), m/s2; and C—calibration rate (according to the source [6] is usually equal 300). The analysis of the vertical acceleration spectrum of the vibratory roller drum is also the basis of calculation of Resonance Meter Value (henceforth RMV), also known as Bouncing Value (henceforth BV) [4]: RMV ¼ BV ¼ C
A0:5f =Af ;

ð2Þ

where A0.5f—the amplitude of a subharmonic of the vertical acceleration spectrum of the vibratory roller drum on frequency of 0.5f (Fig. 1), m/s2. The control value RMV allows to indicate the transition of the vibratory roller drum from the partial uplift into the unwanted double jump.

Fig. 1 Change of the vertical acceleration spectrum of the vibratory roller drum during compaction [8–10]

New Criterion for Continuous Compaction C…

589

Continuous compaction control system, which functions based on calculations of CMV and RMV, is used in the vibratory rollers made by CATERPILLAR, DYNAPAC and VOLVO. North American company Trimble specializing in the production of various land surveying equipment offers a continuous control system called CCSFlex, which is also based on CMV [7]. Oscillator rollers HAMM use the continuous compaction control systems which functionality is based on calculations of Oscillo Meter Value (henceforth OMV), which is equivalent to the control value CMV. To calculate OMV, the accelerations of an axis of the vibratory drum in the horizontal plane are analyzed [6]. Moreover, the calculation algorithm OMV varies depending on the slippage between the vibratory drum and soil [8]. In 2004, the Japanese manufacturer of construction equipment SAKAI proposed its own control value for continuous control systems—Compaction Control Value (henceforth CCV) [9, 10]. This control value is used nowadays in continuous compaction control systems of vibratory rollers made by SAKAI. The calculation of CCV is also based on the analysis of the vertical acceleration spectrum of vibratory roller drum. When calculating CCV in addition to amplitudes of harmonics of the vertical acceleration spectrum of the vibratory drum with frequencies of fluctuation of the driving force f and 2f (Fig. 1), the amplitude of harmonic and the vertical acceleration spectrum with the frequency of 3f, as well as subharmonic with the frequencies of 0.5f; 1.5f, and 2.5f, are also considered [9, 10]:
 CCV ¼ A0:5f þ A1:5f þ A2f þ A2:5f þ A3f =A0:5f þ Af
100%;

ð3Þ

where A1.5f—the amplitude of subharmonic of the vertical acceleration spectrum of the vibratory drum with the frequency of 1.5f, m/s2; A2.5f—the amplitude of subharmonic of the vertical acceleration spectrum of the vibratory drum with the frequency of 2.5f, m/s2; A3f—the amplitude of subharmonic of the vertical acceleration spectrum of the vibratory drum with the frequency of 3f, m/s2. Currently, the continuous compaction control systems, which are based on the calculations of CMV and RMV, are some of the most common systems. However, the systems based on the calculations of CMV and RMV are defined not by the durability of soil, but the indirect criteria based on the analysis of the vertical acceleration spectrum of the vibratory drum in the roller. The main purpose of soil compaction is to ensure the strength of soil, i.e., the ability to withstand the dynamic and static loads from transport, the layers of soil above and the pavement. Therefore, the leading manufacturers of vibratory rollers have decided not to limit themselves with the compaction criteria CMV and RMV and have been developing alternative approaches to implement the continuous compaction control [11]. Control value Machine Drive Power (henceforth MDP), which is used in the continuous compaction control systems of the vibratory rollers made by CATERPILLAR together with the compaction criteria CMV and RMV [7], is one of the alternatives to the criteria which are based on the analysis of the acceleration spectrum of the vibratory drum. The use of MDP for continuous compaction control

590

I. S. Tyuremnov and A. S. Morev

is based on reducing the depth of immersion of the vibratory roller drum in soil and reducing the resistance to movement of the roller, while the density of soil increases during the compaction by the vibratory roller [12]: MDP ¼ Pg  W
V
ðsin h þ a=gÞ  ðb1
V þ b2 Þ;

ð4Þ

where Pg—power required to move the vibratory roller (W); W—the weight of the vibratory roller (N); V—the movement speed of the vibratory roller (m/s); Ѳ—the longitude inclination to overcome (rad); a—the acceleration of movement of the vibratory roller (m/s2); g—free fall acceleration (m/s2); b1, b2—internal loss rates (each roller has a different rate); (b1 V + b2)—the element that shows the power consumption associated with internal losses; W V (sinh + a/g)—the element that shows the power consumption associated with the longitude inclination. In 1982, a well-known German manufacturer of compaction equipment BOMAG suggested a control value Omega to ensure that the continuous compaction control system Terrameter could function. The control value Omega is the energy transferred into soil during the operation of a vibratory roller on a compacted area. The reaction force of soil Fs that occurs during the deformation of soil changes during the compaction depending on the immersion depth of the vibratory roller drum in soil (vertical movement of the vibratory roller drum zd), the operating mode and the soil parameters: its type, density and wetness. The accelerometer mounted on the frame of the vibratory roller drum constantly records the acceleration of the vibratory roller drum during vibrations. Control value Omega is calculated by integrating the curve Fs − zd (soil reaction on the movement of the vibratory roller drum) during the soil compaction with the vibratory roller in two successive vibration cycles [2]: Z ð5Þ Omega ¼ Þ ðmd
€zd þ ðmd þ mf Þ
g þ m0 e0
x2 Þ_zd dt 2T

where zd—the movement speed of the drum (m); €zd —the vertical acceleration of the drum (m/s2); m0 e0 —the eccentric mass moment (kg m); x—circular vibration frequency (rad/s); md—the weight of the vibratory roller drum (kg); and mf—the weight of the frame of the vibratory roller drum (kg). In the late 1990s, BOMAG has introduced the control value Vibration module Evib. Calculation of the phase difference in vibrations of eccentrics of the unbalanced modulus and the drum of the roller, as well as the soil reaction Fs, is performed based on the dual-mass rheological model. After calculating the movement of the vibratory roller drum zd, the graphs Fs − zd are formed for each vibration cycle and the slope of a specific part of the resulting curve Fs − zd to the axis zd. The values Fs, zd, and the slope value are compared to the expected values calculated based on the dependencies suggested by D. Lundberg in 1939 during the study of the static interaction of a rigid cylinder with elastic isotropic half-space with an elastic modulus E (Fig. 2) [4]:

New Criterion for Continuous Compaction C…

591

Fig. 2 Drum on elastic half-space and relationship between stiffness k and modulus E [4]

zd ¼

  2
ð1  m2 Þ Fs L
1:8864 þ ln ; p
E b L

ð6Þ

where m—Poisson’s ratio; Fs—soil reaction force (N); b—the length of a chord that pulls in a part of a circumference of a vibratory roller drum submerged into the soil (m); and L—the width of the vibratory roller drum (m). To determine the control value for a vibrating modulus Evib, the nomogram of dependence of the soil reaction from the movement of the vibratory roller drum is used [4]. In 1990, AMMANN presented a control value Rigidity rate (ks). The value of ks is calculated based on the analysis of the chart that shows the relation of the soil reaction force to the movement of the vibratory roller drum Fs − zd [13]: ks ¼ x2
ðmd þ m0 e0
cos u=zd Þ;

ð7Þ

where x—circular vibration frequency of the vibration exciter (rad/s); md—the weight of the vibratory roller drum (kg); m0 e0 —the eccentric mass moment (kg m); and u—the phase difference of the vibrations of the vibration exciter and the vibratory roller drum (grad). The ks rate is the slope of the line passing through the points of maximum immersion of the vibratory roller drum into the compacted soil (zd max) and the static weight of the vibratory roller drum (Fstatic). The movement values of the vibratory roller drum zd max are calculated by integrating the acceleration values of the drum (accelerometer data mounted on the frame of the vibratory roller drum). The operation of the most advanced systems of continuous control of soil compaction by vibratory rollers created by the leading manufacturers worldwide is based on the simulation of interaction of the vibratory roller drum in the vibratory roller with the compacted soil by using rheological models of various complexity and using the accelerometer data mounted on the vibratory roller drum in the vibratory roller to calculate the density (durability) of the compacted material [3].

592

I. S. Tyuremnov and A. S. Morev

The analysis of the values that are used as a basis for functioning of the system of continuous control of soil compaction by vibratory rollers created by the leading manufacturers worldwide shows that most manufacturers use systems based on the analysis of the vertical acceleration spectrum of the vibratory roller drum. Based on the analysis of the parameters, the construction and road machines department of the Yaroslavl State Technical University has developed a new parameter for the functioning of the system of continuous control of soil compaction by vibratory rollers called Compaction Value (CV) [14]:
CV ¼ K0:5f
A0:5f þ K1:5f
A1:5f þ K2f
A2f þ K2:5f
A2:5f  , þ K3f
A3f =K0:5f
A0:5f þ Kf
Af
K where K—the total calibration constant; K0.5f, Kf, K1.5f, K2f, K2.5f, K3f—the significance ratios of the amplitudes of harmonics of the vertical acceleration spectrum of the vibratory roller drum with frequencies 0.5f, f, 1.5f, 2f, 2.5f, 3f, respectively.

4 Conclusions Using the significance ratios of the amplitudes of harmonics of the vertical acceleration spectrum of the vibratory roller drum in calculation of the CV allows to fine-tune the continuous compaction control systems for a specific type of material. It also allows to increase the sensitivity of this parameter to the variance of the compacted soil. Based on the CV, it is possible to develop the system of continuous control of soil compaction by vibratory rollers in Russia, which is important to solve the problems of increasing the quality of roads and execute the program of import phaseout in the road construction machine building industry of the Russian Federation.

References 1. Tyuremnov IS (2009) The current status and design features of vibratory rollers. Mir dorog 42:64–69 2. Adam D (2007) Standardization, design, quality assurance and monitoring of earth works in road engineering in Austria. Vienna University of Technology Institute for Ground Engineering and Soil Mechanics Budapest, p 33 3. Tyuremnov IS (2016) The review of systems of continuous control of soil compaction for vibratory rollers Part 3. The specifics of the functioning and “intelligent compaction”. Bull PNU 2(41):115–122 4. Mooney MA (2007) Vibratory roller integrated measurement of earthwork compaction: an overview. In: Seventh international symposium on field measurements in geomechanics, pp 1–12

New Criterion for Continuous Compaction C…

593

5. Compactometer, compaction meter for vibratory rollers. ALFA-022R [Digital] / Geodynamik. The company on safe ground service. http://www.geodynamik.com/languages/pdf/. Accessed 05 Dec 2016 6. Sandström AJ, Pettersson CB (2003) Intelligent Systems for QA/QC in soil compaction. Geodynamik HT AB Box 7454 SE-103 92, Stockholm, Sweden, pp 1–17 7. Trimble CCSFlex and GCS900/AccuGrade compaction control system for soil and Asphalt compactors. http://www.webpages.uidaho.edu/bayomy/IAC/52nd/Presentations_52nd/. Acce ssed 05 Nov 2018 8. Tyuremnov IS, Morev AS (2015) The review of systems of continuous control of soil compaction for vibratory rollers. Part 1. Bull PNU 4(39):99–108 9. Scherocman JA, Rakowski S, Uchiyama K, Intelligent compaction, does it exist? Intelligent compaction. [Official website]. http://www.intelligentcompaction.com/downloads/Papers Reports/. Accessed 05 Dec 2016 10. Chang GK, Xu Q, Rasmussen R, Merritt D, Michael L, White D, Horan B (2010) Accelerated implementation of intelligent compaction technology for embankment subgrade soils, aggregate base, and asphalt pavement. IC Data Management, Federal Highway Administration Office of Pavement Technology, p 118 11. Tyuremnov IS, Morev AS (2015) The review of systems of continuous control of soil compaction for vibratory rollers. Part 2. Bull PNU 1(40):69–76 12. White DJ, Vennapusa PKR (2010) A review of roller-integrated compaction monitoring technologies for earthworks. Earthworks Engineering Research Center (EERC), Final Report ER10-04, April 2010, p 31 13. Mooney MA, Rinehart RV, Facas NW, Musimbi OM (2010) Intelligent soil compaction systems. Report 676 Available via DIALOG. http://onlinepubs.trb.org/onlinepubs/nchrp/. Accessed 04 Dec 2014 14. Morev AS (2017) The validation of a soil compaction parameter for a system of continuous control of soil compaction for vibratory rollers. Dissertation, Yaroslavl

Reducing Amplitude of Load Swinging During Operation of Hydraulic Manipulators of Forest Transport Machines P. Popikov, M. Drapalyuk and D. Druchinin

Abstract The problem of loads swinging during the operation of forest transport machines equipped with a hydraulic manipulator is considered. It is noted that the starting and braking modes of hydraulic manipulators operation are characterized by high dynamic loads and significant fluctuations in the pressure of the working fluid. To eliminate these negative events, various hydraulic apparatus (safety valves, throttles, dampers, etc.) are used. At the same time, a promising direction for solving this problem is the use of an energy-saving hydraulic actuator that is able to accumulate energy during the start–stop operation and then return to the system. To determine the factors affecting the reduction of dynamic loading of hydraulic manipulators, and to identify the conditions for reducing the sway of the load, a simulation mathematical model of the manipulator, equipped with an energy-saving hydraulic actuator, was developed. Within the framework of the model, physical processes occurring in the mechanical and hydraulic subsystems of a manipulator equipped with an energy-saving hydraulic actuator and an additional damper are reproduced. On the basis of the model, a theoretical study was conducted in the form of a series of computer experiments in which the main parameters of the energy-saving hydraulic actuator, manipulator and its working conditions were alternately changed. The results of theoretical experiments showed that the use of energy-saving hydraulic actuator can significantly reduce pressure surges in the hydraulic system, and the movement of load in space becomes smoother.







Keywords Hydraulic manipulator Load swinging Energy-saving hydraulic actuator Hydro-pneumatic accumulator Simulation modeling Pressure surge Amplitude fluctuations













P. Popikov (&)
M. Drapalyuk
D. Druchinin Voronezh State University of Forestry and Technologies named after G.F. Morozov, 8, Timiryazev Street, Voronezh 394087, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_63

595

596

P. Popikov et al.

1 Introduction Operating modes of hydraulic manipulators of forest transport machines in the starting-brake modes are characterized by high dynamic loads, as well as the swinging of the gripping devices when they are aimed at timber. Another negative point in the operation of the hydraulic manipulators is swinging of an already captured object during its subsequent laying in a predetermined position, since the productivity of work decreases when the load is swinging due to an increase in the duration of the cycle for a period of gripping body damping. Safety valves, annihilator, chokes, or dampers are used to limit maximum fluctuations in the pressure of the working fluid. Disadvantage of these hydraulic devices is that the limitation of pressure fluctuations of the working fluid occurs due to its flow from one cavity to another through hydraulic resistance, while the hydraulic energy is converted into heat, which leads to overheating of the fluid and energy losses. Energy-saving (recuperative) hydraulic actuators, which are able to accumulate energy during the start-braking mode and return some of the energy back to the system during reverse movement, are the most effective ones [1, 2]. Specific working conditions of various technical devices should be taken into account when justifying the parameters of hydraulic actuator and its operating modes. Workflows of machines occur in unsteady modes due to variable loads. Therefore, it is necessary to consider not only energy and speed indicators but also fluctuation properties of a hydraulic actuator to determine the frequencies of its free oscillations and areas of resonant modes when designing and developing machines with hydraulic drive [3–5].

2 Materials and Research Methods One of the ways to reduce dynamic loads and energy consumption during the operation of manipulators is to use energy-saving recuperative hydraulic actuator equipped with a hydraulic accumulator, due to which braking energy is not transformed into heat but accumulated and then it returns to the system [6, 7]. When analyzing the dynamics of the hydraulic actuator of the working bodies, the Navier–Stokes equations are used in cylindrical coordinates, x-axis of which is directed along the axis of the pipeline, and r coordinate—along the radius of its cross section [2]:  2
 @ux @ux @ux 1 @p 4 @ ur @ 2 ur 1 @ux 1 @ @ur ur þv þ ux þ ur ¼ þ þ ; þ þ q @x 3 @x2 r @r 3 @x @r @t @x @r @r2 r

ð1Þ

Reducing Amplitude of Load Swinging …

597

 2
 @ur @ur @ur 1 @p 4 @ ur 4 @ur 4 ur @ 1 @ux @ur þv þ ur þ ux ¼  þ þ þ q @r 3 @r2 3
r @r 3 r 2 @x 3 @r @t @r @x @x

ð2Þ where ux and ur—fluid velocity projections on the x and r axes; q—fluid density; p—fluid pressure; v—coefficient of kinematic viscosity. The following expressions are obtained taking into account these equations and after transformations Z (s) sh [ϑ ( s )1] + 2 ch [ϑ ( s )1] Z В. Л . ( S ) , Z1 ( s ) = Z В. Л . ( s ) ð3Þ Z 2 (s) sh [ϑ ( s )1] + ch [ϑ ( s )1] Z В. Л . ( s )

Z 2 ( s ) = Z В. Л . ( s )

Z1 ( s ) ch [ϑ ( s )l ] − sh [ϑ ( s )l ] Z В. Л . ( S ) , Z (s) ch [ϑ ( s )l ] − 2 sh [ϑ ( s )l ] Z В. Л . ( s )

ð4Þ

where Z1(s) and Z2(s)—terminal operator resistances; Zv.i.—operator wave impedance; s—variable in the Laplace transform; tðsÞ—Laplace image of the average velocity of the medium in the flow section. The transfer function of the hydraulic system, equal to the ratio of Laplace images in pressure in the output and input sections, has the form W (s) =

p2 ( s,1) 1 , = p1 ( s, 0) Z В. Л . ( s ) sh [ϑ ( s )1] + ch [ϑ ( s )1] Z 2 (s)

ð5Þ

where p(s)—Laplace image of medium pressure flow. The theoretical substantiation of the influence of the dynamics of the hydraulic actuator on the load swinging has been carried out using the Lagrange equation for potential and generalized potential forces 9 2 € P þ ml2 Þ þ mlb ll ðw € cos w  w_ 2 sin wÞ ¼ pdn d ðP1  P2 Þ  Mb = /ðJ b 8 ; € þ lb / € cos w þ g sin w ¼ 0 ; ll w

ð6Þ

where m—mass of the load, JR —total moment of inertia of the boom group and column relative to the axis of rotation of the column and the load relative to figure axis, kg m2: JP ¼ Ib cos2 /b þ Il þ ml2b cos2 /b þ Ik ;

ð7Þ

u—angle of column rotation (deg); w—angle of load deviation from the vertical (deg); ll—length of load installation (m); lb—radius of boom group (m).

598

P. Popikov et al.

However, the influence of the characteristics of dampers, compressibility of the actuating fluid, and flexibility of hydraulic system elements have not been considered in the conducted studies of the dynamic load of carrying and lifting machines with hydraulic actuator [1, 3, 4, 7, 8]. The aim of the work is to determine the factors affecting the reduction of dynamic loading of hydraulic manipulators and to identify the conditions for reducing load swinging during operation of hydraulic manipulators of forest transport machines equipped with energy-saving hydraulic actuator. Recently, abroad, numerical simulation methods are used to study the processes occurring in complex technical systems, which enable to take into account a wide range of parameters of the constituent elements and the actuating medium of the object under study, which provides more accurate calculation of its functioning [9, 10].

3 Theoretical Part The authors have developed a simulation mathematical model of the functioning of the manipulator, equipped with an energy-saving hydraulic actuator [11, 12]. The model reproduces physical processes occurring in mechanical and hydraulic subsystems of manipulator, equipped with energy-saving hydraulic actuator and additional damper, the work of which is described by a system of differential and algebraic equations [13]. The solution of the system of equations is performed by numerical integration while analyzing the time dependences of the main parameters characterizing operation of manipulator and hydro-pneumatic accumulator: pressures in the cavities of the pivot actuators PR(t) and PL(t), tangential and radial load fluctuations. Modeling the actuating process of the column rotation mechanism, rotational motion of the manipulator around the vertical axis, swinging of the load, and progressive motion of the piston of hydro-pneumatic accumulator has been taken into account. The pivot column of the manipulator rotates in the horizontal plane XOY relative to the vertical axis O. The angular position of the column is set by the angle u, measured from the OX direction counterclockwise. In the mathematical model, hydraulic system of the manipulator with the energy-saving system is represented as several separate cavities containing actuating fluid or gas: the cavities of the left and right pivot hydraulic cylinders (denoted by the letters “L” and “R” on Fig. 1), fluid and gas cavities of hydro-pneumatic accumulator “F” and “G”. In addition, “HP” and “D” indices further denote a connection with high-pressure and drain tube, respectively. Figure 1 shows a diagram of the connection of cavities in the mode of braking of manipulator column rotation in clockwise direction. The excess pressure in the “L” cavity of the left cylinder leads to the pumping of actuating fluid into the “F” cavity of the hydro-pneumatic

Reducing Amplitude of Load Swinging …

599

Fig. 1 Main elements of energy-saving hydraulic actuator: design scheme for modeling hydraulic subsystem in the left hydraulic cylinder braking mode

accumulator. The valve used in hydraulic system similarly provides a hydraulic connection of the “R” cavity with the “F” cavity of the hydro-pneumatic accumulator braking the rotation in counterclockwise direction. Complex equations, including the equation of piston motion, equation of consumption at the entrance and at the state of polytropic process in the gas cavity, are used to describe the processes in hydro-pneumatic mechanism: m_ P ¼

1 ½SHPA ðPF  PG Þ  hmP  ðR0fr þ ki PF ÞsighmP  cðxP þ xP0 Þ; mM x_ P ¼ mP ;

0  xP \LHPA  bP ¼ VHPA =SHPA ;

P_ F ¼ ðQF  SHPA mP Þ=kel:F

PG ¼ PG0 ½xP =ðLHPA  bP Þn Pat: ;

ð8Þ ð9Þ ð10Þ

where vP mM SHPA ¼ pD2HPA =4 DHPA PF and PG h R0fr ki = (p/20)DHPA H H c

speed of piston movement which separates gas and fluid cavities; mass of the movable element of accumulator; working area of the (piston) accumulator; piston diameter (internal accumulator cavity); pressure, respectively, in the fluid and gas parts of hydro-pneumatic accumulator; viscous friction coefficient; dry friction force in the absence of pressure; coefficient of proportionality between friction force and pressure in the actuating cavity; lip seal height; hydraulic accumulator rigidity;

600

P. Popikov et al.

xP and xP0

the current coordinate of the piston of hydro-pneumatic accumulator and value of the preliminary compression of the spring; LHPA length of actuating cavity of the hydro-pneumatic accumulator; bP piston thickness of hydro-pneumatic accumulator; VHPA total volume of hydro-pneumatic accumulator; QF the flow rate of actuating fluid entering the cavity with fluid; Kel.F = (DVHPA + xCSHPA)/Ered coefficient of elasticity of a cavity with fluid; Ered = EF/[1 + (DHPA/d)(EF/Ecт)] reduced volume modulus of elasticity of a fluid cavity; DVHPA “dead” volume of actuating chamber; PG0 gas loading pressure (absolute); N polytropic index; Pat atmosphere pressure. When using the developed model, the main computational complexity is associated with integration of differential equations. Focusing on the use of calculating power of the computer, calculation is organized in an iterative way. The current value of the rotation angle of the column u is determined by the position of the piston xG in the left hydraulic cylinder: xG ¼ xG0 þ /
RC ;

ð11Þ

where xG0—piston position at u = 0. As a result of determining the position x of the piston of the left hydraulic cylinder and the position x of the piston of the hydro-pneumatic accumulator, the volumes of the cavities of the hydraulic cylinder (VL, VR) and fluid cavity of the accumulator are calculated (VF): VL ¼ xG

pD2G ; 4

VR ¼ ðLH  xG Þ

ð12Þ

pD2G ; 4

VF ¼ ðLHPA  xP  bP Þ

pD2HPA ; 4

ð13Þ ð14Þ

where LH—the length of the working cavity of the hydraulic cylinder; When changing the volume of cavities due to movements of the pistons of hydraulic cylinders of rotation of the column and piston of a hydro-pneumatic accumulator, the pressures PL, PR, PF in these cavities change, respectively, which are calculated by the formula at the kth integration step

Reducing Amplitude of Load Swinging …

601

Pkm ¼ Pk1 m E

Vmk  Vmk1 ; Vmk

ð15Þ

where m—index means cavity in which pressure is calculated, and it can take values “L”, “R,” and “F”. Some pressures may turn out to be quite small in the previous step, and they can become negative in step k. Therefore, they are adjusted, if Pkm \0, then assignment is made Pkm ¼ 0. Under the influence of a difference in respective pressures, fluid flows from one cavity to another. The following options are possible for the fluid flows for the hydraulic scheme of the column rotation mechanism (Fig. 1). • flow from “L” cavity into “F” cavity when braking the rotation of the column “clockwise”: (

if

pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi VL ¼ VL  kLF PL  PF Dt; then pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi VF ¼ VF þ kLF PL  PF Dt:

PL [ PF ;

ð16Þ

• flow from “R” cavity into “F” cavity when braking the rotation “counterclockwise”: ( PF [ PL ;

if

then

pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi VR ¼ VF  kRF PR  PF Dt; pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi VF ¼ VF þ kRF PR  PF Dt:

ð17Þ

In the main modes (“clockwise rotation” and “counterclockwise rotation”) of manipulator, it is necessary to take into account not only the flow of fluid from one cavity into another cavity but also the flow of fluid from hydraulic pump and the fluid outflow into the drain tube. In “turn the column counterclockwise” mode, the estimated fluid flows are pre-calculated: • Flow from the high-pressure line “HP” in the “L” cavity:

if

PHP [ PL ;

then

pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi QHPL ¼ kHPL PHP  PL :

ð18Þ

If the flow rate QHPL is greater than the nominal supply of the hydraulic pump Qnom, the flow correction is made:

602

P. Popikov et al.

QHPL ¼ Qnom

ð19Þ

And, a new volume of fluid is calculated in the “L” cavity after adjustment: if

PHP [ PL ;

then

VL ¼ VL þ QHPL
Dt:

ð20Þ

Similarly, in “clockwise rotation of the column” mode, the flow from the “HP” line to the “R” cavity is calculated: if

PHP [ PR ;( then pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi kHPR PHP  PR
Dt; VR ¼ VR þ Qnom
Dt;

if if

pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi kHPR PHP  PR \Qnom ; pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi kHPR PHP  PR  Qnom :

ð21Þ

When actuating fluid is fed into one of the cavities from the pressure line (“L” or “R”), the second cavity (respectively, “R” or “L”) is connected to the low-pressure “A” line in the main modes of rotation of the manipulator column. Drainage of actuating fluid is modeled by the following equations: • Fluid outflow from the “R” cavity to the “A” low-pressure line in “counterclockwise” mode:

if

PR [ P;

then

pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi VR ¼ VR  kR PR  PA
Dt;

ð22Þ

• Fluid outflow from the “L” cavity to the “A” low-pressure line in “clockwise” mode:

if

PL [ P;

then

pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi VL ¼ VL  kLA PL  PA
Dt:

ð23Þ

To solve the system of differential equations that make up the mathematical model, and to conduct systematic computer experiments, a computer program in Object Pascal language has been written, which allows producing schematic images of the manipulator and load in three projections, pivot hydraulic cylinders, and pressure graphs in the cavities of hydraulic accumulator and load deviations from the equilibrium position in tangential and radial directions. The developed model has made it possible to test the effectiveness of equipping the manipulator with an energy-saving hydraulic actuator in the form of reducing peak pressure values in the hydraulic system and eliminating unfavorable load swinging. The study of the model has been carried out on the basis of a series of computer experiments, in which the main parameters of energy-saving hydraulic actuator,

Reducing Amplitude of Load Swinging …

603

manipulator, and conditions of its operation have been alternately changed. The experiment was a sequential combination of the following modes: • initial execution time to enter the state of initial equilibrium for the mechanical and hydraulic subsystems is 10 s; • rotation of the column “counterclockwise” from the time t = 10 s to the angular position u = 30°; • rotation braking by connecting the cavity of the right rotation cylinder to the hydro-pneumatic accumulator; • initial execution time for mechanical and hydraulic subsystems for damping the oscillations of load and pressure of actuating fluid is t = 20 °C; • rotation of the column “clockwise” from the time t = 20 s to the initial angular position u = 0°; • rotation braking by connecting the cavity of the left rotation cylinder to the hydro-pneumatic accumulator; • initial execution time for mechanical and hydraulic subsystems for damping the oscillations of load and pressure of actuating fluid is t = 20 °C.

4 Results The simulation model enables both to connect hydro-pneumatic accumulator in the mode of braking the column rotation and to check the option of abrupt locking of supply and drain hydraulic lines. In the second case, pressure spikes appear on PR(t) graph. Moving the boom causes load swinging in tangential and radial directions, which quantitatively characterize the graphs of damped oscillatory processes As(t) and Ar(t). The developed energy-saving hydraulic actuator permits not only to store and further use braking energy but also to reduce fluctuations in the load and pressure in the hydraulic system caused by the process of braking the rotation of the column.

4.1

Reducing Pressure Surges in the Hydraulic System

To evaluate the effective operation of an energy-saving hydraulic actuator in the form of smoothing of pressure surges of the actuating fluid while braking the column rotation, two computer experiments have been conducted. In the first experiment (“without a hydro-pneumatic accumulator”), instantaneous locking of the entrances to the rotation hydraulic cylinders was performed for braking the rotation of the column. In the model, this was implemented by zeroing the throttling coefficients kHPL, kHPR, kLA, kRA, kLF, kRF, responsible for the entry and fluid outflow of actuating fluid from the cavities of the pivot hydraulic cylinders. In the second computer experiment (“with a hydro-pneumatic accumulator”), a hydraulic

604

P. Popikov et al.

cylinder in which an increase in pressure took place was connected with a hydro-pneumatic accumulator when braking a rotation. The results of theoretical experiments have showed that the use of energy-saving hydraulic actuator can significantly reduce pressure surges in hydraulic systems (Fig. 2). Abrupt locking during braking results in a significant surge in pressure of 33 MPa relative to Patm, or 13 MPa relative to operating pressure of 20 MPa (see Fig. 2a). Using the same energy-saving hydraulic actuator enables to almost completely eliminate the pressure surge: AP rises only up to 3 MPa when braking (initial gas pressure in the hydro-pneumatic accumulator) and then actuating fluid is dumped into the accumulator.

4.2

Reducing Load Swinging

Due to the fact that the discharge of pressure in a hydro-pneumatic accumulator reduces surges and fluctuations in pressure in the hydraulic system, the movement of the load in space becomes smoother. This allows the manipulator operator to guide the locking mechanism onto a load quickly or to position an object which is already captured in space.

Fig. 2 Elimination of pressure surges in the hydraulic system when using a hydro-pneumatic accumulator: a significant pressure surges without an accumulator; b small pressure surges when using the accumulator

Reducing Amplitude of Load Swinging …

605

Fig. 3 Reduction of load swinging in tangential direction when using hydro-pneumatic accumulator: a without accumulator; b using accumulator

Load swinging causes a sharp change in the angle of rotation of the boom at the beginning of the turn and braking of the rotation. Figure 3 shows the dotted lines with the points in time at which the turn starts or stops. Comparing Fig. 3a, b, it can be concluded that energy-saving system enables to reduce swinging amplitude in the tangential direction when braking the rotation “counterclockwise” from 29 to 24 cm (by 17%), and when braking the rotation “clockwise” from 30 to 26 cm (by 13%). Load swinging in the radial direction also decreases from 19–28 to 12–13 cm (by 32–54%) when using an energy-saving hydraulic actuator (Fig. 4). However, the radial swinging during rotational movement of the column is insignificant compared with the tangential direction: the amplitude is only about 45 cm and its decrease by 6–15 cm is almost negligible.

606

P. Popikov et al.

Fig. 4 Reduction of load swinging in radial direction when using hydro-pneumatic accumulator: a without accumulator; b using accumulator

4.3

Improving the Oscillation Spectrum of the Load

For a detailed analysis of the nature of the oscillations of the load in the tangential direction, spectra of oscillations As(f) (or often called amplitude-frequency characteristics) are calculated. Figure 5 shows the spectra for the case of instantaneous stopping of the movement of actuating fluid during braking of the column (thin line) and the case of depressurization of hydro-pneumatic accumulator. In accordance with the physical meaning of the amplitude–frequency characteristics, the presence of a peak at a certain frequency on the graph means that load swinging with such a frequency are more pronounced than swinging with other frequencies. Judging by the location of the peaks in the spectrum, the main frequency of tangential oscillations of the load is about 0.5 Hz, i.e., the oscillation period is about 2 s. In addition, vibrations in the frequency range of 0.8–1.5 Hz (the oscillation period is 0.67–1.25 s), corresponding to the load and manipulator shake (as a mechanical system) when starting and stopping the rotation seem to be essential.

Reducing Amplitude of Load Swinging …

607

Fig. 5 Amplitude–frequency characteristics (spectra) of tangential oscillations of the load for the absence and presence of hydro-pneumatic accumulator

The main improvement in the spectrum due to equipping the manipulator with an energy-saving hydraulic actuator is in significantly reduced overall level of peaks in the range of 0.8–1.5 Hz. This means that energy-efficient hydraulic drive significantly reduces load shaking.

5 Conclusion Thus, energy-saving hydraulic actuator enables to virtually eliminate pressure surges (to reduce from 33 to 3 MPa) while braking the rotation of the manipulator column. In this case, pressure surges do not exceed 10 MPa, and swinging amplitude of the load does not exceed 27 cm. Equipping the manipulator with an energy-saving hydraulic actuator enables to reduce 2–3 times the amplitude of load shaking (with a frequency of 0.8–1.5 Hz) with an abrupt change in the mode of boom movement, but it does not allow reducing load gravitational swinging (with a frequency of 0.50–0.55 Hz). This permits to guarantee the reliable operation of the hydraulic system and the use of high-pressure hoses designed for lower pressures, reduce energy costs and improve the performance of loading and unloading operations performed by timber transport machines during timber logging.

References 1. Shcherbakov VF (2011) Jenergosberegajushhie gidroprivody stroitel’nyh i dorozhnyh mashin (Energy-saving hydraulic drives of construction and road machines). Stroitel’nye i dorozhnye mashiny (Constr Road Mach) 11:43 2. Popikov PI (2001) Povyshenie jeffektivnosti gidroficirovannyh mashin pri lesovosstanovlenii na vyrubkah (Improving the efficiency of the hydraulic machines during forest clearings). Dissertation, Voronezh State Forestry Academy

608

P. Popikov et al.

3. Vasilyev S, Eresko S, Eresko T (2011) Zakonomernosti formirovanija pul’sacij davlenija v gidroprivode transhejnyh rotornyh jekskavatorov pri razrabotke merzlyh gruntov (Patterns of formation of pressure pulsations in the hydraulic drive of trench-digging rotary excavators in the exploitation of frozen soils). Stroitel’nye i dorozhnye mashiny (Constr Road Mach) 8: 37–40 4. Gerasimov YY, Siounev VS (2000) Harvester crane key parameters: optimization in European Russian Pines. J For Eng 8:65–67 5. Kovalev AS, Pilyushina GA (2018) Povyshenie rabotosposobnosti gidrooborudovanija lesozagotovitel’nyh mashin (Increased efficiency of hydraulic equipment forestry machines). Aktual’nye problemy lesnogo kompleksa (Actual Probl Complex) 52:9–13 6. Nurmi J, Mattila J (2017) Global energy-optimal redundancy resolution of hydraulic manipulators: Experimental results for a forestry manipulator. Energies 10:647 7. Aranovskiy S, Losenkov A, Vazquez C (2014) Position control of an industrial hydraulic system with a pressure compensator. In: 22nd mediterranean conference on control and automation, art. no. 6961560, pp 1329–1334 8. Rybak AT, Temirkanov AR, Lyakhnitskaya OV (2018) Synchronous hydromechanical drive of a mobile machine. Russ Eng Res 3:212–217 9. Westerberg S, Shiriaev A (2013) Virtual environment-based teleoperation of forestry machines: Designing future interaction methods. J Hum-Rob Interact 2:84–110 10. Liu Y, Handroos H (2003) Modelling and control of a manipulator with hydraulic actuators. Am Soc Mech Eng, Fluid Power Syst Technol Div 10:203–208 11. Dolzhenko SV, Popikov PI, Titov PI (2012) Matematicheskaja model’ rabochego processa gidroprivoda mehanizma povorota kolonny gidromanipuljatora pri pogruzke sortimentov (Mathematical model of the working process of the hydraulic actuator of the hydraulic manipulator column turning mechanism when loading assortments). Lesotehnicheskij zhurnal (For Eng J) 3:84–90 12. Popikov PI, Titov PI, Sidorov AA et al (2011) Matematicheskoe modelirovanie processov v sisteme gidroprivoda lesnyh manipuljatorov (Mathematical modeling of processes in the hydraulic actuator system of forest manipulators) Politematicheskij setevoj jelektronnyj nauchnyj zhurnal Kubanskogo gosudarstvennogo agrarnogo universiteta (Polythematic Netw Electron Sci J Kuban State Agrar Univ) 69:96–106 13. Popikov PI, Oboyantsev DV, Dolzhenko SV (1998) Mehanizm povorota kolonny strelovogo manipuljatora (The mechanism of rotation of the column boom arm). Russian Federation Patent 2479481 (20 April 2013)

Increase in Operational Reliability of Locking Device of Hydraulic Actuator of Vehicle N. A. Fomenko, S. V. Aleksikov and S. G. Artemova

Abstract In modern mechanical engineering, in construction and road, farm, meliorative, timber, industrial, and other land vehicles of special and general purpose hydraulic actuators of working elements, which have advantages regarding speed and smoothness of reverse or separate management, change of position of the car, accuracy of positioning, and dispensing of power impact on working bodies are widely applied. At the same time, the working pressure is optimized for the prevention of destruction of the hydraulic line of high pressure and unauthorized emission of working liquid in a hydraulic system; however, under operating conditions, the probability of emission of working liquid in the atmosphere remains. For the increase in environmental safety, protection systems of a hydraulic actuator are applied. Among the known ways of protection, a technical solution in which the locking device is equipped with the relieving valve is considered. However, the analysis of its work shows that at operation of a protection system of a hydraulic actuator, the locking device has an essential shortcoming. For the purpose of elimination of this shortcoming, it is offered to optimize not only operating pressure in a hydraulic system but also a design of the locking device.













Keywords Hydraulic actuator Hydraulic line Pipe Locking device Plunger Relieving valve Working liquid







1 Introduction The analysis of operation of a hydraulic actuator shows that hydraulic lines of high pressure are exposed to destruction and, as a result, to unauthorized emission in the atmosphere of working liquid, that results in economic and ecological damage. In this regard, the set of actions for elimination of the mentioned shortcoming is developed. However, there is still a problem of their improvement. This paper is N. A. Fomenko (&)
S. V. Aleksikov
S. G. Artemova Volgograd State Technical University, 1, Akademicheskaya Str., Volgograd 400074, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_64

609

610

N. A. Fomenko et al.

dedicated to the parametrical characteristics research of a prototype (Fig. 1). On the basis of the found shortcomings the technical solution is offered. This solution is protected by the patent copyright (Fig. 2). The main idea of this solution is to preserve the functional purpose, to provide reliability and to decrease in material consumption and operating costs. The novelty of technical solution is that the stop of the locking device of a protection system of a hydraulic actuator has the pipe connecting an output cavity of the locking device with the hydraulic line of high pressure and intended for

Fig. 1 Relieving valve of protection system of a hydraulic actuator: 1—tank, 2—pump, 3— pressure head hydraulic line, 4—drain line in a hydraulic tank, 5—hydraulic valve, 6—hydraulic engine, 7—locking device, 8—relieving valve, 9–10—case, 11–13—union, 14—stop, 15— plunger, 16—groove, 17—radial openings, 18—axial pipe, 19—plunger, 20—through opening, 21–22.25—elastic element, 23—cone 23, 24—fillet, 26—piston, 27—sealing ring, and 28—pipe

Increase in Operational Reliability …

611

Fig. 2 Relieving valve of protection system of a hydraulic actuator: 1—tank, 2—pump, 3— pressure head hydraulic line, 4—drain line in a hydraulic tank, 5—hydraulic valve, 6—hydraulic engine, 7—locking device, 8—case, 9, 10, 11—union, 12—stop, 13, 17—plunger, 14—groove, 15, 16, 18—opening, 19, 20—elastic element, 21—cone, 22—fillet, 23—pipe, and 24—hydraulic line

release of the forced-out incompressible working liquid, which is formed at the end of the plunger between plunger and stop, into the pressure head of hydraulic line of a nozzle. Researches show [1–4] that at increase in power of a hydraulic actuator due to increase in operating pressure and application of various ways of protection [5–24], unauthorized emission in the atmosphere of working liquid, which causes significant damage to environmental safety, is still observed.

612

N. A. Fomenko et al.

2 Problem Definition Figure 1 shows the analysis of shortcomings of the protection system of a hydraulic actuator [25] and development of technical solution after its modernization are shown on Fig. 2 [26].

3 Theoretical Part In land traction vehicles, construction and road, industrial vehicles, and other hydroficated working equipment, the hydrostatic actuator is widely used. Working process of such vehicles, as a rule, is followed by sharply expressed dynamic loads. Therefore, in technological working process of vehicles of any appointment the load mode constantly changes. The operator changes the direction of a power stream of working liquid that causes dynamic loads in high pressure hoses, which lead to their destruction and unauthorized emission in the atmosphere of working liquid, up to 800 times per work day by switching of the hydraulic valve of the hydraulic system. At an annual operating mode, for example, of the dozer the number of turning on of the hydraulic valve makes about 200 thousand and recurrence of loading up to 6 s and that causes fatigue tension in hoses. For the purpose of increase in capacity of machines, the developer is forced to increase power and nominal working pressure in the hydraulic system that also leads to fatigue failure of high pressure hoses and unauthorized emission of working liquid in the atmosphere (for power saturated vehicles up to 300 l. for one emission) that causes damage to environmental safety. Therefore, one of the main actions of preventing destruction of high pressure hoses in operation is optimization of operating pressure: for the hydraulic systems of 20 MT equipped with flexible high pressure hoses up to 18 …, and for metal up to 45–60 MPas. However, restriction of operating pressure for the hydraulic systems equipped with flexible hoses does not solve a problem of destruction of high pressure hoses and emission in the atmosphere of working liquid; therefore, improvement of design of protection systems of hydraulic actuators and technological complexes and technologies, search of ecologically safe, and harmless working liquids for the environment are a relevant task. One of the directions of increase in reliability of protection of a hydraulic actuator is improvement of pipelines. The nomenclature of metal pipelines with hinged joints with multiple margin of safety does not guarantee tightness; therefore, search of new technical solutions is required. Researches of authors show that the hose with a gas cover increases its reliability, and however, covering and coiling hoses do not guarantee decrease in cyclic dynamic loads and their destruction with growth of operating pressure. Effective protection of a hydraulic actuator can be considered in case of application of a hydromechanical way. These researches show that the easiest way of

Increase in Operational Reliability …

613

protection of a hydraulic actuator allows to reduce losses at operation of the locking device up to 0.17 l.; however, the shortcoming inherent in it considerably increases these losses in operation. Lack of this system of protection of a hydraulic actuator is that at destruction of high pressure hoses at the end of slide valve course, there is closed incompressible volume of moisture between slide valve and support, which means that there is a “hydraulic pillow,” which interferes with movement slide valve and reliable overlapping of the damaged hydroline. High pressure hoses of a hydraulic actuator of vehicle working bodies [2, 4] not always maintain long cyclic loads with the regulated nominal pressure 18–20 MPas and are often exposed to destruction. However, the ways of protection of a hydraulic actuator given above not always meet the modern requirements. For example, at loss of elasticity of a sealing ring of the relieving valve (Fig. 1) of the locking device or its destruction, there will be a discharge of working liquid in a hydraulic tank, passing the hydraulic valve and the hydraulic engine in its piston cavity, and it will cause pressure drop in an output cavity of the locking device and, as a result, delay of rise or lowering of working body or its stop. As a result, there will be an imitation of damage of the pressure head hydraulic line, that is, false operation of the locking device that reduces overall performance of a system of protection.

4 Practical Importance, Offers and Results of Introduction, and Results of Pilot Studies Technical solution (Fig. 2) allows to significantly increase functional purpose and operational reliability of a protection system of a hydraulic actuator. In this case, the design of locking the device of protection system of a hydraulic actuator has a pipe in the stop, connecting an output cavity of the locking device with the pressure head hydraulic line located between the locking device and the hydraulic valve for transportation of the locking device of the incompressible volume of working liquid, put between the end stop and plunger at the end of its run. It allows to provide discharge of working liquid from an output cavity of the locking device directly to the pressure head hydraulic line, which is forced out from an output cavity. Besides absorption of energy of water hammer in the response time of the device reduces wear of a cone of the valve and oscillatory process of a plunger when closing the pressure line by means of restriction of the forced-out incompressible volume of working liquid via the pipe in the stop, at the same time the design becomes simpler and function of the relieving valve remains. For example, we have the system of protection of a hydraulic actuator including a hydraulic tank, the pump connected by the pressure head hydroline to a hydraulic actuator via distributor, the line of discharge in a hydraulic tank and the locking device including case with input, output and solid cavities, input, output and drain

614

N. A. Fomenko et al.

nozzles with channels, installed in the pressure head hydroline and connected by the input cavity to the output of the pump, and output to the distributor, with the sprung plunger placed in input cavity having the circular groove combined with radial openings and the axial channel connected with an output cavity (Fig. 1), which is blocked by the plunger conic valve, having hydraulic pressure head from a hydraulic pump, and the locking device is in addition equipped with the unloading valve with nozzle, rigidly mounted in boring of support of the locking device, at the same time the unloading valve is executed in the form of a hollow cylinder with placed sprung by the plunger, supplied with a sealing ring from elastic material inside and the plunger divides a cylindrical cavity of the unloading valve into two pieces: The plunger cavity, which is connected with an output cavity of the case of the locking device of the channel, executed in support of the locking device and the spring chamber which is connected with the channel of the nozzle with the drain line. In case of rupture of hydraulic line, power supply of hydraulic engine 6 (Fig. 2) pressure in an output cavity B instantly falls also because of pressure difference of working liquid in cavities A and B. The plunger 13 of input cavity, overcoming resistance of a spring 20, moves to the right from end stop 12, blocking the section of pipe 16 through passage and section h of the output pipe 10, at the same time a groove 14 and radial openings 15 of the plunger 13 are connected with drain pipe G of union 11 and a hydraulic tank 1 and working liquid from pump 2, passing the damaged hydraulic line of high pressure of hydraulic engine 6, through pressure head hydraulic line 3, cavity A, groove 14 and radial openings 15 of the plunger 13 goes to channel G and further through drain hydraulic line 4 in a hydraulic tank 1. At the same time, the volume of incompressible working liquid between end stop 12 and plunger 13 at the end of its run from an output cavity B of locking device 7 through channel 23 put in the end stop 12, hydraulic line 24 comes to the pressure head hydraulic line 3 located between pipe B of union 10 and hydraulic valve 5, at the same time free movement of a plunger of the locking device of the size of its full speed (h + t) is provided, oscillatory process of a plunger is eliminated, and reliable blocking of the section through passage h of channel B of output union 10 is provided. The system of protection of a hydraulic actuator includes a hydraulic tank 1, pump 2, pressure head hydroline 3, drain line 4 into a hydraulic tank 1, hydraulic valve 5, hydraulic engine 6, and locking device 7. Locking device 7 includes case 8 with input 9, output 10, drain 11 nozzle with W and G channels, respectively, and support 12. Inside the case 8 of locking device 7, the sprung plunger 13 with an input cavity is placed A. Plunger 13 is located in the case 8 at distance of its full speed equal (h + t) where h is the section of the channel through passage of output nozzle 10, and t—distance from an end of the support 12 to forming the channel output nozzle 10, with formation of output cavities B between a plunger 13 and an end of the support 12. The external surface of a plunger 13 has a circular groove 14 at distance from a vertical axis of drain nozzle 11—(h + t) equal to a full speed of a plunger 13, and radial openings 15 at distance from a vertical axis of drain nozzle 11—(h t) equal to also full speed of a

Increase in Operational Reliability …

615

plunger 13 is executed. The input cavity A is connected with the output cavity B of case 8 by means of axial channel 16 which is blocked by the sprung plunger 17 with a possibility of its free movement in a through opening of the 18 support 12 with formation of a cavity D. Both plungers 13 and 17 are sprung by elastic elements 19 and 20. At the same time, one of ends of a plunger 17 is cone-shaped 21 with a bead 22 which is the support of a spring 19. The support 12 has channel 23 connected with hydroline 24 output cavity B of the case 8 with the pressure head hydroline 3 of high pressure located between channel W of nozzle 10 and hydraulic valve 5 intended for transportation of the locking device 7 forced out from an output cavity B, the volume of incompressible working liquid between end of the support 12 and a plunger 13 at the end of its course to the pressure head hydroline 3 of the output nozzle 10. With non-working hydraulic actuator, pressure of working liquid in cavities A, B, and W locking devices 7 is identical and equally nominal. At the same time, the plunger 13 of input cavity is propped up by a spring 20 and is located in extreme left position. The plunger 17 of the output cavity B is in an equilibrium state under the influence of pressure force on a plunger 17, created by pump 2 from a solid cavity D and output cavity B, at the same time cavities A and B are connected through axial channel 16 of plunger 13. When turning on the hydraulic valve 5 working liquid via the channel 16 of an input cavity A comes to an output cavity B and through channel W of output nozzle 10 under operating pressure and comes to hydraulic engine 6, which sets working body of the vehicle in motion. At the same time, pressure difference of working liquid in cavities of A and B is insignificant.

5 Conclusion Due to simplification and decrease in material consumption of a design of the locking device of the protection system of a hydraulic actuator its high operational reliability, functional purpose and environmental safety of the environment are provided.

References 1. Pyndak VI, Strokov VL, Lapynin YG et al (1999) Decrease in dynamic loads in hydroficated machines of cyclic action. Science Production. Mechanical engineering, Moscow, no. 10 2. Fomenko VN (2000) Development of protection systems of hydraulic actuators of mechanisms of hinged traction and special transport vehicles. Dissertation, Volgograd, p 166 3. Fedyakin VI, Shevchuk VP et al (2001) Assessment of reliability indicators of a hydraulic system of the machine and tractor unit. In: Agriculture problems in market conditions: materials of the international scientific and practical conference, Volgograd 4. Fomenko NA (2002) Improving the performance characteristic of hydraulic systems in machine-tractor aggregates. Dissertation, Volgograd

616

N. A. Fomenko et al.

5. Burlachenko OV, Serdobintsev YP, Skhirtladze AA (2010) Improvement of quality of functioning of processing equipment (monograph). TNT, Stary Oskol, 411 p 6. Fomenko NA, Tyrnov YA (2013) Research of operability of hydraulic system hoses of machine and tractor units/Increase in efficiency of use of resources by production of agricultural products—new technologies and equipment of new generation for crop production and livestock production: collection of scientific works. The XVII International scientific and training conf, 24–25 September 2013, Publishing house Pershin R.V., Tambov, pp 146–149 7. Fomenko NA, Bogdanov VI, Aleksikov (2014) Resource-saving hydraulic system of construction equipment. In: Collection of scientific articles based on the proceedings of the international scientific and practical conference. Published at Saratov State Technical University, Saratov, pp 221–224 8. Fomenko NA, Bogdanov VI, Sapozhkova NV (2014) Ways of improvement of a hydraulic actuator traction vehicles. Bulletin of the Volgograd state architectural and construction university. Ser: Constr Archit 36(55):218–222 9. Fomenko NA, Aleksikov SV, Bogdanov VI, Sapozhkova NV (2014) Pipeline of a hydraulic system of construction and road vehicles. Messenger Dev Sci Educ 3:115–117 10. Fomenko NA, Bogdanov VI, Burlachenko OV, Aleksikov SV (2015) Decrease in energy of water hammer in the locking device of a hydraulic system of the construction and road equipment. Internet Messenger of IACE of VSTU. Series: Polythematic 1(37). Access mode: www.vestnik.vgasu.ru 11. Fomenko NA, Bogdanov VI, Burlachenko OV, Aleksikov SV (2015) Increase in reliability of the locking device of protection system of a hydraulic actuator of construction and road vehicles (scientific article). Messenger of IACE of VSTU. Ser: Constr Archit 41(60):169–180 12. Fomenko NA, Bogdanov VI, Burlachenko OV, Aleksikov SV (2015) System of protection of a hydraulic actuator of construction and road vehicles (scientific article). Messenger of IACE of VSTU. Ser: Constr Archit 42(61):163–173 13. Fomenko NA, Dubinsky SV, Golobuta GI, Lyshko GP (1987) Protection system of a hydraulic actuator. Patent SU1813937A1F15B20/00 14. Fomenko NA, Perelmiter VI, Fomenko VN (2000) Protection system of a hydraulic actuator. Patent RU 15763U17F15B21/00 15. Fomenko VN, Perelmiter VI, Fomenko NA, Shevchuk VP (2000) Hydraulic system. Patent RU 15764 U1 7 F 15 B 21/00 16. Perelmiter VI (1990) Hydraulic system. Patent SU 1822471 A3 F 15B20/00 17. Gadzhiyev BA, Kirsch BA (1990) The device for elimination of leaks. A.C. No. 1576770 of the USSR 18. Bashta TM (1990) Emergency device. A.C. No. 1596146 of the USSR 19. Shevchuk VP, Bobkov YK et al (1991) Way of protection of a hydraulic actuator. A.C. No. 1661483 of the USSR 20. Manuylov VY, Yershov OB, Kurmambayev AE (1991) System of protection of a hydraulic actuator. A.C. No. 1605046 of the USSR 21. Omenko NA, Bogdanov VI, Burlachenko OV et al (2006) System of protection of a hydraulic actuator. Patent 2549754 C1 Russian Federation MPK F 15 B 20/00 22. Omenko NA, Bogdanov VI, Burlachenko OV et al (2006) System of protection of a hydraulic actuator. Patent 2556835 C1 Russian Federation MPK F 15 B 20/00 23. Fomenko NA, Bogdanov VI, Burlachenko OV et al Protection system of a hydraulic actuator. Patent 2571240 RU C1 F 15 B 20/00 24. Fomenko NA, Bogdanov VI, Fomenko VN (2006) Pipeline of high pressure. Patent RU 2511926 C2 Russian Federation MPK F 15 B 20/00 25. Fomenko NA, Burlachenko OV, Aleksikov SV, Fomenko VN Protection system of a hydraulic actuator. Patent RU 2,583,195 C1 F 15 B 20/00 26. Fomenko NA, Burlachenko OV, Aleksikov SV, Fomenko VN Protection system of a hydraulic actuator. Patent RU 2,642,719 C1 F 15 B 20/00

Influence of Hydrogen Additives on Cycle-to-Cycle Variability of Working Process of Rotary Engine Y. V. Levin, K. V. Prikhodkov and E. A. Fedyanov

Abstract A rotary engine, designed according to Wankel’s scheme, is the one of alternative options for traditional piston ones. However, the Wankel engine has incomplete combustion of an air–fuel mixture in the working chamber which prevents the widespread use of such engines. Hydrogen additive to the main air– fuel mixture helps to decrease the incompleteness of combustion in the volumes near an apex of the rotor. In this paper, the cycle-to-cycle variability of the Wankel engine fueled with additions of hydrogen to the main air–fuel mixture at different excess air ratios was experimentally investigated. The investigation was carried out on the rotary engine VAZ-311 (Russia) equipped with the injection system of hydrogen additive. In this study, the Wankel engine was operated on part loads at the frequency of eccentric shaft 2000 rpm and different excess air ratios varied from 1 to 1.3. The results of the study showed that adding hydrogen to the air–fuel mixture allows increasing the maximum pressure in the rotary engine working chamber and improving the stability of the combustion process. The addition of 5% hydrogen at the excess air ratio of 1.2 improves combustion stability on 59.8% on the mode of the averaged urban cycle.




Keywords Wankel engine Hydrogen Lean combustion Air–fuel mixtures





Gasoline
Cycle-to-cycle variability


1 Introduction A rotary engine (Wankel engine) has some advantages than piston engines, such as low mass and dimension, higher power-to-weight ratio [1–3]. In this regard, the main advantages of the Wankel engine allow us to consider this type of engine as a serious alternative to two-stroke gasoline engines for small boats [4], as well as four-stroke engines of light aviation [5]. The widespread use of the rotary engine in Y. V. Levin (&)
K. V. Prikhodkov
E. A. Fedyanov Volgograd State Technical University, Lenin Avenue, 28, Volgograd 400005, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_65

617

618

Y. V. Levin et al.

the automotive industry is hampered by the fact that these engines have slightly higher fuel consumption, and their exhaust gas contains an increased amount of unburned hydrocarbons [4, 6, 7]. The main cause of these deficiencies is incomplete combustion of air–fuel mixture due to one direction flow. Extended narrow shape of the Wankel engine’s combustion chamber promotes combustion incompleteness of fuel. The incomplete air–fuel mixture combustion in spark ignition engine is highly linked with cycle-to-cycle variability (CCV) [8–11]. These phenomena have long time identified as limiting factor in lean mixture combustion and as a consequence reducing emissions and fuel consumption. As the idling and partial loads modes are the main one for the engine to transportation vehicles in the urban, decreasing of CCV has actuality under these conditions. Causes and effects of CCV are well researched in piston engine [8–10]. As a rule, three major causes are highlights: fluctuation of air–fuel mixture into the cylinder, chaotic turbulent motion of mixture, and spark discharge variability. Undoubtedly, the features of the combustion in the Wankel engine must be taken into account for a complete picture of the CCV. A promising way reducing cycle-to-cycle variability is improved air–fuel mixture burning by addition of promotion gas, hydrogen for example [9, 11, 12]. Recently, hydrogen is considered an alternative for conventional fuels used for transportation sector. Hydrogen is the most energy-intensive fuel. Its combustion heat is 120 MJ/kg, which is 2–3 times higher than for natural (methane) or liquid petroleum gases (LPG), gasoline, and diesel fuel [13–16]. At the same time, running vehicle engine on pure hydrogen creates a complex problem of onboard storage [17, 18], so it is of considerable interest to make relatively small additives of hydrogen to main hydrocarbon fuels for increase flame speed and improve completeness of combustion. The use of minor additives (up to 5% by weight) to the main fuel does not require significant changes in feeding systems and seems to be a more promising approach to the hydrogen using [19–21]. The Wankel engine to a greater extent is adapted to work on hydrogen than the piston engine [1, 4, 19]. The absence of exhaust valve in the design of rotary engine almost completely eliminates the possibility of a reverse flash. Due to the fact that the intake and compression processes in the Wankel engine occur in the stator zone with a low temperature, and the spark plugs are located in special pre-chambers, the probability of pre-ignition of the air–fuel mixture with hydrogen additive is also significantly less than in a piston engine. Moreover for an engine of this type, an increase in the flame propagation speed obtained by hydrogen addition reduces incomplete burning in the vicinity of rotor’s rear apex [19, 21]. Thus, experimental studies of the effect of hydrogen additive on the completeness of combustion of the air–fuel mixture in a rotary engine were carried out.

Influence of Hydrogen Additives on Cycle-to-Cycle …

619

2 Experimental Procedure 2.1

Experimental Setup

Experimental studies of the operation rotary engine with the addition of hydrogen to the air–fuel mixture were carried out under test bench conditions. The test bench was equipped with a control instrumentation, which allows recording the magnitude of the torque, rotational speed of the eccentric shaft, air and fuel flow rates, and other parameters of the rotary engine. Single-section Wankel engine VAZ-311 is used for the experiments. The engine on the stand was equipped with regular lubrication systems, cooling, electrical equipment, etc. The rotary engine specifications are presented in Table 1. Indication of the rotary engine is a difficult task. In order to obtain a complete picture of the pressure change in the working chamber of the Wankel engine, it is necessary to use at least three pressure sensors located in different parts of the stator. However, to study the effect of hydrogen additive on the combustion process in the rotary engine, it is sufficient to use a single pressure sensor installed in the immediate vicinity of top dead center (TDC). In this study, the uncooled piezoelectric sensor “Kistler” (type 6118B) was used to indication of the Wankel engine. The signal from the sensor “Kistler” goes to the amplifier unit. After amplification, the signal from the pressure sensor enters one channel of the analog-to-digital converter (ADC), while the other channels of the ADC receive signals from the TDC sensor and spark timing sensor at the top spark plug (Fig. 1).

2.2

Test Conditions

For experimental research, the Wankel engine was equipped with gasoline and hydrogen injection equipment. The nozzles were installed in the intake manifold in the immediate vicinity of the inlet window. The control of the nozzle supply made it possible to independently change the beginning and duration of the supply of gasoline and hydrogen. The duration of hydrogen injection was calculated based on the accepted mass fraction of hydrogen in the mixture with the main fuel:

Table 1 Wankel engine VAZ-311 technical specifications

Engine manufacturer

Volga Automobile Plant, Russia

Ignition source Displacement (cm3) Compression ratio Maximum power output Maximum torque

Top and low spark plugs 649 9.3 52 kW at 6000 rpm 95 Nm at 4000 rpm

620

Y. V. Levin et al.

Fig. 1 Example of recording signals from sensors of the engine VAZ-311: 1—pressure in the chamber of the rotary engine; 2—TDC; 3—spark timing at top spark plug

G2 ¼

gH2
Gg 1  gH 2

ð1Þ

Equation (1) gH2 is the mass fraction of hydrogen; Gg is gasoline mass flow rate (kg/s). The duration of gasoline injection was determined on the basis of the air flow rate in the studied operating mode of the rotary engine based on the adopted excess air ratio: Gg ¼


Gair

gH2 1gH2


AFH2 þ AFg
k

ð2Þ

In Eq. (2), Gair is air mass flow rate (kg/s); k is excess air ratio; AFH2 and AFg are the stoichiometric air/fuel ratios of pure hydrogen and gasoline, respectively. Researches of influence hydrogen additives on the CCV level in the Wankel engine were performed for the partial loads (2000 rpm) with the effective pressure of 0.16 and 0.2 MPa. To obtain quantitative information about the CCV, pressure traces were processed in at least 30 consecutive cycles. As a quantitative measure of the CCV, the value of the standard deviation of the maximum pressure was chosen: sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi Pn max 2  pmax Þ i ð pi rpz ¼ n

ð3Þ

Influence of Hydrogen Additives on Cycle-to-Cycle …

621

Equation (3) pmax is the maximum pressure of the cycle (MPa); pmax is the mean i value of maximum pressure of the n consecutive cycles (MPa); n is number of the consecutive cycles. It is well known [22] that the value of the standard deviation of the maximum pressure in the cycle sequence strong correlates with the value of the standard deviation of the mean indicator pressure. Influence of hydrogen additives on CCV of rotary engine during its operation on the lean air–fuel mixture was studied on the mode of the averaged urban cycle (n = 2000 rpm, pe = 0.2 MPa). The depletion of the air–fuel mixture was achieved by reducing the supply of hydrogen and gasoline, respectively. Throttle position has not changed. Ignition timing on the upper spark plug was set to 30° of rotation of the eccentric shaft angle before TDC, on the lower spark plug corresponds to 26°. The moment of ignition for both spark plugs during the experiments did not change with the addition of hydrogen and the variation of the excess air ration.

3 Results and Discussion Experimental pressure–time diagrams got from the VAZ-311 bench on a lean air– fuel mixture for 0, 3, and 5% hydrogen fractions (Fig. 2). Figure 2 shows the combined pressure trace in 10 sequential cycles. It clearly illustrates the phenomenon of CCV in the Wankel engine. It could be noted that the combustion CCV is inherent for the Wankel engine as well as for the piston one and takes place in all the studied modes. Similar to piston engines, lean operation leads to growth combustion instability. As can be seen from Fig. 2, the addition of hydrogen to the fuel can increase both the uniformity of the combustion and the maximum pressure of the cycle. Increased amount of hydrogen leads to a shift of the maximum pressure closer to TDC. The earlier position of the point of onset of combustion is explained by the reduction in the forging duration of the initial hearth of burning due to the higher reactivity of hydrogen. The displacement of the point of maximum cycle pressure toward TDC is the result of not only a reduction in the duration of the formation of the initial hearth of burning, but also an increase in the rate of flame propagation [15, 22]. This positive effect is observed both on the lean air–fuel mixtures and on stoichiometric ones. Figure 3a presents relation of the standard deviation with the value of the hydrogen additive for two operation modes for the eccentric shaft speed 2000 rpm on stoichiometric fuel–air mixture. It can be seen, if hydrogen additive is increased, then CCV of combustion is reduced: 37.5% at pe = 0.2 MPa and 30% at pe = 0.14 MPa with the same hydrogen additive. Research of influence hydrogen additives on the CCV level in the Wankel engine was also carried out on lean air–fuel mixtures. As shown in Fig. 3b, CCV of

622

Y. V. Levin et al.

Fig. 2 Indicator diagrams of a rotary engine in the operating mode n = 2000 rpm; pe = 0.2 MPa

Fig. 3 Variation of standard deviation in maximum pressure with hydrogen additives at n = 2000 rpm; a stoichiometric air–fuel mixture; b different excess air ratios varied from 1 to 1.3

burning lean air–fuel mixture increases with a significantly lower rate for air–fuel mixture with additives of hydrogen than without it. The addition of 3% hydrogen increases the standard deviation of the maximum pressure with the rate of 0.02 for every 0.1 units of the excess air ratio and an additive of 5%—with the rate of 0.013

Influence of Hydrogen Additives on Cycle-to-Cycle …

623

for the same units’ air–fuel ratio. For example, operation mode under stoichiometric air–fuel mixture with 5% hydrogen addiction improves combustion stability on 26.9%. Same hydrogen addition under lean conditions with relative air–fuel ratio 1.2 improves combustion stability on 59.8%. It can be concluded that the minor additions of hydrogen to the main fuel can help to achieve a significant reduction of the CCV on wide range engine mode including lean conditions. It is shown that a mixture of gasoline and 5% hydrogen allows achieving stable operation with air equivalence ratio 1.3.

4 Conclusions The results of the study make it possible to state that hydrogen addition to the air– fuel mixture allows increasing the maximum pressure in the rotary engine working chamber and improving the stability of the combustion process. An additive of 5% by mass hydrogen improves combustion stability by 26.9% when the rotary engine is operating on a stoichiometric air–fuel mixture. The use of hydrogen additives also allows to ensure stable operation of the rotary engine in the lean air–fuel mixture under partial load conditions. The addition of 5% hydrogen at the excess air ratio of 1.2 improves combustion stability on 59.8% on the mode of the averaged urban cycle.

References 1. Su T et al (2018) Improving the combustion performance of a gasoline rotary engine by hydrogen enrichment at various conditions. Int J of Hydrogen Energy. https://doi.org/10. 1016/j.ijhydene.2017.11.175 2. Chen H et al (2015) Development of rotary piston engine worldwide. In: AASRI international conference on industrial electronics and applications. https://doi.org/10.2495/feem140201 3. Wang W, Zuo Z, Liu J (2016) Miniaturization limitations of rotary internal combustion engines. Energy Convers Manag. https://doi.org/10.1016/j.enconman.2016.01.002 4. Amrouche F et al (2016) Extending the lean operation limit of a gasoline Wankel rotary engine using hydrogen enrichment. Int J Hydrogen Energy. https://doi.org/10.1016/j.ijhydene. 2016.06.250 5. Boretti A (2015) Modeling unmanned aerial vehicle jet ignition Wankel engines with CAE/ CFD. Adv Aircraft Spacecraft Sci. https://doi.org/10.12989/aas.2015.2.4.445 6. Amrouche F et al (2014) An experimental investigation of hydrogen-enriched gasoline in a Wankel rotary engine. Int J Hydrogen Energy. https://doi.org/10.1016/j.ijhydene.2014.03.172 7. Wakayama N et al (2006) Development of hydrogen rotary engine vehicle. In: 16th World Hydrogen Energy Conference, Lyon, France 8. Masouleh MG et al (2018) Flow and thermal field effects on cycle-to-cycle variation of combustion: scale-resolving simulation in a spark ignited simplified engine configuration. Appl Energy. https://doi.org/10.1016/j.apenergy.2018.08.046

624

Y. V. Levin et al.

9. Yu X et al (2016) Research on cycle-by-cycle variations of an SI engine with hydrogen direct injection under lean burn conditions. Appl Therm Eng. https://doi.org/10.1016/j.applthermaleng. 2016.08.077 10. Kyrtatos P, Brückner C, Boulouchos K (2016) Cycle-to-cycle variations in diesel engines. Appl energy. https://doi.org/10.1016/j.apenergy.2016.03.015 11. Su T et al (2017) Reducing cyclic variation of a gasoline rotary engine by hydrogen addition under various operating conditions. Int J of Hydrogen Energy. https://doi.org/10.1016/j. ijhydene.2017.08.158 12. Reyes-Ramírez I (2018) Symbolic Analysis of the cycle-to-cycle variability of a gasoline– hydrogen fueled spark engine model. Energies. https://doi.org/10.3390/en11040968 13. Verhelst S, Wallner T (2009) Hydrogen-fueled internal combustion engines. J Prog Energy Combust Sci. https://doi.org/10.1016/j.pecs.2009.08.001 14. Singh AP et al (2018) Prospects of Alternative transportation fuels. Springer, Singapore 15. Shivaprasad KV et al (2014) Experimental investigation of the effect of hydrogen addition on combustion performance and emissions characteristics of a spark ignition high speed gasoline engine. J Proc Technol. https://doi.org/10.1016/j.protcy.2014.08.019 16. Gelfand BE et al (2012) Thermo-gas dynamics of hydrogen combustion and explosion. Springer Science & Business Media. https://doi.org/10.1007/978-3-642-25352-2 17. Cipriani G et al (2014) Perspective on hydrogen energy carrier and its automotive applications. Int J Hydrogen Energy. https://doi.org/10.1016/j.ijhydene.2014.03.174 18. Basile A, Iulianelli A (2014) Advances in hydrogen production, storage and distribution. Elsevier, Amsterdam 19. Fedyanov EA et al (2017) Modelling of flame propagation in the gasoline fuelled Wankel rotary engine with hydrogen additives. IOP Conf Ser: Mater Sci Eng. https://doi.org/10.1088/ 1757-899x/177/1/012076 20. Su T et al (2017) Investigation on performance of a hydrogen-gasoline rotary engine at part load and lean conditions. J Appl Energy. https://doi.org/10.1016/j.apenergy.2017.08.049 21. Itkis EM, Fedyanov EA, Levin YV (2019) Experimental and numerical investigation of influence of hydrogen addition to hydrocarbon fuel on wankel rotary engine performance. Ser Nat 22. Glassman I, Yetter RA, Glumac NG (2014) Combustion. Academic press, Cambridge

Efficiency of Usage of Transport and Technological Machines G. V. Redreev, G. A. Okunev and S. A. Voinash

Abstract The efficiency of the usage of transport and technological machines is determined by the number of products produced per unit of funds expended with this. With conditionally constant production volumes, the task of increasing efficiency is reduced to minimizing costs. For production facilities, the volumes of which are strictly related to the calendar delivery time, it is necessary to take into account the risks of reducing production volumes in case of unplanned downtime of technological machines. For machines with a high load, it is also necessary to take into account equipment downtime for maintenance, planned, and unplanned. The above-mentioned type of production includes the production of crop products by agricultural enterprises. The efficiency of production depends on the efficiency of the usage of technological machines during the entire season of the fieldwork— from tillage in spring to harvesting in autumn. In the season of fieldwork, it is customary to distinguish spring, summer, and autumn cycles of work. Each fieldwork cycle consists of a set of field operations, in which any simple technological machines lead to a shortage of agricultural products. We have developed a mathematical model to optimize the use of machine units based on the cost of technical service and downtime losses during the implementation of technical service. For field operations, a variant of a mathematical model is presented that takes into account the loading of energy resources of machine units. Keywords Technological machines Mathematical model


Efficiency of usage
Technical service


G. V. Redreev (&) Omsk State Agrarian University by P.A. Stolypin, 1, Institutskaya square, Omsk 644008, Russia e-mail: [email protected] G. A. Okunev South-Ural State Agrarian University, 75, Lenin Ave., Chelyabinsk 454080, Russia S. A. Voinash Rubtsovsk Industrial Institute (Branch), Altai State Technical University, I.I. Polzunov of, 2/6 Traktornaya str., Rubtsovsk 658207, Russia © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_66

625

626

G. V. Redreev et al.

1 Introduction From the experience of the organization of production processes in industry, effective directions of their functioning are known, based on flow methods, concentration of production, division, and specialization of labor. In contrast to industry, mechanized processes in agriculture are associated with the flora and fauna, with the seasons of the agricultural year, with the movement of machine complexes over large areas. The agrotechnical requirements for the cultivation of agricultural crops, the state of plants, soil and weather conditions determine the modes of operation of the machines and the calendar dates of work. In mass production, the material moves from one group of machines to another. In agriculture, in-line production is more complex, as the units are consistently moved in space [1, 2].

2 The Cost of Technological Machines’ Operation The expression of the total cost function for the operation of technological complexes for the season of fieldwork will have the form [3]: CTC ¼

M X m¼1

CCc t

m1

bm CCt þ

M 1 X m¼0

CCc t

m1

; rub:

ð1Þ

total costs for the implementation of the maintenance process in inter-frame time: CCc t

m1

    t K3 K1 þ K2 þ ¼D þ K4 ðK5 þ K6 Þ ; rub: P t

ð2Þ

under the assumption that the idleness of the machine aggregate does not lead to losses from crop shortages, where D is the duration of fieldwork (operation), hour; t P K1 K2

the duration of maintenance, hour; the frequency of maintenance, hour; the ratio taking into account the maintenance costs, rub./h.; the ratio of operating costs during downtime for maintenance, rub./h.; K3 the cost factor for the movement of mobile maintenance, rub.; K4 = u(t, P) the coefficient taking into account the downtime of the machine unit to eliminate the consequences of failures, hour−1;

Efficiency of Usage of Transport and Technological Machines

627

K5 = x(t, P) the cost accounting factor for eliminating the consequences of failures, rub.; K6 the factor accounting for operating costs from idle machine unit to eliminate the effects of failures, rub. Minimizing the functions of total costs, based on the season of fieldwork, makes it possible to determine the optimal parameters of the process of technical maintenance of machine and tractor units and complexes, and taking into account the importance of each cycle in the field season [4]. Thus, the purpose of the operation of aggregates and complexes determines the choice of technologies for maintenance and for performers of maintenance. If you change the planned volume of fieldwork, the requirements for the operability of the units will also change. In this case, two approaches can be distinguished: at the beginning of the period of operation of the units, ensuring its maximum possible operability with available resources or ensuring the required operability of the units while minimizing the required resources. The choice of approach also influences the selection of technologies for maintenance and for performers of maintenance. For the fieldwork cycle, the total cost function will be: CCt ¼

B X

al1 Ctl ; rub:

ð3Þ

l¼1

where al-1 is the factor of accounting for the decrease in the efficiency of the units with an increase in its total operating time during the fieldwork cycle, a0 = 1, Ct

total costs and losses during maintenance of the unit during the l field operation; l = 1, 2, …, B the number of field operations performed by the machine unit for the cycle of fieldwork. The study of the function of total costs and losses for the fieldwork cycle also allows you to determine the optimal loading of the unit during fieldwork cycle.

2.1

Cost Minimization with Regard to Product Losses

The mathematical expression of the total cost of maintenance and repair and losses from crop shortages during downtime of technological machines for maintenance and the elimination of the consequences of failures can be represented as follows:   t K7 DKCL t K3 þ Þ þ K4 ðK5 þ K6 þ K7 K4 DKCL Þ ; rub: ð4Þ Ct ¼ D ðK1 þ K2 þ P P t

628

G. V. Redreev et al.

where К7 КCL

the tractor performance factor, ha/hour; the generalized coefficient of specific yield losses, rubles/ha.

In the expanded form, the expression for the total cost function will have the following form [4, 5]:

ð5Þ where ccm w g N KN coc КCL Tsd ssf R ctr mmv ctsw Ell tmd Кtt, КtP

ccf Кct, КcP

the unit costs for maintenance, the ruble is on a conditional reference hectare (c.r.ha); the reference tractor performance, c.r.ha/h; the share of wages in the cost structure of maintenance; the number of mobile managers, people; the reduction ratio of the duration of maintenance when servicing the machine unit with several performers of maintenance; the operating costs, rub./c.r.ha; the generalized unit loss rate from aggregate downtime; the shift duration, hour; the shift factor; the mobile unit distance, km; the transport costs, rubl./km; the number of mobile vehicles, pcs.; the tariff rate of a service worker, rub./h.; the efficiency of living labor, rub.; the minimum level of downtime, h/c.r.ha; the coefficients that take into account the change in the specific downtime of the unit depending on the change in the duration (frequency) of maintenance; the minimum level of unit costs for elimination of the consequences of failures, rub./ c.r.ha; the coefficient taking into account the change in the unit cost of eliminating the consequences of unit failures depending on the change in the duration (frequency) of maintenance.

The shape of the region of optimal values of P and t is determined by the features of the mathematical model; it is similar for tractors of various brands, however, occupying different positions on the plane with the coordinates «P − t» (see Fig. 1).

Efficiency of Usage of Transport and Technological Machines

629

Fig. 1 Ratio of the optimal values of P and t 1—wheel tractor 1,4 tf; 2–3,0 crawler tractor 3,0 tf; 3—crawler tractor 4,0 tf; 4—combined schedule

Combining the optimal zones of the “P – t” parameters of several brands of tractors allows us to construct a combined region of optimal values. This will allow us to determine the overall optimal frequency of maintenance, which is important for specialized technical service by mobile teams or units [5, 6]. Duration of maintenance will be individual; it is determined by the organization of technical service, equipment of tractors means of monitoring the technical condition. However, in the presence of agricultural units with different lifetimes and different technical conditions in the fleet structure of the machine park, it is necessary to foresee such a load on the engines that would ensure the planned duration of the trouble-free operation period [7].

2.2

Cost Minimization Taking into Account the Loading of Technological Machines’ Engines

We introduce into our previously developed mathematical model the load factor Kz, taking into account the change in the total cost of technical service, depending on the load of the tractor engine [8, 9]:

ð6Þ where mnr—number of rides. Taking the first derivative of the expression (6) and equating it to zero, we define the expression for the value of the load factor KLF:

630

G. V. Redreev et al.

ð7Þ Changing the values of the components of the right part of the expression (7) will determine the value of short-circuit corresponding to the minimum total cost of technical service. When increasing the value of operating costs, the engine load must be increased (Fig. 2). The degree of increase depends on the duration of the tractor maintenance time. The longer the maintenance time, the more the engine can be loaded. While carrying out maintenance by a centralized method, with field service of 3–4 people, the duration of maintenance in this case determines the completeness of the maintenance operations.

Fig. 2 Dependence of the load factor of the engine from the value of the specific operating costs of the ccm at different values of the duration of maintenance (the surface in the coordinates КLF = f (ccm, t))

Fig. 3 Dependence of the engine load factor on the magnitude of the maintenance time t at various values of the maintenance frequency P (cross section of the load factor at КLF = 0.8, КLF = 0.6, and КLF = 0.4)

Efficiency of Usage of Transport and Technological Machines

631

The analysis of the topographic lines of the surface of the optimal values of the engine load factor reveals its dependence on the amount of maintenance work determined by the duration of maintenance and its periodicity with the averaged values of the other components of the radix in relation (7) (see Fig. 3). In accordance with the expression (7), the engineering services of agricultural enterprises can, based on the performance of the used units, unit costs for their use, the established level of technical service costs, determine the level of tractor engine load corresponding to the minimum amount of technical service costs and the cost of losses from the shortage, focusing on the established values of the frequency and duration of maintenance [10].

3 Conclusion Mathematical modeling of the process of using technological machines and units allows to determine the optimal parameters of technical operation of technological complexes, in which the amount of maintenance costs and elimination of the consequences of failures and losses for the time spent will be minimal. This will ensure the efficient use of technological machines in specific production conditions of agricultural enterprises.

References 1. Redreev GV, Okunev GA (2018) Ensuring the efficiency of technology in the implementation of flow processes in agriculture. N Bull Omsk Atate Agrarian Univ 30:125–132 2. Okunev GA (1988) Flow-cyclic technology of harvesting grain crops: studies. ChSAU, Chelyabinsk 3. Redreev GV, Okunev GA (2015) To the question of the goals achieved in ensuring the operability of machine-tractor units. N Fundam research 6–2:282–286 4. Astafev VL, Ivanchenko PG, Okunev GA et al (2003) Seasonality in agriculture and technical re-equipment of production. N Agric 1:32–33 5. Redreev GV (1997) Justification of the process of maintenance of tractors by a group of performers during field work. Dissertation, State Agroengineering University of Chelyabinsk 6. Redreev GV, Myalo OV, Prokopov SP et al (2017) Machine-tractor aggregates operation assurance by mobile maintenance teams. N IOP Conf Ser: Mater Sci Eng 221:012016 7. Redreev GV (2015) Justification of the concept of ensuring the operability of machine-tractor units. In: Modern problems of science and education: https://www.science-education.ru/ru/ article/view?id=22971. Accessed 25 Nov 2018 8. Shepelev SD, Cherkasov YuB (2015) Justification of the rational level of reliability of technological machines in the harvesting process. N Bull Krasnoyarsk State Agrarian Univ 104:58–63 9. Shepelev SD, Shepelev VD, Cherkasov YuB (2015) Differentiation of the seasonal loading of combine harvester depending on its technical Readiness. N Proc Eng (see in books) 129:161– 165 10. Kirtbaya YuK (1982) Reserves in the use of machine-tractor park. Kolos, Moscow

Differential System of Crane Braking S. V. Streltsov and V. A. Ryzhikov

Abstract The brake system of the crane is one of the main systems responsible for its safety. Due to the presence of a number of factors, such as unequal braking force of the brake spring, uneven wearout of the brake pads, and the presence of pollutants, the friction force is variable and varies widely. As a consequence, there is a different braking torque on the drive wheels of the crane. It leads to the skidding of one of the sides of the crane, the spreading of the drive wheels with the emphasis of the flanges on the crane rail and the skewing of the entire metal structure. There are additional loads in the metal, which reduces the reliability of the crane in the work. The principle of stabilization of braking forces by differential mechanisms on the crane wheels is being developed. The proposed solution provides the same braking force on the wheels with different modes of braking the crane. The kinematic scheme is based on the principle of distribution of power flows between the hydraulic differential, installed on the carts. The design of the hydraulic brake system of the crane is considered. A mathematical model of the differential brake system of a crane and the results of its computer simulation in the Mathcad environment are given. According to the results of mathematical analysis, the parameters of the brake system are determined. Keywords Brake Model Analysis





Differential
Crane
Dynamic loads
Mathematical


S. V. Streltsov Platov South-Russian State Polytechnic University (NPI), 132, St, Prosvesheniya, Rostov Region, Novocherkassk 346428, Russia V. A. Ryzhikov (&) Don State Technical University, 1, Gagarin Square, Rostov on Don 344010, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_67

633

634

S. V. Streltsov and V. A. Ryzhikov

1 Introduction The brake system of the crane is one of the main systems responsible for its safety [1–13]. The brake system of the crane consists of a brake mechanism and a drive. Cranes with a large span have an individual drive mounted on each side of the crane girder for bridge cranes and on the supports for gantry cranes. Due to the presence of a number of factors, such as unequal braking force of the brake spring, uneven wearout of the brake pads, and the presence of pollutants, the friction force is variable and varies widely. As a consequence, there is a different braking torque on the drive wheels of the crane. It leads to the skidding of one of the sides of the crane, the spreading of the drive wheels with the emphasis of the flanges on the crane rail and the skewing of the entire metal structure. There are additional loads in the metal, which reduces the reliability of the crane in the work [14, 15].

2 Differential Brake System To create the same braking force on the traveling wheels of the cranes with a separate drive and increase its efficiency, a differential hydraulic braking system has been developed, shown in Fig. 1 [16]. The shafts of the electric motor 1 through the reducer 2 are kinematically connected to the driving wheels 3. The running wheels 3 are in turn kinematically connected through the multiplier 4 with the corresponding hydraulic pumps 5. The hydraulic distributor 6 with electromagnetic control connects the adjustable throttle 10 to the hydraulic pumps 5 in one position and a hydraulic system with an adjustable throttle 7 and a safety valve 9. In turn, an adjustable throttle 7 is connected to the accumulator 8. The throttle 10 has a mechanical connection with the pedal 11, controlling its flow section. Check valves are installed in the drain Fig. 1 Hydrokinematic scheme of the differential system of braking crane

Differential System of Crane Braking

635

hydrolines of hydraulic pumps 13. Drain hydrolines of hydraulic pumps 12 are connected to the oil tank 14. When the brake system is turned off, under the action of rotation of the engine, the fluid from the hydraulic pump 5 is injected to the hydraulic distributor 6 and passes through the throttle 10, which is in the open state. There is no pressure in the brake system, and a crane without resistance moves along the track. When braking, the drive motor turns off and the driver, pressing the pedal, controls the process of braking the crane. The magnitude of the braking torque in this case is proportional to the force pressing the pedal and can vary within wide limits. When the operator clicks the pedal 11, the flow area of the throttle 10 decreases, increasing the pressure in the hydraulic system. Check valves 13 prevent leakage of the working fluid through the hydraulic pumps in the oil tank, maintaining the necessary pressure in the system. Due to the parallel connection of hydraulic pumps, the pressure of the working fluid in the pressure hydraulic lines is the same. Thus, the braking torque on the driving wheels will also be the same. When the limit switches are triggered or when the power supply is interrupted, the drive unit de-energizes and emergency braking begins. In this case, the valve 6 connects the pressure line of the hydraulic pumps 5 to the adjustable throttle 7 and the safety valve 9. The pressure in the hydraulic system increases with the braking torque to the value set by the safety valve 9 adjusted for the maximum braking. In order to avoid slipping of crane wheels with allowance for the permissible dynamic loads during braking, the rate of increase in the braking moment is selected so that it does not exceed the allowable value of the acceleration of the crane during braking [17]. The rate of rise of the braking torque can be adjusted by changing the flow area of the throttle 7. To ensure a smooth braking process, if necessary, you can adjust the formation time of the maximum braking force by varying the capacity of the accumulator 8.

3 Mathematical Model A crane with a differential braking system is a multi-mass system, the links of which are interconnected by rigid connections. Studies of the dynamics of a crane with a differential braking system are aimed at determining the dynamic loads in the supports of the crane during braking. It also becomes possible to study the operation of the system in transient conditions. The main masses in the system are: m 1, m 2 m3 m 4, m 5 m6

are the reduced masses of driven carts; is the mass of the beam or bridge; are the reduced masses of the non-driven carts; is the mass of the load.

636

S. V. Streltsov and V. A. Ryzhikov

In accordance with this, a crane with a braking system can be represented as a six mass system, with generalized coordinates x1, x2, x3, x4, x5, x6. where x 1, x 2 x3 x 4, x 5 x6

movement of supports with driven carriages; movement of a beam or bridge; moving supports with non-driven carts; movement of cargo.

The weight of the crane is taken as a link of reduction. When developing a mathematical model of the crane braking system, the following assumptions were made: 1. 2. 3. 4. 5. 6. 7. 8.

The system has holonomic, ideal two-way communications. There is no energy dissipation in the system. The metal structure of the crane has a constant mass. During the periods of start-up and braking of the crane, consider the transverse connections of the wheels with the rails as ideal. The crane runway is straightforward. The lateral compliance of the crane runway is not taken into account. Rail joints are taken perfectly flat and do not create shock loads. Elastic properties of supports and lifting rope are taken into account. The remaining elements are taken rigid. The given design diagram of the braking system is shown in Fig. 2 [18]. The kinetic energy of the system is: 1 1 1 1 1 1 TC ¼ m1 x_ 21 þ m2 x_ 22 þ m3 x_ 23 þ m4 x_ 24 þ m5 x_ 25 þ m6 x_ 26 ; 2 2 2 2 2 2

ð1Þ

According to [1], the reduced masses of carts are determined: mi ¼ mci þ mri

Fig. 2 Given the calculated scheme of the braking system

ð2Þ

Differential System of Crane Braking

637

where mci mri

is the static component of the mass of the cart; the mass of the rotating elements of the cart.

Static components of the mass of driven carts: mc1 ¼ md1 þ mp þ mw
n þ mb þ mh1

ð3Þ

mc2 ¼ md2 þ mp þ mw
n þ mb þ mh2

ð4Þ

where md1, md2 mp mw mb mh1, mh2 n

the mass of the corresponding engine; gearbox weight; wheel weight; mass multiplier; the mass of the corresponding hydraulic pump; number of wheels of one bogie.

Static components of the mass of non-driven carts: mc4 ¼ mw
n; mc5 ¼ mw
n

ð5Þ

Given the mass of the rotating elements of the driven carts: m1 ¼

Id1 2 Iw Ih1 i þ n þ 2 i2b R2w p R2w Rw

ð6Þ

m2 ¼

Id2 2 Iw Ih2 ip þ 2 n þ 2 i2b 2 Rw Rw Rw

ð7Þ

where Id1, Id2 moment of inertia of the corresponding engine: Idi ¼ GD2 ip ib Iw Rw Iг

GD2 4

flywheel moment; gear ratio; gear ratio multiplier; the moment of inertia of the running wheel; radius of the running wheel; moment of inertia of the hydraulic pump.

ð8Þ

638

S. V. Streltsov and V. A. Ryzhikov

Given the mass of the rotating elements of the non-driven carts: m4 ¼ m5 ¼

Iw n R2w

ð9Þ

Potential energy of the system: P¼

i 1h cðx3  x1 Þ2 þ cðx3  x2 Þ2 þ cðx3  x4 Þ2 þ cðx3  x5 Þ2 þ cl ðx3  x6 Þ2 2 ð10Þ

where c cl

support stiffness; suspension rigidity; cl ¼

m6 g l

ð11Þ

Here l g

is the length of the rope; acceleration of gravity.

The displacement in the supports is determined by the ratio of the forces applied to the support F to the support rigidity. Using the differential equation of the elastic line known from the resistance of materials, we obtain the rigidity support: c¼

3EI l3

ð12Þ

where L E I

support length; Young’s modulus; moment of inertia of the sectional area.

Taking into account the accepted assumptions, the braking device operation is described by the system of Lagrange second-kind differential equations [18], which can be written the following way: 9 m1€x1  cðx3  x1 Þ ¼ Wg  WC1 > > > > > m2€x2  cðx3  x2 Þ ¼ Wg  WC2 > > > > m3€x3 þ cðx3  x1 Þ þ cðx3  x2 Þ þ cðx3  x4 Þ þ cðx3  x5 Þ þ cl ðx3  x6 Þ ¼ 0 = ð13Þ m4€x4  cðx3  x4 Þ ¼ WC4 > > > > > m5€x5  cðx3  x5 Þ ¼ WC5 > > > > m €x  c ðx  x Þ ¼ 0 ; 6 6

l

3

6

Differential System of Crane Braking

639

Generalized resistance forces operating in the system: Wci—static resistance to movement of the corresponding wheel [17]: WCi ¼ Gi

lr dr þ 2f kp Dw

ð14Þ

where µr dr f kp Dw Gi

coefficient of friction of bearings, reduced to the wheel axle; pin diameter; rolling friction coefficient; coefficient taking into account the friction of the flanges and wheel hubs; wheel diameter; the weight of the crane with the load falling on the corresponding wheel.

For differential brake systems, the basic equation characterizing the energy and force relations is the power balance equation [19]. In this case, the braking force is determined by the formula: P q2h ni¼1 xi Wg ¼ 2 4p gib Rw ðCd þ cg þ V=vÞ

ð15Þ

where Cd Cg V v qh xi η h

throttle flow rate; leakage rate of the ith hydraulic pump; volume of working fluid in the hydraulic system; coefficient of elasticity of the hydraulic system; volume constant of the hydraulic pump; rotation frequency of the hydraulic pump shaft; efficiency multiplier; number of hydraulic pumps.

With the help of the system of differential Eq. (6), it is possible to simulate various modes of braking the crane and determine the load in the elastic links.

4 Results of Computer Modeling The system of differential equations of motion of the crane can be solved using the Mathcad system of mathematical calculations using various functions [20]. For the gantry crane KK-12.5-32, a study of transient processes in the system during braking was conducted. As initial conditions can be given: t ¼ 0; v1 ¼ v2 ¼ v3 ¼ v4 ¼ v5 ¼ v6 ; a1 ¼ a2 ¼ a3 ¼ a4 ¼ a5 ¼ a6 ¼ 0.

640

S. V. Streltsov and V. A. Ryzhikov

In the simulation, braking of a crane with a shoe brake was considered with an uneven braking force of 50% and with a differential hydraulic braking system during working and emergency braking. To study the process of braking a crane equipped with a developed braking system, a situation was simulated in which the load was in the extreme position near one of the supports and when the load was in the center of the span. When the load is positioned near the left support, static resistance to movement on the left side of the KK-12.5-32 crane will be WC1 = WC3 = 465.7 kN, on the right side— WC2 = WC4 = 360.9 kN. As a result of modeling, graphs of changes in displacement, velocity, acceleration, and efforts in the supports are presented in Figs. 3 and 4. The results of the simulation of emergency braking are shown in Fig. 4b. In this case, a constant braking force is created by the safety valve. The rate of increase in the braking force is regulated by changing the filling time of the accumulator. In modeling, the time of filling of the accumulator of the th = 1, th = 2 and th = 3 s. As studies have shown, the use of the developed braking system allows to obtain equal braking forces on the crane wheels. In this case, equal speeds contribute to the equality of movements of the crane supports, which eliminates the formation of a skew. Even with the location of the load in the leftmost position, causing the difference in resistance to movement, the difference between the movements of the right and left crane supports does not exceed 2%. The use of pad brakes does not allow for equal speeds of the traveling wheels during braking, as a result of which a large difference in the displacements of the supports arises, exceeding 10% and causing a distortion of the metal structure. In this case, the acceleration of the supports of the crane may exceed the allowable. By changing the filling time of the accumulator, it is possible to reduce the acceleration value of the crane without a significant increase in the braking time. Also controlling the time of filling the accumulator, you can change the amount of effort in the supports. Thus, modeling using the Mathcad software package confirms the effectiveness of using the differential crane braking system.

Fig. 3 a Changes in the movement of the supports of the non-driven carts of the crane with the using a pad brake; b Change in the speed of the supports of the non-driven of the crane with the using a pad brake

Differential System of Crane Braking

641

Fig. 4 a Change in the acceleration of the supports of the non-driven of the crane with the using a brake pad; b Influence of the filling time of the accumulator th on the braking force of the system during emergency braking

5 Conclusion The proposed differential braking system, providing the same braking forces on the driving wheels of the crane. A mathematical model of the differential brake system of the crane was developed, which allows to simulate its work in transient conditions and analyze the results obtained. It is possible to choose the optimal parameters of the brake system, taking into account the minimization of dynamic loads in its elements. The considered technical solution allows to ensure equality of braking forces on all wheels and to increase the reliability of the crane as a whole.

References 1. Wu JS, Chen CH (2001) Torsional vibration analysis of gear-branched systems by finite element method. J Sound Vib 240(1):159–182 2. Couderc Ph, Callenaere J, Der Hagopian J, Ferraris G (1998) Vehicle driveline dynamic behaviour: experiment and simulation. J Sound Vib 218(1):133–157 3. Crowther A, Zhang N, Liu D, Je yakumaran, JM (2004) Torsional finite elements and non-linear numerical modelling in vehicle powertrain dynamics. J Sound Vib (submitted) 4. Reik W(1990) Torsional vibrations in the drive train of motor vehicles—principle considerations. In: Proceedings of the 4th international symposium on torsional vibrations in the drivetrain Baden-Baden, Germany, 20 April 1990 5. Farshidianfar A, Ebrahimi A, Bartlett H (1999) Hybrid modelling and simulation of the torsional vibration of driveline systems. Proc Inst Mech Eng Part D: J Automobile Eng 215:217–229 6. Gonçalves J, Ambrósio J (2001) Complex flexible multibody systems with application to vehicle dynamics. Multibody SysDyn 6(2):163–182 7. Peng D, Zhang Y, Yin CL, Zhang J (2008) Combined control of a regenerative braking and antilock braking system for hybrid electric vehicles. Int J Automot Technol 6:749–757 8. Bottauscio O, Chiampi M, Manzin A (2008) Modeling analysis of the electromagnetic braking action on rotating solid cylinders”. Appl Math Model 32(1):12–27

642

S. V. Streltsov and V. A. Ryzhikov

9. He R, Liu C, Li N (2010) Fuzzy control of the integrated system of electromagnetic brake and friction brake of car. J Mech Eng 46(24):83–87 10. Lee JrK, Park K (2001) Modeling of the Eddy currents with the consideration of the induced magnetic flux. In: Proceedings of the IEEE Region 10th international conference on electrical and electronic technology, pp 762–768, August 11. Zhengdong L, Man W, Jianying Y (2013) Nonlinear robust control of a hypersonic flight vehicle using fuzzy disturbance observer. Math Prob Eng, Article ID 369092:10 12. Yi L, Chen S, Jun W (2011) A simulation study on fuzzy control of anti-lock characteristic of brake-by-wire system. J Highw Transp Res Dev 10:124–127 13. Zhang W, Ding N, Chen M, Yu G, Xu X (2011) Development of a low-cost hardware-in-the-loop simulation system as a test bench for anti-lock braking system. Chin J Mech Eng 24(1):98–104 14. Sedlmayer F (1967) Bremskrafte der Fahrwerke-ihre dynamische. Wirkung auf die Tragkonstruktion der Krane.-Fordern und heben 4:203–215 15. Zolina TV (2010) Ensuring the safe operation of industrial buildings with crane equipment. Collection Modernization of Russian Regions: Investments in Innovations. Materials of the IV International NPK, 15.10.2010. Astrakhan, pp 16–18 16. Ryzhikov VA, Streltsov SV, Kulyashov SI (2013) Drum brake. RU Patent 2012148944,11, 5 Oct 2013 17. Alexandrov MP (2000) Load-lifting machines. MSTU them. N.E. Bauman, Moscow 18. Komarov MS (1969) Dynamics of mechanisms and machines. Mashinostroenie, Moscow 19. Ryzhikov VA (2012) The scientific basis of differential tensioning devices. South-Russian State University of economics and service, Shahty 20. Dyakonov VP (2007) Mathcad 11/12/13 in Mathematics. A Handbook, Moscow

Thermal Loading Estimation of the Friction Pairs of a Vehicle Automated Brake System V. Dygalo and I. Zhukov

Abstract The paper describes the approach for the evaluation of predesign-thermal load of the braking mechanism for vehicles with anti-lock braking systems (ABSs). The main essence of which is that most of the kinetic energy of the car with ABS is extinguished by friction in the brake mechanism. Overheating of the brake mechanism, namely its friction pairs, leads to the phenomenon of critical fading, accompanied by a sharp decrease in the braking torque. A method is proposed for determining the energy quenched in a braking mechanism with ABS using three approaches. The main of which is the ratio of the longitudinal reaction and the braking torque associated with the dynamic radius of the wheel. Since the rotation speed of the wheel of car with ABS during braking varies according to a complex law. Finding the path of friction as a component of the energy balance is based on the linearization of the speed function. Finding the extinguished energy is necessary for carrying out thermal calculation of details of brake system including with the use of a method of finite elements. The study was the basis for creating a computer model of the temperature field of the braking mechanism, which, on the whole, makes it possible to talk about a system for calculating the thermal loading of braking mechanisms with ABS. Keywords Car


Anti-lock system
Brake mechanism
Thermal load
Fading

1 Introduction The increasing availability of automated braking systems, such as anti-lock braking systems (ABS), in addition to the obvious advantages in terms of active safety, creates a number of problems caused by changes in the working process. In particular, this applies to the popular in recent years, high-speed cars J—class-crossovers and SUVs,

V. Dygalo (&)
I. Zhukov Volgograd State Technical University, 28, Lenin Avenue, Volgograd 400005, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_68

643

644

V. Dygalo and I. Zhukov

which, as a rule, while maintaining high-speed mode have an increased weight compared to conventional cars. This increases the amount of kinetic energy that must be extinguished during braking [1]. However, in contrast to the traditional way of stopping a skid, the main part of the kinetic energy of a car with ABS is extinguished by the operation of friction in the brake mechanism [2], which inevitably leads to an increase in thermal loading, especially when using the manufacturers of brake systems of cars, the traditional elements of basic models.

2 The Energy Balance of the Vehicle During Braking Files must be in MS Word only and should be formatted for direct printing, using the CRC MS Word provided. At the same time, it is known from the literature sources that the overheating of the brake friction pairs contributes to development of a critical feding, accompanied by a sharp decrease (up to 50%) in the coefficient of friction of the brake linings, as well as increased wear of the countertops, with the formation of macrosceles [2–5]. Therefore, it is necessary to assess the thermal load of friction pairs of brake mechanisms of cars equipped with automated systems. The analysis of the effect of the principal change in the working process of braking the car wheel with ABS on the distribution of the work necessary to extinguish the kinetic energy of the car as an example is shown in Fig. 1. The graph

Fig. 1 Distribution of work spent on damping the kinetic energy of the vehicle with ABS during braking

Thermal Loading Estimation of the Friction Pairs of a Vehicle …

645

shows: the change in the kinetic energy of the car with ABS when braking from an initial speed of 60 km/h on dry asphalt during braking (Wk), the work spent on friction in contact with the road tire (ATR) and hysteresis losses in the tire (AFF), as well as absorbed in the brake mechanism (ATM) [6]. The graph shows that the braking of the car with ABS, in contrast to the skid braking, the main part of the kinetic energy is extinguished by the friction work in the friction pairs of the brake mechanism. Therefore, despite the development of design and calculation technologies, the assessment of the probability of possible failure of the vehicle brakes during operation due to fading is an urgent task. It is important that in the design of braking, vehicles with ABS manufacturers have the opportunity to assess the effectiveness of their work at the stage of predesign calculation. The calculation of the average temperature of friction elements in the braking process is of great interest in assessing the temperature of the brake and can be particularly useful in cases where the direct measurement of temperatures at real points of contact and the friction surface as a whole is difficult.

3 Approaches to the Assessment of Thermohardened Brake To solve this problem, it is necessary to have computational methods that allow to analyse such performance characteristics as the change in the braking moment of friction, the temperature of the surface of the brake discs, the speed and duration of braking, as well as the work performed depending on the physical, mechanical and thermal properties of the materials of the friction pair. The average temperature of friction pairs, taking into account the short duration of braking, without taking into account heat transfer to the environment, can be determined by the dependence obtained by Professor AV Chichinadze [7–10]
 WW:B: ð1  aW:B: Þb1 1 # ¼ sN þ sW F01 3 k1 Aa1 tT 

ð1Þ

where WW:B: —full braking operation; a00W:B: —the distribution coefficient of heat flow; Aa1 —the nominal area of friction; tT duration of braking; k1;2 —coefficient of thermal conductivity; sN ; sW —time characteristics of power and operation; F01 — Fourier number of friction elements In turn, the coefficient of distribution of heat fluxes required for the calculation of thermal fields can be found from the expression [2] "

a00W:B:

b1 k2 ¼ 1= 1 þ b2 k1

1 3 sN 1 3 sN

where b1;2 is the thickness of friction elements.

þ FO2 sW þ FO1 sW

!# ð2Þ

646

V. Dygalo and I. Zhukov

Fig. 2 Model of the temperature distribution of the brake disc vehicle with ABS

The given dependences allow us to calculate the average temperature of brake discs with the help of finite element software [11–15] with the known full braking operation. Initial parameters in them are parts and assembly of brake mechanisms, values of heat flows and heat transfer coefficients. Thus, as an example, Fig. 2 shows the model of temperature distribution of the brake disc of the car with ABS at a coefficient of convective heat transfer of 35 W/m2 K and ambient temperature of 294 K [16], obtained by using a finite element software. As mentioned above, for the application of the dependencies obtained by AV Chichinadze [7–10], it is necessary to find the full braking operation based on the braking force on the disc and the friction path of the brake disc. At the stage of predesign analysis to determine the braking force, several approaches are possible. The most promising approach is based on the determination of the braking performance brought to the brake disc, which can be found with sufficient accuracy through the ratio of the longitudinal reaction and the braking torque [17]. To do this, we associate them with the value of the dynamic radius. Despite the conventionality of this ratio, this is enough to estimate the amount of work needed. Brake torque on the wheel is determined by the known dependence [18] Mte ¼ Rx
r@

ð3Þ

Thermal Loading Estimation of the Friction Pairs of a Vehicle …

647

To determine nabs the value of the longitudinal reaction, we will use the degree of using the maximum coefficient of adhesion nabs in the operation of the anti-lock braking system. Then, the expression for the longitudinal reaction of the wheel will take the form Rx ¼ Rz
umax
nabs

ð4Þ

The value is determined from the requirements of standards for the braking performance of the car with ABS. So, for dry asphalt, it should not be below the level of the braking skid nabs ¼ 0:9, wet and compact snow nabs ¼ 0:85 [19]. Introduction to the calculation of this indicator greatly simplifies the calculation process and eliminates the consideration of the dependencies of the algorithm of a particular anti-lock system, since the algorithm and the matrix of values of the control unit of the anti-lock system are often trade secret manufacturers of brake systems. The value of the braking force on the average circumference of the brake disc is found from the expression: Pb ¼ 2

Mte rdisc

ð5Þ

where rdisc is the radius of the brake disc. Taking into account the dependence (4), we obtain an expression for the braking force in the form Pb ¼ 2

Rz
umax
nabs
rdin rdisc

ð6Þ

4 The Friction Path To find the full operation of friction in the brake mechanism, it is necessary to know the friction path in the “brake disc–pad” pair. Theoretical background to determine the path of friction of individual wheels of the vehicle in braking mode is given in [20]. The rotation of each individual wheel of the vehicle with ABS when braking is characterized by a different combination of speed modes due to the operation of the brake system and the contact of the tire with the road surface, which is shown in Fig. 3. At the same time, the following characteristic periods can be distinguished: braking, disinhibition, blocking or use of the wheel, free rolling and quasi-uniform movement in the area x_ k  0. Each period is characterized by a different combination of speed and load performance. Taking into account the short duration of the ABS cycle, we assume that

648

V. Dygalo and I. Zhukov

Fig. 3 Period options during braking: A—braking, B— quasi-uniform rotation, C— disinhibition, D—Yuz, E— free rotation (full release), Va —vehicle speed, Vк—given radial rotation speed of the wheel

each period is characterized by a steady slowdown (acceleration) jconst and an average speed. With the same values of the braking distance, cars with ABS and without it can be implemented different characteristics for friction pairs “drum (disc)-pad” and “bus–road”. To calculate the friction path in each pair of brake mechanisms “disc– pad”, use the dependence LTp ¼ ðVHK þ VKK ÞðtKK  tHK Þðrd =rk Þ=2

ð7Þ

2 2  VK;K Þ=2jconst;K LTp ¼ ðVH;K

ð8Þ

or

where VH,K, VK,K —the radial speed of the wheel at the beginning and end of the time interval; tH,K, tK,K—the moments of time at the beginning and end of the time interval; rd—the radius of the brake drum; rk—the radius of the wheel. The friction paths for each period are determined from the following dependencies. The period of inhibition and disinhibition 2 2 LTp ¼ ðVH;K  VK;K Þ=2jconst;K

ð9Þ

where VH,K, VK,K—wheel speed at the beginning and at the end of the period, jconst,K—the steady-state acceleration of the wheel; rK—wheel radius; rdisc—the average radius of the brake disc. Free wheel rolling LTp ¼ 0, Wheel lock LTp ¼ 0 Constant (quasi-uniform) rotation of the wheel LTp ¼ VH;K
tr

ð10Þ

Thermal Loading Estimation of the Friction Pairs of a Vehicle …

649

where tr—the duration of the period of uniform rotation of the wheel. To find the general friction path in all areas, let us sum the above dependencies LTp;pf ¼

I X i¼0

LTpTi þ

J X

LTpPj þ

i¼0

N X

LTpDn

ð11Þ

i¼0

where Lmp, pf—the friction path of the lining of the drum; I—the number of braking areas; J—the number of areas of disinhibition; N—the number of areas of constant rolling of the wheel; LTpTi LTpPj LTpDn

the friction path for a pair of “disc–pad” on the braking area; way friction pair “disc pad” on the jth section of the release; the friction path for the pair “disc–pad” on the nth section of the wheel constant rolling braking.

Finally, the friction work brought to the brake disc is determined from the expression: WW:B: ¼ LTp Pm

ð12Þ

Due to the phenomenon of non-synchronous rotation of the wheels of the car during braking, due to various factors, both in the presence of ABS and without it, the brake mechanisms of even one axis of the car can have different indicators of work on the friction path. It should be noted that the definition of the full braking performance and, as a consequence, the determination of the energy quenched in the brake mechanism was the basis for the creation of a computer model of the temperature field of the brake mechanism, which generally makes it possible to talk about the system for calculating the thermal loading of the brake mechanisms with ABS.

References 1. Revin AA, Zhukov IS, Shkarupelov VS (2012) Methodology of monitoring the technical condition of the braking system of the car with ABS during operation. Izvestia VSTU Ser Nazemnye Transp Sist 89(5):90–93 2. Bosch (2012) Automotive handbook, 3rd edn. Knizhnoe izdatelstvo Za rulem, Moscow 3. Turbin IV, Epishkin VE, Solomatin NS (2014) Effect of friction coefficient on tribotechnical characteristics of disc brake friction pairs. In: Proceedings of the conference. Perspektivnyye napravleniya razvitiya avtotransportnogo kompleksa: sbornik statey VIII Mezhdunarodnoy nauchno-proizvodstvennoy konferentsii, pp 124–128 4. Kokonin SS, Obizhaev GY, Okulov BS et al (2001) High loaded multidisc brakes and factors determining the efficiency and smoothness of their work. Tyazheloe mashinostroenie, pp 19–26 5. Bezyazychnyy VF, Lyubimov RV (2000) Experimental study of the processes of destruction of the surface layers of the metal in the steady-state process of fretting wear. Collect Sci Pap TSTU Mekhanika i fizika friktsionnykh kontaktov 7:24–28

650

V. Dygalo and I. Zhukov

6. Revin AA, Dygalo VG (2014) Creation of the main operational properties of vehicles in braking mode. Avtomobilnaya promyshlennost 11:3–5 7. Chichinadze AV (1970) Thermal dynamics of friction. Mashinostroenie, Moscow 8. Chichinadze AV (1967) Calculation and investigation of external friction during braking. Mashinostroenie, Moscow 9. Chichinadze AV, Hebda MI (1990) Tribotech handbook, vol 3, t.2. Mashinostroenie, Moscow 10. Chichinadze AV, Hebda MI (1989) Tribotech handbook, vol 3, t.1. Mashinostroenie, Moscow 11. Voloaca S, Fratila G (2012) Concerns regarding temperature distribution obtained by experiments and finite element analyses for types of brake discs. UPB Sci Bull Ser D 74(3):33 12. Gudz GS, Eremenko PI (1979) Investigation of the temperature condition of brake mechanisms by modeling. Avtomobilnaya promyshlennost 10:20–22 13. Pershin VK, Fishbejn LA (2005) Simulation of thermal conditions in the friction interaction of the wheel and brake pad. Transp Urala 4:34–44 14. Starostin IP (2005) Numerical solution of the problem of thermal conductivity in friction pairs with a small overlap coefficient. Matematicheskoe modelirovanie 7:23–30 15. Gudz GS, Zahara IY, Tarapon OG (2009) A new approach to modeling the temperature condition of automotive ventilated brake discs during cyclic braking. Collection of scientific papers modelirovaniya v ehnergetike NANU im. G.E. Puhova: Modelirovanie i inform. Tekhnologii 51:37–42 16. Alekseev GN (2005) General heat engineering. Vysshaya shkola, Moscow 17. Revin AA, Zhukov IS, Shkarupelov VS (2012) Methods for determining the full braking operation carried out by the braking mechanism of the car with ABS. Izv VSTU Ser Nazemnye Transp Sist 21(124)(7):21–24 18. Tarasik VP (2006) The theory of the motion of the automobile: textbook for universities. BHV-Peterburg, Sankt-Peterburg 19. Tumasov AV, Groshev AM, Kostin SY, Saunin MI, Trusov YP, Dygalo VG (2011) Investigation of the properties of active vehicle safety by imitation modeling. Zhurnal avtomobilnykh inzhenerov 2:34–37 20. Revin AA, Poluektov MV, Radchenko MG, Zabolotnyy RV (2013) The influence of the ABS workflow on the durability of the vehicle chassis elements: monograph

Problem of Increasing Tractive Effort of Railway Locomotives in Conditions of Arctic and Continental Shelf Regions A. Keropyan, S. Albul and A. Zarapin

Abstract The completed research proves the relevance of the method and creation on its basis of a device for deicing railway rails for industrial and urban transport in the conditions of the Arctic and continental shelf regions of the Russian Federation, which will increase the tractive effort of operating and newly created locomotives and solve the problem of mineral transportation from the explored fields of the Arctic and continental shelf regions of the Russian Federation. The method and the device for deicing railway rails allow one to revive the Great Northern Railway Project and ensure the strategic security of the RF, according to the Strategy of Scientific and Technological Development Program of the Russian Federation No. 642 approved by the Decree of the President of the Russian Federation dated December 1, 2016. The development of an innovative way to increase the tractive effort of locomotives in the conditions of ice building on the rail working surfaces during the rolling stock operation in the Arctic and continental shelf regions of the Russian Federation will contribute to the creation of a year-round overland communication between Chukotka and the Barents Sea, duplicating the Northern Sea Route.










Keywords The Arctic region Continental shelf Industrial railway transport Locomotive Wheel–rail system Ice deposit Heated sand Friction coefficient













A. Keropyan
S. Albul (&)
A. Zarapin The National University of Science and Technology ‘MISIS’, 4, Lenin Prospekt, Moscow 119049, Russia e-mail: [email protected] A. Keropyan e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_69

651

652

A. Keropyan et al.

1 Introduction It is known that the Arctic region is characterized by below-freezing temperatures, which significantly affect the friction properties of the wheel–rail system of railway transport, and, as a result, its tractive effort. The friction peculiarities of the wheel–rail system during operation in below-freezing temperatures were studied in sufficient detail in [1–3]. The studies shown that under operating conditions, the friction tracks of the wheels are always covered with industrial contaminants, however, in the winter period, in conditions of significant below-freezing temperatures, the solid and liquid components of these contaminants interacting with snow and ice particles deposited on the rails form an ice crust on their surface. In freezing temperatures below minus 20 °C, the use of sand to improve friction usually does not produce sufficient effect due to a significant increase in the hardness of ice and ice deposits. The sand particles, falling on the ice crust due to low friction, move together with the locomotive wheels and thus create conditions for the locomotive wheelset to slip, resulting in the reduction of friction characteristics of the wheel–rail system. The Russian Arctic is extremely rich in minerals (hydrocarbons, raw materials, solid minerals, precious metals, etc.), and their role in the overall balance of the country’s fuel, energy, and mineral resources is so great that in the future the country will not be able to exist and develop without their utilization. The need to develop large areas in the Arctic Circle is determined by their special socioeconomic, geopolitical, and defense value.

2 Supply of Heated Sand in the Wheel–Rail Friction Zone Experiments show that under these conditions the most effective means of increasing the friction of the locomotive wheel–rail system is the additional heat supply in the friction zone [3–8]. Due to experimental studies, it was established that the friction coefficient on ice-covered rails during heating increases on average from w = 0.13…0.17 to 0.26…0.33 [3]. To increase the friction coefficient in the winter period, we (in collaboration with Prof. Yu. M. Luzhnov) proposed the following technical solution protected by the patent No. 2504492 (“Method for increasing the friction of the wheel to the rail”, the patent holder is NUST ‘MISIS’, [5])—in the area of contact of the wheel with the rail, for example, using an induction heater, the preheated dry quartz sand is fed. The present invention allows to increase the friction coefficient and, consequently, to increase the tractive effort of the locomotive. Due to the reduction in wheel slip, the wear of the tires of wheel pairs and rails will decrease, thereby avoiding additional energy losses associated with skidding. In addition, the heating of the sand contributes to reducing the likelihood of its freezing in the sand-feeding system in winter conditions. Due to its versatility, this method can be applied to all

Problem of Increasing Tractive Effort of Railway Locomotives …

653

types of railway vehicles regardless of the type of wheelset drive mechanism: electric and thermal locomotives, various track machines used in railway transport, as well as traction units used in the mining industry, and wheel–rail urban transport (for example, trams). Calculations show that the use of heated sand in below-freezing temperatures can increase the friction coefficient on average up to 10% [9–11]. We consider it necessary to note that this problem has not yet been solved for the main railway transport, as evidenced by the slip of SAPSAN high-speed train, which followed the route ‘Moscow—St. Petersburg’ in October 2016 [12]. The method of increasing the friction between the wheel and the rail and increasing the locomotive pulling traction in winter conditions is as follows (Fig. 1): abrasive material, for example, dry quartz sand heated to the desired temperature, is moved from hopper 1 to inductor 2, then transported through the adjusted for different amount of sand nozzles 3, and fed into the contact zone 4 of the wheel 5 with the rail 6 via pipelines 7, 8, 9 through nozzle 10. After that, the heated sand, having entered zone 4, gets on the ice crust before wheel 5 on the surface of rail 6, melts it intensively and creates conditions for ensuring normal interaction of heated sand particles with the wheel and the rail. Depending on the operation conditions, the required amount of sand is fed into zone 4 by turning on different nozzles using electro-pneumatic valves 11, 12, 13. By pressing buttons 14, 15, 16, the air from supply line 17 through disconnecting valve 18 and the corresponding electro-pneumatic valves is brought to the sand nozzles. According to the available information, when the surface is heated 18…20 °C above the ambient air temperature, the intense dry sublimation of ice occurs. Fig. 1 Scheme of the method of increasing the wheel and the rail friction

654

A. Keropyan et al.

The formula for determining the heated rail temperature tr: tr ¼ 1:16½ta  þ 16

ð1Þ

where [ta] is the absolute value of air temperature in °C. Taking into consideration, the abovementioned, assuming that the heated rail temperature equals the sand temperature, Formula (1) can be represented as ts ¼ 1:16½ta  þ 16Ks

ð2Þ

where ts is the heated sand temperature in °C; Ks is the coefficient, which takes into account heat losses due to the adiabatic expansion of the air-sand mixture coming from pipeline 3 when sand is delivered through sand-feeding pipes 7, 8, 9 from inductor 2 to nozzle 10. Coefficient Ks also depends on the technical characteristics of the device that implements the flow of sand according to this method. On average, Ks can be assumed as equal to 1.18. To increase the friction coefficient in railway transport, dry quartz sand is used, which is fed in a jet of compressed air from the oncoming side of the wheel to the area of contact between the locomotive wheel and the rail. The sand, after being crushed by the locomotive wheel, is substantially ground up, and as a result, the total surface of the dispersed sand grains increases multifold (several thousand times). In [13, 14], it was established experimentally that the friction coefficient when adding fresh sand under the wheel equals 0.4, and when the locomotive wheel passes the rail section covered with crushed sand, it equals 0.18, i.e., more than 2 times less. This can be explained by the fact that all solid materials have surface energy. For example, for quartz, it is equal to 980 erg/cm2, and during dispersion, the surface energy increases in proportion to the newly formed surface [15].

3 Removal of Waste Sand from Rails The use of sand, in addition to the positive effect of increasing the friction of wheels and rails, also has some disadvantages. First of all, the sand remaining on the rails creates additional resistance to the movement of the train, reaching 12%, and the sand trapped on the rubbing parts of the track and rolling stock contributes to their more abrasive wear. All the sand supplied from the sandboxes of the locomotive pollutes the ballast prism, worsens the condition of the railway track, and thus leads to significant material costs associated with the restoration of the environmental characteristics of the career field [3]. Here we have in mind the possible additional costs for the remediation of the quarry after the mining of the mineral. Based on the studies performed, it should be assumed that the sand remaining on the rails after passing of each locomotive wheel should be removed from the rails by suction and collected in a special container installed on the locomotive. This will increase the friction coefficient up to 12%.

Problem of Increasing Tractive Effort of Railway Locomotives …

655

Fig. 2 Scheme of sand removal from the rail surface

To solve this problem, we have proposed a technical solution protected by the RF patent for invention No. 2641957 dated 19.12.2016. Bul. No. 3 dated 23.01.2018 ‘Method to increase the locomotive tractive effort’ [16]. The technical result of the proposed invention is to increase the locomotive tractive effort, reduce pollution of the ballast prism and is achieved by the following—the sand remaining on the rail surface is loosened by multidirectional air jets immediately after passing of the last wheelset of the first moving locomotive cart under which it was fed, after that the sand is removed from the rail surface by suction through exhaust tubes, and then transported through a pipeline system, and collected in a container for waste sand. The scheme of the proposed method is shown in Fig. 2. The proposed method is achieved as follows. Each locomotive wheel trolley 1 has coupled exhaust tubes of forward 2 and reverse 3 motions, and additional tubes for feeding compressed air 4 and 5, equipped with diffusers 6 at the ends, which divide the airflow into several jets and simultaneously increase its speed; moreover, the air jets come into contact with the contaminated rail surface in different directions. When the wheel slip starts, some sand from locomotive sand system 8 is fed under driving wheels 7. At the same time, suction system 9 is launched, and additional tubes 4 or 5 are connected to the general compressed air supply system of the locomotive 10. Moreover, if the locomotive moves forward 11, then the exhaust tubes 3 of reverse motion are connected to the suction system, and if it moves in the opposite direction, then the exhaust tubes 2 of forward motion. Similarly, additional compressed air supply tubes are connected to the general compressed air supply system of the locomotive. To switch the exhaust tubes and additional compressed air tubes, throttle valves 12 are used. After overcoming the wheel slip, the crushed sand remaining on the surface of the rail head is exposed to multidirectional jets of compressed air. As a result, it is separated from the rail and a sand–air mixture is formed above its surface;

656

A. Keropyan et al.

Fig. 3 Device to increase the locomotive tractive effort

this mixture is sucked into the appropriate exhaust tube and transported through a pipeline system to a waste sand container (not shown). The proposed method of waste sand removal from the rail working surface can be implemented using the device produced according to the RF patent for invention No. 2641611 dated 19.12.2016 Bul. No. 2 dated 18.01.2018 ‘Device to increase the locomotive tractive effort’, Fig. 3 [17]. The device works as follows. When the locomotive moves in direction 1 and the wheel slip starts, sand 5 is fed under driving wheels 2, to the place of contact of the wheel with rail 3 from feeding pipes 4 of the locomotive sand system. Waste sand 6, which has lost its properties, is exposed to compressed air from the general supply system installed on the locomotive, and for this purpose, we use compressed air supply tubes 7, equipped with diffusers 8 at the ends, which are fixed to exhaust tubes 9 installed on both sides of the wheeled carts of the driving coupled wheels of the locomotive. To prevent sand from entering ballast prism 10, exhaust tubes 9 are equipped with nozzles 11; at the same time, the rear part of nozzle 12 and rail 3 has a minimum gap between them, the front part of nozzle 13 is raised above the rail surface to a height that does not impede the passage of air jets from diffuser 8, and the side walls of nozzle 14 is lowered below the working surface of rail 3; the rigidity of the structure is ensured by fixing the lower end of the compressed air supply tube in the upper lid of the nozzle and the diffuser, which can change its position relative to the rail working surface, and ends up inside the nozzle, while the air jets are multidirectional and cover the entire rail surface between the contact point of the wheel and the rail and the rear wall of the nozzle. As a result, when exposed to multidirectional air jets inside the suction tube nozzles located on the side of the coupled wheels of the carts opposite to the direction of locomotive movement, air–sand mixture 15 is formed, which is not dispersed to the sides, but under the action of suction force is sent to the exhaust tubes. The feasibility of the proposed innovations is confirmed by the results of the research conducted in [6, 11, 18, 19]. The relevance of the proposed technical solutions can also be justified by the announced competition of the Ministry of Energy of the RF aimed at the development and exploration of the Arctic and continental shelf regions. The competition should contribute to the creation of sustainable socioeconomic development and

Problem of Increasing Tractive Effort of Railway Locomotives …

657

exploration of the Arctic and continental shelf regions, as well as to the conditions for the introduction of developments, which can be of interest for the scientific and technological potential [13]. In addition, it should be noted that, according to the statement of Prime Minister D. A. Medvedev at a government meeting, the volume of the investment program of the Russian Railways for 2018–2020 will constitute more than 1.8 trillion rubles. This funding, in particular, is intended for the development of the northern regions’ transport system, including the so-called Northern Latitudinal Railway, which is synonymous with the name of the railway between Chukotka and the Barents Sea.

4 Conclusions • The proposed method and based on it special device for deicing railway rails of industrial and urban transport in the conditions of the Arctic and continental shelf regions of the Russian Federation will increase the tractive effort of the operated and newly created locomotives. • The method and the device for combating ice deposits on railway tracks allow to revive the Great Northern Railway Project and to ensure the strategic security of the Russian Federation, according to the ‘Strategy of the Russian Federation Scientific and Technological Development’ Program No. 642 approved by the Decree of the President of the Russian Federation dated December 1, 2016. • The development of an innovative way to increase the tractive effort of locomotives when rail working surfaces are covered with ice during the rolling stock operation in the Arctic and continental shelf regions of the Russian Federation will contribute to the creation of a year-round overland communication between Chukotka and the Barents Sea, duplicating the Northern Sea Route. • The implementation of the method and the device for waste sand removal from the rail surface will increase the tractive effort of locomotives and will improve the environmental situation in the railway area in the Arctic region. Acknowledgements The authors are grateful to the Chief Scientific Officer of OAO Russian National Research Institute of Railway Transport ‘VNIIZhT’, Professor, Doctor of Technical Sciences Yury Mikhailovich Luzhnov, for useful advice and methodological assistance when writing the article.

References 1. Kosikov SI (1967) Friction properties of railway rails. Science, Moscow, p 111 2. Luzhnov YuM (1973) Peculiarities of friction on the rails in winter conditions, vol 445. Tr. MIIT, Moscow, 130 p

658

A. Keropyan et al.

3. Luzhnov YuM (2009) Nanotribology of wheels and rails friction. Reality and possibilities. Intext, Moscow, p 176 4. Keropyan A, Gorbatyuk S, Gerasimova A (2017) Tribotechnical aspects of wheel-rail system interaction. In: International conference on industrial engineering, ICIE 2017, Procedia engineering, vol 206, pp 564–569 5. Keropyan AM, Kantovich LI, Voronin BV, Kuziev DA, Zotov VV Influence of uneven distribution of coupling mass on locomotive wheel pairs, its tractive power, straight and curved sections of industrial rail tracks. http://iopscience.iop.org/volume/1755-1315/87http:// iopscience.iop.org/article/10.1088/1755-1315/87/6/062005 6. Keropyan A, Gerasimova A (2017) Connection of the temperature in contact area of the wheel-rail system with the railway slope of industrial railway transport. Izvestia Ferrous Metall Natl Univ Sci Technol MISIS 60(5):355–363. https://doi.org/10.17073/0368-07972017-5-355-363 7. Zarapin AYu, Shur AI, Chichenev NA (1999) Improving the installation for rolling an aluminum strip coated with corrosion-resistant steel. Steel Transl 29:69–71 8. Sedush VY, Sidorov VA, Eron’ko SP, Oshovskaya EV (2003) Optimal methods of maintaining of serviceability of mechanical equipment. Stal 7:77–79 9. Kaminsky VD, Suprunenko OD, Smirnov AD (2016) Mineral resources of the Arctic continental border land of Russia and the prospects for their development. FGUP VNII Okeanology Arctic Ecol Econ 3(15):53–61 10. Keropyan AM, Luzhnov YuM (2014) Method for increasing the friction of the wheel to the rail. Patent of the Russian Federation for the invention 2504492 IPC B61C 15/10, Moscow State Humanitarian University and National University of Science and Technology ‘MISiS’ 11. Keropyan AM (2015) Innovations in career railway transport. Coal 1:31–34 12. SAPSAN high-speed train got stuck on icy rails. News item of VESTI TV program, Russia, dated 27.10.2016. http://www.vesti.ru/doc.html?cid=8&id=2815285 13. Letter of the Ministry of Energy of the Russian Federation from 01.11.2017 No. КM—12186/15 ‘On the competition of scientific, technical and innovative developments aimed at development and exploration of the Arctic and continental shelf regions’ 14. Rengevich AA (1961) Friction coefficient of mine electric locomotives. In vol. ‘Issues of mining vehicles’, ed. Corr. Academy of Sciences of Ukraine Prof. N.S. Polyakova, vol 5. GNTILGD, Moscow, 227–246 15. Khodakov GS (1972) Physics of grinding. Science, Moscow 16. Keropyan AM, Gorbatyuk SM, Bibikov PYa, Kuziev DA (2018) Method to increase the locomotive tractive effort. Patent of the Russian Federation for invention 2641957 IPC B61C 15/08, National University of Science and Technology ‘MISiS’ (RU) 17. Keropyan AM, Gorbatyuk SM, Verzhanskiy PM, Kuziev DA (2018) Device to increase the locomotive tractive effort. Patent of the Russian Federation for invention 2641611 IPC B61C 15/08 National University of Science and Technology ‘MISiS’ (RU) 18. Bardovskiy AD, Gorbatyuk SM, Keropyan AM, Bibikov PYa (2018) Assessing parameters of the accelerator disk of a centrifugal mill taking into account features of particle motion on the disk surface. J Friction Wear 39(4):326–329 19. Zarapin AYu, Levitsky IA, Mokretsov AS, Chichenev NA (1999) Modeling of rolling of three-layer strips with resistance of heating. Steel 7:61–64

New Approach for Experimental Identification of Internal Combustion Engine Power and Performance Characteristics A. Egorov, N. Syutov and V. Belogusev

Abstract This paper proposes a new approach to identify performance and power characteristics of an internal combustion engine on the basis of measurements of angular accelerations of a crankshaft. In contrast to the existing methods for measuring engine torque, power and efficiency, the proposed technique allows the evaluation of its parameters in a wide range of speed values without removing an engine from a vehicle, as well as without the use of brake testing equipment, which, in turn, significantly increases the cost of measurements. The absence of strain gauges in the developed measuring system, which require accurate calibration and have a low level of response speed, makes it possible to carry out measurements in dynamic modes of operation with a higher accuracy. The proposed approach was experimentally evaluated using a four-cylinder engine with a capacity of 1.499 L equipped with a developed software and hardware system. It was determined that the obtained torque values insignificantly differ from the values specified in the engine certificate, which indicates that the proposed method is applicable for monitoring performance parameters and identifying the external characteristics of the subject under research both at the production stage and during its operation.







Keywords Efficiency Torque measurement Friction losses Internal combustion engine Mechanical characteristics







A. Egorov (&)
N. Syutov
V. Belogusev Institute of Mechanics and Machine Building, Volga State University of Technology, 3, Lenin Sq., Yoshkar-Ola 424000, Russia e-mail: [email protected] V. Belogusev e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_70

659

660

A. Egorov et al.

1 Introduction Nowadays, the development of methods for testing internal combustion engines prior to their breaking-in is associated with an introduction of approaches to register their performance parameters and power characteristics, which are characterized by less labor and energy costs and allow measurements without destructing the integrity of a mechanical system [1, 2]. At the present time, brake methods are most widely used for those purposes [3–8]. However, their application can lead to an increased error when used in dynamic operating modes of a vehicle, since the torque is registered with relatively high level of discreteness due to the time required to restore the elastically deformed state of a strain element [9–14]. In addition, such methods need expensive cooling systems and brake equipment, and measurements do not allow determining the power characteristics of an engine itself without removing it from a vehicle, which also limits their use [3, 6, 8, 15–18]. Thus, the current level of methods and tools to measure power and performance characteristics of internal combustion engines requires new approaches, which allow identifying the desired quantities in dynamic modes of operation in a wide speed range and without limitations inherent in the existing techniques. To substitute brake methods for new ones, in turn, is possible when solving the problem of high-precision identification of the moment of inertia of an internal combustion engine without demounting it from a vehicle, which is impossible to perform using existing methods and measuring instruments [19, 20]. The purpose of this paper is to develop and experimentally evaluate a new low-cost approach to determine the moment of inertia, torque, power, and performance characteristics of an internal combustion engine for dynamic operating modes, which allows measurements with a higher accuracy and without removing an engine from a car.

2 Materials and Methods For a detailed description of the proposed approach for identification of power and performance characteristics of an internal combustion engine, Fig. 1 presents a scheme for determining the moment of inertia of its rotating elements using a flywheel with a known value of inertia. During the measurement, the operating temperatures of the gearbox oil, engine oil, and coolant of an internal combustion engine are to be close to the rated values. One of the driving wheels of a vehicle is demounted, and a flywheel 5 with a pre-calculated moment of inertia, Jflywheel, is mounted in its place. Then, an internal combustion engine is started, and, when the direct gear of a gearbox 3 is engaged, a certain angular velocity of a crankshaft is set using the adjustment elements. After that, an internal combustion engine accelerates to the

New Approach for Experimental Identification …

661

Fig. 1 A scheme of vehicle elements to determine the moment of inertia of the rotating parts of an internal combustion engine: 1 is an internal combustion engine; 2 is a clutch; 3 is a gearbox; 4 are transmission units; 5 is a flywheel with a known moment of inertia

required angular velocity, and the angular acceleration of rotating elements, e1, which includes a flywheel 5 with a known moment of inertia, transmission units, a gearbox, and an internal combustion engine, is measured. The average torque value, M, for the selected angular velocity range is calculated according to the following equation: M ¼ e1

Jflywheel J1 þ 2 2 2 kgb kadg kfd

! ð1Þ

where J1 is the moment of inertia of a system of rotating elements including a flywheel with a known moment of inertia, transmission units, a gearbox, and an internal combustion engine (the moment of inertia is specified with respect to the axis of the crankshaft rotation); kgb is the gear ratio between an internal combustion engine and a drive wheel; kadg is the gear ratio of an axle drive gear; kfd is the gear ratio of a final gear. Next, a clutch 2 is disengaged, and the flywheel 5 is demounted. After engaging the clutch 2, the angular acceleration of rotating elements, e2, which includes transmission units, a gearbox, and an internal combustion engine, is determined in the same range of angular velocity of a crankshaft. The average torque value, M, is determined according to the following equation: M ¼ e 2 J1

ð2Þ

Equating the right-hand sides of Eqs. 1 and 2, we can determine the specified moment of inertia of a system of rotating elements including transmission units, a gearbox, and an internal combustion engine according to the following equation: J1 ¼

e1 Jflywheel 2 k2 k2 ðe2  e1 Þ
kgb adg fd

ð3Þ

662

A. Egorov et al.

Then, the clutch 2 is disengaged, and the angular acceleration, e3, of a system of rotating elements including only an internal combustion engine with the moment of inertia JEngine is measured in the same range of angular speeds of a crankshaft. The average torque value, M, is determined according to the following equation: M ¼ e3 JEngine

ð4Þ

Equating the right-hand sides of Eqs. 2 and 4, the moment of inertia of a system of rotating elements including only an internal combustion engine is determined according to the following equation: JEngine ¼

e2 J1 e3

ð5Þ

Substituting the right-hand side of Eq. 3 for J1 in Eq. 5, we determine the moment of inertia of a system of rotating elements of an internal combustion engine 1 according to the following equation: JEngine ¼

Jflywheel e2 e1

2 2 2 e3 e2  e1 kgb kadg kfd

ð6Þ

Thus, by determining the moment of inertia of rotating elements of an internal combustion engine and knowing the gear ratios of transmission units, we can determine the level of change in mechanical losses directly in an internal combustion engine (by comparing the obtained values of the moments of inertia with the values specified by manufacturers), as well as its torque, power, and performance. To experimentally evaluate the proposed approach, a hardware–software system was developed and applied to determine the power and performance characteristics of an internal combustion engine (see Fig. 2).

Fig. 2 Hardware and software system for measuring power and performance characteristics of internal combustion engines: 1 is a laptop; 2 is a high-speed data processing unit; 3 is a modified crankshaft position sensor; 4 is a high-precision digital fuel flow meter

New Approach for Experimental Identification …

663

Fig. 3 Schematic diagram of a software and hardware complex for measuring power and performance characteristics of internal combustion engines

This measuring instrument includes an upgraded crankshaft position sensor 3, a high-precision digital fuel flow meter 4 (11,500 pulse per liter) manufactured by Biotech GmBH, a module 2 for converting and inputting signals into a computer manufactured by National Instruments, a laptop computer Acer Extensa 5210 (licensed development environment of virtual instruments LabView 7.2 with Signal Express 2.5). Two main output contacts of the crankshaft position sensor were connected to the computer’s digital information and input modules. A schematic diagram of the designed measuring system is presented in Fig. 3.

3 Results and Discussion To evaluate the reliability of the results obtained using the proposed approach to determine the power and performance characteristics of an internal combustion engine without removing it from a vehicle, in the experimental part of this work, we tested an engine of type VAZ-2112 (serial number is 0190688) with a displacement of 1.499 L, which was mounted on a vehicle of model VAZ-2112 (identification number is XTA211200Y0002158). The control unit was supplied with software (“firmware”) of type J5V05J16 (dynamic). As a flywheel, we used a steel disk, the moment of inertia of which was 0.1695 kg m2.

664

A. Egorov et al.

Table 1 Average values of the moments of inertia of the test systems of rotating elements Parameter

Value

Average value of the moment of inertia of rotating elements including the test transmission units, gearbox, and internal combustion engine (moment of inertia is specified with respect to the axis of the crankshaft rotation) (kg m2) Average value of the moment of inertia of rotating elements including only the test internal combustion engine (moment of inertia is specified with respect to the axis of the crankshaft rotation) (kg m2)

0.28345

0.19715

The tests were carried out according to the approach described above. In the experimental part of this work in the application of the developed technique, ten measurements were carried out, blunders were eliminated, random errors were determined, and the average values were found. As a result of the measurements, the average values of the moments of inertia of the studied systems of rotating elements were determined. Table 1 presents the results obtained. To evaluate the practical applicability of the proposed approach, a series of experiments to determine the angular accelerations of the crankshaft with and without the flywheel in the range of rotational speeds of 1000–6000 revolutions per minute were carried out, on the basis of which the instantaneous values of the torque developed by the test internal combustion engine were found. Figure 4 presents the instantaneous torque values obtained. Figure 5 presents the average torque values. The obtained values were compared with the values specified by the manufacturer of an engine VAZ-2112 with a displacement of 1.499 L. According to the indicated values, VAZ-2112 develops a maximum torque of 132 Newton-meters at a crankshaft rotational speed of 3700 revolutions per minute (power is 49.7 kw). According to the experimental data obtained, the torque at 3700 revolutions per minute was approximately 137 N m. The discrepancy between the certificate data and the experimental result was approximately 3.6%. The resulting discrepancy may be due to the fact that the “dynamic” firmware to provide a more dynamic mode of operation of a vehicle sends a signal to inject a larger amount of fuel compared to the “standard” firmware, which leads to an increased heat release during the combustion of fuel, which can result an increase in torque value. Fig. 4 Instantaneous torque values developed by the internal combustion engine of type VAZ-2112

New Approach for Experimental Identification …

665

Fig. 5 Average values of the torque developed by the internal combustion engine of type VAZ-2112

4 Conclusion Based on the results obtained during theoretical and experimental study, it can be concluded that the use of the proposed approach to determine the moment of inertia of the system of rotating elements of an internal combustion engine allows, with lower financial and energy costs compared to existing brake methods and with a sufficient accuracy, determining the torque, the power, and the efficiency of an engines equipped with mechanical gearbox without removing it from a vehicle. With the use of the proposed software and hardware system, in contrast to the existing tools, one can measure the performance of an internal combustion engine with a higher accuracy in dynamic operating modes, which is provided by the absence of tensometric elements in the measuring system. Identifying only one parameter (angular acceleration of a crankshaft), it is possible to monitor the deterioration of the technical condition of the engine under test. In the follow-up research, it is planned to apply the developed technique using a flywheel with a predetermined moment of inertia to identify power characteristics and performance parameters of hydraulic and pneumatic motors, as well as of external combustion engines.

References 1. Loskutov AS, Grigoryev AN, Kozhin DV (2007) Ispytaniya dvigateley vnutrennego sgoraniya (Internal combustion engines testing). Mari State Technical University, Yoshkar-Ola 2. Franco J, Franchek MA, Grigoriadis K (2008) Real-time brake torque estimation for internal combustion engines. Mech Syst Signal Process 22:338–361. https://doi.org/10.1016/j.ymssp. 2007.08.002 3. Martyr AJ, Plint MA (2012) Dynamometers: the measurement of torque, speed, and power. In: Engine testing, 4th edn. Butterworth-Heinemann, Oxford, pp 227–258. https://doi.org/10. 1016/b978-075068439-2/50011-6 4. Martyr AJ, Plint MA (2012) Engine testing, 4th edn. Butterworth-Heinemann, Oxford. https:// doi.org/10.1016/b978-0-08-096949-7.00017-0

666

A. Egorov et al.

5. Thompson JK, Marks A, Rhode D (2002) Inertia simulation in brake dynamometer testing. In: Proceedings of the 20th annual brake colloquium and exhibition. https://doi.org/10.4271/ 2002-01-2601 6. Mesonero I, López S, García-Granados FJ et al (2016) Indirect characterisation of indicated power in Stirling engines through brake power measurements. Appl Therm Eng 100:961–971. https://doi.org/10.1016/j.applthermaleng.2016.02.046 7. Öktem H, Uygur I, Çevik M (2018) Design, construction and performance of a novel brake pad friction tester. Measurement 115:299–305. https://doi.org/10.1016/j.measurement.2017.10.058 8. Irimescu A, Mihon L, Pädure G (2011) Automotive transmission efficiency measurement using a chassis dynamometer. Int J Automot Technol 12(4):555–559. https://doi.org/10.1007/ s12239-011-0065-1 9. Guerrero PD, Jiménez-Espadafor FJ (2019) Torsional system dynamics of low speed diesel engines based on instantaneous torque: application to engine diagnosis. Mech Syst Signal Process 116:858–878. https://doi.org/10.1016/j.ymssp.2018.06.051 10. Thor M, Egardt B, McKelvey T et al (2014) Closed-loop diesel engine combustion phasing control based on crankshaft torque measurements. Control Eng Pract 33:115–124. https://doi. org/10.1016/j.conengprac.2014.08.011 11. Thor M, Egardt B, McKelvey T et al (2014) Using combustion net torque for estimation of combustion properties from measurements of crankshaft torque. Control Eng Pract 26:233– 244. https://doi.org/10.1016/j.conengprac.2014.01.014 12. Espadafora FJJ, Villanueva JAB, Guerrero DP et al (2016) Measurement and analysis of instantaneous torque and angular velocity variations of a low speed two stroke diesel engine. Mech Syst Signal Process 49:135–153. https://doi.org/10.1016/j.ymssp.2014.04.016 13. Levintov SD (1984) Beskontaktnye magnitouprugiye datshiki krutyashego momenta (Noncontact magnetoelastic sensors for torque measurements). Energoatomizdat, Moscow 14. Kozlov KE, Egorov AV, Belogusev VN (2016) A new method for control of the efficiency of gear reducers. In: IOP conference series: materials science and engineering, vol 189. https:// doi.org/10.1088/1757-899x/189/1/012003 15. Matějka V, Metinöz I, Wahlström J et al (2017) On the running-in of brake pads and discs for dyno bench tests. Tribol Int 115:424–431. https://doi.org/10.1016/j.triboint.2017.06.008 16. Rajabi-Vandechali M, Abbaspour-Fard MH, Rohani A (2016) Development of a prediction model for estimating tractor engine torque based on soft computing and low cost sensors. Measurement 121:83–95. https://doi.org/10.1016/j.measurement.2018.02.050 17. Walter A, Brummund S, Merz B et al (2007) Estimation of the instantaneous engine torque for vehicles with dual mass flywheel (DMF). IFAC Proc Vol 40:167–174. https://doi.org/10. 3182/20070820-3-US-2918.00024 18. Kee R, Blair G (1994) Acceleration test method for a high performance two-stroke racing engine. In: Motor sports engineering conference proceedings volume 2: engines and drivetrains. https://doi.org/10.4271/942478 19. Yablonskiy AA (1977) Kurs teoretitsheskoy mekhaniki. Tchast’ II. Dinamika (Course on theoretical mechanics. Part 2. Dynamics). Akademiya, Moscow 20. Kozlov KE, Egorov AV, Belogusev VN (2018) A method and instruments to identify the torque, the power and the efficiency of an internal combustion engine of a wheeled vehicle. In: IOP conference series: materials science and engineering, vol 289. https://doi.org/10.1088/ 1757-899x/289/1/012038

Development of Hardware and Software Complex for Increase of Technical Readiness Transport-Technological Machines in Forestry S. V. Lyakhov and S. V. Budalin

Abstract The paper studied information technology in relation to the decision-making process in the management of the technical condition of transport and technological machines in forestry. The export of timber and forest products by road has a seasonality, which greatly limits its working time during the year. During this period, the requirements for the technical readiness of the fleet of transport-technological machines of a logging enterprise increase significantly. To reduce the downtime of transport and technological machines, it is planned to develop a software and hardware complex that works on the basis of an advisory information interactive system, which allows optimizing the time for troubleshooting and eliminating it. The implementation of the software and hardware complex in the system of maintenance and repair of transport and technological machines will increase the speed of the quality of technical decisions to ensure their performance. The paper considers the concept of creating an advising system based on deep neural networks.










Keywords Software complex Hardware complex Transport machines Technological machines Dispatching Automation diagnostics Technical service Information system













1 The Introduction Operation of transport and technological machines (TTM) in forestry is carried out on considerable distance from production and technical base of services of technical service of the enterprise and in such road conditions which can be overcome only S. V. Lyakhov (&) Ural Federal University, 19, Mira Street, Ekaterinburg 620002, Russia e-mail: [email protected] S. V. Lyakhov
S. V. Budalin The Ural State Forest Engineering University, 37, Siberian Tract, Ekaterinburg 620100, Russia © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_71

667

668

S. V. Lyakhov and S. V. Budalin

by a rolling stock of the raised possibility. The use of modern TTM is accompanied by their maintenance and repair by the manufacturer through a network of dealerships. This allows you to maintain the warranty obligations of the manufacturer, which protects the interests of machine owners. A condition for such protection is the obligation to undergo maintenance and repair only in dealerships. The intensive development of the TTM fleet is ahead of the development of the production and technical base of enterprises, which has a significant impact on increasing the share of the fleet with high resource mileage [1]. The insufficient perfection of the production and technical base of enterprises, in particular, has an impact on the premature decommissioning of TTMs, whose share reaches from 3 to 8% annually of their total number [2]. In the conditions of considerable dissociation of the operated TTM and remoteness of places of their specialized maintenance and repair, there is a need for providing remote technical impact for the purpose of ensuring their operability. The greatest impact on the efficiency of transport and technological machines has their technical condition. Direct impact on the efficiency of the fleet has four mutually independent indicators of technical operation: aв—factor of technological machines output, b—loaded mileage proportion, c—cargo load factor, Vэ—running speed, km/h [3]. All but the second indicator depends on the technical condition of the car. Often carrying out the routine maintenance works defined by manufacturer at technical services does not allow to provide trouble-free operation of cars throughout their resource. The operational management of the above indicators allows us to ensure the required performance of the transport department, taking into account the optimal selection of TTM for the specific energy consumption of removal [4–8]. The scope of maintenance work is formed on the basis of regulations on maintenance and repair of the corresponding machine and carrying out diagnostic works performed without disassembly of components and assemblies or with partial disassembly [9].

2 Object and Methodology The object of the study is the process of technical impact on TTM in order to restore or ensure their performance. Managing process of maintenance and repair of the TTM involves remote supervision of a specialist technical service (STS) to the dealership with the actions of the mechanic of the enterprise. Mode lack of the ability to provide a reliable remote link with the specialist of STS mechanic has the ability to obtain auxiliary information through the use of special software is an interactive information consulting system (IICS). The system allows you to optimize the process of troubleshooting, accompanied by its comprehensive audio and video information. The target function is to optimize the time spent restoring or maintaining the health of a node or aggregate. It is proposed to create a hardware and software complex based on information technology, which will provide qualified support of technical impacts on the

Development of Hardware and Software Complex …

669

machine in the mode of dispatching. Thus, a single technician dispatch service can provide support for maintenance and repair of several machines or their components and assemblies remotely. If it is impossible to provide control of actions of workers on maintenance and repair of cars or in the absence of such need, it is possible to use the hardware and software complex on the basis of IICS. It allows to carry out under its management a complex of measures directed on clarification of the reason of refusal and also offers technological options of its elimination. The interaction between the mechanic and IICS is carried out by means of voice control. The system contains reference and information support for obtaining the necessary information about the object of influence (unit, mechanism, machine) by mechanic. IICS is used to carry out the process of training of young specialists in the field of maintenance and repair of transport and technological machines, as well as specialists undergoing retraining. IISS provides automation of maintenance and repair of machines by means of search of the reasons of refusal and the offer of technological routes on its elimination, depending on technical equipment of the zone and repair. Dispatching and automation of maintenance and repair of TTM will significantly reduce the recovery time of their performance, as well as to carry out maintenance without the need to carry out the delivery of machines to the location of the dealer center. Diagnosis of TTM is planned to be carried out using methods of analysis of vibroacoustic signals and the method of studying sound effects. Analysis of vibration parameters is the only method of non-destructive testing, which allows to determine the actual technical condition of the dynamically operating unit without prolonged downtime [10]. The use of methods based on the study of vibration parameters of individual components and assemblies, both separately and together, gives a more complete picture of the technical condition of the unit or assembly. Study of parameters of vibroacoustic signals is the basis for the formation of databases on which it will be possible to determine the limit for the technical condition of individual components and assemblies of machines and equipment. Also, the need to analyze vibroacoustic signals will require the development of mathematical models for the development of the most common defects and the creation of methods for the use of diagnostic equipment [11, 12]. Each of the methods of vibration control has its own limitations on the scope; there is no single universal method that could be equally effectively used both for express diagnostics with regard to vibration parameters and for periodic monitoring of the technical condition of the object being diagnosed [13]. Only the results of the integrated use of several different diagnostic approaches can provide an opportunity for an effective and accurate assessment of the actual state of the components and assemblies of working mechanisms. The conducted vibroacoustic studies prove that the best results can be achieved using a comprehensive diagnostic approach using several methods of vibroanalysis and non-destructive testing. Usually, the basis of this approach is the results of spectral analysis, in addition to which often use kurtosis, analysis of the envelope and the rotor trajectory/procession, quite often

670

S. V. Lyakhov and S. V. Budalin

the shock pulse method can be used [14]. In some cases, it is advisable to add a wavelet transform of the signal and a cepstral analysis to this set [15]. It is this combination of vibroanalysis methods (which depends only on the type of object being diagnosed, its operating modes and measurement features) that is necessary and sufficient to control the vibration parameters of the most different process equipment, allows you to get the maximum useful diagnostic information with minimal time spent on measuring. To find faults and ways to eliminate them in IICS provides for the use of neural networks, which will significantly reduce the downtime of the machine. The use of neural networks will reduce the duration of training for the diagnosis of various components and mechanisms of transport and technological machines by reducing the requirements for their qualification. Hardware-based neural networks will optimize the collection and accumulation of statistical information for individual machines, allowing to predict their technical condition depending on the complexity of the various production tasks. The operation of units and mechanisms is accompanied by characteristic sounds that can be classified. The sounds may vary depending on the technical condition of units and mechanisms, as well as to vary among different models, depending on vehicle class. The object of the study is the internal combustion engine of transport and technological machines. The aim of the research is to design an algorithm that allows sound to classify sounds by belonging to the source of excitation. Recognition of sounds and their classification was carried out on the basis of neural networks that perceive acoustic information in the form of a digital series or vector. This method will allow to solve various problems of classification of sounds. The vector we can consider as normal in a plane and determine relative to other points of the plane: to the left or to the right of the vector. Neural network is based on the principle of combining individual neurons. Thus, neurons are combined in a network when the outputs of one neuron become inputs of the other. This can be a single-layer network consisting of many neurons, thus increasing the number of inputs and outputs, or they can be Daisy-chained together and organize multi-layered network. Such possibilities allow to make a system of hyperplanes, and which allow to limit rather difficult areas and to solve complex problems of classification when points get to one part, or to other part of area concerning a vector. The neural network indicates the location of this point and thus classifies the input information. The neural network must be trained to function properly. Training is possible with or without a teacher. For the method with the teacher, a training sample is used, according to which the neural network shows the correct answers to the information submitted for processing. The problems solved before the 2000s allowed to work with neural networks no more than 5–6 layers thick due to the fact that computing resources were running out. Then, there was a revolution with “cats” and games. At the entrance of the neural network is a photograph, the result is given: if correct, then go to the next set, if not correct, then moving from the end to the beginning of “adjust” the weight of

Development of Hardware and Software Complex …

671

the neurons so that the answer is correct. This procedure is done as long as the entire training set will not be given the correct answer. This is the process of learning a neural network: each neuron is adjusted weight vector to get the right answer. High results of the neural network are obtained if the network has 5–6 layers or more. Neural network training without a teacher is carried out according to the learning algorithm of Hebb, and this article is not considered. After training a neural network on a training sample, it is given a test sample from the training network and the proportion of correct answers is checked. The degree of accuracy in the area of 90–95% of correct answers is considered acceptable. The training sample can be manually typed, or you can use specialized databases with handwritten fonts to train the neural network. A neural network can be trained and used in the future for the processing of graphic or audio information. A neural network can be trained to recognize all objects in graphic images. Databases and large productive capacities of personal computers allow to create deep neural networks that allow to create neural networks in 20–30 layers. At the same time, layers can be specialized, for example, so-called convolutional neural networks, where the image is multiplied by the convolution matrix and a reduced image is obtained, which has already lost some of the data, but some features that were not recognizable by the naked eye were highlighted. If you apply this convolution to a photograph several times, you can get a stylized image, in which the characteristic features are separated. This uses such a feature that the used convolutional matrix that processes the image is selected based on the training sample. Thus, the neural network learns to isolate some characteristic features in the photograph. The task is relevant, for example, for processing images from satellites. One layer separates the dashes, the other—the circles, the third—the corners, the next layer collects some objects from them, and the next figure, contour or forms an image. A new direction for neural networks is to work with sounds. With sounds there is a great difficulty associated with the fact that the sound has both temporal and geometric characteristics. Therefore, it is impossible to apply the same approach as with graphic images. One option, when the amplitude of the sound is depicted, is treated as an image and fed to the input of a neural network trained to work with images and classifies the sound. In this case, the sound should be perceived by the neural network as a whole, that is, the previous neuron in the chain should “remember” information when it is transmitted to the input of the subsequent neuron. As an example, the application “Яндeкc. Conversation for Android™” (Yandex Conversation for AndroidTM), which is positioned as an application for deaf people. Speech is recognized by the program and generates text on the smartphone screen. Thus, this text can be read and give the answer. Works almost instantly and pretty quality, taking into account the additional noise and other voices in the dubbing process. Audio processing requires a more complex network architecture than image processing to work with a neural network. The network must have the ability of the network to remember the previous sound States. This means the creation of recurrent neural networks, when connections between neurons are not only in the same direction but also have the ability to move in the opposite direction. The Program

672

S. V. Lyakhov and S. V. Budalin

“Яндeкc. Conversation for Android™” (Yandex. Conversation for AndroidTM) is based on two recurrent neural networks, one of which divides speech into tokens (syllables), and the other already compares them with text.

3 Research Result Material and technical support of the hardware and software complex involves the availability of audio and video recording in the area of maintenance and repair of TTM (Fig. 1), GSM/GPS/GPRS communication module and the appropriate software for linking devices with each other and coordinated data exchange between them. Mechanical locksmith is equipped with video recording and devices for wireless audio. Collecting information from the area of maintenance and repair is carried out by storing it on a local server. The above devices are combined into a local network. Through communication via the Internet, you can make the monitoring process in the area of maintenance and repair, and make the control mechanic. This task is assumed by the operator STS, which controls the progress of the maintenance and repair process, providing optimization of this process for the time spent on restoring the health of the node or machine unit. In the process of using the hardware-software complex, step-by-step collection of information about the repaired objects is carried out. Information from the previously

Fig. 1 Structure of IP video recording. Local area network—Internet

Development of Hardware and Software Complex …

673

identified object is collected by fixing the diagnostic parameters of the objects of each species group using measurement equipment connected to the local server managing the parameter measurement process. Provide possibility of electronic processing of results of diagnostics in two levels: in the local server of species group and in the central computer collecting information from local servers of all species groups. The possibility of independent evaluation of diagnostic results, and hence the quality of repair work, provide by equipping and local server of each group of objects and computers with autonomous databases. Hardware and software system includes means of identification of the object and the performer. The system for quality control of repair and maintenance is made with the ability to connect to the Internet and allows for a comprehensive, prompt, and timely quality control of repair and maintenance. Diagnose the object by measuring the parameters of the basic assembly units. Depending on the measurement results, the repair actions include repair or replacement of the base assembly unit with its parts. Parts and assembly units with a change factor above the specified threshold are replaced without diagnosis. Time and material costs are reduced during repair [16–20]. Dispatching of maintenance and repair will allow centralized collection, accumulation of information and its further processing. The updated information on the technical condition of individual units and units of a particular machine will provide the possibility of a more accurate prediction of wear and tear of their elements, as well as advance maintenance of repair production with spare parts and materials. Dispatching will allow to develop the direction of determining the technical condition of the object by processing the sound of their work, in particular, the internal combustion engine [21]. The interactive information consulting system (IICS) is applied in case of absence of possibility of remote technical influence on object. Mechanic-mechanic serving TTM works with the system in voice mode on the principle of question– answer (block diagram design). The system allows: • to determine the cause of the fault, forming a request or a list of effects on the machine in order to collect the initial data, narrowing the search. This eliminates the error associated with the omission of the probable causes of the fault; • to create a routing for troubleshooting. The map is formed depending on the technological equipment and qualification of the contractor; • taking into account the flow chart to form a request to collect the necessary statistics on the technical condition of individual elements, nodes, mechanisms, verification of which in other conditions is unacceptable or undesirable; • on the basis of probabilistic and statistical information on the object, graphs of optimal moments of technical impacts on the object are formed in order to reduce the cost of ensuring its performance. Hardware complex, which is a computer (stationary, laptop, touch phone, etc.) with an installed interactive system that helps the user to quickly and accurately determine the fault, as well as to propose measures to eliminate it, depending on a

674

S. V. Lyakhov and S. V. Budalin

number of circumstances. In the memory of the central computer introduced pre-prepared digital video containing step-by-step video that shows the relevant component parts of complex technical systems (STS) in the sequence of technological operations and tools. The program of the system is based on block diagrams (question and answer). The peculiarity of working with the system is voice control. Table 1 shows compound classes of IICS and their functions. The control flow is carried out by the following algorithm: 1. 2. 3. 4.

An instance of the Text to Speech class is initialized. An instance of the Recognizer class is created and started. Another question is asked. When the on partial results event (partial information) analyzes the recognized result: if the answer begins with “Yes” or “no”, then the following appropriate question is asked.

To create a neural network Python programming language was used, which has a number of advantages: • • • •

built-in data structures; a simple and convenient syntax; powerful interface; portability of code between platforms: (automatic generation of documentation for modules and the ability to write self-documented programs);

At the same time, there are already multi-level libraries that allow you to manually configure each neuron, and there are high-level libraries that allow you to create a neural network, the second—to load a series of data, the third—to process the results of the study. In three lines, it is possible to write the program allowing to process thousands of data by a neural network and to give out result. For training, the neural network was used for more than ten recorded sounds of engines of cars of different brands. In view of the insignificance of the number of sound for each imposed different effects: • was cut into pieces of different lengths; • turned the melody; • changed the volume and quality of the melody. The effects applied to the existing audio files have increased the training sample by an order of magnitude. Table 1 Composite classes of the interactive information consulting system Program

Function

Main activity Recognizer Recognizer listener Text to speech Question

Contains instances of other classes, basic system methods Convert speech to text Obtaining a control flow in the event the speech recognition Convert speech to text Storage of a tree of questions and answers

Development of Hardware and Software Complex …

675

In the future, a test sample was fed to the input of the neural network, the results of which determined the proportion of correct answers in percentages. The results of the neural network are presented in the form of a matrix, which vertically postponed the expected answers, and horizontally received. Ideally, all answers must match and be placed on the diagonal (Fig. 2). And, the answers do not fall on the diagonal are errors in the processing of information through a neural network. For processing information, neural network sounds of the engines fed to the input in a different sample. Initially, the neural network processed the whole selection of sounds, which was obtained after applying additional sound effects (Fig. 2). The accuracy of information processing in this case reached 94%. After processing the information by the neural network, the responses are summed up and the library itself allows you to determine the accuracy of the developed neural network. The second sample consisted of four complete audio files of the engines of all four cars (Fig. 3). The accuracy of the answers decreased to 91%. In addition, to check the efficiency of the neural network, the previous four sounds were applied to the input but inverted on the contrary. The processing result is shown in Fig. 4. The accuracy of information processing decreased to 90%. The development of neural networks application in the field of TTM diagnostics will allow creating compact devices and hardware complexes with low cost in the mode of constant time, which control the technical condition of individual units and mechanisms of aggregates according to the level of information quality, not inferior to the work of a qualified diagnostic specialist.

Fig. 2 Error matrix classification of engine sounds of the cars for the whole sample (Quality 0.94)

676

S. V. Lyakhov and S. V. Budalin

Fig. 3 Error matrix classification of the sounds of the engine car for 4 whole sounds (Quality 0.91)

Fig. 4 Error matrix classification of the sounds of the car engine 4 integer reverse sound (Quality 0.90)

Development of Hardware and Software Complex …

677

4 Conclusion The use of information technologies to improve the efficiency of TTM use in forestry will ensure: 1. Restoring the health of the host or assembly of TTM in the shortest possible time. 2. The absence of unnecessary empty runs TTM or additional costs for their delivery to the place of service or repair (dealership). 3. Subject to all requirements of the maintenance and repair system, the maintenance of warranty obligations (if any), both the supplier of the machine and the supplier of equipment installed on the machine. 4. The ability to restore the efficiency of the machine outside the production and technical base of the enterprise, in the case of sufficient technical and technological support. 5. Reduce the complexity of the control and increase the accuracy of information about the technical condition of machines and equipment. 6. Reducing the probability of line failure. 7. Reducing the need for disassembly and assembly operations in the diagnosis of individual units of mechanisms and assemblies. 8. Implementation of quality control of the performed maintenance and repair, and also running-in. 9. Accumulation of statistical information on the intensity and development of failures for the formation of laws and mathematical models of changes in the technical condition of components of mechanisms and assemblies. 10. Maintenance of operation of transport and technological machines on their actual technical condition.

References 1. Terentyev AV (2009) Improvement of the methodology for calculating the production program of maintenance of rolling stock. Dissertation, North-Western State Correspondence Technical University 2. Pukhov EV (2013) Improvement of the system of waste utilization by the enterprise of technical service of transport and technological machines of the agro-industrial complex. Dissertation, Voronezh State Forestry Academy 3. Lyakhov SV (2012) Thesis for the degree of candidate of technical sciences. Increasing the efficiency of timber hauling by the fleet of road trains based on the planning of technical and operational indicators. Dissertation, Ural State Forestry University 4. Budalin SV, Lyakhov SV, Nekrasov DN (2011) Calculation of the specific energy consumption of timber removal by road trains. Nat Tech Sci 6(56):572–575 5. Lyakhov SV (2011) Prediction of technical and operational indicators of timber trucks. Nat Tech Sci 6(56):586–590 6. Budalin SV, Lyakhov SV (2010) Assessment of the condition of trucks of the Sverdlovsk region. Ural Transport 1(24):32–34

678

S. V. Lyakhov and S. V. Budalin

7. Lyakhov SV (2012) Determination of the specific energy consumption of removal of forest raw materials and technical and operational indicators of forest truck trains. Nat Tech Sci 1:78–81 8. Nekrasov DN, Lyakhov SV, Budalin SV (2012) Algorithm for the selection of timber rolling stock for the fleet of timber enterprise. Forests Russia and the Economy in Them 1–2(42–43): 67–70 9. Budalin SV, Lyakhov SV (2011) Analysis of the quality indicators of the operation of timber trucks. Nat Tech Sci 2(52):481–485 10. Klyuev V (2005) Nondestructive testing. Manual in 7 t. Under editing member’s RASV t. 7, Moscow 11. Lyakhov SV, Stroganov YuN (2017) To the actuality of application of the methods of investigation of vibroacoustic signals. Int Acad Agrarian Educ 34:10–14 12. Trompet GM, Aleksandrov VA (2017) Renovation of machinery and equipment. In: Materials of the All-Russian scientific and practical conference. Test results of measuring modules of active control, vol 1, pp 177–181 13. Gericke BL, Gerike PB, Kvaginidze VS, Kozovoy GI, Khoreshok AA (2012) Diagnostics of mining machines and equipment. Tutorial, Moscow 14. Pozhidaeva V (2005) Determining the roughness of contact surfaces of the rolling hearings by the method of shock pulses. In: World tribology congress III, Washington, D.C., 12–16 Sept 2005 15. Tse P, Peng V, Yam K (2001) Wavelet analysis and envelope detection for rolling element bearing fault diagnosis—their effectiveness and flexibilities. J Vib Acoust 123:303–310. https://doi.org/10.1115/1.1379745 16. Molchanov VV, Kamnev MI, Bocharov AG (2009) Method of repair and maintenance and hardware and software package for diagnostics and system for quality control of repair and maintenance used in the method. RUS Patent 2357215, 27 May 2009 17. Vanin AA, Razdelkin ME (2016) Control method repair actions on components and assemblies of automotive vehicles. RUS Patent 2582519, 27 Apr 2016 18. Fears AF, Kalik NA (2012) Automated control system of processes and resources for maintenance and repair. RUS Patent 2450304, 10 May 2012 19. Antonenko IN, Kubrin SS, Matyushin VA, Sukmanov AI (2013) Automation subsystem for maintenance, repair, analysis and evaluation of the technical condition of mining machines and equipment. Mining 1:68–74 20. Lyakhov SV, Stroganov YuN, Tokmancev TB (2017) Automation of maintenance and repair of transport and technological machines. Int Acad Agrarian Educ 37:22–26 21. Lyakhov SV, Stroganov YuN, Tokmancev TB (2017) The use of deep neural networks for the problem of classification of sounds. Int Acad Agrarian Educ 36:172–176

Study of Characteristics of Engine Operation in Stress-Testing Mode of Electric Gasoline Pump A. Vozmilov, D. Vlasov and K. Glemba

Abstract We theoretically determined the estimated model for the fuel supply system, which allows us to investigate the relationship of the maximum engine crankshaft speed with the change in the technical condition of individual elements of the fuel system. The change in the pressure and capacity of the fuel system is associated with the change in the flow characteristics of the electromagnetic injectors, filter resistance, and technical condition of the pump. In connection with various mechanisms of failures of the fuel system elements and changes in their technical condition, we considered a simulation model of the fuel system. We presented the results of studying the output characteristics of electric gasoline pumps of the vehicles’ fuel system. We considered two diagnostic methods to determine the technical condition of the pump: dynamic and static. To control the technical condition of the fuel system elements, we developed a testing method based on load and stress testing. For a pump used in a vehicle, the change in its characteristics with the change in the supply voltage typical for these conditions is of a great importance. The paper presents a study of the relationship between the technical conditions of the gasoline electric pump with its quality indicators of the fuel system operation. We determined that it is possible to monitor the technical condition of the pump by measuring the difference of the values of the crankshaft speed when the sparking and fuel supply are off. We presented experimental results by the dynamic method, which has the most informative and information capacity.







Keywords Fuel system Diagnostics Electric gasoline pump Technical condition Diagnostic parameter





Testing


A. Vozmilov (&)
K. Glemba South Ural State University, 76, Lenin Avenue, Chelyabinsk 454080, Russia e-mail: [email protected] D. Vlasov South Ural State Agrarian University, 75, Lenin Str., Chelyabinsk 454080, Russia © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_72

679

680

A. Vozmilov et al.

1 Introduction The modern fuel system (FS) is saturated with electronic, mechanical, and hydraulic elements. The failure rate of the fuel system is 15–40% of the total number of engine failures. FS failures can be caused by: malfunctioning of the electric gasoline pump (EGP), failure of the EGP triggering device, failure of the EGP control circuit wiring or switch, fuel filter clogging, etc. In the majority of cases, the abnormal performance of the fuel system is explained by the following: use of low-grade fuel; low qualification of personnel engaged in repairs, operation and maintenance of cars; use of poor-quality spare parts. As a rule, FS failures are externally manifested as follows: The engine does not develop power; jerks, dips, twitching; unstable idle operation of the ICE; the engine is sharply gaining momentum (gases); it is impossible to adjust the CO value. It is significantly difficult to find a faulty element in the FS, in particular, in the EGP [1–16]. At the same time, the improvement of the power supply system of engines with gasoline injection is aimed at ensuring high environmental performance, but this is possible only due to a precise dosing of the fuel supply in all engine operating modes. Car manufacturers produce EGPs by the sixth diagnosability category, which creates significant problems in assessing their technical condition: It is necessary to dismantle fuel lines and fuel tank components [17–28]. There arises a need to develop new methods and means of diagnosing EGPs without disassembling the pump itself and the fuel lines, fairly quick and with a high diagnostic accuracy.

2 Theoretical Research The purpose of this work is to study the output parameters of electric pumps of vehicles when simulating changes in their technical condition without disassembling the FS using new diagnostic methods. Two diagnostic methods are considered to determine the technical condition of the EGP—dynamic and static. To control the technical condition of the fuel system elements, we developed a test diagnostic method based on load testing and stress testing [1–4, 12–17]. For the EGP used in a vehicle, the change in its characteristics with the change in the supply voltage typical for these conditions is of a great importance [1–14, 24–26]. To justify the diagnostic modes and parameters, we analyzed possible complex output parameters of automobile EGPs. The estimated (theoretical) pump capacity per revolution (working volume) q (cm3/rev) is equal to the volume described by its pistons [1, 7–9, 27]: q¼

pd 2 h
z 4

ð1Þ

Study of Characteristics of Engine Operation …

681

where d is the cylinder diameter, cm; h is the piston stroke, cm; z is the number of pistons. Given that the piston stroke is equal to double eccentricity (h = 2e), we will obtain an expression for the volume described by the pump pistons [1, 7–9, 27]: q¼

2pd 2 e
z 4

ð2Þ

The minute theoretical capacity Q (cm3/min) will be equal to [1, 7–9, 27]: Q ¼ qn ¼

pd 2 e
z
n 2

ð3Þ

where n is the number of the pump’s shaft rotations, min−1. We will determine the estimated theoretical model for the FS, which allows us to investigate the relationship of the maximum crankshaft speed with the change in the technical condition of individual elements of the fuel system [1, 7–9, 27]:

.

ð4Þ

where K—capacity margin factor; i—number of engine cylinders, pcs; QCmax— maximum cyclic delivery at the maximum engine speed, cm3; smax—pulse duration at the maximum cyclic delivery, s; Tmin—minimum recurrence period of cyclic deliveries, s; nmax—maximum speed of the crankshaft engine, min−1; tu—time of the fuel use by the injector, s; tf—time of fueling the rail and fuel lines, s; li— coefficient of the fuel flow through the injector; fi—section area of the injector, m2; qf—fuel density, kg/m3; Pf.r.—fuel pressure in the rail, MPa; Pi.b.—injection backpressure, MPa. Changes in the pressure and flow capacity of the FS are associated with changes in the flow characteristics of electromagnetic injectors, filter resistance, and the technical condition of the pump. In connection with various mechanisms of failures in the fuel system elements and the change in their technical condition, let us consider the FS simulation model. Thus, for example, clogging of the fuel filter can be represented by a series resistance to the fluid flow in the fuel line, and the EGP wear can be represented by a parallel resistance. Let us write this condition for the fuel system, where the total dynamic capacity (ltot.
ftot.) of all elements of the FS will have the following ratio [1, 7–9, 27]: 1 1 1 ¼ þ ltot:
ftot: lp
fp lcons:
fcons:

ð5Þ

682

A. Vozmilov et al.

where lp
fp is the dynamic capacity of the pump, m2; lcons
fcons. is the dynamic capacity of consecutive elements, m2. In its turn, the dynamic capacity of consecutive elements of the fuel system will be determined from the expression [1, 7–9, 27]: lcons:
fcons: ¼ lf
ff þ liEMI
fiEMI

ð6Þ

where lf
ff is the dynamic capacity of the filter, m2; liEMI
fiEMI is the dynamic capacity of the ith electromagnetic injectors, m2. The pressure value Pf.l. (MPa) in the fuel line is determined by the formula [1, 7– 9, 27]: Pf:l: ¼

Q2 qf 2ðltot:
ftot: Þ2

ð7Þ

3 Methods and Results of Experimental Research The technical condition of electric gasoline pumps can be monitored by measuring the difference in the values of the engine crankshaft speed when the sparking and fuel supply are off. The technical condition of the fuel element under study was determined by the ratio of the ICE speed when working with a new and real (diagnosable) EGP. This ratio allows us to determine the level of reducing the supply by the diagnosed EGP, and as a consequence—the degree of its wear [1]. To study the quality of the EGP functioning using the dynamic method, we made an experiment, in which we tested the engine operation on each of the ICE cylinders when the sparking was off in the other three cylinders. We also reduced the fuel delivery rate to increase the sensitivity of the diagnosable parameter. When conducting the experimental research, we used a research stand (Fig. 1a) based on the ZMZ-4062 engine with the connection of the developed device in the form of an “electronically controlled load” of the engine (Fig. 1b). The stand allows us to record the values of the output experimental data online in the form of engine process values when one, two, and three cylinders are off. As a result of the studies, we obtained a characteristic of the change in the ICE crankshaft (CS) speed n (CV) (Figs. 2, 3) [1, 7–9, 20]. Let us consider the first section of the research, when all the four injectors are still on when the sparking is off, quickly emptying the fuel rail, which leads to a decrease in the ICE CS speed. The limits of the change in the CS speed at an 80, 60, and 40% degree of opening the throttle valve with a serviceable EGP (supply voltage U = 14 V) amounted to 2750–3000 rpm, but at a 20% degree of the throttle valve opening—already 2400–2500 rpm. When the throttle is closed below 20%, the EGP is no longer able to provide the necessary fuel supply for the operating

Study of Characteristics of Engine Operation …

683

Fig. 1 a Research stand for testing the ZMZ-4062 engine systems; b the developed device —“electronically controlled load” of the engine

Fig. 2 Change in the ICE crankshaft speed n (min−1) on each of the four cylinders with a serviceable EGP, depending on the degree of the throttle valve opening (%) at the EGP supply voltage U = 14 V: rows 1–4—parameters of each of the 4 cylinders

injector, as a result of which the engine stalls (see Fig. 2) [1, 7–9, 20]. This characteristic highlights the engine operation only with a new serviceable EGP. In the second section of the research, we made a similar experiment with a serviceable EGP, but with a decrease in its supply voltage to the level of 8 V. The characteristic of the ICE operation in this mode is shown in Fig. 3. The limits of the change in the CS speed with a serviceable EGP (supply voltage U = 8 V) at a 80 and 60% degree of opening the throttle valve were 3100–3750 rpm, at 40%— 2800–2900 rpm, and at 20%—800–900 rpm, then, the engine already stalls [1, 7–9, 20].

684

A. Vozmilov et al.

Fig. 3 Change in the ICE crankshaft speed n (min−1) on one of the four cylinders with a serviceable EGP, depending on the degree of the throttle valve opening (%) at the EGP supply voltage U = 8 V: rows 1–4—parameters of each of the four cylinders

When analyzing the experimental data, it can be concluded that, when the engine is loaded, a decrease in the CS speed is observed already at 30% of the throttle valve opening at the EGP supply voltage U = 14 V (see Fig. 2), and at 40% of the throttle valve opening in case of a decrease in the EGP supply voltage to 8 V (see Fig. 3). The obtained research data confirm the interrelation of the change in the maximum engine crankshaft speed with the change in the technical condition of individual elements of the fuel system (in particular, the condition of the EGP with a reduced power level), which indicates the adequacy of using the estimated theoretical model (4).

4 Conclusion The described methods allow us to determine the technical condition of the EGP. Firstly, in the dynamic mode, the operation of the EGP is tracked due to the change in the maximum possible engine crankshaft speed when the sparking is off in the mode of regulating the EGP supply voltage. Secondly, in the static mode, the operation of the EGP is considered by the developed pressure when the ICE is out of service. These methods are characterized by a significant efficiency of the process of diagnosing the fuel system.

Study of Characteristics of Engine Operation …

685

5 Findings We theoretically determined and experimentally confirmed the estimated theoretical model for the FS, which allows us to investigate the relationship between the maximum crankshaft speed and the change in the technical condition of individual elements of the fuel system. To determine the technical condition of the EGP, it is proposed to use such a diagnostic feature as the limit of the change in the ICE speed. We established that the lower the level of the fuel supply to the FS, depending on the performance characteristics of the EGP, the lower the limits of the change in the ICE speed. The experimental research has shown that the dynamic method used to verify the EGP performance has the maximum informative and informational capacity.

References 1. Gritsenko A (2014) Development of methods for test diagnostics of the efficiency of power and lubrication systems of internal combustion engines. Dissertation, Chelyabinsk 2. Gritsenko A et al (2011) Method for diagnosing the fuel system of internal combustion engines of light vehicles. Bull Chelyabinsk State Agrarian Univ 59:30–32 3. Gritsenko A et al (2013) Diagnosis of electric gasoline vehicle pumps. Mechanization Electrification Agric 4:22–23 4. Gritsenko A et al (2014) Diagnosis of the ICE power supply system. Mechanization Electrification Agric 1:24–26 5. Bakaikin D (2013) Diagnosing electromagnetic injectors of gasoline engines of vehicles operated in agriculture. Dissertation, Chelyabinsk 6. Gritsenko A et al (2013) Optimization of the process of diagnosing autotractor equipment by minimizing costs. Bull Chelyabinsk State Agrarian Univ 63:42–46 7. Stola F, Paolino D, Parotto M, Troina F (2015) Electric feed pump simulation and control for fuel saving and system cost reduction. SAE technical paper 2015-24-2421 8. Ferreira A, Galhardo A, Busani R, Souza R (2016) Development of 1D simulation model of electric fuel pump for flex-fuel application. SAE technical paper 2016-36-0135 9. Edward R, Radu R, Mircea Z (1999) Simulation of the behavior of a flow control device for the high pressure fuel injection pump of a spark ignition engine. SAE technical paper 1999-01-0562 10. Gritsenko A et al (2015) Development of the method and means of diagnosing electric gasoline pumps of the ICE fuel system. Transp Sci Technol Manage 1:40–44 11. Plaksin A et al (2014) Development of methods for test diagnostics of the performance of fuel and lubrication systems of internal combustion engines. Agrarian Bull Urals 7(125):51–58 12. Plaksin A et al (2014) Diagnosing electric gasoline pumps by complex output parameters. Fundam Res 11(part 12):2610–2614 13. Gritsenko A et al (2014) A new method, tool and software environment for testing the vehicle EGP. In: Proceedings of VolgSTU. Energy conversion processes and power plants, vol 6 (145)18, pp 53–56 14. Plaksin A et al (2014) Diagnosing electromagnetic injectors by changing the qualitative composition of the fuel mixture. Fundam Res 11(part 11):2380–2384 15. Gritsenko A et al (2011) Method for diagnosing the engine fuel system. RF Patent 2,418,190, 10 May 2011

686

A. Vozmilov et al.

16. Gritsenko A et al (2011) A method for diagnosing electric gasoline pumps of a vehicle fuel system. RF Patent 2,477,384, 10 Mar 2013 17. Gritsenko A et al (2012) Diagnosing ICE systems on test static modes. Bull Chelyabinsk State Agrarian Univ 61:31–38 18. Gritsenko A et al (2012) Determination of the effectiveness of using technical diagnostic tools taking into account the frequency of failures in ICE systems. Bull Chelyabinsk State Agrarian Univ 60:45–48 19. Gritsenko A et al (2011) Development of effective tools and methods for diagnosing internal combustion engines of vehicles. Bull Chelyabinsk State Agrarian Univ 58:111–117 20. Gritsenko A et al (2011) Diagnostics of internal combustion engine systems using the non-brake method with redistribution of cylinder loads. Bull Chelyabinsk State Agrarian Univ 58:108–110 21. Gritsenko A et al (2011) Justification and development of tools and methods for diagnosing internal combustion engines of vehicles. In: International scientific and technical conference “Achievements of Science—Agro-Industrial Production.” Chelyabinsk, Chelyabinsk State Agrarian University, pp 6–11 22. De Cesare M, Parotto M, Covassin F, Sgatti S (2013) Electric low pressure fuel pump control for fuel saving. SAE technical paper 2013-01-0339 23. Fávero C, Souza L, Pacheco F, Pereira M (2013) Corrosion on electric fuel pump housing at durability test in test bench. SAE technical paper 2013-36-0622 24. Häntsche J, Krause G, Velji A, Spicher U (2005) High pressure fuel pump for gasoline direct injection based on ceramic components. SAE technical paper 2005-01-2103 25. Sarkar S, Kumar S, Singhal A, Kohli S et al (2015) Common design of jet pump for gasoline and diesel based vehicles. SAE technical paper 2015-01-0458 26. Rovai F, Ferrarese A, Sinatora A (2004) Electrical automotive fuel pump resistance against E60 gasoline-ethanol blend. SAE technical paper 2004-01-3277 27. Kunkel R (1999) New fuel pump technology. SAE technical paper 1999-01-0331 28. Felton G (1996) Compact rotary spill pump with electronic control for high speed direct injection engines. SAE technical paper 960865

Technical Level Analysis of Structures of Quarried Excavators H. N. Sultonov and K. Z. Tilloev

Abstract The technical level analysis of structures of products of Russian and foreign producers of career excavators taking into account the demand for this equipment in various industries, depending on their technical parameters, is executed. The following criteria are used to identify trends: available power, theoretical energy intensity, and materials’ consumption. These indicators do not take into account the length of the arrows of compared excavators and thus the amount of work to move the rock. However, the weight of the excavator is proportional to the length of the boom and the bucket capacity, which allows taking the relative weight and energy as an objective criterion characterizing the technical level of the machine. To do correct assessment the technical level of the structures of foreign and Russian excavators, with a bucket capacity from 4.6 to 20 m3, a statistical analysis was performed using the method of Professor E. S. Wenzel. This method reflects the current understanding of the statistical analysis of random variables. The approximating dependences were obtained by the “least squares deviations” method with the maximum value of the Pearson criterion—“v2”.




Keywords Career and construction excavators Technical level power Theoretical energy intensity Materials consumption








Available

1 Introduction In the world, practice of formation of excavator–automobile complexes, the choice of capacity of the bucket of the excavator, and loading capacity of the dump truck carried out on the principle of optimum loading in 3–5 cycles of excavation. H. N. Sultonov (&) Tajik Technical University Named After Academic M.S. Osimi, 10, Street ac. Rajabovho, Dushanbe 734042, Tajikistan e-mail: [email protected] K. Z. Tilloev South Ural State University, 76, Lenin prospekt, Chelyabinsk 454080, Russia © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_73

687

688

H. N. Sultonov and K. Z. Tilloev

Since the 2000s, the production of hydraulic excavators for mining and construction enterprises is 3–5 times higher than the production of bucket loaders. However, in total sales of hydraulic machines more than 70% are excavators with a bucket of less than 20 m3. The limited usage of large standard sizes is explained by the fact that hydraulic excavators lose their technological advantages with a bucket capacity of more than 30 m3 that is short service life (approximately by 2–3 times) and high cost of excavation compared with bucket loaders. What makes a choice in favor of powerful bucket [1, 2]. The main factors that influence the demand of excavator equipment are the dynamics of the volumes of mining and construction work, the development of technological transport in quarries and construction sites, and the effectiveness of the use of excavators of various sizes. The aim of this work is to analyze the technical level of the structures of mining excavators of domestic and foreign companies.

2 Structural Analysis Currently, the production of single-bucket excavators in the CIS (Commonwealth of Independent States) countries is mainly focused on the following enterprises: Uralmashzavod (UZTM) and IZ-KARTEKS LLC (Izhorskiye Zavody), (Russia), as well as Novo-Kramatorsk (NKMZ), Donetsk machine-building plants (DMZ) (Ukraine). Today in the Russian Federation, OJSC “Uralmashzavod” and LLC “IZ-Kartex” named P. G. Korobkova conduct the development and production of quarry excavators. They have created projects of models of career crawler excavators of the new generation EKG-1500R (Fig. 1) and EKG-1500 K with the rack and rope pressure of the handle (this is indicated by the letters “R” and “K”). The structures of these machines take into account a number of suggestions of customers, including the possibility of installing buckets of different capacities, in accordance with geological conditions and available vehicles in the pit [2–7]. Russia’s largest company “IZ-Kartex” goal is the active promotion of excavators’ class 18–60 m3 at the enterprises of Russia and CIS countries at the expense of raising the technical level of machinery, the improvement of consumer properties, and lower capital acquisition costs and operational expenses of the Russian excavators [8, 9]. If we compare the excavators of the new line of LLC “IZ-Kartex” with foreign analogs of production of world leaders Bucyrus (Cat) and P&H Mining Equipment [10–16], the reduction of production costs in the application of the excavators of LLC «IZ-Kartex» will be achieved by: • saving of capital costs (by 62–90%), resulting from the difference in the purchase price of the excavator EXW;

Technical Level Analysis of Structures of Quarried Excavators

689

Fig. 1 Excavator EKG-1500R

• (50–70%), the difference in the cost of delivery (7–15% of the price of the excavator EXW) and the absence of import customs duty (5% of the price of the excavator EXW); • reduction of operating costs by 30% or more when using consumables (teeth, ropes, fuels, and lubricants, etc.) of Russian production and spare parts at a lower (20% or more) cost.

3 Method of Estimation of Dependencies of Criteria of Technical Level of Structures For technical and economic evaluation of models with different weight, bucket capacity, boom length, etc., the indicators or quality factors are used. The simplest indicator of the relative weight is the ratio of the mass of the excavator structure to the capacity of its bucket: G/E (t/m3), the next indicator is relative power indicator, it is the ratio of the drive power to the bucket capacity: N/E (kW/m3). To assess the quality of quarry excavator designs with a bucket capacity of 4.6– 20 m3 with hydraulic and mechanical shovels of the main manufacturers in the Russian Federation, JSC “Uralmashzavod” and LLC “IZ-Kartex” named P. G. Korobkova the criteria of technical level of their structures are used [17–19]:

690

H. N. Sultonov and K. Z. Tilloev

• the available power in the form of the ratio of the installed capacity of the power plant of the excavator—N to its weight—G; WN ¼

N G

ð1Þ

• the theoretical energy intensity of the excavator during its working cycle WA in the form of the ratio of the magnitude of the physical work performed during the cycle to the geometric volume of its bucket—E; WA ¼

N
tc E

ð2Þ

where tc—the duration of the working cycle of the excavator when the passport angle—ur rotation offload equal 0.5p, s. • specific material intensity as a ratio of theoretical energy intensity—WA and available power—WN. WM ¼

G
tc 3:6  103
E

ð3Þ

Taking into account the demand for correct assessment of the technical level of structures of Russian and foreign manufacturers of excavators with bucket capacity from 4.6 to 20 m3, we perform a statistical analysis of the values: available power —WN , theoretical energy consumption of the excavator-WA , specific consumption of materials-WM , the weight of this structural methodology, developed by Professor E. S. Wenzel, reflecting the modern idea of statistical analysis random variables [19]. The method includes: • an analytical approximation of the dependence of available power WN(G), theoretical energy intensity of excavation WA(G), and specific material intensity WM(G) of the excavator from its weight on the maximum size of Pearson’s criterion—v2; • calculation of mathematical expectations M(WN); M(WA); M(WM), dispersions D (WN); D(WA); D(WM), and coefficient of variation kv(WN); kv(WA); kv(WM), by known formulas:

MðjÞ ¼ DðjÞ ¼

Pk

Pk i¼1

i¼1

n


Wi

;

j ¼ 1; 2; . . .; n;


ðWi  Mj Þ2 ; n1

i ¼ 1; 2; . . .; k

j ¼ 1; 2; . . .; n;

i ¼ 1; 2; . . .; k

ð4Þ ð5Þ

Technical Level Analysis of Structures of Quarried Excavators

kv ðjÞ ¼

691

rðjÞ MðjÞ

ð6Þ

where r(j) mean-square deviation is defined as: rðjÞ ¼

pffiffiffiffiffiffiffiffiffi DðjÞ

ð7Þ

4 Results and Discussion Initial data were determined on base on the method of correct assessment of the technical level of structures of Russian excavators with the bucket capacity from 4.6 to 20 m3 developed by Professor E. S. Wenzel. The given method allowed to obtain analytical dependences of the technical level criteria on the weight of excavators on the maximum size of Pearson’s criterion—v2 [19, 20] (see Table 1). The analytical dependencies shown in Table 1 are presented in Figs. 2, 3 and 4, respectively. Table 1 Dependence of the technical level criteria on the weight of excavators The criterion of the technical level of a career excavator

Approximation formula

Pearson criterion —v2

Available power, kW/t Theoretical energy intensity of work, kNm/m3 Specific material intensity of structures, T/m3/h

WN(G) = 0.0018G – 6  10-7G2 + 0.9085 WA(G) = 0.0507G0.3835

0.972 0.353

WM(G) = 0.0549G0.3056

0.237

Fig. 2 Dependence of available power WN on the weight of the excavator G

692

H. N. Sultonov and K. Z. Tilloev

Fig. 3 Dependence of theoretical energy intensity WA on the weight of the excavator G

Fig. 4 Dependence of specific material intensity WM on the weight of the excavator G

The results of the calculation of the statistical parameters of excavators were obtained by Formulas (4), (5), (6), and (7) are presented in Table 2.

Technical Level Analysis of Structures of Quarried Excavators

693

Table 2 The results of the calculation The criterion of the technical level of a career excavator

Mathematical expectation— M(j)

Dispersions —D(j)

Mean-square deviation—r (j)

Coefficient of variation— kv(j)

Available power, kW/t Theoretical energy intensity of work, kNm/m3 Specific material intensity of structures, T/m3/h

1.695 0.544

0.110 0.034

0.332 0.184

0.196 0.338

0.367

0.015

0.124

0.338

5 Conclusion The analysis of dependence of criteria of technical level on weight of the excavator (Table 1) and the analysis of results of calculation of their statistical parameters (Table 2) testify that: • the magnitude of the available power, the theoretical energy intensity of work, and material intensity of structures excavator are not linearly increased with the increase of its weight by a power-law dependence (with Pearson’s criteria equal v2(WN) = 0.972; v2(WA) = 0.353 and v2(WM) = 0.237, respectively); • the greatest value of the criterion characterizes: 1. Available power excavator has a model of ECG-20A (WN = 2.16 kW/t), and a model EKG-5A (WN = 1.27 kW/t) has the smallest value with mathematical expectation M = 1.695 kW/t and coefficient of variation kv = 0.196. 2. Theoretical energy consumption of the excavator operation is the model of EKG 4US (WA = 0.93 kNm/m3), and a model EKG-5A (WA = 0.32 kNm/m3) has the smallest value with mathematical expectation M = 0.544 kW/t and coefficient of variation kv = 0.338. 3. Specific material intensity is the model of EKG 4US (WM = 0.66 T/m3/h) and a quarried excavator model ECG-4,6B (WM = 0.21 T/m3/h) has the smallest value with mathematical expectation M = 0.367 T/m3/h and coefficient of variation kv = 0.338.

References 1. Paladieva NI (2011) Modern trends in the market of excavators for mining enterprises. Min Equip Electromech 5:10–14 2. Kantowicz LI, Saidaminov IA (2001) Experience, current state and prospects of development of structures of mine equipment with hydrostatic transmissions. Min Inf Anal Bull Moscow 11:10–14 3. Paderni RYu (2011) Mechanical equipment career: the Textbook for high schools, 7th edn. Mining of Media Groups, Moscow

694

H. N. Sultonov and K. Z. Tilloev

4. Ivanov IYu, Komissarov AP (2011) Assessment of the intensity of the working process of hydraulic excavator with external circuit loads. Min Equip Electromech 9:45–47 5. Kozlovsky E, Dikonenko E et al (2004) Russian coal encyclopedia. Cartographic fabric VSEGEI, St. Petersburg 6. Vul Yu et al (1978) Single-bucket excavators NKMZ. Nedra, Moscow 7. Anistratov YI, Ilyin SA (1964) Open pit mining in difficult conditions. Nedra, Moscow 8. Trubetskoy KN, Potapov MG, Vinitsky KE et al (1994) Handbook. Open pit mining. Mining Bureau, Moscow 9. Suprun VI et al (1996) Perspective equipment and technology for the production of open pit mining. MGGU, Moscow 10. Ng F, Jennifer A (2016) Harding, Jacqueline Glass Improving hydraulic excavator performance through in line hydraulic oil contamination monitoring. Mech Syst Signal Process 83:176–193. https://doi.org/10.1016/j.jfranklin.2016.04.07 11. Ge L, Quan L, Zhang X, Zhao B, Yang J (2017) Efficiency improvement and evaluation of electric hydraulic excavator with speed and displacement variable pump. Energy Convers Manage 150:62–71. https://doi.org/10.1016/j.enconman.2017.08.010 12. Keith H (2004) Extreme mining machines: stripping shovels and walking draglines. MBI Publishing Company, USA 13. Alexandrov IV, Mikhailov NA, Dzerzhinsky VA (2004) NKMZ—70: from idea to market implementation. Min J 8:28–30 14. United Heavy Machinery Plants (2017). http://www.omz.ru. Accessed 12 Nov 2017 15. Mining Industry Journal (2018). https://mining-media.ru/ru/. Accessed 20 Jan 2018 16. Malyshev Y, Anistratov KYu et al (2005) World mining industry 2004–2005: history of achievements, prospects. Gornoe delo, Moscow 17. Kvaginidze VS, Petrov VF, Koretsky VB (2007) Operation of quarry equipment. Mountain Book, Moscow 18. Kvaginidze VS, Antonov YuA, Koretsky VB, Chupakhina N (2009) Excavators in the quarries. Design, operation, calculation. Mountain book, Moscow 19. Wentzel ES (1964) Probability theory. Science, Moscow 20. Sultonov HN (2016) Justification and selection of dynamic parameters of the single-engine drive of the running gear of the career excavator. Candidate of technical science dissertation research, Moscow

Reliability Analysis of Bus Steering System A. N. Savkin, K. A. Badikov and A. A. Sedov

Abstract Most of steering problems were related with a pitman arm and a steering joint. For these parts, using the example of two buses the bending stress and safety margin were determined with the help of virtual 3D models. A dangerous section and maximum loads were determined, the calculation of bending under static loading was carried out. The proposals were drafted that aimed at improving the operational reliability of the steering. The problem of road safety was one of the most significant in modern conditions. An important contribution in safety was made by the reliability of vehicles, which, in turn, was associated with the development of their design, and also the technical condition. An especially important thing was to ensure the reliability of vehicles that carry passengers. Keywords Steering joint


Steering arm
Bending stress
Strength calculation

1 Introduction The problem of road safety is one of the most significant in modern conditions. A significant contribution to safety is made by the reliability of vehicles, which, in turn, is related to the perfection of their design and technical condition. Especially important is to ensure the reliability of vehicles that carry passengers. In this study, the rolling stock of one of the transport company of the Volgograd region (Russia), carrying out urban and suburban passenger transportation, was analyzed. Motorcade exploits mainly two models (model 1 and model 2 will be indicated below). It was found that a significant part of the total number of bus’s malfunctions is made up of malfunctions by the steering system. It is known that the steering system is the second most important system, after the brake system, which determines the active safety of the vehicle. At the same time, if the brake system of modern cars, in accordance with safety requirements, A. N. Savkin
K. A. Badikov (&)
A. A. Sedov Volgograd State Technical University (VSTU), 28, Lenina St, Volgograd 400005, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_74

695

696

A. N. Savkin et al.

always includes several independent circuits and systems that allow the car to stop even in emergency conditions, there are no duplicate elements for steering [1, 2]. An analysis of the reliability of the buses in use revealed an increase in the number of steering malfunctions in recent years. This could be due to an increase in the transport company’s fleet or an increase in the annual mileage of vehicles.

2 Materials and Methods At the first stage of the study, elements of the steering system were identified that most often failed during the operation of buses [3]. These elements include the pin steering joint and the steering arm. In the literature [4, 5] indicates that the most dangerous section at the steering arm is at section A–A (Fig. 1). The tests carried out showed that on the considered models of buses this dangerous section is located near the hole under the output shaft of the steering wheel reducer (Figs. 2 and 3). The second element, which was examined in detail in this study, is the steering joints. Joints for steering arm and links do not require maintenance due to the sealing cover, filled with a special plastic lubricant. The dimensions of the joints are determined mainly by the diameter of the ball head, which depends on the magnitude of the forces acting on the arm and link in the direction perpendicular to the pin, and on the required angles of deflection. The main part of the steering joint is the pin, the incomplete spherical part (head) of which is located inside the body and is covered by an insert made of a polymeric material [6, 7]. Therefore, in order to reduce wear on the latter, increased

Fig. 1 Design scheme of the steering arm

Reliability Analysis of Bus Steering System

697

Fig. 2 The results of the strength calculation of steering arm of the bus model No. 1 in the APM system FEM for Compass-3D

Fig. 3 The results of the strength calculation of the steering arm of the bus model No. 2 in the APM system FEM for Compass-3D

698

A. N. Savkin et al.

requirements are placed on the surface quality of the ball-head (depending on the car model, the roughness parameter Ra is 0.20–0.45 µm). The most common method of surface plastic deformation of the ball-heads of pins is a planetary rolling. However, this method has the disadvantage of uneven processing of surface areas of the ball-head [8]. The reliability and durability of the steering joints are influenced by the mechanical properties of the materials from which the liners and pins are made [9–11], as well as the stability of their performance over time. According to the data presented in the literature [5, 12], the most dangerous sections for the pins are the sections A–A and BB:BB—the joints of fillets with a ball head and A–A are the section before the arm (Fig. 4). The force experienced by the pin of the steering joint of the arm of the model No. 1 is equal to 6235 N, and that of the model No. 2 is 3485 N. The calculation results are presented in Figs. 5 and 6. To ensure the possibility of performing calculations, a preliminary chemical analysis of the alloy from which the pins were made was carried out using the optical-emulsion method [12]. As a result of the analysis, it was found that the steering joints of the buses under study do not have heat treatment and are made of VSt6ps (BCт6пc) normal quality structural carbon steel, the maximum allowable bending of which is 170 MPa, then for model No. 1 bus at the joint of heads and fillets it does not meet the bending strength requirements. The maximum value of the bending stress in section A–A is 207 MPa [13, 14], therefore, it exceeds the permissible value by 22%, the shear stress is 1%, and in section BB does not exceed the permissible value. The model No. 2 of the excess stress in the considered sections of the steering joint is not observed. However, here the cross section experiencing higher loads is the BB cross section. According to [2], a static calculation was also carried out on the bending of the studied ball joints in dangerous sections using the formula

Fig. 4 The design diagrams of steering arm pin: a bus Model No. 1, b bus Model No. 2

Reliability Analysis of Bus Steering System

699

Fig. 5 The results of the calculation of the steering finger of the bus model number 1 in the system of strength analysis APM FEM for Compass-3D

Fig. 6 The results of the calculation of the steering finger of the bus model number 2 in the system of strength analysis APM FEM for Compass-3D

rmax ¼

32
Pc
k  ½r ¼ 170 MPa p
dn3

ð1Þ

where k and m are the distances from the dangerous section to the center of the head of the pin, m; Pc—force on the ball joint of steering arm, H; dn is the diameter of the pin in a dangerous section.

700

A. N. Savkin et al.

For the first bus model: rmax AA ¼ rmax BB ¼

32
6235
0:011 3:14
ð0:015Þ3 32
6235
0:021 3:14
ð0:02Þ3

¼ 207  106 Pa ¼ 207 MPa [ ½r ¼ 170 MPa; ¼ 167  106 Pa ¼ 167 MPa  ½r ¼ 170 MPa:

For the second bus model: rmax AA ¼ rmax BB ¼

32
3485
0:014 3:14
ð0:027Þ3 32
3485
0:032 3:14
ð0:031Þ3

¼ 25  106 Pa ¼ 25 MPa  ½r ¼ 170 MPa; ¼ 39  106 Pa ¼ 39 MPa  ½r ¼ 170 MPa:

3 Conclusions The obtained results confirm the calculation of 3D modeling for steering joints. They are valid both for the arm’s support and for the steering supports, since the forces acting on them are the same. Using on buses of steering joints made without heat treatment causes accelerated wear of rubbing surfaces: ball-heads, inserts and the inside of the hull and, as a result, additional buses get off the routes. Worn ball joints contribute to the deterioration of controllability, and therefore, reduce the active safety of buses [15–21]. The material the steering joints are made is an important factor that largely determines their operational and other properties, such as tensile strength, corrosion resistance, wear resistance, fatigue strength. From the above, in turn, depends on the resource of the entire steering. The calculation showed that the steel VSt6ps without heat treatment for the ball-head of pin does not allow to achieve an acceptable level of operational reliability according to the criteria of strength and durability. Thus, in the course of the study, it was found that the two considered models of buses have critical elements of the steering system. There are the steering arm and pins of steering joints. The low strength of these elements is associated with the use of steel without heat treatment. To improve the reliability of the system under consideration, measures of a predominantly constructive nature should be used: change the material, add heat treatment or another method of hardening.

Reliability Analysis of Bus Steering System

701

References 1. Lysov MI (1972) Car steering. Engineering, Moscow, p 344 2. Narbut AN (2007) Cars: workflows and payment mechanisms and systems: a textbook for students. Executive Procedure Institutions. Publishing Center “Academy”, Moscow, 256 p 3. Klepik NK (1995) Statistical treatment of experimental problems in the road transport. In: Procedures manual. Volgograd State Technical University, Volgograd, 95 p 4. Standart 52302-2004 (2005) Vehicles. Handling and stability. Publishing House of Standards, Moscow, 27 p 5. Bukharin NA, Prozorov VS, Schukin MM (1973) Cars. Design load conditions, workflows, strength units of the automobile. In: Procedures manual for schools. Engineering, Leningrad, 504 p 6. Badikov KA, Tyshkevich VN, Chernova GA (2013) Calculation of endurance ball stud steering bus “Volzhanin-32901”. Mod Probl Russ Transp Complex 4(4):172–177 7. Pavlyuk AS, Safronov S (2013) Influence of design parameters of the steering gear to the steering mobile machine. Polzunovsky Gaz 4(3):181–184 8. Standard 23207-08 (194) System maintenance and repair of equipment. Selecting and setting parameters maintainability. General requirements. Publishing House of Standards, Moscow, 17 p 9. Habardin AV, Habardin VN, Kuzmin AE, Chubareva MV, Gorbunov TL (2011) Free wheeling steering wheel as a parameter of a technical condition of a steering. Bull Irkutsk State Agric Acad 43:154–160 10. Chernova GA, Badikov KA, Tyshkevich VN (2014) Evaluation of the strength of the ball joint steering bus. Problems of technical operation and service centers of the rolling stock of automobile vehicles: Sat. scientific. tr.: Based on the materials. In: 72th scientific-method. and scientific-issled. Conference MADI, Moscow State Automobile and Road. Technical University (MADI), pp 98–102 11. Larry J (1976) Segerlind: applied finite element analysis, pp 232–239 12. Chernova GA, Tyshkevich VN, Badikov KA, Moseec Yu, Zabolotnyi RV (2014) Evaluation of the strength of steering control bus. In: Proceedings VSTU. A series of “Land transport systems.” vol 8: Hi. Sat. scientific. Art 3(130):47–50 13. Standard 23207-78 (1981) Fatigue resistance. Basic terms, definitions and symbols. Publishing House of Standards, Moscow, 49 p 14. Zaslavsky BV (1986) Short Course materials resistance. Tutorial for aviation specialties. Engineering, Moscow, p 328 15. Migal VD (2013) Technical diagnostics of cars: a handbook in 6 volumes, vol 1, Manufacturing defects and operational faults, 2nd edn, Sr. OOO “CPI” Mask, Moscow, 373 p 16. Migal VD (2013) Technical diagnostics of cars: a handbook in 6 volumes, vol 3. Methods of diagnosing, 2nd edn, Sr. OOO “CPI” Mask, Moscow, 548 p 17. Vernyaev MA, Filkin NM, Shakurov DC (2012) Calculations ball joint suspension of the car. Russian forests and farming in them 1–2(42–43):22–23 18. Abaimov RV (2007) Diagnosis of road transport: Sat. Descriptions of the laboratory. Papers to prepare graduates/comp.; SSI, Syktyvkar, 60 p 19. Menshenin GG (2006) Plastic bags PA workshop on reliability theory and diagnosis part II. In: Procedures manual. VSTU, Volgograd, 230 p 20. Terentiev VF (2002) Fatigue strength of metal alloys. Internet Eng, Moscow, p 288 21. Chernova GA, Tyshkevich VN, Badikov KA, Moiseev YuI, Zabolotniy RV (2014) Estimation of operation capability of the ball pin of the “Volzhanin-32901” bus with relation to fatigue strength. J Mach Manuf Reliab 43(4):294–297

Process Research of Wheel-Rail Mining Machines Traction A. M. Keropyan, D. A. Kuziev and A. E. Krivenko

Abstract During the process of the executed conditions of interaction between working surfaces of wheel-rail mining machines bandages and rails the nature of deformations of interacting rough surfaces ledges in their spot of contact was defined. It is established, that in the interaction zone of a couple of wheel-rail industrial transport peculiar to plastic saturated contact processes occur. At plastic deformation, under the influence of the considerable loadings which are characteristic of the wheel-rail traction cars for industrial use, after all roughnesses of this planimetric platform have received considerable deformation, their base deformation occurs, having a plastic character. These deformations of roughness in the contact zone of their working surfaces contribute to an increase in the true area of their contact and, therefore, leads to ascending traction effort used in mining machinery. Ultimately, available reliable information about deformation characteristics of the processes, occurring in a spot of contact of the interacting surfaces of couple a wheel—rail will afford predicting theoretically expected calculated value of the realizable coefficient coupling and, therefore, the developed traction effort of mining machines. As a result of the executed research it is established, that when grinding the rails to ensure the height of asperities of the interacting surfaces of Rz of 20 lm the settlement coefficient of friction can be increased to 12% and, therefore, the connecting coefficient of a couple wheel-rail and also traction effort the wheel-rail mining machine can be raised.







Keywords Contact between solids Rough surfaces Friction coefficient Interaction zone Contact spot Plastic contact Traction effort













1 Introduction Contact between the surfaces of solids is discrete in all cases and the real surface area of altered surfaces represents an insignificant part of the geometric contact area —a fact that is applied by Russian and foreign experts when considering friction A. M. Keropyan (&)
D. A. Kuziev
A. E. Krivenko NUST “MISIS”, 4, Lenin Street, Moscow 119049, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_75

703

704

A. M. Keropyan et al.

Fig. 1 Scheme of contact between two rough surfaces: 1—nominal, 2—planimetric, and 3—actual

and wear calculations [1–3]. Figure 1 depicts a scheme of contact between two rough surfaces. Planimetric, nominal (geometrical), and actual (physical) zones of contact are distinguished in Fig. 1 [3]. The contact spot of two rough surfaces first occurs not on the highest protrusions, but on those of them that are opposed by such a protrusion that the sum of the heights of the protrusion of the first surface and the protruding protrusion of the second will be the largest. Increasing the load means that all resisting ledges, including those with the lowest heights, will come into contact. The area of contact formed will thus be comprised of platforms located at different elevations, contacting at different angles. Nevertheless, it should be noted that the difference between the actual size of the contact points and their projection onto the plane of the supporting surfaces is small, since the angles between these surfaces do not exceed 3–10° and therefore the difference in the size of these surfaces does not exceed 1.5%. Therefore, for calculation purposes it is possible to assume that all contact surfaces lie on the same plane [1]. Contact under large loads—most probable to occur at interactions between the wheels of railway transport vehicles and tracks, has the following features. If the pressure becomes so great that almost all projections come into contact, then there is a subsequent increase in the contact area between separate ledges. One necessary condition of this process—during plastic contact when all ledges come into contact —is equality of planimetric and nominal areas, Ac = Aa. However, performance of this condition isn’t enough, loading has to be so big that the size of rapprochement exceeded the value calculated as the difference between the average and maximum roughness. For surfaces with high processing purity, the difference between average and maximum heights of roughness is usually greater—practically all ledges will participate in contact only when under very high pressures. This condition is most easily reached for surfaces processed by turning, planning, or milling, that is, with a purity of processing from Rz 20 to Rz 2.5 (3–6 class) [3].

2 Objects and Research Methods Numerous theoretical and experimental studies have established that the size of the friction coefficient between contacting surfaces depends on many of the physical and mechanical properties of the surfaces involved—including their roughness [1, 2, 4].

Process Research of Wheel-Rail Mining Machines Traction

705

For a convenient means of assessing the influence of surfaces roughness [1], the concept of a complex indicator of roughness is introduced ∇: r¼

hmax 1

Rbm

;

ð1Þ

where R is the settlement curvature radius of roughness tops, determined as an average geometrical curvature radius of roughness tops from a profilogram of working surfaces in both longitudinal and cross directions: R¼

pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi rlong rcross ;

ð2Þ

where hmax is the height of greatest roughness; and b and m are the parameters of a basic curved rough surface. Research [2, 4], both theoretical and experimental, has established that at the contact between rough surfaces the ledges of the firmer surface take root in the softer surface. Considering that the hardness of the steel used in railway tracks is slightly higher than the hardness of a tires material, further research is required on steel track samples. Numerous research has established that contact between two surfaces with micro-irregularities is discrete in character—caused by interactions between separate ledges with micro-roughness. The frictional force occurring on each protrusion in areas of surface contact, according to [2, 4], can be presented thus: Fi ¼ Fi mol þ Fi mec ;

ð3Þ

where Fi mol is the molecular component of frictional force and Fi mec its mechanical component. According to [1, 5] the arrangement of ledges in a zone of contact has a casual character and corresponds to the normal distribution law. In this case, the FR of frictional forces will be equal, in any particular zone of contact, to the sum of the forces acting on individual points of roughness: FR ¼ RFi þ RFi mol þ RFi mec ;

ð4Þ

Thus, the total frictional force depends on the actual contact area between interacting surfaces [1, 5]. Consequently, the traction achieved between mining vehicle wheels and track depends on the size of the contact area between interacting surfaces. As a result of theoretical and experimentally confirmed research [1, 2, 6–9], it has been established that when contact occurs between rough surfaces, it takes one of three forms: plastic saturated, elastic non-saturated, and elastic saturated. Even though in many instances (roughness of interacting surfaces, high forces found at contact zones between wheels and tracks, etc.) plastic contact takes place, it is possible to carry out additional calculations for other forms of contact in order

706

A. M. Keropyan et al.

to confirm this hypothesis. According to [3], elastic contact will take place when an inequality exists between contacting surfaces: a\Km rs

ð1  l 2 Þ ; E

ð5Þ

where a is the tilt angle of roughness (rad) a = 2.3° [3, p. 75]; Km is a coefficient varying from 2.5 to 5.0; and µ = 0.3—Poisson’s coefficient. For strengthened steel rs = 1020 MN/m2 (a fluidity limit) and E = 2.1  105 MN/m2 (the given module of elasticity of contacting bodies). For convenience of comparison we can convert a to radians—a = 2.3/ 5.3 = 0.4 rad. We can now calculate the right-hand side of Eq. 5 for two values of Km (2.5 and 5.0): At Km ¼ 2:5; At Km ¼ 5;

Km rs Km rs

ð1  l 2 Þ ¼ 0:011; E

ð6Þ

ð 1  l2 Þ ¼ 0:022; E

ð7Þ

Calculations show that none of the Km values tested satisfied Eq. 4, therefore, we can assume that contact takes place in the form of plastic saturation at zones of interaction between the wheels and track. Characteristic values of coefficients b and m for different values of surface finish are given in Table 1 [10]. Thus, plastic contact should be considered for the wheel–track system, at the same time it is also necessary to consider that contact will happen planimetrically. For deformations of a plastic nature, occurring under the action of high loads, when all uneven-standing contour areas receive significant deformation, plastic deformation of their bases will occur. Therefore, one should assume an increase in the contour area due to the plastic deformation of deeper layers [1, 11–13]. As a result of numerous research it has been established that the basis of traction between wheels and track is the frictional process taking place at the zone of contact and that the adhesion coefficient value is defined by the frictional features of the surfaces of rubbing solids. In turn, the frictional properties of a wheel–track system obviously depend not only on structural conditions—technical characteristics

Table 1 Values of coefficients b and m Processing

Roughness, Rz (lm)

b

m

Whetting, planning, and milling Grinding Polishing

80…20 6.3…1.6 0.8…0.2

1…2 1.5…4 3…10

1.2…2 1.6…3 2…3

Process Research of Wheel-Rail Mining Machines Traction

707

(material, roughness of the interacting surfaces, physico-mechanical properties, etc.)—but also the weather conditions under which rolling stock is operating. As a result of experimental studies executed under real operational conditions, there appears a graphic dependence of the coefficient of locomotive adhesion wn on the change in friction coefficient µ at the time motion starts—measured at the central point of a wheel’s path of a track (Fig. 2). It should be emphasized that experiments were carried out after ensuring the absence of sand and that the frictional parameters of the track for different values of speed showed a wide range of variation in terms of frictional coefficients.

3 Research Results and Discussion The study of experimentally established dependencies presented in Fig. 2 allows us to ascertain that values of w(f) for different types of locomotives are similar in shape and vary in parabolic dependence. The analytical interpretation of the regression equation of the graphs presented in Fig. 2 is approximated by the following function. w ¼ af 2 þ bf þ c;

ð8Þ

where a, b, and c are coefficients characteristic of different types of electric locomotives. The values of the regression coefficients from Eq. 8 for various series of electric locomotives, defined by experimental study, are given in Table 2. Friction coefficients calculated using Eq. 8 are in the range of 0.2–0.5 [10].

Fig. 2 Relationship between the traction coefficient of electric locomotive wheel couples with the track at the time motion starts and the coefficient of friction, measured at the central part of the path of a wheel on the track. 1—for VL8; 2—for VL23; 3—for VL60; 4—for VL60n; 5—according to data from Collins and Prichard; 6—the line of direct proportionality between f and w

708

A. M. Keropyan et al.

Table 2 Regression coefficients for the electric locomotives examined Electric locomotive series

a

b

c

VL8 VL 60 with consecutive agitation VL 60n with independent agitation VL 23

–3.70 –1.48 –1.67 –4.57

3.60 1.50 1.51 4.0

–0.53 –0.04 –0.02 –0.489

Table 3 Largest frictional coefficients for the locomotives in Table 2 Value Friction coefficient Percentage increase Average value of increase f (%)

fcalc 0.535 – 11.96

f1max

f2max

f3max

f4max

(VL8)

(VL60)

(VL60n)

(BL23)

0.486 9.16 –

0.507 5.23 –

0.452 15.51 –

0.437 18.3 –

Equation 8 approximates a parabolic function with a vertex determining the value wmax, which corresponds to a certain value of µ that can be determined after differentiation of its derivative to zero. ðwÞI ¼

dw ¼ 2af þ b ¼ 0; df

then fmax ¼ 

b : 2a

ð9Þ ð10Þ

Table 3 shows the largest frictional coefficient values calculated using Eq. 10. Studying the results of calculations of the greatest friction coefficient values suggests that when processing a track’s surfaces—with micro-roughness values of Rz = 20 lm—the calculated friction coefficient can be increased by 12% and, therefore, the coefficient of traction between a locomotive’s wheels and the track can also be increased.

4 Conclusions • During the course of research it was established that in the area of contact between a locomotive’s wheels and the track an interaction occurs that is peculiar to plastic contact.

Process Research of Wheel-Rail Mining Machines Traction

709

• In the case of deformation that is plastic in nature—occurring under the action of high loads peculiar to wheel–track traction machines, when all the irregularities in the contour site receive significant deformation—base deformation will occur. • Plastic deformation of roughness in the zone of interaction of working surfaces provides an increase in the actual contact area and, as a rule, an increase in the thrust force between the wheels of a mining machine and track. • Analysis shows that when processing track working surfaces that have roughness heights of Rz = 20 lm the calculated value of the coefficient of friction can be increased by 12% and, as a consequence, the coefficient of adhesion between wheels and track and traction force can also be increased.

References 1. Demkin NB (1970) Contacting of rough surfaces. Sience, Moscow 2. Mikhin NM, Lyapin KS, Dobychin MN (1971) Research of the tangential durability of adhesive communication, In the c. “Contact interaction of solid bodies and calculation of friction forces and deterioration”. Science, Moscow 3. Keropyan AM (2016) Features of interaction of the traction wheels of electric locomotive and a diesel locomotive with rails in the conditions of open mountain. Works J Friction Wear 37(1):98–104 4. Kragelsky IV (1968) Friction and deterioration. Mechanical Engineering, Moscow 5. Keropyan AM (2014) Conditions of providing rational geometrical characteristics of working profiles of couple a wheel—a rail of career locomotives. Scientific works of the III International scientific conference “Basic researches and innovative technologies in mechanical engineering”, IMASh RAS, Moscow, 13–15 May 2014, pp 117–119 6. Friedman YB (1972) Mechanical properties of metals. Mechanical Engineering, Moscow 7. Kragelsky IV, Mikhin NM (1984) Frictional units of cars. Reference Book, Mechanical Engineering, Moscow 8. Keropyan AM, Kantovich LI, Voronin BV, Kuziev DA, Zotov VV Influence of uneven distribution of adhesion mass on locomotive wheel pairs, its tractive power, straight and curved sections of industrial rail tracks. Iopscience.iop.org/volume/17551315/87. http:// iopscience.iop.org/article/10.1088/1755-1315/87/6/062005values 9. Keropyan AM (2014) Determination of rational of friction coefficients and roughness of working surfaces of rails and bandages of locomotives. Week of the miner—2014. The collection of scientific works of a seminar “Modern technologies in mountain mechanical engineering”, Moscow, 27–31 January 2014, pp 344–352 10. Luznov YM (2009) Nanotribology of wheel-rail adhesion. Intekst, Moscow 11. Demkin NB, Izmaylov VV (2010) Development of the doctrine about contact interaction of cars details. Messenger Mech Eng 10:28–32 12. Kragelsky IV (1965) Friction and deterioration. Butterworths, London 13. Keropyan A, Gerasimova A (2017) Connection of the temperature in contact area of the wheel-rail system with the railway transport. Izvestia. Ferrous Metallurgy, National University of Science and Technology MISIS 60(5):355–363. https://doi.org/10.17073/ 0368-0797-2017-5-355-363

Development of Double-Sided Summer–Winter Pneumatic Tires I. Voiku and I. Komissarov

Abstract The article substantiates the relevance of improving the design of automobile pneumatic tires. The basic parameters of the utility model “double-sided summer–winter pneumatic tire” (Patent for utility model No. 182679), developed by the fellow workers of the FSBEI of Higher Education Pskov State University, are described. The features and main disadvantages of the closest analog (patent for the invention RU 2028953) are presented: the inability to use it year-round, an increase in the imbalance and the risk of puncture during the operation of the second tread part due to uneven wear of the first and second tread parts. The technical goal, solved by the proposed useful model, is to ensure the possibility of year-round use of pneumatic tires. The goal is achieved by the unique features: The inner part of the tire is a mirror image of the external part; one part of the tire is equipped with spikes; if necessary, the tire can be turned out. The scientific novelty of the developed double-sided pneumatic tires consists of combining two traditional products, which gives a significant synergistic effect: savings of the cost of acquisition and use, reduction of environmental damage.










Keywords Spike Cord Breaker Winter tire pattern


Summer tire
Sidewall
Tread

1 Introduction A car tire is considered to be one of the main components of any car, providing a road grip. The increase in costs of car tires stimulates the development of innovations which have significant functional advantages and are designed to reduce consumption costs.

I. Voiku (&)
I. Komissarov Pskov State University, 2, Lenin Square, Pskov 180000, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_76

711

712

I. Voiku and I. Komissarov

Various domestic and foreign manufacturers of pneumatic tires are working on problem of increasing their service life. The main obstacle they run into while solving this problem is the seasonality of the usage of car tires. This stipulates the relevance of the design of pneumatic tires, combining the advantages of summer and winter tires, at the same time minimizing the disadvantages of their seasonal use. “Double-sided summer-winter pneumatic tire” (Patent for utility model No. 182679) is developed by the employees of the FSBEI of Higher Education Pskov State University. The proposed practical model ensures the possibility of a year-round use of pneumatic tires due to its unique features: The inner part of the tire is a mirror image of an external part, and moreover, one of the parts is equipped with spikes and a key technology allows user to turn out the tire in case of necessity.

2 Relevance Nowadays, motor transport has become a fixture in a daily life of the modern society. Since the early days of automobile, it has been continuously evolving and transforming. At the same time, it resulted in changing the requirements for automobile components. An increase in the costs of automotive components, without which it is impossible to use vehicles, is coming against the background of a decrease in the level of the household income and at the same time stimulates the development of innovations that have significant functional advantages and are designed to reduce consumption costs. A car tire is an elastic shell, located on the rim of the wheel, which is considered to be one of the main components of any car [1]. It is intended for a realization and perception of forces, arising from the contact spot with the road, absorption of vibrations caused by road surface irregularities, as well as to provide a high coefficient of adhesion [2].

3 Goal Setting Manufacturers of pneumatic tires are constantly working on solving the problem of prolonging service life of pneumatic tires [3]. The main obstacle to solving this problem is the seasonality of the usage of a car tire: • The seasonal usage of summer and winter car tires increases the cost of their purchase and storage, requires an additional storage space, and has a negative impact on the environmental pollution.

Development of Double-Sided Summer–Winter Pneumatic Tires

713

• The usage of all-season car tires is not able to provide a high coefficient of adhesion during the winter time [4]. • Anti-skid chains can be used on tires of any size; nevertheless, such devices limit the speed of movement, are dangerous, pose a risk of breakage in case of an improper fastening, can damage the tire, and do not suggest the long-term use. • Retractable spikes are characterized by the unique technology and are used exclusively in high-tech transport facilities. The main drawbacks of retractable spikes could be described as a high cost, difficulties of maintenance, and a narrow consumer segment. Summer car tires are made of a hard rubber and are designed for using at positive ambient temperatures. The tread of a summer tire has a solid rigid structure, ensuring good contact with dry or wet road surfaces [5, 6]. In case of winter car tires, their efficiency increases at negative temperatures up to thirty degrees centigrade and lower [7]. The winter tire tread consists of a large number of lamellae or small sipes in the bulges, providing stability and pass ability of a vehicle [8, 9]. In its turn, winter tires are made of a soft rubber, which loses its effectiveness at positive temperatures of more than 10 °C [10]. Spiked winter car tires have fewer lamellae [11]. Better adhesion and braking are also provided by the asymmetric installation of different types of spikes at different levels [12, 13]. The contact spot of a tire should contain within the limits of 8–12 spikes in order to achieve the effective adhesion to ice-covered road surface. The use of spikes on all-season tires is unacceptable because of their destructive impact on the road surface [14]. Today, the specialists of one of the well-known companies, producing car tires, are developing tires with automatic spikes. This development will be a breakthrough in tires production, involving the possibility of automatic extension of the spikes out of the tire and back at the request of a driver. However, this development can remain solely conceptual due to its very high costs. The low-cost option of car tires with universal characteristics is all-season tires [15]. All-season tires are suitable for any season, because of the combination of summer and winter rubber [16]. The firmness of this type of tire takes an intermediate position between the soft winter tire and hard summer one. The main advantages of using all-season tires are: • No need for seasonal change and, as a consequence, no storage required as for a changeable seasonal tire set; • Efficiency in changeable winter conditions: the optimum temperature for combined rubber is within the range from +5 to −10 °C [17]; • A low noise level. However, the concept on symbiosis of summer and winter car tires has not only advantages but also disadvantages. A regular whole year-round wear of all-season tires reduces their lifetime because the optimal temperature for all-season tires lies

714

I. Voiku and I. Komissarov

in the range from +5 to −10 °C; otherwise, usage of such tires in other temperature conditions reduces their service life [17]. Anti-skid chains are devices, mounted on car tires to ensure a maximum traction when driving on snow and ice. They are manufactured and used only for a specific size of car tires. These chains limit operational capability of a car and do not assume long-term usage [18]. This stipulates the relevance of the design of pneumatic tires, combining the advantages of summer and winter tires and anti-skid devices and minimizing the disadvantages of their seasonal use.

4 Theoretical Part The development of a double-sided summer–winter pneumatic tire is a combination of two types of different season car tires in order to lower consumer costs for their purchase and storage and reduce an environmental damage of their production and disposal. There is a developed model “detachable tread” [Patent on invention RU 2245797 from 10.02.2005 “Detachable tread for pneumatic tires” Thazeplov Hasan Misedovich (RU)], including a profiled rim of elastic material with a pattern on the outer surface. An inner surface of the tread has a pile cover, the inner diameter of which is smaller than the outer one of the main tire by 1–2%, and is designed in such a way that covers the outer and most of the side surface of the main tire. As a result, the reliability of the tire increases. One more model could be mentioned and its name is “Pneumatic tire” [Patent for invention RU 2028953 from 20.02.1995 “Pneumatic tire” Haritonashvili V.A. (RU)], which contains a turned over tread part with the pattern on two opposing surfaces, integrally formed with the sidewalls. One of the surfaces in working condition is turned into a pneumatic cell. An outer tread part wears out while it is operating. After wearing out of the outer tread part to the wear limit, a tire from the wheel rim is removed, and the tire is inverted, so that the outer tread part has become the inner, and the inner tread part has become the outer, i.e., a working tread part. The disadvantage of this device is the fact that its usage involves the following sequence of operations: 1. Operation of the first (outer) tread part to the wear limit; 2. Overturning, in which the first tread part has become internal, and the second one has become the outer working part; 3. Operation of the second tread part to its wear limit. The disadvantage of the known design of the pneumatic tire is an inability to use it year-round. In addition, one more disadvantage is an increase in imbalance and risk of a puncture during the operation of the second tread part due to the uneven

Development of Double-Sided Summer–Winter Pneumatic Tires

715

wearing out of the first and second tread parts (since the operation of the second tread part begins after the first tread part becomes worn to the wear limit).

5 Practical Significance, Proposals, and Results of Implementation, Results of Experimental Studies “Double-sided summer-winter pneumatic tire” (Patent for utility model No. 182679) is developed by the employees of the FSBEI of Higher Education Pskov State University [19]. The utility model belongs to the tire industry and to the designs of pneumatic tires for wheeled vehicles in particular, and it can be practically used for various types of tires. The technical goal, solved by the utility model, is to ensure the possibility of year-round usage of pneumatic tires. This goal is achieved by the fact that the proposed device contains an inverted tread part with a pattern on two opposing similar surfaces connected by a layer of adhesive material, one of which contains anti-slipping spikes and is made of a winter composition of a rubber mixture that retains elasticity at a negative temperature, and the other is made of a summer composition of a rubber mixture that retains elasticity at high positive temperatures. The frequency of turning a pneumatic tire is determined primarily by the time of year, and not by the wear of the outer tread part to the wear limit. The technical essence of the proposed device is explained by the following Figs. 1 and 2. Conceptually, the tire structure consists of the first tread surface 1, which made of a summer composition of a rubber compound, the second tread surface 2, which made of a winter composition of a rubber compound and equipped with anti-slipping spikes 3, the first and the second tread surfaces are connected by layer of the adhesive material 4. Each of the opposite tread surfaces 1 and 2 is characterized by a standard design of the outer surface of a tire. The surface 1 is made of a summer composition of a Fig. 1 Cross-sectional cut of the tire in summer operation [19]

716

I. Voiku and I. Komissarov

Fig. 2 Cross-sectional cut of the tire in winter operation [19]

rubber compound that retains elasticity at high positive temperatures. The surface 2 contains anti-slipping spikes 3 and is made of winter composition of a rubber compound that retains elasticity at negative temperature. The adhesive layer 4 provides adhesion of the surface 1 and the surface 2 with two degree of freedom. Overall dimensions of the proposed device are taken according to existing standards for such constructs. The principle of operation of the proposed device is described by the following mechanism. In the summer season, the outer tread surface 1 operates, before the onset of the winter season, an overturn is performed as a result—the outer tread surface 1 has become internal, and the inner surface 2 has become outer or the working part. Further change of seasons is accompanied by a repeated overturn. The number of overturns is determined by the user of the device. The operation of each tread part is carried out to the wear limit [20]. The technical result of a claimed utility model is an increase in the safety of the operation of a double-sided tire, a reduction of the material consumption for the tire production.

6 Conclusion At present time, a motor transport has become a fixture in a daily life of the modern society. Since the very beginning of the automobile invention, a technical vehicle has been continuously evolving and transforming; therefore, the requirements for automobile components also have to change. One of the main components of any car is a tire or an elastic shell, located on the rim of the wheel. Manufacturers of pneumatic tires are constantly working on solving the problem of prolonging service life of pneumatic tires. The main obstacle to solving this problem is the seasonality of the usage of car tire.

Development of Double-Sided Summer–Winter Pneumatic Tires

717

In the course of research work, the patented scientific and technical developments, aimed to solve this problem, were identified; however, all of such developments are characterized to have both advantages and disadvantages. “Double-sided summer-winter pneumatic tire” (Patent for utility model No. 182679) is developed by the employees of the FSBEI of Higher Education Pskov State University. The utility model belongs to the tire industry and to the designs of pneumatic tires for wheeled vehicles in particular, and it can be practically used for various types of tires. Conceptually, the structure of innovative solution consists of the first tread surface, which is made of summer composition of a rubber compound, the second tread surface, which is made of winter composition of a rubber compound and equipped with anti-slipping spikes. A key principle of operation of the proposed device is described in the following algorithm. In the summer season, the outer tread surface is operated. Before the onset of the winter season, an overturn is performed, in which the outer tread surface has become internal, and the inner surface has become outer or working part. Further change of seasons is accompanied by a repeated overturn. The number of overturns is determined at the user’s discretion. The operation of each tread part is carried out to the wear limit. The development of a double-sided summer–winter pneumatic tire allows reducing the consumer costs for their purchase and storage, lowing environmental damage of their production and disposal.

References 1. Tarkovskij VN, Gudkov VA, Tret’yakov OB (1990) Car tires. Transport, Moscow, p 217 2. GOST 4754-97. Pneumatic tires for light vehicles, trailers to them, light trucks and buses of extra small capacity. Technical regulations. http://docs.cntd.ru/document/1200017825. Date of access 29 Dec 2018 3. Chemerinskij VB (2012) Procedure for improving tires and their production. Transport Business of Russia 6–2:253–256 4. Kulikov GM (1990) Design of pneumatic tires on the basis of generalized broken-line hyrothesis. Strength Mater 22(2):272–277 5. Voronin VV, Kondrashov VN, Timaev DM (2010) Mechanical characteristics of car tires. In News of the Moscow State Technical University MSTU, vol 2, pp 20–23 6. Kolbasov AF (2011) Some topical issues of the duty of car tire. Fundamental Research 8– 1:128–130 7. Biderman VL, Slyudikov LD, Levin YuS et al (1970) Influence of structural and operational factors on wear, adhesion and rolling resistance of car tires. CNIITEH neftekhim, Moscow, p 106 8. Knoroz VI (1976) Duty of car tire. Transport, Moscow, p 238 9. Sokolov SL (2010) Prediction of the fatigue life of pneumatic tires. J Mach Manuf Reliab 39 (5):459–465 10. Kvasha EhN, Tkacheva VV (2009) Calculation method of the temperature fields of pneumatic tires. Bull Pridniprovian State Acad Const Archit 1(132):27–31

718

I. Voiku and I. Komissarov

11. GOST 22374-77 Pneumatic tires. Design. Terms and definitions. http://docs.cntd.ru/ document/gost-22374-77. Date of access 29 Dec 2018 12. Lomakin VV, Karpuhin KE, Voronin VV, Kondrashov VN (2008) Perspective directions of development of car tires. In News of the Moscow State Technical University MSTU, vol 2, pp 64–67 13. Tarasov VN, Boyarkina IV (2018) Method of sections in analytical calculations of pneumatic tires. J Phys: Conf Ser 944:012116 14. Grigolyuk EI, Kulikov GM, Plotnikova SV (2004) Contact problem for a pneumatic tire interacting with a rigid foundation. Mech Compos Mater 40(5):427–436 15. Kvasova A, Gerike B, Murko E, Skudarnov D (2017) Forecasting of a thermal condition of pneumatic tires of dump trucks. Forecasting of a thermal condition of pneumatic tires of dump trucks. In E3S Web of Conferences the second international innovative mining symposium 16. Belkin AE, Bukhin BL, Mukhin ON, Narskaya NL (1997) Some models and methods of pneumatic tire mechanics. In Vehicle System Dynamics Proceedings of the 1997 2nd international colloquium on tyre models for vehicle dynamic analysis. Berlin, Germany 17. Hodes IV (2007) Methodology for forecasting the controllability of a wheeled vehicle. Dissertation, Volgograd, 33 p 18. Vantsevich VV, LomakoG Opeiko A (2003) Interaction between autonomous vehicles and road surface. Int J Veh Auton Syst 1(3–4):291–308 19. Voiku IP, Komissarov IA, SHmanaj IA (2018) Double-sided summer-winter pneumatic tire. Patent for utility model RUS182679 28.08.2018 20. Shifrin BM (2006) An analytic model of a rolling pneumatic tire. Int Appl Mech 42(4): 479–485

Asphalt Concrete as Object of Destruction by Operating Units of Milling Machines D. V. Furmanov, V. A. Nikolayev and N. N. Klochko

Abstract The chapter presents the analysis of methods for determining the strength of road asphalt concrete in terms of cutting resistance. Existing methods, which are included in the list of standards of many countries and are determined for assessing the quality of asphalt concrete coatings, evaluate only the operational strength of the material. In order to determine the strength of asphalt concrete as an object of destruction by the working bodies of milling machines, it is necessary to use a different system of methods. As one of these methods, it is proposed to use the method of static penetration with the use of profiles of different shapes. The results of strength tests of some types of asphalt concrete by the method of static penetration are presented, and the description of this method, which is provided for the use of press equipment together with cylindrical profiles and linear profiles, is given. A comparative evaluation of the results obtained by static penetration with the standard method is given. It is noted that the obtained values of contact stresses significantly exceed the values of stresses in the destruction of standard samples for uniaxial compression. The influence of the strength of the structure-forming components of asphalt concrete is also noted. From the point of view of similarity of processes of destruction and visual results, the method of static penetration is the closest to the process of milling of asphalt concrete.







Keywords Rotary hoes Linear compression Asphalt concrete penetration Contact strength Uniaxial compression








Static

1 Introduction Road resurfacing requires removing the previous, old and worn asphalt concrete. Currently, cold milling is widely used for these purposes. The existing milling machines and equipment designed to remove old pavement are complex, expensive D. V. Furmanov (&)
V. A. Nikolayev
N. N. Klochko Yaroslavl State Technical University, 88, Moskovsky prospect, Yaroslavl 150023, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_77

719

720

D. V. Furmanov et al.

and energy-consuming. However, new types of road asphalt concrete of increased durability are developed continuously, which requires organizing the material depending on the durability and the answers to the question of the performance anticipated, and, perhaps, the capability of milling machines to work with this material. Asphalt concrete milling is accompanied by the destruction of its structure, and the grains themselves are being destroyed together with the links between them. Given this, the destruction of asphalt concrete is the task opposite to that, which it was created for, specifically to maintain high durability regardless of the loads and the impact of the environment. All the positive qualities of asphalt concrete prevent its destruction—high durability, density, viscosity, elasticity and plasticity combined with a wide variety of physical and mechanical properties depending on temperature and grain size composition. Finding out the asphalt concrete durability as an indicator of destruction resistance during milling is one of the important tasks toward the creation of new milling and other machines for combing asphalt concrete, as well as the development of the existing milling machines and equipment.

2 Method of Research The methods and tools to evaluate the durability of asphalt concrete were improved continuously together with the development of transport infrastructure, the increase of requirements to reliability and longevity of road pavements, the increase of loads on road networks, and so on. Also the new durability, shear resistance, aging, cracking resistance and abrasion tests have appeared for asphalt concrete pavements. The first asphalt concrete stability test (Hubbard-Field test), which appeared in 1920, marked the beginning of mechanical testing in the quality control system. The method was to determine the force required to penetrate a standard sample of asphalt concrete through a hole of a certain size; see Fig. 1c. Some other methods have received certain popularity, such as Hveem’s complex approaches [1–3], where asphalt concrete mix in the form of cylindrical samples that is selected by a special method was tested for a uniaxial compression under the Duriez procedure; see Fig. 1a. The Marshall method [4, 5] allowed to obtain an internal friction rate at the split (see Fig. 1b) using the following dependencies: tga ¼

3
ðA M  A C Þ ; 3
AM  2
AC

ð1Þ

where AM and AC—the forces of deformation toward the sample under the Marshall method and the uniaxial compression, accordingly. AM and AC can be determined by the following formula:

Asphalt Concrete as Object of Destruction by Operating Units …

721

Fig. 1 Schemes of asphalt concrete samples tests

AM ¼

F
l ; 2

ð2Þ

where F—destructive force, kN; l—critical deformation, mm. The adhesion C can be found depending on the resulting angle of internal friction and the critical durability of compressed asphalt:

722

D. V. Furmanov et al.

1 C ¼ ð3  2
tguÞ
Rcomp ; 6

ð3Þ

where Rcomp—the critical durability of the asphalt concrete sample under the uniaxial compression, MPa. A progressive approach based on a set of tests is realized in the Superpave method [6, 7], which is gradually replacing the previous methods. There is also a special place for penetration methods (Fig. 1d) based on the analysis of deformation properties of asphalt concrete under load [8–10]. The methods described above were developed specifically to analyze the long-term durability and stability of asphalt concrete pavements. However, this property is unacceptable for power and energy analysis of milling. It is known that the main reason for the loss of operational durability is the disturbance of bonds between individual structure-forming components of the material. It causes flaws such as cracks, chips and dents in the pavement. Milling is also accompanied by the active crushing of the stone fraction, which increases energy costs considerably.

3 Research Result That is why the methods to determine the durability of road asphalt concrete should consider both the strength of bonds between the grains and the durability of the grains themselves. The closest from this point of view is the static penetration method, because by using this method the destruction of the stone fraction appears at the same time as the destruction of bonds between the grains. Such phenomena are typical for the asphalt concrete milling. The temperature of the asphalt concrete has a significant impact on its durability. Regardless of the testing method, changing the temperature from zero to +50 °C changes the durability enormously. Therefore, the resulting durability value must always be tied to the temperature. As part of the study of determining the durability value for asphalt concrete of different grades, a series of tests was conducted that allowed to evaluate the impact of various factors on the durability value. The durability values that were obtained by the destruction method of uniaxial compression of a cylindrical standard sample under GOST 12801-98 [11] and the durability values that were obtained by static penetration of the same asphalt concrete samples were compared. The last method is not standard and implemented on a hydraulic press (see Fig. 2) when using indenters of cylindrical and linear forms, displaying and storing values on a computer. During the study, the samples with 0, +20 and +50 °C were used and they were subjected to thermal control previously. The following grades of asphalt concrete were chosen: A1, B1 and D3 according to GOST 9128-2009 [12] as well as

Asphalt Concrete as Object of Destruction by Operating Units …

723

Fig. 2 A test stand to determine the durability value of asphalt concrete by static penetration

Table 1 Asphalt concrete sample compositions used in the research Attribute

Asphalt concrete grade D3 SMA-15

Classification

Sand asphalt

Crushed stone Bitumen

Not used BND 60/90 [15]

A1

Granular asphalt concrete consisting of crushed stone by up to 80%

B1

Granular asphalt concrete consisting of crushed stone by 50–60% Crushed granite, M1200 brand [14]

Granular asphalt concrete consisting of crushed stone by 40–50%

BND 60/90, modified

BND 60/90

BND 60/90

SMA-15 according to GOST 31015-2002 [13]. Table 1 shows the composition of these types of asphalt concrete. The critical contact stresses were determined for the static penetration, and they are defined by their maximum force of indentation of a cylindrical and a linear mold. The values of maximum contact stresses r (MPa) are calculated with these formulas: • for a cylindrical mold: where F—force needed to press the indenter in, H; d—indenter diameter, mm;

724

D. V. Furmanov et al.

• for a linear mold: r¼

F ; d
a

ð5Þ

where a—mold length, mm; d—mold width, mm. Below are the test results of four types of asphalt concrete widely used for solving various road and airfield construction tasks, differing in purpose, durability values and composition. Figure 3 shows a dependency graph showing the dependence of durability on temperature when tested for a uniaxial compression and obtained by pressing-in of a cylindrical indenter with a diameter of 13 mm for D3 sand asphalt. The significant difference of the durability value prohibits using the standard method to determine the resistance to milling. Figure 4 shows the graphs obtained by testing the samples of granular asphalt A1 and B1, for which the durability values were obtained not only with a cylindrical, but also with a linear indenter. Since certain granularity of asphalt concrete causes an uneven resistance to the pressing-in of a cylindrical indenter, the results of such test have significant variation. The linear indenter affects several grains at the same time, providing a more stable result. For granular asphalt, there is also a specific significant difference in durability of standard samples under the uniaxial compression and durability under static penetration.

Fig. 3 Dependency graph of the critical durability of the D3 asphalt on temperature

Asphalt Concrete as Object of Destruction by Operating Units …

725

Fig. 4 Dependency of B1 (a) and A1 (b) grade asphalt durability on temperature

4 Conclusion The most efficient outer diameter of the perforated tube is the diameter with the values closest to each other. • The closest asphalt concrete resistance evaluation method similar to the material destruction process would be the static penetration method. • Numerical values of critical contact stresses obtained by static penetration are several times higher than the critical durability obtained by the destruction of standard samples with uniaxial compression. • Finding the durability value by static penetration of granular asphalt with higher repeatability is possible when using a linear indenter. • For both the penetration and uniaxial compression methods, there is a significant increase in the durability value when the temperature decreases. This is especially true for low temperatures. This clearly indicates that the asphalt cutting forces concentrated on the teeth of a working milling unit and, thus, the power consumption of milling depends on the temperature of asphalt. Thus, the rational design of working milling units and milling machines overall inevitably depends on the durability of materials. To promptly predict the loads on the working milling unit, it is possible to establish the similarity between the cutting

726

D. V. Furmanov et al.

forces applied to the cutters of a cutter drum and the critical contact stresses defined by static penetration. This method also suggests the way to determine the approaches to find the rational use cases of milling equipment considering the weather conditions, time of year and time of day to carry out the milling operation.

References 1. Radovskiy BS (2006) Doroznaya Tehnika: Asphalt composition design methods in the USA 6:68–81 2. Kiryukhin GN (2005) Avtomobilniye dorogi i mosty: Asphalt composition design and its testing methods. Overview, Issue 6, 96 p 3. Radovskiy BS (2008) Doroznaya Tehnika: Modern state of development of the American asphalt mix design method Superpave 6:12–22 4. Asphalt Institute (1989) The Asphalt handbook MS-4 5. Asphalt institute (1997) Asphalt cold mix manual, manual series no. 14, Lexington, KY 6. Radovskiy BS (2007) Doroznaya Tehnika: Asphalt mix composition design by the Superpave method, pp 86–99 7. Final Report of the TRB Superpave Committee (2005) TRB. Superpave, pp 1–56 8. Eryomin VG, Eryomin AV, Volokitin VP (2003) Scientific Herald of the Voronezh State University of Architecture and Civil Engineering. Road and Transport Construction series: No. 1. The results of the study of deformation and durability properties of asphalt by pressing-in of a spherical indenter 9. Ladygin BI et al (1972) Nauka i Tehnika: Durability and longevity of asphalt. In Ladygin BI, Yatsevich IK (ed) Minsk, 288 p 10. GOST 12801–98, (1999) Materials on the basis of organic binders for road and airfield construction. Test methods. Publ, MNTKs, Moscow 11. GOST 9128-2009 (2010) Asphaltic concrete mixtures for roads and aerodromes and asphaltic concrete. Specifications. Introduced from 1.01.2009. Standard in form, Moscow 12. GOST 31015-2002 (2003) Bituminous stone mastic mixtures and stone mastic asphalt. Specifications. Introduced from 1.05.2003. Standard in form, Moscow 13. GOST 8267-93 (1994) Crushed stone and gravel of solid rocks for construction works. Specifications. Introduced from 10.11.1993. Standard in form, Moscow 14. GOST 22245-90 (2005) Viscous petroleum road bitumens. Specifications. Introduced from 1.01.1991, IPC Publishing House of Standards, Moscow

Improving Automotive Torque Converter Quality N. N. Trushin, V. Y. Antsev and A. A. Obozov

Abstract Hydrodynamic torque converters are commonly used in self-propelled vehicle transmissions. Torque converters are most efficient in trucks, haulers, tractors and other heavy and utility vehicles operating in variable road conditions and off-road. Torque converters can automatically control the engine torque, but their efficiency is lower compared to mechanical gears. Many automotive hydromechanical transmissions use single-stage, four-wheeled torque converters containing two reactor wheels. In such torque converters, two reactor wheels form a single reactor. This design enables to slightly increase the efficiency as compared to a three-wheeled torque converter, but in practice four-wheel torque converter performance is not much different from that of three-wheel torque converters. In order to improve the efficiency of automotive hydromechanical transmissions, the study considers methods for online control of the four-wheel torque converter load capacity. The design of a torque converter with two reactors is proposed, providing for the online reactor switching depending on the vehicle driving conditions. This solution makes it possible to obtain two ranges through a joint operation of the transmission with the vehicle engine. Keywords Self-propelled vehicle Transparency factor


Torque converter
Quality improving


1 Introduction Transmissions in self-propelled machines of various functionalities are commonly assembled with hydromechanical transmissions (HMT) with hydrodynamic torque converters. One of the first HMT for Mercedes cars and buses was developed by a N. N. Trushin (&)
V. Y. Antsev Tula State University, 92, Lenin av, Tula 300012, Russia e-mail: [email protected] A. A. Obozov Bryansk State Technical University, 7, Bulvar 50 let Oktiabrya, Bryansk 241035, Russia © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_78

727

728

N. N. Trushin et al.

German engineer Rieseller (1925), whose HMT was a two-stage torque converter connected to the planetary gearbox [1, 2]. Currently, the world and Russian automotive park is highly saturated with HMT-provided cars and buses. HMT is used in heavy-duty trucks and cross-country vehicles. HMT is also used in forklift trucks, road construction machinery, tracked vehicles and tractors, military machines and railway locomotives. According to statistics, the automotive industry has seen steady growth in the production of cars with automatic hydromechanical transmissions. The most rapid pace has been seen in equipping new cars with automatic transmissions. According to AUTOSTAT Russian analytical agency (http://www.autostat.ru), Russia witnesses a steady growth in sales of cars with automatic transmissions. While in 2010 only 33.5% of the cars sold in Russia had automatic transmissions, in 2017 the market share of such cars reached 54%. Accordingly, the proportion of torque converters used in car transmissions increases. The reasons for extensive use of HMT with torque converters in vehicles are as follows: • automatic ratio adjustment controlled by the drive train load as the engine torque is continuously transferred; • damping drive train torsional vibration that extends service life and ensures smoother operation of the engine and drive train components; • high capacity with relatively compact size and low weight compared to mechanical transmissions; • better drive train controllability and general vehicle handling, specifically in demanding operating conditions [3, 4]. Typically, automotive HMTs use a hydrodynamic torque converter, an automated gearbox and a hydraulic control system. The mostly widespread HMTs are the single-stage (or single-turbine) complex torque converters, capable of operating at modes of torque transformation and hydraulic coupling. Such complex torque converters contain the impeller (or pump) of the centrifugal type, the turbine wheel (or turbine) of the centripetal type and one or two reactor wheel axis type [5, 6]. Despite its advantages, the HMT still has a number of drawbacks in comparison with traditional manually operated mechanical transmissions, such as complex design, high cost, lower coefficient of performance (COP) and higher performance requirements [7, 8]. Therefore, research work in the field of design and technological HMT improvement for self-propelled machine transmissions remain relevant for many years [9, 10].

2 Problem Statement Torque converters have the property of automatic control of the engine torque, but are characterized by lower values of efficiency compared to mechanical gears. The lowest efficiency values of the torque converter are seen in the range of small gear

Improving Automotive Torque Converter Quality

729

ratios (from 0 to 0.5). The reduction of efficiency is also observed in the range of high gear ratios preceding the transition of the complex torque converter to the hydraulic coupling mode. In this regard, the actual task of designing the HMT with torque converters is to provide high values of efficiency in the entire control range of the transmission. In order to extend the regulation range in the field of high gear ratios, complex torque converters were developed, in which the reactor is connected to the torque converter housing by means of a freewheel unit. With a transmission ratio of 0.75– 0.85, this torque converter automatically switches to hydraulic coupling mode. Complex tricycle torque converters are very widespread in the hydrodynamic transmission of self-propelled vehicles, as they have a simple design. However, when designing the transmission of a self-propelled vehicle, the torque converter parameters have to be chosen from compromise considerations. If a tricycle torque converter should have a high torque ratio (about 3–3.5), the efficiency in the field of high gear ratios is inevitably reduced; if the torque converter is selected according to high efficiency in the entire control range and in the hydraulic coupling, then the maximum value of the torque ratio is not high (not more than 2.5) [11, 12]. In the late 1940s, in order to further increase the efficiency in the range of high transfer ratios, complex torque converters equipped with two reactor wheels were designed [13]. In particular, such torque converters were developed by Chrysler in 1948 (patents US 2551746, 2548207, 2613503, 2616310, 2609706, etc.). Four-wheel torque converters have a two-stage reactor, consisting of two impellers, each of which is mounted on a separate freewheel unit. Thanks to this solution, the maximum efficiency of the torque converter transfers into the range of average values of the transmission, and switching to hydraulic coupling mode occurs at a higher transmission ratio compared to the torque converter with one reactor. Torque converters with two-stage reactor were originally widely spread in hydromechanical transmissions of heavy self-propelled vehicles of the world and Russian transport engineering. For example, Allison Transmission used four-wheel torque converters in transmissions for heavy-duty trucks and tanks. In the 1960s, the USSR developed unified complex torque converters with two reactors and cast impellers, intended for trucks and city buses [14]. Experience in the production and operation of torque converters with two reactors showed that the gain from increasing the efficiency at low maximum values of the torque ratio does not always justify the design complexity. Therefore, many manufacturers have stopped equipping transmissions of self-propelled vehicles with torque converters having two reactors. At present, Russia continues to produce four-wheel torque converters of GT type (GT-390, GT-543, etc.), which are used to equip hydromechanical transmission of heavy trucks, dump trucks, road construction machines, tractors and locomotives [15]. However, the potential of the classic four-wheel automotive torque converter is not fully exhausted yet. The operational capabilities of this torque converter can be expanded by quick changing its load capacity, depending on the conditions of movement of the self-propelled machine. Let us consider the proposed method [16].

730

N. N. Trushin et al.

Torque converter properties largely determine the properties of the entire drive train of the vehicle. Torque converters are characterized by the property of transparency, which consists in changing the operation mode of the engine when the torque or angular velocity of the driven shaft is changed. The property of the complex automotive torque converters to load the engine is characterized by the ratio of the pump torque coefficient k1max at the stopped turbine (turbine angular velocity x1 is 0) to the pump torque coefficient k1M corresponding to the transition of the torque converter to the hydraulic coupling mode (i.e.Q at the torque converter transformation coefficient K is 1). This value is called the transparency factor: Y

k1max : ð1Þ k1M Q The non-transparent torque converter has ¼ 1 and is characterized by the fact that at change of the turbine operating mode, the loading mode of the pumpQ and the engine does not change. The direct-transparency torque converter ð [ 1Þ increases the load on the pump and then on the engine as the turbine load increases. Q In automotive torque converters, the value of the transparency coefficient is usually 1.2–1.5 for the semitransparent ones and is 1.8–2.5 for the transparent ones. Transparent torque converters are usually used in passenger cars, and non-transparent and semitransparent ones are used in city buses, trucks, wheeled and tracked transporters and tractors [17, 18]. As a rule, the parameters of HMT torque converter are selected on the basis of the average operating conditions of a self-propelled machine. However, the various operating conditions of many self-propelled machines require rapid changes in the torque converter transparency. The expediency of the transparency operational regulation of the torque converter is caused by the following considerations. When choosing the joint operation of the engine and the torque converter, various options for combining their characteristics are possible. For example, to obtain good efficiency at steady driving in good road conditions, it is advisable for the speed entrance n10 , developed by the motor and the pump wheel of the torque converter at full throttle and braked the turbine wheel to be lower. In this case, the slip during the operation of the integrated torque converter in the hydraulic coupling mode is reduced and the efficiency of the hydraulic transmission is increased. On the contrary, for fast acceleration and better use of engine power, it is desirable to choose as high rpm input as possible. High n10 values may also be required when driving a self-propelled vehicle in adverse road and off-road conditions. Thus, in torque converters used in cars, the n10 value is selected as a compromise and typically varies in the range from 1400 to 2000 min−1. On the other hand, the momentum input of the converter should be changed in the range from 1000 to 3000 min−1 depending on the driving conditions of the vehicle. In 1950s, the General Motors Company (GMC) proposed to use rotary blades mounted on the reactor wheel in automotive torque converters in order to regulate inlet revolutions (US patents Nos. 2882684, 2910832, etc.). It is possible to execute the turning vanes at the pumping wheel, but it is structurally difficult. In addition, ¼

Improving Automotive Torque Converter Quality

731

the rotation of the pump wheel blades, having a complex shape, is difficult due to kinematic reasons. This purpose is easier achieved with the reactor wheel, as a more accessible element strongly influencing the load properties of the torque converter. When using the rotary blades, the angle of the working fluid outlet from the reactor wheel can vary from 15° to 85°. The position of the blades of the reactor is usually installed depending on the feed rate of fuel to the engine: full pressing of the fuel supply pedal corresponds to the small angle of the blades, and minor pressing corresponds to the setting of the blades of the reactor at a greater angle [16]. It should also be noted that the change in the position of the reactor blades allows, under certain conditions, to reverse the rotation of the turbine wheel of the torque converter when reversing the rotation of its impeller [14]. Torque converters with rotary blades of the reactor are not widely spread in the HMT due to the significant complexity of the design. To overcome this drawback, the blade of a unified four-wheeled torque converter with two reactors is used. In this case, the torque converter acquires the ability to operate in two modes corresponding to different n10 values. The torque converter control system ensures operation of only one of the two reactors in each operating mode, while the inactive reactor rotates freely in the flow of the working fluid [19].

3 Results of Engineering Figure 1 presents first version of the fundamental kinematic schemes of the adjustable complex torque converter for automotive HMT, based on inventions [20, 21]. Let us consider the first variant of the converter (Fig. 1). Torque converter Pos. 1 is installed in the housing not indicated in the diagram, in which the drive Pos. 2 and slave Pos. 3 shafts are, respectively, pump Pos. 4 and turbine

Fig. 1 Kinematic diagram of the first embodiment of an adjustable complex torque converter

1 4

5

6

7

10 11

2

3

8

9

732

N. N. Trushin et al.

Pos. 5 wheels. The first reactor wheel Pos. 6 and the second reactor wheel Pos. 7 are installed, respectively, on the freewheeling mechanisms Pos. 8 and Pos. 9, which ensures the transition of the torque converter to the mode of the hydraulic coupling in each range of joint operation of the torque converter with the drive motor. For the first version of the variable torque converter, both freewheeling mechanisms are coupled to the body by means of the Pos. 10 and Pos. 11 controlled brakes. The adjustable torque converter works as follows. The rotation of the drive shaft Pos. 2 is transmitted to the pump wheel Pos. 4, which creates the flow and pressure of the working fluid. The turbine wheel Pos. 5 converts the energy of the working fluid into the rotational motion of the output shaft Pos. 3. The liquid coming out of the turbine wheel then enters the blades of the reactor wheels Pos. 6 and Pos. 7 and returns to the pump wheel Pos. 4. In the first mode of operation of the torque converter, one of the brakes, for example, brake Pos. 10, is switched off. The reactor wheel Pos. 6 will rotate freely in the flow of the working fluid. The active brake Pos. 11 in this case includes the reactor wheel Pos. 7. In the second operating mode, the brake Pos. 10 is switched on, and the brake Pos. 11 is switched off. The reactor wheel Pos. 6 will be switched on, and the reactor wheel Pos. 7 will rotate freely. The operation of the torque converter in the mode of the hydraulic coupling is provided by means of free running Pos. 8 and Pos. 9. Manual or automatic activation and deactivation of brake Pos. 10 and brake Pos. 11 are achieved by means of the HMT hydraulic system. Figure 2 presents the second version of the fundamental kinematic schemes of the adjustable complex torque converter for automotive HMT [22]. In the second variant of the adjustable torque converter, only one brake Pos. 10 is used to automatically switch the reactors. The blades of the reactor are profiled in such a way that the exit angle of the working fluid from the first reactor is equal to the angle of the liquid entrance in the second reactor. Since in this design the torque converter reactors do not work at the same time, the angles of entrance of

Fig. 2 Kinematic diagram of the second embodiment of an adjustable complex torque converter

1 4

5

6

7

2

10

3 8

9

Improving Automotive Torque Converter Quality

733

the working fluid in reactors Pos. 6 and Pos. 7 and the angles of the liquid outlet of the reactors are calculated on the basis of torque converter operation as a three-wheeler. In order to ensure the automatic switching of reactors with only one brake, the following condition must be met: the angle of the working fluid outlet from the blades of the first reactor Pos. 6 must ensure the rotation of the second reactor Pos. 7 towards the rotation of the pump wheel Pos. 4. The specified condition is illustrated in Fig. 3, where the scan is shown on the plane of the blades of the reactor wheels (arrows showing the direction of the working fluid flow along the axes of the shafts of the torque converter) [8]. The blades of the reactor wheel Pos. 6 have the angles of the inlet and outlet of the working fluid b11 and b12 , respectively, and the blades of the reactor wheel Pos. 7 have the angles b21 and b22 , respectively. The angles are related to each other by the following relations: 

b12 6¼ b22 : b12 \b21

ð2Þ

For comparison, Fig. 4 presents the development drawing of reactor blades for a conventional single-stage four-wheel torque converter, which shows that the angle of the blade outlet (or the angle of the working fluid outlet from the blades of the first reactor) is equal to the angle of the second reactor entry into the blades (b12 ¼ b21 ). In the first torque converter operational mode, the brake Pos. 10 is switched off, and the reactor wheel 6 rotates freely in the working fluid flow. In this case, it is the reaction wheel Pos. 7 that works. In the second operating mode, the brake Pos. 10 is switched on and the reactor wheel Pos. 6 is switched on. The working fluid exiting

Fig. 3 Plane development of the adjustable four-wheel torque converter blades

Output flow

β22

Second reactor 7

β21 β12

First reactor 6

β11

Input flow

734 Fig. 4 Plane development of the conventional single-stage four-wheel torque converter blades

N. N. Trushin et al. Output flow

β22

Second reactor

First reactor

β11

Input flow

the blades of the reactor wheel Pos. 6 interacts with the reverse side of the blades of the reactor wheel Pos. 7 and causes this wheel to rotate in the rotation direction of the pump wheel Pos. 4. The freewheeling mechanism Pos. 9 is automatically switched off.

4 Conclusion The proposed design of the torque converter makes it relatively easy to solve the problem of regulating the HMT load capacity in relation to the changing conditions of the movement of the self-propelled machine. In implementing this technical solution, it is possible to use the blade system of the unified Russian four-wheel complex torque converters of “GT” series. Modification of the original torque converter will mainly include changing the attitude position of the reactor impellers, and only two new elements are added to the control system of the torque converter hydromechanical transmission. At the same time, the lockup clutch can be completely borrowed from the original torque converter. Acknowledgements The authors thank the personnel of the Research Library at Tula State University and of the Tula Region Research Library for their assistance with the references and patent search, and innovative design development.

References 1. Kudryavtsev AP (1934) Fundamentals of hydrodynamic energy conversion (turbodrives). Leningrad 2. Prokof’ev VN (1947) Automotive hydraulic transmission. Mashgiz, Moscow

Improving Automotive Torque Converter Quality

735

3. Kosenkov AA (2003) Arrangement of automatic gearshifts and transmissions. Fenix, Rostov-on-Don 4. Kharitonov SA (2003) Automatic transmissions. Astrel, Moscow 5. Naunheimer H, Bertsche B, Ryborz J, Novak W (2011) Automotive transmissions, 2nd edn. Springer, Berlin, Heidelberg 6. Nanney MJ (2007) Light and heavy vehicle technology, 4th edn. Elsevier, Oxford 7. Fischer R, Küçükay F, Jürgens G, Najork R, Pollak B (2015) The automotive transmission book. Springer, Heidelberg 8. Petrov AV (1966) Planetary and hydromechanical transmissions of wheel and track-type vehicles. Mashinostroeniye, Moscow 9. Laptev YN (1973) Automotive torque converters. Mashinostroeniye, Moscow 10. Zhang Y, Mi C (2018) Automotive power transmission systems. Wiley, New York 11. Lapidus VI, Petrov AV (1961) Hydromechanical transmissions of motor vehicles. Mashgiz, Moscow 12. Narbut AN (1966) Torque converters. Mashinostroeniye, Moscow 13. Lapidus VI (1971) Torque converters. Mashinostroeniye, Moscow 14. Trusov SM (1977) Torque converters of motor vehicles. Mashinostroeniye, Moscow 15. Sergeev LV, Kadobnov VV (1980) Hydromechanical transmissions of high-speed track-type vehicles. Mashinostroeniye, Moscow 16. Mazalov ND, Trusov SM (1971) Hydromechanical Gearboxes. Mashinostroeniye, Moscow 17. Gavrilenko BA, Semichastnov IF (1969) Hydraulic couplings and transformers. Mashinostroeniye, Moscow 18. Hydromechanical Transmissions. Design, Production, Operation (1980) Mashinostroeniye, Moscow 19. Stesin SP (1996) Parameter optimization of hydrodynamic drives of construction and road machines. Mashinostroeniye, Moscow 20. Narbut AN, Brimmer AA (1972) Reverse torque converter. SU Patent 331202, 7 Mar 197 21. Narbut AN (1975) Torque converter. SU Patent 356377, 25 Oct 1975 22. Trushin NN, Orlov AB (1996) Controlled complex torque converter. RF Patent 2065103, 10 Aug 1996

Stabilization of Biaxial Trailer Motion Yu. Stroganov, A. Popova and D. Zhelev

Abstract The paper studies the problem of stabilizing the motion of wheels of a biaxial tractor-trailer. Design solutions are proposed of stabilization devices for wheeled bogies of tractor-trailers. Force responses were determined, which impact the stabilization of wheels of a bogie equipped with a swivel which rotation axis (a kingpin) is mounted atilt to the vertical in the vertical longitudinal plane (kingpin caster). The relations were obtained between the moments of reactions acting on the bogie wheel innermost to the center of cornering and rotation angles of the running axis of its wheels. The analysis of these relations shows that the kingpin caster of a biaxial trailer bogie helps to generate such reactions and moments of force which stabilize the motion of front wheels—i.e., to keep the neutral position corresponding to linear motion, with the cumulative moment of stabilization getting the maximum value when the motion is linear and, thus, improving such an operating property as the stability of linear motion. The increase of that stability ensures a safer operation of tractor-trailers and provides the opportunity of fuller utilization of velocity and power characteristics of modern powerful tractors for transportation operations.














Keywords Tractor-trailer Biaxial trailer Bogie Swivel Kingpin caster Wheel stabilization Stabilization device Moment of wheel stabilization







1 Introduction The use of tractor-trailers for cargo transportation in agriculture industry, wood industry, and other sectors of national economy is justified in many cases—it substantially increases capacity and reduces costs of transportation. It is achieved owing to the fact that tractor-trailers have an advantage over motor cars when it comes to cross-country or off-road driving [1–3]. Tractor transport characteristics are improved mainly by increasing load-bearing capacities or velocities. Potentially Yu.Stroganov (&)
A. Popova
D. Zhelev Ural Federal University, 19, Mira Street, Yekaterinburg 620002, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_79

737

738

Yu. Stroganov et al.

more powerful tractors with higher velocity performance come into service [4]. But those capabilities cannot be fully realized due to significant velocity lag of their trailers. Tractor-trailers are utilized in various situations, including their usage on highways along with other vehicles. The maximum allowed value of a trailer’s lateral deviation due to its transverse vibration (wandering) is a very important requirement for the traffic safety. That kind of motion instability increases the overall width of the lane of a vehicle, leads to its sideslip and, therefore, may create a threat to the traffic safety—both for oncoming and passing vehicles and the tractor-trailer itself [5–8]. There are two types of mechanisms which are commonly used to steer the front wheels: (1) a wheel turning mechanism by means of a steering linkage, (2) and the one for wheels which are mounted on a rotating bogie connected to the trailer frame with a swivel. One reason for the loss of motion stability could be a destabilization of the trailer’s steering wheels [9–11]. The stabilization of the wheels of a biaxial trailer’s leading bogie means the ability to maintain a neutral position, which corresponds to linear motion and reverts to it by itself after driving out of a road bend (cornering) [12, 13]. For the trailers with the steering linkage, that stabilization is provided by a kingpin caster, for the trailers with the swivel, the stabilization is not applicable due to low-speed motion. The control of tractor-trailers with a low stabilization of the trailer bogie is difficult [14], and the movement is unstable.

2 Object and Method According to GOST P 52746-2007 “Tractor-trailers and semitrailers”, a full-laden trailer or semitrailer combined with a traction-providing tractor must not override its lane (0.5 m in width more than the tractor-trailer’s width) at any velocity while driving in linear motion on a leveled hard-surfaced road. The motion of the steering-linkage trailers [15] is more stable at higher velocities than that of the bogie-based (swiveled) ones due to the wheels of the former being mounted at more efficient angles. There are several types of mechanic, hydraulic, and other devices which are recommended for the stabilization of the wheels of the bogie-based trailers [16–21]. But stabilization devices described in papers or patents are rather difficult to make and not reliable enough, so they have failed to be widely used. As a result, the prevention of destabilization-caused accidents for bogie-based trailers is provided, nowadays, only by limiting operating speeds of their tractors. A simpler design of the stabilization device for a wheeled bogie with a swivel could be proposed. It is patented by the authors of the paper [17] and uses a slewing unit with a kingpin caster, i.e., with a longitudinal tilt of the axis of the swivel, which connects the trailer frame and the bogie frame. Figure 1 shows the kinematic diagram of a biaxial trailer with the slewing mechanism which delivers the

Stabilization of Biaxial Trailer Motion

739

Fig. 1 Slewing unit of biaxial trailer with kingpin caster (1—swivel (slewing platform); 2—trailer frame; 3—towbar; 4—bogie frame; 5—kingpin; 6—axis of trailer’s leading wheels)

specialization of bogie wheels by means of the kingpin caster. This slewing mechanism differs from conventional designs in that its slewing platform (swivel) 1 is mounted on the frame 6 of the bogie atilt to the frame 2 of the trailer in its vertical longitudinal plane; and the axis of the kingpin 5 is arranged at an angle µ to the vertical line (which goes through the middle of the axis of the bogie wheels), crossing the horizontal motion plane at a forward displacement in relation to the bogie’s wheel axis. As a result of using the kingpin caster, when the bogie and its wheel axis rotate, there occurs a stabilizing moment, which helps to return the wheels of the bogie into a neutral position corresponding to linear motion. Figure 2 presents a diagram of forces which affect the wheels of the bogie while it is cornering. The kingpin axis BC is tilted to the motion surface and forms an angle µ with the vertical line. The projection of the running axis on the motion plane forms an angle h with the projection of this axis during linear motion. The method of determining the stabilizing moment about the point of intersection of the kingpin axis and the road plane for different transport means is described by various authors in [9–11, 15]. When a tractor-trailer makes a turn, the wheels of its bogie are affected by the following reactions: • X—tangential reaction of the road (ground); • Y—lateral reaction of the road (ground); • ZR—vertical reaction of the road (ground). Let us determine the lateral wheel reaction, which acts along the running axis of the bogie wheels. To facilitate the procedure of determining the reaction Y, components of the reaction ZR, and their force levers, consider the motion of the running axis in a formalized form when a turn is made to the right by a h angle from the neutral position (Fig. 3). The turning of the running axis occurs in a plane tilted at a l angle to the horizontal longitudinal axis of the trailer. To determine the moment MY from the reaction Y, it is necessary to evaluate that reaction and the lever Bd.

740

Yu. Stroganov et al.

Fig. 2 Forces affecting the bogie wheel innermost to the center of cornering

Fig. 3 On determination of the moments MY, MX, MZR from the reaction of the road (ground) to the bogie wheel innermost to the center of turning

Stabilization of Biaxial Trailer Motion

741

When the running axis is rotated by the h angle, it is tilted at a a angle to the motion surface and elevated to a h height from the neutral position. Let us draw a segment KM from the point of contact K between the wheel and the road and normal to the neutral position of the running axis. Then, from the KNM triangle, it follows: h ¼ KN
sin l:

ð1Þ

Regarding that KN = L sin h, where L is the length of the semiaxis of the bogie wheels, we get h ¼ L
sin h
sin l:

ð2Þ

From the KCO triangle, sin a could be found (a is the angle at which the running wheel axis is tilted to the horizontal plane): sin a ¼

h L
sin h
sin l ¼ ; L L

a ¼ arcsinðsin h
sin lÞ:

ð3Þ ð4Þ

The lever Bd between the intersection point of the kingpin axis and motion plane and the direction of the Y-reaction component would be determined from the OBd triangle: Bd ¼ OB
sin c;

ð5Þ

where c is an angle between the projection of the running axis at a road bend and a segment OB, which connects the intersection point of the kingpin B and the projection of the center of the running axis O on the motion surface. Regarding that c ¼ 90  h, OB could be found: OB ¼ h
tgl;

ð6Þ

Bd ¼ h
tgl
sinð90  hÞ;

ð7Þ

Bd ¼ h
tgl
cos h:

ð8Þ

then

Substituting h in this formula with the expression from Eq. 2, we get Bd ¼ L
sin h
sin l
tgl
cos h:

ð9Þ

742

Yu. Stroganov et al.

The reaction Y along the running axis would be determined as a component of the reaction Z acting vertically: Y ¼ Z
sin a:

ð10Þ

The component Y1 of the reaction ZR, which acts along the projection of the running axis on the motion plane and causes the moment about the point B, is determined as Y1 ¼ ZR
sin a
cos a:

ð11Þ

Regarding the expressions in Eqs. 3 and 4, we get: Y1 ¼ ZR
sin a
cosðarcsinðsin h
sin lÞÞ:

ð12Þ

To determine the moment of the component of the ZR reaction, which acts in the motion plane on the wheel about the point B and transversely to the semiaxis L, let us decompose it as a force Z′R (Fig. 3), parallel to the kingpin (and therefore not developing a moment of force in relation to its axis), and a force ZRl, located in the road plane and parallel to the OB line (which connects the intersection point of the kingpin and the motion surface at a tilt and the projection of the center of rotation of the wheel axis on the road). Then, we would have: MZR ¼ ZRl
L
cos h;

ð13Þ

where L is a half of the running axis of the wheels. ZRl could be determined as ZRl ¼ tgl
ZR ;

ð14Þ

The moment from the reaction ZR is evaluated as MZR ¼ L
ZR
cos h
tgl:

ð15Þ

To find the moment of the reaction X, which acts in the motion plane on the wheel about the point B, let us determine the lever Kd: Kd ¼ od þ oK:

ð16Þ

From the triangle oBd, od would be found as qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ðOBÞ2 ðBd Þ2 ;

ð17Þ

qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ðL
sin h
sin lÞ2 ðL
sin h
sin l
tgl
cos hÞ2 :

ð18Þ

od ¼ od ¼

Stabilization of Biaxial Trailer Motion

743

OK, equal to the projection of the semiaxis L, is determined as OK ¼ L
cos a;

ð19Þ

which gets us the lever Kd as Kd ¼

qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ðL
sin h
sin lÞ2 ðL
sin h
sin l
tgl
cos hÞ2 þ ½L
cosðarcsinðsin h
sin lÞÞ:

ð20Þ The tangential reaction of the road X is evaluated as X ¼ ZR
f ;

ð21Þ

where f is the coefficient of road resistance. Regarding the expressions obtained for the reactions XZY and their force levers in relation to the point B, we get the following moments: MY ¼ L
ZR
sin2 h
sin2 l
cosðarcsinðsin h
sin lÞÞ
tgl
cos h; 0 qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 ðL
sin h
sin lÞ2 ðL
sin h
sin l
tgl
cos hÞ2 A MX ¼ ZR f
@ ; þ L
cosðarcsinðsin h
sin lÞÞ MZR ¼ L
ZR
cos h
tgl:

ð22Þ ð23Þ ð24Þ

A wheel equipped with an elastic tyre is also affected by the moment of cornering resistance of the tyre MT, which occurs due to the displacement of the X and Y reactions from the center of contact when wheels slip. The moment of cornering resistance MT is taken from experimentally obtained data [9–11]. A diagram showing the relation between MT and the rotation angle is in Fig. 4. Taking a trailer 2-PTS-4 (with a swivel enabling the rotation of the leading bogie up to a 90° angle in relation to its platform) as an example, let us consider the behavior of stabilizing moments MZR, MX, MY, MT. While doing so, we make some basic assumptions: an absolutely rigid suspension of the leading bogie, lack of torsional deformation of the trailer frame, lack of wheel slip, full usage of the trailer’s bearing capacity. The caster of an imaginary kingpin is 5°. Under these assumptions, rotation of the leading wheeled bogie due to the inclination of the running axis by an a angle is accompanied by lifting of the bogie wheel innermost to the center of cornering from the road surface. Then, the whole load from the trailer weight applied to the bogie is redistributed toward the wheel innermost to the center of cornering (it is not a typical case in the operation of tractor-trailers, since their spring suspensions, frame torsion, elasticity of wheels usually ensure a fixed contact with the road).

744

Yu. Stroganov et al.

Fig. 4 Stabilization moment and its components

Figure 4 demonstrates relations of the moments MZR, MX, MY, MT, along with the cumulative moment of wheel stabilization MST, to the angle h of the bogie cornering relative to the trailer frame. Assume that the moment of the reaction Y (acting along the running axis) is negative since the direction of its force causes a moment opposing the moments MZR, MX, MT (which form the cumulative stabilization moment MST). The analysis of the graphs above allows to make the following conclusions: – A kingpin caster of the bogie swivel for a biaxial trailer helps to generate such reactions and moments of force which stabilize the motion of front wheels—i.e., to keep the neutral position of linear motion. – At a road bend, the kingpin caster contributes to redistributing of the load from the trailer weight toward the wheel innermost to the center of cornering, which increases the trailer stability against tilting due to centrifugal forces. The calculated data shows that an increase of the folding angle between the bogie and the trailer platform leads to a decrease of the stabilization moment, but this effect is a rather positive one as not hindering the cornering at greater angles (when lower velocities are usually used).

Stabilization of Biaxial Trailer Motion

745

References 1. Grebnev VP, Polivaev OI, Vorohobin AV (2011) Automobiles and tractors. Theory and operational propeties, Knorus, Moscow 2. Ignatov VD (1978) Organization of cargo transportation in collective and state farms. Rosselhozizdat, Moscow 3. Turaevskiy IS (2011) Motor transportation. Izdatelskiy dom “Forum”: Infra-M, Moscow 4. Bazhenov EE (2000) Teoriya avtomobilya i traktora (Automobile and tractor theory). UGTU-UPI, Ekaterinburg 5. Protas AY (1972) Influence of the hook power on the motion stability of a “Belarus” tractor trailer. In: Proceedings of the scientific and technical conference on increasing of towing abilities and passabilities of 1.4-ton wheeled tractors, Gorki 6. Petrenko AM (2013) Ustoychivost specialnyh transportnyh sredstv (Stability of special transport). MADI, Moscow 7. Chudakov EA (1950) Konstrukciya i raschet avtomobilya (Design and calculation of automobiles). Mashgiz, Moscow 8. Shinkarenko AA, Kuyukov VV (2013) Active tractor-trailer safety in the traffic. J Vestnik SibADI 4(32):35–41 9. Ivanov VV et al (1977) The basic automobile and tractor theory. Visshaya shcola, Moscow 10. Litvinov AS, Pharobin YE (1984) Automobile: Theory of operational properties. Mashinostroenie, Moscow 11. Zakin YH (1967) Prikladnaya teoriya dvizheniya avtopoezda (Applicable theory of tractor-trailer motion). Transport, Moscow 12. Ivanov AM, Narbut AN, Parshin AS et al (2014) Avtomobili: Teoriya ekspluatacionnyh svoystv (Automobiles: Theory of operational properties). Izdatelskiy centr “Akademiya”, Moscow 13. Vahlamov VK (2010) Avtomobili: Ekspluatacionnye svoystva (Automobiles: Operational properties). Izdatelskiy centr “Akademiya”, Moscow 14. Zakin YH (1986) Manevrennost avtomobilya i avtopoezda (Maneuverability of automobile and tractor-trailers). Transport, Moscow 15. Balakina EV (2004) K voprosu o velichine ugla prodolnogo naklona osi shkvornya upravlyaemogo kolesa avtomobilya (On value of the caster of an automobile steering wheel). J Izvestiya VolgGTU pp 81–85 16. Stroganov YN, Zhelev DY, Stroganova OY (2018) Tyagovo-scepnoe ustroystvo avtomobilnogo polupricepa (Towing coupler of a semitraler). RF Patent 177788 17. Stroganov YN, Popova AI (2017) Oporno-povorotnoe ustroystvo pricepa (Rotary support of a trailer). RF Patent 170879 18. Stroganov YN, Zhelev DY (2018) Stabiliziruyushchee ustroystvo povorotnoy telezhki pricepa (Stabilizer of a trailer bogie). RF Patent 181371 19. Stroganov YN, Popova AI, Stroganova OY (2017) Podkatnaya telezhka dlya buksirovki avtomobilya metodom chastichnoy pogruzki (Dolly for towing an automobile by means of partial loading). RF Patent 174781 20. Stroganov YN, Popova AI, Lyahov SV, Stroganova OY (2018) Oporno-povorotnoe ustroystvo dvuhosnogo pricepa (Rotary support of a two-axle trailer). RF Patent 182885 21. Stroganov YN, Zhelev DY, Stroganova OY (2018) Tyagovo-scepnoe ustroystvo polupricepa s naklonnoy povorotnoy platformoy (Towing coupler of a semitrailer with an inclined rotary platform). RF Patent 185183

Development of Optimization Algorithm to Control Open-Pit Excavator Operation O. Lukashuk, A. Komissarov and K. Letnev

Abstract It is shown that in the process of rock excavation, the joint action of the main actuating mechanisms (lifting and thrusting mechanisms) of an open-pit excavator with the operational equipment of a front shovel type forms a leverage mechanism connecting the main mechanisms of the excavator and its bucket and transforming their operating parameters into parameters of energy and force realized on the cutting edge of a bucket (its teeth) in accord with properties of the mechanism kinematics. As a result of the analysis of kinematics and forces of the leverage, expressions were obtained for kinematic transfer functions of that mechanism, which define relations between the velocity of bucket’s cutting edge (excavation velocity) and velocities of operating motions (lifting and thrusting) in the case of the bucket moving within the working area of the excavator. An optimization algorithm to control the operation of an open-pit excavator was developed, which allows obtaining desired values of the operating parameters of its main mechanisms when they are utilized in the process of excavation on an excavated face. Keywords Main actuating mechanisms of open-pit excavators anism Transfer functions Optimization algorithm of control








Leverage mech-

1 Introduction Operating modes of engines used in main actuating mechanisms of an open-pit excavator with the operational equipment of a front shovel type are characterized by many switches on and off, considerable variation of loads and velocities of working motions. Controlling the operation in those conditions (a lot of information, lack of time) is a quite difficult task to accomplish.

O. Lukashuk (&)
A. Komissarov
K. Letnev Ural Federal University, 19, Mira Street, Yekaterinburg 620002, Russia e-mail: [email protected]; [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_80

747

748

O. Lukashuk et al.

The analysis of how the process of excavation is controlled shows that the control logic is quite complex, with a lot of data and lack of advanced technology to obtain it. Nowadays, progress is made toward developing such control systems which combine both operator- and PC-based approaches [1–14]. Increased quality and efficiency of control could be based on the formalization of the rock excavation process and determination of relations between parameters of energy and force realized on the cutting edge of a bucket (its teeth) and operating modes of the main mechanisms while taking into consideration mining conditions of excavation and parameters of an excavated face.

2 Aim of the Research The main focus of the research was on increasing the efficiency of controlling the working process of excavation by identifying operating modes of the main mechanisms of an excavator. The problems solved in the study are: • determination of relations between the operating modes and the parameters of excavation—path (trajectory) which the bucket (its top of the cutting edge) follows, excavation velocity, etc.; • definition of transfer functions for the leverage formed by the joint action of the main mechanisms in the process of excavation; • development of an optimization algorithm to control the working process, helping to minimize the time needed to move the bucket from one position to another.

3 Solution of the Problems The research analyzes the operation of excavation carried out by an open-pit front shovel excavator. The excavation is accomplished by the main mechanisms (lifting and thrusting mechanisms) which act jointly to move the bucket (its top of the cutting edge) along equidistant (face-slope parallel) paths and, at the same time, help it to separate layers (“chips”) of the rock mass (Fig. 1). The object of the research is the analysis of operating modes of the lifting and thrusting mechanisms while excavating the rock mass from an excavated face. The process of excavation involves a two-crank leverage formed in the operational equipment of an excavator, which consists of the links of its lifting mechanism (head block of a boom and its lifting rope) and its trusting mechanism (a saddle bearing and arm with a bucket) and connects the main mechanisms and the bucket (Fig. 2). In that case, the operating parameters of the mechanisms (forces and velocities of lifting and thrusting) are transformed, in accord with the properties

Development of Optimization Algorithm to Control Open-Pit …

749

Fig. 1 Path of digging for the bucket of an open-pit front shovel excavator: a mechanical shovel; b hydraulic excavator

Fig. 2 Diagram of the mechanism of the operational equipment: 1, 4—two cranks; 2—arm and bucket; 3—lifting rope; VL, VT, VE—velocities of lifting, thrusting and excavating

of the leverage mechanism kinematics, into the parameters of energy and force which characterize the bucket operation. Based on a mathematical model of the excavation, expressions for kinematic transfer functions of the leverage mechanism were obtained, which allow to evaluate the velocities of operating motions (lifting and thrusting) on the condition of a moving bucket maintaining a desired velocity and duration of excavation [15–18].

750

O. Lukashuk et al.

In their general form, those expressions for the kinematic transfer functions look like: UL ¼ VL =VE ¼ f1 ðXK ; YK ; w; lln k Þ;

ð1Þ

UT ¼ VT =VE ¼ f2 ðXK ; YK ; w; llnk Þ;

ð2Þ

where VL, VT, VE are velocities of lifting, thrusting, and excavating, respectively; XK, YK—coordinates of the point K (top of bucket’s cutting edge); w—angle of the tangent to a bucket path at K; llnk—geometric dimensions of the leverage links. Figure 3 shows graphs of the transfer functions for velocities of lifting and thrusting in the case of a bucket moving with a constant velocity of excavation within the work area of an EKG-20 excavator manufactured by JSC “Uralmashplant”. The proposed method of estimating the velocities of operating motions would allow to coordinate those motions of lifting and thrusting and their corresponding velocities in the process of excavation. An optimization algorithm to control an excavator which works in an open pit was developed with the purpose of obtaining the desired values of the operating parameters of the main mechanisms for a bucket moving with a given velocity of excavation within the work area.

Fig. 3 Hodographs of velocities of operating motions for the bucket moved up to a calculated height of excavation: VL, VT—lifting and thrusting velocities; A, B, C—trajectories of the bucket moved forward until the boundaries of the work area (A and C), and in the middle of the area (B)

Development of Optimization Algorithm to Control Open-Pit …

751

Fig. 4 Algorithm to control the bucket which excavates while moving along linear paths and with a constant velocity of excavation

The algorithm (Fig. 4) determines the specifics and sequence of running the following operations, which allow to move the bucket along a given path: • computer-based calculation of lifting and thrusting velocities for a bucket in its initial (or other) excavating position, or in its following positions corresponding to the bucket moved forward with a given stride;

752

O. Lukashuk et al.

• determination of average values of the velocities in two positions and transmittance of commands proportional to those values onto the input of the system controlling the engines of the mechanisms; • relocation of the bucket to the next position.

4 Conclusion Development of an optimization algorithm of controlling the bucket motion of an excavator allows to obtain optimal values of its performance criteria (minimized time of moving a bucket along a given path and minimized total energy costs). The control system for the process of excavation based on calculated values of kinematic transfer functions for a leverage mechanism of an open-pit excavator’s operational equipment utilized in given mining conditions ensures increase in efficiency of its operation owing to the coordination of its working motions and lifting and trusting velocities on the condition of maintaining a desired velocity and duration of excavation.

References 1. Bender FA, Sawodny OA (2014) Predictive driver model for the virtual excavator. In: Proceedings of the 13th international conference on control, automation, robotics and vision (ICARCV), Nanyang Technological University, Singapore, 10–12 Dec 2014 2. Berns K, Proetzsch M, Schmidt D (2010) Simulation and control of an autonomous bucket excavator for landscaping tasks. In: Proceedings of the IEEE international conference on robotics and automation (ICRA), Anchorage, Alaska, USA, 3–8 May 2010 3. Lee B, Kim HJ (2014) Trajectory generation for an automated excavator. In: Proceedings of the 14th international conference on control, automation and systems (ICCAS), KINTEX, Seoul, 22–25 Oct 2014 4. Le QH, Jeong YM, Nguyen CT et al (2013) Development of a virtual excavator using simmechanics and simhydraulic. J Drive Control 10:29–36 5. Liu J (2009) Integrated mechanical and electrohydraulic system modeling and virtual reality simulation technology of a virtual robotic excavator. In: Proceedings of IEEE 10th international conference on computer-aided industrial design & conceptual design, Wenzhou, China, 26–29 Nov 2009 6. Tao N et al (2008) A low-cost solution for excavator simulation with a realistic visual effect. In: Proceedings of IEEE conference on robotics, automation and mechatronics, Chengdu, China, 21–24 Sept 2008 7. Babakov SE, Pevzner LD (2012) Algoritmizaciya upravleniya dvizheniem kovsha ekskavatora v rezhime cherpaniya s primeneniem nechetkoy logiki (Algorithmizing the control of excavator bucket motion in excavation via fuzzy logic). J Gornoe oborudovanie i elektromehanika 9:8–17 8. Druzhinin AV (2016) Povyshenie kachestva upravleniya odnokovshovymi ekskavatorami na osnove multiagentnogo podhoda (Improving the control of single-bucket excavators using a multiagent approach). J Novye ogneupory 3:11–12

Development of Optimization Algorithm to Control Open-Pit …

753

9. Pevzner LD (2014) Avtomatizirovannoe upravlenie moshhnymi odnokovshovymi ekskavatorami (Automated control of high-duty single-bucket excavators). Gornoe delo, Moscow 10. Poderni RY (2013) Mehanicheskoe oborudovanie karierov (Mechanical equipment of open pits). Mining Media Group, Moscow 11. Korchagin PA, Korchagin EA (2018) Razvitie nauchnyh osnov proektirovaniya vibrozashhitnyh sistem zemleroynyh mashin (Advances in research principles to design vibration protection systems of earth-moving machinery). SibADI, Omsk 12. Korchagin PA, Teterina IA (2015) Matematicheskaya model slozhnoy dinamicheskoy sistemy «vozmushhayushhie vozdeystviya—mashina—operator» (Mathematical model of a complex dynamic «perturbation action—machine—operator» system). J Vestnik SibADI 5:118–123 13. Malafeev SI, Tihonov YV (2015) Intellektualizaciya kariernogo ekskavatora (Intellectualization of open-pit excavators). J GIAB 11:107–115 14. Malafeev SI, Serebrennikov SA (2018) Povyshenie energeticheskoy effektivnosti kariernyh ekskavatorov na osnove modernizacii elektrooborudovaniya i sistem upravleniya (Improving the energy efficiency of open-pit excavators by modernizing electrical equipment and control systems). J Ugol 10:30–34 15. Gafurianov RG, Komissarov AP, Shestakov VS (2009) Modelirovanie rabochego processa kariernyh ekskavatorov (Modeling of the working process of open-pit excavators). J Gornoe oborudovanie i elektromehanika 6:40–45 16. Lagunova YA, Komissarov AP, Shestakov VS et al (2011) Gornye mashiny. Enciklopediya (Mining machines. Encyclopedia), vol 4, issue 24. Mashinostroenie, Moscow 17. Ivanov IY, Komissarov AP, Lagunova YA et al (2015) Intensifikacija processov ekskavacii gornyh porod (Intensification of the processes of rock excavation). J Izvestija vuzov. Gornyj zhurnal 3:94–99 18. Komissarov AP, Lagunova YA, Shestakov VS (2017) Proektirovanie kariernyh ekskavatorov (Design of open-pit excavators). Innovacionnoe mashinostroenie, Moscow

Complex for Inspection of Crane Rails Design V. Yu. Antsev, P. V. Vitchuk and K. Yu. Krylov

Abstract Technical condition of the crane rails has a direct impact on the performance, dependability and noiseless operation of the load-lifting crane. Therefore, to ensure the correct and safe operation of the load-lifting crane, systematic surveys of crane track are carried out. Crane rails inspection is associated with significant risks since is carried out at high altitude. This encourages researchers to automate known methods using various devices and complexes, as well as to develop new methods of control to virtually eliminate the works at high altitude done by humans. There are known various designs of automatic and semiautomatic systems for determining different parameters of the crane rails. Such complexes include a wide range of equipment by different principles of operation. They are not widely used, mainly due to the fact that acceptable measurement accuracy is provided along with a significant increase in the cost of the complex. This determines the relevance of the development of complex for crane rails inspection, which provides sufficient accuracy and has a relatively low cost. Requirements for the developed complex can be divided into two parts: requirements for measurement accuracy and requirements for ensuring the normal functioning of control and measuring equipment. The implementation of these requirements will provide a relatively inexpensive design that secures acceptable measurement accuracy. The developed complex will reduce the cost of works on a comprehensive survey and will allow more efficiently monitor the technical condition of crane rails and therefore—to increase the level of industrial safety as a whole. Keywords Load-lifting crane Dependability


Defect
Inspection
Complex
Crane rail


V. Yu. Antsev FSBEE “Tula State University”, 92, Lenin Avenue, Tula 300012, Russia P. V. Vitchuk (&) Kaluga Branch of FSBEE “Moscow State Technical University, After N.E, Bauman (National Research University)”, 2, Bazhenov Street, Kaluga 248000, Russia e-mail: [email protected] K. Yu. Krylov Regional Engineering and Technical Center, 52, Glagoleva Street, Kaluga 248021, Russia © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_81

755

756

V. Yu. Antsev et al.

1 Introduction Rail-mounted lifting cranes are an important part of the transport and logistics process in industrial enterprises and construction sites, dockyards, etc. Technical condition of the crane track has a direct impact on the performance, durability and noiseless operation of the load-lifting crane. Therefore, to ensure the correct and safe operation of the load-lifting crane, systematic surveys of crane rails are carried out. During the survey of crane rails, it is necessary to control its planned high-altitude position, the wear of the railhead, the condition of the track joint fastenings and intermediate fastenings, and other parameters [1]. Various geodetic devices are used for this purpose [2, 3]: theodolites; level rulers; leveling rods; levels of various designs; and laser distance meters. Inspection of the crane track is associated with significant risks since, as a rule, is carried out at high altitude. This encourages researchers to automate known methods using various devices and complexes, as well as to develop new methods of control to virtually eliminate the works at high altitude done by humans. There are known domestic [4–8] and foreign [9–11] automatic and semiautomatic complexes for determining different parameters of the crane rails. Such complexes include a wide range of equipment by different principles of operation: laser and screen brand, theodolite, tacheometer, etc. They are not widely used for a number of reasons, mainly due to the fact that acceptable measurement accuracy is provided along with a significant increase in the cost of the complex. This determines the relevance of the development of complex for crane rails survey, which provides sufficient accuracy and has a relatively low cost.

2 Development of Requirements to the Complex Forthe Crane Tracks Inspection In general, requirements for the developed complex can be divided into two parts: requirements for measurement accuracy (provided by the appropriate control and measuring equipment) and requirements for ensuring the normal functioning of control and measuring equipment. When establishing the requirements for the measurement accuracy, it is necessary to analyze the measured parameters. Permissible deviations of the rail crane track from the design position in plan and profile views are regulated in the Russian Federation [12] and are given in Table 1. There are several ways to move from the permissible deviation d to the root-mean-square deviation m of the measurements. An analysis of these methods is carried out, and a formula is proposed in the item [3]:

Complex for Inspection of Crane Rails Design

757

Table 1 Permissible deviations of the rail crane track from the design position in plan and profile views [12] Parameter

The difference in grades of railheads in cross-direction The difference in grades on tracks along the crane track Narrowing and widening of the track gage Mutual displacement of the ends of the joined rails in the plan and in height Clearances in rail joints The difference of altitude marks of railheads at 10 m length of the crane track

Indication

Permissible deviation, mm Crane type Overhead Tower Portal traveling crane bridge crane crane

Portal slewing crane

Overload crane

P1

40

45–60

40

40

50

P2

10

–

–

–

–

P3

15

10

15

15

20

2

3

2

2

2

40

30

20

30

P4 ; P04

P5 P6

6 –

m¼

0:4d ; mm t

ð1Þ

where t—normalization factor, which corresponds to a defined value of probability belief. Table 2 shows the results of the calculation of the root-mean-square deviation of measurements on the example of overhead traveling cranes. As can be seen in Table 2, the requirements for the accuracy of measurements of the crane rails parameters are quite high. In addition to ensuring the requirements for measurement accuracy, the complex should function properly in a wide variety of application conditions. Experience has proven that the actual technical condition of the crane rails may differ significantly from the established one. It means that defects that significantly exceed the standard values may occur on the crane rails in real operating conditions. This leads to increased requirements for rigidity and pass ability of the complex. To assess the actual technical condition of crane rails, the analysis of more than 50 survey results of the overhead traveling crane rails operated in the Kaluga region was carried out. The results are presented in Table 3. As can be seen from Table 3, in some cases, the defects that significantly exceed the standard values were detected on the crane rails. The presence of such defects can lead to the fact that the complex will get stuck on the crane rails or fall from a considerable height.

758

V. Yu. Antsev et al.

Table 2 Results of calculation of root-mean-square deviation in measurements (by the example of overhead traveling cranes) using a formula (1) Parameter

P1 P2 P3 P4 ; P04 P5

Permissible deviation, mm

40 10 15 2 6

Probability belief 0.89 0.955 0.988 Root-mean-square deviation of measurements, mm 10.00 2.50 3.75 0.50 1.50

8.00 2.00 3.00 0.40 1.20

6.40 1.60 2.40 0.32 0.96

0.997

5.33 1.33 2.00 0.25 0.80

Table 3 Defects of bridge crane rails Defect type The difference in grades of railheads in cross-direction (P1), mm The difference in grades on rails along the crane rail (P2), mm Narrowing and widening of the rail gage (P3), mm Mutual displacement of the ends of the rails in the plan and in height P4 ; P04 , mm Clearances in rail joints (P5), mm Deviation of the guide rail from the vertical axis, deg Absence/weakness of intermediate fastenings Absence/weakness of Raul joint fastenings Defects of stop bars and shock-absorbing elements Defects of current-flow elements General dust on the crane rail, the presence of foreign objects Other defects Total

Number of defects

Defect dimension

12

41…60

20

12…36

9 8

16…35 5…11

7 3 10 6 33 5 3 10 126

6…25 6…13 – – – – – – –

3 Selection of Control and Measuring Equipment Despite the fact that the parameters measured by all known complexes are the same, these complexes contain different control and measuring equipment by the principle of operation. While developing of the complex for measuring of the crane rails defects, it is necessary to make a reasonable choice of control and measuring equipment, which includes the best combination of characteristics such as accuracy, measurement convenience, convenience of the desk study of measurement results, and cost.

Complex for Inspection of Crane Rails Design

759

This task solution can be achieved by using the methods of qualimetric estimation as an example. Processing of the results of the qualimetric estimation is proposed to be carried out using the criterion of G. Taguchi signal/noise S=N. In this case, the determination of the value of the signal/noise criterion is based on the results of a survey of experts on the applicability of the equipment type for the control
of various defects. The response ðyi Þ—is estimation done by an expert, and factor xj , which impact on expert’s estimation (response) is equipment used for the crane rails survey. The criterion signal/noise is to be defined by formula: S=N ¼ 10
lg

 2 y ; mm D

ð2Þ

Pn 1 yÞ 2 — where y—mathematical expectation of the response; D ¼ n1 i¼1 ðyi   dispersion; yi —response (estimation by expert); n—number of experts. The higher the signal-to-noise ratio, the more preferable the equipment type for crane rails inspection. Expert estimation of the equipment applicability is proposed to be carried out for three evaluation groups: planned altitude survey, control of joints and intermediate fastenings, and control of the distance passed. For each group of the measurements, the corresponding list of equipment can be determined and the corresponding values of the criterion S=N can be calculated (Table 4). The table shows several equipment types that can be used to carry out the selected groups of measurements.

Table 4 List of equipment for various groups of measurements of crane tracks parameters Equipment

Laser, screen, distance meter Theodolite Tacheometer Laser tracker Level ruler, distance meter Camera Laser scanner Set for visual and measuring control Odometer Laser rangefinder

S=N Group of measurements Planned Control of joints and altitude intermediate fastenings survey

Control of the distance passed

16.46

–

–

11.87 21.32 19.83 16.37

– – – –

– 29.59 25.41 –

– – –

28.32 17.46 16.52

– – –

– –

– –

29.33 21.14

760

V. Yu. Antsev et al.

The criteria for the estimation of the applicability of the equipment type for each considered group were the duration of the measurements, the duration of the desk study of measurement results, accuracy, measurement convenience, convenience of the desk study of measurement results, cost, the possibility of installing an automated complex on the carriage frame, and safety of measurement works. Expert estimation of the equipment applicability for the crane rails inspection by listed criteria was based on the proposed rating scale: 2 points—full compliance with criterion; 1 points—partial compliance with criterion; 0 points—noncompliance with criterion. Seven experts competence in the solution of the set task were involved to carry out an expert estimation. According to the results of the calculation of the signal-to-noise criterion, the tacheometer is the most preferred type of equipment for equipping an automated complex and for carrying out the first group of measurements “Planned altitude survey”. According to the expert’s opinion, this type of equipment most fully meets the selected evaluation criteria. For the second group of measurements “control of joints and intermediate fastenings” implementation, the atomized complex should be equipped with a camera. According to the results of calculating the signal-to-noise criterion, there was no definitive device for the implementation of the third group of measurements “Control of the distance passed.” Therefore, control of this group of defects can be made by any of them depending on the existing technological, design, economic, and other constraints.

4 Providing the Rigidity When the complex passes along the crane rail, its oscillations occur due to the dynamics of movement and various rail defects (wear of the railhead, the presence of burrs, cracks, slip marks of the crane wheels, etc.). These oscillations create errors in the operation of control and measuring equipment. Besides, the presence of significant defects in the joints determines the likelihood of the complex getting stuck or falling from a considerable height. This necessitates the development of measures aimed at ensuring the rigidity of the complex. Table 5 provides the list of methods to ensure the rigidity in known complexes [8–11]. Let us analyze the ways of providing rigidity, used in the known measurement complexes [8–11]. Table 5 shows that the side rollers or levers on which they are mounted are adjustable. This is necessary to ensure the operation of the complex on various nominal sizes of rails or other rails used as part of the crane rail.

Complex for Inspection of Crane Rails Design

761

Table 5 Methods to provide the rigidity in known measurement complexes [8–11] Measurement complex MSTU n.a. N.E. Bauman [8]

LMS Demag [9]

KONE RailQ [10]

ARTIS [11]

Spring-loaded adjustable levers with concave shape side rollers

Adjustable cylindrical shape side rollers

Location of the center of gravity below the railhead level, adjustable cylindrical shape side rollers

Adjustable levers with cylindrical shape side rollers

The measurement complex [8] includes concave shape side rollers. Such rollers are used to secure the rigidity of various mechanized track tools. The roller bends around the railhead and based on that ensures the rigidity of the tool on the rail. Crane rails of CR type, in contrast to railways, have a reduced neck height, as a result of which the joint cover passes directly under the railhead. Besides, various pipes, rolling profiles, etc., can be used as the guides for the crane track. Thus, the correct operation of such rollers in the measuring complex is uncertain. Cylindrical shape side rollers are used in the complexes [9–11]. The exact value of the diameter of these rollers is unknown, but it can be assumed that it ranges from 25 to 35 mm according to the available schemes, photographs, and video materials on the operation of these complexes. If there is a clearance in joint of the crane track with 15…25 mm size, it is likely that the complex will not be able to manage this defect. Besides, the side rollers do not have a device for fixing them to the rail in the considered complexes [8–11]. In case of local reduction in the width of the rail (for example, as a result of its wear), it leads to a clearance between the side rollers and the rail causing an error in measuring the crane rail parameters. Thus, none of the considered methods of rigidity providing is the best. As prototypes for designing a device for the rigidity providing of the complex, a device for measuring rail deviations, made in scissors form, with springs and side rollers [13] at both ends, and a guide roller device for cranes with separate adjustment of the distance between the rail and side rollers were taken due to the eccentric and splines [14]. As a result, a device was developed (Fig. 1). The device works as follows. The operating ends of the scissors 4 are installed with the help of angle regulators 9 so that the spring 6 of the scissors handles 5 creates a significant force of the side rollers 3 on the side surfaces of the rail 2. Then the two-wheeled carriage frame is moved along the rail 2. Counterweights 8 provide a stable balance to the two-wheel carriage frame 1. The scissors axis 7 keeps the balance of the upper and lower scissors. There are splines in the operating end of the scissors 4 and the scissors handle 5 of the angular regulator 9. To change the tilt angle of the operating end of the scissors 4 relative to the scissors handle 5, it is necessary to unscrew the nut 11, remove the operating end 4, turn it one or more splines of the scissors handle 5, then to fix the splined shaft 10 with the nut 11.

762

V. Yu. Antsev et al.

Fig. 1 Device for providing the complex rigidity: 1—two-wheeled carriage frame; 2—rail; 3—side roller; 4—operating end of scissors; 5—scissors handle; 6—spring; 7—scissors axis; 8—counterweight; 9—angle regulator; 10—splined shaft; 11—nut

The use of angle regulators with independent adjustment of the pressing force of the upper side rollers and scissors allows you to keep the direction of movement of the control and measuring equipment (mounted on a two-wheeled carriage frame of the measurement complex) strictly along the rail axis. It increases the accuracy of measurement of the rail axis tilt and measurement of deviation in railhead width. In the well-known literature on load-lifting machines [15–21], the following types of surfaces for running wheels and side rollers tread are given cylindrical, conical, and barrel-shaped. At the same time, depending on the combination of parameters of the running wheels or rollers and rails, the contact can be point and linear. In case of a point contact, the contact stresses are higher, but the running characteristics are better (resistance to imbalance, no sliding, etc.). In case of linear contact, the contact stresses are lower, but the running characteristics are worse, as well as there is a possibility of the side roller “dragging” onto the rail. A scheme with a point contact was chosen since there are no significant forces in the measuring complex. The only form of side rollers that guarantees precise contact at any position is barrel-shaped. Therefore, barrel-shaped rollers were used to secure the rigidity of the developed complex. The outer diameter of the rollers (45 mm) was chosen based on the analysis of defects in the joints of crane rails (Table 3). The length of the rollers was chosen based on the size of the railhead which made in accordance with GOST R 53866-2010 (crane) and GOST R 51685-2013 (rail). In order to reduce the weight, the rollers were made of fluorine plastic. There is a hole in the center of the roller in order to mount the roller to the control bar of the operating end of scissors (pos 4. at Fig. 1).

5 Conclusion Thus, the correct choice of control and measuring equipment, along with ensuring the rigidity and pass ability of the complex, will help to reach a relatively inexpensive design that provides acceptable measurement accuracy.

Complex for Inspection of Crane Rails Design

763

The developed complex will make it possible to reduce the operating cost of a comprehensive survey and monitor the technical condition of the crane rails in a more effective way and thus to increase the level of industrial safety in general.

References 1. Federal rules and regulations in the field of industrial safety “Safety regulations for hazardous industrial facilities that use lifting devices” 2. Seroshtan V, Ogar Yu, Golovin A et al (1992) Diagnosis of lifting machines. Moscow 3. Shekhovtsov G (2018) Modern methods of geodetic control of running gear and bridge crane tracks. Nizhniy Novgorod 4. Zharnikov V, Nagorny Yu (1988) Calculation of the accuracy of the automated installation for geodetic control of crane tracks. In: Proceedings of Siberian State University of Geosystems and Technologies vol 37, pp 73–78 5. Chernikov V (1963) Some devices for alignment of crane tracks by geodetic method. In: Proceedings of Siberian State University of Geosystems and Technologies vol. 16, pp 103–111 6. Lambin N (1978) Crane tracks survey using a semi-automatic device. Eng Geod 21:21–25 7. Antsybor V (1985) Laser tools for surveying works. Moscow 8. Dmitriev A, Kondratenko A (2009) Measuring buggies for the crane track diagnosis. In: Proceedings of the XIII International Conference Lifting and Transport, Construction, Road, Track Machines and Robotic Systems. Moscow 9. Tseboev A (2000) Laser calibration system “Mannesman Dematic”. Lifting and transport industry, vol 1, p 21 10. Complex KONE RailQ. Official site of the company KONE official website. Available via DIALOG. http://www.konecranes.ru/servis-i-zapchasti/konsultacionnye-uslugi/proverkasostoyaniya-kranovyh-putey-railqtm 11. Dennig D, Bureick J, Link J, Diener D, Hesse C, Neumann I (2017) Comprehensive and highly accurate measurements of crane runways. Profiles Fastenings Sens 17:1118. https:// doi.org/10.3390/s17051118 12. RD 10-138-97 (1997) Comprehensive survey of crane tracks of lifting machines. General requirements 13. Molˇcan V (2014) Deformation surveying of crane track. Master’s Thesis. Brno University of Technology. Brno. Czech Republic 14. Golder M (2013) Crane rolling guide. RU Patent. 2475443, 20 Feb 2013 15. Alexandrov M (2000) Hoisting machines. Moscow 16. Alexandrov M, Reshetov D, Bajkov B (1987) Hoisting machines. Moscow 17. Kazak S, Dusie V, Kuznetsov E (1989) Course design of hoisting machines. Moscow 18. Ermolenko V (2013) Calculation of hoisting machines mechanisms. Moscow 19. Alexandrov M, Ivashkov I, Kazak S (1993) Calculations of hoisting machines mechanisms and details. Ref. in 2 books. Book. 1. Moscow 20. Alexandrov M, Ivashkov I, Kazak S (1993) Calculations of hoisting machines mechanisms and details. Ref. in 2 books. Book. 2. Moscow 21. Vainson A (1989) Hoisting machines. Moscow

Study of Impact of Amount of Shock Absorbers on Parameters of Vibrations of Drum and Frame of Vibrating Roller I. S. Tyuremnov, D. V. Fyodorova and A. S. Morev

Abstract The article presents the results of mathematical modeling of the oscillations of the roller and the frame of the roller of the vibratory roller with a change in the characteristics of the soil compacted and the number of shock absorbers. The rheological model used makes it possible to analyze the features of the interaction of the vibrating roller drum with the material being compacted in the continuous contact, partial uplift, and double jump modes. The paper presents the results of a computational experiment to study the effect of the number and characteristics of absorbers on the oscillation characteristics of the drum and the drum frame as part of the modernization of the DM-614 vibration roller produced by the Rybinsk Plant of Road Machines with the aim of increasing the efforts of the vibratory pathogen from 215 to 280 kN. The computational experiment was carried out for different numbers of shock absorbers with different values of the coefficient of elastic resistance of the soil. The objectives of the research were to determine the magnitude of oscillations of the vibrating roller and the frame of the drum of the vibrating roller DM-614 in the process of soil compaction. The obtained results make it possible to justify the number and characteristics of the vibration roller shock absorbers to ensure a safe level of vibrations transmitted to the operator’s cabin, considering the use of a vibrating roller in compacting various road construction materials in a wide range of variations in compaction coefficients.





Keywords Soil Compaction Vibratory drum Drum frame


Vibration
Impact force
Vibrating roller


1 Introduction Vibrating rollers are the most common means of mechanical soil and stone compaction in road construction. The main technical characteristics of the vibrating soil rollers that define their technological capabilities during compaction of road I. S. Tyuremnov
D. V. Fyodorova
A. S. Morev (&) Yaroslavl State Technical University, 88, Moskovsky Prospect, Yaroslavl 150023, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_82

765

766

I. S. Tyuremnov et al.

construction materials are operating weight, mass per module of the vibratory drum; the driving force and the vibration frequency; the diameter and width of the vibratory drum; engine capacity, and some other parameters. Despite many years of experience in the production and operation of vibrating rollers, manufacturers still can’t find common ground on numerical values of the above-mentioned parameters, which lead to a significant dispersion of values of the above-mentioned parameters of models of vibrating rollers with similar mass produced by different manufacturers [1, 2]. It is even more complicated with the values of the vibration amplitude of a vibratory drum of a roller, which depends not only on the mass of the vibratory drum, the driving force and the vibration frequency, but also on the parameters of the compacted material, as well as the amount and parameters of shock absorbers of the vibratory drum [3, 4]. Thus, along with the basic parameters of the vibrating roller, the amount and the parameters of shock absorbers have a significant impact on the mode of drum vibrations (continuous contact, partial uplift, and double jump [5–7]) and, therefore, on the compaction efficiency of road construction materials. The amount and the parameters of shock absorbers will affect the level of vibration transmitted onto the frame of the vibratory drum and the whole roller, as well as the vibration parameters at the working space of the driver.

2 Methods The study of the impact of the amount and the parameters of shock absorbers on the vibration parameters of the drum and the drum frame can be carried out using a rheological model proposed in [8]. A feature of this model is that it is now possible to account for the uplift of the vibratory drum from the compacted material during its compaction, which allows to analyze the specifics of interaction of the vibratory drum of a roller with the compacted material in the modes continuous contact (the drum never loses contact with the compacted material during vibrations), partial uplift (the drum loses contact with the compacted material during each vibration cycle, but the amplitude of adjacent uplifts of the drum stays the same), and double jump (the drum makes an uplift from the compacted material, but the amplitude of adjacent uplifts of the drum varies, and that causes increased load spikes on bearings and shock absorbers of the vibratory drum and reduces the efficiency of compaction) [7, 9]. The computational experiment to study the impact the amount and parameters of shock absorbers of a vibratory drum have on the vibration parameters of the drum and the drum frame was conducted during improvement of the vibrating roller DM-614 manufactured by DM Road Construction Machinery Factory, Rybinsk (operating weight 14,000 kg, vibratory drum weight—4000 kg, drum frame weight —4000 kg, exciter vibration frequency—30 Hz), and one of the tasks was to increase the driving force of the vibratory drum from 215 to 280 kN.

Study of Impact of Amount of Shock Absorbers …

767

Based on the study [10], the value of the elastic resistance rate of one shock absorber U150.030 (this type of absorbers is installed in rollers DM-614) was equal kf = 0.362 MN/m. The value of the viscous resistance rate of one shock absorber (bf) was determined according to [11], MN
s/m:  bf ¼ k f
g x

ð1Þ

where η—loss factor (η = 0.16 [11]); x—the angular vibration frequency of the exciter, rad/sec. The computational experiment for the vibrating soil roller DM-614 with 20, 22, 24, and 26 shock absorbers U150.030 mounted on the drum was conducted with different elastic resistance rate values of soil ks. The viscous resistance rate of soil was constant and equal to bs = 0.21 MN s/m2 [11].

3 Results and Discussion Figures 1, 2, 3, 4, 5, 6, 7, 8, 9, 10, 11 and 12 present the calculating results of vibratory drum movement and the movement of vibratory drum frame of the DM-614 roller with 20 and 26 shock absorbers. During the mounting of 20 absorbers, the vibration mode in continuous contact is the operating mode for the primary and the secondary stages of soil compaction (ks = 20 and 58 MN/m) with the vibration swing of vibratory drum in the range of 3.1–4 mm and the vibration swing of the frame of vibratory drum in the range of 0.18–0.22 mm (see Figs. 1 and 7; Tables 1 and 2). When ks = 100 MN/m, the vibrations are in the partial uplift mode with the vibration swing of vibratory drum equals 5.2 mm and the vibration swing of the frame of vibratory drum equals 1.2 mm (see Figs. 2 and 8; Tables 1 and 2). When ks = 145 MN/m, the vibrations are in the undesirable double jump vibration mode with the vibration swing of

Fig. 1 Movement of the vibratory drum of the roller DM-614 (20 shock absorbers U150.030, ks = 58 MN/m)

768

I. S. Tyuremnov et al.

Fig. 2 Movement of the vibratory drum of the roller DM-614 (20 shock absorbers U150.030, ks = 100 MN/m)

Fig. 3 Movement of the vibratory drum of the roller DM-614 (20 shock absorbers U150.030, ks = 145 MN/m)

Fig. 4 Movement of the vibratory drum of the roller DM-614 (26 shock absorbers U150.030, ks = 58 MN/m)

vibratory drum up to 7.2 mm and the vibration swing of the frame of vibratory drum up to 1.75 mm (see Figs. 3 and 9; Tables 1 and 2). Tables 1 and 2 show the calculated values for the vibration swings of vibratory drum and its frame obtained during the installation of 20, 22, 24, and 26 shock absorbers.

Study of Impact of Amount of Shock Absorbers …

769

Fig. 5 Movement of the vibratory drum of the roller DM-614 (26 shock absorbers U150.030, ks = 100 MN/m)

Fig. 6 Movement of the vibratory drum of the roller DM-614 (26 shock absorbers U150.030, ks = 145 MN/m)

Fig. 7 Movement of the frame of the vibratory drum of the roller DM-614 (20 shock absorbers U150.030, ks = 58 MN/m)

Thanks to this experiment, it is possible to establish the amount and parameters of shock absorbers in a vibratory drum to ensure a safe level of vibrations transmitted to the operator’s cabin while considering the use of the vibrating roller during compaction of various road construction materials in a wide range of compaction factors.

770

I. S. Tyuremnov et al.

Fig. 8 Movement of the frame of the vibratory drum of the roller DM-614 (20 shock absorbers U150.030, ks = 100 MN/m)

Fig. 9 Movement of the frame of the vibratory drum of the roller DM-614 (20 shock absorbers U150.030, ks = 145 MN/m)

Fig. 10 Movement of the frame of the vibratory drum of the roller DM-614 (26 shock absorbers U150.030, ks = 58 MN/m)

Study of Impact of Amount of Shock Absorbers …

771

Fig. 11 Movement of the frame of the vibratory drum of the roller DM-614 (26 shock absorbers U150.030, ks = 100 MN/m)

Fig. 12 Movement of the frame of the vibratory drum of the roller DM-614 (26 shock absorbers U150.030, ks = 145 MN/m)

Table 1 Results of the computational experiment to determine the parameters of vibrations of the drum of the roller DM-614 Amount of absorbers

Vibration swing of vibratory drum, mm ks = 58 MN/m ks = 100 MN/m ks = 20 MN/m (Кy = 0.90) (Кy = 0.95) (Кy = 0.97)

ks = 145 MN/m (Кy = 1.00)

20 22 24 26

3.1 3.3 3.2 3.4

7.2 7.4 7.7 7.5

4.0 4.0 4.0 4.1

5.2 5.4 5.5 5.3

Similarly, it is possible to obtain data to study the impact of roller performance on the force values in shock absorbers and bearing of the vibrating module of the roller.

772

I. S. Tyuremnov et al.

Table 2 Results of the computational experiment to determine the parameters of vibrations of the frame of the roller DM-614 Amount of absorbers

Vibration swing of vibratory drum, mm ks = 58 MN/m ks = 100 MN/m ks = 20 MN/m (Кy = 0.90) (Кy = 0.95) (Кy = 0.97)

ks = 145 MN/m (Кy = 1.00)

20 22 24 26

0.18 0.19 0.22 0.23

1.75 4.50 2.50 2.75

0.22 0.24 0.26 0.28

1.22 1.18 1.18 1.65

4 Conclusions The data analysis presented in Tables 1 and 2 shows that the vibration swing of vibratory drum decreases with 24 or 24 absorbers compared to the values corresponding to 22 or 26 absorbers, which show the feasibility of using 20 and 24 absorbers U150.030 during the improvement of the vibratory drum to the driving force of 280 kN, compared to 20 absorbers used in DM-614 with the driving force of 215 kN.

References 1. Tyuremnov IS, Ignatyev AA, Filatov IS (2014) Statistical analysis of performance of vibrating soil rollers. Bull PNU 3(34):81–88 2. Tyuremnov IS, Ignatyev AA (2012) Soil compaction with vibrating rollers: monograph. Publishing House of YSTU, Yaroslavl, p 140 3. Tyuremnov IS, Chabutkin EK, Okulov RD (2008) “Intelligent” rollers—“intelligent” compaction. Constr Road Mach 8:2–7 4. Miheev VV, Savelyev SV (2016) About criteria of efficiency of process of vibration compaction of soil rollers with adaptive working bodies. In the collection: the North of Russia: strategies and prospects of development materials II all-Russian scientific-practical conference, pp 98–105 5. Adam D, Kopf F (2004) Operational devices for compaction optimization and quality control (Continuous compaction control & light falling weight device). In: Proceedings of the international seminar on geotechnics in pavement and railway design and construction, Athens, Greece, pp 97–106 6. Mooney MA, Rinehart RV, Facas NW, Musimbi OM (2010) Intelligent soil compaction systems. Report 676 Available via DIALOG. http://onlinepubs.trb.org/onlinepubs/nchrp/. Accessed 04.12.2014 7. Tyuremnov IS (2016) A review of soil compaction continuous control systems for vibrating rollers. Part 3. The specifics of functioning and smart compaction. Bull PNU 2(41):115–122 8. Morev AS (2017) Study of the compaction value for soil compaction continuous control system for vibrating rollers. Dissertation, Yaroslavl State Technical University 9. Tyuremnov IS, Filatov IS, Morev AS (2012) Rheological model of interaction between the vibratory drum of a roller with compacted soil. Mir Dorog 59(March):86–88

Study of Impact of Amount of Shock Absorbers …

773

10. Tyuremnov IS, Fyodorova DV, Morev AS, Tarasova NE (2017) Experimental determination of numerical values of the elastic shear resistance rate of the shock absorbers U150.030 for vibrating rollers. Mech Stroit 4:42–45 11. Paul J van Susante, Mooney MA (2008) Capturing nonlinear vibratory roller compactor behavior through lumped parameter modeling. J Eng Mech ASCE, pp 684–693

Optimization of High-Power Belt Conveyor Parameters G. G. Kozhushko, O. A. Lukashuk and T. A. Roscheva

Abstract Belt conveyors become more and more widespread in many industrial sectors of the Russian Federation and other countries. They vary by their design, performance, length, spatial configuration of the route, and other technical specifications. Current trends of increasing their performance, belt velocity, and length attest even more to the advantages of the conveyor transport. 40% of the cost of a belt conveyor of 240–480 m long goes to its belt, with longer conveyors having this indicator rise up to 75%. Increasing the service life of such an expensive element of a conveyor is obviously a pressing problem. Among the main factors which impact the service life of a belt are dynamic processes which accompany the operation of belt conveyors in both transitional and steady-state modes. The goal of the research is to study longitudinal vibrations in the conveyor belt using a modified method of Bubnov–Galerkin. The analysis of dynamic processes was carried out with the allowance for the dissipation of vibrational energy in its complex form. A minimal belt tension near the load tensioner was determined on the condition of excluding a loss of stability of the belt along the conveyor route. Keywords Conveyor belt


Tensioner
Dynamic force
Internal friction

1 Introduction A modern belt conveyor is a so-called large mechanical system; that is why it is necessary to carry out a more thorough study of the main physical processes which accompany the operation of the conveyor and its main elements to develop, on that basis, scientifically justified methods of calculation. Existing trends of proposed improvement in designs of belt conveyors are cited in [1–7], while further development of the methods of traction calculation and analysis of starting–breaking modes could be found in [8–11]. The problems of G. G. Kozhushko
O. A. Lukashuk (&)
T. A. Roscheva Ural Federal University, Mira Street 19, Yekaterinburg 620002, Russian Federation e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_83

775

776

G. G. Kozhushko et al.

optimizing the parameters of drives used in belt conveyors on the basis of computer-aided modeling are discussed in [12], while optimizing the power intensity of transportation—in [13]. One perspective route which a researcher can take is the synthesis of dynamic systems—which, in this case, means the design of belt conveyors with their dynamic characteristics is given in advance. A conveyor belt is a system of distributed parameters. When the belt is affected by a longitudinal power momentum, waves of elastic deformation occur in it. The dynamic force in the belt is formed by incident, reflected, and refracted waves. When the drive is switched on, distant segments of the belt begin to move only after a certain period of time needed by a wave to get from the driving drum to the analyzed cross section of the belt. In a conveyor with a flexible tension mechanism, when a power momentum is applied to the belt, a wave of deformations is spreading from the drive along the belt’s working branch, whose tension is increased. In the descending branch, the wave decreases the tension and could be dumped by adjusting the travel of the tension drum. In a conveyor with a rigid tension mechanism, when a power momentum is applied to the belt, a wave of deformations is also spreading—increasing the tension along the working branch and decreasing it along the unloaded branch. The counter running waves superimpose, make a round, and reflect from the drive and boundaries of the branches. The wave process in a conveyor belt which operates in nonstationary modes could lead to undesirable loss of motion stability and change the form of the belt’s cross section [14–16]. Along with the transported material, another important characteristic of the belt would be the level and behavior of external and internal dissipation forces, which affect the form and frequency of vibrations in a working belt [17–20]. Introducing the coefficients of internal friction in the equation of vibration could be achieved using a real or complex form. Since those coefficients are too high for lengthier conveyors, only a direct wave of deformations is considered. Increased coefficients of internal friction lead to speedier waves of deformation.

1.1

Theoretical Part

Let us consider a belt conveyor whose design diagram is shown in Fig. 1. Draw an X-axis along the belt with its reference point x = 0 set on the tension drum. The masses of rotating parts of rollers and the driving drum are reduced to the belt and included into the distributed mass m(x). Regard the rigidity of the belt as depending on a static tension—that is being a function of the coordinate (x). In particular, the rigidity could be constant along the whole length. A differential equation of the belt motion directed along the conveyor with allowance for its internal friction in a complex form [7] is written as

Optimization of High-Power Belt Conveyor Parameters

777

Fig. 1 Design diagram of a conveyor with a load tensioner

ð1 þ icÞ


 @ @uðx; tÞ @ 2 uðx; tÞ Bð xÞ ¼ dðx  l1 Þf ðtÞ;  mð xÞ @x @x @t2

ð1Þ

where uðx; tÞ is a complex function for the elastic displacement of the belt cross section, with (x) as its coordinate and (t) as a moment of time; B(x)—variable pffiffiffiffiffiffiffi rigidity of the belt; ðcÞ—coefficient of internal friction, i ¼ 1; mð xÞ—function of mass distribution along the belt length; x, t—a coordinate on the belt and time; x½0; l; l1 —coordinate of the driving drum; F ðtÞ—peripheral force on the driving drum; l—the whole length of the conveyor belt. Boundary conditions are defined by the following expressions:     

 @u @u I1 @ 2 z @ 2 u ð1 þ icÞ Bð xÞ  Bð xÞ  þ ¼ 2 ; @x x¼0 @x x¼l 2r1 @t2 x¼0 @t2 x¼l    

2   @u @u M2 @ u @ 2 u  ¼ ; ð1 þ icÞ Bð xÞ  þ Bð xÞ  @x x¼0 @x x¼l 2 @t2 x¼0 @t2 x¼l

ð2Þ

where I1 ; r1 are inertia moment and radius of the tension drum; M2 —mass of the tension load. Initial conditions depend on the mode of conveyor operation and could be expressed as Re uðx; 0Þ ¼ lð xÞ; Re

@uðx; 0Þ ¼ #ð xÞ: @t

ð3Þ

To solve Eq. 1 under boundary conditions in Eq. 2 and initial conditions in Eq. 3, one could use a modified Bubnov–Galerkin algorithm, having Eq. 2 rewritten in an operator form, as in mcp l2

d2 L2 u þ ð1 þ icÞBcp L1 u ¼ dðx  l1 Þf ðtÞ; dt2

ð4Þ

778

where

G. G. Kozhushko et al.

  L2 u ¼ l12 qð xÞuðx; tÞ; L1 ¼  ddx k ð xÞ ddx ; kð xÞ ¼ BBðavxÞ ; qð xÞ ¼ mmðavxÞ ; Bav ; mav

are an averaged rigidity of the belt and its mass per unit length. Let us find an approximated solution of Eq. 4 under boundary conditions in Eq. 2 and initial conditions in Eq. 3 in the form of n X

un ðx; tÞ ¼

ðnÞ

ak ðtÞuk ð xÞ;

ð5Þ

k¼0

where fuk ð xÞgðk ¼ 0; 1; 2; . . .Þ is a system of functions within the domain of an ðnÞ operator L1 ; ak ðtÞ—unknown functions needed to be defined. Using the Bubnov–Galerkin method, we get n  P k¼0

  ðnÞ ðnÞ L2 uk ; uj mav l2 a€k ðtÞ þ ð1 þ icÞ L1 uk ; uj Bav ak ðtÞ  ¼  dðx  l1 Þ; uj f ðtÞ; ðj ¼ 0; 1; 2; . . .; nÞ;

ð6Þ

where 

1 L2 uk ; uj ¼ 2 l

Zl qð xÞuk ð xÞuj ð xÞdx; 0



Zl



L1 uk ; uj ¼ 


_ ð x Þ d du k ð xÞ k uj ð xÞdx; dx dx

ð7Þ

0



dðx  l1 Þ; uj ¼

Zl dðx  l1 Þ; uj ð xÞdx ¼ uj ðl1 Þ: 0

Integrals in Eq. 7 would be taken by parts, which, under boundary conditions in Eq. 2, gets us n  X

uk ; uj

k¼0

L2

  ðn Þ ðn Þ mav l2 €ak ðtÞ þ ð1 þ icÞ uk ; uj L Bav ak ðtÞ ¼ uj ðl1 Þf ðtÞ; 2

ð8Þ

where 

uk ; uj

L2

¼

1 l2

Zl qð xÞuk ð xÞuj ð xÞdx þ

 q1 ðu ð0Þ þ uk ðlÞÞ uj ð0Þ þ uj ðlÞ 4l k

0

 q2 ðu ð0Þ  uk ðlÞÞ uj ð0Þ  uj ðlÞ ; 4l k Zl _ ð xÞ duj ð xÞ du dx: ¼ k ð xÞ k dx dx

þ 

uk ; uj

L1

0

ð9Þ

Optimization of High-Power Belt Conveyor Parameters

779

Let us introduce two scalar products for the set of elements in the domain of the operator L1 : 

 uk ; uj ¼ uk ; uj L ;

ð10Þ

1



   uk ; uj ¼ uk ; uj L þ uk ; uj L : 1

ð11Þ

2

Equation 8 is written in a vector form as mav l2 A1 €aðnÞ ðtÞ þ ð1 þ icÞBav A2 aðnÞ ðtÞ ¼ uðl1 Þf ðtÞ; 

   where A1 ¼ uk ; uj ; A2 ¼ uk ; uj  uk ; uj ¼ uk ; uj L ðnÞ

1

ð12Þ

are matrices, and

a ðtÞand uðl1 Þ are column vectors. Solving an inhomogeneous equation of Eq. 12 requires finding a fundamental matrix. To do that, a solution which corresponds to a homogeneous equation of mav l2 A1 €aðnÞ ðtÞ þ ð1 þ icÞBav A2 aðnÞ ðtÞ ¼ 0

ð13Þ

has to be found in the form of aðnÞ ðtÞ ¼ C ðnÞ ekXt ;

ð14Þ

where CðnÞ is a columnvector; X2 ¼ ð1 þ icÞ mBavavl2 ; k a nondimensional parameter which defines the frequency of free vibrations. Substituting the solution in Eq. 14 into the homogeneous equation of Eq. 13 and equating the system determinant to zero, we get a characteristic equation for the values of the parameter k   detA2  k2 A1  ¼ 0:

ð15Þ

Every value of k has a corresponding fundamental solution of C ðnÞ . A value of every p-th frequency of free vibrations could be calculated via this formula: rffiffiffiffiffiffiffiffiffiffiffi Bav : Xp ¼ kp ReX ¼ kp mav l2

ð16Þ

Let H ðtÞ be a fundamental matrix of solutions. Then, a general solution of the inhomogeneous equation of Eq. 12 is written as ðnÞ

a ðtÞ ¼

n X j¼0

ðnÞkXjt Dj

Zt þ

H ðt  sÞuðl1 Þðf ðsÞÞds 0

ð17Þ

780

G. G. Kozhushko et al.

The coefficients Dj are determined on the initial conditions in Eq. 3. Separating real and imaginary parts in Eq. 17, we obtain a solution in a real form:  T Rezn ðx; tÞ ¼ ReaðnÞ ðtÞ uð xÞ

ð18Þ

where (T) denotes an operation of transposition, that is ReaðnÞ ðtÞ is a row vector and uð xÞ ¼ ui ð xÞ is a column vector. Thus, in a vector form, real solutions of Eq. 1 under boundary conditions in Eq. 2 and initial conditions in Eq. 3 are scalar products from Eq. 18. Consider the problem of selecting a minimal tension of the conveyor belt on the tension drum with allowance for dynamic processes in transitional modes of operation. Let the law of distribution of forces in the belt be written as S ðx; tÞ ¼ S0 þ W ð xÞ þ Sg ðx; tÞ;

ð19Þ

where S0 is the belt tension near the tension drum, W(x)—resistance to belt motion over the segment between the driving drum and the cross section (x).  T ðx;tÞ Sg ðx; tÞ ¼ cð xÞ @u@x ¼ cð xÞ ReaðnÞ ðtÞ dudðxxÞ is a dynamic force. Let us find the minimal force S0 on the condition of the total belt tension S ðx; tÞ at no cross section of the belt falling below a given critical level of Scr , whose value could be chosen, for example, with the purpose of maintaining a fluted form of the working branch of the belt [21]. It is logical to search for the function minimum S ðx; tÞ within the segment of ½0; T1 , where T1 is a period of free vibrations, which corresponds to the lowest nonzero frequency. Given that, a necessary condition of the extremum is for partial derivatives of the first order on (x and t) to be equal to zero @Sg ðx;tÞ @S ðx;tÞ @W ðxÞ ¼ @x @x ¼ @x þ @Sg ðx;tÞ @S ðx;tÞ ¼ 0: @t @t ¼

0

ð20Þ

These equations allow us to determine such stationary points which are «suspicious» in terms of the extremum. Then, evaluating the function S ðx; tÞ at those points and choosing among them the minimum S(x0, t0) as a critical tension Scr at which flattening of the belt or loss of the stability could occur, we get an equation of the minimal belt tension near the driving drum: S0 ¼ S2  W ðx0 Þ  Sg ðx0 ; t0 Þ;

ð21Þ

where x0 ; t0 are a coordinate and moment of time at which the studied function takes on the maximum value.

Optimization of High-Power Belt Conveyor Parameters

781

In a general case, finding the x0 ; t0 point is a quite difficult task to accomplish. But as to particular designs of conveyors, while only the belt vibrations in the form which corresponds to the lowest nonzero frequency are taken into account, this task of determining the minimum S0 becomes significantly simpler. Similar to the problem discussed above of the minimum force of preliminary belt tension, one could consider another problem of finding a cross section at which total forces are maximal.

2 Conclusion Practical importance of the dynamic analysis completed lies in the possibility of preventing an emergency related to a loss of stability and characterized by formation of a so-called crease as a result of a slackened tension of the belt in transitional modes of operation below the critical level.

References 1. Galkin VI, Sheshko EE (2017) Lentochnye konveyery na sovremennom etape razvitiya gornoy tekhniki (Belt conveyors at the modern stage of development of mining machines). Mining J 9:85–89 2. Tarasov YD (2013) Povyshenie tyagovogo usiliya lentochnyh konveyerov s uvelichennymi dlinami, uglami naklona i proizvoditelnostiyu (Raising the traction power of belt conveyors with increased lengths, inclinations and performances). J Min Equip Electromechanics 5:46– 48 3. Papoyan RL (2012) Tekhnicheskie usovershenstvovaniya na konveyernom transporte (Technical improvements in conveyor transport). J Min Inf Anal Bull 8:228–233 4. Wang BH, Liu JP, Lu S (2014) Analyzed development of the high-power and high-speed belt conveyor. J Min Mach 42/1:27–30 5. Trufanova IS (2014) Promezhutochnyy privod kak sredstvo sovershenstvovaniya lentochnogo konveyera (Using an intermediate drive as a way of improving a belt conveyor). J Min Equip Electromechanics 6:13–16 6. Roberts AW (2001) Recent developments in belt conveying—bulk solid and conveyor belt interactions. In: Levy A, Kalman H (eds) Handbook of Conveying and Handling of Particulate Solids. Elsevier Science, pp 225–233 7. Yardley ED, Stace LR (2008) Belt conveying of minerals. Woodhead Publishing Ltd., Cambridge 8. Korneev SV, Dolgih VP (2016) Metodika tyagovogo rascheta shahtnyh lentochnyh konveyerov na osnove kompyuternogo modelirovaniya soprotivleniy dvizheniyu tyagovogo organa (Method of traction calculation of belt conveyors used in mines on the basis of computer-aided modeling of tractive resistance). News High Education Min J 3:81–87 9. Dmitriev VG, Cherednik PN (2016) Programmnyy kompleks dlya tyagovogo rascheta i analiza puskovyh i tormoznyh rezhimov lentochnyh konveyerov (Software package for traction calculation and analysis of starting and breaking modes of belt conveyors). J Min Inf Anal Bull 2:25–35

782

G. G. Kozhushko et al.

10. Voytyuk IN, Proskuryakov RM, Kopteva AV (2012) Perspektivy primeneniya sovremennyh informatsionnyh tekhnologiy i tekhnicheskih sredstv izmereniya v sistemah upravleniya lentochnym konveyernym transportom (Prospects of using modern information and measuring technologies in control systems for belt conveyor transport). J Min Inf Anal Bull 6:223–227 11. Nordel LK, Ciozda ZP (1984) Transient belt stresses during starting and stopping: elastic response simulated by finite element methods. J Bulk Solids Handl 4(1):93–98 12. Kuleshov MV, Syromyatnikov VS (2017) Optimization of the conveyor drive parameters under stochastic loads. In: Proceedings of Higher Educational Institutions. Mach Build 10/ 691:69–76 13. Monastyrskiy VF, Monastyrskiy SV, Kiriya RV (2008) Optimizatsiya energoemkosti transportirovaniya nasypnyh gruzov lentochnymi konveyerami (Optimization of power intensity of transporting bulk loads by means of belt conveyors). In: Papers of Seminar 21 of the “Mining Week-2007” Symposium, pp 304–309 14. Podporin TF (2016) Metodika rascheta prodolnoy ustoychivosti dvizheniya lenty v rezhime ustanovivshegosya dvizheniya na bremsbergovom konveyere so slozhnym po prochnostnym harakteristikam konturom lenty (Method of calculating the longitudinal stability of belt used in a brake incline conveyor which functions in the mode of steady motion and utilizes a belt of a complex profile characterized by its strength impact). J Min Equip Electromechanics 6:20– 29 15. Galkin VI, Sazankova ES (2013) Vliyanie parametrov prostranstvennoy trassy lentochnogo konveyera na ustoychivost dvizheniya lenty (Impact which parameters of the space route of a belt conveyor has on the stability of belt motion). J Min Equip Electromechanics 7:6–9 16. Galkin VI, Sheshko EE, Sazankova ES (2015) Vliyanie tipov i harakteristik lent na ekspluatatsionnye parametry spetsialnyh lentochnyh konveyerov (Impact of types and characteristics of belts on operating parameters of specialized belt conveyors). Mining J 8:88– 91 17. Li J, Pang X (2018) Belt conveyor dynamic characteristics and influential factors. J Shock Vib 11:1–13 18. Chen LH, Zhang W, Liu YQ (2007) Modeling of nonlinear oscillations for viscoelastic moving belt using generalized Hamilton’s principle. J Vib Acoust 129(1):128–132 19. Serpil K, Gerdemeli I, Cengiz C (2012) Analysis of belt conveyor using finite element method. In: Scientific proceedings of the IX International congress on machines, technologies, materials 20. Zür TV (1986) Viscoelastic properties of conveyor belts; modeling of vibration phenomena in belt conveyors during starting and stopping. J Bulk Solids Handling 6(3):553–560 21. Kozhushko GG (1992) Mehanika deformirovaniya i prognozirovanie resursa rezinotkanevyh lent konveyerov gornorudnyh predpriyatiy (Mechanics of deformation and prediction on lifespan of rubberized-fabric belts of conveyors at mining enterprises). Dissertation Abstract, Ural State Mining University, Yekaterinburg

Intelligent Control, Correction, and Adaptability of Output Parameters of Vehicles’ Intake System A. V. Gritsenko, I. V. Makarova and G. N. Salimonenko

Abstract Today, the modern motor industry is developing in the direction of significant intellectualization, automation, and robotics at the system operation. It provides for the creation of a comprehensive system of system control algorithms. However, when the technical condition of individual electronic components varies, the control algorithms can significantly change. Thus, correction and adaptation of the operation algorithms to the changing technical condition of the vehicle systems is necessary for a further qualitative object operation. Thus, failures of mass airflow sensors and idle air control valves comprise 14 and 12%, respectively, of the number of failures of all the elements of the microprocessor-based engine control system. A failure of these components leads to a serious change in fuel consumption and exhaust toxicity. We developed an experimental research procedure to elaborate the modes of correction and adaptation to the changing technical condition of these elements of the car electronics. It lies in the formation of stress test actions consisting in the cutout of the nth number of cylinders and individual operating cycles and subsequent loading of the kth number of cylinders. As a result of experimental studies, we built diagrams of the dependencies of the said control parameters on the cylinder load. We analyzed the diagrams, based on which we determined the inflection points corresponding to the prelimit and limit change in the technical condition of the mass airflow sensors and idle air control valves. The developed method of stress test actions allows us to avoid purchasing expensive load banks.













Keywords Idle control Airflow sensor Diagnosis Testing Cylinder cutout Speed Toxicity Fuel efficiency










A. V. Gritsenko (&)
G. N. Salimonenko South Ural State University (NRU), 76, Lenin Avenue, Chelyabinsk 454080, Russia e-mail: [email protected] I. V. Makarova Kazan Federal University, 10a, Syuyumbike Avenue, Naberezhnye Chelny 423800, Russia © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_84

783

784

A. V. Gritsenko et al.

1 Introduction One of the main mechanisms of cars, which give 14.2% of engine failures, is the mass airflow sensor (MAFS). Meanwhile, the process of MFFS diagnostics in actual operating conditions has low reliability, and the information obtained in this case does not allow us to determine the required technological effects to maintain its operability and, consequently, to control its condition. The control of the technical condition of the ICE systems is basically reduced to the fact of the presence of a malfunction or failure, which ultimately ends with the correction of the ICE operating parameters. The actual determination of a failure epicenter is often difficult, and therefore, currently, despite the availability of onboard self-diagnostic systems and a wide variety of diagnostic equipment, there arises an acute problem of developing effective ICE diagnostic means and methods. The lack of efficient diagnostic means and methods results in errors in assessing the technical condition of mobile machines, which is a significant problem leading to a decrease in the labor productivity, increased environmental pollution, and consumption of operating materials [1–5]. One of the ways to solve this problem is a rational organization and management of ICE operation processes based on the complete information on the targeted object. At the same time, all the missing information on the technical condition of the ICE systems can be obtained as a result of technical diagnostics, which plays a significant role in improving the management of the technical condition of the ICE systems and maintenance and repair processes, reducing the costs of their implementation and improving the operational safety of vehicles [6–9]. The first part of the hypothesis lies in an unambiguous reliable assessment of the technical condition of the mass airflow sensor (MAFS) by forming stress test interactions consisting in the cutout of three cylinders and loading one cylinder remaining in operation with the friction horsepower obtained in the cylinder cutout mode. The load formation process passes stepwise by controlling the following parameters: position angle of the throttle valve, ICE crankshaft (CSh) speed, mass airflow, fuel consumption, and multimeter readings. The second part of the hypothesis lies in an unambiguous reliable assessment of the idle air control valve (IACV) by forming stress test actions consisting in the cutout of two cylinders and loading of one of the remaining ones with the friction horsepower obtained in the cylinder cutout mode. The load formation process passes stepwise by controlling the following parameters: position of the IACV, fuel consumption, airflow through the IACV, adaptive fueling, and ignition advance angle. The purpose of the research is to increase the efficiency of diagnosing the vehicles’ MAFS and IACV. The research tasks include: • analyze failures of the intake system elements, analyze the existing diagnostic methods and means; • theoretically justify the diagnostic parameters and modes for assessing the technical condition of MAFS and IACV;

Intelligent Control, Correction, and Adaptability …

785

• develop a method and means for diagnosing the technical condition of MAFS and IACV; and • carry out operational and production check of the method and means for MAFS and IACV diagnostics. The object of the research is the MAFS and IACV process of vehicles with a microprocessor-based engine control system. The subject of the research is the interrelation of the technical condition of the MAFS with the parameters of the mass airflow and fuel when forming a test loading by a full and partial cutout of cylinders. Check of the speed of the IACV corrective actions and the degree of adaptability to the stress test.

2 Theoretical Research The studies have revealed serious shortcomings in the receipt and analysis of diagnostic information on the MAFS when using various diagnostic instruments and complexes at service stations. A comprehensive approach to the design and development of the MAFS diagnostic instrument made it possible to create a mobile highly effective diagnostic means. The imperfection of diagnostic methods and means entails a malfunction and even a complete failure of mechanisms and systems that require significant resources for restoration. A comparative analysis of the competitive device models has shown the necessity of creating a device meeting higher requirements of ergonomics and accuracy [10–14]. The intake system is represented by a row of consecutive and parallel elements. Some of them practically do not change their properties and output parameters during operation, while others constantly change their technical condition. Thus, for example, the air filter is one of the most dynamically changing elements, which changes the parameters of the entire system as far as it is polluted. A test action is proposed to determine its technical condition, which consists in bringing of the ICE to the operation mode on a single cylinder with a fully open throttle valve with or without the air filter. The larger the degree of the air filter clogging, the lower the CSh speed of the ICE with the filter in relation to the speed without the air filter. It is supposed to use this feature to determine the technical condition of the air filter [15–17]. Thus, following and substituting the values of the resistance coefficients, we obtained the dependence of the change in the amount of the supplied air on the resistance of the air filter, showing high sensitivity of the change in the air input to the increase in resistance. Thus, the increase in the resistance from 0.05 to 52.6 resulted in the air input change from 620 to 205 kg/h. Such a change in the filter resistance causes serious failures in the ICE operation and a power decrease. A further growth of the filter resistance from 52.6 to 760 causes a decrease in the amount of air from 205 to 57 kg/h, which results in the ICE shutdown and inoperability.

786

A. V. Gritsenko et al.

One of the most unreliable elements of the intake system is the MAFS. Its failure is manifested in the incorrect presentation of data on the amount of air by the microprocessor control system. There often appear failures that give a significant deviation in the air supply. When the reference MAFS is used, which is installed in sequential order after the regular one, it is possible to determine head losses at the cylinder inlet and the amount of air passing through it. The introduction of a reference MAFS into the intake path will be reflected by a slight increase in resistance. It is calculated that, at the maximum possible discharge in the inlet manifold, the addition of a reference MAFS will be reflected in the 0.16% decrease in the air supply, which is very insignificant and is within the limits of the MAFS error value [14, 18, 19]. Thus, it is established that the head losses at the cylinder inlet, when a reference MAFS is installed, are insignificant and cause a decrease in the amount of supplied air within Q = 621…620 kg/h. That is, at the maximum possible discharge in the intake manifold, the addition of a reference MAFS will be reflected in the 0.16% decrease in the air supply, which is very insignificant and is within the limits of the MAFS error value. In case of a successive installation of the diagnosed and reference MAFS, the head losses on them are almost zero, and the application of the claimed diagnostic tools does not result in any measurement error. The IACV becomes unserviceable quite often. The degree of rod extension (the number of steps) is observed, and then, the speed of the IACV corrective actions and the degree of adaptability to the stress test are checked to detect its failure during the ICE start-up [18, 20]. To check the speed of the IACV corrective actions, it is proposed to cut out two cylinders simultaneously and to check the response time before adding the number of steps.

3 Methods We developed a general methodology to reduce the fuel consumption of gasoline internal combustion engines by a full and partial cylinder cutout to solve these problems [21]. In the experiment, two cylinders were cut out, since when three cylinders were cut out, it was impossible to maintain a stable engine operation at the idle speed [21, 22]. Two cylinders remained in the working position, one of which was artificially loaded using an algorithm of the developed program and the implementation on the patented electromagnetic injector circuit breaker. When the IACV was tested, the CSh speed was kept at the same level and comprised 880 min−1. The test object is the electronically controlled ZMZ-406.10 engine. To check the reliability of the obtained data, we had to know the technical condition of the tested engine, the wear parameters of the cylinders of which were measured with K-69 M air tester.

Intelligent Control, Correction, and Adaptability …

787

4 Experimental Research Experimental research was carried out on ZMZ 406.10 engine, only one cylinder of the four remained in operation. For example, in case of a 4-cylinder engine, the first cylinder remains in operation, and the second, third, and fourth ones are completely cut out. Under these conditions, the position of the throttle valve is changed with a 5% step, and the following parameters are measured: ICE CSh speed (min−1); mass airflow (k/h); fuel consumption (l/h); and multimeter reading (V) (Table 1). To create the necessary ICE loading modes, we used our method and the electromagnetic injector circuit breaker allowing us to create any load on the diagnosed cylinder [21]. The reliability of the obtained results was checked by comparing the experimental diagnostic data with the analogous ones on other ICEs. During the research, it has been found that when the cooling radiator fan was turned on, the experimental data changed. Later, when the cooling radiator fan was turned on, the experimental data were not recorded. Based on the experimental data and the known parameters of the ICE technical condition, we built graphic dependencies, according to which it is clearly possible to judge on the technical condition of the cylinder-piston group. Figure 1a shows the dependence of the change in the ICE CSh speed on the degree of the partial cylinder load. An analysis shows that for the second operating cylinder, the diagram of the CSh speed has a linear form, and when the cylinder is loaded, the speed does not exceed 4100 min−1. For the fourth operating cylinder, the CSh speed when the cylinder is loaded does not exceed 3700 min−1, and the diagram itself is linear. For the third cylinder, when the load on the cylinder reaches 1/6 of the share, there is a sharp drop in the CSh speed. For the first cylinder, when the cylinder load reaches 1/5 of the share, a sharp decrease in the CSh speed is observed. For the first and third cylinders, as it can be seen in the diagram, it is impossible to load the cylinder by over 1/4 of the share, since in this case the engine cannot operate. The largest value of D is observed when the cylinders are loaded by 1/4 of the share and is 730 min−1. Considering that the wear condition of the fourth cylinder on the tested engine is 14%, of the second—22%, of the first—32%, and of the third—29%, the experimental dependences are completely accurate. Figure 1b shows the dependence of the change in mass airflow (kg/h) on the degree of the partial cylinder load. An analysis shows that for the second operating cylinder, the diagram of the mass airflow has a linear form, and when the cylinder is loaded, the airflow does not exceed 280 kg/h. For the fourth operating cylinder, the mass airflow when the cylinder is loaded does not exceed 260 kg/h, and the diagram itself is linear. For the third cylinder, when the load on the cylinder reaches 1/7 of the share, there is a sharp drop in the airflow. For the first cylinder, the decrease in the airflow from 189.5 to 226.1 kg/h has a linear form. The largest value of D is observed for the second and third cylinders when the cylinders are loaded by 1/4 of the share, making the value of 104.5 kg/h.

ICE CSh speed (min ) Mass airflow (kg/h) Fuel consumption (l/h) Voltmeter readings (V)

−1

Position of the throttle valve (%)

840 43.9 4 2.83

5

Table 1 Experimental research parameters

1400 72.5 6.5 3.12

10 2080 113.2 9.4 3.39

15 2520 147.1 13 3.61

20 3040 180.2 16.2 3.79

25 3360 215.9 19.5 3.95

30 3680 239.4 23.9 4.04

35 3840 252.2 25.8 4.09

40 3920 274.6 28 4.16

45 3960 258.6 29.6 4.2

50

4000 299 30.4 4.27

55

… … … …

…

4080 314.9 31.2 4.18

90

4120 313 31.6 4.18

95

4120 292.2 32 4.24

100

788 A. V. Gritsenko et al.

Intelligent Control, Correction, and Adaptability …

789

Fig. 1 a Diagram of the dependence of the change in ICE CSh speed (min−1) on the degree of the partial cylinder load (share) of alternatively switched cylinders. b Diagram of the dependence of the change in the mass airflow (kg/h) on the degree of the partial cylinder load (share) of alternatively switched cylinders

Fig. 2 a Diagram of the dependence of the change in fuel consumption (l/h) on the degree of the partial cylinder load (share) of alternatively switched cylinders. b Diagram of the dependence of the compliance of the IACV position on the degree of the cylinder load (share) of alternatively switched cylinders

Figure 2a shows the dependence of the change in fuel consumption (l/h) on the degree of the partial cylinder load. An analysis shows that for the second operating cylinder, the diagram of the fuel consumption has a linear form and when the cylinder is loaded, it does not exceed 30.5 l/h. For the fourth operating cylinder, the fuel consumption when the cylinder is loaded does not exceed 27.2 l/h, and the diagram itself is linear. For the third

790

A. V. Gritsenko et al.

cylinder, when the load on the cylinder reaches 1/6 of the share, there is a sharp drop in fuel consumption. For the first cylinder, when the load on the cylinder reaches 1/5 of the share, there is a sharp decrease in the fuel consumption. For the first and third cylinders, as it can be seen in the diagram, the cylinder cannot be loaded by over 1/4 of the share, since in this case the engine cannot operate. The largest value of D is observed when the cylinders are loaded by 1/4 of the share and is 13.8 l/h Figure 2b shows a diagram of the dependence of the compliance of the IACV position (the number of steps) on the degree of the cylinder load (share) of alternatively switched cylinders. The IACV position when the second and fourth cylinders are in operation is lower than when the first and third cylinders are in operation. It is explained by a better technical condition of the second and fourth cylinders as compared to the other ones. The largest value of D is reached when the cylinders are loaded by 1/4 of the share and between the second and third cylinders and is eight steps. The inflection point of the diagrams is also the place of loading by 1/4 of the share, before which the value of D is five steps with a constant magnitude. The IACV position equal to 149 steps is reached on the fourth cylinder with the load of 1/2 of the share, which is characterized by its technical condition. Figure 3a shows a diagram of the dependence of the change in the airflow through the IACV (kg/h) on the degree of the cylinder load when the cylinders are partially loaded with work cycles. To ensure a preset ICE operation under loads, due to the IACV operation, the air supply to the combustion chamber is increased. To ensure a given load, the first and third cylinders need more air as compared to the second and fourth cylinders, which are also explained by their technical condition. The inflection point is the moment of loading by 1/4 of share. The largest value of D is 4 kg/h and is reached at the moment when the cylinders are loaded by 1/5 of the share. The airflow of 33 kg/h is achieved at the moment of loading by 1/2 of the share by the first and third cylinders due to their worst technical condition. At the same time, the airflow on the second and fourth cylinders is 31 and 30 kg/h, respectively. Figure 3b shows a diagram of the dependence of the change in the idling fuel consumption on the degree of the cylinder load when the cylinders are partially loaded with work cycles. Due to the increased air supply to the first and third cylinders, the electronic control unit of the ICE controls the fuel-and-air mixture by increasing the fuel supply, which results in higher fuel consumption on the more exhausted cylinders. At the moment of loading by 1/3 of the share, the value of D is 0.4 l/h between the first and fourth cylinders. The fuel consumption of 2.8 l/h is reached at the moment of the first cylinder loading by 1/2 of the share and the fully loaded third cylinder, which is characterized by their technical condition.

Intelligent Control, Correction, and Adaptability …

791

Fig. 3 a Diagram of the dependence of the change in airflow through the IACV (kg/h) on the degree of the cylinder load (share) of alternatively switched cylinders. b Diagram of the dependence of the change in airflow through the IACV (kg/h) on the degree of the idling fuel consumption (l/h) on the degree of the cylinder load when the cylinders are partially loaded with work cycles

5 Conclusions The analysis of the obtained dependencies allows us to judge on the technical condition of the cylinder-piston group by the indirect parameters, which allows to carry out diagnostics without large labor costs and without direct access to the motor engine, namely through a standard diagnostic connector. When operating vehicles in the test mode, the control of inflection points corresponding to the prelimit and limit change in the technical condition of the MAFS and IACV allows us to monitor the level of the change in the technical condition, to correct and adapt the sensors in any intermediate moment. We developed the method of stress test actions, which allows us to avoid purchasing expensive load banks. The creation and implementation of the gasoline engine loader built into the engine’s electronic control unit make it possible to obtain an economic effect of about 3.9 thousand rubles when diagnosing per one car a year. Acknowledgements The work was supported by Act 211 Government of the Russian Federation, contract № 02.A03.21.0011.

792

A. V. Gritsenko et al.

References 1. Khan MI, Yasmeen T, Muhammad M (2016) Research progress in the development of natural gas as fuel for road vehicles. Renew Sustain Energy Rev 66:702–741 2. Berdov EI, Fedoseev SYu (2015) Cutout of some engine cylinders as a way to improve the fuel and economic indicators of the tractor-transport unit. Agric Ind Complex Russ 72(2): 20–24 3. Karavalakis G, Short D, Russell RL et al (2014) Assessing the impacts of ethanol and isobutanol on gaseous and particulate emissions from flexible fuel vehicles. Environ Sci Technol 48(23):14016–14024 4. Gurgenci H, Aminossadati SM (2009) Investigating the use of methane as diesel fuel in off-road haul road truck operations. J Energy Resour Technol 131(3) 5. Stein RA, Anderson JE, Wallington TJ (2013) An overview of the effects of ethanol-gasoline blends on SI engine performance, fuel efficiency, and emissions. SAE Int J Engines 6(1):470– 487 6. Gonçalves M, Jiménez-Guerrero P, Baldasano JM (2009) Emissions variation in urban areas resulting from the introduction of natural gas vehicles: application to Barcelona and Madrid greater areas (Spain). Sci Total Environ 407(10):3269–3281 7. Milkins EE, Allen RG, Edsell VD (1990) Gaseous fuel injection system for the operation of heavy duty engines on natural gas. In: Society of automotive engineers, eighteenth Fisita congress—The promise of new technology in the automotive industry, Torino, Italy, 7 May 1990 8. Pasechnik DV (2004) Gas fuel supply system for injection engines ZMZ. J Automot Ind 5:12–15 9. Gaidar SM, Svechnikov VN, Usmanov AYu et al (2013) Improving the engine performance characteristics based on the use of nanotechnologies. Proc GISNITI 111:4–8 10. Eck C, Konigorski U, Cianflone F et al (2011) Fault detection system for the air path of common rail diesel engines with low pressure EGR. SAE Technical Papers, World Congress and Exhibition, Detroit, MI; United States; 12 April 2011 11. Kimmich F, Isermann R (2002) Model based fault detection for the injection, combustion and engine-transmission. IFAC Proc 15(1):203–208 12. Hajari SC (1996) Diagnosis and repair of excessively emitting vehicles. J Air Waste Manag Assoc 46(10):940–952 13. Gumus M, Ugurlu A (2011) Application of phase change materials to pre-heating of evaporator and pressure regulator of a gaseous sequential injection system. Appl Energy 88(12):4803–4810 14. Grebennikov AS (2002) Diagnosis of car-and-tractor engines by the intra-cycle changes in the crankshaft angular speed. Dissertation, University of Saratov 15. Ivanov RV (2010) ICE diagnosis by the parameter of the mechanical loss power. Dissertation, University of Volgograd 16. Leshakov IA, Kravchenko IN, Erofeev MN (2013) Mathematical model for calculating the basic parameters of catalyst converters. Mod Sci-Intensiv Technol 5:76–80 17. Isaenko PV, Isaenko VD (2006) To the calculation of the hydraulic resistance of the neutralizer-purifier for exhaust systems of vehicles of the motor transport complex. Bul. of Tomsk State University of Architecture and Constr. 1:95–106 18. Khimchenko AV, Mishin DG, Buzov AV (2013) Reduction of the unevenness of the engine’s torque with the cylinder cutout in partial loading modes. Intern Combust Engines 1:46–51 19. Isaenko VD, Isaenko AV, Isaenko PV (2007) Fundamentals of the reliability theory and car diagnostics. Teaching guide. Federal Agency for Education, State Educational Institution of Professional Education Tomsk State University of Architecture and Construction, Tomsk

Intelligent Control, Correction, and Adaptability …

793

20. Gritsenko AV, Kukov SS, Tsyganov KA et al (2011) A method for diagnosing the exhaust path of the reciprocating internal combustion engine. RUG01 M 15/04, 2011139288, 26 Sept 2011 21. Gritsenko AV, Plaksin AM, Shepelev VD (2017) Studuing lubrication system of turbocompressor rotor with integrated electronic control. Procedia Eng 206:611–616 22. Patrakhaltsev NN, Strashnov SV, Kornev BA et al (2011) Diesel control by the method of switching on-off of cylinders or cycles. Engine Building 3:7–12

Duration of Ignition Delay of Fuel–Air Mixture in Diesel Engines V. F. Guskov and A. N. Gots

Abstract Empirical and semi-empirical dependences for determination of ignition delay time of the fuel–air mixture si in piston engines with compression ignition were analyzed. While the mathematical simulation of the piston engine cycle, it must be determined the ignition delay time of mixture. It allows not only to forecast the start process of combustion and determine the location of the area of rapid pressure increase relative to the top dead center, but also to improve the angle of start of fuel injection by energy, economical, or environmental indicators. This task is solved at any complexity of the mathematical model of the cycle and is the main one when we were compared calculated and test indicator diagrams of the piston engine cycle. The interdependences obtained by various researchers to determine the ignition delay time of the fuel–air mixture in diesel engines were analyzed. Estimates suggest that a number of formulas give values si unreal for automobile and tractor engines. Analysis of indicator diagrams of tractor diesel 3CH10.5/12 allowed to rank the determining factors (pressure and temperature of the working medium, speed mode, delivery ratio of air) by the degree of their influence on the ignition delay time of the fuel–air mixture. On this basis, dependence and recommendations can be given for calculating the duration of the ignition delay time of the fuel–air mixture.




Keywords Ignition delay time Fuel–air mixture Temperature Working medium





Diesel
Pressure


V. F. Guskov
A. N. Gots (&) Vladimir State University Named After Alexander and Nikolay Stoletovs (VlSU), 87, Gorky Street, Vladimir 600000, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_85

795

796

V. F. Guskov and A. N. Gots

1 Introduction While the mathematical simulation of the piston engine cycle, it is important to determine the ignition delay time si of the fuel–air mixture, which allows to forecast the start process of combustion, the duration of which is determined by the indicator diagram in coordinates p-u. In this case, the ignition delay time of the fuel–air mixture is determined either by the angle of rotation of the crankshaft u, or by the turnaround time si of the crankshaft, from fuel injection point for the engine with ignition from compression to the point of deviation of the line of pressure during combustion from the compression pressure curve. The region of the indicator diagram is shown in Fig. 1 (on the compression stroke and the expansion stroke near the top dead center (TDC)). At point F while rotation of the crankshaft by an angle uf-. Then the start of fuel injection. The point D is taken as the ignition start due to compression when the crankshaft is rotated by an angle ud (see Fig. 1)—(the compression line is displayed as a dotted line). The point D is taken as the start process of combustion, about the angle of rotation of the crankshaft ud, i.e., is determined as the first recorded time when the pressure on the indicator diagram exceeds the compression pressure [1–3]. Point C on the indicator diagram of operation process determines the cylinder pressure when the piston is in TDC and point Z is the maximum cycle pressure. Injection timing angle of fuel for engines with compression ignition relative to TDC Fig. 1 Diagram of the cycle in coordinates p-u (rotation angle of crankshaft)

Duration of Ignition Delay of Fuel–Air Mixture in Diesel Engines

797

is /af ¼ /c  /f . Angles /i ¼ /d  /f and /ad ¼ /c  /d have an influence on the rise angle of pressure. Angles /i and /ad have an effect on the combustion process. While angle /ac is high then combustion of the mixture starts very early and it created a high back pressure in the piston stroke to the TDC (DC line on the Fig. 1) and greatest pressure is too high and increases with big speed, then the engine is overloaded with gas forces and overheated, and power of diesel engine does not reach optimal values [4, 5]. In the mathematical model of the piston engine cycle, the angle of rotation of the crankshaft fi or the ignition delay time is calculated, and the orientation of the points D and Z relative to the TDC is achieved by adjusting the position of the point F—the beginning of fuel injection. In experimental research, the optimal on economic and energy marks of indicated values of angles are provided with a choice of the advance injection angle, certain value si is determined by on the indicator diagram. Observed value si is individual for this type of piston engine, working conditions, and other factors. It is suitable in mathematical modeling to create one of the most suitable empirical dependence to determine si ignition delay time relative from main factors: pressure and temperature of the working medium at the point of injection, speed rate, activation energy, etc. Strong correlation of these factors was held by A. Tolstov [6–8]. Therefore, when you made forecast the start of the process combustion in the synthesis of the indicator diagram it is important to determine si ignition delay of the fuel–air mixture. The duration of the ignition delay time of the fuel–air mixture is si = 0.4…1.5 ms or 7…12% by the total time of the combustion [1–3]. The duration of the ignition delay and the dynamic load determines the rate of pressure rise in the rapid combustion phase and the position of the maximum pressure in the cylinder relative to the top dead center (TDC). The complexity of taking into account the influence of various factors on the duration of the ignition delay time of the fuel–air mixture gave rise to a wide variety of empirical or semi-empirical formulas for calculating the time ti, which should have a minimum deviation from the experimental and allow to forecast the start of the combustion process. It is especially important at the stage of engine development. The variety complicates their choice for a particular mathematical model. A lot of researches are devoted to the solution of this problem, among which the most complete and qualitative analysis of methods for determining the ignition delay time given in the book by Kavtaragze [9]. Researches devoted to self-ignition of heterogeneous combustible mixtures suggest the solution of two problems: • to determine the duration of the ignition delay time of the mixture si [1, 10–13]; • to study the physicochemical processes carried in the working medium for si [4–9].

798

V. F. Guskov and A. N. Gots

The solution of the first problem allows to determine the optimal location with help the indicator diagram of the rapid combustion area relative to the TDC while calculation of the cycle. The solution of the second problem is to set the connections to the characteristics of heat input and fuel supply. The classical model of self-ignition of a drop of fuel is based on the steady-state approximation by Semenov–Frank-Kamenetsky [5, 10]. Analysis of the development of such models, as well as the model of combustion of liquid hydrocarbon drops with the calculation of si values for the homogeneous stoichiometric mixture of h-decane-air at different temperatures and pressures, is given in the source [11, 12]. In the mathematical model of the piston engine cycle, the duration si is calculated, and the orientation of the points of the takeoff pressure line from the compression line on the indicator diagram piston engine cycle, and the maximum pressure relative to the TDC is achieved by adjusting the fuel injection angle.

2 The Purpose of the Study To perform analysis of the published results on the calculation of the ignition delay time of the fuel–air mixture in diesel engines and give recommendations on determining of its duration. While the calculation of the cycle it allows to determine the optimal by indicated work location of the rapid pressure area relative to TDC (line D–C-Z, see Fig. 1). It is solved with any complexity of the mathematical model of the cycle and is the main one when comparing the calculated and indicator diagrams of the piston engine cycle.

3 Research Methods The calculation-theoretical method is used for the research of ranking the efficiency of the main factors on the duration of the ignition delay time of the fuel–air mixture in diesel engines and study recommendations for its calculation. According to statistics, on the operating condition (full-load operation) of automobile and tractor engines angle /af ¼ 12. . .35 (uf = 325…348° crank angle when it is rotating after the admission stroke) (see Fig. 1) and for high-speed diesel engines (ne = 2000… 0.4000 min−1) should not exceed si  0.0025 s, and for spark-ignition engine (ne  4000 min−1) (/ad ¼ 4. . .7 , see Fig. 1). It is sensible to find a universal dependence for si, by determining the degree of influence of determining factors on the ignition delay time of the mixture.

Duration of Ignition Delay of Fuel–Air Mixture in Diesel Engines

799

This will allow the development of the diesel engine at an early stage to choose the parameters of the working process, which will provide the specified effective and environmental performance.

4 The Results of the Study and Their Discussion The most complete and qualitative analysis of the calculation formulas si is given in book Kavtaradze [9]. Most of the formulas match to the dependence, proposed by Semenov [5, 14–16]: ~T si ¼ B pn exp E = R





ð1Þ

where B = const—some constant value; p, T—pressure and temperature at the point ~ of fuel injection; E—activation energy of the working medium; and R—universal gas constant. The conversion of time values of the ignition delay time of the fuel–air mixture si match to the correspondence given in book [9] to angle units /i ¼ 6 ne si (ui— angle of turn crank) proves that a number of formulas give unreal values si (/i ) for automotive and tractor engines (for example, in the formula of A. I. Tolstov). Only this can establish the reason for such a large spread of values si. Using formula (1) or si ¼ B
pn
T m

ð2Þ

find out the degree of influence determining factors on the ignition delay time of the fuel–air mixture according to the formulas from the book [9], which structure is close to (1). In Table 1, dependences arrange in order to increase of the magnitude of an ignition delay time si of the fuel–air mixture. Formulas of Sitkei [17], H. Hardenberg [18], and Neugebauer [19] were not taken into account, as they contain an algebraic addition of the components as well as the formulas of V. S. Semenov, A. I. Tolstov, and others, where speed rate and other factors are taken into account. It is noteworthy that there is quite a difference in indicators degrees (see Table 1) for pressure pn —from 0.35 (G. Vashni) to 2 (L Spadacini). It is unusual for the duration of physicochemical processes during the ignition delay time of the fuel–air mixture. The exponent degree at T or the number a in the common factor exp(a/T) in the above correspondences do not change significantly. Note that the calculation by the formulas given in Table 1 of the delay time of ignition of the mixture si showed that the difference in the value of these numbers reaches 22. Let us find average values of the degrees of influences pf and Tf on the duration of the ignition delay time of the fuel–air mixture.

800

V. F. Guskov and A. N. Gots

Table 1 Influence of various factors on the ignition delay of the mixture in diesel 3CH10.5/12 T, К

si, ms

/i CA

dB dp dT Degree of influence on si , %

No.

pf , MPa

1

si ¼ 12:7=ne ð40=C Þ0:69 p0:386 expð4644=T Þ; B. Knight. C = 52—cetane number. 1.5550 713 2.406 28 36.9 28.9 34.2

2

si ¼ 1:3 p0:35 expð990= T Þ; G. Vashni, F. Anisits 1.5550 713 3.543 42 24.2

59.3

16.5

3

si ¼ 4646
104 p0:7 T 2:22 ; H. Åberg. 1.5554 709 1.011 12

41.8

23.6

34.6

4

si ¼ 1:76 p0:866 expð2490= T Þ; R. Z. Kavtaradze 1.5552 705 0.262 3 3.4

75.1

21.5

5

si ¼ 0:55 p1:3 expð4400= T Þ; R. Z. Kavtaradze 1.5553 708 0.976 11

3.1

64.8

32.1

6

si ¼ 0:0405 p0:757 expð5473=T Þ; F. Stringer 1.5552 712 1.814 21

14.8

49.8

35.4

7

si ¼ 78:07 p1:66 expð2055=T Þ; M. Tougue. 1.5552 704 0.159 1

18.7

68.8

12.5

si ¼ 0:44 p1:19 expð4650=T Þ; X. Wolter (p, mbar) 1.5551 713 3.067 36 4.4 Medium value 1.5552 710 1.655 20 18.4

61.0 53.9

34.6 27.7

8

For this purpose, the calculation program of a cycle in which mathematical model was used and is stated in articles [6, 7]. It synthesizes the indicator diagram for the given values of the average piston speed, index of excess air ratio, and other factors. Using the calculation cycle of the diesel 3CH 10.5/12 (D-130) at the rotary speed n = 2000 min−1, determine the effect of factors B, p, and T on si. The structure of formulas of two types was used: by (1) or (2). If Eq. (1) or (2) finds the absolute values of the logarithms of the factors 

 ~T jlg si j ¼ lg B j þ lgðpn Þ j þ lg exp E = R or jlg si j ¼ jlg B j þ jlgðpn Þ j þ jlgðT m Þ j; this allows us to calculate the degree of impact d of each factor B, p, and T on the value si:

Duration of Ignition Delay of Fuel–Air Mixture in Diesel Engines




lg exp E= R ~T jlg B j jlgðpn Þj jlgðT m Þj dB ¼ ; dp ¼ ; dT ¼ or dT ¼ jlg si j jlg si j jlg si j jlg si j

801

ð3Þ

Formulas and results of calculations on Eq. (3) are given in Table 1. Values of n and m in p−n and T−m calculated with the average values dB ; dp ; dT (Table 1) and for time average siav = 0.001655 s. The exponent m by dependence (3) for dT with an average value of T = 710 K is equal m  −0.27. Analogically, calculated the exponent n at pressure p, Pa from correspondence (3): n  −0.24. Cofactor B in formula (2) according to the analyses given in source [9, 20] should be taken into account: • makeup of the working medium—value excess air factor k because when k increase then it should increase ignition delay time of fuel–air mixture; • turbulation of the working medium by means of average piston speed wp , its increase reduces the ignition delay time of the fuel–air mixture. To take into account the impact of other factors, we introduce another cofactor Bi, and value B is presented in the form (4) B ¼ Bi k wk p :

ð4Þ

The indicator k will take Bi ¼ 1:0 and calculate cycle diesel 3Ч10,5/12 the value k = 1.5 and wp ¼ 8 m/s. Based on the formulas (3) and (4) with an average value of dBmed = 0.184, we will receive k  0:754. On the bases of our research, we offer the following formula for calculation of ignition delay time of the mixture si: n m si ¼ Bi k wk p pf Tf ;

ð5Þ

where Bi ¼ 1:0—factor, which, when adjusted by experimental data /i , may have a value other than one; p; T—pressure, Pa and temperature, K; and wp—the average speed of the piston, m/s The values of the exponents n, m, k were taken as follows [21–25] k ¼ 0:754; n ¼ 0:242; m ¼ 0:270:

5 Conclusion Obtained results allow: • To rank the influence degree of determining factors on the duration of the ignition delay time of the mixture: pressure pf —dp  50%; temperature Tf —dT  25%; average piston speed wp —dw  20%; and other factors 5%;

802

V. F. Guskov and A. N. Gots

• Set approximate ranges of possible values when correlating the formula for determining the ignition delay time of the mixture in the calculation of piston engine cycles: dp ¼ 0:540. . .0:535; dT ¼ 0:275. . .0:280; dwp ¼ 0:18. . .0:19:

ð6Þ

It is recommended to take the exponents in formula (5) in the ranges: n ¼ 0:236. . .0:242; m ¼ 0:267. . .0:270; k ¼ 0:754

ð7Þ

Refinement values of n and m in formula (5) can be performed in the presence of an experimental indicator chart or data from the engine prototype: • To determine the ignition delay time si of the fuel–air mixture when the angle ui is measured on the indicator diagram and the time is calculated si ¼ /i =ð6 ne Þ; • Pressure pf, PA and temperature Tf, K at point F are determined (see Fig. 1); • The values of the degrees of influence on the time  si are taken from (7);  • Indicators n and m are calculated n ¼ ðdp lg si Þ lg pf ;m ¼ ðdT lg si Þ lg T f ; • Calculation of the cycle is performed and the coincidence of the calculated and experimental values of the ignition angle of the mixture is estimated. Checking the obtained dependence to determine the ignition delay angle of the working mixture on the engines produced in production showed good convergence of the calculated and experimental values. This will improve the process of debugging experimental diesel.

References 1. Varbanets RA, Ivanovsky VG (2009) Model of working process in problems of increase of efficiency of operation ship diesel power installations. Materials of the III Ukrainian scientific and technical conference “Modern problems of a engine building: a condition, ideas, decisions.” Pervomaisk, pp 6–10 2. Gavrilov AA, Gots AN (2013). Influence of external factors on pressure of a working body in a cycle piston engine Materials of the V Ukrainian scientific and technical conference “Modern problems of a engine building: a condition, ideas, decisions.” Pervomaisk, pp 52–59 3. Gavrilov AA, Gots AN (2013.) Model of a turbo-supercharging in a cycle of the engine with variable pressure of air upon an admission. Basic Res 8(part 1):24–28 4. Lebedev ON, Somov VA, Kalashnikov SA (1990) Engine of combustion of river courts. Transport, Moscow, p 320 5. Semenov NN (1934) The chain reactions. ONTI, Leningrad, p 110 6. Slavutskij VM, Salykin EA, Slavutskij VV (2012) Analys of the factors determining a delay of ignition of fuel in diesel engines. News of VolgSTU. no. 2 (99), (Processes of transformation of energy and power installations) VolgSTU. Volgograd, Release 4:51–54 7. Portnov DA (1963) A high-speed turbomondial engines with ignition from compression. Mashgiz, Moscow, p 640

Duration of Ignition Delay of Fuel–Air Mixture in Diesel Engines

803

8. Tolstov AI (1955) Display the period of delay of ignition and dynamics of a cycle of the high-speed engine with ignition from compression. Works NILD Research of working process and submission of fuel in high-speed diesel engines 1:5–55 9. Кavtaradze RZ (2008.) Theory of piston engines. Special the chapter: the Textbook for high schools. Publishing house of MSTU name N.E.Bauman, Moscow, 720 p 10. Frank-Kamenetsky DA (1967) Diffusion and a heat transfer in chemical kinetic. M.: Publishing house AN of the USSR, 1987, Nauka, Moscow 11. Frolov SM, Basevich VJ, Frolov FS, Borisov AA, Smetanjuk VA, Avdeev KA, Gots AN (2009) Correlation between drop vaporization and self-ignition. Russ J Phys Chem 3:333–347 12. Gavrilov AA, Gots AN (2014) Duration of the ignition delay of the fuel–air mixture in piston engine. J Fundam Res 6–4:703–708 13. Skripnik AA, Frolov SM, Kavtaradze RZ, Efros VV (2004) Simulation of ignition in a jet of liquid fuel. J Fundam Res 23(1):54–61 14. Frolov SM, Scripnik AA, Kavtaradze RZ (2003) Modeling of diesel spray ignition. Semenov memorial. combustion and atmospheric pollution. Torus Press Ltd., Moscow, pp 220–227 15. Kavtaradze ZZ, Zelinger Zitzler GK (2005) The ignition Delay in a diesel engine using different fuels. Russian Academy of Sciences. Thermophys High Temp 43(6):947–965 16. Voinov AN (1977) Combustion in high-speed diesel engines. Machines Making, Moscow, p 346 17. Sitkei G (1964) Kraftstoffaufbereitung und Verbrennung bei Dieselmotoren. Springer-Verlag, Berlin, p 224 18. Harbenberg H, Wagner W (1976) Der Zunderverzug in direktteinspritzenden Dieselmotoren. MTZ 7:240–248 19. Neugebauer S Das instationare BBetriebsverhalten von Ottomotoren—exherimantelle Erfassung bnd rechnerische Simulation. Dissertation, TU Braunschweig 20. Svirodov YB (1972) Mixture formation and combustion in diesel engines. Engineering, Leningrad, p 345 21. Fujito H et al (1980) Illumination delay in diesel spray. JASE Paper, no. 800–13, 149 p 22. Boehman AL, Le Corre O (2008) Combustion on syngas in internal combustion engines. Combust Sci Technol 6(180):1193–1206 23. Heywood JB (1988) Internal combustion engine fundamentals. McGrawHill, New York, pp 345–349 24. Lata DB, Misra A (2011) Analysis of ignition delay period of dual Diesel engine with hydrogen and LPG as secondary fuels. Int J Hydrog Energy 36:3746–3756 25. Alekseev VP (1974) Equation for the determination of the ignition delay period in the combustion chamber of a diesel engine. News of Universities. Engineering, pp 106–109

Modeling of Maximum Cycle Pressure Based on Engine External Speed Performance A. N. Gots and V. S. Klevtsov

Abstract When developing a draft design of a piston engine, it is necessary to carry out a calibration calculation of its main parts. To do this, it is necessary to know how the loads on parts change when the indicators vary according to the external speed characteristic. A technique is proposed for simulating the maximum pressure of a piston engine cycle using an external velocity characteristic at the design stage, when it is impossible to obtain experimental data. For this, dimensionless coordinates are proposed for describing the performance of the engine being designed. The transition to dimensionless indicators allows the use of these prototype engines. The proposed dependences allow one to model an external velocity characteristic, by which one can trace the change in the main parameters of a piston engine. The laws of changes in the average effective and indicator pressures on the external velocity characteristic are analyzed and the law of change in the maximum cycle pressure is proposed based on the change in the dimensionless coordinates. Keywords Piston engine performance


Maximum cycle pressure
Modeling
External speed

1 Introduction When calculating the loads on the elements of the details in piston engines, one has to know the variation of the maximum cycle pressure (combustion pressure) pmax with the engine speed n. Theoretical dependencies of pmax on the parameters of irreversible thermodynamic cycles are given by the following relationships [1]: Spark-ignition engines under assumption of constant-volume combustion

A. N. Gots
V. S. Klevtsov (&) Vladimir State University Named After Alexander Grigorjevich and Nikolay Grigorjevich Stoletovs (VlSU), 87, Gorky Street, Vladimir 600005, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_86

805

806

A. N. Gots and V. S. Klevtsov

pmax c ¼

h iVs n

1 n2 1




120Pi ðe  1Þ 
i ; 1  en211  kðn111Þ 1  en111

ðpmax r ¼ 0:85pmax c Þ

When applying the above equations for calculating the thermodynamic cycle of a specific engine, the following parameters are kept constant: compression ratioe, number of cylinders i, cylinder volume Vs, and engine speed n. Other parameters: rated power Pi, combustion pressure ratiok, preliminary q and subsequent d expansion ratios [2], mean polytropic exponents at compression n1 and expansion n2 strokes—are varied as they depend on the engine operation mode. Exponents n1 and n2 depend mainly on the charge temperature and intensity of heat transfer with the cylinder walls [3–6]. At a fixed position of the fuel supply actuator, the effect of engine speed n on the characteristic temperatures of compression and expansion strokes is relatively small and, therefore, the assumption on the constant mean values of polytropic exponents n1 and n2 for all speed ranges is expected to result in insignificant errors in determining pmax. The most complicated task is to find the dependence of parameters k, q, and d on the engine speed as their effect on the mean indicator pressure pmi and power Pi is the most significant. Indicator parameters pi and pmi may be determined based on the mechanical efficiency ηm or the power of mechanical losses pm as Pi ¼ Pe =gm pmi ¼ pme =gm

or or

Pi ¼ Pe þ Pm : pmi ¼ pme þ pm ;

ð1Þ

where pmi, pme, and pm are the mean indicator, effective, and mechanical loss pressures, respectively, which vary with the engine speed n. The values of parameters ηm and pme for the engine being designed can be chosen according to statistical data available for other engines already under production or taken equal to those for the prototype engine in a nominal operating mode. Their variation with the engine speed n is dependent essentially of engine design, type, and destination. For determining mechanical losses, various empirical correlations have been proposed [7], e.g.: pm ¼ a þ bvm ;

ð2Þ

where a and b are the empirical coefficients depending on engine design, type, and destination; vm ¼ ðsnÞ=30 is the mean piston speed; s is the piston stroke; xi (i = 1, 2, 3, 4) are the exponents depending on the engine type. At variable engine speed n, parameter pmi is mainly affected by the engine load and fuel burning conditions. These latter parameters are determined by the variation of: supercharging or reverse outburst which increases with decreasing n at fixed phases of gas distribution; the value of hydrodynamic drag at the inlet port which grows with n;

Modeling of Maximum Cycle Pressure Based on Engine …

807

the ignition crank angle which increases with decreasing n, in particular, at fixed ignition or fuel-injection angles, resulting in elevated cylinder pressure at the compression stroke and therefore in diminishing pmi; other parameters. As a result, with increasing n along with the external speed performance, mean pressure pmi first increases and then decreases. Increase in engine power with n is thus possible only if the relative drop in pmi is less than relative increase in n. In this case, the drop in the mean indicator pressure is compensated for by the increase in the engine speed [8–10]. Analysis of indicator diagrams shows that variations of pmi and maximum combustion pressure pmax with engine speed n are rather close to each other, whereas variations of effective mean pressure pme and engine torque Ttq are affected by mechanical losses, and therefore, the maxima of pme and pmi as functions of engine speed do not in general coincide. Engine speed npmi max , at which pressure pmi attains maximum, is always higher than speed npme max , at which pme is maximal.

2 The Purpose of the Study Develop a mathematical model based on the external speed characteristics to determine the pressure of the gasoline engine cycle.

3 Research Methods Take as an example a gasoline engine for VAZ vehicle with a cylinder bore of d = 0.079 m; piston stroke of s = 0.08 m; number of cylinders i = 4; and single cylinder volume of Vs = 0.3945 l. Table 1 shows the variation of relevant engine parameters along with the external speed performance (ESP). Let us demonstrate that the dependence of the maximum combustion pressure versus engine speed pmax x ¼ f1 ðnÞ may be obtained on the basis of the engine torque curve Ttq x ¼ f2 ðnÞ or the mean effective pressure curve pme x ¼ f2 ðnÞ. Here, index x labels the current value of effective engine torque Ttq and mean pressure pme, as well as the maximum cycle pressure pmax. To obtain pmi and pm in Table 1, the correlations of Eqs. (1) and (2) were used, respectively, with a = 0.049 and b = 0.0152. Note that effective engine torque Ttq max (and mean effective pressure pme) attains its maximum value at nTtq max ¼ 3200 min−1, whereas maximum pressure and mean indicator pressure attain their maxima at npmax = 3500 min−1 and npimax = 4000 min−1, respectively (maximum values are shown in bold in Table 1). It is hardly possible to determine the value of npmax theoretically because of the reasons mentioned above. As a rule, the values of pmax for a gasoline engine are determined

808

A. N. Gots and V. S. Klevtsov

Table 1 Parameters of ESP for VAZ vehicle engine No.

n (min−1)

Ttq x (N m)

Pex (кW)

pme x (MPa)

pm (MPa)

pmi (MPa)

pmax x (MPa)

pmax c (MPa)

1 1 2 3 4 5 6 7 8 9 10 11 12

2 1500 2000 2500 3000 3200 3500 3700 3900 4000 4500 5000 5400

3 98.5 108.1 113.6 115.9 116.1 115.6 115 114.1 113.6 110.6 107.3 105

4 15.47 22.64 29.74 36.41 38.90 42.37 44.55 46.60 47.58 52.15 56.18 59.37

5 0.7894 0.8663 0.9104 0.9288 0.9304 0.9264 0.9216 0.9144 0.9104 0.8864 0.9599 0.8415

6 0.1098 0.1301 0.1503 0.1706 0.1787 0.1909 0.1990 0.2071 0.2111 0.2314 0.2517 0.2679

7 0.8991 0.9963 1.0607 1.0993 1.1091 1.1172 1.1205 1.1214 1.1215 1.1177 1.1115 1.1093

8 5.18 5.65 6.02 6.21 6.23 6.24 6.23 6.2 6.19 6.08 5.93 5.85

9 5.51 5.86 6.17 6.25 6.26 6.24 6.22 6.19 6.17 6.06 5.93 5.85

after the calculation of engine cycle in two operating modes—in the mode with maximum torque and in the nominal mode. For a diesel, these values are specified based on the considerations of engine material strength. Despite parameters pme, pmi and pmax are of the same dimension, the resultant dependencies pmax x ¼ f1 ðnÞ, pme x ¼ f2 ðnÞ and pmi x ¼ f3 ðnÞ are difficult to compare with each other even after their approximation since the coefficients entering the corresponding mathematical models are different [11, 12]. Therefore, it is instructive to transform these parameters to the dimensionless variables and compare their behavior in terms of dimensionless coordinates nn and ηp. The origin of the coordinates will be set to the points of maxima at curves pmax x ¼ f1 ðnÞ, pme x ¼ f2 ðnÞ and pmi x ¼ f3 ðnÞ. For this purpose, the following correlations will be used: For the engine speed, nx (X-axis) [13–15]: nn ¼ ðnx  npsum Þ=ðnnom  npsum Þ;

ð3Þ

for pressures pmax x , pme x , and pix (Y-axes): gp ¼ ðpx  psum Þ=ðpnom  psum Þ;

ð4Þ

where nx is the current value of the engine speed; npsum is the frequency at which pressures pmax x , pme x , and pmi x attain their maximum values; nnom is the nominal engine speed; px is the current value of maximum, mean effective, and indicator pressure; psum are the maximum values of these pressures attained along the ESP: pmax esc , pme max , and pmi max ; pnom are their values in the nominal operating mode (pmax nom , pme nom , and pmi nom ).

Modeling of Maximum Cycle Pressure Based on Engine …

809

Table 2 Dimensionless parameters for VAZ gasoline engine No.

nnpme

gpme

nnpmi

gpmi

nnpmax

gpmax

1 2 3 4 5 6 7 8 9 10 11 12

−0.7727 −0.5455 −0.3182 −0.0909 0 0.1364 0.2273 0.3182 0.3636 0.5909 0.8182 1

−1.5856 −0.7207 −0.2252 −0.0180 0 −0.0450 −0.0991 −0.1802 0.2252 −0.4955 −0.7928 −1

−1.7857 −1.4286 −1.0714 −0.7143 −0.5714 −0.3571 −0.2143 −0.0714 0 0.3571 0.7143 1

−18.2680 −10.2820 −4.9960 −1.8160 −1.0182 0.3483 0.0773 0.0038 0 −0.3101 −0.8178 −1

−1.0526 −0.7895 −0.5263 −0.2632 −0.1579 0 0.1053 0.2105 0.2632 0.5263 0.7895 1

−2.7189 −1.5128 −0.5641 −0.0769 −0.0256 0 −0.0256 −0.1026 −0.1282 −0.4103 −0.7949 −1

After applying Eqs. (3) and (4) for calculating the parameters of ESP for the VAZ [16–18] gasoline engine (Table 1), their dimensionless analogs are obtained which are presented in Table 2. As the calculation of these dimensionless parameters was based on the engine speeds at which pmax x , pme x , and pmi n attain their maximum values, the origin of coordinates for each of the parameters is different (zeros in Table 2). Figure 1 shows the plots of some cycle parameters in dimensionless coordinates obtained by the use of the data in Table 1. Curve 1 corresponds to the mean effective pressure gpme ¼ f1 ðnnpme Þ; curve 2 to   the mean indicator pressure gpmi ¼ f2 nnpm ; and curve 3 to the maximum comi   bustion pressure gpmax ¼ f3 nnpmax [19]. n n

n n

x pmi max x pme max Here, in accordance with Eqs. (3) and (4), nnpme ¼ nnom ; nnpmi ¼ nnom ; np np

n n

me max

mi max

pme x pme max pmi x pmi max pmax x pmax esc x pmax esc nnpmax ¼ nnom npmax esc ; gpme ¼ pme max pme nom ; gpmi ¼ pmi max pmi nom ; gpmax ¼ pmax esc pmax nom . It follows from Fig. 1 that coordinates, the curves for mean  in dimensionless 

effective pressure gpme ¼ f1 nnpme and maximum combustion pressure gpmax ¼   f3 nnpmax eventually coincide, whereas the curve for mean indicator pressure   gpmi ¼ f2 nnpmi differs considerably from the other two curves, in particular, in its initial part. If one assumes with some reserve that the largest value of the maximum combustion pressure pmax esc for the gasoline engine is attained in the operating mode with maximum engine torque at nTtq max , then the actual and experimental values differ by only 1%.   The dependence gpme ¼ f1 nnpme

polynomial [20]:

can be approximated by the third-order

810

A. N. Gots and V. S. Klevtsov

Fig. 1 Plots of ESP parameters in dimensionless coordinates: 1—mean effective pressure  gpme ¼ f1 nnpme ; 2—mean indicatorpressure  gpmi ¼ f2 nnpmi ; and 3— maximum combustion   pressure gpmax ¼ f3 nnpmax

gpme ¼ 0:9693n3n  1:9356n2n  0:0288nn  0:0045:

ð5Þ

Let us assume that as a result of engine cycle calculation, the value of pmax nom = 5.85 MPa has been obtained for the nominal operating mode. If one further assumes that the maximum combustion pressure varies similar to Ttq x ¼ f2 ðnÞ [or pme x ¼ f2 ðnÞ], then the value of pmax Ttq max relevant to the operating mode with maximum engine torque can be readily obtained from the rated torque margin:
 lTtq ¼ Ttq max  Ttq nom =Ttq nom ;

ð6Þ

where Ttq max is the effective maximum torque and Ttq nom is the effective torque in the nominal operating mode. Similar to Eq. (6), one can determine the maximum combustion pressure pmax, assuming that it attains the largest value in the mode of maximum power:
 lpmax ¼ pmax Ttq max  pmax nom =pmax nom ;

ð7Þ

where pmax Ttq max and pmax nom are the maximum combustion pressures in the operating mode with maximum engine torque and nominal mode, respectively. Besides other parameters, the effective engine torque at the ESP is affected considerably by mechanical losses. Therefore, as these losses do not affect the maximum combustion pressure, one obtains lpmax \lTtq . Assume that lpmax ¼ k1 lTtq ;

ð8Þ

where k1 ¼ pm Ttq max =pm nom ; pm Ttq max and pm nom are the mean pressures of mechanical losses in the operating mode with maximum engine torque and in nominal mode, respectively.

Modeling of Maximum Cycle Pressure Based on Engine …

811

4 The Results of the Study and Discussions From Table 1 and Eq. (7), one obtains: lTtq ¼ ð116:1  105Þ=105 ¼ 0:1057;

k1 ¼ 0:1787=0:2679 ¼ 0:6670

Assuming now that the dependence of pmax x on the engine speed is the same as that of Ttqx (or pmex), and pmax = 5.85 MPa in the mode of nominal power, one comes to the estimate: lpmax ¼

pmax Ttq max  5:85 ¼ 0:1057
0:6670 ¼ 0:0705: 5:85

Thus, in the mode of maximum engine torque, the maximum combustion pressure will be equal to pmax Ttq max = 6.26 MPa. Comparison of this value with the experimental result presented in Table 1 gives for the estimation error of pmax Ttq max the value of ((6.26–6.24)/6.24)
100% = 0.3%.

5 Inferences If one assumes that curve gpmax ¼ f2 ðnn Þ for VAZ engine is given in dimensionless coordinates by Eq. (5), then the use of Eq. (4) yields: pmax x  6:26 ¼ 0:9693n3n  1:9356n2n  0:0288nn  0:0045 6:26  5:85 Column 9 in Table 1 presents the calculated values of pmax c. Analysis of these values indicates that at the initial part of ESP, the relevant error is only of about 5%.

References 1. Vyrubov DN, Ivaschenko DN, Ivin VI et al (1983) Internal combustion engines. Theory of piston and hybrid engines, 4th edn. Mashinostroenie, Moscow, 372p 2. Dyachenko NKh, Dashkov SN, Musatov VS et al (1962) High-speed piston internal combustion engines. Mashgiz, Moscow, 359p 3. Vyrubov DN, Ivaschenko NA, Efimov SI et al (1985) Internal combustion engines: systems of piston and combined engines. Mashinostroenie, Moscow 4. Kolchin AI, Demidov VP (2010) Calculation of automobile and tractor engines: a manual for universities. Higher School, Moscow 496p 5. Denisov AS, Baskov VN (1985) Assessment of the degree of influence of operating factors on the operating modes of a diesel engine. Dvigatelestroyeniye 11:39–41 6. Vzorov BA, Adamovich AV et al (1981) Tractor diesel engines: a handbook. Mashinostroenie, Moscow 535p

812

A. N. Gots and V. S. Klevtsov

7. Diachenko NK, Dashkov SN, Musatov VS et al (1962) High-speed piston internal combustion engines. Mashgiz, Moscow 360p 8. Portnov DA (1963) High-speed turbo piston engines with compression ignition. Mashgiz, Moscow 640p 9. Lukanin VN et al (2005) Internal combustion engines. Book 1. Theory of working processes: textbook for universities. High school, Moscow, 479p 10. Lenin IM (1969) Theory of automotive and tractor engines. Mashinostroenie, Moscow 367p 11. Gots AN, Gavrilov AA (2003) Modeling of engine cycle parameters based on the external speed performance. In: Fundamental and applied problems of improving piston engines. Proceedings of the 9th international conference. Vladimir State University, Vladimir, p 52–57 12. Gots AN, Kudelya IN (1998) Modeling of external speed performance of Diesel with a regular branch. J. Tractor Agricu Mach 9:21–22 13. Gotz AN (1997) Model of fatigue durability of engine parts. Engine-97: Material international scientific—technical conference. MSTU. N.E. Bauman, Moscow, p 82 14. Kavtaradze RZ (2008) Theory of piston engines. Special chapters: textbook for universities. Publishing House of Moscow State Technical University. N.E. Bauman, Moscow, 720p 15. Marchuk GI (1973) Methods of computational mathematics. Science, Novosibirsk 312p 16. Balyuk BK, Azhippo NA (1961) A new method for testing engine friction bearings for chipping fatigue. Dvigatelestroyeniye 9:60–61 17. Azhippo NA, Balyuk BK (1985) Prediction of the durability of sliding bearings of tractor engines at the design stage. Dvigatelestroyeniye 8:17–20 18. Wiechert MM, Dobrogayev RP et al (1964) The design and calculation of automotive engines. Mashinostroenie, Moscow 552p 19. Kavtaradze RZ (2011) Thermophysical processes in diesel engines covered with natural gas and hydrogen. Moscow State Technical University Publishing House. N.E. Bauman, Moscow 238p 20. Gots AN, Efros VV (2007) The design of automotive and tractor engines: a study guide. VlSU, Vladimir 148p

Effectiveness of Road Transport Technology in Modern Housing Systems A. V. Kulikov and S. Y. Firsova

Abstract In the present paper, it is recommended to use a systematic approach to organizing road transport in the housing industry. A proposal was made to consider the transportation process in the construction industry at three levels: micro-level, meso-level, and macro-level. The operations, which make up the transportation process of building materials, should be considered as a process flow schemes of transportation. The paper defines the total costs of circulation of logistical support in housing construction. The paper presents the performance evaluation of road transport when transporting construction cargoes. The transport will be considered at the macro-level when analyzing the organization of road transport in the construction industry of residential facilities within the framework of state housing programs (city, region, and country). To reduce transport costs, it is planned to create specialized trucking companies, which would work under these programs and would be the unstructured element of the system. The boundary of the system in this case will be a city, state, and country.







Keywords Micro-level Meso-level Macro-level Transportation costs Economic effect





Transportation technology


1 Introduction At present, the improvement of housing conditions of the population is an important element of social policy influencing the demographic and socioeconomic development of society. Housing construction is one of the important sectors of construction in Russia [1]. Housing construction in Russia falls into four categories depending on the materials used for the construction: wood, bearing-wall, monolithic, and brick. A. V. Kulikov
S. Y. Firsova (&) Federal State Budget Educational Institution “Volgograd State Technical University”, 28, Lenin Avenue, Volgograd 400005, Russia e-mail: fi[email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_87

813

814

A. V. Kulikov and S. Y. Firsova

Also, there are mixed types of house building, which combine the above categories in one way or another (frame and monolith-brick). Housing construction is one of the most resource-demanding spheres of production. Certain volumes of transport and cargo-handling works are performed during the construction of any building or structure. These works are connected with transportation from the place of production to the construction site of materials, half-finished products, and finished products [1]. Currently, building cargoes include a variety of materials, constructions, components, technological equipment, as well as the cargoes, which come out of construction works (soil, construction waste, etc.). Based on the organization of loading and unloading, building cargoes are divided into the following groups: piece cargoes; small-pieces, break-bulk cargoes; bulk materials; cementing material [2]. The largest share falls on soil (35%). It is connected with large amounts of earthworks in the construction. Significant shares fall on inert materials (22%), concrete (20%), and reinforced concrete products (14%), as they are the main building materials. Transportation costs are sometimes higher than the cost of extraction or production in the cost of some building materials. About 80% of all transportations of construction cargoes are carried out by road transport. The advantages of road transport are high speed, high maneuverability, and the ability to deliver a variety of goods directly to the object of construction. This type of transport is the most widely used in conditions of housing construction. The main task of the organization of transportation of building cargoes is the timely delivery of materials and designs to the construction site with minimum cost of transportation. The paper uses a systematic approach to the organization of transportation of building cargoes. The operations, which make up the transportation process, are presented in the form of process scheme of building cargoes transportation using one mode of transport or several modes of transport. The cost structure of the circulation system of logistical support in housing construction is considered. The paper presents ways to reduce the transport component in the construction of residential facility.

2 A Systematic Approach to Organizing Road Transport The main task of road transport is timely, quality, and full satisfaction of customer. The organization of transport is the organization of activities related to the preparation for loading, loading, transportation to destination, and unloading of various cargoes. The task of organization of transport is the effective linkage between stages of the transportation process. The basic concept of the systematic approach to organization as a process is the interconnection of parts or subsystems of enterprise. Such approach envisages a goal setting and focusing on the creation of a whole, in contrast to the creation of the components, steps or subsystems. Organizational systems are designed to

Effectiveness of Road Transport Technology …

815

achieve simultaneous work of separate but interrelated parts, which provide higher overall effectiveness than the total effectiveness of the parts taken separately [3, 4]. Organization of transportation is connected with the establishment of the procedure of preparation and transportation of cargoes with proper settlement system, accounting, and control. It should focus on achieving high performance of the vehicle with minimum cost of transportation [2]. To obtain a more objective description of such a system using minimum information, the transportation process should be considered at three levels: the micro-level, meso-level, and macro-level. The level of the system depends on what variables describe the system. It is necessary to determine the system boundaries and the structure of unstructured element of the system. The unstructured element of the system is its smallest part, mode of behavior of which is subjected to the laws of all structural level of the system. When determining the unstructured element, it is necessary to consider that there is a minimum level of dimensions, below which the properties of the system are no longer appearing. To select a part of the system, which could be unstructured element, it should be borne in mind that elements of the system interact with each other. The transport will be considered at the macro-level when analyzing the organization of road transport in the construction industry of residential facilities within the framework of state housing programs (city, region, and country). To reduce transport costs, it is planned to create specialized trucking companies, which would work under these programs and would be the unstructured element of the system. The boundary of the system in this case will be a city, state, and country. At the micro-level, the work of transport is considered during the construction of one residential facility, and the unstructured element of the system will be cargo-car-driver-road. The boundaries of micro-level of the system are set by route of transportation of construction cargoes from the supplier to the construction facility. At the meso-level, the unstructured element will be routes of cargo delivery from the supplier to the various construction sites of one builder. Boundaries of the system are determined by territory under construction. It is necessary to reduce transport costs at all three levels. At the micro-level, costs can be reduced by selecting the optimal type of vehicle and cargo-handling mechanisms; preparation and rational placing of cargo on the pallet, platform or in a truck body [5–7]; selection of optimal scheme of transportation; improvement of performance of the truck; coordination of the work of the objects of production, consumption, transport. At the meso-level, transport costs can be reduced through the effective use of vehicle and creation of routes for transportation of small parties of cargoes and cargo shipping using one truck. At the macro-level, transport costs can be reduced through the coordinated work of the transport complex of town (district, region) within the social housing construction programs [8]. At the micro-level, the output parameters of the system include the following elements: cargo description; volume of transportations (year, month, day); classification rating; value of cargo; point of departure and destination; carrying distance; time of departure and delivery of cargo; type, model and the number of units of vehicle [9]; technical and operational parameters of vehicle (capacity, capacity

816

A. V. Kulikov and S. Y. Firsova

utilization rate, loaded mileage proportion, technical speed, downtime during loading and unloading); carrying capacity of vehicle [10]; irregularity coefficient of transportations (by months, days of the month, and hours of the day); prime cost of transportation; transportation costs; share of transportation costs in the costs of cargo; reserve of the carrying capacity of vehicle; utilization rate of carrying capacity. It is not necessary to consider so many parameters for many purposes, it is sufficient to consider the transportation process at the meso-level. The list of output parameters of freight traffic at the meso-level includes the following elements: volume of transportation; value of cargo; prime cost of transportation; transportation costs; share of prime cost of transportation in the costs of cargo; organization level of transportation; effectiveness ratio of the transportation process; carrying capacity of vehicle; reserve of the carrying capacity; and share of transportation costs in the gross domestic product of enterprises. At the macro-level, the output parameters of freight traffic include the following elements: volume of transportation; prime cost of transportation; transportation costs; reserve of the carrying capacity; and share of transportation costs in the gross domestic product of enterprises. It is important to determine performance indicators for each level of the system, which most fully describe its state at a particular time. Many details about the work of transportation systems are lost when moving from micro-level to meso-level and macro-level. For example, descriptions and characteristics of transported cargoes, characteristic of vehicle, etc. It is necessary to focus attention on parameters such as the organization level of transportation, the share of transportation costs in the gross domestic product of enterprises, and determination of rational reserve of carrying capacity of transport means.

3 Estimation of the Transportation Cost The transportation of building cargoes starts at the point of production and finishes at the point of consumption. The transportation process starts with the process of cargo preparation for transportation (storage, packaging, labeling, etc.). The accumulation process (for example, at the factory or construction warehouse) is necessary to get the right amount of cargo forwarded to one consumer. The process of loading and delivery from the supplier to the construction site using road transport follows previous stage. Transportation of building cargoes is a complex process of sequential, interconnected, and cross-operations, regulating all steps to transport materials from their production site to place of consumption. The method of delivery is defined as transport and technological scheme, which is a set of sequential process steps and complex of technical means, ensuring the fulfillment of these operations and established procedure of transportation of cargo using optimal scheme [9].

Effectiveness of Road Transport Technology …

817

The rational shipping method is considered to be one, which is the most effective in accordance with the optimality criterion. The process operations, which make up the process of transportation, are heterogeneous and differ greatly in their duration. Some operations form certain stages of the transport process, when they are combined, all of which performs a specific task. The process schemes of transportation of building cargoes are shown in the articles [11]. The transportation volumes of building cargoes in housing construction are based on cost estimates for the construction of residential buildings. Planning of traffic volumes of building cargoes is carried out in technological order in accordance with the schedule of construction of a particular object. The role of road transport is to provide transportation of necessary construction materials within fixed time limits, in proper amounts and to the “right” price. The transportation costs are an important component, which forms the cost of construction. Improving transportation technology of building materials will allow reducing transportation costs. Development of technological schemes of transportation of building materials provides a coherent work of transport and cargo-handling facilities, thereby reducing production losses [12–14]. The economic indicators are important elements. They reflect economic interests of the construction sector, and they should be systematically upgraded. Proper accounting of the costs associated with the shipping process is important not only for the transport industry, but primarily for industries and businesses served by transport. When determining the costs related to implementation of the transportation process, it is necessary to take into account: technical and economic parameters of the vehicle; transportation distance; costs associated with loading and unloading, damage and loss of cargo, delay in delivery of cargo. The total costs of circulation of the system of logistic support in construction are defined by the following formula [12, 15]: X Cls ¼ Cproc þ Cmtr þ Ctr þ Ctrack þ Cstor þ Cscar ð1Þ where Cproc —order placement and order processing costs; Cmtr —cost of material and technical resources; Ctr —transport costs; Ctrac —costs of material resources tracking; Cstor —storage costs; Cscar —costs associated with resource scarcity. The economic effect DEtr1 from the optimization of the volume of a single shipment is defined by the following formula [2, 12, 15]. The economic effect from changes in regularities of distribution of service time of vehicle at the point of loading is determined by the formula [12, 15]: DEmp2 ¼ ðt1  t10 Þ
Ca
Ax ;

ð2Þ

where t1 —waiting period at the point of loading of vehicle with an exponential distribution of service time, h; t10 —waiting period at the point of loading of vehicle

818

A. V. Kulikov and S. Y. Firsova

with Erlangian or regular distribution of service time, h; Ca —prime cost of vehicle usage, rub/h; Ax —number of cars, u. The economic effect from optimization of the number of vehicles running with loading mechanism is defined by the following formula [12, 15]: DEtr3 ¼ Etl  Etlmin ;

ð3Þ

where DEtr3 —effect from optimization of the number of vehicles running with loading mechanism, rub/h; Etl —losses connected with downtime of loading mechanisms and vehicle due to non-uniformity of their work, rub/h; Etlmin —minimum losses related to downtime of loading mechanisms and vehicle due to non-uniformity of their work, rub/h. The economic effect of the choice of optimal type of vehicle and cargo-handling mechanisms is defined by the following formula [12, 15]:    opt   Cl
Mxopt þ Cvopt
Aopt ; DEtr4 ¼ Clcur
Mxcur þ Cvcur
Acur x x

ð4Þ

where DEtr4 —economic effect of the choice of optimal type of vehicle and cargo-handling mechanisms, rub/h; Clcur —prime cost of using the current loading mechanism, rub/h; Cvcur —prime cost of using the current vehicle, rub/h; Clopt — prime cost of using the selected loading mechanism, rub/h; Cvopt —prime cost of cur using the selected vehicles, rub/h; Acur x —number of used cars, u; Mx —number of opt used loading mechanisms, u; Ax —optimal number of vehicles, u; Mxopt —optimal number of loading mechanisms, u. The expected economic effect of transportation with the highest values of technical speed can be calculated by the formula [12, 15, 16]: DEmp5 ¼

j  X nri
lrc
Ca i¼1

VT1 i



 nri
lrc
Ca ; VT2i

ð5Þ

where DEtr5 —economic effect of transportation with the highest values of technical speed, rub; j—recommended number of time periods per day, which provides speed increase, u; nri —number of haulages over time period, u; lrc —length of haulages with a cargo, km; VT1i —technical speed in the current time interval of shipments, km/h; VT2i —technical speed in the proposed time interval of shipments, km/h. A rationalized method of delivery in relation to specific customers, transport network, possible forms of transport service, and types of construction cargoes reduces the transport costs in the construction based on the feasibility studies in compliance with system, structural, technological, and organizational requirements. According to our evaluations, the transport costs are reduced from 25 to 15% of the total cost of housebuilding [14, 17–19].

Effectiveness of Road Transport Technology …

819

4 Conclusion A proposal was made to consider the transportation process in the construction industry at three levels: micro-level, meso-level, and macro-level. The paper considers a systematic approach to transport of building materials. The operations, which make up the transportation process of building materials, should be considered as a process flow schemes of transportation. These schemes include the following stages: preparation of cargo for transportation; loading; transportation; transfer of cargo from one mode of transport to another; unloading; and storage of cargo at the construction site [20]. The paper considers the structure of distribution costs of logistical support in housing construction. The transportation costs in the construction of a residential facility can be reduced through: optimization of the volume of one shipment; changing the regularities of distribution of vehicle service time at the loading point; optimizing the number of vehicles operating with one loading mechanism; choosing the optimal type of vehicle and loading and unloading mechanisms; and organization of transportation with the highest values of technical speed. At the meso-level, transport costs can be reduced through the effective use of vehicle and creation of routes for transportation of small parties of cargoes and cargo shipping using one truck. At the macro-level, transport costs can be reduced through the coordinated work of the transport complex of town (district, region) within the social housing construction programs. The transport will be considered at the macro-level when analyzing the organization of road transport in the construction industry of residential facilities within the framework of state housing programs (city, region, and country). To reduce transport costs, it is planned to create specialized trucking companies, which would work under these programs and would be the unstructured element of the system. The boundary of the system in this case will be a city, state, and country.

References 1. The program “Affordable housing” in Volgograd 2013–2020. http://www.dostupnoe-zhile.ru/ volgograd.html/, free 2. Velmozhin AV, Gudkov VA, Mirotin LB, Kulikov AV (2006) Freight transport by road: a textbook. Supp. EMA on education in the field of transport machines and transporttechnological complexes. Hotline-Telecom, Moscow, 560p 3. Velmozhin AV, Kulikov AV, Firsova SY (2010) The adequacy of the use of mathematical modeling in the description of the experiment in production and transport systems. News of VSTU. Series “land transport systems”. Issue 3: intercollegiate collection of scientific articles 10:136–138 4. Kulikov AV, Velmozhin AV, Firsova SY (2010) Mathematical models and their adequacy in production and transport systems. Technology, organization and management of road transport: collection of scientific papers no. 3. Siberian State Automobile and Highway Academ (SibAHU), Omsk, pp 159–164

820

A. V. Kulikov and S. Y. Firsova

5. Firsova SY, Kulikov AV (2010) Determination of the optimal variant of placement of reinforced concrete products on a semi-trailer platform. Technology, organization and management of road transport: collection of scientific papers no. 3. Siberian State Automobile and Highway Academ (SibAHU), Omsk, pp 164–168 6. Firsova SY, Kulikov AV (2012) Determination of the optimal scheme of goods placement on a car platform. Youth and scientific and technical progress in the road sector of southern Russia: materials VI of the international scientific-technical conference of students, graduate students and young scientists, 15–17 May 2012, Volgograd State University of Architecture and Civil Engineering, Volgograd, pp 295–299 7. Firsova SY, Kulikov AV (2013) Determination of the optimal layout of floor slabs on a semi-trailer platform. Youth and scientific and technical progress in the road sector of southern Russia: materials VII of the international scientific-technical conference of students, graduate students and young scientists, 14–16 May 2013, Volgograd State University of Architecture and Civil Engineering, Volgograd, pp 279–282 8. Firsova SY, Kulikov AV (2015) Systematic approach to organization of work of road transport in the housing industry. SWorld J Int Period Sci J 17(1(8)):33–36. http://www. sworldjournal.com/e-journal/j11517.pdf 9. Velmozhin AV, Kulikov AV, Firsova SY (2010) On the question of determining the minimum number of car rides during transportation of concrete goods to the object under construction. News of VSTU. Series “land transport systems”. Issue 3: intercollegiate collection of scientific articles 10:134–135 10. Kulikov AV, Firsova SY (2012) Ratio of demand for transport products and carrying capacity of rolling stock in housing construction. Problems of quality and operation of vehicles: materials VII international scientific and technical conference, 16–18 May 2012, FSBEI HVE “Penza State University of Architecture and Construction”, Automobile and Road Institute, Penza, pp 95–100 11. Kulikov AV, Firsova SY (2011) Technology of carriage of goods by road in building systems. Transport and transport-technological systems: materials international scientific and technical conference, 20 Apr 2011, SEI HVE “Tyumen State Oil and Gas University”, Tyumen, pp 153–158 12. Fadeeva IG, Kulikov AV, Metelev IS, Nuzhnova YuA, Firsova SY (2015) Innovative approaches to the development of enterprises, industries, complexes: monograph. SWorld project, Odessa, Book 1, 202p 13. Kulikov AV, Firsova SY (2012) The basic principles of the preparation of technological schemes for the carriage of goods in residential construction. Problems of quality and operation of vehicles: materials VII international scientific and technical conference, 16–18 May 2012, FSBEI HVE “Penza State University of Architecture and Construction”, Automobile and Road Institute, Penza, pp 100–104 14. Kulikov AV, Firsova SY (2010) Ways to reduce transport costs in the construction of residential facilities. Technology, organization and management of road transport: collection of scientific papers no. 3. Siberian State Automobile and Highway Academ (SibAHU), Omsk, pp 155–158 15. Firsova SY, Kulikov AV (2015) Performance evaluation of road transport when transporting construction cargoes. SWorld J Int Period Sci J (Ukr). Transp J11511:13–32 http://www. sworldjournal.come-journal/j11511.pdf 16. Kulikov AV, Firsova SY (2011) Study of the impact of technical speed on transport costs in the organization of construction cargo transportation. News of VSTU, Series “land transport systems”. Issue 4: intercollegiate collection of scientific articles 12:93–96 17. Kulikov AV, Firsova SY (2013) Reducing transport costs through the use of an effective technological scheme for the transport of construction goods. News of VSTU. Series “land transport systems”. Issue 6: intercollegiate collection of scientific articles 10(113):72–75 18. Kulikov AV, Firsova SY (2012) Existing opportunities to reduce the cost of construction of residential facilities. Actual problems of the development strategy of Volgograd: dig. art. Administration of Volgograd, MUE “City News”, Volgograd, pp 35–38

Effectiveness of Road Transport Technology …

821

19. Firsova SY, Kulikov AV (2013) Reducing transport costs due to the choice of the optimal type of pallet for transportation of construction materials. News of VSTU. Series “land transport systems”. Issue 6: intercollegiate collection of scientific articles 10(113):86–88 20. Kulikov AV, Firsova SY (2012) The use of rational technological schemes for the transport of construction goods as one of the ways to reduce the cost of housing construction objects. Actual problems of the development strategy of Volgograd: dig. art. Administration of Volgograd, MUE “City News”, pp 32–34

Robot Manipulator Control with Efforts Stabilization in Capture of Object with Fuzzy Geometrical Characteristic V. I. Chizhikov, E. V. Kurnasov and A. B. Petrov

Abstract Modern production is becoming “smart” and able to flexibly adapt to customer needs. In high-variety low-volume production, robot manipulators play an important role that has extended functionality for reliable capture of various shape objects, taking into account the uncertain and rapidly changing conditions of the production process. The authors solved the synthesis problem of the regulator control system with efforts stabilization for capturing. The movement of the robot gripper mechanism links has been investigated at the site of the effort development from the smallest value in the object tactile detection case to the maximum permissible condition, which is determined by the possibility of the object catching without the object permanent deformation. The reaction stability in contact of the gripper with the object is ensured by the selection of the numerical values offree parameters of the control law in order to minimize the amplitude of the error signal in the steady-state mode. The control law which implements the controller contains a playback error of the control signal, the first and second derivatives of the output signal. The control object in the control system is a mechanism where the control signal in the form of a gas mass flow rate is fed to controlled elastic kinematic connections. We proposed a control system for stabilizing the gripping effort may be subject to external disturbance.







Keywords Actuator Controller Control system gripper Capture Fuzzy information








Robot
Robot arm
Robot

1 Introduction Modern society requires the production of the high-variety products with a constant change in their functions, quality, and other consumer properties. Therefore, an increasing number of enterprises are focusing on high-variety low-volume production V. I. Chizhikov (&)
E. V. Kurnasov
A. B. Petrov MIREA—Russian Technological University (RTU MIREA), 78, Vernadsky Avenue, Moscow 119454, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_88

823

824

V. I. Chizhikov et al.

of products with a small number of the production series repeatability of the same product now. This circumstance requires the development of both new intellectual methods of the production processes automation [1–4] and the new development approaches to the organization of production in accordance with the concept of Industry 4.0 [5, 6]. At the same time, the basis for building digital production and an important driver for the Industry 4.0 concept implementation is robotics and automation technology. More industrial robots are evolving with the latest technological innovation to facilitate the Industrial Revolution within the concept of Industry 4.0 [7]. In high-variety low-volume production, manipulation robots must quickly adapt to a new product type. The flexibility can be achieved by reducing the diversity of the robot tongs used, i.e., by unifying the gripper design and developing a new solution that ensures reliable products capture of various shapes, taking into account the specifics of the production process. To date, there are several solutions that provide a stable capture of the object by a robot manipulator. For example, an interesting solution is [8], which allows capturing objects of an indefinite form based on visual information from a 3D sensor and heuristic characteristics of a scene with a set of possible grippers. To eliminate uncertainty regarding the location and the object geometry, tactile feedback from finger sensors is provided in the control system. In [9] proposed a robot gripper controller with a kinesthetic method of controlling the gripper fingers are based on the tactile sensation inherent in humans and not on visual information about the object. The controller processes measurements from the fingers pressure, selects an appropriate initial gripping effort, detects when the object is slipping from the gripper and increases the gripping effort as needed. If the gripper involves a change in the manipulation object relative position, then for reliable fixation in the gripper the controller requires the implementation of additional solutions that optimally meet the intended conditions for the object retention. The task of the robot fingers controlling implemented on the sliding movement of their sensitive elements along the object surface is solved in [10]. The proposed control law implies the development of control actions for given transition and steady-state tracking behavior by capturing an object. The method is interesting because it is computationally simple. However, it does not take into account the robot dynamics, which can lead in some cases to an unstable object capture. The active elements of controlling task of a robot gripper through flexible couplings with parametric uncertainty and unknown constraints of external disturbances were considered in [11]. To solve this problem, robust control is used [12]. The stability of the control system is proved by Lyapunov functions. The object stable capture is possible only in the absence of significant force fluctuations when manipulating the complex shape object. Analysis of the proposed solutions showed that the dynamics of interaction with the manipulation object under difficult contact conditions with its surface is usually not considered. It also doesn’t take into account the fact that in the multi-link actuating system, kinematic errors occur, which are subject to mandatory correction in the robot control system.

Robot Manipulator Control with Efforts Stabilization …

825

It is important not only to ensure the stability of the contact for this object reliable holding but also to take into account the dynamic interaction with the robot gripper, when a robot interacts with an object of indefinite shape and its surface physical and mechanical properties. This article proposes a solution to these problems.

2 The Actuation Device Features of the Robot Gripper Effort Stabilization and the Controller Synthesis Problem Formulation The operation system of the robot gripper considered in this paper fundamentally excludes the kinematic pairs use [13, 14], which are replaced by controlled elastic kinematic connections. The actuator mechanism construction has five kinematic chains, and each of them consists of three links and has one degree of freedom. The object-capturing process involves the simultaneously using at least two links (fingers) and subsequent involvement of other links at the process as reported by the corresponding information and measurement system. Such gripper functionality is close to the human hand actions [15], and it is possible to manipulate objects with a fuzzy geometric characteristic with using information and measurement system. When the manipulation subject geometry is changed, the gripper mechanism involves through the sensory system the required number of active kinematic chains with an appropriate adaptive justification in the capture process. Such connection kinematic error is estimated only by the elastic hysteresis, which is practically absent within the constraints of Hooke’s law. A controlled elastic element can be built on open-type shells (bellows, Bourdon tubes, and Nagatkin springs) filled with gas, which the overpressure determines the required gripping effort. We considered the conditions under which the reaction on the actuator links from the gripper’s effort is balanced by an elastic kinematic connection. The elastic reaction on the kinematic connection relative to the equilibrium position is determined by the equation: Rðk; DGÞ ¼ ðF0 þ DF ÞðDPÞ;

ð1Þ

where k; DG are, respectively, the movement of the gripper movable links, corresponding to the incoming air mass, and the mass of circulating air in the controlled elastic element cavity between the space in the equilibrium state and in a deflected state; DP is the pressure difference in the cavity between the air pressure in the equilibrium position and the current pressure value; F0 and DF are the area and the cross-sectional area increment of the controlled elastic element. The total absolute error of the reaction can be found from the expression:

826

V. I. Chizhikov et al.

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi  2  2 ffi @Rðk; DGÞ @Rðk; DGÞ Dk þ DðDGÞ : DRR ðk; DGÞ ¼  @k @ ðDGÞ

ð2Þ

Assume that all components have a normal distribution law. It considers the definition of a mass flow rate of a gas stream by standard narrowing device. In this case, we assume that the random component of the error is absent, and the correction for the flow compressibility is equal to unity. Then we represented the flow by the expression: DG ¼ af10 e10

pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi pffiffiffiffiffiffiffiffi 2qðp1  p2 Þ ¼ af10 e 2qh;

ð3Þ

where f10 is the area of the narrowing device; e10 is the correction factor for the substance compressibility, the consumption of which is measured; q is the stream density in front of the narrowing device; and h is the static pressure drop on the narrowing device. In addition, write the error in consumption in the form: ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi s 2  2ffi @ ðDGÞ @ðDGÞ DðDGÞ ¼  Dq þ Dh ; @q @h where

@ðDGÞ @q

ð4Þ

pqffiffiffiffiffiffi : ¼ af0 e10 phffiffiffiffiffiffi ; @ðDGÞ @h ¼ aF0 e10 2qh

2qh

Take into account the errors in determining the flux density. In accordance with p the equation of the gas state q ¼ RT , where p and T are, respectively, the absolute pressure and temperature of the gas in front of the narrowing device. The absolute error of determining the flux density without taking into account, the error of the gas constant will be: s ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2  2 @q @q Dp þ DT ; Dq ¼  @p @T

ð5Þ

p 1 @q where @q @p ¼ RT ; @T ¼  RT 2 : Thus, the reaction on the surface of the manipulation object is an integral assessment of the gas parameters, which is associated with the accuracy of creating the required effort on the surface of the object by elastic kinematic compounds. The playback error of the control signal DRðk; DGÞ ¼ eðtÞ at the current time is estimated as the difference between the input and output signal, which cause a reaction on the object surface. It is obvious that the mass flow rate of air (3) should be used as an input signal xðtÞ, the developed reaction (1) on a moving element of the gripper as an output signal yðtÞ, and the change in gas density (5) as a disturbance f ðtÞ.

Robot Manipulator Control with Efforts Stabilization …

827

The formulation of the problem of controller synthesis is formulated as follows. 1. To determine the values of the free parameters of the control law of the gripper effort in such a way as to minimize the amplitude of the error signal eðtÞ in the steady-state mode, i.e., find e ¼ minðmaxest ðtÞÞ under the control signal constraint jUðtÞj  U0 . 2. To find the expression for yst ðtÞ with free parameters of the controller xðtÞ ¼ 0, f ¼ f0 cosðxtÞ and checking for the amplitude constraint of the control signal U0 . 3. To determine est ðtÞ with f ðtÞ ¼ 0 and xðtÞ ¼ C1 t þ C0 .

3 The Controller Synthesis of Control System of the Capture Efforts Stabilization Consider a system for automatic control of the capture effort by a robot, the functional diagram of which is shown in Fig. 1. The equation of a controlled object that can experience accelerations and pulse loads: a3 y000 þ a2 y00 þ y0 ¼ kU U þ kf f ;

ð6Þ

where ku  0; kf  0—the object’s gain factors in relation to the control signals and disturbance, respectively.

3.1

The Block Diagram of the Control System of the Capture Efforts Stabilization

The control object contains second- and third-order derivatives, which imposes corresponding requirements on the controller, which must parry them [9]. As a result, the control law that the controller must implement is:

Fig. 1 Functional diagram of the automatic control system of the capture effort: x(t) is the input signal to be reproduced by the output signal y(t); f(t)—is the disturbance; U(t)—control signal implemented by the controller

828

V. I. Chizhikov et al.

UðtÞ ¼ ke ðeðtÞ  b1 y0 ðtÞ  b2 y00ðtÞÞ;

ð7Þ

where eðtÞ ¼ xðtÞ  yðtÞ is the reproduction input signal error at the current time. Write the expression (7): UðtÞ ¼ ke ðxðtÞ  y1 ðtÞÞ; M ðyðtÞÞ ¼ y1 ðtÞ ¼ yðtÞ þ b1 y0 ðtÞ þ b2 y00 ðtÞ;

ð8Þ

ke ; b1 ; b2 are free parameters are selected by the developer. With a reasonable choice of free parameters, the structure of the control law (2) allows increasing the input signal reproduction accuracy (the main task). Write the equation of the controlled object in operator form: yðpÞða3 p3 þ a2 p2 þ pÞ ¼ kU UðpÞ þ kf f ðpÞ: The operator form of Eq. (8) has the form: UðpÞ ¼ ke ðxðpÞ  y1 ðpÞÞ; y1 ðpÞ ¼ yðpÞð1 þ b1 p þ b2 p2 Þ ¼ yð pÞM ð pÞ Using the proposed recording form, obtain the transfer functions of the controlled object and the controller. Functions for controlling, disturbing influences and feedback have the form: Wy;U ðpÞ ¼

yðpÞ ku ¼ ; UðpÞ pða3 p3 þ a2 p2 þ 1Þ

ð9Þ

Wy;f ðpÞ ¼

yðpÞ ku ¼ ; f ðpÞ pða3 p3 þ a2 p2 þ 1Þ

ð10Þ

y1 ðpÞ ¼ MðpÞ ¼ 1 þ b1 p þ b2 p2 : yðpÞ

ð11Þ

Wy1 :y ðpÞ ¼

Thus, the managed object has astatism of the first order. Now the functional diagram in Fig. 1 corresponds to the block diagram in Fig. 2. As mentioned above, it’s imposed constraints on the control signal in the form jUðtÞj  U0 . Next, determine the transfer function of the closed system (see Fig. 2) taking into account (10) and (11) by the control signal. Taking into account simple transformations, we have the following expression:

Robot Manipulator Control with Efforts Stabilization …

829

Fig. 2 Block diagram of the control system stabilization efforts

Uy;x ð pÞ ¼

Y ðP Þ ¼ X ðP Þ

a3 ke kU

p3

þ



a2 ke kU



1

þ b2 p2 þ



1 ke kU

 : þ b1 p þ 1

ð12Þ

The transfer function of a closed-system disturbance signal is: Uy;f ð pÞ ¼

a3 p3 ke kU

þ



kf ke kU

a2 ke kU

   : þ b2 p2 þ ke1kU þ b1 p þ 1

ð13Þ

The transfer function of a closed system by error is: Ue;x ð pÞ ¼

E ð pÞ X ð p Þ  Y ð pÞ ¼ ¼ 1  Uy;x ð pÞ: X ð pÞ X ð pÞ

ð14Þ

Substituting in (14) the expression (12), it’s got: Ue;x ð pÞ ¼

Eð pÞ 1    ¼1 3  : a p a2 1 3 X ðP Þ 2 k e k U þ k e k U þ b2 p þ k e k U þ b1 p þ 1

ð15Þ

ð pÞ The transfer function of the control signal Uu;x ðpÞ ¼ UXðpÞ can be represented taking into account the operator form for the controller. It is possible to represent a control signal through the following sequence:

U ð pÞ ¼ ke e1 ¼ ke ðxð pÞ  y1 ð pÞÞ ¼ ke ðxð pÞ  yð pÞM ð pÞÞ ¼ ke xð pÞ  M ð pÞWy;U ð pÞU ð pÞ :

830

V. I. Chizhikov et al.

There is a rewrite form of the control signal taking into account simple transpÞ . Then the transfer function of formations of the last equality U ð pÞ ¼ 1 þ ke MkeðxpðÞW y;U ð pÞ the control signal will be: UU;x ðpÞ ¼

ke k e ð a3 p3 þ a2 p2 þ pÞ ¼ : ð16Þ 1 þ ke M ð pÞWy;U ð pÞ a3 p3 þ a2 p2 þ p þ ke kU ðb2 p2 þ b1 p þ 1Þ

Next, transform the last equality to the following form: UU;x ðpÞ ¼

a3 kU ke

p3 þ



a2 kU ke

ða3 p3 þ a2 p2 þ pÞ kU   ; þ b2 p2 þ kU1k þ b1 p þ 1

ð17Þ

e

that is convenient for further analysis. Further, find the transfer function of the disturbing signal for the controller WU;f ð pÞ ¼ Uf ððppÞÞ with the input signal xðtÞ ¼ 0. In this case, the playback error of the input signal at the current time is equal to e1 ðtÞ ¼ y1 ðtÞ. Then the control signal in the operator form can be represented by the equality, which after simple transformations, taking into account (8)–(10) will have the form: WU;f ¼

U ð pÞ ke M ð pÞWy;f ð pÞ ke kf ðb2 p2 þ b1 p þ 1Þ ¼ ¼ : f ð pÞ 1 þ ke M ð pÞWy;U ð pÞ a3 p3 þ a2 p2 þ p þ ke kU ðb2 p2 þ b1 p þ 1Þ

There is the last expression in the form: WU;f ¼¼

ke M ð pÞWy;f ð pÞ ¼ 1 þ ke M ð pÞWy;U ð pÞ

k

 kUf ðb2 p2 þ b1 p þ 1Þ    : a3 3 a2 1 2þ p þ þ b þ b p p þ 1 2 1 ke kU ke kU ke kU 

ð18Þ Express in an operator form, the output signal of the control system yð pÞ with the presence of input signal xð pÞ and disturbance f ð pÞ: yð pÞ ¼ Uy;x ð pÞxð pÞ þ Uy;f ð pÞf ð pÞ: Taking into account (10) and (12), write the output signal yð pÞ in the form k

y ð pÞ ¼

xð pÞ þ ke kf U f ð pÞ    : a3 3 a2 1 2 k e k U p þ k e k U þ b2 p þ k e k U þ b1 p þ 1 

ð19Þ

Robot Manipulator Control with Efforts Stabilization …

3.2

831

Definition of the Controller Free Parameters

Use the method of approximate transfer functions for determining the controller free parameters, coefficients ke ; b1 ; b2 . A monotonous transient process can be provided a third-order transfer function for a closed system Qð pÞ, and the transient process there depends on the degree of stability g, with the growth of which the process time decreases. The constraint on the control signal jUðtÞj  U0 does not allow assigning an infinitely large value g. We have the following expression of the  3 transfer function Qð pÞ ¼ 1g p þ 1 : Write the last expression in expanded form: Qð p Þ ¼

1 3 3 3 p þ 2 p2 þ p þ 1: 3 g g g

ð20Þ

Convert the transfer function Uy;x ð pÞ to the form: Uy;x ð pÞ ¼

1 at Qð0Þ ¼ 1: Qð pÞ

ð21Þ

Taking into account (12) and (20) as a result of substitution in equality (21) and comparison of coefficients with the same degrees, obtain the expression:

a3 ke kU

p3 þ



a2 ke kU



1

þ b2 p2 þ



1 ke kU

 ¼ 1 3 þ b 1 p þ 1 g3 p þ

3 g2

1 ; p2 þ 3g p þ 1

ð22Þ

from which the system of three linear equations follows, and on its basis, we find parameters ke ; b1 ; b2 with known values kU ; a3 ; a2 : ke ¼

a3 g3 ; kU

b2 ¼

3 a2  ; g2 a3 g3

3 1 b1 ¼  ; g a3 g3

where g [ 0:

ð23Þ

Found parameters are functions of an unknown quantity g. This approach is due to the requirement for the transition process of the transfer function of a closed system Qð pÞ and the degree of stability g. For a sufficiently wide class of external influences, with the g increase the reproduction, the error of the input signal and maximum deviation in the stabilization mode decrease.

3.3

Calculation of Amplitude-Frequency Characteristics of the Controller

Define the amplitude-frequency characteristics Ae;x ðxÞ, AU;x ðxÞ. Since in the steady-state mode with xðtÞ ¼ x0 cosðxtÞ, the error value and the control signal will be:

832

V. I. Chizhikov et al.

  est ðtÞ ¼ Ae;x ðxÞx0 cos xt þ /e;x ðxÞ ;

ð24Þ

  Ust ðtÞ ¼ AU;x ðxÞx0 cos xt þ /U;x ðxÞ ; than max est ¼ Ae;x ðxÞx0 ; max Ust ¼ AU;x ðxÞx0 :

ð25Þ

Taking into account the above, add a degree of stability as an argument in the amplitude-frequency characteristic and, further, we will be denoted the amplitude-frequency characteristic as Ae;x ðx; gÞ, AU;x ðx; gÞ. A sufficiently small value of the function Ae;x ðx; gÞ ensures a fairly accurate reproduction of the signal xðtÞ ¼ x0 cosðxtÞ in the steady-state mode, and a sufficiently small value of the function AU;x ðx; gÞ at the condition jUðtÞj  U  . The transfer function of a closed system by error (15) taking into account (22) can be written in the form Ue;x ð pÞ ¼ 1  1 p3 þ 3 p12 þ 3p þ 1. Turning to the frequency specg3

g2

g

trum, obtain the following expressions for the transfer function of the amplitude-frequency characteristic Ae;x ðx; gÞ:    g32 x2 þ j 3g x  g33 x3   ; Ue;x ðx; gÞ ¼  1  g32 x2 þ j 3g x  g33 x3 vffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi  2 u u 9 4 3 3 3 x þ x  x 4 3 u g g g Ae;x ðx; gÞ ¼ Ue;x ðx; gÞ ¼ u 2  2 : t 3 2 3 1  g2 x þ j g x  g33 x3 Introduce a new variable # ¼

ð26Þ

ð27Þ

 2 x g

. Then, get after the obvious transformations

and the substitution of a new variable #: rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi #3 þ 3#2 þ 9# Ae;x ðx; gÞ ¼ : #3 þ 3#2 þ 3# þ 1

ð28Þ

Similarly, get for UU;x ðx; gÞ and AU;x ðx; gÞ taking into account (17), without giving intermediate transformations:    aku2 x2 þ j k1u x  aku3 x3 ; UU;x ðx; gÞ ¼  3 1 þ 1g ðjxÞ

ð29Þ

Robot Manipulator Control with Efforts Stabilization …

qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 2 2 2 x ð1  a3 x Þ þ a2 x AU;x ðx; gÞ ¼ jUðx; gÞj ¼ !3 : r ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ku  2 1 þ xg

833

ð30Þ

Considering the amplitude-frequency characteristic Ae;x ð#Þ, it is easy to show that the function has a single maximum max# Ae;x ð#Þ ¼ Ae;x ð# Þ ¼ 1:286 at the   2 pffiffi  point # ¼ xg ¼ 0:75, i.e. xg ¼ 23 and, that, lim#!1 Ae;x ð#Þ ¼ 1, Ae;x ð0Þ ¼ 0: It is known that for a wide class of external influences the maximum error in the reproduction of the input signal and the maximum deviation in the stabilization mode decreases with increasing g. In the steady-state mode when xðtÞ ¼ x0 cosðxtÞ there is eycT ðtÞ ¼ Ae;x ðx; gÞx0 cosðxt þ /e;x ðxÞÞ and to have the minimum value of the maximum error in the steady-state mode, it is necessary to choose the smallest  2 possible value # ¼ xg , that is a larger value g. However, as the degree of stability g increases, the value max AU;x ðx; gÞ may increase. As a consequence, the condition jU ðtÞj  U0 may be violated, since with a harmonic input signal xðtÞ ¼ x0 cosðxtÞ the steady-state mode for the signal has the form U ðtÞ ¼   AU;x ðx; gÞx0 cos xt þ /U;x ð/Þ and AU;x ðx; gÞx0  U0 , where AU;x ðx; gÞ is determined by (30). Ensure that the inequality AU;x ðx; gÞ0  Ux00 resulting from the constraint on the control signal is equivalent to the inequality: 1 1  WðzÞ ¼ 2 g z

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ! zx20 3 2 2 ðð1  a3 zÞ þ a2 zÞ  1 : ðu0 ku Þ2

ð31Þ

Denote for the evidence z ¼ x2 and rewrote the expression (30) as pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi pffi ð1a3 zÞ2 þ a22 z z AU;x ðx; zÞ ¼ ku rffiffiffiffiffiffiffiffiffiffiffiffiffiffi  3 . Raise this expression will be in a square and divide 1þ

by

z g2

U02 , x20

which takes into account the inequality and transformations will have the rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi  zx20 3 2 2 z  1  1 . Denoting the left-hand side of the form 1z ð 1  a z Þ þ a 3 2 g2 U 2 k2 0 u

inequality as WðzÞ, obtained inequality (31), which was required to be proved. We have the following expression for g from (31): 1 g  pffiffiffiffiffiffiffiffiffiffi : W ðzÞ

ð32Þ

834

V. I. Chizhikov et al.

4 Analysis of the Numerical Solution of the Synthesis Problem Consider the nature of the behavior of the function WðzÞ with the following values of the parameters: ku ¼ 0:25; U0 ¼ 30; a3 ¼ 0:4; f0 ¼ 0:65; c0 ¼ 3; c1 ¼ 0:4:

a2 ¼ 0:6;

x0 ¼ 8:5;

x0 ¼ 0:7;

Taking into account the specified parameters, the investigated function takes the following form (Fig. 3): 1 WðzÞ ¼ z

! rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 3 0:49z 2 ð1  0:4zÞ þ 0:36z  1 : 56:25

The graph increases on the half-open interval z ¼ ð0; z0 , where z0 ¼ x20 ¼ 8:52 ¼ 72:25: Consequently, the point z0 is the desired maximum point WðzÞ ¼ 0:25 (see Table 1). From the expression (32), we have a valid value of the degree of stability 1 1 ¼ pffiffiffiffiffiffi ¼ 2. According to the found maximum value of the degree of g  pffiffiffiffiffiffiffi 0:25 WðzÞ

stability, find the numerical values of the free parameters ke ; b1 ; b2 : Using the initial data of the example and the three equalities (23), obtain the numerical values of the desired controller parameters:

Fig. 3 Function of the control signal WðzÞ in determining the acceptable degree of stability g

Robot Manipulator Control with Efforts Stabilization …

835

Table 1 Calculated values of the inverse function of the stability coefficient z WðzÞ

0.1 −7.9

ke ¼

0.5 −1.67

0:4g3 ¼ 12:8; 0:25

5 −0.08

b1 ¼

20 0.2

30 0.211

3 1  ¼ 1:19; g 0:4g3

40 0.22

b2 ¼

50 0.225

60 0.23

72 0.25

3 0:6  ¼ 0:56; g2 0:4g3

and therefore the equation of the controlled object: 0:125y000 þ 0:75y00 þ 1:5y0 þ 1 ¼ xðtÞ þ 0:18f ðtÞ:

ð33Þ

Using the Hurwitz criterion, it is easy to verify that the system is stable. To accomplish paragraph 2 of the task with the found parameters of the controller, it is necessary to find expressions yst ðtÞ where xðtÞ ¼ 0, f ¼ f0 cosðxtÞ when checking for constraint of the amplitude of the control signal U0 . Use the transfer functions Uy;f ð pÞ and UU;f ð pÞ for this task, which are defined above. Find   yst ðtÞ ¼ Ay;f ðxÞf0 cos xt þ /y;f ðxÞ . Substituting the numerical values into the transfer functions, find their absolute values and the corresponding phase angles for x ¼ x1 ¼ 0:2: 0:6 0:18 0:25ke Ay;f ðxÞ ¼ Uy;f ðxÞ ¼  ¼ pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ¼ 0:2;  3 1 ðjxÞ þ 1 0:7225 þ 0:0894 2 2:4ð0:56x2 þ 1:19jx þ 1Þ AU;f ðxÞ ¼ UU;f ðxÞ ¼ ¼ 2:48; 1 3 jx þ 1 2 x ¼ x1 ¼ 0:2; /y;f ðxÞ ¼ arctg

1:5x  0:125x3 ¼ 19:2 1  0:75x

The steady-state value of the output signal in the presence of a disturbance can be found yst ðtÞ ¼ 0:2
0:65 cosð0:2t  19:2Þ. The steady-state value of the control signal in the presence of a disturbance is calculated without a phase angle /U;f ðxÞ, since we find the maximum value Ust ðtÞ: Therefore, Ust ðtÞ ¼ AU;f ðx; gÞf0 ¼ 2:48
0:65 ¼ 1:6\U0 ¼ 30: The value of the steady-state error est ðtÞ where f ðtÞ ¼ 0 (i.e., with the absence of disturbances) and the linear input signal xðtÞ ¼ c1 t þ c0 is determined taking into account k ¼ kU ke ¼ 0:25
12:8 ¼ 3:2. Then the value of the steady-state error will be est ¼ ck1 ¼ 0:4 3:2 ¼ 0:13; that corresponds to the accuracy of the control system with a polynomial signal. With constant exposure the error in the system with first-order astatism is equal by zero. The steady-state control signal will be Ust ¼ est ke ¼ 0:13
12:8 ¼ 1:66.

836

V. I. Chizhikov et al.

5 Conclusion At the work result, the authors proposed an actuating system for the robot manipulator gripping with controlled elastic kinematic connections on the basis of an open-type loop, which makes it possible to ensure the control implementation when performing capturing operations of objects having a fuzzy geometric characteristic. We considered the conditions under which the reaction at the links of the robot gripper actuator from the gripping effort is balanced by the excess gas pressure on the elastic controlled kinematic connection. The control object is a mechanism where the control signal in the form of a gas flow rate is fed to controlled elastic kinematic connections made on open-type loops. We assumed that the system may be disturbed, which is estimated by a change in gas density. Our research group solved the problem synthesis of the control system controller for gripping effort stabilization which increases the accuracy of the input signal. We obtained parameters of the control law, which allow minimizing the amplitude of the error signal under the constraints on the control signal during the development of the mentioned action. Besides, we obtained the expression of the output signal in the steady-state mode (reaction on the object surface) for the found free parameters of the controller in the presence of a disturbance with a check for constraining the amplitude of the control signal. Also we determine the value of steady-state error in the absence of a disturbing signal. Acknowledgements This work was supported by the Russian Foundation for Basic Research (project no. 19-08-00775).

References 1. Zheng P, Wang H, Sang Z, Zhong Ray Y, Liu Y, Liu C, Mubarok K, Yu S, Xu X (2018) Smart manufacturing systems for Industry 4.0: conceptual framework, scenarios, and future perspectives. Front Mech Eng 13(2):137–150. https://doi.org/10.1007/s11465-018-0499-5 2. Kashirskaya EN, Kurnasov EV, Kholopov VA, Shmeleva AG (2017) Methodology for assessing the implementation of the production process. In: Proceedings of 2017 IEEE 2nd international conference on control in technical systems, CTS 2017, IEEE, pp 232–235. https://doi.org/10.1109/ctsys.2017.8109533 3. Kashirskaya EN, Kholopov VA, Shmeleva AG, Kurnasov EV (2017) Simulation model for monitoring the execution of technological processes. In: Proceedings of 2017 IEEE 2nd international conference on control in technical systems, CTS 2017. IEEE, pp 307–310. https://doi.org/10.1109/ctsys.2017.8109553 4. Holopov V, Kushnir A, Kurnasov E, Ganichev A, Romanov A (2017) Development of digital production engineering monitoring system based on equipment state index. In: Proceedings of the 2017 IEEE Russia section young researchers in electrical and electronic engineering conference, ElConRus 2017, IEEE, pp 863–868. https://doi.org/10.1109/eiconrus.2017. 7910692

Robot Manipulator Control with Efforts Stabilization …

837

5. Thoben K-D, Wiesner S, Wuest T (2017) “Industrie 4.0” and smart manufacturing—a review of research issues and application examples. Int J Autom Technol 11(1):4–19. https://doi.org/ 10.20965/ijat.2017.p0004 6. Kholopov VA, Kashirskaya EN, Kushnir AP, Kurnasov EV, Ragutkin AV, Pirogov VV (2018) Development of digital machine-building production in the Industry 4.0 concept. J Mach Manuf Reliab 47(4):380–385. https://doi.org/10.3103/s1052618818040064 7. Bahrin MAK, Othman F, Azli NHN, Talib MF (2016) Industry 4.0: a review on industrial automation and robotic. J Teknol (Sci Eng) 78(6–13):137–143. https://doi.org/10.11113/jt. v78.9285 8. Hsiao K, Chitta S, Ciocarlie M, Gil Jones E (2010) Contact-reactive grasping of objects with partial shape information. In: 2010 IEEE/RSJ international conference on intelligent robots and systems. Taiwan, pp 1228–1235. https://doi.org/10.1109/iros.2010.5649494 9. Romano JM, Hsiao K, Niemeyer G, Chitta S, Kuchenbecker KJ (2011) Human-inspired robotic grasp control with tactile sensing. IEEE Trans Rob 27(6):1067–1079. https://doi.org/ 10.1109/TRO.2011.2162271 10. Okur B, Aksoy O, Zergeroglu E, Tatlicioglu E (2015) Nonlinear robust control of tendon-driven robot manipulators. J Intell Rob Syst 80(1):3–14. https://doi.org/10.1007/ s10846-014-0141-7 11. Droukas L, Doulgeri Z (2016) Rolling contact motion generation and control of robotic fingers. J Intell Rob Syst 82(1):21–38. https://doi.org/10.1007/s10846-015-0255-6 12. Gao X, Dawson D, Qu Z (1992) On the robust control of two manipulators holding a rigid object. J Intell Rob Syst 6(1):107–119. https://doi.org/10.1007/BF00314701 13. Manko SV, Shestakov EI (2018) Automatic synthesis of gait scenarios for reconfigurable mechatronic modular robots in the modification of the walking platform. Ross Tekhnol Zh (Russ Technol J) 6(4):26–41 14. Vorob’ev EI, Dorofeev VO (2017) Orientation mechanism with linear drives for robot manipulators and limb prostheses. Russ Eng Res 37(6):475–478. https://doi.org/10.3103/ s1068798x17060235 15. Chizhikov VI, Kurnasov EV, Vorob’ev EI (2018) Capture of an object on the basis of tactile surface recognition. Russ Eng Res 38(4):251–255. https://doi.org/10.3103/s1068798x1804 0044

Minimization of Energy Costs for Movement Resistance of Ground for Walking Device by the Control of Support Points Motion V. V. Chernyshev, V. V. Arykantsev and I. P. Vershinina

Abstract Tests show that they are simple and reliable and can work effectively in extreme conditions, for example, on waterlogged and underwater soils. The disadvantage of cyclic walking mechanisms is the constant “program” trajectory of the support point. This limits the abilities of the walking machine on the shape passability. In the underwater walking machine MAK-1, the possibility of adjusting the program movements of the legs was achieved by introducing an additional controlled degree of freedom into the mover. Thus, the control (within a limited range) of the movement of the control points is achieved. Changing the law of motion of the support point leads to a change in the structure of energy consumed by the movement. The paper considers a possibility of mutual compensation of energy consumption for pressing the soil with other components of energy consumption by controlling the movement of the support points of the walking mover. A comparative analysis of energy consumption for the movement of the wheel and walking mover in the low-speed range is also carried out. The analysis is based on the results of dynamic modeling and generalization of experience in the development of experimental walking machines units. It is shown that via controlling the movement of the points of the walking mover, it is possible to achieve a reduction in energy consumption for the ground resistance to the movement of the walking device. The reduction appears at the cost of recuperation of energy spent on lifting the machine body in each cycle of movement.










Keywords Mobile robots Seabed robots Robots mechanics Movement dynamics Controlled movement Energy consumption Energy recuperation Mathematical modeling













V. V. Chernyshev (&)
I. P. Vershinina Volgograd State Technical University, Lenin Avenue 28, Volgograd 400005, Russia e-mail: [email protected] V. V. Arykantsev Robotics Development Center, Innopolis University, Universitetskaya str. 1, Innopolis City 420500, Russia © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_89

839

840

V. V. Chernyshev et al.

1 Introduction Tests in real conditions of experimental units of walking machines with cyclic movers show that they are simple and reliable and can work effectively in extreme situations [1–7]. Such machines can be in demand when carrying out the ground works in underwater conditions. The practice of underwater technical works sets a number of tasks related to the conduct of the ground works on the bottom. Existing machines moving along the bottom (underwater bulldozers, self-moved bottom mining units, cable layers, etc.) have, as a rule, a tracked mover [8–12]. However, in the conditions of their exploitation, characterized by soils with the low-bearing ability and rough bottom shape, traditional types of movers are inefficient [13–15]. The walking mover, with higher soil and shape passability, is more suitable for use in such conditions. Walking machines also have decreased traction force on the resistance to movement [16]. For walking movers, unlike wheeled and tracked, the ground is not an obstacle for movement, but only requires the necessary power costs to press it. When vehicles move on weak and waterlogged soils, the costs of overcoming the ground resistance can reach up to 20–40% of all power costs. In the work, on the basis of the analysis of the structure of energy consumption of the walking type of movement, the possibility of mutual compensation of the energy consumption for the pressing of soil with other components of energy consumption due to control the movement of control points of the mover was investigated. Walking machines with cyclic type of mover were considered. A comparative analysis of energy consumption for the movement of the wheel and walking mover in the low-speed range is also carried out. The analysis is based on the results of dynamic modeling and generalization of the experience of development and testing of the robotic complex “Vosminog” [17, 18] and underwater walking device MAK-1 [19].

2 Adjustment of the “Program” Leg Movements in Walking Machines with Cyclic Movers Walking robotic complex Vosminog (Fig. 1a) and underwater walking device MAK-1 (Fig. 1b) developed in VSTU. The Vosminog is designed to work on waterlogged grounds and soils with low-bearing ability. There is an experience of using the robot in swamp areas. Walking device MAK-1 designed for the testing of methods of motion control of underwater walking robotic and optimization of parameters of walking mechanisms at the development stage. Both walking robots have been tested in conditions of waterlogged and underwater soils. Tests have confirmed their increased traction properties and high soil and shape passability [19–22]. The use of cyclic movers allows not to care about the saving and stability of gait and eliminates the need of sophisticated controlled adaptation system. As a result,

Minimization of Energy Costs for Movement Resistance …

841

Fig. 1 a Walking robotic complex “Vosminog” and b underwater walking device MAK-1

machines have a minimum number of controlled degrees of freedom and become significantly more reliable, easier and cheaper than analogues with adaptive control. For example, the mover of the MAK-1 (Fig. 1b) consists of three kinematically connected cyclic walking mechanisms of one side. Outer walking mechanisms work in the same phase and the average in the opposite phase. The movers of the right and left sides have independent actuators. As a result, the robot has only two controllable degrees of freedom. The limbs of the robot Crabster, for comparison, have 30 controlled actuators [23]. To improve the adaptability to the ground roughness in the walking mechanism implemented system of passive foot control [24]. The walking mechanism provides rising of the toe of the foot with hinge joint both in straight and reverse movement (without additional actuator). With that, modeling of dynamics of similar walking machines movement presents that for full realization their abilities on the ground passableness and maneuverability the ability for combining and correction of legs motion is necessary. In the underwater walking machine MAK-1, the possibility of adjusting the program foot movements was achieved by introducing an additional controlled degree of freedom into the mover. In walking mechanism (Fig. 2a) was implemented an additional turning link O1 O2 . The control performs as discrete alteration of its angular position (as gearshifting in traditional vehicles). It leads to shifting of arm point of support of the walking mechanism (for example, from O1 to O01 ) and to transformation of forward movement base trajectory (Fig. 2b) into one with increased height and length of step (Fig. 2d). The points on the trajectories are located at equal intervals. It is possible to implement a large number of intermediate paths, for example, Fig. 2c. Thus, the control (within a limited range) of the movement of the control points is achieved. Moreover, there is a change in the trajectory itself, as well as a change in the law of motion of the support point along the trajectory. Changing the law of motion of the reference point leads to a change in the structure of energy consumed by the movement.

842

V. V. Chernyshev et al.

Fig. 2 a Scheme of walking mechanism of the device MAК-1 and b–d transformation of it’s supporting point: 1—main winch; 2—supporting link; 3—arm; 4—foot; 5—linear electrical drive; O1O2—controlled turning link of movement supporting point

3 The Structure of Energy Consumption of a Walking Machine with Cyclic Type of Mover In the analysis of energy costs for movement, the power for passing of the ground resistance was calculated as: W f ¼ kf vx G

ð1Þ

where vx—velocity; kf—movement resistance coefficient; G—weight of walking machine. For walking mover, approximate values of kf, without experimental data, were calculated as kf ¼ Af GS, where Af—mechanical work for the ground pressing at; S—transfer length of the machine at cycle. In our case, pressing of the ground at cycle (2 steps) happens twice, that’s why Af ¼ 2Gh, where h—depth of the track. The analysis showed that in hard movement conditions (swamp, sand, etc.) swapping of wheeled mover to walking provides lower Wf approximately by 2–3 times. However, in walking machines movement besides power costs for the ground pressing there are some other power costs. During walking, there is irregularity of straight movement and oscillations of the body in each step. This requires an additional energy costs, which include power for overcoming of gravity force W0G , power for overcoming of inertial forces of body W0U and power for overcoming of inertial forces in mover, which consist of power costs for overcoming of inertial forces in links of walking mechanisms WMU :

Minimization of Energy Costs for Movement Resistance …

WMU ¼

843

W0G ¼ Gvz

ð2Þ

W0U ¼ U0x vx þ U0z vz

ð3Þ

4 X 4 X 2   X Uijk x x_ Cijk þ Uijk z z_ Cijk þ MijUy xij y

ð4Þ

i¼1 j¼1 k¼1

where vz—vertical machine speed; U0x ¼ m0 ax , U0z ¼ m0 az —straight and vertical components of the main vector of body inertial forces; ax, az—its straight and vertical acceleration; m0—body weight; Uijk;x , Uijk z —straight and vertical inertial forces of ith link jth walking mechanism kth board side; x_ Cijk , z_ Cijk —straight and vertical velocities of center of mass of links; MijUy and xij y —moment of inertial forces of link and his angular speed. To determine the velocities and accelerations of the nodal points of the walking mechanisms, the relative motion of the links of the walking mechanism and the absolute motion of the body were studied. The mechanisms of walking were considered as flat four-segment mechanisms. In making of differential equations of motion of the links of the mechanism, their angular velocities are expressed in terms of the speed of the points, which are superimposed on the external connections. Equations for the velocities of the nodal points, including the control points, are obtained sequentially, from link to link, by considering the movements of solids. For the point M, which is the nodal point m and m + 1 links, the expression for the velocity vector in the frame of reference associated with the body has the form: ðrÞ

vM ¼ vO þ

m X

xi  li ;

ð5Þ

i¼1 ðrÞ

ðrÞ

where vO —pole speed (in the selected frame of reference vO ¼ 0); xi ; li —the angular velocity vector and the vector connecting the nodal points of the ith link of the mechanism. For the points of the mechanism on which external relations are imposed, the projections of the vector Eq. (1) on the coordinate axis are converted into the equations of relations. For the considered four-link flat walking mechanism, three such equations were made. Solving them together with respect to the angular velocities of the links of the mechanism, analytical expressions of the form were obtained u_ 1 ¼ f1 ðtÞ; u_ 2 ¼ f2 ðu1 ; u2 ; u3 ; x1 Þ; u_ 3 ¼ f3 ðu1 ; u2 ; u3 ; x1 ; x2 Þ;

ð6Þ

844

V. V. Chernyshev et al.

where f1(t)—the law of rotation of the driving crank; u1 ; u2 ; u3 —the angles formed by the links of the mechanism with the axis Ox; x1 ; x2 ; x3 —angular speeds of links. Equations (5) and (6) were also supplemented with formulas for relative coorðrÞ dinates zCj of support points Cj (j = 1, 2), of mechanisms working in the antiphase, necessary to determine the moment of feet changing. Acceleration links were by differentiating of their speeds. As a result, a system of equations defining the motion of walking supports was obtained. The movement of the sides of the body of the walking machine was given by kinematic equations (it was believed that the ground is solid, and there is no feet slipping): vk ¼ 

J X j¼1

ðrÞ

Ujk vCjk ;

ak ¼ 

J X

ðrÞ

Ujk aCjk ;

ð7Þ

j¼1

where vk, ak—velocity and acceleration vectors of the kth (right and left) side of the ðrÞ ðrÞ walking machine (k = 1, 2) in absolute motion; vCjk ; aCjk —relative velocities and accelerations of control points of jth walking mechanism of mover of kth side; Ujk—a unit function describing the state of jth leg, which is equal to 1 in contact phase and to 0 in transfer. It was assumed that in the contact phase is a walking ðrÞ mechanism with support point with the smallest coordinate zCj in the reference frame associated with the body.

4 Results A mathematical modeling of the dynamics of the walking machine MAK-1, determined its structure of energy consumption in different speed ranges. Calculations have shown that in the range of speeds 3–7 km/h the main amount of additional energy costs W is associated with the power going to overcome the gravity and inertial forces of the body of the walking machine. The structure of power consumption per cycle (two steps) for a walking device with weight 500 kg at a speed of 5 km/h is shown in Fig. 3. In the figure on the abscissa axis, there are ordinal numbers of points of the relative trajectory (Fig. 2b) the support point of the walking mechanism (points on the trajectory are located after 1/24 cycle period). In walking machines, unlike wheeled and tracked, the loss of power to overcome the ground resistance there are only at the part of the cycle—at the time of the feet changing (during step). Energy costs for pressing the soil depend on its rheological properties and the character of the part of relative trajectory of the support points, corresponding to the moment of feet changing. Character of alteration Wf for walking device MAK-1 for linearly elastic soil at the depth of the track h = 0.1 m presented in Fig. 3 curve 4. Feet changing with that track depth in straight

Minimization of Energy Costs for Movement Resistance …

845

Fig. 3 Structure of power consumption per cycle at speed 1 km/h: WMU (curve 1); W0U (curve 2); W0G (curve 3); Wf (curve 4)

movement mode (Fig. 2b) will be at range 7–10 and 19–22 of trajectory, where the power Wf (curve 4, Fig. 3) will be increasing together with the ground reaction underfoot, which starts the contact phase, from zero to maximum value. In feet changing, the power Wf and W0G (Fig. 3) have different signs. Power to overcome the ground resistance is almost completely provided by the recuperation of energy, spent on lifting the body in the previous part of the cycle. In result, an additional costs W have a small influence on the average value of the power for the movement at cycle and walking mover more energy-efficient than wheeled and tracked. Because Wf and W0G are proportional to straight velocity, the same situation will also be observed with low movement speed. The power for overcoming of cyclic inertial forces of the body and mover in presented speed range is comparably small. With an increasing of the speed, power W0U and WMU , proportional to the cube of the forward speed, become the most significant in the structure of energy costs for movement and the walking mover begins to yield in energy costs to the wheeled and tracked movers. When increasing the length of the step, for example, when moving from the trajectory Fig. 2b to trajectory Fig. 2d, the average value of the power required for motion must be reduced by reducing the number of deformations of the soil per unit of the way. On the other hand, due to the higher step height, the feet changing time is reduced approximately by 1.5–2 times. The feet changing will occur not on the sections 7–10 and 19–22 of the trajectory, as in the considered case, but on the sections 9–11 and 22–24 of the trajectory Fig. 2d. Since the area of the triangle under the curve 4 (Fig. 3), characterizing the work on pressing the soil for half a cycle does not change, and the base of the triangle has decreased by 1.5–2 times, then the same growth of the peak of the curve 4 will take place. As a result,

846

V. V. Chernyshev et al.

the power to the soil pressing is provided by energy recuperation to a much lesser extent. Also, from a practical point of view, it is important not only the average value of the required power for the movement, which determines the energy consumption, but also its maximum value, which determines the drive power. In the latter case, the maximum value of the required power for the movement will coincide with the maximum curve 4 and will be about 11 kW, which is several times higher than with the wheel drive. Unreasonable increasing of the power of the motor is undesirable. By correcting the law of motion of the support point of the walking mechanism, aimed at increasing the area corresponding to the change of feet and, accordingly, the basis of the triangle characterizing the work on pressing the soil, the specified drawback can be partially eliminated. The required trajectory correction can be achieved by shifting the suspension point of the rocker arm of the walking mechanism to one of the intermediate modes of motion or by changing the law of motion of the support point along the trajectory, for example, by controlling the rotation of the main link in each cycle of motion. As a result, the maximum value of the required power for the movement, which determines the power of the electric drive, will be reduced to the level of power, required for the movement with a wheel or tracked mover.

5 Conclusion The analysis of the structure of the energy consumption of the walking method of movement is based on the results of dynamic modeling and generalization of experience in the development and testing of a number of experimental units of walking machines. It is shown that via controlling the movement of the support points of the walking mover, it is possible to achieve a significant reduction in energy consumption for the ground resistance to the movement of the walking device. The reduction appears at the cost of recuperation of energy spent on lifting the machine body in each cycle of movement. In difficult movement conditions, for example, on waterlogged or underwater soils, walking machines at low speeds (up to 5–7 km/h) can surpass wheeled and tracked by energy efficiency. At speeds up to 3 km/h, the walking mover can surpass the traditional types of movers by several times. The results of the work may be demanded in the development of walking machines and robots designed for underwater technical works, for new industrial technologies of seabed resources development, to ensure anti-terrorist and technological security of underwater infrastructure and other works. Also, the results of the paper can be used in the development of walking devices for agriculture. Acknowledgements Work partially supported by RFBR and Administration of Volgograd region, research projects No. 19-08-01180, 18-41-340010, 18-48-343005 and the scholarship of president RF-SP5102.2018.1.

Minimization of Energy Costs for Movement Resistance …

847

References 1. Bessonov AP, Umnov NV, Korenovsky VV, Silvestrov EE, Khoborkov SV (2000) Six link mechanism for the legs of walking machines. In: Theory and practice of robots and manipulators. ROMANSY 13: proceedings of the 13-th CISM-IFToMM symposium, international centre for mechanical sciences. Wien, New York, pp 347–354 2. Briskin ES, Chernyshev VV, Maloletov AV et al (2001) On ground and profile practicability of multi-legged walking machines. In: Proceedings of the 4th international conference on climbing and walking robots (CLAWAR 2001), Germany, pp 1005–1012 3. Genta G, Amati N (2004) Mobility on planetary surfaces: may walking machines be a viable alternative? Planet Space Sci 52:31–40 4. Briskin ES et al (2009) The investigation of walking machines with movers on the basis of cycle mechanisms of walking. In: The 2009 IEEE international conference on mechatronics and automation (Changchun, Jilin, 9–12 Aug 2009): conference proceedings. China, pp 3631–3636 5. Comanescu A, Comanescu D, Dugaesescu I, Ungureanu LM (2013) Optimal inverse models for bi-mobile mechanisms of walking robot legs. In: 24th DAAAM International symposium on intelligent manufacturing and automation, pp 417–430 6. Chernyshev VV, Arykantsev VV, Kalinin YaV, Gavrilov AE, Sharonov NG (2015) Development of the walking mover for underwater walking vehicle. In: Proceedings of the 26th international DAAAM symposium “intelligent manufacturing & automation”. Zadar, Croatia, pp 1143–1148 7. Kubota T, Katoh H, Toyokawa T, Nakatani I (2004) Multi-legged robot system for deep space exploration. In: Robotics: trends, principles, and applications—proceedings of the sixth biannual world automation congress, pp 203–208 8. Verichev S, de Jonge L, Boomsma W, Norman R (2014) Deep mining: from exploration to exploitation. In: Minerals of the ocean—7 & deep-sea minerals and mining—4: abstracts of international conference. VNIIOkeangeologia, St. Petersburg, pp 21–24 9. Nautilus Minerals Inc (2018) Nautilus minerals. http://www.nautilusminerals.com. Accessed 12 Nov 2018 10. Komatsu Ltd (2018) Amphibious bulldozer. http://www.komatsu.com/CompanyInfo/views/ pdf/201312/Views_No20_amphibious_bulldozer.pdf. Accessed 12 Nov 2018 11. Hannington M et al (2017) Subsea mining moves closer to shore. Nat Geosci 10:158–1592 12. Isaka K, Tadami N, Fujiwara A, Nakatake T, Sugesawa M, Yamada Y, Yoshida H, Nakamura T (2018) Water jetting excavation and consideration of earth auger shape to reduce drilling torque for seabed robotic explorer. In: 2018 IEEE/ASME international conference on advanced intelligent mechatronics (AIM), pp 887–892 13. Hong S, Kim HW, Choi JS (2002) Transient dynamic analysis of tracked vehicles on extremely soft cohesive soil. In: The 5th ISOPE pacific/asia offshore mechanics symposium, pp 100–107 14. Kim HW, Hong S, Choi JS (2003) Comparative study on tracked vehicle dynamics on soft soil: single-body dynamics vs. multi-body dynamics. In: ISOPE, OMS-2003, Tsukuba, Japan, pp 132–138 15. Chernyshev VV, Arykantsev VV (2016) Dynamic measures of the force of foot breakaway from ground of deepwater walking machines. In: 2016 2nd international conference on industrial engineering, applications and manufacturing. IEEE, 4p 16. Chernyshev VV, Gavrilov AE (2014) Traction properties of walking machines on underwater soils with a low bearing ability. In: Minerals of the ocean—7 & deep-sea minerals and mining—4 (St. Petersburg, Russia, 02–05 June 2014): abstracts of international conference. VNIIOkeangeologia, St. Petersburg, pp 21–24 17. Briskin ES, Zhoga VV, Chernishev VV et al (2010) Walking Machines. In: Elements of theory, experience of elaboration, application, emerging trends in mobile robotics, pp 769–776

848

V. V. Chernyshev et al.

18. Briskin ES, Chernyshev VV, Maloletov AV (2003) Power efficiency and control algorithms of walking machine with cycle propellers. In: Climbing and walking robots: and their supporting technologies, pp 861–870 19. Chernyshev VV, Arykantsev VV, Gavrilov AE, Kalinin YaV, Sharonov NG (2016) Design and underwater tests of subsea walking hexapod MAK-1. In: Proceedings of the ASME 2016 35th international conference on ocean, offshore and arctic engineering OMAE2016. Pusan National University, Busan, 9p 20. Chernyshev VV, Arykantsev VV, Kalinin YaV (2017) Underwater tests of the walking robot MAK-1. In: Human-centric robotics. proceedings of the 20th international conference on CLAWAR 2017. Professional Engineering Publishing, pp 571–578 21. Chernyshev VV, Arykantsev VV, Vershinina IP (2018) Fuzzy control of underwater walking robot during obstacle collision without pre-defined parameters. In: Industrial Engineering, Applications and Manufacturing, IEEE, 5p 22. Ksenzenko AYa, Prysev EA, Pryanichnikov VE, Chernyshev VV (2017) Design the contactless charger and contactless data transfer between underwater robotsatellits and underwater 6-Legged vehicle. In: Annals of DAAAM and proceedings of the international DAAAM symposium, pp 1197–1201 23. Yoo SY, Jun BH, Shim H (2014) Design of static gait algorithm for hexapod subsea walking robot: crabster. Trans Korean Soc Mech Eng A 38(9):989–997 24. Chernyshev VV, Arykantsev VV, Kalinin YaV (2017) Passive foot control in cyclic walking mechanism. In: Industrial engineering, applications and manufacturing: international conference on proceedings. St. Petersburg Polytechnic University of Peter the Great. IEEE (Institute of Electrical and Electronics Engineers). IEEE, 5p. https://doi.org/10.1109/icieam.2017. 8076189

How Different Autonomous Cutter Cooling Methods Affect Machining Performance D. Yu. Dubrov

Abstract This paper covers the effects of various autonomous cooling methods on the cutting performance. Heat phenomena considerably affect the wear of tools induced by finishing structural materials of low thermal conductivity. Aside from being environmentally unfriendly, conventional cooling methods sometimes fail to solve the problem. This is why this paper proposes a cooling method based on first-order phase transitions, which is successfully employed in other industries. Based on the proposed thermal cooling diagram, the researchers have designed a novel indexable cutter that combines the use of heat pipes and consumable media. It has been found out that evaporative cooling requires equipping various elements of a standard indexable cutter with porous inserts as well as special adaptations in the design of the cutting inserts. Besides, the method is not autonomous as such, as the operator must periodically replenish the coolant. According to the studies presented herein, the most preferable option is an integrated cooling system based on first-order phase transitions. This method is shown to make cutters 1.8–2.6 times more durable when finishing 110G13L steel.




Keywords Wear Cooling cutter Porous insert





Cutting tools
Machining
Dry cutting
Indexable

1 Introduction One of the most important challenges of modern mechanical engineering is finding the efficient methods for finishing a variety of hard-to-machine materials that feature low thermal conductivity. The prevalent factor affecting the machining performance is the thermal processes occurring therein and indicated by the cutting temperature. D. Yu. Dubrov (&) Rostov State Transport University, 2, Rostovskogo Stelkovogo Polka Narodnogo Opolcheniya Pl, Rostov-on-Don 344038, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_90

849

850

D. Yu. Dubrov

When turning structural materials of low thermal conductivity, the cutter wedge receives much more heat, which increases the surface and bulk temperatures in the contact spot, speeding up the cutter wear. The conventional methods to reduce such temperatures consist in using cutting fluids (CF), which may be inefficient and inadequate to the modern industrial requirements while also carrying extra costs associated with the design, operation, and recycling of the CF. These may account for 12–17% of the total production costs. CF also causes considerable environmental damage as well as human pathologies like neural depression, intoxication, and cancers. Conventional cooling methods require the operator’s involvement and regular replacement/replenishment of consumables, making it impossible to classify such systems as automated or autonomous [1–11]. This is why the global trend consists of ceasing to use the CF and transitioning to “dry” cutting, where the challenge of cooling down the tool remains unresolved. The results of the studies carried out by the research team behind this paper enable complementing the tool cooling methods (Fig. 1) with additional heat withdrawal technology based on altering the state of matter (open or closed evaporative cooling, use of low-MP substances, etc.) [6, 7, 12]. The experience of applying first-order phase transitions in other industries shows that such approaches are well-applicable to indexable cutters [9, 12–18].

Fig. 1 Classification of cooling methods to improve machining performance (DEC for dry electrostatic cooling; MMS, MQL for minimum use of the cutting fluid)

How Different Autonomous Cutter Cooling Methods …

851

2 Materials and Methods The research team has tested the effects of evaporative cooling by heat pipes vs integrated cooling system (ICS) on the durability of indexable lathe cutters. The experiments involved the use of BK 6 replaceable multifaceted inserts (RMI) selected to have equivalent thermal EMF (TEMF) characteristics. 110G13L steel pieces were machined by a 16K20F3 CNC machine. During the experimental studies of how evaporative cooling and heat pipes would affect the tool durability, the research team designed and made novel RMI prototypes and toolholder elements to enable using standard cutter types (Figs. 2a, b, 3 and 4), as well as a novel anchor plate design (Fig. 3). Experiments to study the effects of heat pipes on the cooling performance involved an upgraded standard tool holder with a tailored anchor plate (Fig. 4). In order to study the effects of an integrated cooling system, the researchers made a special cutter with a cavity filled with a low-MP substance (Fig. 5).

Fig. 2 Cutter design with porous plate: a Cutter with the evaporative cooling of the open type (1-holder; 2-reservoir; 3-pipe; 4-supporting plate; 5-RMCP; 6-porous finger; 7-hold-down tool). b 1-holder; 2-porous substrate; 3-porous finger

852

Fig. 3 RMCP with the porous insert

Fig. 4 Cooling the indexable cutting plate with a heat pipe

Fig. 5 Cutter with an ICS

D. Yu. Dubrov

How Different Autonomous Cutter Cooling Methods …

853

3 Results The proposed methods for cooling indexable cutters are suitable as an alternative to the conventional use of costly and environmentally unfriendly consumable CF. The presented histogram of durability for the three types of first-order phase transition cooling (open evaporative cooling, heat pipes, and ICS) shows that the use of these methods results in 1.8–2.6 times better tool durability, see Fig. 6. However, evaporative cooling has its drawbacks: (a) it requires equipping various elements of a standard indexable cutter with porous inserts as well as special adaptations in the design of the cutting inserts; (b) this method is not autonomous as such, as the operator must periodically replenish the coolant; (c) the method is not cost-effective, as it requires special RMI plates with porous inserts. Closed heat-pipe cooling method is an autonomous and environmentally friendly method that allows using standard pipes in combination with upgraded tool holders. The most advisable method is the ICS, which can be implemented on a wide scale at CNC machines equipped with indexable cutters carrying the removable modules of our design (Patent RU No. 183364). This is apparent in Fig. 7, which shows how this method affects durability as compared to dry cutting for two comparable cases

Fig. 6 How autonomous cooling affects the tool durability as compared to dry cutting (110G13L steel, BK 6 m S = 0.1 mm/rev; t = 0.5 mm: a evaporative cooling (1-evaporating, 2-dry cutting). b Heat pipes (HT) (1-heat pipe cooling, 2-dry cutting). c ICS (1-integrated cooling systems, 2-dry cutting)

854

D. Yu. Dubrov

Fig. 7 Cutter durability improvement attained by different autonomous cooling methods for 110G13L steel, BK 6, S = 0.1 mm/rev; t = 0.5 mm): 1 for ICS, 2 for heat pipes

KT1 ¼

Tkco ; Tdry

KT2 ¼

TTT Tdry

The ICS advantage is obvious, as the coefficient Kt for the curve 1, which equals the RMI durability ratio (ICS vs no cooling, or Kt = Tics/Tdry), is higher than that of the heat pipes.

4 Conclusion • The proposed methods for cooling indexable cutters are proven suitable as an alternative to the conventional use of costly and environmentally unfriendly consumable CF. Analysis of these methods reveals that the ICS is the most promising autonomous cooling technology. • Based on the developed thermal cooling diagram and its performance criteria, the research team has designed and produced a novel indexable cutter with an ICS that combines the use of consumable media and heat pipes, see Patents RU No. 99363, 111787, and 183364. This enables the operator to control the CNC cutting performance by using the temperature curves as the criteria for adjusting the cutting speed. • ICS-based autonomous cooling can be used to improve the performance of other machining methods, e.g., milling (Patent RU No. 183517).

How Different Autonomous Cutter Cooling Methods …

855

References 1. Vasin SA, Vereshchaka AS, Kushner VS (2001) Cutting of materials: thermomechanical approach to the interconnection system during the cutting process: textbook for high schools. MSTU, Moscow 448p 2. Grigoriev SN (2009) Cutting tool durability increasing techniques: textbook for high technical school. Mashinostroenie, Moscow 368p 3. Grigoriev SN, Volosova MA (2007) Coating and tool surface modification: study guide. PC MSTU Stankin, Ianus-К, Moscow 324p 4. Latishev VN, Naumov AG, Asnos TM et al (2003) The influence of the air activated by the corona discharge on the cutting process. Fiz Khim Mekh Tribos 2:14–16 5. Ordinartsev IA, Philippov GV, Shevchenko AN et al (1987) Toolmaker’s: handbook. Mashinostroeniye, Leningrad 846p 6. Pankin AV (1961) Metal processing by cutting. Mashgiz, Moscow 522p 7. Proskuriakov YG, Petrov VN (1962) Cutting tools cooling by of thin-layered mist spray. Mashgiz, Moscow 112p 8. Advancements in metal cutting (1967). Mashinostroeniye, Moscow, 416p 9. Reznikov AN, Reznikov LA (1990) The heat processes in the technological systems. Mashinostroeniye, Moscow 288p 10. Reznikov AN, Zhivogliadov NI (1989) Cutting tools with autonomous cooling system. Stanki Instr 6:40–41 11. http://engjournal.ru/articles/1591/1591.pdf 12. Achmarov RG, Shuster LS, Migranov MS et al (2010) The heatsink technique from the cutting tool contact surfaces. RF patent 2390401, Bul. 15, 27 May 2010 13. Alekseev VA (1975) Electronics cooling with the help of melting substances. Energia, Moscow 88p 14. Alekseev VA, Arefiev VA (1979) Heat pipes for electronics cooling and thermostabilization. Energia, Moscow 128p 15. Alekseev VA, Karabin AE (2011) The new kind of heat accumulator for space electronics cooling. In: Trudy MAI, vol 49. http://www.mai.ru/science/trudy/published.php?ID=28050 16. Dubrov DYu, Dubrov YuS (2014) The use of cutters with two-phase cooling devices in the processing of titanium alloys. In: Second international conference “heat pipes for space application”. Moscow, 15–19 Sept 2014 17. Dubrov DYu (2015) The split cutter efficient life rising during the dry cutting with the help of the vapor-phase cooling. Dissertation, Briyansk, 19p 18. Dubrov DYu, Dubrov YuS Chukarin AN (2017) The ways to increase the cutting tools assemblies durability. Vestnik of P. A. Solovyov Rybinsk State Aviation Technical University 2(41):290–296

Calculation of Thermodynamic Parameters of Geometrically Complex Parts at Abrasive Globoid Gear Machining V. A. Spirin, V. F. Makarov and O. A. Khalturin

Abstract Globoid gear honing made by abrasive tool has such special features as considerable length of the line of contact, profile contact along several toothed surfaces simultaneously, dependence of the stressed-deformed state in the work area on machining conditions, and characteristics of an abrasive layer. All these facts lead to the considerable modification of thermodynamic parameters in the work area. Tool setting accuracy relative to the component part influences greatly the character of these modifications. This paper is devoted to the development of a thermodynamic mathematical model and determination of basic laws of thermodynamic parameters formation in the process of abrasive globoid honing of geometrically complex toothed surfaces. The temperature in the work area, the distribution of temperatures along the line of contact, and change of temperatures depending on the parameters of globoid tool setting could be defined by this model. Numerical experiments have been carried out on the basis of the proposed model. The results have been represented as a diagram.












Keywords Thermal source Temperature field Thermal power density Thermal conductivity Finite difference method Numerical experiment Heat exchange model




1 Introduction Output quality improvement is the indispensable condition for its competitiveness on the domestic and foreign markets. Such improvement is stipulated by permanently increasing demands for the operation reliability. The requirements of accuracy and quality of the surface layer assume the greatest importance in manufacturing geometrically complex screw parts such as power section of the screw downhole motor [1–4]. V. A. Spirin
V. F. Makarov
O. A. Khalturin (&) Perm National Research Polytechnic University, 29, Komsomolsky Prospekt, Perm 614990, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_91

857

858

V. A. Spirin et al.

2 Current Importance Introduction of globoid gear honing process is of current importance due to the advantage of globoid action. It has been developed the technique of characterizing abrasive layer of globoid hone and method of defining conditions for processing by simulation of stressed-deformed state of the rotor and globoid hone working surfaces in the contact region [5–15]. It makes possible to investigate thermal physic parameters of globoid gear honing depending on the characteristics of globoid hone location [16–24]. To carry out this research, it was set a problem of developing mathematical model describing the process of heat exchange at globoid gear honing during the machining of screw downhole motor rotor. The model should provide the calculation of temperature in the supporting points of rotor adjoining the contact zone of globoid hone with the machining area.

3 Problem Statement Basic data for modeling: • Thermal and physical characteristics of the rotor material; • Temperature in the rotor points before the beginning of gear honing; • The length of globoid hone and work surface contact zone. Conceptual model: 1. The object of modeling is the temperature field appearing in rotor in the process of gear honing. Rotor is considered as the rigid body in the end section (Fig. 1). 2. Material of the rigid body is uniform and isotropic. Its thermal and physical characteristics (mass heat capacity, density, and coefficient of thermal conductivity) are temperature-independent. 3. In the process of heat exchange phase transformations do not take place in the body material; deformation induced by temperature change is negligible comparing to the body dimensions; there is a lack of internal sources of heat in the body; 4. Cutting zone is considered as infinitely long in X direction flat source of heat having nonuniformly distributed density of heat generation along X, Y axes; then, stationary mode of gear honing is considered. 5. On the grinding finish out-of-cut zone, there appears the place of the body’s heat exchange with the environment. The density of heat flow corresponds to the Newton–Rikhman law. There is no heat exchange with the ambient in Y = 0 area. 6. Temperature field in the surface layer of the plate is considered to be two-dimensional since the source of heat has the constant density of heat

Calculation of Thermodynamic Parameters of Geometrically Complex … Fig. 1 Rotor end section

859

z q X

Y

generation along Z-axis (assumed originally) as well as due to the assumption of homogeneity and isotropy of the rotor material. 7. Uniform distribution of temperature equal to the ambient temperature takes place up to the moment of globoid hone and work surface contact. On the assumption of the listed hypotheses, conceptual problem statement could be as follows: To define temperature field appearing in the points of the rotor surface layer as result of action of the source of heat with nonuniformly distributed density of heat generation. Density of heat generation by the source of heat is predetermined by: q¼k
r

W=m2

where r—is stress in contact zone, Pa (N/m2); k—reduced velocity, m/s.

4 Research Methodology 4.1

Mathematical Problem Definition

According to the assumed hypotheses 1–4 and 6, the process of heat distribution in the points of the rotor surface layer in X-axis direction is described by the next differential equation: @2h @2h þ 2 ¼ 0; @x2 @y

x0

ð1Þ

860

V. A. Spirin et al.

Entry conditions for the solution of this equation are: hðx; yÞ ¼ h0

ð2Þ

Boundary, when y = 0, conditions are as follows: (

k @h @y ¼ q; 0\x\l k @h @y ¼ 0; x\0 [ x [ l

ð3Þ

Equations (1)–(3) correspond mathematical model of thermal process arising in the process of rotor globoid gear honing. Mathematical problem of thermal process modeling by this method is reduced to the following boundary problem: to find the solution of differential Eq. (1) obeying the initial (2) and boundary (3) conditions.

4.2

Computational Problem Definition

Solution of the stated boundary problem is realized by the finite difference method. Temperature field testing in the points of the rotor surface layer is fulfilled in rectangular area (Fig. 2).

4.3

Digitization and Algebraization of the Problem

Spatial cross-grating in increments of hY along the Y-axis and in increments of hXi along the X-axis by the coordinate curves Yj = j ∙ hY (j = 0, 1, 2, …, m) and Xi = i ∙ hXi (i = 0, 1, 2, …, n).

Fig. 2 Distribution of density of heat generation made by the source of heat

Calculation of Thermodynamic Parameters of Geometrically Complex …

861

Entry conditions: hi;j ¼ h0 ;

i ¼ 0; 1; 2; . . .; n; j ¼ 0; 1; 2; . . .; m

Temperature modification is realized, first of all, in the points of the source of heat contact with rotor surface (Y = 0) and in the points of this surface located behind and in front of the source. Temperature change in these points in Y coordinate makes up:  @h hi;1  hi;0 ¼ ;  @y ði;0Þ hY

i ¼ 0; 1; 2; . . .; n;

Using this formula, we change boundary conditions (3) differential: hi;1  hi;0 ¼ q; hY hi;1  hi;0 k
¼ 0; hY k


0\x\l x\0 [ x [ l

After transformation, it is obtained the following: hi;0 ¼

q
hY þ hi;1 ; k

hi;0 ¼ hi;1 ;

0\x\l; i ¼ 0; 1; 2; . . .; n;

x\0 [ x [ l; i ¼ 0; 1; 2; . . .; n

Use the next formula for the derivatives:  @ 2 h hi þ 1;j  2
hi;j þ hi1;j ¼ @x2 ði;jÞ h2Xi  @ 2 h hi;j þ 1  2
hi;j þ hi;j1 ¼ @y2 ði;jÞ h2Y Substituting these formulas in differential equations of thermal conductivity, we obtain: h2Y
ðhi þ 1;j þ hi1;j Þ þ h2Xi ðhi;j þ 1 þ hi;j1 Þ ; 2
ðh2Y þ h2Xi Þ i ¼ 0; 1; 2; . . .; n; j ¼ 0; 1; 2; . . .; m:

hi;j ¼

862

4.4

V. A. Spirin et al.

Solution of the Problem

Solution of the derived set of equations is made programmatically. hi;0 ¼

q
hY þ hi;1 ; k

hi;0 ¼ hi;1 ;

0\x\l; i ¼ 0; 1; 2; . . .; n;

x\0 [ x [ l; i ¼ 0; 1; 2; . . .; n;

h2
ðhi þ 1;j þ hi1;j Þ þ h2i ðhi;j þ 1 þ hi;j1 Þ ; 2
ðh2 þ h2i Þ i ¼ 0; 1; 2; . . .; n; j ¼ 0; 1; 2; . . .; m:

hi;j ¼

5 Procedure of Numerical Experiment Temperature in the surroundings of the cutting zone with different fluctuations from the assumed parameters of globoid tool setting is considered. The next variant of globoid hone setting relative to the work piece is selected as an example of numerical experiment: DA = −0.3; DZ = −0.3; DG = −0.3. Temperature values for every from 71 points in the section and inside the metal at the depth of 12 layers with each of 0.05 mm depending on the value of globoid hone deflection relative to the work piece are derived from the formulas. Examination of temperatures has shown that the character of their distribution on the section and inside the metal depends on the value of setting parameters deflection (Fig. 3).

Fig. 3 Distribution of temperatures at DA = −0.3; DZ = −0.3; DG = −0.3 design conditions

Calculation of Thermodynamic Parameters of Geometrically Complex …

863

6 Conclusion Considering the cited data, it may be concluded that it is necessary to take into account the deflection of setting parameters and their reasonable combination in order to avoid overheating in the cutting zone since the process of globoid gear honing is prohibitive at the highest temperatures.

References 1. Resnikov AN (ed) (1977) Abrazivnaia i almaznaia obrabotka materialov. Spravochnik (Abrasive and diamond machining of materials. Reference book). Mashinostroenie, Moscow 2. Anuryev VI (1992) Spravochnik konstruktora-mashinostroitelia (Reference book of designer-mechanic) In: 3 vol—7th edn, Mashinostroenie, Moscow 3. Spirin VA (1988) Povyshenie kachestva obrabotki slozhnoprofilʹnykh zubchatykh detaleĭ (Quality improvement of geometrically-complex toothed parts machining). Dissertation, Odessa Polytechnic Institute 4. Tsepkov AV, Spirin VA, Serebrennik YuB (1984) Finishnaia obrabotka rotorov vintovykh zaboĭnykh dvigateleĭ (Finish machining of the rotors in screw downhole motors). In: Abstracts of the zonal STC “Methods of productivity and quality improvement of parts machining at engineering enterprises of the Urals”. Sverdlovsk, pp 74–75 5. Kozlov GV, Sanditov DS (1994) Angarmonicheskie ėffekty i fiziko-mekhanicheskie svoĭstva polimerov (Anharmonic effects and physical-mechanical properties of polymers). Nauka, Novosibirsk 6. Garshin AP, Gropyanov VM, Zaitsev GP, Semyonov SS (2002) Keramika dlia mashinostroeniia (Ceramics for Engineering). Nauchtekhlitizdat, Moscow 7. Richard MK (1982) Introduction to mechanics of composite materials (trans from eng: Beil AI, Zhnud NP). Mir, Moscow 8. Naerman MS, Popov SA (1971) Pretsizionnaia obrabotka detaleĭ almaznymi i abrazivnymi bruskami (Precision machining of parts by the diamond and abrasive bars). Mashinostroenie, Moscow 9. Novosyolov YuK (2012) Dinamika formoobrazovaniia poverkhnosteĭ pri abrazivnoĭ obrabotke (Dynamics of surfaces generation of geometry during abrasive machining). Publishing House SevNTU, Sevastopol 10. Pankov AA, Cokolkin YuV, Tashkinov AA (1988) Novye modeli prognozirovaniia ėffektivnykh svoĭstv kompozitov (New techniques of forecast of effective composite properties). In: Moshev VV (ed) Mechanics of micro nonuniform structures. Ural Branch of the Academy of Sciences of the USSR, Sverdlovsk, pp 4–22 11. Svistrova LA (2003) Fizika polimerov: Ucheb. posobie (Physics of polymers: Manual). Publishing House of Perm State Technical University, Perm 12. Sokolkin YuV, Tashkinov AA (1984) Mekhanika deformirovaniia i razrusheniia strukturno neodnorodnykh tel (Mechanics of deformation and structurally non-uniform bodies). Nauka, Moscow 13. Volkov SD, Stavrov VP (1978) Staticheskaia mekhanika kompozitnykh materialov (Static mechanics of composite materials). Publishing House of Belarusian State University, Minsk 14. Kryzhanovsky VK, Burlov VV, Panimatchenko AD, Kryzhanovskaya YuV (2003) Tekhnicheskie svoĭstva polimernykh materialov: Uch. – sprav (Technical parameters of polymer materials: Manual). “Professija” Publishing House, Saint Petersburg

864

V. A. Spirin et al.

15. Bochkaryov SV (1999) Tekhnologiia proizvodstva polimernykh kompozitnykh materialov i konstruktsiĭ na ikh osnove: Ucheb. posobie (Production technology of polymer composite materials and construction on their basis: manual). Publishing House of Perm State Technical University, Perm 16. Reznikov AN, Reznikov LA (1990) Teplovye protsessy v tekhnologicheskikh sistemakh: Uchebnik dlia vuzov (Thermal processes in technological systems: textbook for universities). Mashinostroenie, Moscow 17. Reznikov AN (1981) Teplofizika protsessov mekhanicheskoĭ obrabotki (Thermal physics of machining processes). Mashinostroenie, Moscow 18. Sipailov VA (1978) Teplovye protsessy pri shlifovanii i upravlenie kachestvom poverkhnosti (Thermal processes in grinding and surface quality control). Mashinostroenie, Moscow 19. Tsaplin AI (2008) Teplofizika v metallurgii: ucheb. posobie (Thermal physics in metallurgy: manual). Publishing House of Perm State Technical University, Perm 20. Tashkinov AA, Vildeman VE, Gazizov RYa, Ivanov SG (compilers) (1991) Ėffektivnye termouprugie svoĭstva anizotropnykh kompozitov: Metod. ukazaniia dlia samost. raboty (Effective thermoelastic properties of anisotropic composites: instructional guidelines for independent work, vol I: Theory and design formulae). Perm Polytechnic Institute, Perm 21. Huo WG et al (2008) The finite element analysis of surface temperature on dry belt grinding for titanium alloys. Adv Mater Res 53–54:219–224. https://doi.org/10.4028/www.scientific. net/AMR.53-54.219 22. Zhang X, Cabaravdic M, Kneupner K, Kuhlenkoetter B (2004) Real-time simulation of robot controlled belt grinding processes of sculptured surfaces. Int J Adv Rob Syst 1(2):109–114 23. Li CH, Li JY, Wang S, Zhang Q (2013) Modeling and numerical simulation of the grinding temperature field with nanoparticle jet of MQL. Adv Mech Eng 5. https://doi.org/10.1155/ 2013/986984 24. Wang X, Yu T, Sun X et al (2016) Study of 3D grinding temperature field based on finite difference method: considering machining parameters and energy partition. Int J Adv Manuf Technol 84:915. https://doi.org/10.1007/s00170-015-7757-z

Providing of Surfaces’ Geometry at the Design Stage of Profile Milling Operation of Off-Grade Workpiece A. A. Fomin, V. G. Gusev and N. F. Timerbaev

Abstract In the article, the technological operation of a wood profile milling and a large waste (off-grade workpieces), formed as a result of logs machining, is observed. Forecasting of geometrical characteristics of profile products from the off-grade workpieces on the basis of results of design calculation allows to divide out the volume of experimental works on maintenance of demanded geometry to stages of the development of products’ manufacture. The development of an algorithm for calculating these output parameters required the mathematical support of design calculations in the form of analytical adequate models of the cut-off layer; the main, radial components of the cutting force; the profile milling power and the geometric errors of the machined surfaces. Models determine the regularity of the change in the power characteristics of profile milling, the geometry of the milled surface as a function of the elements of the cutting mode, the characteristics of the shaped cutting tool, the peripheral segment resulting from the cutting of the logs. The developed algorithm allows calculating the parameters of profile milling of stem wood and peripheral segments (off-grade workpieces) with curved surfaces and knots. Iterative calculations of profile milling parameters, performed using the developed algorithm, allow determining the processing conditions, under which the required accuracy of the treated surfaces is provided at the maximum possible productivity of the profile milling.







Keywords Profile milling Off-grade workpiece Technological operation Geometrical accuracy Milling cutter Cutting force










A. A. Fomin
V. G. Gusev (&) Institute of Engineering and Automobile Transport, Vladimir State University, 87, Gorky Street, 600000 Vladimir, Russia e-mail: [email protected] N. F. Timerbaev Kazan State Power Engineering University, 51, Krasnoselskaya Street, Kazan 420066, Russia © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_92

865

866

A. A. Fomin et al.

1 Introduction Our planet has rich forest resources, which are used to produce wood products by processing wood and its waste [1–6], including mechanical processing by milling [7–12]. The dynamics of an overmeasure removal [8], kinematics of formation of the processed surface [9], its microgeometry, waviness, formation of a chip [10– 12], influence of milling conditions on a quality of the processed surfaces [13], and also cutting force [14], spindle knot rigidity at a milling of off-grade workpieces are investigated [15]. The thickness of a chip [16] and stability of milling operation [17] are analyzed, and optimum parameters of milling [18] are defined. In the papers [19–21], a profile, the surfaces’ microstructure, and the quality of control of functional properties of created objects are considered. The influence of cutting tools, wood’s properties, cut methods, of a technological equipment, etc., on the efficiency of cylindrical and angular milling of wood are studied in sufficient detail, but profile milling of wood off-grade workpieces, a dynamics of the forming surfaces were not investigated, as a result of which the productivity and accuracy, the cost of produced profile products and other efficiency indicators of machining processes depend on subjective factors. There are expensive machining systems, which provide complex opening of logs (at the same time and of peripheral segments into small elements). These small elements of peripheral segments stick together, dry and subject to machining, what leads to a deterioration of the ecological situation (because of application of glues), to an elongation of a technological process and to considerable rise in price of a finished product. In this connection, the creation of low-cost equipment and of non-polluting processes of profile milling of off-grade workpieces is very important. On the basis of this equipment, a production of qualitative profile products of type “block house” is real. Now, the pilot equipment with a numerical control and an automatic control system by the speed of the working feed was developed. Equipment allows to carry out a profile milling of off-grade workpieces. High efficiency of use of this equipment can be attained on a base of authentic data of cutting regimes and of products’ parameters; therefore, the scientific base for machining processes of off-grade workpieces is very important.

2 The Technological Characteristics of Off-Grade Workpieces A long curvilinear concave and convex surfaces of products make by form milling cutters from edged board and from the large waste of wood yet do not make. At logs machining in size H (Fig. 1a) are formed two peripheral segments S1 , which have one processed plane, other surfaces are corticated.

Providing of Surfaces’ Geometry at the Design Stage of Profile …

867

Fig. 1 a Scheme of a log sawing up; b a segment with knots; c milling of a segment with an increment overmeasure

At machining in size E in the log, two peripheral segments S2 and a bar are formed. Segments S2 also have a bark 1 and peripheral fibres 2, and the pith 3 is present in the bar center. The most adverse cutting conditions appear at the milling of accrete knots 1, 2, 3 (Fig. 1b) and segments 4 with an increasing overmeasure hi (Fig. 1c). Hardness of spruce knots is in 2.3–3.7 times more, and hardness of pines and larches is in 2.2–3.5 more, and birches hardness is in 1.3–1.5, and aspens is in 1.25–1.50 times more than deckman wood’s hardness. Arrows Dr and Ds show the directions of working movements of cutting tool and workpiece.

3 Characteristics of Initial Data The initial data for calculation of parameters of the profile milling process of stem wood and peripheral segments with continuously changing overmeasure are presented in the first operator of algorithm (Fig. 2), which includes four connected to each other blocks, performed as the rectangles 1–4. They are separated between each other by straight dashed lines. There are entry data about sizes and other characteristics of the wood product in the operator 1 of the first block: the width Bo , the thickness ho ; species of wood Por (1–7) seven names total, initial moisture of wood W(%), knots Suc , modules of elasticity and shear Ex , G; values of geometric quality indicators of the treated surface, which must be observed in profile milling; geometrical quality indicators of the processed surface, which have been to get during the profile milling (a roughness Ra , waviness HB , a deviation in a longitudinal dpr , transverse dp sections of a product), and a power Nd of the drive electric motor of the cutter main motion. In the operator 2 of the first block, there are the width Bn , the thickness hn of an off-grade workpiece, the sizes of a milling cutter (height B, minimum Rmin , the maximum Rmax radii, and the profile radius Rrk ), the main vector Dst of unbalance, the number of teeth z, the spindle speed n, the cutting depth t, the longitudinal feed max rates vmin s , vs , and the cutting speed v. In the operators 3 and 5 of the second block, a wood species is assigned, with each current value i, corresponding to one name of the breed, for example, pine, spruce, oak, birch, ash.

868

A. A. Fomin et al.

Fig. 2 Algorithm block scheme for calculating of the milling parameters of the off-grade workpieces

4 Calculation of the Parameters of Cut Layer In operators 4 and 6 (Fig. 2), the thickness amax , the area Fsl , and the volume Vsl of the cut layer are calculated, respectively, by the formulas (1)–(3): amax ¼ e1  e2 ; where e1 ¼ Rrk ð1  cosðarcsinðzi =Rrk ÞÞÞ þ Rmin ;

ð1Þ

Providing of Surfaces’ Geometry at the Design Stage of Profile …

869

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi  0;5   6:28vs 2 12:56vs 2e1
e3  e23 2 ; e3 ¼ t þ vs s
tgbsb ; e2 ¼ e1 þ  x
Z x
Z Fsl ¼ Sz ½e3 þ Rrk cosðarcsinðzi =Rrk ÞÞ sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi     2    zi 16:7pvs 2 þ ðRmin þ Rrk Þ  þ Rmin þ Rrk  Rrk cos arcsin Rrk x
Z ð2Þ Vsl ¼ 2Fsl Rrk arcsinðB=2Rrk Þ

ð3Þ

where Rrk , Rmin —respectively, the radius of the milling profile and its minimum radius; zi —the distance of the cross section of the cutter from its geometric center; vs — working feed; x, z—angular velocity and number of teeth of the milling cutter, respectively; t—the depth of cut, determined by the milling mode; s—treatment time: bsb —average statistical angle, characterizing the amount of run; B—cutter height.

5 Calculation of External Workload on the Technological System The cutting power of the profile milling of the workpiece with the positive increment of the overmeasure is determined in operator 6 by formula: Pr ¼ 0:033 Kt
apop ðt þ vs s
tgbsb Þvs Rrk ;

arcsinðB=2Rrk Þ

ð4Þ

where Kt —tabular value of the specific work of cutting; apop ¼ ap
aw
ar
ad
ay —overall correction factor; ap ; aw ; ar ; ad ; ay —correction factor for the species, wood moisture, dull blades, cutting angle, and cutting speed, respectively. In operators 7 and 8 of the third block, the conformity of the cutting power to the main drive motor power of the main movement of the milling cutter is checked and if the condition Pr  0:85Pd is not met, the working feed vs decreases by one step. In operators 9 and 10, the calculation of the main Pz the radial Py components of the cutting force, as well as the unbalanced centrifugal force Q, due to the action of the main vector of unbalances of the shaped cutter, is made using the formulas: Pz ¼ 1020
60Pr =v

ð5Þ

Py ¼ m
Pz

ð6Þ

Q ¼ Dst x2 ¼ mi x2 qi

ð7Þ

870

A. A. Fomin et al.

where m—is the coefficient, that is a function of the initial cutting conditions: the sharpness of the cutter blade, the average thickness of the cut layer and the cutting angle, the values of m are given in the woodworking reference books; Dst —the main vector of imbalances of the spindle unit with the installed cutter, caused by inaccurate installation and fastening of cutting inserts in the cutter body, as well as the error of setting the milling cutter on the spindle; x—the angular velocity of the mill; n—rotational speed of the milling cutter, mi —unbalanced mass; qi —the radius of the center of the unbalanced mass.

6 Calculation of Geometrical Errors of the Treated Surfaces In the operator 11, the maximum elastic displacements of the workpiece are calculated under the action of the work load by the formula: ymax

   l3B Py þ DPy cosð0:5ðe1 þ e2 ÞÞ  ðPZ þ DPZ Þ sinð0:5ðe1 þ e2 ÞÞ ¼ 48Ex I

ð8Þ

where lB —the distance between adjacent rolls, which communicate the workpiece with the speed of the working feed; DPy , DPz —respectively, the variable component of the forces Py and Pz , due to the runout of the workpiece; (e1 þ e2 )—the exit angle; Ex —modulus of elasticity of wood; I—the moment of inertia of the cross section of the workpiece. The error of the machined surface, caused by the kinematics of the profile milling and the parameters of the shaped cutter, is determined by the formula: ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ffi    s    2  zi zi pvs 2  Rrk cos arcsin þ Rmin þ Rrk  dk ¼ Rrk cos arcsin Rrk Rrk x
z þ Rmin þ Rrk :

ð9Þ Maximum geometrical error, caused by the action of cutting force, equals to: dd1 ¼



ðJB þ Jz ÞðP2y þ P2z Þ0:5
cos 0:5p  arctgðPy =Pz Þ  0:5ðe1 þ e2 Þ J6 Jz

;

where JB , Jz —respectively, the rigidity of the spindle knot and workpiece.

ð10Þ

Providing of Surfaces’ Geometry at the Design Stage of Profile …

871

The maximum error, caused by unbalance of milling cutter, is: dd2 ¼ Dst x2 =JBmin

ð11Þ

where JBmin —the minimum stiffness of the spindle knot with the installed milling cutter. In operator 13, the total geometric error of the machined surface is calculated: dR ¼ dk þ dd1 þ dd2 :

ð12Þ

7 Ensuring the Required Geometric Accuracy of the Machined Surfaces and the Maximum Productivity of Profile Milling In operator 14, the total geometric error dR of the machined surface is compared with the error dch , allowed by the product drawing. If dR  dch , then, in order to use the equipment more efficiently, operator 15 checks the condition dR  0:7dch , under which the operator 16 increases the work feed of the workpiece by one step and starts a recurring cycle of calculations. If the condition dR  0:7dch is not observed, the operator 15 issues a command to print 17 the most important output parameters of the milling process with a run out: the milling power, the main and radial components of the cutting force, and also the geometrical characteristics of the accuracy of the machined surfaces. The algorithm provides for the steps to ensure the required values of the quality indicators of the treated surfaces at the maximum possible working feed rate and, consequently, the maximum possible productivity of the profile milling. Results of researches were used in the development of a woodworking machine with software control of the PFP-100 model, technical and technological solutions, used in the machine tool, are protected by patents of the Russian Federation [22, 23]. The real machine model was certified, presented at the 6th Moscow International Salon of Innovations and Investments, where it was awarded a diploma and a silver medal.

8 Conclusion 1. In woodworking enterprises, as a result of cutting logs, a large number of off-grade workpieces (peripheral segments) are formed with a constantly changing one-way overmeasure. Due to the lack of developed base surfaces of peripheral segments, the dispersion of hardness and dimensions, and also the curvilinear contours of the knotty surface, covered by bark, mechanical processing of the segments causes serious difficulties, therefore to get the

872

A. A. Fomin et al.

qualitative products it is necessary to define optimum conditions of machining at a designing of a technological operation. 2. The algorithm and mathematical models of a cut-off layer, power characteristics, and geometrical parameters of accuracy of the processed surfaces, on which basis iterative calculations of parameters of profile milling are executed, forecasting of geometrical accuracy of products from off-grade workpieces are developed and the conditions of machining, providing demanded geometry of profile surfaces, are defined. 3. Results of researches were used in the development of a woodworking machine tool with software control of the PFP-100 model for profile milling of peripheral segments and deckman wood. The real milling machine was certified, presented at the 6th Moscow International Salon of Innovations and Investments, where it was awarded a diploma and a silver medal.

References 1. Safin RG (2016) A mathematical model of thermal decomposition of wood in conditions of fluidized bed. J Acta Facultatis Xylologiae Zvolen res Publica Slovaca 58(2):141–148. https:// doi.org/10.17423/afx.2016.58.2.15 2. Safin RG (2017) Technology of wood waste processing to obtain construction material. J Solid State Phenomena 265:245–249. https://doi.org/10.4028/www.scientific.net/SSP.265. 245 3. Tuntsev DV (2016) The mathematical model of fast pyrolysis of wood waste. In: Proceedings of 2015, J Meacs 2015. Art no. 7414929. https://doi.org/10.1109/meacs.2015 4. Saldaev VA (2016) Equipment for the production of wood-polymeric thermal insulation materials. IOP Conf Ser Mater Sci Eng 142(1):012097. https://doi.org/10.1088/1757-899X/ 142/1/012097 5. Sadrtdinov AR (2016) The development of equipment for the disposal of solid organic waste and optimization of its operation. IOP Conf Ser Mater Sci Eng 142(1):012095 6. Sadrtdinov AR (2016) The mathematical description of the gasification process of woody biomass in installations with a plasma heat source for producing synthesis gas. IOP Conf Ser Mater Sci Eng 124(1):012092. https://doi.org/10.1088/1757-899X/124/1/012092 7. Konovalov S, Chen X, Sarychev V et al (2017) Mathematical modeling of the concentrated energy flow effect on metallic materials. Metals 7(1):4 8. Gusev VG, Fomin AA, Sadrtdinov AR et al (2017) Dynamics of overmeasure removal in profile milling process by shaped tool. J Procedia Eng 206:279–285 9. Fomin AA (2013) Kinematics of surface formation in milling. J Russ Eng Res 33(11):660– 662. https://doi.org/10.3103/S1068798X13110099 10. Fomin AA (2017) Microgeometry of surfaces after profile milling with the use of automatic cutting control system. In: Proceedings of 2017 international conference on industrial engineering, applications and manufacturing, ICIEAM 2017, Art. no. 8076117. https://doi. org/10.1109/icieam.2017.8076117 11. Gusev VG, Fomin AA et al (2017) Multidimensional model of surface waviness treated by shaping cutter. J Procedia Eng 206:286–292 12. Sharkov OV, Koryagin SI, Velikanov NL (2018) Shaping cutter original profile for fine-module ratchet teeth cutting. IOP Conf Ser Mater Sci Eng 327:042102

Providing of Surfaces’ Geometry at the Design Stage of Profile …

873

13. Nekrasov RY, Tempel YA, Starikov AI, Proskuryakov NA (2018) Fuzzy controllers in the adaptive control system of a CNC lathe. Russ Eng Res 38(3):220–222. https://doi.org/10. 3103/S1068798X18030188 14. Bardovsky A, Gerasimova A, Aydunbekov A (2018) The principles of the milling equipment improvement MATEC web of conferences, 224. https://doi.org/10.1051/matecconf/ 201822401019 15. Fomin AA, Gusev VG (2013) Spindle rigidity in milling blanks with nonuniform properties. J Russ Eng Res 33(11):646–648. https://doi.org/10.3103/S1068798X13110087 16. Fomin AA et al (2017) Determining undeformed chip thickness models in milling and its verification during wood processing. J Solid State Phenom 265:598–605 17. Yemelyanov V, Tochilkina T, Vasilieva E, Nedelkin A, Shved E (2018) Computer diagnostics of the torpedo ladle cars. AIP Conf Proc 2034:020008. https://doi.org/10.1063/1. 5067351 18. Rezchikov AF, Kochetkov AV, Zakharov OV (2017) Mathematical models for estimating the degree of influence of major factors on performance and accuracy of coordinate measuring machines. MATEC Web Conf 129:01054 19. Sharkov OV, Koryagin SI, Velikanov NL (2016) Design models for shaping of tooth profile of external fine-module ratchet teeth. IOP Conf Ser Mater Sci Eng 124:012165 20. Gromov VE, Kormyshev VE, Glezer AM et al (2018) Microstructure and wear properties of Hardox 450 steel surface modified by Fe–C–Cr–Nb–W powder wire surfacing and electron beam treatment. IOP Conf Ser Mater Sci Eng 411(1) 21. Grechnikov FV, Rezchikov AF, Zakharov OV (2018) Iterative method of adjusting the radius of the spherical probe of mobile coordinate-measuring machines when monitoring a rotation surface. Meas Tech 61:347–352 22. Fomin AA, Gusev VG (2010) The rig for slab preprocessing. RF Patent 95, 589, 7 July 2010 23. Fomin AA, Gusev VG (2012) A way of a roughing-out of a slab. RF Patent 2, 443, 547, 27 Feb 2012

The Method of the Combined Flat Peripheral Grinding A. V. Morozov and V. G. Gusev

Abstract The article deals with the method of combined grinding of materials, providing simultaneously preliminary and final surfaces treatment on one machine tool. The analysis of existing grinding methods showed that depending on volume of details exhaustion, a preliminary and final grinding can be executed at one or two technological operations. In the first case, the preliminary and final grinding is carried out on the same workplace with the replacement of coarse-grained with a fine-grained grinding wheel. In the second case, after the preliminary operation, the workpiece is transported to another workplace for final processing. In both variants, a large amount of auxiliary time is required, which leads to a decrease in the productivity of the technological operation. The noted deficiencies are eliminated by using the method of combined grinding. Possible combined peripheral grinding schemes were developed and analyzed, elastic displacements of the spindle knot with a tool under the action of cutting force were determined, on the basis of which a scheme was chosen, that ensures rational arrangement of coarse-grained and fine-grained wheels on the machine spindle. The technological possibilities of the combined grinding were determined, The comparative experimental data of the microgeometry of surfaces, processed by proposed and traditional methods, are given. Combined grinding allows reducing the machine and auxiliary time for the performance of technological operation and to increase the processing productivity with the same microgeometry of the surface, typical for traditional grinding.







Keywords Combined grinding Workpiece Elastic displacements knot Fine-grained wheel Coarse-grained wheel








Spindle

A. V. Morozov
V. G. Gusev (&) Vladimir State University, 87, Gorky Street, Vladimir 600000, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_93

875

876

A. V. Morozov and V. G. Gusev

1 Introduction Grinding among other processes of machining plays a most important role in the formation of geometric accuracy and quality of the treated surfaces [1]. The scientific works directed on raise of efficiency of grinding processes by a discretization of the cutting surface [2–6] are known. Researches of dynamics of processes [7–9]; of self-organizing of discrete grinding wheels [10]: modeling of the processed surfaces [11, 12] and trajectory of abrasive grain [13]; of combined grinding wheels [14]; of pulsing pressing and highly temperature caking of wheels [15]; of cooling technological means [16, 17]; of highly-porous grinding wheels [18] and of a loading on the tool [19] are executed. Existing grinding processes are characterized by differentiation of preliminary and final grinding of workpieces. Processing is carried out on two grinding machine tools or on one machine tool with the obligatory replacement of coarse-grained with fine-grained wheel. This leads to a significant increase of the main and auxiliary time for doing a technological operation, consequently, to a decrease of productivity and to an increase of the technological cost of the product. Elimination of these shortcomings of existing grinding processes is possible by developing and using so-called combined grinding. In this direction, separate studies were carried out, in particular, of the segmental tools for preliminary and final grinding [4]. These grinding tools solve the problem of reducing heat in the cutting zone, but do not consider the issues of providing small roughness, waviness of the surface and increasing the processing productivity.

2 The First Scheme of Combined Grinding The proposed tool for combined grinding consists of a wheel 1 (Fig. 1) for preliminary grinding, in which abrasive grains 2 have a sizes F30–F36 or more, and of a grinding wheel 3 for final grinding with the abrasive grains F90–F120. Coarse grinding wheel 1 removes the overmeasure zpr for pretreatment, the fine-grained wheel 3—the overmeasure zo for the final treatment and serves to ensure the required quality of the machined surface layer. The wheel for preliminary grinding has an outer diameter d pr , which is smaller than the diameter d o of the wheel for final grinding. Between the grinding wheels 1 and 3, there is a gasket 4, which eliminates the contact of the wheels and forms a gap d, which allows the diamond pencil to be positioned at correcting wheels with different diameters. Coarse-grained wheel 1 performs the main work to remove the overmeasure, at the same time, wheel 3 forms the required quality of the surface layer. The wheel 1 is located closer to the operator 8, serving the machine tool, and the wheel 3—between the wheel 1 and the front support 5 of the spindle 6. The diameters of both grinding wheels are related by the equation:

The Method of the Combined Flat Peripheral Grinding

877

Fig. 1 First scheme of combined grinding

do ¼ dpr þ 2ðRmax þ TÞ;

ð1Þ

where do ; dpr —respectively diameter of the grinding wheel for final and preliminary grinding: Rmax —the maximum of the surface roughness after preliminary grinding; T—defective layer, formed in detail as a result of the force and thermal action of preliminary grinding.

3 The Calculation of the Elastic Displacements of the Spindle Axis at the First Scheme The grinding process is performed at rotation of the wheels 1 and 3 in direction of arrow Dr, longitudinal Dspr and cross Dsp feed of workpiece in direction from the operator 8 (Fig. 1) and is finished with a fine-grain wheel and a cross-feed in the direction to the operator with removal of the allowance, determined by the formula: zf ¼

PR Po þ Ppr ¼ ; j j

ð2Þ

where PR —a total cutting force; Po ; Ppr —respectively the cutting force, caused by the wheels of final and preliminary grinding; j—the rigidity of the spindle knot in a direction, perpendicular to the machined surface. Allowance zpr several times exceeds the allowance zo , therefore the cutting force Ppr much more of the cutting force Po (Fig. 2). Cutting force Po , created by the wheel 3 for final grinding, is located from the front spindle support 5 at a distance: lo ¼ 0:5Bo þ f ;

ð3Þ

where Bo—a height of wheel 3; f—a distance from the right-end face of the wheel 3 to the front support 5 of the spindle 6 (Fig. 1).

878

A. V. Morozov and V. G. Gusev

Fig. 2 Elastic displacements of spindle axis at the first scheme

Cutting force Ppr is located from the front support 5 in the distance: lpr ¼ f þ Bo þ d þ 0:5Bpr ;

ð4Þ

where Bpr —a height of wheel 1 for preliminary grinding, d—a thickness of the gasket 4. It agrees (2) the total cutting force is equal: PR ¼ Po þ Ppr :

ð5Þ

The system of concentrated forces Po ; Ppr with known coordinates of application points can be replaced by an equivalent system, in which the total cutting force PR acts. The condition of equivalence of the two considered systems is the equality of the bending moments, created by the concentrated forces relative to the origin of the rectangular coordinate system XOY (Fig. 2). From the condition of equivalence of two systems, the coordinate of the application point of force PR is determined by the formula:
ð0:5Bo þ f ÞPo þ Bo þ f þ 0:5Bpr þ d Ppr l1 ¼ Po þ Ppr

ð6Þ

Maximum elastic displacements of the spindle axis under the action of the total cutting force PR in the point of its application x = l1 are equal: YM1 ¼

PR l31 ; 3EI

ð7Þ

where E = 2.1  105 MPa—is the modulus of elasticity of steel, from which the pd 4

spindle 6 is made; I ¼ 64pr ; dpr —the moment of inertia of the cross section and the reduced diameter of the spindle console. Angle of turn of the spindle cross section, located at a distance x = l1 (Fig. 2), is:

The Method of the Combined Flat Peripheral Grinding

u1 ¼

PR l21 2EI

879

ð8Þ

Elastic displacements of the spindle axis under the action of force PR on length 0  x  l1 are described by a curve OM1 of third order, then from the point M1 — by straight line segment M1 N1 ; tangent at a point M1 to the curve OM1 . Maximum elastic displacements of the left end face of the wheel 1 (Fig. 1) under the action of force PR at the point with the coordinate x ¼ ðf þ Bo þ d þ Bpr Þ are: Ymax1 ¼


PR l31 þ f þ Bo þ d þ Bpr  l1 arctgðu1 Þ 3EI

ð9Þ

4 The Second Scheme of Combined Grinding Combined grinding can be also realized in the second scheme, when the fine-grained wheel 1 is located at the maximum distance from the spindle front support 5, and the coarse-grained wheel 3—at the minimum. In this case, the cutting forces Po and Ppr change they places, but the values of these forces do not change. Taking into account the foregoing, the second scheme of combined grinding has the form, shown in Fig. 3. The wheel 1 is located from the front support 5 at a distance f, and the wheel 3 at a distance ðf þ Bpr þ dÞ. As with the first combined grinding scheme, the wheel 1 removes the main allowance Zpr, and wheel 3—allowance Zo. The frequency of rotation of the spindle with the tool, the magnitude of the longitudinal, and transverse feeding of the workpiece in the second processing scheme (Fig. 3) remain the same, as in the first scheme (Fig. 1). Fig. 3 Second scheme of combined grinding

880

A. V. Morozov and V. G. Gusev

Fig. 4 Elastic displacements of spindle knot at the second scheme

In the second scheme, the cross-feed of the workpiece is directed to the operator, controlling the machine tool, but in the first scheme it is from the operator. Distance of an application point of the force PR from the point O in the second scheme (Fig. 4) is:
ð0:5Bo þ f ÞPpr þ Bo þ f þ 0:5Bpr þ d Po l2 ¼ : Po þ Ppr

ð10Þ

5 The Calculation of the Elastic Displacements of the Spindle Axis at the Second Scheme Elastic displacements of spindle under action of force PR in the point x = l2 (Fig. 4) are determined by formula: YM2 ¼

PR l32 : 3EI

ð11Þ

Angle of turn of the spindle cross section, located at a distance x = l2 from the point O, under action of the total cutting force PR is determined by formula: u2 ¼

PR l22 : 2EI

ð12Þ

Elastic displacements of the spindle axis under the action of force PR on length 0  x  l2 are described by a curve OM2 of third order, then from the point M2 — by straight line segment M2 N2 ; tangent at a point M2 to curve OM2 . Maximum elastic movements of the left
end face of wheel 3 (Fig. 3) under the action of force PR at x ¼ f þ Bo þ d þ Bpr , that is in the point N2 , are equal:

The Method of the Combined Flat Peripheral Grinding

Ymax2 ¼


PR
l32 þ f þ Bo þ d þ Bpr  l2 arctgðu2 Þ 3EI

881

ð13Þ

Let us analyze the obtained formulas from the position of geometric accuracy of the surfaces, machined according to the first and the second scheme.

6 The Analysis of the Elastic Displacements of the Spindle Axis in the Combined Grinding Analyses (6) and (10) show, that in (6) before a relatively small force Po there is
a multiplier ð0:5Bo þ f Þ, which is less, than the multiplier Bo þ f þ 0:5Bpr þ d ; standing in front of the force Ppr , considerably exceeding the force Po . This fact
indicates that the member Bo þ f þ 0:5Bpr þ d Ppr in (6) much more, than ð0:5Bo þ f ÞPo . In (10), the forces Po и Ppr swapped places, while in front of a large force Ppr , there is a multiplier ð0:5Bo þ f Þ, which is less, than the multiplier
Bo þ f þ 0:5Bpr þ d , standing in front of a small force Po . As a result, the numerator in (6) is larger, than the numerator in (10), and consequently, the coordinate l1 of total force PR at the first scheme more coordinate l2 at the second scheme. On the basis of it follows that at the first scheme the total force creates the bigger bending moment in a front support 5 of spindle 6 and the bigger elastic movements of a spindle with the tool, than at the second scheme. This conclusion is also confirmed by inequality YM1 [ YM2 . Dividing (7) by (11), we obtain the value of the coefficient, indicating how many times the elastic displacements in the point of application of force PR at the first scheme more in comparison with the second scheme: Ky ¼

YM1 ¼ YM2

 3 l1 : l2

ð14Þ

For grinding machine tool 3M71 the ratio l1 =l2  1.5–2.0 and the coefficient Ky  3.4–8.0.

7 The Method of Practical Realization of the Combined Grinding Scheme The proposed method of combined grinding [20] was realized by using the second scheme (Fig. 3). Set and fix the grinding wheels 1, 3 on the spindle 6. The workpiece 7 is mounted on the magnetic table of the flat grinding machine so that it

882

A. V. Morozov and V. G. Gusev

is located behind the coarse-grained wheel 1 (see from the operator). Include rotation of the spindle 6 with the grinding wheels 1, 3 in the direction of arrow Dr and longitudinal feed in the direction of the arrow Dspr . Fail a coarse-grinding wheel 1 to the workpiece 7 to the first spark, and then the wheel haul down on magnitude of an allowances zpr , to be removed at pre-grinding. After adjusting the coarsegrained wheel 1 for size include the cross feed in the direction of the arrow Dsp . At high requirements to a surface roughness in addition carry out grinding by fine-grained wheel at the cross-feed in a direction from the operator. The coarse-grained wheel 1 for preliminary grinding does not touch the processed surface and does not worsen the small roughness, formed by fine-grained wheel 3. In the known grinding processes, this cannot be achieved. Experiments have been done on the machine tool 3M71, processed the plates from a steel 40X by the combined grinding tool, and also separately by coarse-grained and fine-grained wheel of the above-stated characteristic of an abrasive material. The cutting speed for all was 35 m/s, the longitudinal feed of the workpiece—15 m/min, and the cross feed—4 mm/double course of table. It is established, that the roughness of a surface after the combined grinding is 2.5 times less in comparison with traditional processing by a coarse-grained wheel, and is 5–7% more, when processing by a fine-grained wheel. At machining by the new tool with use of only fine-grained wheel, the roughness of surfaces is equal to compared alternatives. Thus, from the position of providing a geometric accuracy of ground surfaces, it is necessary to use a second scheme of combined grinding, which provides a smaller bending load on the spindle and smaller elastic displacements of the grinding tool.

8 Conclusion 1. A method of combined grinding by a tool, containing coarse-grained and fine-grained abrasive wheels, mounted on a spindle, was worked out. The tool allows to carry out simultaneously preliminary and final processing on one machine tool. 2. Possible schemes of combined grinding were worked out and the elastic displacements of the spindle under the influence of the cutting force were analyzed, and on this basis, the second combined grinding scheme to practical realization was chosen. 3. The combined grinding tool has been tested at a flat grinding of workpieces and allows to reduce the main and axillary time for performing of technological operations, what leads to increase of machining productivity.

The Method of the Combined Flat Peripheral Grinding

883

References 1. 2. 3. 4. 5. 6. 7.

8. 9. 10. 11.

12. 13.

14. 15. 16. 17. 18. 19. 20.

Suslov A (2000) Quality of a blanket of machine components. Mech Eng, Moscow Gusev V, Morozov V (2007) Flat discrete grinding technology. RF, Vladimir Gusev V, Morozov A (2012) Flat peripheral grinding with discrete wheels. RF, Yoshkar-Ola Morozov A, Gusev V (2016) Discrete end face grinding, Moscow Morozov AV et al (2016) Deterioration of the diamond tool at editing of discrete grinding wheels, J Sci. Eng Indus Bul 3:59–64 Gusev VG, Morozov AV et al (2017) Discretization technology of abrasive wheels operating surfaces with laser and hydro-abrasive jet. J Sci Sci Intensive Tech Mech Eng 9(73):20–27 Fomin AA 2017 Microgeometry of surfaces after profile milling with the use of automatic cutting control system. In: Proceedings of 2017 international conference on industrial engineering, applications and manufacturing, ICIEAM 2017, Art no 8076117. https://doi.org/ 10.1109/icieam.2017.8076117 Fomin AA (2017) Limiting product surface and its use in profile milling design operations. Solid State Phenom 265:672–678. https://doi.org/10.4028/www.scientific.net/SSP.265.672 Gusev VG, Morozov AV, Shvagirev PS et al (2009) Evaluating discrete wheels and their influence on grinding dynamics. J Sci Russ Eng Res 29(8):835–837 Salov B, Kravchenko M (2001) Principles of self-organising of grinding wheels discretization. RF, Samara Novoselov YuK, Kainov DA, Bratan SM et al (2010) The construction of the model of a ground surface on the basis of a stochastic description of the properties of the tool. Nauk pr Donez naz tech un-ty 7:17–31 Novoselov Iu (2012) Dynamics of surfaces forming at abrasive machining. RF, Sevastopol Stepanov IuS, Barsukov GV, Mikheev AV et al (2006) Modelling of a mechanical trajectory of abrasive grain at hydroabrasive cutting of structural materials. J Sci News Oryol State Tech un-ty Eng Indus Instrum Mak 1:58–61 Rudetskii AV, Khudobin LV (2009) Method of supplying coolant in internal grinding. RF Patent 2,359,809, 27 Jun 2009 Muslina GR, Obshivalkin MYu et al (2000) Combined grinding tool. RF Patent 2,155,129, 27 Aug 2000 Poljanchikov Ju (2002) Scientific bases of creation and application of the unicomponent abrasive tool, formed by pulse pressing and high-temperature sintering. RF, Saratov Khudobin L, Babichev A, Bulyzhev E (2006) Lubricant-cooling technological means and their application at machining by cutting, Moscow Starkov V (2007) Grinding by high-porous wheels, Moscow Khudobin L (2007) Minimisation of a loading of grinding wheels. RF, Ulyanovsk Gusev VG, Morozov AV, (2017) The Method of grinding. RF Patent 2,606,143, 10 Jan 2017

Defect Analysis of Operating Hydro-Gasified Piping System Min Ko Hlaing, Phone Htet Kyaw and B. N. Maryn

Abstract Pipelines and hydraulic systems are widely used in all branches of engineering; their failure can lead to not only economic losses, but also result in serious environmental disasters. To assure the safety of a piping system, it is very important to carry on a routine inspection of pipelines to find out the defects or corrosions and the cause of defects as early as possible. Therefore, this article analyzes the causes of defects in pipelines and hydro-gasified systems in their manufacture, assembly and installation units, as well as in the service. A detailed classification of the fault lines and hydro-gasified systems is reported in this paper. The factors of reliability of a pipeline and communications units are determined. Further improvement of repairing a pipeline system technology is suggested. The research investigated the impact of piping configuration and its material on the system reliability in whole processing. The defects, which can be occurred when connecting pipelines with other elements during their manufacture, installation and in the service, are also presented in this article.







Keywords Classification of defects Pipelines Hydro gasified system Deformation Flaw detection Leakage Connecting elements













1 Introduction Pipelines and hydro-gas systems are especially important in the construction and service of oil and gas platforms in the northern regions near the coastlines, where working conditions are harsh on temperature fluctuations, and the environment is aggressive, causing corrosion (seawater). Pipelines of hydro-gas systems are designed to connect energy sources (pumps, accumulators, cylinders) with energy consumers or actuators (power cylinders, hydraulic boosters, dampers, pumps’ M. K. Hlaing (&)
P. H. Kyaw
B. N. Maryn Komsomolsk-on-Amur State University, 27, Lenin Prospect Street, Komsomolsk-on-Amur 681013, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_94

885

886

M. K. Hlaing et al.

hydraulic motors, etc.). This connection is carried out by supplying sources of the working substance (liquid or gas) under pressure through control and distribution devices (taps, valves, spools, etc.) in the pipeline system. The quality of the treatment of external and internal surfaces has a great influence on the performance of pipelines, since the slightest surface defects become fatigue stress concentrators. In such a situation, for example the condition of the soil surrounding the pipeline, it can be subjected to corrosion that generates surface defects and reduces pipe wall thickness. Additionally, other defects in the pipe can be caused during installation and the pressure conveying of fluid or gas should be reduced to avoid rupture of the pipeline in the region of the defects that could affect the conveying capacity. Another solution could be the interruption of operation to repair the defective part. Both cases involve financial implications. Thus, it is necessary to develop a methodology to assess the effect of limiting the operating pressure in defective pipes with more accuracy, without being too time consuming [1]. The other important problem in the aircraft industry is the manufacture of high-resource, reliable pipelines, because due to the destruction of pipelines, 15–20% of accidents and disasters occur. The parts, such as curved pipe, tees, connectors, nozzles, fittings, are widely used in pipeline systems of aircraft. The nomenclature of such parts on the aircraft of the light class is in the hundreds and on the aircraft of the heavy class—in the thousands. Stamped-welded nozzles make up 20–40% of the total production of parts of pipeline systems, on which the reliability in operation and the life of aircraft largely depend. There are many publications about the defect of piping system caused by many factors. For example, in this paper [2] reported about the information of defects crossing pipelines, which buried under the rivers. These crossing pipelines will be easy to float in a flood. The impact of floods on the exposed pipes can cause significant bending deformation and even ruptures. Therefore, the deformation processes should ensure high and stable mechanical properties of the pipe material, high quality of the inner and outer surfaces, minimal wall thinning and distortion of the cross-sectional shape of the pipelines. In this works, [3, 4] reported the advanced and proven methods of calculation and design and describes modern technological processes for the manufacture, control and testing of pipeline communications that contribute to the production of aircraft with high-reliability indicators.

2 Feature of Failures and Defects in Pipeline Defects on pipelines and hydro-gas systems can lead to fatality and ecological catastrophes. Therefore, the study of defects on pipelines and hydro-gas systems will ensure their quality and production, installation, and in service [3]. In Fig. 1, reliability of hydro-gas system can be obtained by analyzing the statistical failures and malfunctions, which are detected during the process of repair and new products of the aircraft equipment [5, 6].

Defect Analysis of Operating Hydro-Gasified Piping System

887

Fig. 1 Distribution diagrams of failures and defects: a for the system as a generalized; b for pipeline systems and the nature of their failures: 1—assemblies; 2—pipelines in places of flaring, soldering and welding (in the seals); 3—filters, 4—hydraulic boosters, hydro accumulators, etc.; 5 —connecting fittings (nipples, nuts, fittings, cross connections, elbows, etc.); 6—pipelines due to metallurgical defects; 7—fuel and hydraulic systems; 8—pipeline connections

The most common failure in pipelines is (Fig. 1b) leakage due to nipple joint and destructions of pipelines, which also led to the leakage of systems. The graphs of changes in the intensity of defects in the pipelines, which are built for new and repair products in aircraft construction, show that up to 70–80% of all defects are detected in the first 50–60 h after the manufacture of products or their release from repairs (Fig. 2). The results of the qualitative and quantitative analysis, which performed according to the data of the in service and repair organizations, indicate that the characteristic of the occurrence, the causes of the main facts, and dangerous defects

888 Fig. 2 Probability of failures of pipelines in the aircraft construction

M. K. Hlaing et al. P(t) 1 0.8 0.6 0.4 0.2 0

20

40

60

80

100

t,h

of the pipelines are relevant. Therefore, it is possible to classify defects of pipelines according to the main features with predicted performance of typical methods for detecting and eliminating defects, as well as recommendations for their prevention. This work should be systematically carried out while the process of repair and service simultaneously with the improvement of methods of inspection, repair, and manufacture of pipelines. The most dangerous failures of pipelines and hydro-gas systems are mechanical damage to the pipes, which are causing leakage of the systems. The leakage of systems also can be caused by corrosion of pipes. Due to the defects of pipe can occur jamming of the valve and piston of hydraulic systems. The main factors determining the reliability of pipeline are showing at the following factors: • The quality of the manufacture and installation of pipelines in the production and repair; • Operational conditions of the pipelines (vibration, temperature, pressure pulsations, etc.); • The quality of anticorrosive protection of pipes and fittings; • The quality of technical service (compliance with the rules of care, using of quality tools, etc.) [7, 8]. To ensure a high level of reliability of pipeline during in the service and repair, it is necessary to know the main types of pipeline failures, their causes and prevention methods, as well as systematically improve the technical service of pipelines and methods for their repair, as well as classify the failures pipeline and hydro-gas system [9].

3 Classification of Failures in Pipelines The classification of failures in pipelines is shown in Fig. 3. This classifier also includes failures resulting from system breakdowns, during fire and other deviations from service conditions of the systems. The occurrence failures of pipeline can be

Defect Analysis of Operating Hydro-Gasified Piping System

889

Pipeline (Defects)

Destruction Necking Сorrugations

Manufacturing, Installation

Destruction Necking Сorrugations Corrosion

Destruction Necking

In Service

The main causes of defects

Methods of defect detection

Methods of defect elimination

Repairs

Сorrugations

Fig. 3 Classification of pipeline failures

due to a number of factors. Each failure has its own demonstration character, distinctive signs, and features of detection. Below is a brief description of faults, which are adopted in the classification. Pipelines’ failures occurring during service, installation, and repair of assembly workpieces are divided into the following main types: transverse fatigue cracks; longitudinal cracks; destruction from the impact of external forces (loads); destruction from exposure to high temperatures. In addition, destruction of the pipe walls (formation of holes) during their corrosion damage could be included in the main types of defects in pipelines.

3.1

Transverse Fatigue Crack

This type of pipe destruction is one of the most common defects. Cracks occur in the area of connection of pipes with fittings tools and connecting pieces, in the area of welded pipes with nipples, as well as in the area of jointing points. The main causes of transverse fatigue crack of pipelines in cross section can be: • High-level vibrations of bending pipes, the driving source, which is usually as vibrations of structural elements of the assembly workpieces and pressure vibrations of the working fluid; operating variable loads; • the main sources of stress force are the mounting force inaccuracies in the jointing of pipes and fittings:

890

M. K. Hlaing et al.

• repeat reloading of the pipe material in the sealing area, which can be in the condition of periodic touching of moving structural elements acting on pipelines; • the occurrence of transverse fatigue cracks of the pipe walls is also possible with their radial oscillations due to pressure pulsations of the working fluid. The occurrence of a working fluid leak from the connector during operation is a typical external sign of the leakage of the joints due to the formation of a through fatigue crack in the pipe flaring zone. In an attempt to eliminate leaks by additional tightening the nut, the cracks are usually increased. In a welded joint, fatigue crack can be caused, besides the above reasons, by the presence of abrupt transitions in the joint zone, material burnout, welding effects, etc.

3.2

Longitudinal Cracks and Failure

This type of defect includes fatigue cracks and failure due to internal pressure [10, 11]. Fatigue longitudinal destruction of pipes most often occurs in sections with a large ovality, which is mainly formed during the bending of pipes. The source of vibration of the distorted section of the pipe during bending is the vibration of the pressure of the working fluid. The repetition of maximum loading that exist in the curvature of the pipe walls also occurs in the condition of its periodic bends in the process of operation of mechanisms (e.g., articulated joints) or during vibrations of the bent pipe in the flexure plane. By comparison of the cracks, which can occur from the weakness of the pipe wall by metallurgical defects with the longitudinal fatigue cracks, the distinctive signs are blockages of the outer edges of the crack and their tortuosity. In additionally, the edges of cracks, which are occurred by metallurgical defects, have sharp edges and there is no tortuosity [12, 13]. The rupture of the pipe wall due to its local weakness (wall thickness, wall thinness) or due to excessive overpressure is caused a significant beginning crack and plastic deformation of the material of the walls in the rupture zone. When exposed to excessive pressure, a tube is expanded throughout its length. In the case of insufficient strength of the pipe wall in the area of destruction, the expansion occurs only near the site of destruction.

3.3

Destruction Due to External Forces (Loads)

Because of breakdowns of hydro-gas systems or falling into the area of moving parts of other objects, the following pipe failures can occur: longitudinal fracture; destructions from transverse; or longitudinal bending.

Defect Analysis of Operating Hydro-Gasified Piping System

3.4

891

Destruction Due to Tensile Loads

Destruction due to tensile loads is characterized by the formation of a neck (narrowing) in the rupture area. If tensile forces act on a pipeline, which consists of several pipes, then, as a rule, its destruction occurs due to the failure of one of the pipes among them. Damage of the pipe from bending occurs when subsequent repeated straightening and bending due to the secondary movements of the sealing of the pipe.

3.5

Destruction from Exposure to High Temperatures

Pipe failures can occur in case of fire and breakthrough of hot gases from the combustion chambers of the engines, short circuit of the electric wire on the pipe, and other types of high-temperature exposure. The destruction of steel pipes from the effects of internal pressure of the working fluid (gas) during their local heating is a significant opening of the crack and plastic deformation of the material of the pipe wall in the heating zone. Some aluminum alloys pipes when heated to temperatures of not more than 300–350 °C are destroyed in the same way as steel pipes. At higher temperatures, due to the loss of plasticity of the material, the destruction of pipes occurs when lower internal pressure. When the metal is melting or short circuit of the electric wire on the pipe wall, the destruction of pipe walls surface will be become with the hole that have the significant form of melting edges. That feature can be possible to distinguish it from holes that can form during corrosion and then burning the pipeline in case of fire [14].

3.6

External Damage to Pipes

The most common are the following groups of external damage to the pipes: • Damage of the protective coating (without destruction of the surface of the pipe); damage to the protective coating can be the formation of scuffs when fastening pipe with clamp, various kinds of defects during assembly and disassembly works. In addition, damage to coatings is their fissure, swelling, flacking due to external factors (aggressive environment, high temperatures, etc.), or poor quality of their application in the manufacture of pipelines. When exposed to high temperatures can cause darkening and carbonization protective coatings. • Damage of the surfaces of pipes without protective coating and with protective coating; destruction of the pipe geometry.

892

M. K. Hlaing et al.

External damage to pipes can occur from friction between them or elements’ structure, disobey of installation technology, the usage of poor quality tools, etc. Damage to pipe surfaces of damage includes risks, scratches, small nicks (no more than 0.1 mm), deep nicks (more than 0.1 mm) and dimple, scuffs due to friction between structural elements or conjugated pipelines with the hole.

3.7

Destruction of the Pipe Geometry

When the manufacturing of pipelines, assembly and service, repair, fracture of the pipe geometry [15–18] are possible: • • • • •

Sharp bends of pipes; Inadmissible (more than 5–10%) ovalization of pipes during their bending; Deformation of the pipe with clamp; Distortion of the pipes due to pipe connector when tightening screwed nuts; Expansion of the pipe during hydraulic shock or excessive internal pressure.

4 Corrosive Damage of Pipes and Fittings The following main forms of corrosion types are distinguished: • Uniform (surface) corrosion; • Local (pitting) corrosion; • Internal crystalline corrosion. In addition, there are such varieties of local corrosion as transcrystalline, subsurface, etc. By the type of corrosion content, it is possible to determine the mechanism of its occurrence. For example, when electrochemical corrosion, contents of corrosion can be easily removed by mechanical method. Chemical corrosion content (gas, corrosion due to the content of chemically active impurities and components in fuels, oils, and other liquids) has a more dense structure. The corrosion area of pipes is area with a broken protective coating, as well as areas affected by corrosive environments. Internal surfaces of pipelines, which are located in the lowest part of the system and bent in a vertical plane (U-shaped sections), where the accumulation of moisture, that are contributing to the occurrence of corrosion, are the most susceptible to corrosion. During the given time of processing system, destruction of the pipe may occur with the hole. When pitting or pitting corrosion is detected on the inner surface of the pipes, selective monitoring (with cutting) of the pipes of this type (10% of Al– Mg and Al–Mn pipes and 5% of steel pipes) is performed.

Defect Analysis of Operating Hydro-Gasified Piping System

893

Connecting parts for pipeline fittings (nuts, nipples, bush fitting, etc.) are most susceptible to mechanical stress during assembly and disassembly of connections. Collapses, nicks, abrasion of threads, and other defects cause damage to the protective coating and, as a consequence, corrosive damage of fittings. The main consequences of corrosion on the fittings are: – a sharp crack growth (“welding”) in the threaded connection and in nipple–pipe pairs, nipple nut, resulting in possible damage to pipes and parts, such as twisting, crushing of faces, etc.; – systems’ contamination with corrosion during assembly and dismantling.

5 Conclusion As a result of the qualitative analysis of defects, can be solved causes of the formation of defects and the methods for their prevention. Destruction of pipelines is the most dangerous type of pipeline system failures (especially such as fuel, oil, oxygen), even in isolated cases should be carefully analyzed and investigated with the involvement of research institutions, if necessary. Based on the results of the analysis of pipeline failures related to repairs on operational repair service, proposals are being developed for further improvement of the repair technology of pipeline systems. According to the results of the analysis and research of pipeline defects, during operation, there is a real possibility of choosing a configuration and pipelines.

References 1. Shang HY (2016) A methodology for analysis of defective pipeline by introducing stress concentration factor into beam-pipe finite element formulation. Acta Scientiarum Tech 38 (3):313–320. https://doi.org/10.4025/actascitechnol.v38i3.27988 2. Li S, Duan Q, Zhang H, Wang J (2017) Failure analysis of the floating pipeline with defect under flooding load. Eng Fail Anal 77:65–75. https://doi.org/10.1016/j.engfailanal.2017.02.01 3. Maryin BN (1998) The manufacture of pipelines of hydro-gas systems of aircraft. Moscow, pp 5–45 4. Yang H, Li H, Zhang ZY, Zhan M, Liu J, Li GJ (2012) Advances and trends on tube bending forming technologies. Chin J Aeronaut 25:1–12 5. Kolykhalov DG, Sysoev OE, Ivanov IN (2016) Evaluation of the manufacturability of aircraft piping systems in the early design stages. Works MAI 90:26–30 6. Maryin BN, Feoktistov SI, Kolykhalov DG, Kuriniy VV, Ivanov IN (2016) Investigation of combined processes in the manufacture of aircraft parts. Sci Notes KnASU 2(26):34–41 7. Maryin BN (2008) Method of bending thin-walled pipes with filler. RU Patent 2337779 C2, 10 Nov 2008 8. Maryin BN (1990) Method of manufacturing of steeply curved thin-walled adapters. RU Patent, 1581411 A1

894

M. K. Hlaing et al.

9. Sapozhnikov VM, Marin BN (1995) Intensification of technological processes of forming parts from pipes. Moscow, p 175 10. Gu RJ, Yang H, Zhan M, Li H, Wang GX (2005) Effect of mandrel on cross section quality of thin-walled tube numerical controlled bending. Trans Nonferrous Met Soc China 15: 1264–1274 11. Li H, Yang H, Zhan M, Gu RJ (2007) The interactive effects of wrinkling and other defects in thin-walled tube NC bending process. J Mater Process Technol 187–188:502–507 12. Paulsen F, Welo T (2003) An analytical model for prediction of tube ovalization in bending. In: Brucato V (ed) ESAFORM 2003: proceedings of the sixth ESAFORM conference on material forming, Salerno, Italy, pp 775–778 13. Lin Y, Yang H, Li H, Zhan M (2003) Influence of forming parameters on the wrinkling during the NC thin-walled tube bending process. Acta Aeronaut Astronau Sin 24:456–461 14. Maryin SB, Aung PW, Hlaing MK, Timoshinin MA (2018) Device for modern hardening and dispensing heated to a temperature for quenching the tubular billet of heat-treatable aluminum alloy. RU Patent, 2649102 15. Maryin BN (2007) Method of bending thin-walled pipes. RU Patent, 2322320 C2, 20 Apr 2007 16. Maryin BN (1993) Horn-shaped core for the manufacture of taps with flanges of pipe blanks. RU Patent, 1787617 A1 17. Murata M, Kuboki T, Takahashi K, Goodarzi M, Jin Y (2008) Effect of hardening exponent on tube bending. J Mater Process Technol 201:189–192 18. Yang JB, Jeon BH, Oh SI (2001) The tube bending technology of a hydroforming process for an automotive part. J Mater Process Technol 111:175–181

Contact Zone Effect Analysis onto the Deforming Rollers Geometrics During the Surface Plastic Deformation Treatment Y. N. Oteniy, O. V. Martynenko and N. I. Nikiforov

Abstract The research studies the problem-solving methods of the parts geometrics in the contact zone depending on the deforming rollers geometrics. The direct and indirect methods are used to solve this problem. Different forms of contact zones were studied. Research results showed that the contact zone form and size do not provide the same working environment. Maximum depth of the roller feeding-in into the part surface and the contact zone volume can be used as such data. This problem solution determines the surface quality influence factors during the parts’ treatment with the surface plastic deformation. The direct method is commonly used to solve the problems about the different deforming rollers geometrics influence and the treatment factors influence on tension in the contact zone and the surface layer quality during the surface plastic deformation. The method consists of determining the contact zone geometric parameters depending on the specific parameters of the deforming roller used in the deformation process. During the indirect method, the arbitrary shape and the size of the contact are defined first, after that, the contact zone correspondence to the type and the size of the roller is determined. Keywords Contact zone


Deforming roller
Geometrics

1 Introduction The direct method is commonly used to solve the problems about the different deforming rollers geometrics influence and the treatment factors influence on tension in the contact zone and the surface layer quality during the surface plastic deformation. The method consists of determining the contact zone geometric parameters depending on the specific parameters of the deforming roller used in the deformation process. During the indirect method, the arbitrary shape and the size of Y. N. Oteniy
O. V. Martynenko (&)
N. I. Nikiforov Kamyshin Technological Institute (branch) of Volgograd State Technical University, 6A Lenin st, Volgograd reg, Kamyshin 403874, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_95

895

896

Y. N. Oteniy et al.

the contact are defined first, after that, the contact zone correspondence to the type and the size of the roller is determined. The initial choice of the contact zone is necessary to model the required mode of deformation in its limits as well as to determine the contact optimum geometric features and necessary treatment process conditions and also to provide the required quality of the surface layer. Besides, the contact zone specifics pressure grading analysis impact onto the surface quality level can be initially separated from detail roller interaction real chart and then, after the necessary deformation process condition choice, revert to the question about the tension correspondence to the roller type. In some cases, it is not possible to solve the individual questions using the direct method, in particular, to determine the rollers size effect onto the contact tensions with the desired contact shape and size, minimum amount of the practice rollers, instrument retrieval, and its centralized enterprise production [1–18].

2 Main Part and Mathematical Model The contact zone is definitely determined by setting up the contour line equation and by the roller’s depth of penetration alteration along its maximum loading (maximum tensions). During the deforming rollers geometrics determination that secures the desired contact, the detail diameter and the type of the treated surface should be also considered (arbor, bore or plain face). Let us assume that the roller’s and detail’s arbors are parallel to each other. Omitting conversions, on the basis of the roller feeding-in into the detail surface chart shown on Fig. 1, the following expression to calculate roller’s radius and depth of its penetration along the maximum load disturbance line. During the arbor treatment rpв =

C C2 + + B 2 + 4 ⋅ D 2 ⋅ zk 2 ; 2 4

hk = rpв − rpp

ð1Þ ð2Þ

The following symbols are used to reduce the entry C ¼ 2B þ 4D2 ; D ¼ Rd  rpp ; B ¼ R2d  D2 ; where Rd—treated surface radius, rpp—roller’s radius. During the bore size treatment compatible to the start of the contact zone (heading radius); zk—contact zone half-width predetermined transformation

rpo

vffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi u  2 u   zk t 2 1 ¼ ðRo þ rpp Þ þ R2o  2Ro Ro  rpp Ro

where Ro—treated bore radius.

ð3Þ

Contact Zone Effect Analysis onto the Deforming Rollers …

897

Fig. 1 Design diagram for roller’s radius re-proportioning determination by the set shape and contact zones sizes during the bores and arbors treatment

The initial roller’s radius r is chosen arbitrarily and depends on the instrument design characteristics, the quantity of the simultaneously set rollers along the detail circumference and other considerations, determined by the set or chosen criteria of the technological or constructive nature. There are no difficulties to establish the contact zone half-width changes by its length. Any circumference equations can be considered. For example, ellipse and drop-shaped contacts that are commonly used in manufacturing practice are shown in Fig. 2. Drop-shaped contact (Fig. 2a) consists of two parts: lead-in area located on the length L1 and the runout area located on the length L2. Let us suppose that the set area represents the ellipse half with the differential axles L1 и zkmax, the contact half-width will be determined from the relation:

898

Y. N. Oteniy et al.

Fig. 2 Commonly practiced contact zones types. a Drop-shaped contact zone; b ellipse contact zones

zK1

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi   lk  0:5L1 2 with 0  lk  L1 ; ¼ zkmax 1  0:5L1

ð4Þ

On the runout area zK2 ¼

zkmax
ðLk  lk Þ ðLk  L1 Þ

with L1  lk  Lk

ð5Þ

For the ellipse contact, half-width contact equation along its length will be presented as follows. sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi   lk  0:5L1 2 zk ¼ zkmax 1  0:5L1

with 0  lk  Lk

ð6Þ

For the circular contact in the last equation, let us assume that zkmax = L1. For the contact zone volume determination (part of the metal, displaced by roller), the author obtained the relationship ZLj  Vj ¼

rp2
arcsin

     Zj  2
zj
rp  hj
dlj rp

ð7Þ

0

where rp ; zj ; hj —rollers radius, half-width, and the depth of the rollers feeding-in variations along the contact length, lj и Lj —current contact length coordinate and the complete contact length. During calculations as applied to the ellipse contact was determined that with the same set contact zone, the depth of the rollers feeding-in into the detail surface and the contact zone volume depend on the treated arbors radii and the deforming rollers diameters: the bigger the roller’s radius the less is the contact zone volume and the depth of the roller feeding-in (Figs. 3 and 4).

Contact Zone Effect Analysis onto the Deforming Rollers …

899

Fig. 3 Roller’s radius variations relations (a) and the depth of the roller’s feeding-in (b), providing ellipse contact from the contact length to solve the reverse contact problem

Fig. 4 Contact zone volume deviations and the roller’s depth feeding-in from the rollers initial diameters rpp with the different treated detail radii values: Rd = 15, 40 and 100 mm

That explains why, with the same contact half-width, roller with the bigger diameter to provide the same contact width should be fed-in deeper. The scheme on Fig. 5 demonstrates this statement. Fig. 5 Rollers direction and details with two different radii, which circuits simultaneously cross two points on the detailed surface, belonging to the contact contour line

900

Y. N. Oteniy et al.

3 Conclusion The maximum stress in the most loaded contact zone part will be increasing, this will entail linear magnification tension increase of the rest of its part, consequently, the resulting deformation force should increase. As a result, contact zone shape and size do not provide singularly similar treatment conditions. Rollers maximum feeding-in into the detailed surface and the contact zone volume can simultaneously serve as such parameters, because, in turn, the contact zone volume and the depth of feeding-in depend on the roller’s and detail’s size.

References 1. Martynenko OV (2017) Determination of plastic deformation depth distribution during the plastic surface deformation. Mod Technol Achievements 4:20–22 2. Martynenko OV (2017) Physical-mechanical effects influence in the contact zone onto the hardening characteristics and detail surface quality. In: PKK polytechnic, innovative technologies in study, Volgograd, p 304 3. Martynenko OV (2014) Management automation constructive-technological parameters of treatment by surface plastic deformation by rollers. Achievements Mod Nat Sci 11:41–43 4. Nikiforov NI, Martynenko OV, Moroz VU (2017) Determination of plastic deformation depth expansion when surface is plastically deformed by rollers. Achievements Mod Nat Sci 4:48–50 5. Martynenko OV (2013) The study of stress state influence in the area of contact on the points kinematics of deformed surface when the surface is plastically deformed by rollers. Achievements Mod Nat Sci 12:63–65 6. Otenyi YN (2006) Technological support of the quality of machine parts when processing surface plastic deformation rollers. In: Reference book. Engineering Magazine, 23p 7. Suslov AG (2013) Mechanical engineering. Kno Rus, Moscow, 336p 8. Pleshakov VV (2006) Surface plastic deformation. Stankin, Moscow, p 192 9. Zaiydes SA (2015) Encyclopedia of surface plastic deformation. Irkutsk Publishing House, Irkutsk National Research Technological University, 395p 10. Markin AA (2013) Thermomechanical elastoplastic deformation. Physics and Mathematics Faculty, Moscow, p 319 11. Ivanov NB (2013) Theory of the rigid body deformation. KNRTU, Kazan, p 124 12. Oteniy YN (2005) Machine components technological quality assurance by surface plastic deformation: study. Volgograd Technological Institute, Volgograd, p 224 13. Masterov IA, Berkovsky VS (1989) The theory of plastic deformation and metal treatment under pressure. Metallurgy, Moscow, p 324 14. Odintsov P (1987) Plastic deformation surface strengthening and treatment. In: Reference book. Machine construction, Moscow, 329p 15. Papshev DD (1978) Strengthening and conditioning using surface plastic deformation. Machine Construction, Moscow, p 152 16. Shneider UG (1989) The technology of metal finish treatment by pressure. Polytechnic, SPb, 411p 17. Joson W (1978) Engineering plasticity. London Yar Nostrand Reinhold Company, 577p 18. Martynenko OV The study of influence of deforming rollers geometric parameters on the quality of surface layer when treating with plastic deformation

Increasing Resistance of Cutting Tool with Diamond Burnishing N. Papsheva and O. Akushskaya

Abstract The paper presents the data from a study of a diamond burnishing effect on the physico-mechanical characteristics of the surface layer and the operational characteristics of the cutting tool. Diamond burnishing is made with a tool with a radius of up to 4 mm, so the burnishing force will be small up to 250 N, which makes it possible to process low-rigid and thin-walled parts. The previous operation is grinding (sharpening) the tool. As a result of diamond burnishing, a new microrelief with a characteristic brilliance is formed. It has been established that the height of microroughness is influenced by the processing modes: smoothing force, feed, radius of the diamond sphere. In this case, the height of microasperities is significantly reduced. An intense-oriented plastic deformation of the surface layer and a change in its volume due to the phase c ! a transformation contributes to the formation of residual compressive stresses. With diamond smoothing, intensive crushing of grains, an increase in the density of dislocations, a decrease in the amount of residual austenite, and the development of microdistortions are observed. The initial specific wear of the tool reaches its maximum values at the breaking-in stage, and then the wear rate decreases sharply. Studies have shown that diamond smoothing doubles the durability of the cutting tool. Keywords Diamond burnishing

 Plastic deformation  Wear resistance

1 Introduction Diamond burnishing relates to the methods of finishing-strengthening processing. The method consists of surface plastic deformation of a part or tool with a diamond tool having a spherical shape. At the same time, during the processing, the surface irregularities remaining after the previous treatment are smoothed, surface hardness increases, residual compressive stresses occur. This leads to an increase in the N. Papsheva  O. Akushskaya (&) Samara State Technical University, 244, Molodogvardeiskaia, Samara 443100, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_96

901

902

N. Papsheva and O. Akushskaya

operational characteristics of the service life of machine parts and tools, which are largely determined by the state of the surface layer. A particular feature of this method is the use of a diamond as a deforming tool with such properties such as high hardness, thermal conductivity, low index of friction on metal [1–12]. Due to its high hardness, a diamond can treat various materials, including tool steels, hardened to HRC 60–65 [13, 14].

2 The Effect of the Diamond Burnishing on the Physico-Mechanical Characteristics of the Surface Layer Since diamond burnishing is done with a tool with a small radius up to 4 mm, the smoothing efforts will be small (up to 250 N), which makes it possible to process low-rigid and thin-walled parts [15]. While processing, diamond sphere smoother, pressed with a certain radial force Py to the workpiece surface, glides over it, causing plastic deformation of the surface. As a result, the initial microroughness is smoothed and a new microrelief is characterized by a much smaller height of asperities [16, 17]. The previous treatment before burnishing the cutting tools is grinding (sharpening). At the same time, the polished surface contains traces of cutting of individual grains; after diamond smoothing, the treated surface acquires a characteristic brilliance. The main factor affecting the roughness parameters is the effort of burnishing Py, with increasing of which increases the depth of implementation and the contact area of the smoother with the surface that is being treated, which leads to a significant reduction in the arithmetic mean height of Ra as microirregularities (Fig. 1). When the value of Py = 200 N is reached, the roughness reduction stops and reaches 0.05 l. The feed is the second most important factor affecting surface roughness. With an increase in the feed, an increase in Ra occurs, which is caused by a decrease in the degree of plastic deformation due to the low loading ratio. The roughness of the hardened surface is also affected by the tool radius r, with an increase in which the values of Ra decrease. So with an increase in the radius r from 1.5 to 3 mm, Ra decreases from 0.2 to 0.08 mm (Fig. 1). As the results of research have shown, the main reduction in roughness occurs during the first pass; with an increase in the number of passes to 3, the value of Ra decreases slightly. When grinding high-speed steels, tensile residual stresses of up to 600 MPa occur in the surface layer. Residual diagrams after diamond burnishing have a qualitatively different nature. Intense plastic deformation of the surface layer and a change in its volume due to phase c ! a transformation contributes to the occurrence of residual compressive stresses [18]. It has been established that a change in Py from 150 to 200 N is accompanied by an increase in residual compressive stresses from 150 to 700 MPa (Fig. 2a).

Increasing Resistance of Cutting Tool with Diamond Burnishing

903

Fig. 1 Dependence of arithmetic mean height of asperities Ra on the effort of smoothing Py

Fig. 2 Effect of the burnishing force Py and the feed S on the tangential residual stresses: a (1) Py = 50 N, (2) Py = 100 N, (3) Py = 200 N, (4) Py = 300 N; b (1) S = 0.02 mm/rev, (2) S = 0.07 mm/rev, (3) S = 0.15 mm/rev

A decrease in supply from 0.15 to 0.02 mm/rev leads to an increase in residual stresses to 950 MPa (Fig. 2b). One of the main characteristics of the quality of the surface layer, significantly affecting the performance of the cutting tool, is the surface hardness [19]. The study of the effect of diamond burnishing on the surface hardness was performed by the method of multivariable planning. The adopted methodology allowed minimizing the number of experiments required to establish the most likely dependence of process indicators on its parameters, and to obtain more information about the interaction of individual factors. The value of the maximum surface hardness HV was chosen as the optimization parameter. The variables were the radial force Py, the radius of the diamond sphere

904

N. Papsheva and O. Akushskaya

r, the longitudinal feed S, the rotation speed of part V, the number of passes i. The dependence of HV on the parameters of smoothing was approximated by the equation y ¼ b0 þ b1 x 1 þ b2 x 2 þ b3 x 3 þ b4 x 4 þ b5 x 5 ; where y is the optimization parameter, b0 ; b1 ; b2 ; b3 ; b4 ; b5 are regression coefficients, x1 ; x2 ; x3 ; x4 ; x5 are independent variables. To solve the equation, an orthogonal central composition plan with 1/4 replicas was chosen; the results of the experiments were used to determine the regression coefficients b. The constructed mathematical model has the following form: y ¼ 910 þ 27:5x1  7:5x2  5x3  5x4  2:5 x5 : Analysis of the equation shows that the radial force Py and the tool radius r have the greatest influence on the surface hardness. The results of microhardness determination showed that with an increase in the radial force of Py from 50 to 200 N, the maximum value of microhardness Hd increases from 9.8 to 10.2 GPa. A further increase in Py leads to a slight decrease in Hd, which can be explained by the excessive strain hardening (Fig. 3a). The amount of the feed, determining the frequency of application of the deforming force to each point of the treated surface, affects the microhardness. At the feeds of 0.02– 0.07 mm/rev, the maximum increase in the microhardness is observed (Fig. 3b). The intensity of strain hardening amounted to 15–18%, depending on the processing modes. Analysis of the statistical processing of microhardness measurements showed that the deviations are reduced compared with grinding. Accordingly, the scattering field of Hd values decreases, that is, the structure of the surface layer becomes more uniform. With diamond burnishing, the fragmentation of grains (D = 4.3  10−2 lm) and the formation of a finely dispersed structure in the surface layer up to 200 lm in depth are observed (Fig. 4a).

Fig. 3 Effect of burnishing force Py and feed S on microhardness distribution processed material R9К5: a (1) Py = 50 N, (2) Py = 100 N, (3) Py = 200 N, (4) Py = 300 N; b (1) S = 0.02 mm/rev, (2) S = 0.07 mm/rev, (3) S = 0.15 mm/rev

Increasing Resistance of Cutting Tool with Diamond Burnishing

905

Fig. 4 Effect of the radial force Py on the fragmentation of the grains (a) and the dislocation density q (b). (1) Py = 50 N, (2) Py = 200 N, (3) the fault strength

There is a growth of microdistortion, the fine carbides separation and phase c ! a transformations, which contributes to the improvement of the wear resistance of the tools [18, 20]. And with increasing Py, the maximum values of microdistortion grow, being located at a certain distance from the surface, which is explained by the local heat generation during smoothing (Fig. 5a). In this case, the density of dislocations increases (q to 0:78  1012 ) (Fig. 4b), and the amount of residual austenite A decreases from 27 to 15% (Fig. 5b).

Fig. 5 The change of the fine crystal structure during diamond burnishing (a microdistortion; b residual austenite A) (1) Py = 50 N, (2) Py = 200 N, (3) the fault strength

906

N. Papsheva and O. Akushskaya

To assess the stress–strain state, a functional relationship was established between stress and strain, strain and hardness. As can be seen from Fig. 6, the dependence of the intensity of plastic deformation ei on the radial force of Py has an extreme nature. The maximum value of ei takes place at Py = 200 N. A further increase in the force leads to a decrease in ei , which is associated with excessive strain hardening and a discontinuity of the surface layer. The maximum value of ei is at some distance from the surface. This is confirmed by the results of X-ray structural analysis, which states the highest density of dislocations at a certain distance from the surface (Fig. 6). During diamond burnishing, friction forces that affect the heating of the treatment area, the plastic deformation, and the surface quality parameters arise in the contact zone as a result of overcoming the adhesive bonds that occur between the surfaces of the tool and the workpiece. In this case, the friction coefficient ff includes the deformation fd and adhesive fa component: ff ¼ fd þ fa : The deformation component of the friction coefficient can be determined by the formula rffiffiffi h fd ¼ 0:55 ; r where h is the depth of introduction of the smoother, mm; r is the radius of the working part of the smoother, mm. Adhesive component for steel R9К5 is fa = 0.036. The calculation results presented in Fig. 7 show an increase in the friction coefficient ff from 0.042 to 0.07 and the deformation component fd from 0.015 to 0.038 with an increase in the smoothing force from 50 to 3000.

Fig. 6 The intensity of the deformation in diamond burnishing: (1) ei ¼ /ðhÞ; (2) ei max ¼ f ðPy Þ

Increasing Resistance of Cutting Tool with Diamond Burnishing

907

Fig. 7 The influence of burnishing force on the friction coefficient: (1) fd ; (2) ff

3 The Effect of the Diamond Burnishing on the Wear and the Durability of Cutting Tools Research has shown that the wear resistance of smoothed tools increases 1.5–2.5 times. The greatest increase in the wear resistance is observed for the instrument, the surface layer of which is characterized by high values of residual compressive stresses, higher microhardness. This is most effectively achieved by increasing the hardening force Py. Resistance tests were carried out using the method of radioactive isotopes, which allows to obtain additional information about the nature of the wear of the cutting tool. This method is characterized by high sensitivity and makes it possible to drastically reduce the time of testing. Neutron irradiation was chosen as the activation method. Analysis of the graphs (Fig. 8) shows that the initial specific wear for all the cutters reaches maximum values (breaking-in stage), and then the wear rate decreases sharply. In the area of the steady-state wear after 30 min of operation, the specific wear of non-reinforced cutters was Dq = 185 imp/s, and the smoothed Dq = 65 imp/s. The investigation of the effect of cutting speed on specific wear showed that the least wear of hardened cutters is observed at speeds up to 35 m/min. With an increase in cutting speed, the specific wear of hardened cutters grows less intensively, and at a speed of 80 m/min Dq = 185 imp/s and for the non-reinforced cutters Dq = 350 imp/s. When turning, the wear of the cutting tool occurs mainly on the back surface; therefore, in order to increase productivity, it is advisable to expose only the back surface for the smoothing. Figure 9 shows a graph of tool resistance T versus machining speed V, from which it follows that diamond smoothing increases the tool life by a factor of two compared to grinding, while with decreasing processing speed, the resistance increases.

908

N. Papsheva and O. Akushskaya

Fig. 8 Effect on diamond burnishing specific wear (R9K5 material, (1) Py = 200 N, (2) the fault strength)

Fig. 9 The effect of diamond burnishing on the durability of the cutting tool: (1) after diamond burnishing, (2) after grinding

4 Conclusion Diamond burnishing greatly reduces the height of microroughness, promotes the formation of residual compression stress, increases microhardness, improves fine crystal structure, reduces wear, resulting in increased tool life by a factor of two. When the tool is re-brushed on the non-reinforced surface, the hardening effect is preserved. This allows to use the cutting tool during the entire period of operation without an intermediate diamond burnishing after each regrinding.

Increasing Resistance of Cutting Tool with Diamond Burnishing

909

References 1. Smelyansky V (2002) Mechanics of hardening of parts by surface plastic deformation. Mechanical Engineering, Moscow 2. Torbilo V (1972) Diamond burnishing. Mechanical Engineering, Moscow 3. Dvoivnev A, Barats F, Aleksandrov S (2007) Improvement finishing and hardening treatment surface diamond burnishing. Autom Mod Technol 4:36–39 4. Gubanov V (2010) Roughness parameters of ironed surfaces. Strengthening Technol Coat 11:6–9 5. Mihin N (1968) Friction in a plastic contact. Science, Moscow 6. Chepa P (1988) Operational properties of hardened parts. Science and technology, Minsk 7. Usov Y (2017) The study of contact interaction in the working area of diamond smoothing. In: Research and development in the field of engineering, energy and management. p 82–86 8. Odintsov L (1984) Finishing parts diamond burnishing and vibrovyglazhivaniem. Mechanical Engineering, Moscow 9. Torbilo V, Plotnikov A (1983) Diamond smoothing low-hard and thin-walled parts. Abrasive Diamond Processing, Perm, pp 25–32 10. Papsheva N (2016) Methods of diamond burnishing machine parts and tools. Mod Probl Theor Mach 4:102–104 11. Gubanov V (2013) Smoothing: quality, technology and tools. Academy of Natural Sciences, Moskow 12. Ryazanov-Khitrovskaya N, Pyzhov I, Kryukova N (2015) Some ways to improve the efficiency of the diamond smoothing process. High technology in mechanical engineering. Kharkov 25:173–182 13. Luoa H, Liub J, Wangb L, Zhonga Ϙ (2006) Study of the mechanism of the burnishing process with cylindrical polycrystalline diamond tools. J Mater Process Technol 3:9–16. https://doi.org/10.1016/j.jmatprotec/2005.03.041 14. Hongyun L, Jianying L, Lijiang W, Ϙunpeng Z (2006) The effect of burnishing parameters on burnishing force and surface microhardness. Int J Adv Manuf Technol 7(8):707–713 15. Korhonen H, Laakkonen J, Hakala J, Lappalainen R (2013) Improvements in the surface characteristics of stainless steel workpieces by burnishing with amorphous diamond-coated tip. Mach Sci Technol 4:593–610 16. Barats A, Kochetkov V (2013) Optimization of finishing-hardening treatment with a diamond indenter. Strengthening Technol Coat 9:45–47 17. Barats, A, Kochetkov V (2013) Modes and quality of processing of tools from high-speed steels by the method of SPD. Automobile industry, 2 18. Kurdyukov V, Ostapchuk A, Ovsyannikov V, Rogov V (2010) Features of formation of the microrelief at burnishing thermostrengthened materials. Bull Kuzbass State Tech Univ 78:7–9 19. Barats A, Kochetkov V (2013) Residual stress investigation at finishing and hardening treatment of high-speed steel. J Technol Mech Eng 1:21–25 20. Barats F, Arzhanukhina S, Kochetkov A (2013) Ensuring the constancy of the radial force in finishing and hardening treatment. In: Abstracts of the Materials 11 All-Russian scientific-practical conference, NSTU, Novosibirsk, 27 Mar 2013

Determining Coordinates of Cutting Force Application Point in Grinding Zone V. A. Nosenko and M. V. Danilenko

Abstract The article is devoted to the method of determining coordinates of the resulting cutting force application point for wheel periphery flat grinding. The solution is based on the method of summing cutting forces from all grain vertices in contact with the treated surface. The number of contacting grains is determined by the probability of contact. The actual depth of grain vertex cut, located in an arbitrary micro-volume of the contact zone, is calculated from the contact probability level equal to 0.0027. This level is taken as the boundary defining the line of the removed material in the cutting zone. On the basis of the developed mathematical model for determining the cutting force and the created software, an example of determining the coordinates of the cutting force application point for specific grinding conditions is shown. The results of the study may be used to solve problems related to determining the intensity of heat generation in the contact zone of the grinding wheel and the billet, calculation of cutting power, tool wear, and other issues related to the study of the grinding process. Keywords Mathematical models Contact probability


Grinding
Abrasive tool
Cutting force


1 Introduction Grinding is the most common method of machine parts final surface treatment. Improving the efficiency of the grinding process to ensure observance of the necessary requirements for quality of the treated surface is a priority in abrasive processing. Cutting force is considered a general indicator of the process, depending on a large number of factors. Therefore, the design and study of the cutting force can be V. A. Nosenko (&)
M. V. Danilenko Volzhsky Polytechnic Institute (Branch) of VSTU, 42a Engels Street, Volzhsky, Volgograd Region 404121, Russian Federation e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_97

911

912

V. A. Nosenko and M. V. Danilenko

most effectively carried out with the help of computer technology, based on the relevant mathematical models, which reflect the characteristics of the billet, grinding wheel, and treatment. The cutting force mathematical model reliability largely determines the reliability of the grinding process design, including the quality parameters of the treated surface. A large number of research papers are devoted to the study of cutting force in grinding and the development of its mathematical models [1–6]. In most of the known models, the solution is based on the method of summing the cutting forces from all the grain vertices contacting with the treated surface. To do this, we need to examine the wheel working surface profile part, located in the interaction zone at a given point in time. Breaking this part of the profile into layers, we count the number of contacting grains in each of them [7–10]. Assuming that the cutting force per grain depends on the depth of cut, we summarize their values within a single layer. Upon making calculations for all layers, we are able to determine the total cutting force [11–15]. Considering the distribution density of grain vertices as a continuous function of the distance from the apex of the peak and the thickness of the layer as the infinitely small value, in a number of model’s researchers go from summation to integration. The difference between the existing models is the method of determining the density distribution of grain vertices, the number of contacting vertices within a single layer, and the calculation of the abrasive grain vertex cutting force.

2 Problem Statement To solve a number of problems, for example, for calculations related to the determination of machining accuracy, machine power, rigidity of the technological system, the intensity of heat generation in the grinding wheel and billet contact zone, wear of the wheels, etc. It is necessary to determine the cutting force and its direction. The resulting force application point can be conditionally assigned to different areas of the grinding wheel and billet contact zone. The methods proposed in this article allow to determine its coordinates and the direction for different treatment conditions.

3 Main Body The grinding wheel and billet contact zone is a certain area V, bounded on the bottom by the conditional outer surface of the grinding wheel, and on the top- by the COZ plane, see Fig. 1, drawn through the billet surface roughness maximum protrusion. Considering the radial wear of the grinding wheel and metal removal Dr, the contact zone upper limit of which varies along the curved arc C′O′A′, and the actual contact zone will be in the form of a curved segment A′B′C′O′.

Determining Coordinates of Cutting Force Application …

913

Fig. 1 Abrasive tool and billet contact zone at flat cut-in grinding

Divide the contact zone area V by horizontal and vertical planes into elementary micro-volumes, see, Fig. 2. The distance between the planes is equal to the radial wear of the wheel per one circuit DR. Elementary volume normalized to grinding wheel height unit: DV = Dy  Dz, where Dy = Dz = DR. Position of the center of each micro-volume DVij is determined by two coordinates yi, zj. The micro-volume DVij can be considered infinitely small, so the density of the grains distribution within a given volume q(yi; zj) is constant. Then, the number of vertices in the micro-volume determined (Eq. 1).
 Dnij ¼ q yi ; zj DyDz

ð1Þ

The cutting force of grain vertices Pij located in the micro-volume DVij was determined taking into account the probability of their contact with the treated material pk(yi;zj), which will also be deemed constant within the considered micro-volume (Eq. 2):


 Pij ¼ Pg yi ; zj
q yi ; zj
pk yi ; zj DzDy

ð2Þ

where Pg(yi; zj) is grain vertex cutting force. According to the data presented in [16, 17], the pattern of allowance removal is determined by line CO′A′ (see Figs. 1 and 2), having a contact probability level pk  0.0027. For each of the coordinates (yi, zj), the coordinate yi0 can be found, in

Fig. 2 Scheme of dividing contact zone into micro-volumes to build the cutting force model

914

V. A. Nosenko and M. V. Danilenko

which the probability of contact takes the value of pk  0.0027. Therefore, the actual grain vertex cut depth will always be less than the nominal value by the removed allowance: yif = yi − yi0, which is considered when determining the components of the grain vertex cutting force. Cutting force P reduced to the grinding wheel height unit (hereinafter referred to as the cutting force) at some point in time s is determined by the number of elementary volumes in the considered area of the contact zone (Eq. 3): P¼

n X l X i¼1 j¼1

Pij ¼

n X l X


 q yi ; zj
pk yi ; zj
Pg yi ; zj DzDy

ð3Þ

i¼1 j¼1

To find the coordinates of the cutting force application point, we use the data on the distribution of the elementary values of the cutting force in the selected radial layers of the abrasive tool and the treated material contact surface. The presented method of calculation allows to determine the resulting cutting force application point in the layer and in the contact zone. Its coordinates are determined by formulas:  P
Pi
zj Zi ¼ P ; Pi

P Yi ¼

ðP
y Þ Pi i Pi

ð4Þ

where i is the layer number. The resultant force in the micro-volume, layer and contact zone for the flat cut-in grinding is the geometric sum of its two components: the tangent Pz and the radial Py.

4 Practical Implementation of the Proposed Methodology As an example, consider the implementation of this method for the case of wheel periphery flat grinding. Initial data for the calculation: processed material—steel 45 GOST 1050 (HRC 42…45); grinding wheel—1 200  20  76 25AF60K6B 50 m/s 1 cl. GOST R 52781-2007. Cutting mode: grinding speed—37 m/s; table speed—12 m/min; feed to the depth of grinding—0.01 mm/stroke. The initial density distribution of grain vertices in layers of the grinding wheel working surface is determined according to the characteristics of the wheel and the billet material, as well as the adjustment conditions (see Fig. 3). During the grinding period, density distribution of the grain vertices can change as a result of wear of the grinding wheel working surface. Not all the grain vertices, which at some point of time in the grinding zone contact with the billet surface. Their number is determined by the probability of contact of the grain vertex pk with the treated material.

Determining Coordinates of Cutting Force Application …

915

Fig. 3 Density distribution of grain vertices n across the depth of the grinding wheel working surface u

Using the mathematical model presented in [16, 17], for the considered grinding conditions, considering the initial distribution of grain vertices, the probability of contact of the grain vertex with the treated material is calculated (see Fig. 4). The most protruding grain begins to interact with the treated material at a distance of z  −1.8 mm. With approaching the axial plane and increasing the cutting depth, the probability of contact increases, reaching 1 at z  −1.2 mm. The presence of a linear section (curve 1) to z  0.5 mm corresponds to the removal of the metal by the grain in the zone of no influence of the previous passes and the intrusion of the cutting vertex into the metal area untouched by other grains. The probability of contact is then reduced to near zero at z  1.42 mm. The contacting vertices are distributed unevenly in the grinding zone (see Fig. 5). In layers located at a distance of u  5 µm from the conventional outer surface of the grinding wheel (curves 1, 2, 3, 4), the number of contacting grain vertices is low, despite the high probability of contact (see Fig. 4). The reason for this is the low-density distribution of vertices in these layers, not exceeding 0.15 1/mm2 (see

Fig. 4 Change in the probability of contact pk along the trajectory of the grain vertex movement at different distances from the conditional outer surface of the grinding wheel to the layer under consideration: 1—0.5 µm; 2—1.5 µm; 3—2.5 µm; 4—3.5 µm; 5—4.5 µm; 6—5.5 µm; 7—6.5 µm; 8—7.5 µm; 9—8.5 µm; 10—9.5 µm; 11—10.5 µm; 12—11.5 µm; 13—12.5 µm

916

V. A. Nosenko and M. V. Danilenko

Fig. 5 Change in the number of grain vertices nk along the length of the contact arc projection on the Z-axis in different layers of the wheel working surface (see Fig. 4 for lines designation)

Fig. 3). The maximum number of contacting grain vertices corresponds to curve 5. With a further increase in the distance u, there is a decrease in the number of contacting grain vertices in the layer. The cutting force of the grain vertex Pg is determined taking into account the actual cutting depth according to the mathematical model proposed in [18, 19]. Pg values vary significantly, both along the trajectory of the grain vertex in the contact zone and along the depth of the grinding wheel working surface (see Fig. 6). By multiplying the number of grain vertices in contact at the microscopic volume Vij, the cutting force of the abrasive grain vertices, we find the total cutting forces Pij in the given micro-volume (see Fig. 7) [20]. As we approach the vertical axial plane and increase the distance from the conventional outer surface to the layer under consideration, the cutting force Pij increases (curves 1–5), reaching the greatest value along the trajectory of the curve 5 at z  –0.5 mm (see Fig. 7). The reason is that for this micro-volume, the product of the number of contacting vertices (see Fig. 5) and the cutting force of the grain vertex (see Fig. 6) is maximum. With further increase in u (curves 6–13), force Pij decreases. The obtained data on the distribution of the elementary values of the cutting force in the selected layers of the abrasive tool and the treated material contact surface are used to determine the coordinates of the resulting cutting force

Fig. 6 Change in the cutting force of the grain vertex Pg in the grinding zone along the layers of the wheel working surface (see Fig. 4 for designation of lines)

Determining Coordinates of Cutting Force Application …

917

Fig. 7 Changes in the total cutting force Pij in the micro-volume along the length of the contact arc projection Z-axis (see Fig. 4)

Fig. 8 Dependence of the total cutting force in the Pi layer on the coordinates of the Yi layer

application point in the layer (Eq. 4) (see Fig. 8). The graph is plotted in the following coordinates: horizontal axis—coordinate Yi of the cutting force application center in the ith layer; vertical axis—value of the total cutting force in such layer Pi. The total cutting force is shown for the first 12 layers. In the overlying layers, the total cutting force approaches zero and is not shown in the figure. The maximum cutting force is formed in the fourth layer. From layer 1 to layer 4, the total cutting force will increase rapidly then decreases gradually from layer 4 to layer 12. Thus, for the considered case, the maximum cutting force is formed at a distance of about 4 µm from the most distant grain vertex. Coordinates of the total cutting force application center are shown with a marker in the form of a circle in Fig. 9. The point of the total cutting force application is shifted along the Z-axis to the right to the central axis of the wheel, and along the Y-axis, it is at a distance of about 5 µm from the vertex of the most protruding grain.

918

V. A. Nosenko and M. V. Danilenko

Fig. 9 Coordinates of the total cutting force application center in the layer and in the grinding zone

5 Conclusion The proposed method allows us to investigate the nature of the cutting force distribution in the grinding zone along the layers of the abrasive tool working surface. For this example, the resulting cutting force application point in the layers of the wheel working surface is removed from the axial plane of the wheel to the left at a distance of about 0.2 Lmax. The resulting cutting force application point at the height of the grinding wheel working surface profile is at a distance of about 0.3 tф from the most distant grain vertex. The results of the study may be used to solve problems related to determining the intensity of heat generation in the contact zone of the grinding wheel and the billet, the calculation of cutting power, tool wear, and other issues related to the study of the grinding process. In order to automate the calculation of cutting force, special software for computers has been developed.

References 1. Balashov VN, Mishin VN (2011) The interaction of the circle with the part and the estimated estimate of the forces during grinding. Autom Ind 3:26–35 2. Dyakonov AA (2012) Development of a scientific and methodological base for improving the efficiency of abrasive machining processes based on a multifactor assessment of the workability of materials. Dissertation, South Ural state university 3. Gorlenko OA, Bishutin SG (1999) Model of the working surface of the abrasive tool. STIN 2:25–28 4. Maslov EN (1974) Theory of grinding metals. Mechanical Engineering, Moscow 5. Rechenko DS (2008) Investigation of cutting force at high-speed grinding. STIN 10:37–38 6. Hou ZB, Komanduri R (2003) On the mechanics of the grinding process. Part I. Stochastic nature of the grinding process. Int J Mach Tools Manuf 43:1579–1593 7. Nosenko VA (2000) Grinding of adhesion active metals. Mechanical Engineering, Moscow 8. Nosenko VA, Danilenko MV (2010) A mathematical model of the formation the working surface of the wheel during grinding. Tool and technology 27:64–67 9. Nosenko VA, Fedotov EV (2003) Theoretic probability model of the formation a working surface of an abrasive tool during grinding. Tool Technol 15–16:58–61

Determining Coordinates of Cutting Force Application …

919

10. Nosenko VA, Fedotov EV, Nosenko SV et al (2009) Probabilities of abrasive tool grain wearing during grinding. J Mach Manuf Reliab 38(3):270–276 11. Belov AG (2004) Development of mathematical models of cutting force components at grinding the stepped shaft edges by the method of direct profile wheel axial feed. Proc Chelyabinsk Sci Center 1:122–126 12. Nosenko VA, Danilenko MV (2017) Dynamic theoretical and probability model of cutting force in grinding. Izvestiya VSTU Ser Adv Technol Mech Eng 12:31–33 13. Ostrovsky VI (1981) Theoretical bases of grinding process. Publishing House of Leningrad University, Leningrad 14. Zubarev YM (2004) Mathematical description of the grinding process. Tool Technol 17– 18:55–56 15. Zubarev YM, Priemyshev AV (2010) Theory and practice of improving the efficiency of materials grinding. Lan Publishing House, Saint-Petersburg 16. Korolev AV (1975) Research of tool and part surface formation processes at abrasive treatment. Publishing House of Saratov University, Saratov 17. Novoselov YK (2012) Dynamics of formation of Surfaces in abrasive machining: monograph. Publishing house SevNTU, Sevastopol 18. Filimonov LN (1979) High-speed grinding. Mechanical Engineering, Leningrad 19. Korchak SN (1976) The performance of the grinding process of steel parts. Mechanical Engineering, Moscow 20. Danilenko MV (2018) Development of dynamic mathematical model of cutting force considering the influence of wear of the grinding wheel. Dissertation, Volgograd State Technical University

Improving Efficiency of Machining of Grooves on Shafts of Increased Hardness Structural Steel S. V. Grubyi and P. A. Chaevskiy

Abstract The method of calculating forces, temperatures in the cutting zone when turning grooves in workpieces of various structural materials with grooving cutters, is disclosed. As a general case, the standard groove cutters with soldered plates and the geometry of which corresponds to GOST 18884-73 are considered. The verification strength calculation of the cutting wedge is performed. The geometric parameters and strength characteristics of the tool material are justified. The wear-resistant coating to reduce temperature and increase tool durability was analyzed and selected. The conducted study has shown the ineffectiveness of the standard soldered turning groove cutters use at modern machine-building enterprises in the machining of workpieces of increased hardness structural steel. Design features and conditions for the use of built-up tool, equipped with changeable multifaceted plates (CMP) of the developed shape and size, are justified. The manufacture and subsequent implantation of the developed built-up cutters equipped with CMP of high-strength hard alloy will ensure an increase in the reliability of the tooling system of the machine-building enterprise and eliminate the breakage of the cutting tool and associated equipment downtime. The use of high-temperature wear-resistant coating on CMP will improve the performance of the machining of grooves on the workpieces of increased hardness structural steel and increase tool durability.








Keywords Grooving cutter Cutting forces Wear-resistant coatings temperature Built-up cutter Changeable multifaceted plate





Cutting

S. V. Grubyi Bauman Moscow State Technical University, 5/1, 2-Aya Baumanskaya St., Moscow 105005, Russia e-mail: [email protected] P. A. Chaevskiy (&) LLC “Company RITS”, 40, Bolshaya Semenovskaya St., Moscow 107023, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_98

921

922

S. V. Grubyi and P. A. Chaevskiy

1 Introduction Modern products manufactured by machine-building enterprises contain a large number of parts with grooves. These parts have a wide range of shapes and sizes and are made from various materials. At the majority of machine-building enterprises, technical measures are being taken to reduce costs and increase in productivity of output product. One of the ways to increase production efficiency is to use of progressive design tools. In particular, on parts such as shafts, the grooves are machined with solid, composite, or built-up hard alloy cutters. Tool companies offer tools of various designs for the machining of grooves. Thus, the catalogs of tool companies Iscar, SANDVIK Coromant, SimTec, Paul HORN, Carmex, WIDIA, and DENITOOL offer various cutter designs with changeable multifaceted plates (CMP) for the grooves turning. The tools of the companies under consideration have various design features and can be used in the conditions of machine-building enterprises. However, the main disadvantage of these tools is the relatively high cost, which, when it used, leads to an increase in the cost of corresponding part’s machining. Therefore, the issues of development, manufacture, and use of built-up cutters of domestic designs are relevant for those enterprises and companies that pose and solve tasks to reduce production costs and operating expenses. As a base case, the standard groove cutters with soldered plates, the geometry of which corresponds to GOST 18884-73, are considered. These cutters are used in some enterprises for this type of operations. The cutters of this design have a number of serious disadvantages: need for regrinding to restore cutting properties, which implicates the need for additional qualified personnel and specialized equipment; relatively low tool durability between regrints due to the inexpediency of using wear-resistant coatings on soldered plates; lack of special geometry of the front surface to ensure an optimal chip formation process; increased auxiliary time to change the tool in comparison with designs with CMP. In particular, it should be noted the absence of wear-resistant coating on the soldered plates. It is impractical to apply a protective coating on the soldered tool after each regrinding since in addition to restoring the geometry of the cutting wedge after wear, the application of a wear-resistant coating will significantly increase the tool manufacturing cost. Also, it should be noted that the soldered plates for the machining of grooves do not have holes. In some sets for the wear-resistant coating application, the absence of a hole on the plate complicates the process of coating due to the need to fix the plates in a special device and reduces the total volume of loaded plates in the chamber of set. When machining grooves on parts made of difficult-to-machine materials, in particular, of increased hardness structural steel, high temperatures occur in the cutting zone, that lead to a weakening of the hard alloy, resulting in intense wear of the cutting edges and an increased probability of breakage of the cutting wedge. To

Improving Efficiency of Machining of Grooves …

923

reduce the cutting temperature and increase the wear resistance of the tool material, the use of built-up tool equipped with CMP with wear-resistant coatings is recommended. RITS Company LLC is engaged in the implementation of efficient designs of cutting tools at the modern machine-building enterprises. In partnership with Moscow State Technical University named after N. E. Bauman, studies to justify the introduction of turning cutters of progressive designs with CMP and replacement of soldered cutters are conducted. The followings were the objectives of the study: • Development of the method and force calculating, temperatures in the cutting zone when turning grooves in workpieces of various structural materials with grooving cutters with a given geometry. • Verification strength calculation of the cutting wedge, justification of geometric parameters and strength characteristics of the tool material. • Selection of wear-resistant coating to reduce the temperature and increase wear resistance of the tool; justification of design, conditions of use, and implementation of built-up tool with CMP.

2 Calculation of Cutting Forces and Test Strength Calculation When machining difficult-to-machine materials, including increased hardness structural steel, the failures of the groove cutters in the form of failures due to insufficient tool strength occur. According to the results of the strength calculation, recommendations can be given on the geometrical parameters of the tool, properties of the tool material, and design features of the tool. The force interaction of the tool and the workpiece were considered by a number of authors [1–5]. It has been justified that in order to develop the design of grooving cutters with CMP, it is necessary to analyze the cutting forces and stresses arising in the cutting wedge of the tool and calculate corresponding safety factors. The fundamentals of modern calculation of cutting forces and stresses are reflected in a number of textbooks, monographs, and scientific articles. For example, in a generalized form with references to literary sources, calculation methods are given in the national textbook [6]. Foreign authors, in particular in the works [7, 8], also calculate cutting forces using similar algorithms. In the conducted research, for the calculation of cutting forces, it was used technique developed in Moscow State Technical University named after N. E. Bauman and cited in article [9]. The input variables for the calculation are the tool geometric parameters, cutting speed, cutting layer thickness and width, and known mechanical characteristics of machined and tool materials. Output parameters are: shear angle, chip shortening factor, relative shear, friction factors and

924

S. V. Grubyi and P. A. Chaevskiy

force component at front and rear surfaces of tool, and total force components in machine coordinate system. The developed technique makes it possible to calculate forces in a wide range of shear layer thicknesses. At the same time, in relation to the strength calculation of the groove cutters, the thickness of cutting layer, which exceeds the value of the cutting edge rounding diameter, is of interest. The following initial data and parameter values are taken in the calculations: • The material of the cutter cutting part is VK8 hard alloy, back angle is 8°, front angle is 10°; • Machined material is hardened steel 40X grade, HRC50 hardness; 2 mm wide groove, machined on the outer cylindrical surface of the workpiece. For comparison, the standard representatives for the group of carbon and low-alloy structural steels, and corrosion-resistant steels—grades 45 and 12X18H10T, respectively, were selected. The scheme of rectangular free cutting, for which the main component of cutting force Pz matches in direction with the cutting speed vector, radial component of force Py matches with the feed vector of the groove cutter. For this scheme, the feed value is equal to corresponding value of the layer thickness being cut. Figure 1 shows the diagrams of main component of cutting force for three structural steels considered dependence on thickness of cutting layer. Figure 2 shows similar dependencies of the radial component of the cutting force. The obtained dependences of the component cutting forces reflect significant influence of the hardness and strength of the material being machined and serve as the basis for the strength calculation of the tool.

Fig. 1 Diagrams of main component of cutting force dependence on thickness

Improving Efficiency of Machining of Grooves …

925

Fig. 2 Diagrams of radial component of cutting force dependence

The calculation of the cutting tool strength is made by the permissible stresses of the first kind. The stresses at the boundaries of the cutting wedge outside the contact zone are calculated using formulas borrowed from the works of M. P. Vadachkoriy and given in monograph [10]. The radial stresses on the front surface from the total action of normal forces and friction acting on the front and rear surfaces of the tool are analyzed. Diagrams of radial stresses dependences on thickness of layer being cut are shown in Fig. 3.

Fig. 3 Diagrams of radial stresses at front surface dependences on thickness of layer being cut

926

S. V. Grubyi and P. A. Chaevskiy

The strength condition on the front surface outside the contact zone takes the form: rr \rb ;

ð1Þ

those, the radial stress values should be less than tensile strength of hard alloy. The corresponding diagrams of safety factors for considered grades of machined steels are shown in Fig. 4. The cutter of hard alloy grade VK8 when machining hardened steel grade 40X has a safety factor of either slightly more than one for small values of cut thickness or less than one. The results suggest that the use of such cutters is ineffective for modern machine-building enterprises because an increase in feed decreases the safety factor that leads to an increase in the probability of tool breakage. According to the study results, the main necessity of replacing grade VK8 hard alloy with a modern hard alloy with increased strength was justified. In particular, it has been proposed to use A10 grade hard alloy manufactured by Kirovgrad hard alloys plant [11] for manufacturing CMP of built-up grooving cutters. The specified hard alloy has a strength of 2–2.3 times exceeding the strength of grade VK8 hard alloy. Increasing in strength of tool material in proportion will increase safety factor that will prevent the breakdown of groove cutters when machining increased hardness structural steel.

Fig. 4 Diagrams of safety factors’ dependences on thickness for various steel grades

Improving Efficiency of Machining of Grooves …

927

3 Justification and Selection of Wear-Resistant Coating Increased temperature in the cutting zone negatively affects tool durability. With a strong temperature effect on the hard alloy tool, it is softening and a significant decrease in durability occurs, and the probability of breakage increases. The heat-resistant wear-resistant coating can increase the temperature resistance. The criterion for choosing a coating will be its heat resistance and versatility of use. Having estimated theoretically maximum cutting temperature, the required heat resistance was determined. For the same materials and machining conditions, the temperatures occurring at front and rear surfaces of hard alloy tool were calculated. Temperature on the front surface is higher, therefore, we will consider it as the maximum possible under these modes. The calculation results in form of diagrams are shown in Fig. 5. The maximum design temperature is 1029 °C. Based on the calculations made and taking into account the capabilities of modern nanocomposite coatings for a significant increase in the machining efficiency, it is proposed to use PLATIT’s naCo-type coating [12]. Nanocomposite coating nACo® (on the base of Ti, Al) is a high-tech hard coating consisting of nanoparticles residing in a binder amorphous matrix. The nACo® is applied by physical way when using powder paints in one or several layers. The coating is unique due to the combination of physical parameters: at increase in hardness, its plasticity increases simultaneously. The basic physical characteristics of the coating are shown in Table 1. The selected coating has a thermal resistance above the calculated and universal application area.

Fig. 5 Diagrams of temperature at front surface of hard-alloy groove cutter

928 Table 1 Physical properties of wear-resistant coating nACo®

S. V. Grubyi and P. A. Chaevskiy Physical properties nACo® Micro-hardness (GPa) Thickness (µm) Frictional factor on steel Max. operating temperature (°C)

35 1.0–4.0 0.45 1200

Fig. 6 Model of developed built-up cutter

4 Design Features of the Developed Built-up Cutter After analyzing and justifying the need for using CMP with a wear-resistant coating, the line of turning grooving cutters was designed for RITS Company LLC. Based on the experience of foreign companies in design of groove cutters, the design that eliminates disadvantages of the previously used design is proposed. The model of the developed built-up cutter is shown in Fig. 6. The main features of the developed built-up cutter are the following: • Production of plates from standard CMP form workpiece. • Possibility of manufacturing of a plate with non-standard width of cutting wedge and also a shaped profile. • Plates for various operations can be mounted on one toolholder. • In case of damage to one cutting edge, all other edges remain operational. • Fastening the plate on the toolholder of increased rigidity and reliability.

5 Conclusion The conducted study has shown the ineffectiveness of the standard soldered turning groove cutters use at modern machine-building enterprises in the machining of workpieces of increased hardness structural steel. Recommendations on the use of modern hard alloy and wear-resistant coatings were developed. The design of the turning grooving tool with CMP was also developed. The manufacture and subsequent implantation of the developed built-up cutters equipped with CMP of high-strength hard alloy will ensure an increase in the

Improving Efficiency of Machining of Grooves …

929

reliability of the tooling system of the machine-building enterprise and eliminate the breakage of the cutting tool and associated equipment downtime. The use of high-temperature wear-resistant coating on CMP will improve the performance of the machining of grooves on workpieces of increased hardness structural steel, increase tool durability.

References 1. Zubkov NN, Ovchinnikov AI, Vasil’ev SG (2016) Tool–workpiece interaction in deformational cutting. Russ Eng Res 36:209–212. https://doi.org/10.3103/s1068798x16030217 2. Petrushin SI, Proskokov AV (2010) Theory of constrained cutting: chip formation with a developed plastic-deformation zone. Russ Eng Res 30:45–50 3. Kabaldin YuG, Kuzmishina AM (2016) Kvantovo-mekhanicheskoe modelirovanie deformatsii I razrushenia srezaemogo sloya pri rezanii (Quantum-mechanical deformation simulation and destruction of the cutting layer during cutting). Vestnik mashinostroeniya, Moscow 4. Grubyi SV (2014) Optimizatsiia protsessa mekhanicheskoi obrabotki i upravlenie rezhimnymi parametrami (Optimization of the machining process and control regime parameters). Optimizatsiia protsessa mekhanicheskoi obrabotki i upravlenie rezhimnymi parametrami, Moscow 5. Grubyy SV (2017) Raschet parametrov struzhkoobrazovaniya i sil rezaniya plastichnyh materialov (Calculation of chip formation parameters and cutting forces of ductile materials). Mach Plants Des Exploiting 1:25–37. https://doi.org/10.24108/aplts.0117.0000058 6. Rosenberg YuA (2007) Rezanie materialov (Cutting materials). Rezanie materialov, Kurgan 7. Toenshoff HK, Denkena B (2013) Basics of cutting and abrasive processes. Springer, Berlin 8. Ravi Shankar M, Verma R, Rao BC, Chandrasekar S, Compton WD, King AH, Trumble KP (2007) Severe plastic deformation of difficult-to-deform materials at near-ambient temperatures. Metall Mater Trans A 38:1899–1905 9. Grubyi SV (2018) Calculation of the cutting forces when processing plastic materials with a wide range of thicknesses of the cutting layer. https://doi.org/10.18698/0536-1044-2018-2-310 10. Loladze TN (1982) Prochnost’ i iznosostojkost’ rezhushchego instrumenta (Strength and wear resistance of cutting tools). Prochnost’ i iznosostojkost’ rezhushchego instrumenta, Moscow 11. Jsc KZTS (2017) TU 48-4205-81-2017. JSC KZTS, Kirovgrad 12. Platit AG (2018) PLATIT COMPENDIUM ev60. Ideal-lab, Bern

General Patterns in Formation of Surface Layer of Machine Parts Treated by Combined Electro-technical Methods S. V. Usov, P. A. Davydenko and D. S. Sviridenko

Abstract The humanity has always needed new products and shorter time to incorporate them. These needs can be met by new technological processes based on a non-mechanical impact of an instrument on a blank, for example, on combined electro-technical methods. Combined electro-technological methods are the most popular in the technologies utilized in machine parts surface treatment based on sophisticated processes of metal dissolution. Rig testing allowed identifying certain physical and mechanical patterns in the increase of durability under the condition that strengthening and sizing combined electro-technical methods were used. In order to meet the requirements of the test, the suitable equipment, technological gear, and instruments were used, and the electrolyte environment class was ensured. The test confirmed the possibility of combining different methods of machine parts treatment that endures high tribological pressure. The best combinations and regime support were researched and identified. Keywords Machine parts Tribology Wear





Combined treatment
Electrochemical treatment


S. V. Usov Technological Systems for Protective Coats LLS, 27, the 16th Parkovaya, Moscow 105484, Russia P. A. Davydenko Federal State Autonomous Educational Institution of Higher Education, RUDN University, 6, Miklukho-Maklaya Str., Moscow 117198, Russia D. S. Sviridenko (&) Federal State Unitary Enterprise All Russian Scientific Research Institute of Aviation Materials, 17, Radio Street, Moscow 105005, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_99

931

932

S. V. Usov et al.

1 Introduction In order to solve the most difficult technological problems in the field of designing new machines, modern methods of treatment of materials are used and developed, based on non-mechanical impact of an instrument on a blank. In most cases, spatially localized electromagnetic fields with high-density power flux are utilized to implement these methods. All these treatment methods are called electrochemical machining (ECM). ECM is most popular in dimensional processing technologies (sizing) of working surfaces of machine parts, based on sophisticated electrochemical processes of metal dissolution that in comparison with cutting is a technologically difficult task to handle. That is why ECM machine parts help to increase the accuracy of machine parts processing and thus enhance the reliability of a product as a whole [1, 2].

2 Formulation of the Problem Experimental research sought to identify the patterns in forming quality parameters of the surface layer by specifying favorable ranges of combined electrotechnological methods regimes. The results of the rig testing (tribological, thermo-cycle, alternating cyclic testing, friability testing of electrolyte coating) were used as a performance assessment of best regimes that provide approximate initial conditions under which typical machine parts function. In this regard, the following connections between technological methods and durability parameters were identified [3]: (a) combined consecutive method of electrochemical (fixed cathode-instrument, electrochemical drawing) and electrolyte chromium coating; (b) combined parallel-consecutive method of diamond electrochemical honing and electrolyte chromium-coating; (c) combined consecutive method of shot peening and electrochemical polishing; (d) combined consecutive method of impulse-cycling sizing and shot peening.

3 Research Results Let us consider the nature of the connection between quality parameters and the produced results in enhancing durability achieved during the rig testing in the identified best range of the regimes of electro-technological combinations [4]. Securing durability under the conditions of thermo-mechanical wear and cyclic alternating loading is influenced by a number of special characteristics of the

General Patterns in Formation of Surface Layer …

933

formation process of quality parameters complexes. Higher micro- and macro-geometrical accuracy indexes ensure better durability up to 15% and at the same time, the friability of electrolyte coating falls by 0.5–7% and that provides for better durability of parts when the following methods are used: diamond electrochemical honing and electrolyte chromium coating; electrochemical sizing and electrolyte chromium coating. These combinations implement all the three identified methodological principles of combination (superposition, exclusion, improvement). The implementation of the first principle (diamond electrochemical honing) enhances the durability by 5–10% by forming adhesive-cohesive wear mechanisms with the best zone of regime maintenance creating less extensive areas of cohesive wear that can be explained by stronger effect of strengthening that is achieved thanks to the chosen best value of the time of diamond electrochemical honing burnishing. The implementation of the improvement principle (diamond electrochemical honing and electrolyte chromium coating) helps to considerably reduce the wear by 10–20% and to altogether exclude adhesive and cohesive effects [5]. In this case, the main specific features of the fracture patterns are the areas of warm gray background with the thicker electrolyte coating (up to 70 lm) helping to form a crack on the surface of the sample, that is, within the zone of best regimes that ensure lower levels of electrolyte coating friability and better hardness, and the emergence of a compression curve (diamond electrochemical honing). The implementation of exclusion and improvement principles under the conditions of the development of the consecutive method of electrochemical sizing (fixed cathode-instrument, electrochemical drawing, electrolyte chromium coating) excludes the possibility of the development of macro- and micro-effects of the surface layer. It is estimated to be 10–15% and lowers the friability of the electrolyte chromium coating by 0.5–1% and thus enhances the durability of the samples by 5–10%. It must be said that thicker electrolyte coating (up to 200– 220 lm) causes numerous cracks and chips during tribological tests provoked by plastic edging effect under the conditions of tribological interaction. The reliable functioning under the conditions of cyclic alternating loading, shock, and fatigue wear is made possible by two mixed consecutive combinations [6, 7]: • electrochemical impulse-cycling sizing (ECM imp.) and shot peening (SP); • shot peening (SP) and electrochemical polishing (ECP). The formation of the first mixed combination makes micro- and macro-geometry lower by 10–15%, enhances hardness by 45–50%, and forms a compression curve with the growth up to 40–50%. At the same time, the second combination enhances micro-geometrical indexes by 10–15%, hardness up to 75% and forms a compression curve that augments the parameters up to 60–75%. This change in the researched complex of quality parameters provides for better durability under the conditions of cyclic alternating loading up to 10–30%. Higher levels of thermo-cyclic durability have not been detected. The applied principles (exclusion and combination) are also quite specific for the aforementioned combinations of technological methods under the conditions of tribo-technical rig tests. The effect of

934

S. V. Usov et al.

the combination of ECM and SP as a hardening combination of SP and ECP under the conditions of tribological tests feature the mechanism of cohesive and adhesive wear. Moreover, the shift of the wear mechanism to adhesive for the hardening combination and vice versa, the cohesive effect pertains to the sizing combined method.

4 Technological Support (Equipment, Gear, Instruments) In order to use the developed combinations of electro-technological methods, it is crucial to design the relevant equipment, technological gear (gadgets), instruments, the relevant environment class, electrolytes and to optimize the regimes. This said that the list of implemented technical combinations provides for the aforementioned composition of quality parameters for the surface layer. Technological treatment for the durability of machine parts has the following sections [8]: • the development of customized high-capacity quick-setting equipment; • the creation of multi-purpose reliable instrumental gear (devices) that ensures the required quality parameters; • the design of customized durable instruments for applying combined electro-technological methods; • the creation of the relevant electrolyte environments with special qualities that ensure the necessary macro- and micro-geometrical accuracy; • the creation and design of support equipment, the relevant paperwork to introduce sections of electro-technical equipment; • the optimization of regimes of combined electro-technological methods that help to implement the required list of quality parameters providing for the necessary durability level of machine parts. The equipment used for electrochemical methods of drawing and processing by a fixed cathode-instrument [9, 10]: this method is applied at the series-produced plant “ЭXC-1У” and the method using the fixed cathode method is implemented at a special plant. The equipment for the combined method of diamond electrochemical honing [11, 12]: the designed equipment for the combined method of diamond electrochemical honing includes modernized vertical honing semi-automatic machines, models 3M83, 3К82У, 3821. The equipment is used based on the principle of incorporation into the operating manufacturing plan and creation of a new special section. In both cases, the equipment has the necessary ventilation system, electrolyte delivery system, cleaning and passivation facilities for machine parts. The facilities have operating chambers made of plexiglass that allow to monitor the process and minimize electrical leakages. The hatch of the chamber is rotating on vertical hinges and rubber-sealed. The electrolyte that trickles down the hatch when

General Patterns in Formation of Surface Layer …

935

it opens, and collects in a drainage conduit and is discharged. In order to make the change of the gear easier in the walls, the electrolyte is delivered to a processed part from below. That is why a device that stops the electrolyte flow from the chamber to the place, where the instrument is withdrawn from it, is necessary. Good results were produced when for this purpose a hydrodynamic seal was used that stopped the electrolyte from flowing up from the working zone by a support flow that forms an 110°–130° angle to the first one. In order to protect the equipment from accidental spill of electrolyte that has not been stopped by the hydrodynamic seal, a readily removable rubber sleeve limiting the space between the hydraulic seal and the spindle head is used. The goffered rubber sleeve is readily removable with the bonnet being attached to brush carrier case by a bayonet lock so that the sleeve does not block the access to the attachment fitting of the tool in the socket. The proposed design of the seal effectively protects the equipment from electrolyte and allows to use instruments with various diameters. In order to fix the instrument in the spindle and to supply electricity, a joint socket is used that is connected by flexible conductors to the collector that operates in contact with six copper–graphite brushes. So as to lower the heat regime in the current feedthrough, there is a circular cavity in the brush carrier case with water in it. The “flotation” of the half coupling that fixes the instrument is possible because it is connected with the half coupling with a swing joint. The reequipped machine ЗК82У has a different design solution for supplying electricity to the instrument. This machine has a revolving quill with a moving spindle inside instead of a movable spindle head. That is why the brush device operates in steam friction with the spindle that is in seesaw and rotary motion. Two customized graphite brushes are used to supply electricity to the spindle. They are cooled by water that flows in canals inside the brushes and is supplied via connecting tubes. The brushes are attached to the brush carrier and are pressed to the spindle by springs. Copper plates that deliver electricity from the power source are attached to the brushes by screws. Electrically isolated from the machine Plate 1 made of 30X13 steel serves for installation, deployment, and fixing the device. The plate is attached to the machine table by interposed steel supports and insulation liners that go through the openings in the bottom of the chamber. The gap between the plate and the bottom of the chamber is filled with PVC paste. It is not recommended to install the plate directly on the bottom of the chamber because of the difficulty of its exposure and lower structure rigidity. Boreholes for the screws on the fixed plate have epoxy glue in them. A filler neck for electrolyte supply is attached to the bottom of the plate, and it goes through an opening cut in the table. Electricity is supplied to the blank by a separated contact element of the device that is connected by a flexible conductor through the wall of the chamber to the power source cable. It was essential for diamond electrochemical honing with considerable sizing defects to improve the system of radius supply of honing bars that 3821 machine

936

S. V. Usov et al.

has. The management of bars supply on the machine is based on the principle of current monitoring that the rotational electric drive of the spindle consumes. When the machine is turned on, the accelerated separation of bars starts. In this case, the consumed current is on its minimal level. When bars touch the blank, the current gets stronger and the signal from the shunt on the motor phase becomes strong enough for the polarized relay to work. The relay closes the circuit with its contacts that then blocks itself and signals the transition from accelerated to measured graded radial feed of the bars that marks the beginning of the honing process. In case, the load during the honing process gets higher up to “overload”, the relay goes off and signals to stop feeding the bars. The closed contacts connect to one of the reference winding of their relay that is adjusted to the measuring winding to make the reset coefficient higher. It allows the relay to turn off when the signal differs from the switch signal and to continue honing with dosed supply after overload was eliminated and the load is close to normal. However, when blanks with considerable sizing errors in the honed opening are processed, the said radial supply system does not work. If the transition from accelerated to dosed feed was made when the honing head processed an area with a big diameter, the instrument jammed. In order to solve this problem, a technical solution was proposed and implemented based on the change in cyclograms of the transition from accelerated to dosed feed of bars. The solution was to introduce the negative reverse connection of the location of the bars with the overload. In case of an overload, the circuit of dosed feed goes off and breaks and an additionally introduced circuit closes that signals to execute an accelerated retraction (“bounce”) of the bars from the processed surface. The “bounce” is executed in the way that is enough to eliminate the overload and is adjusted by a resistor setting. Then, the machine goes back to normal that allows to continue diamond electrochemical honing without the instrument getting jammed. The centralized system of electrolyte supply includes a divided tank with current dampers and a coiler to cool the solution, two sedimentation tanks, four stainless centrifugal pumps with fine filters and a communication system. The purification of electrolyte is executed by a mixed method: by filtration, floating, and sedimentation. The floating reagent is a technical cleansing agent 0.02–0.4% of which is added to the electrolyte, and the slime that comes to the surface is eliminated at the end of the shift. Every pump in the system supplies the machines with electrolyte. The output pressure in the pumps is adjusted by valves and is monitored by manometers. The level of filter clogging is identified by the difference in the manometers’ indications. The valves can cut the supply of electrolyte to the machines and adjust it. The electrolyte is discharged from the operating chambers into the tank via pipes. After non-floatable slime has accumulated in the tank, the contaminated solution is pumped from the tank into the empty sedimentation tank. The tank is filled with the purified electrolyte from the sedimentation tank by gravity when the valves are opened.

General Patterns in Formation of Surface Layer …

937

The main electrolyte supply system data: • • • •

volume of the tank: 2.5 m3; volume of the sedimentation tanks: 2 items  2.5 m3; one pump feed: 200 L per minute; electrolyte pressure: 6 MPa.

Machines connected to the centralized electrolyte supply system and the power source form devices that allow to operate in manual and semi-automatic regimes. Manual regime is intended for adjustment and experimental operations that allows to: (1) turn on and off: • • • •

seesaw and rotary motion of the spindle; rotation of the spindle; electrolyte supply; electricity supply.

(2) adjust the system of active control for the set size (only 3821 machine). The equipment for electrochemical method impulse-cycling sizing [13]: the plant of the model “ЭXП-I” is used for impulse-cycling sizing. The device includes: the machine itself; pumping stations for electrolyte supply; electrolyte storage tanks, power source serial number “BAК 3200” modernized to be used with an impulse circuit. Constructively, the machine has a chamber; a vibrator; supply device (by moving the carriage on a router table); a stand; management system; table supply control system. Functionally, it ensures the movement of the cathode-instrument, the lifting and lowering of the operating chamber, the transportation of electrolyte to the processing zone, and the discharge of electrolyte from the operating chamber to the buffer tank. The motor driver with a capacity of 4.5 kW sends operating vibration to the instrument. Accelerated and operating movements are powered by a motor driver with a capacity of 1.6 kW. The electrolyte supply system consists of a main tank. In order to protect the socket against electrolyte that is under excessive pressure in the processing zone and can spill through the rod, the joint honing head case-expanding cone is strengthened by a rubber ring. The basic target value when designing the honing head is the length of its cathode area. Mechanically fixed insulators are most durable as they are made of plastics that do not change their qualities in electrolyte: caprolon, fluoroplastic. Equipment for the combined method of shot peening and electrochemical polishing: equipment for electrochemical method of impulse-cycling sizing that is used for the experimental selection of regimes for the considered method with a special shot peening device and a conventional tank for electrochemical polishing with a power source.

938

S. V. Usov et al.

Technological support of combined consecutive method of electrochemical sizing (fixed cathode-instrument, electrochemical drawing) and electrolyte chromium coating [14, 15]: combined method of electrochemical sizing with a moving and fixed cathode-instrument provides the required quality parameters by using custom design electrode-instruments, optimized regimes that in combination with further technological method of electrolyte chromium coating (9–70 lm) helps to ensure the necessary durability level. Moreover, the durability of the parts has been enhanced by 20–30% [16]. Let us consider the design and technological differences of the developed cathode-instruments and ranges of optimized regimes of the mixed methods. In order to exclude the possibility of macro-defects, it is suggested that an electrode-instrument should be used with an inner canal to supply electrolyte that has guide and insulator bushes that are fixed on a rod. There are lengthwise spline grooves that form ledges that mirror the form of pits. A head that is a cylinder with radial openings is installed in the zone of radial output electrolyte canal that opens into the dissolution zone. The instrument operates in the following way: the instrument is installed with the help of a cathode-holder when the cathodeinstrument is fixed in this position, the pumps are turned on and electrolyte is supplied to the sizing zone and is then discharged via tanks. After that, the power source is activated. Electrolyte goes through the head by the rod canal into the zone of anodic dissolution. As the results of field tests showed, certain geometry of cutting inside the pipe ensures the required durability and enhances it [17, 18]. This is why a custom design of cathode-instrument was developed that ensures the required geometry, this said, in terms of design the cathode-instrument consists of a brass rod, caps, insulating plates that are cast under pressure from polyamide-type material. The brass case has oriented spiral grooves corresponding to the geometry of the cutting being made. Processing of the said geometrical profile is done on a specialized device. The cathode-instrument moves inside the processed part with the help of a special rod fixed on the cathode near the front spindle head. The relevant electrolyte supply system supplies it when the technological power source is turned on. The formation of cutting profile is done when the cathode rotates and former finger moves on the template at the same time. Moreover, the designed cathode ensures the required geometry of the cutting because the insulation inserts are made as a flexible beam that is fixed at both ends with a deflection that stabilizes the distribution of the electrolyte flow as required by the width of the sizing cutting. Custom-made cathode allowed to change the values of inner and outer cutting radius and width and provided the required level of durability of the parts by changing current density. Optimization of the regimes of electrochemical drawing of parts like “body of revolution” together with the electrolyte coating of corresponding thickness promotes durability stabilization. Technological support of combined parallel-consecutive method of diamond electrochemical honing and electrolyte chromium coating: diamond electrochemical honing (together with technological method of electrolyte chromium coating) of machine parts, for example, of bodies of revolution, enhances durability by 10%. This required level of durability helps to achieve better performance of the surface layer quality and improves the endurance of the electrolyte chromium coating that

General Patterns in Formation of Surface Layer …

939

is applied later. In case of diamond electrochemical honing, the aforementioned machines use special instruments and devices and their design is unified and interchangeable. The instrument is a honing head with a fixture. The fixture consists of a brass rod with epoxide insulation coating reinforced by kapron thread. Stainless steel to manufacture the rod is not recommended as it has low conductivity. The rod is installed into the socket with its cone shank and fixed. The pusher connects the bar supply mechanism with the expanding cone of the honing head. Honing heads for diamond electrochemical honing of through-holes have two versions: the second one differs from the first in the design of blocks that allows to use unified bar holders. In order to protect the socket against electrolyte that is under excessive pressure in the processing zone and can spill through the rod - the joint honing head case-expanding cone is strengthened by a rubber ring. The basic target value when designing the honing head is the length of its cathode area. Mechanically fixed insulators are most durable as they are made of plastics that do not change their qualities in electrolyte: caprolon, fluoroplastic. Insulation coating made from thermosetting material on the basis of phenolformaldehyde resin and glass fiber as a filler is very durable [19]. The coating is applied by the method of injection molding and has bending strength of at least 120 MPa and impact resistance not less than 30 MPa-mm. In order to improve the adhesiveness of press materials with the part, it is recommended to make double knurling and drill holes. From the point of view of friction and wear theory, it is preferable to make gap stops from high hardness materials. However, the attempts to manufacture the stops using oxide ceramics and polycrystal superhard materials were unsuccessful, as when they came into contact with the processed surface scratches appeared and surface finish was affected. This is why honing heads have plastic stops. The normal operation of the honing head depends on the production accuracy and materials employed. Strength elements of the honing head (case, expanding cone, wedge-type column) are made of heat-treated (with low temper hardening) steel 30X13, and the installation of movable joints is executed in accordance with 7–8 accuracy degree [20]. Devices for diamond electrochemical honing are mechanical, and their design takes into account the specific details of their operation. During downtime, saline deposits form in the gaps of movable joints. In order for the devices to remain movable, accurate running fits are not used and the area of the coupling face is assumed to be minimal. As the test showed the surface of the contact “power supply circuit–blank” that is in electrolyte is prone to anodic dissolution. This is why: the current is supplied to a blank through an easily replaceable special contact element that is part of the device; the possibility of electricity leakages is eliminated where the device contacts the blank, especially when it comes to basing surfaces. If the blank has holes or cracks that cross the processed surface, plugs are used in the device that stops the leakage of electrolyte. Thanks to the easy process of equipment readjustment (the instrument and the device is simply changed and cam controls of the seesaw and rotary motion of the hone are installed), simple

940

S. V. Usov et al.

manufacturing process of new gear, the section of the machines is suited for serial and small volume manufacturing, change of stock lists and processed parts. Technological support of the combined method of electrochemical impulsecycling sizing and shot peening: this combined method is allowed to enhance the durability of parts by 10–15%. The equipment necessary to implement this method is discussed above. Technological support of the combined method of electrochemical polishing and shot peening: this method allowed to enhance the durability of parts by 20–30%. Shot peening method is implemented by shots made from 1 class wire National Standard 9389-75. Electrochemical polishing of the aforementioned parts is done in a tank with electrolyte solution. As a rule, the part is placed into a special device and is clamped with the cathode being oriented strictly against the inner surface of the spring. At the same time, the surface of the cog of the rammer is located so that to ensure the implementation of the electrochemical polishing method.

5 Conclusion Thus, the rig testing allowed to identify certain physical and mechanical patterns in enhancing testbed durability when strengthening template combined electrotechnical methods are used. Moreover, best matches of the identified combinations’ regimes were discussed and formed “Chart 1” (Table 1)[21].

Table 1 Regime support for the combined methods №

Method name

Best regime support

1

Combined consecutive method of electrochemical generation of geometry (fixed cathode-instrument, electrochemical broaching) and galvanic chromium coating Combined parallel-consecutive method of diamond electrochemical honing and galvanic chromium coating Combined consecutive method of shot hardening and electrochemical polishing

Current density—30–40 A/cm2, voltage—14–20 V, time—900–1200 s, electrolyte—10–15% solution, the speed of the moving cathode-instrument—186– 225 mm/min

2

3

4

Consecutive combined method of electrochemical impulse-cycling generation of geometry and shot peening

Current density—20–30 A/cm2, Pressure per unit of area of the honing instrument—0.4–0.5 MPa, Pressure of the electrolyte—0.6–0.8 MPa, voltage—14–20 V, time—90–200 s, electrolyte—10–15% solution Time of shot hardening—400–800 s, Time of electrochemical polishing—up to 480 s, Current density during electrochemical polishing—0.6– 0.7 A/cm2, voltage—14–20 V Time—600–900 s, ECM current density—35– 40 A/cm2, ECM voltage—8–12 V, ECM electrolyte —10–15% solution, Time of shot hardening—1080– 1200 s

General Patterns in Formation of Surface Layer …

941

References 1. Usov SV, Vyacheslavova OF et al (2010) Untraditional methods of machine parts treatment based on information technologies. Slavyanskaya shkola, Podolsk, p 150 2. Popov AP, Popova TA et al (2013) Reliability enhancement:methods. Bull Moscow State Open Univ Mach Technol Ser 4:32 3. Popov AP, Popova TA (2015) Technological processes optimization. World Transport 5:16– 25 4. Usov SV, Polyakov SV et al (2011) Scientific basis for the development of progressive methods of machine parts treatment. Slavaynskaya shkola, Podolsk, p 209 5. Usov SV, Sviridenko DS et al (2012) The research of the influence of combined treatment methods on tribological characteristics of the surface layer of machine parts. Bull Voronezh State Tech Univ 5:138–141 6. Boldyrev AI, Usov SV et al (2014) Experimental research of technological possibilities of electro-chemical sizing. Bull Voronezh State Tech Univ 10(3–1):120–122 7. Usov SV, Vyacheslavova OF Theoretical basis for the implementation of combined electro-technological methods of machine parts treatment. Tech Res J 2(6):3 8. Albagachiev AY, Ambrosimov SK et al (2015) Progressive machine-building technologies, equipment and instrument, vol VI. Spektr Publishing House, Moscow, p 464 9. Bakhvalov VA (2006) Methods of honing deep holes and the instrument to apply it. STIN 4:39–40 10. Zvontsov IF, Serebrenitsky PP et al (2016) Technological methods of quality management in the production of parts with deep holes. Metal Process 4:11–25 11. Batyshev KA, Sviridenko DS et al (2018) Design of manufacturing technology for machine parts using diamond electro-chemical honing and electrolyte chromium coating. In: Prospects of processing technologies and equipment development in machine building: abstract. National scientific and technical conferences with an international element, Kursk, pp 30–33, 15–16 Feb 2018 12. Gavrish AP, Roik TA et al (2014) New designs for honing over-alloyed composition materials in friction parts of printing machines. Bicник Xмeльницькoгo Haцioнaльнoгo yнiвepcитeтy. Texнiчнi нayки 6:39–43 13. Smolentsev VP, Boldyrev AA (2010) Magnetic and rheological liquids in metalwork. Untraditional processing methods: Interuniversity collection of research papers. Is. 9. Machine-Building, Moscow, pp 120–129 14. Kuzovkin AV, Smolentsev VP (2000) Dimensional sizing of parts with complicated profile using hard electrically conductive filler. VSTU Publishing House, Voronezh, p 176 15. Sviridenko DS, Batyshev KA et al (2018) Calculation of instruments for electro-technological methods of machine parts processing. In: Prospects of processing technologies and equipment development in machine building: abstract. National scientific and technical conferences with international element, pp 297–299, 15–16 Feb 2018 16. Zvontsov IF, Serebrenitsky PP Electrical diamond honing of deep holes. http://mirprom.ru/ public/elektroalmaznoe-honingovanie-glubokih-otverstiy.html. Accessed 22 Nov 2017 17. Ryabtseva AV, Sviridenko DS (2015) The basis of forming small diameter holes through electro-technological treatment and their control. Inform Technol Design Prod 4:55–57 18. Domnin PV, Timofeeva AA (2016) The impact of various methods of surface treatment in deep holes on surface finish index. Modern Mach Technol 3:16–20 19. Bilgi DS, Jain VK (2004) Electrochemical deep hole drilling in super alloy for turbine application. J Mater Process Technol 149(1–3):445–452 20. Pa PS (2007) Electrode form design of large holes of die material in ultrasonic electrochemical finishing. J Mater Process Technol 192–193:470–477 21. Usov SV, Sviridenko DS (2007) Electro-chemical methods of machine parts treatment. Saturn-S, Podolsk, p 122

Evaluation of Tool Life Equation of Single-Point Cutting Tool by Accumulation Model A. V. Antsev, N. I. Pasko and A. V. Khandozhko

Abstract The effective use of modern machines with numerical control is impossible without a stable cutting process. An unexpected tool failure leads to high production costs. Various models are proposed to predict the wear value and tool life of the cutting tool. In this paper, the tool life equation is understood as a set of the law of distribution of the tool life and the parameters of the law, depending on the cutting mode. It is assumed that the spread of the tool life is associated with the spread of hardness and the allowance for the machining of workpieces. The case of asymptotically normal wear distribution for a given operating time is considered. The method of the assessment of the tool life equation parameters by the maximum likelihood method by statistics “speed-feed-depth-operating time-wear” is offered. The average intensity of the wear depends on the parameters of the cutting mode in the form of an exponent from the polynomial which is not higher than the third power of the logarithms of the noted parameters of the cutting mode. The method is illustrated by a numerical example with the statistics “speed-feed-operating time-wear”. It is shown that the average tool life calculated by the proposed method differs from the mathematical expectation by an amount of 0.4–5%, so in practical calculations, you can use a simpler formula to calculate the mathematical expectation of the tool life.










Keywords Tool life Wear Cutting tool Tool life equation Normal distribution Maximum likelihood method





Wear rate


A. V. Antsev (&)
N. I. Pasko Tula State University, 92, Lenin Ave, Tula 300012, Russia e-mail: [email protected] A. V. Khandozhko Bryansk State Technical University, 7, Blvd. 50 October, Bryansk 125993, Russia © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_100

943

944

A. V. Antsev et al.

1 Introduction The development and improvement of metalworking technologies is inextricably linked with the use of numerically controlled machines. Their effective use is impossible without ensuring a stable cutting process, which is evaluated comprehensively on the following criteria: the accuracy, the finished surface quality, and the reliability of the cutting tool. Modern machine-building enterprises are characterized by a very wide range of used cutting tools [1]. The cutting tool is the most vulnerable link in metal-cutting systems. An unexpected tool failure leads to high production costs due to the increased costs of prevention and maintenance of cutting tools and the increased costs of repairable and irrepairable reject due to the high costs of workpieces for finishing operations, so a large number of papers in the modern literature are devoted to the study of the wear mechanism of cutting tools, the diagnosis of their conditions, and the prediction of their tool life. When diagnosing the state of cutting tools, the methods which use the machine learning methods [2] are widely used. For this purpose, various indirect parameters of the cutting process are used: acoustic emission [3], vibration [4], temperature [5], cutting force [6], as well as their combinations [7]. The optical control is also used to control the condition of cutting tools [8]. However, despite scientific research, an extensive industry experience in this area and the economic importance of the cutting operations, the cutting process needs in-depth study, which is confirmed by a low predictive ability of the known cutting models and their efficiency only in relatively narrow and constant operating conditions [9]. Various models are proposed to predict the wear value and tool life of the cutting tool. For example, De-Jun Cheng et al. propose an approach for calculating the wear overlap (WO) geometry [10]. Anton Panda et al. use Taylor’s equation to predict the cutting tool life [11]. Also there are known approaches using the finite element method (FEM) [12], the partial least-squares regression (PLSR) [13] and a general model of wear [14]. However, these models do not take into account the stochastic nature of the cutting tools wear, which depends on a large number of factors: cutting modes, cutting properties of the tools, type of machining, hardness of machined parts, value of machining allowances, pre-existing mode of deformation, vibration, geometric errors of machine, etc. [15–19]. Therefore, in this paper, the tool life equation is considered in a generalized form as a set of the distribution law of the cutting tool life T and the dependence of the parameters of this law on the cutting mode [20, 21]. When turning the mode parameters are the cutting speed V, the feed S, and the depth of cut h. The tool life is mean time between failures (a dulling or a breakage of the cutting blade). Here, we consider the case when the reason for the spread of the tool life is the spread of hardness, machining allowance, and other parameters of workpieces [20].

Evaluation of Tool Life Equation of Single-Point …

945

2 Materials and Methods The failure of the cutting tool is considered to have occurred if the wear of the cutting tool Y from the moment of installation after processing a batch of T workpieces for the first time exceeds the maximum permissible value L. If DYi is the increment of the wear after machining of the ith workpiece, then after P machining the T workpieces the total wear of the cutting tool is Y ¼ Tt¼1 DYt . Because of the marked spread, the DYi wears are random variables and their sum after machining t workpieces according to the central limit theorem has an asymptotically normal distribution with the density of distribution " # 1 ðy  Y
t Þ2 ft ðyÞ  pffiffiffiffiffiffiffiffiffiffi exp  : 2Dt 2pDt

ð1Þ

Formula (1) with practice-relevant accuracy can be used when t [ 10. The average wear after machining t workpieces is Y
t ¼ u
t;

ð2Þ

and the dispersion of the wear, if the increments DYi are statistically independent, is Dt ¼ r2 t;

ð3Þ

where u is the average wear on one workpiece (wear rate), a r is the standard deviation of the wear on one workpiece. If there is a correlation between the increments, then the dependence of Dt on t is more complicated and the formula (3), in this case, must be considered as an asymptotic formula for t. Taking into account the formulas (2) and (3), we obtain that " # 1 ðy  u
tÞ2 ft ðyÞ ¼ pffiffiffiffiffiffiffiffiffiffiffiffi exp  : 2r2 t 2pr2 t

ð4Þ

The probability of non-failure operation during the operating time, that is, the probability that YðtÞ\L is ZL PðtÞ ¼ 0

" #     1 ðy  u
tÞ2  Lu
t  u
t p ffi p ffi pffiffiffiffiffiffiffiffiffiffiffiffi exp  dy ¼ U  U ; 2r2 t r t r t 2pr2 t

ð5Þ

Rx where U ðxÞ ¼ 1 p1ffiffiffiffi expðx2 =2Þdx is the cumulative distribution function of 2p the normal When t [ 10 you can use the approximation  distribution.   Lu
t pffi : PðtÞ  U r t

946

A. V. Antsev et al.

The mathematical expectation T
and the coefficient of tool life variation KT are calculated using the following formula T
¼

vffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi u Z1 u u
KT ¼ t2 PðtÞtdt  T
2 =T:

Z1 PðtÞdt; 0

ð6Þ

0


the average tool life, Along with the mathematical expectation of the tool life T, which is calculated by the formula (7), may be of practical interest: ^

T ¼ L=u:

ð7Þ

^

It should be noted that T
and T are close in value, but do not coincide. When optimizing the cutting modes in the case of tool replacement on failure, we should
use the value T
since the costs per unit in this case are directly dependent on T. Since the tool life T is a random variable, then to determine the tool life equation, we must also know in addition to the distribution law T the dependence of the parameters of this distribution on the cutting modes. In this case, you should know the dependence of the parameters r and u on the cutting mode parameters V; S; h. The coefficient of wear variation during machining one workpiece k ¼ r=u;

ð8Þ

depends on the coefficient of variation of the workpiece hardness and the coefficient of variation of the machining allowance, but it does not depend on the cutting modes. As for the wear rate u, it essentially depends on the cutting modes and mainly determines the tool life equation.

3 Assessment of Tool Life Equation Parameters To assess the parameters of the tool life equation from tests, we use the statistics ðti ; Vi ; Si ; hi ; Yi Þ; i ¼ 1; . . .; N, where ti is the number of the machined workpieces when the cutting mode parameters are Vi ; Si ; hi , and Yj is the total wear during the time of machining ti workpieces. i is the test number with the cutting mode parameters Vi ; Si ; hi , N is the number of tests. For the tests adequacy, it is necessary that at each test, the values of the cutting mode parameters are changed, for example, as in a two-level factorial experiment. We will look the wear rate as a function of the cutting mode parameters in the form of power dependence, for example, of this type:

Evaluation of Tool Life Equation of Single-Point …

 uðV; S; hÞ ¼ exp b0 þ b1
ln V þ b2
ln2 V þ b3
ln3 V þ b4
ln S þ b5
ln2 S þ b6
ln V ln S þ b7
ln h;

947

ð9Þ

where b0 ; . . .; b7 are the required coefficients to be estimated using the statistics noted above. The number of components in the formula (9) may vary depending on the available statistics, the factors taken into account, and the required accuracy of the dependence (9). To estimate the coefficients b0 ; . . .; b7 , let us use the maximum likelihood method based on the distribution (4) with taking into account (8). The optimal values of the parameters b0 ; . . .b7 and k maximize the likelihood function " # 1 ðYi  ui ti Þ2 pffiffiffiffiffiffiffiffi exp  Prðb0 ; . . .; b7 ; kÞ ¼ 2k 2 u2i ti 2pti kui i¼1 N Y

ð10Þ

It is easier to find the maximum of the likelihood function’s logarithm, that is, the maximum lnðPrðbo ; . . .; b7 ; kÞÞ ¼

N X i¼1

pffiffiffiffiffiffi ðYi  ui ti Þ2 ½ lnð 2pÞ  lnðkÞ  lnðui Þ  lnðti Þ  : 2k2 u2i ti ð11Þ

In the formulas (10), (11), the notation is accepted ui ¼ uðVi ; Si ; hi Þ. The maximum of the function (11) in our example is searched by the differentiation method together with the random search. From the equation @ lnðPrðb@k0 ;...;b6 ;kÞ ¼ 0 we obtain that k2 ¼

N 1X ðYi  ui
ti Þ2 : N i¼1 u2i
ti

ð12Þ

Taking into account (12), the expression (11) will be simplified and take the form: N N pffiffiffiffiffiffi 1 X 1X 1 lnðti Þ  lnðkÞ  lnðui Þ  : lnðPrðbo ; . . .; b6 ; kÞÞ ¼ N½lnð 2pÞ  2N i¼1 N i¼1 2

ð13Þ We are searching for the maximum of function (13) for b0 ; . . .; b7 using a random search method. Random search is implemented as follows. The parameter value options b0 ; . . .; b7 are generated by the formula

948

A. V. Antsev et al.

bj ¼ b0j þ ðb00j  b0j Þ
random;

j ¼ 0; . . .; 7;

ð14Þ

where b0j ; b00j are the search boundaries, which are defined from a priori considerations, random is a function generating each time a pseudorandom number uniformly distributed in the interval from 0 to 1. The parameter value options b0 ; . . .; b7 are generated in the search cycle. For each option, the value of lnðPrðb0 ; . . .; b7 ; kÞ is calculated by (12) and (13). The best option parameters b0 ; . . .; b7 ; k and the maximum achieved lnðPrðb0 ; . . .; b7 ; kÞÞ will be stored in the computer memory.

4 Illustration of the Method Now let us consider a specific example of the calculation of the tool life equation on statistics ðti ; Vi ; Si ; Yi Þ; i ¼ 1; . . .; N. Such statistics are shown in Table 1. The statistics was gathered during test with material steel 1X18H9T, cut depth 0.5 mm, tool material hard alloy T30К4, and maximum wear 0.4 mm. The wear rate function will be searched in the form similar to (9), but with the feature that the cut depth h during the tests did not vary, so the dependence for the wear rate on h is not included, that is, ui ¼ uðVi ; Si Þ. In addition, we will check ten options of this function as uðV; SÞ according to the likelihood ration test. The simplest form for uðV; SÞ is option 1, where the maximum powers at ln V and ln S are 1 and 1, which is indicated in the Table 2 as 1 1. Hereby, uðV; SÞ ¼ expðb0 þ b1 ln V þ b4 ln SÞ. In option ten, these powers are 2 and 2 and besides there is a member ln V
ln S, which is indicated in Table 2 as 2 2 1*1. This means that uðV; SÞ ¼ expðb0 þ b1 ln V þ b2 ln2 V þ b4 ln S þ b5 ln2 S þ b6 ln V ln SÞ. Similarly, other dependency options uðV; SÞ, marked in the Table 2, are defined. The option based on uðV; SÞ with the maximum likelihood Prðb0 ; . . .; b6 Þ is considered preferable. The calculation results for all ten options are summarized in Table 2. Table 1 Experimental data: speed-feed-operating time-wear (V  S  T  Y) V

S

T

Y

V

S

T

Y

V

S

T

Y

37 70 100 150 200 37 70 100 135 150

0.1 0.1 0.1 0.1 0.1 0.15 0.15 0.15 0.15 0.15

41 45 62 25 2.5 55 75 80 70 13

0.4 0.4 0.4 0.4 0.4 0.4 0.4 0.4 0.4 0.4

170 210 37 70 100 140 210 45 70 95

0.15 0.15 0.2 0.2 0.2 0.2 0.2 0.3 0.3 0.3

6 1.8 45 54 37 14 1.8 37 60 27

0.4 0.4 0.4 0.4 0.4 0.4 0.4 0.4 0.4 0.4

110 120 210 45 70 100 110 155 210

0.3 0.3 0.3 0.4 0.4 0.4 0.4 0.4 0.4

27 25 1.2 44 65 60 31 5 0.9

0.4 0.4 0.4 0.4 0.4 0.4 0.4 0.4 0.4

Evaluation of Tool Life Equation of Single-Point …

949

Table 2 Results of the calculation №

Structure UðV; SÞ

1

11

−9.52

2

21

−0.19

−3.42

3

2 1*1

−0.19

4

31

−3.96

5

3 1*1

−3.96

6

32

7

3 1 1*1

8

b0

b1

b2

b3

b4

b5

b6

k

Pr

0

0

0.087

0

0

5.015

0.0003

0.533

0

−0.02

0

0

5.273

0.0332

−3.42

0.533

0

0

0

−0.02

4.77

0.0462

8.54

−4.17

0.49

−0.01

0

0

2.341

1791

8.54

−4.17

0.49

0

0

−0.01

2.323

953

8.913

−2.50

0.242

0.461

−0.06

0

3.356

0.018

−5.62

6.243

−2.91

0.345

−0.60

0

0.08

2.731

2.143

3 2 1*1

−4.24

6.326

−3.11

0.405

−1.57

1.074

1.151

2.510

0.379

9

22

−1.747

−2.75

0.425

0

−1.47

−0.53

0

4.987

0.016

10

2 2 1*1

0.795

−3.97

0.665

0

−1.38

0.585

0.662

4.106

0.032

−15.1

1.175

Fig. 1 Graphs of tool life dependence on the cutting speed at different feeds. 1—experimental values, 2—mathematical expectations. a S ¼ 0:1 mm/rev; b S ¼ 0:2 mm/rev; c S ¼ 0:3 mm/rev; d S ¼ 0:4 mm/rev

It follows from Table 2 that option four is preferable in terms of likelihood. In this case, the wear rate is  uðV; SÞ ¼ exp 3:96 þ 8:54
ln V  4:17
ln2 V þ 0:49
ln3 V  0:01
ln S : ð15Þ

950

A. V. Antsev et al.

As it follows from the analysis of Table 2, the agreement of the actual wear distribution with the theoretical one can be judged by a simpler indicator calculated by the formula (12). The graphs shown in Fig. 1 are constructed using the dependence (15). The mathematical expectation of the tool life T
was calculated by the formulas (5), (6).

5 Conclusion The proposed tool life equation as a set of the distribution law of the cutting tool life T and the dependence of the parameters of this law on the cutting mode can accurately predict tool failure taking into account the stochastic nature of the cutting ^

tools wear. The average tool life T , calculated by the formula (6) using the expression for uðV; SÞ (15), differs from the mathematical expectation of the tool ^ life by an amount of 0.4–5%. The difference between T
and T is not big and in ^

practical calculations, you can use the approximation T
 T and a simpler formula
(7) to calculate T.

References 1. Vasin SA, Antsev VY, Dolgov DV (2002) Information assistance in the tool assurance system on the machine-building plant. Avtom Sovr Tekhnol 2:3–10 2. Ademujimi TT, Brundage MP, Prabhu VV (2017) A review of current machine learning techniques used in manufacturing diagnosis. IFIP Adv Inform Commun Technol 513:407– 415 3. Filonenko S (2015) The effect of wear of a cutting tool with a controlled depth of cut on the acoustic emission. East -Eur J Enterp Technol 6(9):47–50 4. Ghorbani S, Kopilov VV, Polushin NI et al (2018) Experimental and analytical research on relationship between tool life and vibration in cutting process. Arch Civ Mech Eng 18 (3):844–862 5. Inţǎ M, Muntean A (2018) Researches regarding introducing temperature as a factor in cutting tool wear monitoring. MATEC Web of Conferences 178:01013. https://doi.org/10.1051/ matecconf/201817801013 6. Martinov GM, Grigor’ev AS (2013) Diagnostics of cutting tools and prediction of their life in numerically controlled systems. Russ Eng Res 33(7):433–437 7. Zhang B, Shin YC (2018) A multimodal intelligent monitoring system for turning processes. J Manuf Processes 35:547–558 8. García-Ordás MT, Alegre-Gutiérrez E, Alaiz-Rodríguez R et al (2018) Tool wear monitoring using an online, automatic and low cost system based on local texture. Mech Syst Signal Process 112:98–112 9. Astakhov VP (2006) Tribology of metal cutting. Elsevier, London 10. Cheng D, Oh Y, Kim S (2018) A new approach for predicting wear overlap geometry in ball end finish milling. Int J Adv Manuf Technol 98(9–12):2677–2691 11. Panda A, Duplák J, Vasilko K (2013) Comprehensive identification of durability for selected cutting tool applied on the base of taylor dependence. Adv Mater Res 716:254–260

Evaluation of Tool Life Equation of Single-Point …

951

12. Malakizadi A, Gruber H, Sadik I et al (2016) An FEM-based approach for tool wear estimation in machining. Wear 368–369:10–24 13. Wang G, Guo Z, Qian L (2014) Tool wear prediction considering uncovered data based on partial least square regression. J Mech Sci Technol 28(1):317–322 14. Zhu K, Zhang Y (2019) A generic tool wear model and its application to force modeling and wear monitoring in high speed milling. Mech Syst Signal Process 115:147–161 15. Erenkov OY, Ivakhnenko AG, Radchenko MV (2013) Oscillatory process of production systems during turning of caprolon blanks. Chem Pet Eng 49(5–6):411–417 16. Kuts V, Ivakhnenko A, Khandozhko A (2016) Investigations of cutting force effect upon shaping error of surfaces with double curvature in technological systems with mechanisms of parallel structure. In: Proceedings of 2015 international conference on mechanical engineering, automation and control systems (MEACS 2015) 2015:7414887. https://doi.org/10.1109/ meacs.2015.7414887 17. Bannikov AI, Kurchenko AI, Makarova OA et al (2010) Tool wear in turning corrosion-resistant steel. Russ Eng Res 30(7):734–735 18. Muratov KR (2016) Influence of rigid and frictional kinematic linkages in tool-workpiece contact on the uniformity of tool wear. Russ Eng Res 36(4):321–323 19. Chernikov PP, Sharipov BU (2008) Influence of metal lubricant on tool wear. Russ Eng Res 28(2):194–195 20. Pasko NI, Antsev AV, Antseva NV et al (2017) The generalized mathematical model of the failure of the cutting tool IOP. In: Conference series: materials science and engineering 177:012052. https://doi.org/10.1088/1757-899x/177/1/012052 21. Antsev AV, Pasko NI, Antseva NV (2018) Assessment of wear dependence parameters in complex model of cutting tool wear. IOP. In: Conference series: materials science and engineering 327:042005. https://doi.org/10.1088/1757-899x/327/4/042005

Ultra-precision Machining of Surfaces of Elements of Devices from Optical Materials S. V. Grubiy, M. A. Shavva and V. V. Lapshin

Abstract Optical materials are widely used in the composition of modern engineering products and aerospace and electronics industries. At the same time, optical surface processing quality is ensured only with the use of diamond cutting tools and ultra-precision ultra-rigid equipment. In addition, it is necessary to provide cutting conditions in the nanoscale thickness range of the cut layer. The results of theoretical and experimental studies aimed at improving the performance, accuracy, and quality of machining surfaces of elements of devices from optical materials (reducing the roughness and depth of fractured layer), potassium dihydrophosphate (KDP), Sitall, quartz glass, by using diamond cutter and abrasive diamond tools, are presented in the chapter. Kinematic schemes of ultra-precise machining of optical surfaces and the basic theoretical dependences for calculation of cutting layer thickness are considered. Recommended cutting modes and conditions are presented. The examples of diamond milling and abrasive processing on ultra-precise experimental stands and results of metrological control of treated surface quality parameters are also given. Keywords High-accuracy machining micromilling Diamond grinding





Optical materials
Diamond

1 Introduction The traditional technology of treatment of the surface of elements made of optical materials provides diamond grinding with a successive reduction of abrasive grit and chemical mechanical polishing. The main disadvantages of traditional technology are poor process performance, complexity of automation and control, and S. V. Grubiy (&) BMSTU, 5/1, 2-aya Baumanskaya Street, Moscow 105005, Russia e-mail: [email protected] M. A. Shavva
V. V. Lapshin VNIIINSTRUMENT, 49, B. Semenovskaya, Moscow 103027, Russia © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_101

953

954

S. V. Grubiy et al.

damage to the surface layer by abrasive particles. In Bauman Moscow State VNIIINSTRUMENT JSC, Resurs Tochnosti LLC in cooperation with other organizations carried out a set of technological researches and engineering developments aimed at increasing the productivity, accuracy, and quality of treatment of the surface of the elements made of optical materials by diamond edge and abrasive tools. Technological experimental studies were carried out on high-accuracy stands according to the diagram of flat surface milling by diamond single-cutter milling head and flat diamond grinding with the grinding wheel shaft tilted. Samples for treatment are made of optical materials: Sitall, quartz glass, and potassium dihydrogen phosphate. It is theoretically substantiated and experimentally confirmed that it is possible to create conditions for treatment that ensure cut layer thickness in the nanometer range providing that the mechanism of fragile interaction “tool-treated material” is changed to the plastic deformation [1–5]. The transition from brittle chipping to plastic deformation allows to obtain a surface with minimal damaged layer and roughness Ra less than 0.01 lm. The treated surface form error is ensured by the high-accuracy machine kinematics [6].

2 Diamond Single-Point Cutting To ensure optical material plastic deformation in the cutting area, the diamond single-crystal tool shall have the following characteristics: front angle in the range of 0° to −35° [7]; cutting edge corner radius: 50–100 nm; tool blade top radius: 1– 10 mm; rear angle: 7°–15°. The cut layer thickness in course of the edge treatment depends on the cutting modes and parameters of the cutting tool. Figure 1 shows the dependencies of the cut layer thickness a on the longitudinal feed of the workpiece per one revolution of the tool S at 1 mm radius and 2 lm cutting depth, and Fig. 2 shows the above dependencies on cutting depth t at 1 mm top radius and 2 µm/rev workpiece feeding. The analysis shows that the cut layer thickness decreases as the workpiece feeding and cutting depth decrease [8–14]. It has been experimentally proven that the cut layer thickness should not exceed 100–120 nm to ensure KDP.

Fig. 1 Graph of dependence of the cut layer thickness on the workpiece feeding

Ultra-precision Machining of Surfaces of …

955

Fig. 2 Graph of dependence of the cut layer thickness on the cutting depth

Experimental studies on the treatment by a single-cutter diamond milling head were carried out on the high-accuracy experimental stand for diamond micromilling with the following constructional features: Stand spindle box and linear saddle are made on aerostatic bearings; the stand is mounted on antivibration mounts; the spindle box motor cooling system is provided. The main parameters of the high-accuracy experimental stand are given in Table 1. The kinematic diagram of treatment by an edge diamond tool on the high-accuracy experimental stand is shown in Fig. 3. The objects of research were workpieces made of potassium dihydrogen phosphate (KDP) with dimensions of 180  180  10 mm. The parameters of the diamond cutter, cutting modes, and cut layer thickness are given in Table 2. Leica DCM3D 3D profilometer was used to control the treated surface roughness. The treated surface profilogram is shown in Fig. 4. The treated surface roughness was Ra 3 nm, Rz 12 nm. The form error of the treated flat surface of KDP workpiece was measured by the laser interferometer produced by the RAS Applied Physics Institute (Nizhny Novgorod). Figure 5 shows the treated surface interferogram. The flatness deviation was 280 nm. The treated part form error is determined by the machine slideway straightness error, poor alignment accuracy, and vibrations occurring during treatment [15].

Table 1 Parameters of high-accuracy experimental stand for diamond micromilling No.

Parameter Name

Size

1. 2. 3.

The largest dimensions of the workpiece, mm Spindle speed range, min−1 Number of axles, pcs. X-axis—tangential saddle C-axis—spindle rotation Tangential saddle stroke, X-axis, mm Resolving power along X-axis, nm Feeding range along X-axis, mm/min

450  450 50–600 2

5. 6. 7.

800 1 0.002–150

956

S. V. Grubiy et al.

Fig. 3 Kinematic diagram of treatment on high-accuracy experimental stand for diamond micromilling

Table 2 KDP workpiece treatment modes Parameter

Size

Single-cutter milling head Rml, mm Diamond cutter radius r, mm Cutting depth t, lm Feed per revolution S, lm/rev Spindle speed n, rpm Cut layer thickness a, nm

325 1 2 2 340 124

Fig. 4 KDP workpiece treated surface profilogram

Ultra-precision Machining of Surfaces of …

957

Fig. 5 KDP workpiece treated surface interferogram

Results of experimental studies on edge diamond blade treatment show that optical and microelectronics components made of potassium dihydrogen phosphate, silicon, and germanium can be treated with the roughness Ra that does not exceed 5 nm by providing plastic deformation conditions for the material in the cutting area [16–18].

3 Diamond Grinding To ensure plastic deformation in course of the diamond grinding—the cut layer thickness in the nanometer range, the cutting tool shall have the following parameters: diamond bearing layer abrasive grit in the range of 2–3 lm; grain concentration in the diamond bearing layer—not less than 150%; use of organic bonds; periodic adjustment of the grinding wheel; it was also proposed to use the flat grinding diagram with the wheel shaft tilt in relation to face-plate spinning axis with workpieces—Fig. 6 [19]. In Fig. 5, the grinding wheel with radius R is tilted to an angle b and rotates at a frequency n1. The workpiece circular feed at Sprod speed is set by the rotation speed n2 of the rotary table with Dzag diameter. Cutting depth is indicated by t. The wheel transverse feed is carried out at Spop speed. Changing the grinding wheel shaft tilt allows to change the direction of the crack development in the damaged layer. Optimum tilt of the wheel that provides plastic deformation of the material in the cutting area and minimum depth of the fractured layer (less than 50 nm) is experimentally confirmed and located in the range of 1°– 3°. Table 3 provides parameters for calculating the cut layer thickness by single grain cutting when grinding with the wheel shaft tilted. Figure 6 shows a graph of the dependency of the cut layer thickness by single grain cutting on the longitudinal feed. Calculations show that the cut layer thickness depends on the cutting modes, cutting tool parameters, and wheel shaft tilt. The kinematic diagram for treatment by diamond grinding wheel with shaft tilt (Fig. 7)

958

S. V. Grubiy et al.

Fig. 6 Graph of dependence of the cut layer thickness by single grain cutting on longitudinal

Table 3 Parameters for calculating the cut layer thickness by single grain cutting Parameter

Size

Coefficient that takes into account the grains in the bundle ɛ Average diameter of diamond grain x, lm Concentration of grains in diamond bearing layer K, % Wheel rotation speed n1, rpm Cutting depth t, lm Width of diamond bearing layer B, m Radius of grinding wheel R, m Wheel shaft tilt b, deg

0.5 2 100 1000 5 0.01 0.05 1.5

Fig. 7 Topograms of treated optical surfaces: a Sitall SO-115 M; b quartz glass KU1

was implemented on a high-accuracy experimental stand. The parameters of the stand are given in Table 4. The structural features of the stand are: Workpiece spindle and grinding wheel spindle are based on gas-lubricated bearings, which provides the value of the operating surfaces radius runout not exceeding 0.5 lm and roughness not less than 100 N/lm; longitudinal and transverse saddles are made with porous choking for

Ultra-precision Machining of Surfaces of …

959

Table 4 Parameters of the high-accuracy experimental stand to examine the optical surface grinding process No.

Parameter Name

Size

1. 2. 3. 4. 5.

Frequency range of grinding spindle n1, min−1 Workpiece spindle speed range n2, min−1 Longitudinal saddle stroke, Z-axis (cutting saddle), mm Tangential saddle stroke, X-axis, mm Number of axles, pcs. X-axis—tangential saddle Z-axis—longitudinal saddle (cut-in) S-axis—grinding wheel rotation; S1-axis—workpiece rotation Working feed of longitudinal saddle, mm/min Tangential saddle movement discretibility, lm Power of electric motor grinding wheel, kW

50–3000 0.01–100 100 200 4

6. 7. 8.

5–200 0.1 1

gas lubrication feeding and have additional damping by extra viscous liquid; the machine display system allows to set the value of the tool cutting within the accuracy of 0.1 lm. The objects of research for diamond grinding were workpieces made of Sitall brand SO-115 M and quartz glass brand KU1. The workpiece treatment in course of the diamond grinding was carried out in the following modes: cutting depth t— 1–2 lm, grinding wheel rotation speed n1—1000 rpm, workpiece spindle speed n2—0.2 rpm, and Sprod longitudinal feed—0.0075 m/min. The workpiece rotation diameter was 120 mm. Diamond wheel used for grinding: shape 12A2 45°, wheel dimensions: diameter—100 mm, height—21 mm, bore diameter—20 mm, diamond bearing layer width—6 mm, diamond bearing layer height—3 mm, ASM brand diamond powder with the abrasive grit of 3/2 microns, concentration—100%, and bond grade—B1. Analysis of the experimental studies results shows that diamond grinding of optical materials allows to provide conditions for material plastic deformation in the cutting area and obtain the treated surface roughness Ra not less than 10 nm. In addition, the plastic material deformation in the cutting area and using of the grinding diagram with the wheel shaft tilted provide the reduction of the damaged layer depth to the value less than 50 nm. The fractured layer minimum depth allows to reduce the dispersion value directed to the treated optical surface of dispersion, and the general efficiency, durability, and accuracy of the instruments [20–22].

4 Conclusion The obtained experimental data on diamond edge and abrasive treatment allow us to conclude that upon conditions of plastic deformation for optical materials in the cutting area, the treated surface roughness Ra is less than 10 nm, and the fractured

960

S. V. Grubiy et al.

layer depth does not exceed 50 nm. The obtained quality parameters of the treated surfaces of KDP crystals, Sitalls, and quartz glass show that during treatment of the considered materials by diamond milling and grinding on high-accuracy machines, it is possible to minimize or exclude polishing operations from the technological process. The necessary recommendations for cutting tool parameters and treatment modes shall be observed to ensure plastic deformation of the optical material in the cutting area. At the same time, the treated surface form accuracy is ensured by kinematic accuracy of the equipment used and can reach up to 0.3 lm on the aperture 180  180 mm.

References 1. Xianqun H, Chaoshui X (2015) Specific energy as an index to identify the critical failure mode transition depth in rock cutting. Rock Mech Rock Eng 49:1461–1478 2. Masahiko Y, Sivanandam A, Matsumura T (2008) Critical depth of hard brittle materials on nano plastic forming. J Adv Mech Des Syst Man 2(1):59–70 3. Muhammad A, Mustafizur R, Wong Y (2011) Analytical modelling of ductile regime machining of tungsten carbide by end-milling. Int J Adv Man Techn 5:53–64 4. Bifano T, Fawcett S (1991) Specific grinding energy as an in process control variable for ductile-regime grinding. Prec Eng 4:256–262 5. Masahiko Y, Sivanandam A, Matsumura T (2008) Critical depth of hard brittle materials on nano plastic forming. J Adv Mech Des Syst Man 1:59–70 6. Borovskiy G, Shavva M, Grubiy S et al (2015) Ultraprecision machining of brittle optical materials. Rus Eng Res 91:883–889 7. Bifano T, DePiero D, Golini D (1993) Chemomechanical effects in ductile regime machining of glass. Prec Eng 4:238–247 8. Qiangguo W, Hang G, Zhijian P (2013) An experimental investigation on slicing of potassium dihydrogen phosphate crystal. J Eng Man 227:890–897 9. Zhang J, Zhang Y, Xu K (2008) Young’s modulus surface and Poisson’s ratio curve for tetragonal crystals. J Chem Cryst 17(5):1565–1573 10. Fang T, Lambropoulos J (2002) Microhardness and indentation fracture of potassium dihydrogen phosphate (KDP). J Am Cer Soc 85(1):174–178 11. Chunpeng L, Hang G, Jinghe W (2010) Mechanical properties of potassium dihydrogen phosphate single crystal by the nanoindentation technique. Mat Man Proc 25:740–748 12. Zhang Y, Zhang L, Liu M (2016) Revealing the mechanical properties of potassium dihydrogen phosphate crystals by nanoindentation. J Mat Res 31(8):1056–1064 13. Chen H, Dai Y, Zheng Z (2011) Effect of crystallographic orientation on cutting forces and surface finish in ductile cutting of KDP crystals. Mach Sci Techn An Int J 15:231–242 14. Hocheng H, Hsieh M (2004) Signal analysis of surface roughness in diamond turning of lens molds. Int J Mach Tools Man 44:1607–1618 15. Blake P, Scattergood R (1990) Ductile regime machining of germanium and silicon. J Am Cer Soc 73(4):949–957 16. Goel S, Luo X, Comley P (2013) Brittle-ductile transition during diamond turning of single crystal silicon carbide. Int J Mach Tools Man 65:15–21 17. Yan J, Yoshino M, Kuriagawa T (2001) On the ductile Machining of silicon for micro electro-mechanical system (MEMS). Opto-electronic and optical applications. Mat Scien Eng 297:230–234

Ultra-precision Machining of Surfaces of …

961

18. Zaharevich E, Lapshin V, Shavva MA et al (2016) The Experimental Determination of the Ductile-Fracture Transition Boundaries when Cutting Brittle Materials. News of High Ed Inst Eng 7:64–71 19. Shavva M (2017) The Methods of calculation of cutting forces for diamond grinding of brittle optical materials. News High Ed Inst Eng 2:61–69 20. Inkson B, Wu H, Steer T et al (2001) 3D mapping of subsurface cracks in alumina using FIB. Mat Res Soc 649:Q.7.7.1–Q.7.7.6 21. Salvati E, Sui T, Lunt A, Korsunsky A (2016) The effect of eigenstrain induced by ion beam damage on the apparent strain relief in FIB-DIC residual stress evaluation. Mat and Des 92:649–658

Identification of Deformations and Errors in Flat Thin Workparts During Grinding T. N. Ivanova

Abstract The research on measurement errors characterizing the accuracy of grinding of thin workparts by surface grinders was carried out. It was established that the geometrical error of the magnetic base exerts the most considerable influence on the accuracy of the shape of workparts during surface grinding. Temperature errors result in the growth of bulge, the absence of straightness of worktable movement and buckling of mounting surface of magnetic base. It was identified that during surface grinding the highest influence on accuracy of shape of workpart to be machined is exerted by its own thermal deformations which occur due to the action of heat source in cutting area. The analysis of the given solution demonstrates that the value of non-flatness due to temperature deformations is directly proportional to initial temperature difference between workpart and mounting surface, linear expansion coefficient of a workpart and its squared diameter. At the same time, the value of non-flatness is inversely proportional to the workpart thickness. The lower linear expansion coefficient of the machined material is, the higher possibility of fluctuations in the workpart temperature is. The use of obtained dependences of temperature deformations will make it possible to solve practical tasks of identification of errors in flat workparts made of any machinable material. Keywords Deformation Error Device





Non-flatness
Thin workpart
Mounting surface


1 Introduction The process of grinding of workparts made of thin plates is entailed by a huge amount of various factors, which cause a deviation of real values of geometric characteristics from their nominal values—machining errors [1–20]. T. N. Ivanova (&) Tchaikovsky Branch Perm National Research Polytechnic Institute, 73 Lenin St., Tchaikovsky 617764, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_102

963

964

T. N. Ivanova

2 Structure All in all, a total error of workpart construction can be demonstrated as a scheme (Fig. 1). The total error, characterizing the accuracy of machining of workparts by surface grinders, is expressed as a functional dependence: D ¼ f ðDy; Dn; Dc; Di; DT Þ

ð1Þ

where Dy, Δf, Dc, Di, DT—errors that occur due to elastic displacements, mounting of workpiece, geometric deflection of machining station, wear of a tool and thermal deformations. The machining allowance in case of grinding is 0.01–0.3 mm. Workparts are fastened by hard devices and machining stations of an ultrahigh accuracy; the diamond grinding wheels, which decrease cutting forces and elastic compression, are employed. By using spark-out of a machined surface, the error Dy can be minimized. The error of mounting Δf of a workpart during grinding exerts an insignificant influence on the accuracy of machining because the allowance for non-straightness of a mounting surface of a magnetic base of the device should not exceed 0.005 mm per 1000 mm. The error Dc makes an impact on the workpart accuracy through nonstraightness of worktable movement and a spindle runout. In view of the fact that surface grinders are made with increased and high accuracy, these errors do not limit the process of machining. The error from dimensional wear of a cutting grinding tool in the process of face grinding by diamond grains or by borazon wheel is reduced to an insignificant value, considering the small size of workparts and minimal machining allowance. Fig. 1 Generalized classification of the factors governing the total error of machining

Identification of Deformations and Errors …

965

The highest influence on the total error is exerted by the error ΔT, which is caused by thermal deformations of the machining station-device-tool-workpart system. Therefore, ΔT can be expressed as a result of an action of four components: DT ¼ f ðDTO ; DTC ; DTP ; DTD Þ

ð2Þ

where DTO—an error used by thermal deformation of the machining station-device-tool-workpart system; DTC—an error caused by thermal deformation of the machining station units and parts; DTP—an error resulting from thermal deformation of a magnetic base; ΔTD—an error resulting from thermal deformation of a workpart. The error DTO occurs due to thermal deformation of the machining station-device-tool-workpart system under an action of the environment temperature. It decreases when the workshop temperature remains constant. The error DTC connected with an evolution of heat in machining station units (engines, bearings, friction units, hydraulic system) is insignificant due to an application of thermal insulation, forced cooldown of heat-emitting units, relocation of hydraulic system outside of machine stand. The error resulting from thermal deformation of a magnetic base DTP depends on the following factors. Under action of the heat, emitted when a current passes through magnetic coils, the base is warmed up to 32–36 °C. If the workpart is grinded with an application of a lubricating–cooling fluid, the warming up reaches 25–27 °C. Under the influence of the temperature of magnetic coils of the base, the air expands and creates the pressure around the case. This pressure causes deformations of high magnitude in the center of magnetic base EP–21G(Russian State Standard GOST) because the stiffness in central points of the base is lower than along the edges. The experimental research conducted demonstrated [1–3] that only under the action of the heat emitted by magnetic coils, the bases of model EP–21G are buckling up to 3–5 l in the center. Besides an internal heat source, a magnetic base can receive heat from the mounted workpart, which will influence the accuracy of the shape of thin workparts. Thus, the geometric error of the magnetic base has the most significant influence over the accuracy of the shape of workparts during surface grinding. Temperature errors result in the growth of a bulge, non-straightness of worktable movement and buckling of the mounting surface of the magnetic base. The error ΔTD occurs due to thermal deformation of a workpart under an action of evolution of heat in the grinding area. Depending on the machined material, cooldown conditions and wheel characteristics, up to 80% of the heat emitted in the cutting area can transfer inside the workpart. So, the biggest influence on the accuracy of a shape of machined part during surface grinding is exerted by thermal deformations of the part, which occur due to an action of a heat source in the cutting area. In this case, an occurrence of even insignificant temperature gradient in a workpart can cause such value of deformations of the workpart that its geometric dimensions will exceed the tolerance range for deviation from straightness. Shape errors are connected with the workpart:

966

T. N. Ivanova

DTD ¼ DDY þ DDT þ DDJ þ DDE

ð3Þ

where DDY—an error caused by non-flatness of mounting surface of a workpart; DDT—an error from temperature deformations of a shape of a workpart; DDJ—an error from workpart deformations occurring due to varying stiffness; DDE—an error caused by inconstancy of a cutting force due to non-uniform distribution of allowance along the machined surface. The errors a used by non-flatness of amounting surface of a workpart DDY is possible in case of a deviation of the shape of mounting surface, resulting from an action of the attraction force, which leads to the deformation of machined surface. The error of a mounting surface is transferred to the machined surface with opposite sign: A concavity of the mounting surface results in bulging of the machined surface and vice versa. This error can be transferred to the machined surface of the part fully or partly depending on the workpart stiffness and attraction force. For instance, when the part was 200 mm long and 3.5 mm wide and was fastened with the attraction force equal to 3  105 N/m2, the concavity of the mounting surface reaching 0.08 mm was fully transmitted to the machined surface [3]. The shaper or resulting from varying stiffness of a workpart DDJ occurs due to constancy of the stiffness of the machining station and the magnetic base. Non-rigid flat parts, in turn, are pressed to the magnetic base without any clearance and do not change the stiffness of the machining station-device-tool-workpart system. That is why the influence of this error on the accuracy of the shape of a workpart can be disregarded. The error caused by inconstancy of the cutting force due to non-uniform distribution of the allowance along the machined surface DDE has an insignificant influence on the total error of a workpart. The process of grinding is multipass, so the error from non-uniform distribution of the allowance of a workpiece is removed from passage to passage. The last passages are made under real force of cutting, which does not exceed several Newtons, so the value of this error will be small. The error resulting from thermal deformation of a workpart DDT in case of surface grinding is determined not by the amount of the heat absorbed by the part and its warming up, but by the distribution of the temperature along the cross section of the part. In case of surface grinding, a workpart receives the heat from two sources: from the cutting process carried out (machined surface) and from the magnetic base (mounting surface). As the result of non-uniform heating, a workpart becomes deformed bending toward metal layers, which have higher temperature, to the following value f ¼ a
DT
L2 =ð8hÞ

ð4Þ

where f—the workpart deflection value, h—the workpart thickness, ΔT—the temperature difference between upper and lower surface, L—the length of machined workpart and a—coefficient of linear expansion of workpart material.

Identification of Deformations and Errors …

967

If we use Eq. (4) to estimate temperature deformation, which leads to the error of the shape, we can make the conclusion that its value cannot be ignored. For example, for the part being 200 mm long and 3 mm wide, the temperature difference between opposite surfaces reaching only 1 °C causes deformation equal to 0.6 l [1]. From the point of view of temperature deformation of shape, dry surface grinding is unfavorable. In this case, the upper layers are warmed up relatively to the lower ones, which are in contact with the magnetic base, so the part, deforming, bends toward the grinding wheel. Enhanced metal removal occurs in the central part of the machined surface. After the allowance has been removed, the temperatures of machined and mounting surfaces stabilize. The temperature stabilization reduces the value of workpart deformation and, consequently, machined surface sags in the center, acquires a concave nature, though the quality parameters do not change. However, if the machined surface is constantly washed by a lubricating–cooling fluid flow, this contributes to the removal of heat from it. Due to the fact that the mounting surface of a workpart, contacting with the magnetic base, is located in less favorable conditions, an application of a lubricating–cooling fluid during grinding results in higher temperature of the mounting surface in comparison with the machined one. Relative elongation of metal layers adjacent to the mounting surface will lead to part bending away from the grinding wheel, being torn away from the magnetic base along the edges. Enhanced metal removal occurs along the edges of the machined surface. When the grinding process is over, the machined profile acquires a convex character during cooling. If due to small material volume and optimal cutting mode the amount of heat, transferring to the part per time unit, is turned out to be not enough to keep the temperature of the machined surface higher than the temperature of the mounting surface, an application of cooling will significantly reduce the magnitude of temperature deformations of the shape. By studying dependence (4), it was identified that overall dimensions of a workpart influence the magnitude of temperature deformation of the shape. Temperature deformations increase disproportionately. It is explained by the fact that an increase in the length of the part causes the growth of the volume of the machined material. In this case, an amount of heat per volume unit of material decreases, heat emission to the environment increases. As a result, temperature difference between mounting and machined surfaces is reduced, temperature deformation of the shape decreases as well. The value of temperature deformations of the shape also decreases in case of increasing the width of a machined part. Temperature deformations depend on the thickness of a flat part and on the time of grinding. An increase in the thickness leads to drop in a volume of heated material and temperature deformations reduction according to dependence (4). A decrease in the thickness affects the fact that the heat from grinding area comes from the machined surface to the mounting surface, the temperature difference between surfaces drops, and, consequently, the value of temperature deformation of the shape grows. An increase in grinding time results in the growth of temperature deformation of the shape.

968

T. N. Ivanova

Occurring temperature deformation of the shape increases real cutting depth, which, in turn, causes growth of temperature and, consequently, deformation. It was experimentally observed that the value of temperature deformation of the shape according to grinding time in case of installation of the part with fastening along the edges was insignificant during first working strokes (4–10 strokes) and did not exceed 1 l. If a workpart is mounted on magnetic base, the grinding is possible with or without cooling. Firstly, let us observe the case of machining of a part without cooling. Grinding of the magnetic base was carried out in a cold condition of the machining station. After stabilization of temperature deformations, non-straightness of a profile of the magnetic base did not exceed 2 l. The mounting surface of the part also excluded the possibility of deformation of the part in case of its fastening to the magnetic base. During machining of the part without cooling, tearing of the part from the magnetic base in the center and edge points was observed. It is explained by relative heating of the upper layers of metal. Figure 2 demonstrates the dependence of the value of temperature deformation of the shape on grinding time for different cutting modes in case of its fastening to magnetic base EP–21G (Russian State Standard GOST).

Fig. 2 Dependence of temperature deformation of a workpart shape on grinding time in different cutting modes: a St = 1 mm per stroke, t = 0.02 mm, lines 1, 2—without cooling, 3, 4—with cooling, lines 1, 3—steel 9XC/01T31507, lines 2, 4—steel 12XH3A/3415; b V = 10 m per minute, t = 0.02 mm, lines 1, 2—without cooling, 3—with cooling, line 1—steel 9XC/01T31507, lines 2, 3—steel 12XH3A/3415; c V = 10 m per minute, St = 1 m per minute lines 1, 2—without cooling, 3—with cooling, line 1—steel 9XC (HB197–241) (Russian State Standard GOST)/01, T31507 Quenched HRC 58–63, lines 2, 3—steel 12XH3A (HRC 57–63) (Russian State Standard GOST)/3415, wheel 150x32x40 AC6 100/80-M04-100% (Russian State Standard GOST)/ Abrasive grit ISO 8486-86 FEPA 42L 160/125–80/63, F12/F16–F24/F30, grain ISO 8486-86 AC20, post ISO 8486-86 M1, M13, ПM1, MК

Identification of Deformations and Errors …

969

The impact of the time of heating on the accuracy of the machined surface can be reduced due to a decrease in the temperature of device elements that can increase the speed of heat removal from electromagnetic coils and enhance the power of the device. Let us consider the four possible variants occurring during machining of a workpart mounted on a magnetic base. Variant 1. A clearance through mounting surface and magnetic base along the center. Grinding without cooling. If the value of temperature deformation of the shape exceeds deformations that occur under an action of the attraction force of the magnetic base, non-flatness of mounting surface of the part and the surface of the base does not affect the accuracy of the shape of the machined surface of the part. The value of a deformation of the shape of a part is determined according to Eq. (4). If the value of temperature deformation of the shape is insignificant, the errors of temperature surfaces of the part and the magnetic base partly or fully influence the accuracy of the machined surface. Variant 2. A clearance in the center, machining with cooling. The errors, caused by the temperature deformation of the shape, non-flatness, mounting surfaces of the part and the magnetic base, are summarized within the limits of machining allowance. Variant 3. The clearances along the edges, machining without cooling. The errors, caused by temperature deformations of the part, are summarized with the errors, resulting from non-flatness of mounting surfaces. Variant 4. The clearances along the edges, abundant cooling. In this variant of machining, the attraction force of the magnetic base counteracts temperature deformations of a workpart as well as in case of machining of the workpart without cooling and with a clearance along the edges. If the values of temperature deformations of the shape are significant, non-flatness of the mounting surface does not influence the accuracy of the machined surface. Consequently, depending on certain conditions of machining, the errors of the mounting surface of the machined part and the magnetic base can both influence and do not influence on the accuracy of the machined surface. The deviation of the shape of the machined surface during surface grinding is determined by a total action of errors with their intrinsic signs. The importance of taking into account the summation of errors considering their signs in designing of technological process of surface grinding can be shown in the following practical example. Let us examine total errors from non-straightness of worktable movement with shape errors due to temperature deformations of a workpart. We carried out the research on the accuracy of the shape of the parts machined by eight surface grinders of 3G71 (Russian State Standard GOST) type with different exploitation life. These grinders are installed in the tool and overhaul workshops of Udmurt Republic enterprises. As for new machining stations, the positive value of geometric errors was typical, though it decreased with the growth of exploitation life, as for machining stations with a long exploitation life—negative value of the errors.

970

T. N. Ivanova

During the examination without cooling, the errors from temperature deformation of the shape were identified to have “+” sign. Consequently, the improvements of the accuracy of the shape of machined parts can be expected during the growth of exploitation of the machining station. The flat parts with size 200  40  3.5 mm made of steel 12XH3A/3415 were machined during the experiment. The parts were mounted on the magnetic base. Grinding mode had the following parameters: td = 12 m per minute, t = 0.02 mm, St = 0.1 mm per stroke. We used the tool 150x32x40 AC6 100/80-MO4-100% (Russian State Standard GOST)/Abrasive grit ISO 8486-86 FEPA 42L 160/125–80/63, F12/F16–F24/F30, Grain ISO 8486-86 AC20, Post ISO 8486-86 M1, M13, ПM1, MК with continuous cutting surface. Four working strokes were fulfilled, and allowance removed was 0.08 mm. Machining time was equal to 8 min. After allowance removal, spark-out was conducted until the full stop of sparks. The profile of the machined part was measured with a microscope. All the parts were in a cold condition before the experiment. The temperature error of the grinder was accepted as zero. To exclude an influence of a machining station error, each grinder machined only one part. The magnetic base was grinded before the machining process. The results of conducted research demonstrated that there is a tendency for the accuracy of the machined shape of part to increase with the growth of an exploitation life. The mentioned statements are connected with non-straightness of the profile in the longitudinal direction. As for the dependence of a change of non-straightness in cross section, it had a random character. The regularity of an increase in the accuracy with the growth of an exploitation of machining station was also observed for non-flatness. That is, during grinding of the given workparts on new machining stations, the geometric error was positive and temperature deformations of the shape of the part had “+” sign, that is why their overlapping and summarizing of errors occurred. Worn machine slideways have the negative geometric error, while temperature deformations of the shape of the part have “+” sign, so in this case there is partial mutual compensation. When workparts were grinded with cooling, the other character of change of the shape error of parts was distinguished with the growth of exploitation time of machining station. The accuracy of the shape was higher for new grinders. The error caused by the temperature deformations of the part during machining had “−” sign and mutual compensation was observed during grinding with new machining stations. Thus, in case of identifying shape errors, geometric error, temperature errors of machining station and temperature deformations of shape cause different errors on the machined surface, but with opposite sign. The so-called mirror reflection occurs. Observed regularities prove the expediency of considering error signs in designing of technological operations of surface grinding of thin plates. Existing measures for termination of errors caused by thermal deformations of parts during surface grinding include the use of abundant cooling; artificial deforming of the machined part in the opposite direction to temperature deformation before grinding; setting the trajectory of relative movement of the machined part; and the grinding wheel which compensates temperature deformation. They

Identification of Deformations and Errors …

971

have significant drawbacks, as they require considerable material expenses or can compensate only one certain value of thermal deformation, which changes considerably depending on applied grinding conditions, including cutting modes. The value of non-flatness is equal to sag of span of a thin part and is determined according to formula (4). Resulting value of non-flatness is compared to the allowance for non-flatness of a part. By changing geometric parameters, for example, decreasing the diameter and the thickness under the same conditions, deviation of a plate can exceed the allowance for non-flatness of the part. The analysis of the given solution demonstrates that the value of non-flatness due to temperature deformations is directly proportional to initial temperature difference between the workpart and the mounting surface, linear expansion coefficient of workpart and its squared diameter. At the same time, the value of non-flatness is inversely proportional to the workpart thickness. The lower linear expansion coefficient of machined material is, the higher possibility of fluctuations in workpart temperature is. The research on errors during grinding of flat thin workparts allowed us to give the recommendations for identifying, controlling and reducing the error from temperature deformations. It is recommended to take into account the geometric size of a machined plate, linear expansion coefficient of a machined material and the allowance for non-flatness from temperature deformations.

References 1. Dementyev VB, Ivanova TN (2016) Investigation of the surface layer structure of high-chromium and high-strength steels at the variation of the heating temperature. Mater Sci Forum 870:431–436. https://doi.org/10.4028/www.scientific.net/MSF.870.431 2. Ivanova TN, Dolganov AM (2007) Up-to-date equipment in the technology of diamond face grinding of flat surfaces. Institute of Economics Press, Ural Branch of RAS, Yekaterinburg– Izhevsk 3. Ivanova TN (2013) Advanced technologies of the XXI century. Kuprienko, Odessa 4. Ivanova TN, Svitkovsky FYu (2004) Investigation of the regularities and the enhancement of the efficiency of the grinding process for components from hard-to-machine materials. Metal Treat Technol Equip Tools 1(22):22–24 5. Ivanova TN (2005) Investigation of the surface layer structure at grinding. Metal Treat Equip Tools 3(28):30–32 6. Ivanova TN, Galikhanov DD (2012) Investigation of machinability of hard-to-machine steels at grinding. Nat Sci Tech 3(59):173–176 7. IvanovaTN Dementyev VB (2012) Formation of the properties of the surface layer of components from hard-to-machine materials simultaneously at heating and cooling. Chem Phys Mesoscopy 15(4):587–598 8. Dementyev VB, Ivanova TN (2012) Investigation of the physical-mechanical properties of steel 8CrV after high-speed thermal process. Chem Phys Mesoscopy 4(4):7–18 9. Zakharenko IP (1985) Ultra-hard materials in tool production. Higher School, Kiev 10. Zverovshikov VZ (2005) Dynamics of the centrifugal treatment of components with discrete grinding material. Penza State University Press, Penza 11. Volkov DI, Koryazhkin AA (2012) Finite-difference calculation of the temperature in belt grinding. Russ Eng Res 32(3):296–298

972

T. N. Ivanova

12. Volkov DI, Koryazhkin AA (2012) More precise machining by regulating tool-blank contact in the belt grinding of gas-turbine blades. Russ Eng Res 32(9–10):698–701 13. Kozlov AM (2005) Model-based determination of the parameters of the working surface of an abrasive tool. Proc Inst High Educ 1:51–55 14. Nosenko VA, Nosenko SV (2013) Technology of metal grinding. TNT, Stary Oskol 15. Podzei DV (1973) Technological residual stresses. Machine-building, Moscow 16. Popov SA, Malevsky NP, Tereshenko LM (1977) Diamond-abrasive treatment of metals and hard alloys. Machine-Building, Moscow 17. Sulima AN, Shulov VA, Yagodkin YuD (1988) Surface layer and service properties of machine components. Machine-Building, Moscow 18. Yasheritsyn PI, Martynov AN (1983) Finishing of metals in machine-building. Higher School, Minsk 19. Trent EM (1990) Metal cutting. Butterworth, London 20. Kalpakjian S (1989) Manufacturing engineering and technology. Addison-Wesley Publishing Company

Novel Method of Single-Pass Threading by Cutter N. N. Zubkov

Abstract The technological aspects and some results of a new method intended for producing threads and projections of the triangular profile on plastic metals are considered. The method is implemented according to the schemes of traditional machining with a simple tool such as a cutter. The method is intermediate between the threading based on the cutting process with a removal of the workpiece material and the threading based on metal cold flow with a redistribution of the workpiece material and combines both methods. The new principle and high value of the tool rake angle improve metal flow conditions. This leads to a significant reduction in the amount of material removed in the form of chips, reducing the required depth of cut, high surface quality of the thread, possibility of using smaller diametrical dimensions of the workpiece and possibility of using thin-walled tubes for threading. The authors theoretically justified and experimentally confirmed the selection of a manufacturing parameter to obtain a threaded profile of the required geometry.













Keywords Thread Threading Cutting Thread cutter Deformational cutting Ductile metals




1 Introduction Technological methods for the forming of threaded profiles consist of the process of cold flow and the process of cutting [1]. Cold flow by thread-rolling dies and thread rolls is highly productive and produces no waste. Cold flow for external and internal threading is widely used in mass and large-scale production; it is a well-studied technology [2]. However, the tools for external threads are specialized and difficult to manufacture; they have high N. N. Zubkov (&) Bauman Moscow State Technical University, 2-ya Baumanskaya Str. 5, Moscow 105005, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_103

973

974

N. N. Zubkov

cost. Thus tools are designed to produce thread only of given size. The threads are formed by using the plasticity of material. Another disadvantage is large radial loads, which make it impossible to use this method for forming threads on thin-walled tubes [3, 4]. Medium- and small-scale production facilities normally use threading taps, dies, thread mills and thread cutters. A thread mill is a universal tool, but it costs a lot and its use requires special equipment [4, 5]. Another problem of thread milling is obtaining high-precision threads [6]. Threading with thread cutter on the lathes is another widely used method. This method allows using one tool to obtain threads with different pitches on the workpieces of different diameters. Low productivity is the disadvantage due to the required multi-pass processing. For ductile metals, the thread surface quality achieved with this method is insufficient. This disadvantage arises due to the impossibility of providing a large rake angle at the same time on both cutting edges of the tool [7]. Obtaining satisfactory quality of the thread surface is difficult for such tough and ductile metals as copper, deformable aluminium alloys, pure titanium and others, even with many cutter passes. The work presented here employs deformational cutting (DC) method [8] in order to produce threads on ductile metals. In DC, the chips remain on the workpiece as a functional part of it. This is the main difference compared to traditional cutting. DC method is multifunctional, non-waste and implementable according to the schemes of traditional machining using standard metal-cutting equipment. DC tool is a cutter with a working part formed by three intersecting planes. Differences from the conventional cutter (e.g. turning tool) consist in special geometric parameters of the tool sharpening, by virtue of which the cutting process is possible only on the main cutting edge. The auxiliary edge of the tool does not destroy and only plastically deforms the undercut layer so that it remains in place on the processed surface in the form of projection [9]. DC method has a wide range of applications [10]. DC machining allows to increase the surface area up to 12-fold, which is essential in heat exchange processes [11]. DC is used in manufacturing of boiling surfaces and capillary structures for heat pipes [12], alongside a range of applications in electrical connectors [13] and slotted screen pipes [14]. DC method can also be used for surface quenching [15, 16]. A simple modification in the implementation scheme of the DC method, namely changing the tool feed direction to the opposite, leads to a qualitatively new process of surface machining. This process allows cutting triangular projections with symmetrical and asymmetrical profile, both on flat and cylindric surfaces. For round workpieces, the resulting spiral triangular projections are similar to threads. The purpose of this paper is to estimate the technological parameters for one-path IDC threading.

Novel Method of Single-Pass Threading by Cutter

975

2 Theory Behind the Process Consider a single-pass tool for DC on smooth surface, e.g. in planning (Fig. 1a), with only the main tool movement. The material is undercut by the cutting edge 1. The deforming edge 2 of the tool has large negative rake angle and not capable of cutting. The cut layer (vertical hatching in Fig. 1a) is extruded by the tool face in the form of a ridge on the workpiece surface. The volume of extruded ridge material must be equal to the material volume of the embedded tool part. Traditional DC process suggests that in addition to the main cutting movement 3, the tool must have a feed movement 4 on the distance p, and the direction of the feed must be towards the cutting edge 1. Figure 1b shows the DC process on the third pass of the tool. The undercut layer (vertical hatching on Fig. 1b) with parallelogram-shaped cross section bent into a previously formed groove, forming a fin. The feed distance p in DC process is typically smaller than the cutting depth t. Consider a variant when the tool feed 4 moves in the direction towards the deforming edge 2 (Fig. 1c). The feed value p is greater than the cutting depth t and

Fig. 1 Influence of feed direction generated by the DC profile. a Single pass of the tool, b feed direction 4 is to the cutting edge 1, c feed direction 4 is to the deforming edge 2

976

N. N. Zubkov

selected so that the cutting process takes place at the maximum length (AC) of the cutting edge 1. In this case, the cutting edge (its full-length AB) will partially cut previously formed ridge, and section BC of the cutting edge will undercut a layer of workpiece material. The tool will simultaneously extrude the material on the surface of the workpiece in the form of projection the same way as shown in Fig. 1a. with a single pass. Repeated passes will generate on a blank surface a relief in the form of triangular projections and depressions. Projections rise above the initial surface on size DH, i.e. the size after machining increases. This process is the basis of thread forming called inverted deformational cutting (IDC). Unlike wasteless DC method, the IDC process is accompanied by partial formation of chips. In contrast to traditional cutting, when all section of the cut layer leaves the blank as a chip, the IDC part of the undercut layer remains on the blank surface. The scheme of IDR on a flat surface is shown in Fig. 2. Triangular profile sides inclination angles determined by the projection angles of cutting and deforming edges, i.e. by side cutting edge angle u and end cutting edge angle u1. With their equality, the method provides symmetrical threaded profiles. For cylindrical surfaces, the diameter of the blank, measured at the tops of thread profile, is more than initial workpiece diameter. At a predetermined ratio of the cutting depth t and the feed value p, the minimum volume of the previously obtained protrusion is cut, if the complete threaded triangular profile is formed. The calculation scheme is shown in Fig. 3, where p—the

Fig. 2 Forming the triangular ridge profile on the plane by IDR method. 1–4 the same as in Fig. 1, 5—tool, 6—workpiece, 7—formed relief, 8—extruded ridge, 9—the part of earlier formed ridge removed as chip

Novel Method of Single-Pass Threading by Cutter

977

required thread pitch, d—the required outer diameter of the thread, H—the required height of the thread profile, u and u1—the angles of the threaded profile. The criterion to obtain dependencies was forming a full thread profile with minimum waste removed from the blank. The following assumptions used: • straightness and perpendicularity of protrusion-free edge at the stage before the next tool pass (line AB in Fig. 3), • free side of the protrusion has no radius transition to the treated surface (no radius at point A). The common case is asymmetrical triangular threaded profile when u 6¼ u1. The solution for selecting the initial diameter of the workpiece d0 and cutting depth t is based on the equality of material volume interacting with the embedded part of the tool, and the volume of the ridge is extruded by the tool on the treated surface. To obtain a non-symmetrical threaded profile with the specified parameters by the IDR method, the cutting depth must be: pffiffiffiffiffiffiffi tgu t ¼ p
pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi pffiffiffiffiffiffiffi ðmmÞ: ðctgu þ ctgu1 Þ þ ðctgu þ ctgu1 Þ
tgu if the resulting profile height is

Fig. 3 Scheme for IDR parameters’ calculation

978

N. N. Zubkov

H¼

p
tgu
tgu1 ðmmÞ; tgu þ tgu1

workpiece initial diameter d0 determines from the expression:   p
tgu
tgu1 d0 ¼ d  2
ðH  tÞ ¼ d  2
 t ðmmÞ: tgu þ tgu1 General cases are required to obtain a symmetrical threaded profile (u = u1), for which the calculation formulas are simplified. For metric threads: t¼

p
tgu pffiffiffi ðmmÞ; 2þ 2

p
tgu ðmmÞ: d0 ¼ d  p
tgu þ 2t ¼ d  pffiffiffi 21 For pipe and inch threads, when the pitch is set by the number of threads z per inch: t¼

tgu pffiffiffi ðinÞ; z
ð2 þ 2Þ

d0 ¼ d 

tgu pffiffiffi ðinÞ: z
ð 2  1Þ

For metric threads u = u1 = 60°, and for pipe and inch threads u = u1 = 62.5°.

3 Experiments and Results And for the usual DC process, the tool face position for IDR must meet the conditions ensuring the cutting process on the main cutting edge (large positive rake angle) and not separating the undercut layer from the workpiece on the deforming edge (large negative rake angle) and ensuring sufficient strength of the cutting tool [17]. Theoretically, it is not possible to find a solution that satisfies these interdependent conditions, so the optimal position of the tool face was found based on experimental results. The first aim of the experiments was the determination of the ranges for rake angle c and cutting edge inclination angle k, which can be used in IDC with sufficient result. Those two angles determine the position of rake face. The position of the tool face is changed by cutter grinding. As a sufficient result was the fact of

Novel Method of Single-Pass Threading by Cutter

979

projection formation with the diameter larger than original without any chip separation and without tool breakage. The experiments carried out under the following conditions. Equipment is turning lathe 16K20. The tool material was cemented carbide (92% WC and 8% Co). Workpiece material was copper M2 (impurity not more than 0.3%), cutting depth t = 1.1 mm and feed value S0 = 2.0 mm/rev. Constant tool geometrical parameters were as follows: • side a and end relief a1 angles were a = a1 = 1.5°, • side cutting edge angle u and end cutting edge u1 angle were u = u1 = 60°. Experiments founded that IDR process in machining copper M2 is feasible with a change in the rake angle c in the range 43°–57° and when cutting edge inclination angle k is within 25°–41°. The tool with values c and k in excess has insufficient strength of the cutting edge. A tool with smaller angles works like the traditional cutter and forms chips. The most stable IDC occurs when the angles are c = 53°– 57° and k = 34°–39°. When testing the method on stainless steel X12CrNiTi18-9 and steel St.3 (analogue steel 304), the formation of the ridge on a treated surface is achieved only at thread small pitches (p up to 1 mm). With pitch increasing, the plasticity of these materials is not enough to compensate for the increased deformations, which leads to the formed ridge separation as a chip. A set of experiments were carried out when cutting threads on copper grade M2. Thread pitches were from 0.5 to 2 mm, and the angles at the top of the threaded profile were 55° and 60°. High speed steel tools were used in the experiments. According to the selected technological parameters, a threaded profile was cut, after which the characteristics of the profile of the obtained thread were measured. The photograph of the thread M20  1.5 obtained on the copper tube is shown in Fig. 4. When cutting depth t is less than the calculated t, the thread profile is not fully formed. This is clearly seen on the right side of the photograph (Fig. 4). Forgetting pipe thread G 1/2″ used copper pipe billet with outer diameter 20.0 mm and 2.0 mm wall thickness. Side cutting edge angle u and end cutting edge u1 angle were u = u1 = 62.5°. Before threading tube was to calculated diameter d0 = 19.14 mm. The lathe was set on the threading with a pitch, 14 threads per inch. IDC was carried out with the calculated cutting depth t = 0.833 mm. The cutting speed was V = 11.3 m/min (spindle rotated 200 rpm). For receiving metric thread 25  1.5 used copper pipe billet with outer diameter 24.0 and 1.5 mm wall thickness. Side cutting edge angle u and end cutting edge u1 angle were u = u1 = 60.0°. Fig. 4 Crosscut of M20  1.5 thread on copper tube 20  1.5 thread made on copper tube

980

N. N. Zubkov

Table 1 Comparison of the obtained thread parameters with the standards G 1/2″

M25  1.5

d1 d2 d3 d1 d2 d3

(mm) (mm) (mm) (mm) (mm) (mm)

ISO R228

Measured

20.955 19.793 18.631

20.9 19.7 18.4

ISO 724:1993

Measured

25.00 24.026 23.160

24.9 23.9 23.0

Fig. 5 Obtained threads: G 1/2″ (on the left) and 25  1.5 (on the right)

Before threading tube was cut to calculated diameter d0 = 23.64 mm. The lathe was set on the threading with a pitch 1.5 mm. IDC was carried out with the calculated cutting depth t = 0.62 mm. The cutting speed was V = 14.2 m/min (spindle rotated 200 rpm). Comparison of the obtained thread parameters with the standards is presented in Table 1, where d1, d2 and d3 are nominal, pitch and minor diameters, respectively. The parameters of the obtained threads are very close to the standard ones. Photographs of the tube samples with obtained threads are presented in Fig. 5.

4 Discussion The method of obtaining threads by IDC is intermediate between the threading based on the cutting process with the removal of the workpiece material and the threading based on metal cold flow with redistribution of the workpiece material and combines both methods. The new principle and high value of the rake angle improve metal flow conditions. This leads to a significant reduction in the amount

Novel Method of Single-Pass Threading by Cutter

981

of material removed in the form of chips, reducing the required depth of cut, the high surface quality of the thread, availability of using smaller diametrical dimensions of the workpiece and thin-walled tubes for threading. At present time, the industry, due to the significant cost of copper, is moving to the use of heat exchangers with thin-walled copper tubes having a wall thickness of 1.0–1.5 mm. Tube ends’ expanding of thin-walled tubes in tube plate decreases the reliability of the connection compared with tubes having a wall thickness of 2.0 mm used before. An alternative variant for fixing the tubes in tube boards is a threaded fastening (soldering and welding are practically not used due to many technical problems). At the same time, the small wall thickness does not allow to form a thread by known methods. For example, the method of IDC at the ends of heat exchange tubes with the most widely used outer diameters of 16.0 and 20.0 mm allows obtaining pipe threads 3/8″ and 1/2″, respectively. The method provides a standard profile with a diameter at the tops and profiles’ height 0.86 and 1.16 mm, respectively. With an initial wall thickness of 1.5 mm, the residual wall thickness of the tube will be 1.0 mm (for a tube with a diameter of 16.0 mm) and 0.8 mm (for a tube with a diameter of 20.0 mm).

5 Conclusion 1. The new method of thread forming proposed, which makes it possible to obtain a threaded profile in a single pass of the tool on materials having high plasticity. 2. Theoretically justified and experimentally confirmed the manufacturing parameter choice to obtain a threaded profile of the required geometry. 3. The method can be used for threads forming at the ends of thin-walled heat exchange tubes. Acknowledgements Work was supported by the Ministry of Education and Science of Russia (Grant № 9.5617.2017/VU) and RFBR according to the research project № 16-08-00489.

References 1. Franklin DJ (2017) Thread-cutting methods: a treatise on the operation and use of various tools and machines for forming screw threads, including the application. Chizine Pubn 2. Henryk Czarnecki H et al (2016) Changes in the properties of the surface layer of the internal thread made by cold forming. Tribologia 270(6):11–20. https://doi.org/10.5604/01.3001. 0010.6704 3. Degarmo JT et al (2011) Materials and processes in manufacturing. Wiley, New York 4. Franklin DJ (2018) Thread-cutting methods: a treatise on the operation and use of various tools and machines for forming screw threads, including the application of lathes, taps, dies, standard and special attachments, thread-milling machines, and thread-rolling machines. Creative Media Partners, Sacramento

982

N. N. Zubkov

5. Danilenko BD (2015) Cutting conditions for thread mills. Russ Eng Res 35(1):76–77. https:// doi.org/10.3103/S1068798X15010098 6. Mal’kov OV (2013) Precision of the external thread profile in thread cutting. Russ Eng Res 33 (3):172–175. https://doi.org/10.3103/S1068798X1303012X 7. Klocke F (2011) Manufacturing processes 1: cutting. Springer, Berlin. https://doi.org/10. 1007/978-3-642-11979-8 8. Zoubkov N, Ovtchinnikov A (1998) Method and apparatus of producing a surface with alternating ridges and depressions. US Patent 5,775,187. 07 Aug 1998 9. Zubkov N, Sleptsov A (2016) Influence of deformational cutting data on parameters of polymer slotted screen pipes. J Manuf Sci Eng 138(1):011007. https://doi.org/10.1115/1. 4030827 10. Kukowski R (2003) MDT—micro deformation technology. In: Proceedings of ASME 2003 international mechanical engineering congress and exposition. Washington, pp 305–308 11. Thors P et al (2013) Method for making enhanced heat transfer surfaces, US Patent 8,573,022, 05 Dec 2013 12. Yakomaskin et al (2013) Investigation of heat transfer in evaporator of microchannel loop heat pipe. J Heat Transfer 135(10):101006. https://doi.org/10.1115/HT2012-58503 13. Solovyeva et al (2012) Novel electrical joints using deformation machining technology-Part II: experimental verification. IEEE Trans Compon Packag Manuf Technol 2(10):1718–1722. https://doi.org/10.1109/TCPMT.2012.2199755 14. Zubkov NN, Sleptsov AD (2010) Production of slotted polymer filter tubes by deformational cutting. Russ Eng Res 30(12):1231–1233. https://doi.org/10.3103/S1068798X10120117 15. Zubkov N et al (2017) New method of quench surface turning. Solid State Phenom 265:696– 701. https://doi.org/10.4028/www.scientific.net/SSP.265.696 16. Zubkov N et al (2018) Steel case hardening using deformational cutting. J Manuf Sci Eng 140 (6):061013. https://doi.org/10.1115/1.4039382 17. Zubkov N et al (2016) Tool-workpiece interaction in deformational cutting. Russ Eng Res 36 (3):209–212. https://doi.org/10.3103/S1068798X16030217

Chip-Forming Processes at High-Speed Grinding V. M. Shumyacher, O. G. Kulik and S. A. Kryukov

Abstract This article presents the results of microanalysis obtained at high abrasive and metallic sludge increase, which can be used to explain in the mechanisms of chip formation, as well as the course of physical and mechanical processes occurring on the surface layers of the ground blanks. By controlling these processes, it is possible to achieve a reduction in thermal stress in the pair “abrasive grain—workpiece,” which, in turn, will allow obtaining the required quality of the surface layer of the workpieces and the specified dimensional accuracy with the highest productivity of abrasive processing. The reduction of thermal stress of the process in the condition of highest productivity will allow obtaining the required quality of the surface layer of the workpieces and the specified dimensional accuracy. The reduction of thermal stress of the process can be achieved by controlling the chip formation processes at high-speed grinding and selecting the appropriate ratio between the cutting speed and other processing parameters. Keywords Grinding


Chip formation
Chip types
Dispersing products

1 Introduction Improving the product quality at the enterprises of the engineering industry is associated with a constant increase in requirements for the accuracy of the geometry of the specified parameters of machine parts and the quality of their working surfaces. In this regard, a particularly important issue is to increase the efficiency of finishing processes, the main of which is grinding. This is a complex physical, mechanical, and chemical process, accompanied by heat generation, deformation of V. M. Shumyacher
O. G. Kulik (&)
S. A. Kryukov Volzhsky Polytechnic Institute (branch) Volgograd State Technical University, 42a Engels Street, Volzhsky, Volgograd Region 404121, Russia e-mail: [email protected] V. M. Shumyacher e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_104

983

984

V. M. Shumyacher et al.

the metal, and tool material wear. Knowledge of the regularities of the chip formation process and the allocation of priority factors allow you to efficiently manage this process and ultimately produce parts more efficiently, productively and economically.

2 Theoretical Part The surface defects prevention is an intensification of the grinding process, which, in turn, is possible only if the negative effect of the thermal factor is eliminated. According to numerous experimental and theoretical studies, the grinding process is characterized by the following temperature indexes [1]: • the average temperature of the Tav part, fluctuating within 20  350 °C in the cutting zone; • maximum contact temperature Tmax = 200  1100 °C; burns and cracks may appear at this temperature if Tmax > 500 °C; • instantaneous temperature T0 at the surface point, where at the moment cutting with one abrasive grain occurs. The value of this temperature, depending on the grinding conditions, ranges from 1000 °C to the melting temperature of the treated metal [2]. As a result, uncontrolled chemical reactions, diffusion processes, and phase transformations occur, and burning, thermal deformations, microcracks, and residual stresses appear on the treated surface. However, the assessment of the temperature distribution and time of their action when grinding in a particular local volume of the workpiece part requires detailed knowledge of the distribution of the density of heat fluxes on the treated surface at any point in time. According to numerous experimental and theoretical studies [3–5], the maximum temperature Tmax of the metal mass in the grinding zone cannot exceed 500 ° C, since at temperatures Tmax > 500 °C, burns, cracks, and other defects associated with thermophysical phenomena during steel heating can occur. At the same time, according to theoretical concepts, expressed by D. Jager’s formula [6], the instantaneous temperature T0 in the contact “abrasive grain—workpiece” under the conditions of an adiabatic process can exceed 2000 °C: T0 ¼ k

f
p
ðl
mj Þ0:5 ðk
c
Þ0:5

;

ð1Þ

where k is the coefficient; f is the friction coefficient of the abrasive grain on the ground metal; P is the specific pressure in the contact; l is the contact length; Vк is the speed of rotation of the circle; k is thermal conductivity of the metal; C is the heat capacity of the treated metal.

Chip-Forming Processes at High-Speed Grinding

985

The calculation of the actual temperatures by Formula (1) is difficult. Therefore, the abovementioned formula should be considered as the sum of factors affecting the metal instantaneous heating in the contact microzone of the abrasive grain with the metal. In this case, the values f, P, l, and V reflect the energy consumption of the process, and the coefficients k, c, and c are the thermophysical properties of the material being grinded. The friction coefficient f multiplied by the force of the normal pressure P on the displacement value l characterizes the work done with the V speed. In this case, cofactors f
P, in our opinion, are more legitimate to imagine, as the sum of the efforts R(Z, Y), reflecting the total resistance to displacement from plastic and elastic deformations during shear, plus friction of dispersed elements and grains on fixed surfaces. At the same time, the yield tensile strength changes from the initial one depending on its temperature. At h > 2000 °C heating temperatures exceeding the melting temperature of steel at 1350  1500 °C, Formula (1) loses its meaning, since the resistance forces to deformations determined by the melt viscosity are several orders of magnitude lower than the initial shear strength of metal and friction forces between solids.

3 Hypothesis Assessment Experimental determination of the thermal and physical characteristics of the shaped metal (k, c, c), which is in a solid state of aggregation at temperatures above the melting point, is problematic. Therefore, there are limitations in the application of this formula for practical calculations of abrasive treatment zone temperatures. However, our structural analysis of its cofactors gives a theoretical model of the physical processes occurring in the contact zone of the abrasive grain with the treated material, which is valuable when there is a sensible choice of technological grinding modes during their intensification. Due to the increase in specific removal rates, there is an increase in pressure along the contact arc of the metal with the grain moving along the path of the circle with the moving center in the direction of the longitudinal feed. Instantaneous temperature increases adequately to pressure at each point of contact. From practical experience of steel heat treatment, it is known that flash heating (HFC) shifts the points of onset of austenite formation compared with equilibrium (furnace) heating by more than 25%. This can explain the physical probability of the absence of burns and cracks in the presence of instantaneous temperatures exceeding the point of onset of thermal transformations during high-speed abrasive processing.

986

V. M. Shumyacher et al.

From the Clapeyron–Clausius differential equation: dP Q ¼ dT T
DV

ð2Þ

where Q is the heat of transformation; V is volume change during the transformation; T is the temperature; P is pressure follows that when ddTP [ 0, the melting point rises with increasing pressure [4, 7–10]. By such pressure increase, we explain the displacement of the melting–hardening phases at the contact points (Fig. 1a, b). Accelerated heating, adequate to an increase in pressure, with a peak temperature comparable to the equilibrium melting point Tф  TL, creates conditions for overheating and a shift of the melting phases at the point of contact of the grain with the metal (Fig. 1b). This means that when grinding in the contact zone of the cutting grain with the metal, there may be a compacted superheated volume of the metal (cutting core) in the solid phase at a temperature above the melting temperature determined in the equilibrium state, that is, at the minimum overpressure and minimum heating rate, with a time sufficient for phase transformations. Since the amount of heat released during the core shift of the compacted metal exceeds the amount of heat required to melt this core, at the extremely short time of passage of the contact zone (4.3  10−5 s), the core has the potential to melt outside this zone, using the accumulated heat. In this case, the higher the rate of passage of the metal removed through the contact zone, the higher the probability of formation of melted balls in the sludge and the lower the probability of sticking of the softened metal to the surface of the abrasive grains and intergrain space.

Fig. 1 A fragment of the constitutional diagram iron– carbon: a equilibrium heating; b accelerated heating; TL is the melting point on the liquidus line (1500 °C); TS is solidus line solidification point (1450 °C); TF is the actual heating temperature; TL1 is point on the liquidus line with accelerated heating; TS1 is point on the solidus line with accelerated heating

Chip-Forming Processes at High-Speed Grinding

987

This explains the existence of a superheated metal (core) at an instantaneous temperature exceeding the melting point under the conditions of an adiabatic process in the contact zone “abrasive grain—workpiece” [11, 12]. The higher the rate of passage of the removed metal through the contact zone, the higher the probability of formation of melted balls in the sludge and the lower the probability of the softened metal sticking to the surface of abrasive grains and intergrain space. Thus, the presence of melted metal balls with chips ejected from the treatment area can be explained by the participation in the grinding process of an intermediate link—a cutting core made of compacted processed metal (Fig. 2).

4 Experimental Confirmation of the Hypothesis The diameter of spherical particles, as a rule, is 1–80 lm, [13, 14]; spherical chips with a diameter of up to 300 lm are encountered [15, 16] (Fig. 3a, b). The size and number of spherical shavings depend on the chemical composition of the material being processed [14], heat treatment, and cutting conditions. The number of spherical chips in the sludge correlates with the contact temperature in the cutting zone. Flow grinding chips that appear during machining steel, as noted by different authors [16, 17], are long curved ribbons of different thickness and width with thin torn edges (see Fig. 4c, d, i, f). Rounded tops of edges (Fig. 4d, f) are formed as a result of the particles reaching the temperature of red heat in the air after their separation from the material being processed and subsequent hardening. An increase in cutting speed from 35 to 100 m/s leads to a noticeable change in chip size. At high speeds in the composition of the chips, a significant percentage of small chips appears.

Fig. 2 Structure of the grinding sludge: 1—spherical chips; 2—flow chips; 3— chipped grain

988

V. M. Shumyacher et al.

Fig. 3 Chips during microcutting with a single abrasive grain from cubic boron nitride: a steel Х18Н10Т Vk = 100 m/s; b steel 45 Vk = 100 m/s

Fig. 4 Pictures of polished steel chippings 45 (NIS 48 … 52) around 24A25SM17K43: a Vk = 35 m/s (1000); b Vk = 35 m/s (4000); c Vk = 100 m/s (1760); d) Vk = 100 m/s (2100); e Vk = 35 m/s (1760); f Vk = 35 m/s (800) [16]

Chip-Forming Processes at High-Speed Grinding

989

The appearance and structure of the chips at cutting speeds of 35–100 m/ s and at high magnifications (Fig. 4b, d) allow us to draw the following conclusions: • There is a distinct difference between the inner cutter and the outer side of the chip; • The bearing surface side is a continuous tape, sometimes slightly concave due to the convex shape of the front surface of the cutting edges; • Longitudinal grooves from plastic deformation, shelling of metal particles, tears and voids, traces of burns and melts, indicating high temperature and severe plastic deformation during chip movement along the front surface of the abrasive grain are well visible; • A thin continuous layer of plastically deformed metal (up to 0.05 lm) of the inner side as if holding onto itself the elements of shear entering on the outer side of the chip. The results of electron microscopic studies clearly showed the presence of chips, the so called worn places, on the tops of the aluminum oxide and cubic boron nitride grains (Fig. 5) in the process of microcutting of workpiece materials at a speed of 35 and 100 m/s. Micrographs (Fig. 5) clearly show these worn places. They are covered with light layers of the secondary structure. Moreover, the density of the resulting structure is directly proportional to the speed of processing. Accordingly, the higher the density, the lower the thermal conductivity of the abrasive grains under this formed structure. In the photographs (Fig. 5a, b), on both surfaces of the aluminum oxide grains, at both speeds of microcutting, relatively smooth faces of chipped grain are visible, covered with layers of secondary structures adhering to them with typical grooves 2–20 µm wide, which indicate phase transformations of the secondary structures that occurred after the grain leaves the zone of its contact with the treated metal, i.e., structural and chemical modulation of the core, from overcompacted treated metal remaining on the grain surface. On the surface of the cubic boron nitride grains with an increase in cutting speed from 35 to 100 m/s, adhered layers of the secondary structure became more dense and solid (see Fig. 5d). The main surface of the wore places in this case becomes covered with microcracks, indicating a high brittleness of the formed secondary structure under cutting conditions at a speed of 100 m/s [16]. The polished surface of the workpiece is a set of grinding marks left after high-speed microcutting. The formation of each grinding mark occurs as a result of the successive process of penetrating the grain cutting edge into the processed surface, the elastic slip of the formed blunting site of the top of the abrasive grain (cleavage), the formation of treated metal beneath it over the compacted core blade, which provides plastic edging of metal with the formation of sagging and by cutting off the metal layer when reaching the critical depth of penetration. The photographs (Fig. 6) show that there are grooves inside single grinding marks, the width of which, 2 … 20 lm, practically coincides with the width of grooves observed earlier on the surface of abrasive grains (see Fig. 5a, b).

990

V. M. Shumyacher et al.

Fig. 5 Surface areas of abrasive grains from aluminum oxide (a, b) and cubic boron nitride (c, d) after grinding P18 steel with a speed of 35 m/s (a, c) and 160 m/s (b, d): a 1200; b 700; c 1200; d 1100

When microcutting at a speed of 35 m/s, scratches and shelling of fibers of highly deformed metal layers appear on the surface of the marks, especially after microcutting with aluminum oxide grain (Fig. 6a). Along the sides of the grinding marks, there is an influx of metal with ragged edges. With a microcutting speed of 100 m/s, the surface of the marks becomes more uniform (Fig. 6b), and the microhardness of the bottom of the marks, their width and relative height of sagging, decreased [16, 18].

Chip-Forming Processes at High-Speed Grinding

991

Fig. 6 Risks formed by microcutting of a single aluminum oxide grain 24A on a sample of steel Х18N10T: a at a cutting speed of 35 m/s; b at a cutting speed of 100 m/s—2000

5 Theoretical Analysis of the Experiment Results A similar pattern on the metal surface processed by grinding can be achieved as a result of plastic impact [19]. According to the data [21], at the time of cutting the abrasive grain into the metal, flow ranges appear in it. The source of forces acting during grinding is wave shock pulses. As a result, during a high-velocity microcutting in front of a moving abrasive grain, creating a periodic stress concentration, an unloading plastic deformation wave is formed, ensuring its relaxation, as a result of local structural and phase transformations of the deformable crystal, which generally remains structurally stable. The period of deformation localization (chip segment size) depends on the elastic and plastic state, on the thermal and kinetic conditions of high-velocity microcutting, and on the relaxation capacity (structural and concentration) of the workpiece-treated material. The dissipative processes in the contact zone of the workpiece contribute to the periodic formation of high-strength secondary structures that increase the protective wear resistance of the material of the friction pair. Thus, to effectively increase the tools wear resistance, it is necessary to create conditions for the energy balance between the supplied thermal energy and its dissipation due to structural and phase transformations in the workpiece material. In [19], it was shown that the abrasive grain surface experiences pressure that exceeds the temporary resistance of the material being processed. The specific energy of the abrasive grain interaction and the metal being processed during grinding can be determined from the dependence: Nsp ¼ jVp
s
sin a;

ð3Þ

where KVp is coefficient taking into account the pressure increase in the contact “abrasive grain—workpiece”; rB is the tensile strength brake of the treated metal; a is the average inclination angle of the abrasive grain contact to the part surface.

992

V. M. Shumyacher et al.

After transformations (3), we obtain the dependence for calculating the specific grinding energy: Esp ¼ jVp
sB :

ð4Þ

It follows from (4) that the specific grinding energy decreases with a reduction of the tensile strength to fracture of the treated metal according to the coefficient КVp. From (3) and (4), it can be seen that between KVp and V, there is an inverse proportion. When grinding, the interaction of an abrasive grain and workpiece can be described by of three successive stages: impact penetration, displacement, and contact exit. The time of interaction between the grain and the metal is *10−5 s, which corresponds to the implementation of the conditions under which the explosion reaction occurs [4, 7–10]. According to the chip formation model [20], the removal of a processed metal piece is set out by the ratio h/q (h is the depth of penetration of the abrasive grain, and q is the grain rounded radius) and V/J (the rate of interaction of the “abrasive grain—workpiece” pair). Most of the energy of the abrasive grain interaction and the treated metal is expended on the plastic deformation of the treated material, which leads to its heating and, accordingly, to a change in the physical and mechanical properties. Most of the interaction energy of the abrasive grain and the treated metal is expended on the plastic deformation of the treated material, which leads to its heating and, accordingly, to a change in the physical and mechanical properties. Comparing the data of calculations performed according to the results of our experiments when grinding 65G steel with the data of [21–23], it can be noted that the increase in the interaction speed between the “abrasive grain—workpiece” pair leads to a decrease in the specific energy of dispersion of the treated material. The explanation for this, in our opinion, is the effect of an increase in the energy density in the microvolume in front of the grain, which causes a change (increase in plasticity) of the physical and mechanical properties of the metal being processed. Manifestation of this effect is in the reduction of the h/q ratios, in which the microchip with increasing grinding speed.

6 Findings The results obtained in the course of the research can be used to explain the mechanisms of chip formation, as well as the course of the physical and mechanical processes occurring on the surface layers of the grinded workpieces. Chip formation during grinding is represented as a complex of simultaneously occurring processes: Along with the formation of grooves and piles with the front face of the grain, located at an obtuse angle to the direction of its movement, shock compaction and hardening of the growing volume of metal occur with the back face (or blunting area), followed by shearing and separating the compacted volume from

Chip-Forming Processes at High-Speed Grinding

993

the main mass of the treated metal. The separated part of the metal, possessing increased strength and hardness, is itself a cutting element moving with the grain. The instantaneous temperature in the zone of contact of the grain with the metal may exceed the melting point of steel, since the instantaneously increasing pressure, according to the Clapeyron–Clausius law, shifts the melting point in the direction of its increase. In this case, the superheated metal remains solid until it leaves the contact zone. Melting, oxidation, and other processes take place in the free space of the ejection fan and other dispersion products. Thus, studies of chips show that the elemental character of chip formation, characteristic of the adiabatic shear, predetermines the resistance of steel to plastic deformation. As a result, the formation of chips during grinding at speeds up to 100 m/s occurs under conditions of unstable cutting, which leads to the release of a large number heat, the emergence of macroscopic residual stresses of the first kind in the metal of the surface layer of the workpiece, intensive wear of abrasive grains, and dynamic instability in the closing subsystem—the cutting zone. The process of chip formation is regulated by the intensity of the interaction of the pair “abrasive grain—workpiece” and is the result of the accumulation of energy in front of the abrasive, leading to the development of flow zones in the deformable volume, which have significantly changed physical and mechanical characteristics of the metal being treated. By controlling chip formation processes at high-speed grinding, by optimally selecting the appropriate ratios between cutting speed and other processing parameters, a reduction in process thermal density can be achieved, which, with the highest productivity, will allow to obtain the required quality of the surface layer of the workpieces and a given dimensional accuracy.

References 1. Shalnov VA (1972) Grinding and rationing of high-strength materials. Mechanical Engineering, Moscow, p 272 2. Chandrasekar S, Farris TN, Hebbar RR, Hucker SA, Bulsara VH (1994) Thermal aspects of surface finishing processes. ASM MetalsHandbook: Surf Eng 5:152–157 3. Filimonov LN (1979) High-speed grinding. Mashinostroyenie, Leningrad, p 248 4. Bykadorova OG (2008) More efficient deep grinding. Russ Eng Res 28:836–837. https://doi. org/10.3103/S1068798X08080261 5. Shumyacher VM, Slavin AV (2008) The mechanical and chemical principles of abrasive processing management with the help of LCTM. In: Digest of the All-Russia Scientific— Technical conference: new chemical technologies: production and application. Technology of Mechanical Engineering, pp 29–32 6. Carslaw G, Jager D (1964) Thermal conductivity of solids. Nauka, Moscow 7. Bykadorova OG, Shumyacher VM (2003) The mechanism of metal dispersion of abrasive grain. In: Proceedings of international scientific and engineering conference “Shlifabraziv-2003”, Volzhsky, pp 38–39

994

V. M. Shumyacher et al.

8. Bykadorova OG, Shumyacher VM (2004) Relationship management of abrasive grain and metal to be treated in order to increase the efficiency of the grinding. In: Proceedings of XXIV conference of the russian school on the problems of science and technology, Dedicated to the 80th Anniversary of Academician Serbin V.P. vol 2, pp 118–125 9. Bykadorova OG, Shumyacher VM (2004) The mechanism of contact interaction of the abrasive grain and the treated surface during grinding. In: Proceedings of VIII All-Russian scientific-practical conference: modern technologies in mechanical engineering, Penza, pp 54–58 10. Bykadorova OG, Shumyacher VM (2005) The role of the wear platform of tops of abrasive grains in the metal dispersed during grinding. In: Proceedings of III-rd international scientific and engineering conference: materials and technologies of the XXI century, Penza, p 82–85 11. Kulik OG, Demenkov VA (2017) Kinematics and dynamics of chip formation during grinding. Proc Eng 206:210–215. https://doi.org/10.1016/j.proeng.2017.10.462 12. Kulik OG, Demenkov VA (2017) The management of kinematic factors for the improvement of the deep grinding performance. Proc Eng 206:216–221. https://doi.org/10.1016/j.proeng. 2017.10.463 13. Shumyacher VM (1988) Research and development of the sludge processing technology for use in the construction industry and other industries: research record no. 627–87, Volgograd Institute of Civil Engineering, Volzhsky, p 31 14. Ardashev DV (2011) The influence of the chemical composition of grinded steel on the particle size distribution of chips. Metalworking 5(65):41–44 15. Tsokur AK, Tsokur VP (2007) The role of physico-chemical phenomena in the heat balance during grinding. Donetsk National Technical University, Machine Building and Engineering, pp 56–61 16. Zubarev UM, Priemishev AV (2018) Theory and practice of improving the efficiency of grinding of materials. Lan, Saint Petersburg, p 304 17. Ostrovsky VI (1981) Theoretical basis of the grinding process. LSU Publishing House, Leningrad, p 142 18. Zubarev Yu M (2010) Theory and practice of improving the efficiency of grinding materials: teaching medium. Lan, p 303 19. Shumyacher VM, Kadilnikov AV (2007) Model of the interaction of abrasive grain and the material to be processed during grinding. Chip pattern. Manuf Eng 4:18–22 20. Repko AV, Kirianov AG (2004) Experimental determination of the magnitude of the impact force of diamond grains during grinding. Abrasive processes, abrasive tools and materials. In: Collected articles of the international scientific—technical conference “Shlifabraiv - 2004”. Institute of Construction and Technology (branch) Volgograd State University of Architecture and Civil Engineering, Volzhsky, pp 198–203 21. Pushkarev OI, Shumyacher VM (2004) Methods and means of monitoring the physical and mechanical characteristics of abrasive materials: monograph. Volgograd State University of Architecture and Civil Engineering, Volgograd, p 1441 22. Kulik OG, Illarionova ED (2019) Intensification of technical grinding modes by managing physical processes of chip formation. In: Radionov A, Kravchenko O, Guzeev V, Rozhdestvenskiy Y (eds) Proceedings of the 4th international conference on industrial engineering. ICIE 2018. Lecture Notes in Mechanical Engineering. Springer, Cham 23. Shumyacher VM, Slavin AV, Kulik OG (2019) Chip formation mechanism at metal grinding. In: Radionov A, Kravchenko O, Guzeev V, Rozhdestvenskiy Y (eds) Proceedings of the 4th international conference on industrial engineering. ICIE 2018. Lecture Notes in Mechanical Engineering. Springer, Cham

Express Control of Abrasive Tool Operational Characteristics V. M. Shumyacher, S. A. Kryukov and O. G. Kulik

Abstract The analysis of evaluation method of abrasive wear resistance and the correlation welding of their operational and physical and mechanical characteristics has been carried out. Studies of the relative wear resistance of abrasive tools samples in conjunction with the technological and operational characteristics have been performed. A method for determining the relative wear resistance of abrasive tools samples using the “Schlif” instrument of the Volzhsky Research Institute of Abrasives and Grinding Scientific-Technical Center is given. A linear relationship has been established between the wear of a sample of an abrasive tool on the “Schlif” instrument and its operating characteristics during grinding. A method is proposed for monitoring the wear resistance of a grinding tool on model samples that is sensitive to technological and operational factors, which makes it possible to use it for an express analysis of the quality of a tool at the stage of its development. Keywords Wear resistance Structure


Operational characteristics
Abrasive tool


1 Introduction Evaluation of the quality and operational characteristics of an abrasive tool in the process of its design is carried out either under the working conditions of a specific consumer or at research centers of manufacturing companies. This contributes to the growth of unproductive costs in the factories producing abrasive tools and insufficient objectivity of control due to the impossibility of analyzing a large sample of test circles. In the absence of objective methods and means for the operational control of the production technology of grinding wheels, it is difficult to guarantee the stability of V. M. Shumyacher (&)
S. A. Kryukov
O. G. Kulik Volzhsky Polytechnic Institute (Branch) Volgograd State Technical University, 42a Engelsa Street, Volzhsky, Volgograd Region 404121, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_105

995

996

V. M. Shumyacher et al.

its quality, the most important characteristic of which is the wear resistance of the tool that determines the durability and grinding coefficient. When choosing a foreign-made abrasive tool, domestic consumers face a serious problem, the essence of which lies in the fact that the world’s leading manufacturers usually specializes in a specific operation performed on given equipment. As a result, the costs of abrasive tools are increased dramatically, while productivity and quality of processing are reduced. When fulfilling an order for the grinding wheels production, enterprises of the abrasive industry have a minimum of information on its operating conditions and are usually guided by recommendations for the manufacture of universal products. Attempts to adjust their production based on the results of bench tests that are made by manufacturers of grinding wheels are largely unsuccessful, because the discrepancy in the abrasive indices in the laboratory and production conditions is quite significant. The grinding wheel is a complex multi-parameter system that interacts with the processed workpiece and provides a specified amount of metal surface roughness and the required physical and mechanical properties: micro-hardness, wear resistance, and vibration resistance. In accordance with the existing concepts, the basis of the abrasive tool is the abrasive material and efficiency of processing depends on its interaction with the workpiece surface. Information about the principles of selection of the structural and mechanical characteristics of the grinding wheel—hardness, structure number, and type of binder—has an indefinite general nature. Depending on the requirements for the tool performance indexes, various structural and mechanical characteristics of the grinding wheel should be implemented. For profile grinding of parts for fuel, hydraulic, pneumatic equipment, grinding wheels with high hardness and density, with good edge resistance are necessary. Sharpening of the blade tool is performed by wheels with low hardness and density. Creep feed grinding with high speeds is performed using a highly porous and durable tool with low hardness. Finishing processes are performed with tools with high elasticity, strength, and wear resistance. In accordance with the existing concepts, the formation of the abrasive tool structure is made by the molding and heat treatment operations. When forming a grinding wheel blank, along with the shaping from the abrasive mixture, its components are distributed: binders and abrasive grains, i.e., fixing the structure. Coagulation reversible contacts are formed between the particles of the dispersed phase, which under the influence of the firing temperature are transformed into condensation (irreversible). As a result, the instrument acquires the final structural and mechanical parameters: density, strength, and porosity. Based on the analysis of the abrasive processing, it can be noted that the technological approach currently being implemented at the choice of rational characteristics of the grinding wheel is based on bench testing in a wide range of grinding conditions using samples produced on the basis of recommendations that do not take into account changes in machine tool building and abrasive industry.

Express Control of Abrasive Tool Operational Characteristics

997

In actual practice, during a specific grinding operation under set modes using the exhaustive method for tools with different characteristics, a grinding wheel is selected that provides the required productivity and qualities. It is obvious that this method is very expensive, and most importantly does not allow to respond to the requests of manufacturers of abrasive tools, who want to have information about the required characteristics of the grinding wheel, which must be made. That means that the information obtained during the monitoring of the actual process should allow you to assign the required tool characteristics to ensure the performance and quality of grinding.

2 Relevance The reduction of costs during the design of grinding wheels at manufacturers of abrasive tools, as well as at enterprises-consumers with an objective control of its effectiveness requires the creation of a method of express quality control of the tool.

3 Formulation of the Problem Within the framework of existing ideas, the wear of the grinding wheel is realized in the form of torn abrasive grains from the ligament, local and bulk destruction, and mechanical wear by abrasion (surface damage of the grain) [1–8]. The wear resistance of the grinding wheel depends on many factors: pressure on the actual contact area, friction coefficient, the environment, and physical, mechanical, and chemical properties of the tool and the workpiece. The abrasive materials’ wear resistance is determined by the micro-cutting method with a single grain under conditions close to its actual work in the grinding wheel: cutting speed and cutting depth, processed material. In order to implement this method, the abrasive grain is embedded in the holder, which is attached in the periphery of a metal disk that simulates a grinding wheel of a given diameter and is mounted on the faceplate of a circular grinding machine. Index lines are applied on a cylindrical sample at a given depth of cut with an abrasive grain. The process of micro-cutting is carried out until the complete cessation of micro-cutting. The path traveled by the abrasive wheel to complete wear is a criterion of its wear resistance. In the works [9, 10], it was shown that abrasive grains’ wear resistance during micro-cutting is consistent with the operational characteristics of grinding wheels. The described method for assessing abrasive grain wear resistance correlates with the finishing grinding process, because the conditions of interaction between “abrasive–metal” pair are identical.

998

V. M. Shumyacher et al.

In conditions of grinding with a large cross-feed tool, specified method is ineffective. A known method for assessing the operational properties of an abrasive grain for given conditions is based on the limiting thickness of a metal cut, at which its destruction occurs [11]. For all the results obtained, the abovementioned methods do not allow to reliably determine the relative wear resistance of an abrasive, which characterizes its resistance to mechanical abrasion without the influence of high temperatures, pressures, and physicochemical processes in the abrasive–metal contact. In the work [10], the possibility of determining the grinding materials wear resistance by abrading the counterface on the lapping tool with a suspension of abrasive powder of a certain grain was shown. The most important characteristic of the grinding tool, which determines its quality, is durability. According to current standards, the quality indicators of grinding materials are impact resistance of grinding grain, abrasive (cutting) ability, strength of single grains, as well as their micro-hardness, micro-fragility, and micro-strength. In the work [12], it was shown that of all the known mechanical properties of grinding materials—micro-brittleness, micro-hardness, and micro-strength—the latter is most closely associated with the wear of the grinding wheel. As the analysis of the results of works [13–22] in the group of corundum and carbides shows, there is a relationship between the micro-mechanical properties of abrasives and their operational properties. The higher the micro-strength of the abrasive grain is, the higher its wear resistance during micro-cutting is and the greater the specific productivity of the tool from this abrasive is. However, it should be noted that the above data are related to the wear resistance of abrasive grains in the tool and not to the durability of the tool itself. In response to this problem and also for operational quality control of an abrasive tool at the stages of its design, as well as production, the development of an appropriate methodology and means for its practical implementation is required.

4 Theoretical Part As a result of the analysis of the methods for express evaluation of abrasives wear resistance and correlation values with operational indicators, the possibility of practical application of the method of relative wear resistance implemented using the “Schlif-2” unit designed by Volzhsky Research Institute of Abrasives and Grinding has been established. According to the existing method, 22  22  5 mm samples from a grinding wheel were cut out from the cutting part of the grinding wheel with a diamond disk. A suspension (10  15 drops) consisting of a mixture of water and glycerin and boron carbide in a ratio of 1:1 with an average particle size of *160 lm was applied on the counterface from the modified cast iron. The choice of boron carbide powder was due to the need to fulfill the

Express Control of Abrasive Tool Operational Characteristics

999

requirement that the hardness of the abrading body should be 1.5–2 times higher than the hardness of the abradable. Micro-hardness of boron carbide is 39 GPa, and of corundum is 22 GPa. Since abrasive tool wear during grinding is implemented by abrading abrasive grains (finish grinding), and by tearing them out of the binder or chipping (rough grinding), it is not known a priori that the ratio of these processes is due to both factors. The cutting edges of boron carbide particles with a grain size of 160 lm are much larger, both local micro-volumes of the instrument grains, and smaller in most cases, the sizes of its grains, since for a significant number of grinding wheels, an abrasive grain (electrocorundum, silicon carbide) with an average size of more than 160 microns is used. The weight of the boron carbide particles established after numerous experiments is 0.04  0.012. The contact pressure is 0.01; the test time is 300 s [10, 21]. The wear of the samples was determined by the amount of the change in the height of the 22  22  5 mm bar before and after the experiment, measured using a microcator with an accuracy of ±0.5 lm. According to the data [10], there is a linear relationship between the sample wear and grinding wheel wear with a correlation coefficient of 0.86–0.92, which allows us to recommend this method for express evaluation of the tool quality of a made of abrasive materials for grinding operations. In the course of experimental studies, it was found that with a decrease in the samples’ hardness, the wear increases and practically does not depend on the grain of the abrasive. To clarify the reasons for this fact, studies have been conducted. The “Schlif-2” unit was upgraded by connecting the electric motor drive to the wattmeter, which allowed us to fix the energy consumption during wear of each sample during 300 s. Due to the fact that the strength of abrasive grains and abrasive powder sample (boron carbide) is greater than Al2O3 (corundum) and SiC (silicon carbide), and even more so of the binder, the most likely scenario for the process development should be considered as tearing out Al2O3 or SiC grains. The abrasive grain retention force in a binder depends on a number of factors: volume fraction in the grain circle, binder and pore, grain size, tensile and compressive strength of the binder material, the amount of binder bridges per abrasive particle, and its protrusion magnitude above the binder. According to the test results, the volume of the worn part of the sample and the amount of energy consumed during 300 s were recorded (E). By dividing the energy E by the volume of the worn layer of the sample, the value of the specific energy of destruction was determined. It has been established that with an increase in the volume of the binder, the specific energy of destruction of the sample increases (Fig. 1). As follows from Fig. 1, with an increase in the amount of binder from 20 to 35%, the specific fracture energy of the composite increases in *3 times. Analysis of the grain size of the sample wear products indicates a decrease in the grain content of the main fraction (250 lm) and an increase in the fine fraction. These data indicate a transition from wear of the sample due to the tearing of the abrasive grains to micro-cleaving.

1000

V. M. Shumyacher et al.

Fig. 1 Dependence of the specific crushing energy of abrasive sample on the quantity of binder (grain material Al2O3, main fraction size 250 microns)

Comparison of the specific wear energy of the samples of grinding wheels obtained by express analysis and the actual grinding process indicates a linear relationship between them with a correlation coefficient of *0.84.

5 Conclusions On the basis of the conducted research, a method of express control of the most important characteristics of the grinding wheel wear resistance has been developed. It has been established that there is a direct linear relationship between the wear of the grinding wheel in the actual process and the sample wear determined by the express method. The determination of the energy indicators of the process of wear by the express method can be useful in designing new tools for the given conditions of their operation, in particular, if it is necessary to reduce energy costs.

References 1. Bogomolov NI, Sayutin GI, Kharchenko IV (1970) Study of abrasive wear during microcutting. Phys Chem Mech Mater 6(2):27–31 2. Bogomolov NI, Sayutin GI, Kharchenko IV (1970) Investigation of abrasive materials wear under high contact pressure. Phys Chem Mech Mater 6(2):31–39 3. Bokuchava GV (1984) Tribology of grinding process. Sabchota Sakartvelo, Tbilisi, p 238 4. Korolev AV, Novoselov Yu K (1987) Probability—theoretical bases of abrasive processing. Part 1, Condition of the working surface of the tool. Publishing House of Saratov University, Saratov, p 160 5. Korchik SN (1974) Performance of the grinding steel parts process. Mechanical Engineering, Moscow, p 280

Express Control of Abrasive Tool Operational Characteristics

1001

6. Pushkarev OI, Schumyacher VM (2006) Micromechanical and operational characteristics of abrasive grains during grinding. Mech Eng 12:29–32 7. Schumyacher VM, Grishin IV, Kulikova LN et al (2006) Study of the mechanism of wear of abrasive material in the process of microcutting. Bull Saratov State Tech Univ 1(2):56–59 8. Schumyacher VM, Kadilnikov AV (2007) Model of the interaction of abrasive grain and the material being processed during grinding. Chip control circuit. Mech Eng Technol 4:18–21 9. Shumyacher VM, Pushkarev OI. Strength of the abrasive grain during grinding. Mech Eng Technol 7:26–27 10. Pushkarev OI, Shumyacher VM (2004) Methods and means of physical and mechanical characteristics of abrasive materials monitoring. Monograph, Volgograd State University of Architecture and Civil Engineering, p 144 11. Kharchenko IM, Sayutin GI, Bogomolov NI (1972) Method and device for determining the wear resistance of abrasive grains. Abrasives 4:14–19 12. Pushkarev OI, Shumyacher VM (2004) Prediction of working capacity of grinding materials during abrasive processing. In: Dynamics of technological systems, collection of scientific works from the international conference. Saratov State Technical University, Saratov, pp 308–309 13. Berdikov VF, Pushkarev OI, Leonidov LD (1981) Comparative evaluation of some physical and mechanical properties of grinding materials. Abrasives 8:11–14 14. Pushkarev OI, Shumyacher VM (2004) Quality control of grinding materials by impact resistance. Abrasive production, collection of scientific works. South Ural State University, Chelyabinsk, pp 135–142 15. Pushkarev OI, Shumyacher VM, Torshin DD (2005) Micromechanical and operational characteristics of abrasive grains at grinding. Abrasive production, collection of scientific works. South Ural State University, Chelyabinsk, pp 95–102 16. Kulik OG, Demenkov VA (2017) Kinematics and dynamics of chip formation during grinding. In: Procedia Engineering. 3rd International Conference on Industrial Engineering (ICIE-2017), vol 206, pp 210–215. https://doi.org/10.1016/j.proeng.2017.10.462 17. Shumyacher VM, Pushkarev OI, Slavin AV (2017) Energy efficient technology of obtaining advanced composite grinding materials and tools. In: “Silicon Carbide – Corundum” system from aluminium-containing residual products. In: Procedia Engineering. 3rd International Conference on Industrial Engineering (ICIE-2017), vol 206, pp 228–231. https://doi.org/10. 1016/j.proeng.2017.10.465 18. Shumyacher VM, Bochkarev PU, Slavin AV (2017) Monitoring operating parameters of abrasive materials. In: Procedia Engineering. 3rd International Conference on Industrial Engineering (ICIE-2017), vol 206, pp 232–235. https://doi.org/10.1016/j.proeng.2017.10.466 19. Shumyacher VM, Slavin AV, Kryukov SA (2016) The influence of ceramic binder grinding wheel structural and mechanical characteristics on its durability. In: Procedia Engineering. 2nd International Conference on Industrial Engineering, vol 150, pp 916–919. https://doi.org/ 10.1016/j.proeng.2016.07.057 20. Shumyacher VM, Slavin AV, Brzhozowski BM (2016) On the mechanism of interaction of abrasive tools and a workpiece at grinding. In: Procedia Engineering. 2nd International Conference on Industrial Engineering (ICIE-2016), vol 150, pp 920–923. https://doi.org/10. 1016/j.proeng.2016.07.059 21. Shumyacher VM, Slavin AV, Bochkarev PY (2019) Monitoring of grinding process with help of automatic measuring complex. In: Radionov A, Kravchenko O, Guzeev V, Rozhdestvenskiy Y (eds) Proceedings of the 4th International Conference on Industrial Engineering. ICIE 2018. Lecture Notes in Mechanical Engineering. Springer, Cham 22. Shumyacher VM, Slavin AV, Kulik OG (2019) Chip formation mechanism at metal grinding. In: Radionov A, Kravchenko O, Guzeev V, Rozhdestvenskiy Y (eds) Proceedings of the 4th International Conference on Industrial Engineering. ICIE 2018. Lecture Notes in Mechanical Engineering. Springer, Cham

Study of Physical and Chemical Processes Occurring During Polycondensation of Bakelite Binder in Order to Adjust Technological Process of Abrasive Tool Production I. Yu. Orlov, T. N. Orlova and I. V. Bashkirtseva

Abstract Abrasive–bakelite compositions after molding are subjected to heat treatment (bakelization) to cure the ligament. The process of binder curing at temperatures up to 190 °C is accompanied by the release of gaseous substances, mainly water vapor, which is contained in liquid bakelite (up to 27%) and ammonia, which is formed at 120 °C during the heat treatment of the instrument during the decomposition of hexamethylenetetramine (urotropine), which is contained in powder binder. Water vapor, withdrawing from the volume of the wheel during heat treatment under pressure, destroys the structure of the product, and increases the volume, which manifests itself in the form of bulges, shells, and bundles. This in turn leads to a decrease in the density, and physical, mechanical, and operational properties of the abrasive. A technological adjustment is needed to increase the mechanical strength of the instrument. The emission of volatiles in the bakelization process in the raw materials and molding sand mixture of face grinding wheels was determined. The processes occurring in certain temperature ranges of bakelization were also analyzed. Keywords Bakelite tool


Adhesion
Polycondensation
Volatile substances

1 Introduction During bakelization of a bakelite instrument with a phenol–aldehyde binder, which is a part of the molding sand mixture, physical and chemical processes take place, which were studied in the laboratory of the Research and Development Center “Volzhsky Research Institute of Abrasives and Grinding” (Volzhsky Polytechnic Institute). After analyzing the data obtained, the following conclusion was made I. Yu. Orlov (&)
T. N. Orlova
I. V. Bashkirtseva Volzhsky Polytechnic Institute (Branch) Volgograd State Technical University, 42a, Engels Street, Volzhsky, Volgograd Region 404121, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_106

1003

1004

I. Yu. Orlov et al.

about the need to remove volatiles even before the polycondensation reaction, i.e., avoid its critical points, namely to remove the maximum amount of volatile substances in the operation of the previous bakelization—mixing [1–8].

2 Relevance To obtain high performance of bakelite tools, the effect of temperature increase in the mixing process was studied, to remove volatile substances from the molding sand mixture already at 60 °C, thereby reducing their content in the wet wheel, which creates conditions for a fuller polycondensation reaction during bakelization.

3 Research Objective The emission of volatiles in the bakelization process in the raw materials and molding sand mixture of face grinding wheels was determined. The processes occurring in certain temperature ranges of bakelization were also analyzed (Table 1). Using these studies, it is possible to determine the qualitative and quantitative content of volatile substances during the cross-linking of the binder in different temperature ranges [1]. In the production of abrasive tools on a bakelite bond, a powdered binder with a different content of urotropine is used, and accordingly, the content of volatile substances will also be different [9–11]. When studying this process, in the laboratory of Research and Development Center “Volzhsky Research Institute of Abrasives and Grinding,” the amount of volatile substances was determined depending on the temperature, according to the heat treatment conditions (Table 2). From Table 2, it is seen that with an increase in the percentage of urotropine in the binder, the volatile yield increases, and up to a temperature of 100 °C, the volatile yield is approximately the same for all batches of Pulverbakelite, and at higher temperatures, the higher the percentage of urotropine in the binder. This is probably due to the fact that decomposition of hexamethylenetetramine begins at a temperature of 120 ° C, although pure urotropine completely decomposes at a temperature of 263 °C and a curing reaction takes place, and therefore, the volatile yield depends on the amount of hexamine in the binder. It can be seen from graph in Fig. 1 that at the temperature of 180 °C for Pulverbakelite with 12% content of urotropine, the volatile yield increases by 90% compared to the serial Pulverbakelite volatile yield. Therefore, it should be noted that when using a Pulverbakelite binder with a high urotropine content during heat treatment, an additional exposure of 2–3 h at a temperature of 120 °C is advisable for the best yield of volatile. The use of Pulverbakelite with 12% content of urotropine is undesirable [1, 2, 12–16].

Study of Physical and Chemical Processes Occurring During …

1005

Table 1 Processes occurring during the bakelite binder polycondensation Bakelization temperature ranges (°C)

Ongoing processes

60–80

Viscosity breaking of the humidifier Liquid bakelite-3, slow polycondensation of Liquid bakelite-3, and removal of 20% of volatile. The viscosity of the molding sand mixture is high enough; do not impose the special requirements for the process Further viscosity breaking of the humidifier and polycondensation of Liquid bakelite-3, removal at 1000 °C of 30–70% of all volatile. The viscosity of the molding sand mixture is low, which creates conditions for the deformation and the appearance of defects (pulled surface, cracks, melts, etc.). The rate of temperature increase is the lowest duration of the interval 25–31% of the entire conditions Viscosity breaking of Liquid bakelite-3, final polycondensation of Liquid bakelite-3; phenolic powdered binder has not yet reacted with urotropine, the amount of volatiles is insignificant; phenolic powdered binder has the lowest viscosity, melted. The rate of curing reaction of Liquid bakelite-3 is the highest. The rise in temperature is slowed down or stopped (an extract is given) The decomposition of hexamine occurs; the viscosity of phenolic powdered binder increases, the remaining volatile components stand out. It is recommended to reduce the rate of temperature elevation The curing reaction of phenolic powdered binder continues. The viscosity of the molding sand mixture continues to grow rapidly. Volatile emissions are reduced to a minimum, 3–5% of the total. In this interval, a faster rate of rise of temperature is possible than in the previous one The binder (phenolic powdered binder + Liquid bakelite-3) goes through the final stage of the polycondensation reaction and goes into a solid insoluble state

80–100

100–120

120–140

140–180

180–200

Table 2 Number of volatile in the Pulverbakelite binder at different temperatures Temperature (°C)

Amount of urotropine in the binder (%) 2 4 6 8

10

12

20 90 100 110 125 140 150 170 180

0 0.92 1.4 1.4 1.48 1.52 1.6 1.6 1.75

0 1.15 1.35 1.55 1.75 2.03 2.11 2.23 2.35

0 0.96 1.28 1.44 2.16 2.36 2.56 2.92 3.28

0 1.4 1.64 1.64 1.68 1.72 1.84 1.88 1.96

0 1.4 1.64 1.76 1.80 1.84 1.92 1.96 2.08

0 1.27 1.51 1.67 1.99 2.03 2.07 2.07 2.19

1006

I. Yu. Orlov et al.

Fig. 1 Dependence of the amount of volatiles on the percentage of hexamine in Pulverbakelite binder at a 180 °C temperature

The reduction in ammonia content was achieved by additional exposure during heat treatment. The work addressed the issue of how to ensure the removal of volatile substances from the molding sand mixture during the mixing process, since a significant role in shaping the quality of the finished product in the production of an abrasive tool belongs to the process of preparing an abrasive mixture—the mixing process. Blending is a fundamental operation of the manufacturing process of abrasive tools. In order to increase the temperature during the mixing process, a heating source was chosen—an electrohelix for a mixing machine with a volume of 50 L. This method of obtaining a free-flowing molding sand mixture is characterized by the fact that the mixture consists of individual granules, which are grains covered with liquid bakelite with partially dissolved phenolic powdered binder and various fillers. Such a mixture has good flowability and does not crack, and a uniform distribution of components by volume occurs. The molding sand mixture is prepared in the technological laboratory. 5 kg of the molding sand mixture is prepared by cold mixing, and 5 kg of the molding sand mixture is prepared by “warm” mixing according to the developed recipe.

4 Experiment According to this recipe, two types of molding sand mixtures were prepared. The experimental procedure consisted of two stages.

4.1

Determination of Heat Losses During Molding Sand Mixture Heating

Polycondensation reactions are characterized by the release of a by-product, water, in this case. The binder is the part of the mixture that is contained in the final

Study of Physical and Chemical Processes Occurring During …

1007

Table 3 Number of volatile losses in the Pulverbakelite binder at different temperatures temperature conditions of mixing Mixing time (minutes)

1

2

3

4

5

6

7

8

Temperature (°C) Volatile loss (%)

20 0

25 0.3

30 0.6

35 0.8

40 1.8

45 1.9

50 2.0

60 2.0

product. In the case where bakelite binder wheels are used, the binder consists of two or more components [17–19]. They are the following components: • dry resin—2.5 kg • liquid resin—1.1 kg • inorganic filler—0.4 kg. The determination of the amount of volatile components in the mixture at different temperatures was carried out using the unit that consists of a counterbalance, a drying cabinet, and a thermometer. The determination was made as follows: The sample of the studied bakelite mixture (100 g) was evenly distributed over the foil and placed on a horizontal platform of counterbalance, installed in a drying cabinet. The balance of the balance was established, and the temperature was raised according to the temperature conditions. The number of volatile losses in the Pulverbakelite binder at different temperatures was reflected in Table 3.

4.2

Determination of Mechanical Strength (Breaking Stress Under Tension)

Tensile strength is a very structure-sensitive property and depends largely on the technology of the material [20, 21]. 8-shaped prototypes are made by the formation of abrasive molding material and subsequent bakelization. For the test, three samples of the 8-shapes prepared from the molding sand mixture by cold mixing and three samples of the 8-shapes prepared from the molding sand mixture by warm mixing are taken. Further tests of bakelite samples are produced on a UMM-5 universal machine. The prototype is placed in the grippers, one of which, being active, moves in the direction opposite to the stationary gripper and, turning on the machine, smoothly increases the load applied to the sample, up to its destruction. The test results are listed in Table 4. Table 4 Value of the indicator “Tensile stress at break”

Mixing method

r1 (MPa)

r2 (MPa)

r3 (MPa)

ravg (MPa)

Cold Warm

120 140

140 160

130 150

130 150

1008

I. Yu. Orlov et al.

5 Conclusions Based on the conducted experimental studies, it was found that: • The reduction of ammonia content was achieved by additional curing during heat treatment. • The content of volatile substances in the molding sand mixture prepared by “warm” mixing is 0.25 times less than that in the molding sand mixture prepared by cold mixing, which positively affects the quality of the molding sand mixture; it has good flowability, does not crack, and provides high cutting performance laps. • In prototypes of 8-shapes prepared by “warm” mixing, mechanical strength is 1.25 times greater than in prototypes of 8-shapes prepared by cold mixing. • We propose to use the technology of warm mixing, which allows to increase the adhesion between the abrasive grain and the binder, for the production of an abrasive tool on a bakelite binder.

References 1. Orlov IYu (2015) Improving the efficiency of the operation of grinding steel with a bakelite tool by selecting fillers that reduce the thermal stress of the process (article). Reference Book. Eng J Appl 6:219 2. Orlov IYu (2007) The impact of technological methods on operational performance. Materials science and technology of construction materials are the most important components of the competence of a modern engineer. Problems of quality technological training. In: Collected works of All-Russian meetings of heads of materials science and technology of construction materials departments. Volzhsky Institute of Construction and Technology, Branch of Volgograd State University of Architecture and Civil Engineering - Volgograd, Volzhsky 3. Orlov IYu, Orlova TN (2007) The method of calculating the specific amount of volatile substances during the heat treatment of an abrasive bakelite tool materials science and technology of structural materials are the most important components of the competence of a modern engineer. Problems of quality technological training. In: Collected works of All-Russian meetings of heads of materials science and technology of construction materials departments. Volzhsky Institute of Construction and Technology, Branch of Volgograd State University of Architecture and Civil Engineering - Volgograd, Volzhsky 4. Orlov IYu, Orlova TN (2003) The study of the processes occurring during the polycondensation reaction of phenol-formaldehyde resins. Theory, technology and equipment for the production of abrasive tools. In: Collected research works. South Ural State University Publishing, Chelyabinsk 5. Pushkarskaya OYu, Orlov IYu (2004) Research and development of technology for processing (using) slimes of bearing plants for use in the production of abrasive tools. Abrasive production. In: Collected research works. South Ural State University Publishing, Chelyabinsk

Study of Physical and Chemical Processes Occurring During …

1009

6. Orlov IYu, Shumyacher VM (2008) Study of the thermal conductivity of prototypes for the fillers choice, increasing the strength characteristics of face grinding wheels (with the expansion of environmental problems). Prospective directions for the development of engineering technology and metalworking. In: Materials of the international scientific and technical conference. Don State Technical University, Rostov on Don, September 29–October 3 7. Orlov IYu (2010) Physical and mathematical model of the process of face grinding of rails with abrasive wheels, science and education: problems, solutions and innovations. In: Scientific and practical conference of higher-education teaching personnel of Volzhsky Institute of Construction and Technology, Volzhsky, December 9–10. 2010. In: Collected works: in 2 parts—Volgograd: Volgograd State University of Architecture and Civil Engineering Publishing, Part 1 8. Kurdyukov VI (2000) Scientific basis for the design, manufacture and operation of abrasive tools. Dissertation of Doctor of Engineering Science, Kurgan 9. Orlov IYu (2011) Improving the cutting properties of abrasive tools for grinding of railroad rails (article) abrasive machining processes, abrasive tools and materials. In: Shumacher VM (ed) Collected works of the international scientific and technical conference “Shlifabraiv-2011”. Volgograd State University of Architecture and Civil Engineering, Volgograd 10. Ostrovsky VI (1981) Theoretical foundations of the grinding process. LSU Publishing, Leningrad 11. Nosenko VA, Orlov N, Shegay AA (2006) To the question of the structure of the abrasive tool. Inzhenerny Zhurnal Handbook 12. Baidakova NV, Kryukov SA (2014) On the influence of the shape of the abrasive grain on the cutting ability of the tool during roughing operations. In: New materials and technologies: state of the issue and development prospects a collection of materials of the All-Russian youth scientific conference 2014 13. Orlova TN (2003) Investigation of processes occurring during the phenol-formaldehyde resin olycondensation reaction theory and technology and equipment for the production of abrasive tools. In: Chaplygin BA et al (eds) Collection of scientific works. South Ural State University Publishing House, Chelyabinsk 14. Orlova TN, Orlov IYu (2004) Study of the influence of the two-stage mixing of the components of the molding sand mixture on the manufacturability of the mixture (obtaining free-flowing molding sand mixtures) and the physicomechanical properties of the abrasive tool on the bakelite bond. Abrasive processes, abrasive tools and materials. In: Collected works of the international scientific and technical conference “Shlifabraiv-2004”, Volzhsky Institute of Construction and Technology, Branch of Volgograd State University of Architecture and Civil Engineering, Volzhsky 15. Kryukov SA, Slavin AV, Baidakova NV (2013) Prediction of mechanical strength of abrasive composite materials. Mechanical Engineering Technology, 10 16. Bagayskov YuS (2003) Forming a uniform material structure abrasive tools. In: Collected conference works abrasive machining processes, abrasive tools and materials, Volzhsky Civil Engineering Institute, Volgograd, Volzhsky, September, pp 8–14 17. Trofimova TV, Shumyacher VM (2001) Technological principles of creating an abrasive tool on a bakelite binder. In: Collected materials of the All-Russian scientific and technical conference: materials and technologies of the XXI century, Penza, part, vol 3, pp 84–87 18. Trofimova TV, Shumyacher VM (2003) Operational indicators of abrasive tools on bakelite binder. Abrasive processes, abrasive tools and materials. In: Collected conference works, Volzhsky Civil Engineering Institute, Volgograd, Volzhsky, September, pp 8–14

1010

I. Yu. Orlov et al.

19. Orlova TN, Pushkarskaya OYu (2002) Study of the organic bonded abrasive tool composition. In: Collected conference works abrasive machining processes, abrasive tools and materials, Volzhsky Civil Engineering Institute, Volgograd, Volzhsky, 11–17 Sept 2002 20. Orlova TN, Orlov IYu, Pushkarskaya OYu (2004) Research and development of research and development of technology for processing (using) slimes of bearing plants for use in the production of abrasive tools. In: Collected research works: abrasive tools and metalworking, UNIIASh OJSC, Chelyabinsk, 20041 21. Orlova TN, Orlov IY, Hvan NS (2019) Physical and mathematical model of the process of face grinding of rails with abrasive wheels operating at speed of 50 m/s. In: Radionov A, Kravchenko O, Guzeev V, Rozhdestvenskiy Y (eds) Proceedings of the 4th International Conference on Industrial Engineering. ICIE 2018. Lecture Notes in Mechanical Engineering. Springer, Cham. https://doi.org/10.1007/978-3-319-95630-5_152

Optimal Sequencing at Selection of Abrasive Tools’ Characteristics for Their Improvement I. Yu. Orlov, N. V. Baidakova and P. Yu. Bochkarev

Abstract The problems of optimal sequencing at the selection of abrasive tools’ characteristics for the improvement of their quality characteristics have been considered. It has been established that the grain composition of the abrasive material affects the metal removal rate to the greatest extent, the second most important characteristic is hardness, the third is porosity, and the fourth is the structure of the tool. The obtained results serve as a basis to recommend the developed methodology for researching and evaluating ways and options for improving abrasive tools for practical use. As a result of scientific research, the optimal sequence of improving the basic characteristics of abrasive tools has been determined. Keywords Abrasive tools composition Hardness





Improvement
Optimum characteristics
Grain

1 Introduction Today, the improvement of abrasive tools is carried out in several areas, due to the increasing demands of consumers and the conditions and technologies of their production, in order to obtain tools for various purposes and improving their quality and efficiency.

I. Yu. Orlov (&) Volzhsky Polytechnic Institute (Branch) Volgograd State Technical University, 42a, Engels Street, Volzhsky, Volgograd Region 404121, Russia e-mail: [email protected] N. V. Baidakova “Moscow Power Engineering Institute” in the City of Valzhsky, Branch of National Research University, 69, Lenin Avenue, Volzhsky, Volgograd Region 404110, Russia P. Yu. Bochkarev Yuri Gagarin State Technical University of Saratov, 77, Politechnicheskaya Street, Saratov 410054, Russia © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_107

1011

1012

I. Yu. Orlov et al.

The most common ways to improve the tools are to develop the ways to control their structural, mechanical properties, and operational characteristics due to the rational choice of physical, mechanical, and structural characteristics determined by the ratio of grain, binder and pores; to develop new types of tools such as highly porous, composite for high-speed grinding, etc.; to use heat treatment and impregnation with various compositions; variation recipes, etc. Despite the large amount of theoretical and experimental works on the improvement of abrasive tools, the management of their characteristics and performance has not been studied sufficiently. The problem of the mechanism of cracking in the tools shard based on the analysis of internal stress fields requires solving. Improving the efficiency of the wheels, their heat treatment and impregnation are implemented empirically in the absence of reasonable temperature-time conditions, as well as technological principles and methods for their regulation. In addition, the principles of changing the abrasive tool characteristics and parameters under the influence of the external environment are not well understood, which significantly affects the grinding performance, especially during computer-integrated manufacturing. One of the most promising ways to improve the efficiency of tools and the grinding quality is to justify the choice of the grain composition of abrasive materials. Nowadays, however, this issue is also given insufficient attention. Experimental studies, as a rule, don’t take the presence of small fractions of the tested instruments into account, which can lead to contradictory conclusions of the study’s results. The content of individual fractions of the grain composition and their ratio among themselves has a direct impact on the structural and mechanical characteristics of the tool and on the indexes of the grinding process, determining the removal of metal, quality of the surface to be machined, wear, and tool life. Meanwhile, when choosing the characteristics of a tool for a given grinding operation, these decisive factors for the machining process are not fully taken into account. In the scientific literature on the issues of grinding, the influence of the grinding materials grain composition on the characteristics and performance of tools, despite its top priority, has not been adequately evaluated. Experience in abrasive tools operating shows that in order to improve the quality and effectiveness of the grinding process, it is often necessary to perfect them by targeted modernization or modification. Getting to this point, it is necessary, first of all, to solve the question of which tool characteristics need improvement and which operational factors should be improved and how these changes will generally affect the quality and performance of grinding.

Optimal Sequencing at Selection of Abrasive Tools …

1013

2 Relevance The issues of abrasive tools improvement are highly relevant, allowing us to achieve the necessary technical level of grinding processes and, as a consequence, the higher quality of finished products. However, the solution to such problems is often unsystematic. Optimizing the sequence in choosing the characteristics of abrasive tools aims to solve this problem.

3 Research Objective The choice and optimization of structural and technological ways of regulating the tools’ characteristics can be performed based on the method of obtaining an overall assessment of the technical level indicators of the product that is upgraded by combining individual indicators into generalized ones using weight coefficient [1–20]. For these purposes, it is necessary to investigate the function of many variables in order to determine such a characteristic, the indicator of which gives the greatest increase in the function ∂P when it changes to DP.

4 Theoretical Part This problem belongs to the class of problems of value engineering by the method of gradient ascent. The following formula is used:
 @P0 @P0 @P0 gradðP0 Þ ¼ ; ; . . .; ; @x1 @x2 @xn

ð1Þ

where P0—is a composite index; x1, x2,…,xn—integral characteristics of the object. Then, to determine the amount of change in DP, a matrix (2) is composed (2):   2 2  @ P0 P0    @x2 ; . . .; @x@1
@x n   1 ð2Þ  2 . . . 2 ;  @ P0 @ P0  ; . . .;  @xn @x2  n

the norm of which is expressed by the following equality: vffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi u
 uX @ 2 P 0 2 t N¼ : @xi @xj i;j

ð3Þ

1014

I. Yu. Orlov et al.

In this case, DP value providing the search for the P0 maximum is selected on the basis of the provision of the following condition: DP 

1 : N

ð4Þ

In this case, it is necessary to study the dynamics of changes in the composite index and the sequence of individual properties (characteristics) that have the greatest influence on the change in P0 when searching the gradient at each point of the function f (x1, x2, …, xn). Studying the gradient of the composite index function, one can simulate the optimal way to improve the tool and the sequence of structural and technological regulation and stabilization of its characteristics [1, 19–22]. The optimization of the structural and technological regulation of the abrasive tool characteristics by the proposed method is carried out as follows. Initially, the significance and priority of the tool characteristics, their interval and variation limit, as well as the formula for the function f (x1, x2, …, xn) are established at x0j  xj  xjmax. Then the composite index P0 is selected, and its required value PTP 0 is set. Further, the values of the function are calculated at the points with P0(x0,1, …, x0,n) coordinates, and its new values are determined provided that each of the arguments is given a DP increment. As a result, the function increment DP0,i is discovered when each index changes by DP value according to the following expression:    DP0;1 ¼ P0;1  P0      ...     ...    DP1;N ¼ P0;N  P0 : 

ð5Þ

After that, the functions deduced are ranked in accordance with the DP0, j. increment. For all j = 1, …, N, the calculations are repeated until P0j  PTP 0

ð6Þ

At the end of the calculations, a graph of the choice of optimized tool characteristics is plotted, while the values of P0 are plotted on the Y-axis, and the values of the tool characteristics are plotted on the X-axis. To solve this problem, ceramic-bonded aluminum oxide abrasive tools were taken as objects of study. Table 1 presents the characteristics of the tools and the limits of their changes according to expert evaluations. The grinding performance Qм was taken as a composite index. Processing modes: wheel speed Vd = 35 m/s, longitudinal feed S = 20 m/min, transverse feed St = 1.5 mm/ stroke, and grinding depth t = 0.01 mm. Taking into account the data presented in Table 1, all the necessary calculations were carried out at the choice of optimized instrument characteristics. The results of this calculation are summarized in Table 2, according to which the graph is plotted as shown in Fig. 1.

Optimal Sequencing at Selection of Abrasive Tools …

1015

Table 1 Characteristics of abrasive tools and a composite index Characteristics and their change limits

1

2

3

4

Grain composition of abrasive tools with Z = 16  40

Structure C = 4, …, 9

Hardness T = M2, …, CM2

Porosity П = 30, …, 60%

Composite index (metal removal, cm3/min)

0.03–0.08

0.07–0.10

0.06–0.09

0.04–0.08

Table 2 Calculated data for selecting optimized tool characteristics Modified argument number U

1

2

3

4

5

6

7

8

9

10

Composite index P0 Rank of function increments j

0.079

0.081

0.105

0.110

0.116

0.124

0.127

0.139

0.142

0.147

I

I

I

III

III

III

IV

II

II

II

Fig. 1 Graph of the selection of the sequence of tool characteristics improvement

1016

I. Yu. Orlov et al.

5 Experimental Findings On the graph, each vertical section of the curve shows an increase in grinding performance due to a change in the corresponding tool characteristic, and arrows show the transition of one optimized characteristic to another. As can be seen from the graph, the grain composition of the abrasive material x1 affects the metal removal to the greatest extent. The second most important characteristic is the tool hardness x3, the third is the porosity x4, and the fourth is the structure x2. In this sequence, it is necessary to improve the abrasive tool by target-specific adjusting their structural and mechanical characteristics [18–22].

6 Conclusions To summarize, the results above indicate a great potential for applying a certain sequence of improving the tools’ characteristics and indicators in order to increase their efficiency, reliability, and quality. The obtained results serve as a basis to recommend the developed methodology for researching and evaluating ways and options for improving abrasive tools for practical use. As a result of scientific research, optimal sequence of improving the basic characteristics of abrasive tools has been determined. It has been established that the grain composition of the abrasive material affects the removal rate of the metal to the greatest extent, the second most important characteristic is hardness, the third is porosity, and the fourth is the structure of the tool.

References 1. Ananyan RV (1978) Study of the operational properties of grinding wheels made of aluminum oxide materials with a given porosity. Dissertation of Candidate of Engineering Sciences, Moscow, 196p 2. Baidakova NV, Kryukov SA (2012) Influence of grit and shape of abrasive grain on grinding efficiency. In: Collected works from international scientific and technical conference “Shlifabraiv-2011”. Abrasive processes, abrasive tools and materials, Volgograd State University of Architecture and Civil Engineering, Volgograd, pp 96–100 3. Bochkov AP, Gasyuk DP, Filyustin AE (2003) Models and methods of managing the development of technical systems, Soyuz Publishing House, St. Petersburg, p 288 4. Dvoryankin AM, Polovinkin AI, Sobolev AN (1997) Methods of synthesis of technical solutions. Nauka, Moscow, p 104 5. Ivashinnikov VT (1967) The cutting properties of grinding wheels of different grain content with constant hardness. Abrasives 2:28–31 6. Yulikova YuF et al (1981) Manufacturing of grinding wheels of increased hardness and strength by using grinding materials of various grit. Abrasives: express information, 1 edn. Moscow, pp 6–12

Optimal Sequencing at Selection of Abrasive Tools …

1017

7. Karaulov EV, Kryukov SA (2004) The influence of the structure of the abrasive tool on the technological parameters of the grinding process. In: Collected works from international scientific and technical conference “Shlifabraiv-2004”. Abrasive processes, abrasive tools and materials. Volgograd State University of Architecture and Civil Engineering Volzhsky Polytechnic Institute (branch), Volgograd, pp 73–74 8. Karachentseva TG, Yulikova YuF, Zaitsev GP. Methods of design, manufacture and testing of highly porous abrasive tools. In: Works of VNIIASh, vol 6, pp 7–21 9. Kovalchenko MS (1993) Mechanical properties of isotropic porous materials. Powder Metall 3:89–96 10. Kryukov SA (2000) Analysis and development of ways to control the operational properties of an abrasive tool. In: Inter-University scientific and practical conference of young scientists and students V Part 3, VolSU Publishing House, Volgograd, pp 104–105 11. Kryukov SA, Anokhin VI, Shevchuk VP (1998) Improving the abrasive tool in the production of bearings. In: Advanced technologies in mechanical engineering: Inter-University collected research works, VolgGTU, Volgograd, vol 1, pp 23–25 12. Kryukov SA, Guselnikov DA (2004) Modeling and experimental study of the coordination of grains in abrasive tools of various structures. In: Modern trends in the development of transport engineering and materials: collected works from IX international scientific and technical conference, Penza, pp 113–116 13. Kryukov SA, Guselnikov DA (2004) Management of structural and mechanical properties of abrasive tools. In: Abrasive machining processes, abrasive tools, materials: collected works from international scientific and technical conference, Volzhsky, pp 71–73 14. Kryukov SA (2005) Dependence of the strength of abrasive tools on their structural characteristics. In: Abrasive machining processes, abrasive tools and materials. In: Collected works from international scientific and technical conference Shlifabraziv-2005, Volzhsky, Volgograd, p 216 15. Kryukov SA, Baidakova NV, Nazarenko VA, Baidakova TA (2009) Improved productivity and product quality when using controlled—grain grinding tools. Russ Eng Res 5:509–511 16. Krukov SA (2015) Grinding tools based on multicomponent abrasive. Russ Eng Res 35 (1):19–21 17. Shumyacher VM, Slavin AV, Krukov SA (2016) The influence of ceramic binder grinding wheel structural and mechanical characteristics on its durability. Proc Eng ICEE 150:916–919 18. Shumyacher VM, Slavin AV, Krukov SA (2016) Phenomenological model of the mixing process of competents of a moldable mixture for manufacturing of abrasive tools. Proc Eng ICEE 150:911–915 19. Krukov SA, Krukova AS (2017) Determining the parameters of grinding wheels working surface Profile. Proc Eng ICIE 206:204–209 20. Krukov SA, Tkach MA (2017) Phenomenological model of abrasive tool components mixing process. Proc Eng ICIE 206:200–203 21. Kryukov SA, Baidakova NV, Bochkarev PY (2019) State of problem of technological support of workpiece surface quality during grinding. In: Radionov A, Kravchenko O, Guzeev V, Rozhdestvenskiy Y (eds) Proceedings of the 4th international conference on industrial engineering. ICIE 2018. Lecture Notes in Mechanical Engineering. Springer, Cham 22. Kryukov SA, Baidakova NV (2019) Dependence of roughness parameters of workpiece surface on structural characteristics of grinding discs. In: Radionov A, Kravchenko O, Guzeev V, Rozhdestvenskiy Y (eds) Proceedings of the 4th international conference on industrial engineering. ICIE 2018. Lecture Notes in Mechanical Engineering. Springer, Cham

Change in Microhardness of Metal Depending on Wetting Ability of Lubricating Coolant I. V. Bashkirtseva and T. N. Orlova

Abstract On the basis of the theoretical and experimental studies, a new technological direction for the rational use of the composition of lubricating coolant with the contact interaction of the lubricating coolant with the metal, including the determination of the physicochemical parameters of the lubricating coolant and the evaluation of their quality by performance, has been proposed. The prerequisites for the active interaction of the medium and the metal as the abrasive grain moves through it arise under conditions of significant normal and shear stresses. At the same time, in the deformed metal volumes, dislocation friction emerges on the surface, cracks form and microdefects develop due to surface activation. The decrease in the shear resistance of the boundary lubricant film in the “grain-tometal” contact, an improvement in its mobility and an increase in the bearing capacity are the reasons that an increase in the abrasive sliding speed on the metal when working with hydrocarbon and emulsion lubricating coolant leads to a decrease in the friction coefficient. Keywords Lubricating coolant


Abrasive treatment
Microhardness of metals

1 Introduction In the contacting surfaces of solid bodies with external friction, energy dissipation occurs. It makes sense to consider the process of friction of two bodies as a surface activation of the contacting bodies. In this case, a change in the physical and mechanical characteristics of friction bodies is caused by the influx of energy into the surface layers. The nature of the reactions in the near-surface layers and the dynamics of the process are influenced by the presence of a lubricant in the contact [1–6]. I. V. Bashkirtseva (&)
T. N. Orlova Volzhsky Polytechnic Institute (Branch) Volgograd State Technical University, 42a Engels Street, Volzhsky, Volgograd Region 404121, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_108

1019

1020

I. V. Bashkirtseva and T. N. Orlova

2 Theoretical Part For the frictional boundary layer in the molecular and kinetic friction theory, the concept of a “third body” which is formed by friction of two bodies at the point of contact is introduced. In this case, the physical and mechanical properties of the wearing surfaces under the influence of the lubricating coolant determine the nature of the two-body contact [4, 5, 7, 8]. The prerequisites for the active interaction of the medium and the metal as the abrasive grain moves through it arise under conditions of significant normal and shear stresses. At the same time, in the deformed metal volumes, dislocation friction emerges on the surface, cracks form and microdefects develop due to surface activation [9–11].

3 Practical Part The determination of the microhardness changes during the interaction of the metal with the lubricating coolant (Fig. 1) was performed by indentation of the diamond pyramid into the metal. The decrease in microhardness when using hydrocarbon

Fig. 1 Influence of the composition of the lubricating coolant on the microhardness of steel 40X during their contacting: H0—the initial microhardness (1.0); H—microhardness, determined in the presence of lubricating coolant

Change in Microhardness of Metal Depending on Wetting Ability …

1021

Fig. 2 Diagram of the contact of the diamond pyramid with the metal surface during microindentation

lubricating coolant, under experimental conditions, was 57%. When contacted with the surface of the metal, the emulsion lowered the metal microhardness by 3  5%. An aqueous solution of 0.3% sodium carbonate (Na2CO3) and synthetic liquid Aquol-10 had practically no effect on the microhardness of the metal. Let us consider the physical meaning of microhardness to explain the reasons. The effective force during microindentation is decomposed into two components (Fig. 2). The resistance of plastic deformation of metal is characterized by the normal component of the N force. The friction force arising on the verge of the pyramid during its sliding is characterized by the component T. When certain conditions are created, an effect of adsorption decrease in the strength of a solid body is observed [6, 9, 12]. The nature and intensity of the action of stresses in a deformed volume become important. Under conditions of compression of the metal adjacent to the face of the diamond pyramid, the action takes place in the contact surface layer, since the molecules of surface-active substances (surfactants) from the medium cannot penetrate into its volume [13]. Microindentation is also accompanied by the movement of layers of material adjacent to the edges of the pyramid, which creates conditions for manifestation of the adsorption effect of reducing the strength of a solid body due to the development of significant shear stresses in the surface layer. The friction force decreases due to a shear stress drop caused by the ease of movement of dislocations in the metal layer adjacent to the pyramid because of the high penetrating and wetting ability of hydrocarbon lubricating coolant, OSM3 and kerosene. In this case, surfactant molecules penetrate into the “face-metal” contact zone and further into the microdefects of the metal layer deformed [14, 15]. A solution of sodium carbonate in water and Aquol-10 synthetic liquid has low wetting ability, as a result of which their penetration into the pre-fracture zone of the metal is insignificant. Emulsions have slightly better penetrating and wetting properties as compared with a solution of 0.3% Na2CO3 in water and Aquol-10 liquid, they are more capable of moving in the face-metal contact, and therefore, they have a greater effect on the microhardness of the metal [12, 15, 16].

1022

I. V. Bashkirtseva and T. N. Orlova

Due to the strength reduction through adsorption and facilitating the shear of its surface layer during microindentation, there is an increase in the depth of the pyramid penetration into the metal (reduction in microhardness) and, as a result, a decrease in the friction force of the “face—metal” sliding. The result obtained coincides with the conclusions of the work of academician V. D. Kuznetsova that the quality of the lubricant on the metal surface affects the depth of the indenter penetration [17]. The use of Aquol-10 liquid as a lubricant with an increase in the level of energetic activation of the friction surface (an increase in the sliding speed of the abrasive) leads to instability of the friction force, and in the medium of hydrocarbon lubricating coolants and emulsions, the process of abrasive friction on the metal is stable (Fig. 3). The use of hydrocarbon lubricating coolants and emulsions does not cause an increase in the friction coefficient with increasing slip speed due to the fact that the layer of surfactant molecules from hydrocarbon lubricating coolants and emulsions adsorbed by the metal surface has significant elasticity and strength, as well as mobility and thickness, and under boundary friction of abrasive grains on metal in the lubricating coolant medium along with the process of adsorption lowering the strength of its surface layer on the dynamics of contact interaction influence the rheological properties of the lubricant layer [16–18]. When using hydrocarbon coolants and emulsions, the sub-microrelief of friction surfaces of abrasive grains on metal is characterized by homogeneity (Fig. 4a, c), while there are defects which are characteristic for the processes of adhesive setting (pulling, sagging) on the friction surface when using aqueous cutting fluids (Fig. 4d, e) [19, 20].

Fig. 3 Influence of the lubricating coolant composition on stability of the friction force of abrasive grain on metal (I′ = 0.15 m/s, P = 40H): lubricating coolant: □—distilled water; ●— OSM3; x—kerosene 85%, industrial oil 20-15%; ж—B25; ■—40К5U 5% concentration; ○— Aquol-10, 10% concentration; ▲—IHP 45E, 5% concentration; ♦—IHP 175

Change in Microhardness of Metal Depending on Wetting Ability …

1023

Fig. 4 Sub-microrelief of friction surfaces obtained by sliding abrasive grain Al2O3 on steel 40X in different lubricating coolant (1500): a kerosene–oil mixture, b OSM3, c NSC-5U, d B25, e Aquol-10

The result of a decrease in the wetting properties of lubricating coolants is an increase in the coefficient of friction of the abrasive on the metal (Fig. 5) [21]. The dependences of the friction coefficient on the metal depending on the composition of the lubricating coolant, cutting speed Vp and the normal component of the force N are presented in Table 1. An increase in the normal pressure of an abrasive grain (indenter) leads to an increase in the friction coefficient for all lubricating coolants. The growth of the molecular component of the friction force occurs with an increase in the actual pressure in the contact in accordance with the Binominal Molecular Friction Law of Deryagin [3, 22, 23]. This is especially clearly seen in water lubricating coolants V25, Aquol-10, 0.3% Na2CO3 in water. The mechanical properties of frictional boundary layers cause a difference in the effect of the normal pressure of the grain on the metal on the coefficient of friction in the environment of different lubricants. When using OSM3, IHP-45E, NSK5U lubricating coolants, the lower sensitivity of the friction coefficient of the abrasive on metal is due to the fact that hydrocarbon lubricating layers have a high compression modulus [1, 24, 25].

1024

I. V. Bashkirtseva and T. N. Orlova

Fig. 5 Influence of the wetting angle H lubricating coolant on the friction coefficient of abrasive grain Al2O3 on steel 40X: coolant: ж—distilled water, x—OSM3, ●—kerosene-85%, industrial oil 20-15%; ▲—B25, ○—40К5U, 5% concentration, ♦—IHP175

Table 1 Functional relations of the friction coefficient Al2O3 on 40X steel example of a table Lubricating coolant

Regression equation

Confidence span

Kerosene 85%, industrial oil 20  15% OSM3 NSK-5U IHP-45E Aquol-10 B25 0.3% Na2CO3, water—99.7%

ƒ = 0.054 − 0.036Iʹp + 0.011 N

±0.003

ƒ ƒ ƒ ƒ ƒ ƒ

±0.005 ±0.007 ±0.006 ±0.004 ±0.008 ±0.01

= = = = = =

0.056 – 0.032Iʹp + 0.008 N 0.078 − 0.061Iʹp + 0.04 N 0.1 − 0.087Iʹp + 0.095 N 0.12 − 0.1Iʹp + 0.12 N 0.137 − 0.12Iʹp + 0.1 N 0.15 − 0.127Iʹp + 0.13 N

4 Results of the Study Obviously, a decrease in the shear resistance of the boundary lubricant film in the “grain-to-metal” contact, an improvement in its mobility and an increase in the bearing capacity are the reasons that an increase in the abrasive sliding speed on the metal when working with hydrocarbon and emulsion lubricating coolant leads to a decrease in the friction coefficient (Table 1) [2, 25].

Change in Microhardness of Metal Depending on Wetting Ability …

1025

5 Conclusion At contact interaction of a cutting fluid with a metal, a change in its microhardness occurs. The higher the wetting properties of the lubricating coolant and its surface activity are, the more effective the reduction of the microhardness of metals is; reduction of friction, that is, shear resistance of metal layers adjacent to the indenter occurs due to the appearance of the effect of adsorption lowering of its strength, which underlies the mechanism of lowering the microhardness of the metal when it contacts the lubricating coolant; the use of aqueous liquids is characterized by instability of the forces that arise, and when abrasive is rubbed on a metal in a medium of hydrocarbon and emulsion liquids, the opposite is true; the increase in sensitivity of the friction coefficient of the abrasive over the metal in the medium of aqueous liquids compared with hydrocarbon is due to differences in the mechanical properties of boundary lubricant films; a decrease in the friction coefficient of an abrasive grain over metal with an increase in the sliding speed is due to an increase in the mobility of the film of adsorbed hydrocarbon molecules of surface-active substances and a decrease in the resistance to their shift; the main disadvantage of hydrocarbon cutting fluids in comparison with aqueous fluids is their non-ecological compatibility and high fire hazard.

References 1. Berdichevsky EG (1984) Lubricating and cooling technological means for processing materials: a handbook. Mechanical Engineering, Moscow, p 224 2. Berzin VR (1986) Application of lubricating coolants during internal grinding of steel alloy blanks. Publishing House of Saratov State University, Saratov, pp 32–35 3. Bilik ShM (1960) Abrasive water metal processing. Mashgiz, Moscow, p 198 4. Volkov MP (1977) Study of the influence of the physicochemical properties of lubricating coolants on the quality and dynamics of the grinding process. Abrasives 4:1–3 5. Dusko OV Shumiacher VM, Bashkirzeva IV (2005) Micromechanics of contact interaction of an abrasive with a material in the presence of lubricating coolant. In: Collected articles of the international scientific—technical conference “Shlifabraiv - 2005”, Volzhsky, pp 198–203 6. Eshchenko EP (2003) Effective implementation of lubricating coolants in metalworking manufacturing. World Eng Technol 10:64–65 7. Kascheev VN (1978) Processes in the zone of frictional contact of metals. Mechanical Engineering, Moscow, p 213 8. Korchak SN (1974) The performance of the process of steel parts grinding. Mechanical Engineering, Moscow, p 280 9. Latyshev VN (1985) Improving the efficiency of cutting fluids. Mechanical Engineering, Moscow, p 64 10. Makushin VM (1968) Deformation and stress state of parts in places of contact. Mechanical Engineering, Moscow, p 464 11. Maslov EN (1974) Grinding materials theory. Mechanical Engineering, Moscow, p 320 12. Orlov PN (1988) Technological quality assurance of parts methods of refinement. Mechanical Engineering, Moscow, p 383

1026

I. V. Bashkirtseva and T. N. Orlova

13. Kiselev ES et al (1990) The use of lubricating coolants for grinding materials: technical guidance Material TG 1.4.1928-83. National Institute of Aviation Technologies, Moscow, p 102 14. Shumyacher VM, Bashkirzeva IV (2004) Physical and chemical processes at finishing processing. Monograph. Volgograd State University of Architecture and Civil Engineering, Volgograd, p 161 15. Shumyacher VM, Bashkirtseva IV (2005) Method of determining the specific surface of the products of abrasive dispersion of metal and the study of the kinetics of sedimentation of the products of dispersion of metal in lubricating coolants. Volgograd State University of Architecture and Civil Engineering 16. Bashkirtseva IV (2009) Increasing of the efficiency of superfinishing by rational application of lubricating and cooling liquid (theses), Volgograd, 112p 17. Dementieva NA, Evgrafov YuV, Dorokhov AA, Egorov SA, Marshalov MS (2016) Cooling-and-lubricating liquid “Invetix” for machining metals and alloys 6:40–43 18. Pushkarev OI, Bashkirtseva IV, Brazhnikov DV (2014) Cooling of superhard tools in the finishing of grinding-wheel blanks. Civil Eng 10:653–654 19. Pushkarev OI, Bashkirtseva IV, Brazhnikov DV (2014) Finishing of grinding-wheel blanks by power burnishing. Civil Eng 12:773–774 20. Umut U (2012) How to work with lubricating and cooling liquids correctly. Other Eng Technol 5:39–42 21. Shumiacher VM, Bashkirzeva IV (2014) Rational application of lubricating-cooling liquid in superfinishing processes. Processes of abrasive processing, abrasive tools and materials, pp 10–14 22. Shumiacher VM, Dushko OV, Bashkirzeva IV (2006) Study of technological efficiency of lubricating-cooling liquids in the processes of abrasive processing using methods of mathematical planning. XXVI Russian School on Science and Technology, pp 347–349 23. Shumiacher VM, Dushko OV, Bashkirzeva IV (2005) Micromechanics of contact interaction of an abrasive with a material in the presence of a lubricating-cooling liquid. Processes of abrasive processing, abrasive tools and materials, pp 198–203 24. Shumiacher VM, Dushko OV, Bashkirzeva IV (2005) Investigation of the influence of lubricating-cooling liquids on finishing abrasive processes. Materials and technologies of the XXI century, pp 85–87 25. Shumiacher VM, Dushko OV, Bashkirzeva IV (2005) Investigation of the effect of lubricating-cooling liquids on the friction coefficient of abrasive for metal. Technology of mechanical engineering, pp 39–41

Studies on Titanium Alloy Turning Rate Improvement A. V. Savilov, V. M. Svinin and S. A. Timofeev

Abstract The chapter deals with methods for improving the turning of titanium alloy parts applied in the aircraft construction industry. The factors preventing cutting performance from improvement are analyzed. The influence of temperature on the state of the machined material and cutting tool was assessed. The influence of cutting data on the output turning parameters was studied. In machining of aircraft titanium alloy parts, cutting speed restriction was observed. Turning cutters with carbide inserts were used. During the experiment, cutting forces and temperature were measured. The influence of cutting parameters on machining performance and temperature was studied. The relationship between cutting forces and cutting parameters was identified. Calculation of relative energy consumption for machined material removal proved to be correct. Methods for improving the turning productivity for VT20 titanium alloy parts were developed. Potential areas for studies on a high productivity turning of titanium alloys were suggested. The studies can improve part’s machining performance. Keywords Turning


Productivity
Cutting forces
Cutting temperature

1 Introduction Machining performance improvement is an urgent research task as it can increase the competitiveness of machine-building enterprises by reducing the technological costs of products. The greatest cutting improvement effect can be achieved in the aircraft-, shipbuilding, and other sectors where parts of complex shape are made using hard-to-work materials and high-performance tools. Titanium alloys, nickel, and some grades of stainless and high-alloyed steels are hard-to-machined aviation materials. To machine them, significant cutting forces are applied. These forces cause intensive heat generation. In cutting zones, temA. V. Savilov (&)
V. M. Svinin
S. A. Timofeev Irkutsk National Research Technical University, 83 Lermontov St., Irkutsk 664074, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_109

1027

1028

A. V. Savilov et al.

perature values are high [1]. High-temperature values cause negative changes in the microstructure and properties of tools and materials. High values of cutting force and temperature cause increased wear of cutting edges of the tool. As a result, allowable cutting speed rates and productivity decrease. To increase machinability, the geometry of the cutting edge of the tool has to be improved [2]. Simulation of the turning process is of great importance [3–9]. In machining of some aviation materials, maximum allowable temperatures prevent cutting parameters from being optimized [10]. For example, in the machining of titanium alloys, the cutting speed should not exceed 100 m/min. The cutting speed which is higher than the specified limit can increase the temperature to the level at which structural and phase changes affecting mechanical properties of the material occur. An increase in cutting depth and feed rates can also increase cutting temperature values. Thus, titanium alloy machining performance improvement is a task requiring a comprehensive assessment. Machine tools used by enterprises have specific technical parameters which determine cutting performance. These are power and torque of the spindle. Cutting parameters are calculated in such a way that cutting power and torque are lower than power and torque of the spindle. This prevents equipment from overloading. However, in most cases, it is ignored that the volume of the removed material depends on various combinations of cutting parameters, and different cutting forces and cutting power and torque values correspond to this volume. Thus, assessment of energy consumption per unit removed material is important for improving machining performance in the aerospace industry. Therefore, minimization of energy consumption contributes to maximum performance for a specific technological system. The purpose of this study is to identify optimal methods for increasing titanium alloy turning performance which does not increase cutting temperatures.

2 Materials and Methods Titanium alloy VT20 (Ti-6%Al-2%Zr) was used for the research purpose. For the experiment, a DMG NEF 400 turning machine center (Fig. 1)and a Sandvik Coromant DCLNL 2020 K 12 turning cutter with CNMG 12 04 08-MM inserts (Fig. 2) were used. The tool has the following geometrical parameters: entering angle u = 95°; lip angle e = 80°; rake angle c = 3°; and relief angle a = 6°. Rod-shaped workpieces 90 mm in diameter and 410 mm in length were machined. The workpiece was installed in a three-jaw gripper. Cutting force values were measured with a three-component Kistler 9129AA dynamometer. The surface temperature was measured with a FLIR SC7700 thermal imaging camera. Due to the high rigidity of the technological system, vibrations were not taken into account. However, they have negative effects on the tool life, surface finish quality, and machine tool life [11, 12]. For non-rigid workpieces, the influence of vibrations on cutting performance has to be taken into account.

Studies on Titanium Alloy Turning Rate Improvement

1029

Fig. 1 NEF400 machine area with measuring tools

Fig. 2 DCLNL 2020 K 12 turning cutter

Cutting parameters were changed in the following ranges: depth of cut ap = 0.5…4.5 mm; cutting speed Vc = 15…70 m/min; and feed rate fn = 0.05… 0.35 mm/rev. The experiments were carried out according to the non-compositional plan suggested by Box and Behnken which is the samples of lines from the full factorial 3k-type experiment. Changeable cutting parameters were feed fn(x1), cutting speed Vc(x2), and cutting depth ap(x3). Table 1 shows the levels of changing parameters.

1030

A. V. Savilov et al.

Table 1 Values of experiment factor levels Factors

Feed rate fn, mm/rev x1

Cutting speed Vc, m/min x2

Depth of cut ap, mm x3

Upper level (+1) Main level (0) Lower level (−1) Variability interval

0.3 0.2 0.1 0.1

70 50 30 20

4 2.5 1 1.5

3 Results and Discussion Figures 3, 4, and 5 show the dependences of the cutting force and temperature values on the cutting parameters determined in the experiments carried out in accordance with the Box–Behnken non-compositional plan (see Table 1). Figure 3 shows the cutting force (Fig. 3a) and temperature (Fig. 3b) values determined by varying the feed and cutting speed rates. The nature of the dependences allows us to determine the ranges of cutting speed and feed rates at which cutting force and temperature values are maximum. It is evident that a minimum feed rate increases a temperature value and decreases the cutting performance. An increase in a feed rate up to maximum rates decreases temperature values and increases the cutting performance. Figure 4 shows the cutting force (Fig. 4a) and temperature (Fig. 4b) values determined by varying the cutting depth and feed rates. The surface shape corresponding to the cutting force (Fig. 4a) is indicative of the direct relationship between this output parameter and variable cutting parameters. The maximum value is reached at maximum cutting depth and feed rates. At the maximum cutting depth value, temperature increases at feed rates varying from 0.08 mm/rev to 0.16 mm/ rev and decreases at a feed rate of 0.32 mm/rev.

Fig. 3 Dependence of the output parameters on feed rate fn and cutting speed Vc when turning VT20 titanium alloy: a cutting force; b temperature

Studies on Titanium Alloy Turning Rate Improvement

1031

Fig. 4 Dependence of the output parameters on cutting depth ap and feed fn when turning VT20 titanium alloy: a cutting force; b cutting temperature

Fig. 5 Dependence of the output parameters on depth of cut ap and cutting speed Vc when turning VT20 titanium alloy: a cutting force; b cutting temperature

Figure 5 shows the cutting force (Fig. 5a) and temperature (5b) values determined by varying cutting depth values and speed rates. The surface corresponding to the cutting force has a parabolic shape (Fig. 5a). At the maximum cutting depth value of 4.5 mm, maximum cutting force values are achieved at minimum and maximum cutting speed rates. The minimum cutting force value is achieved at the cutting speed rate of 55 m/min. The temperature dependence is different (Fig. 5b). At the maximum cutting depth value, the temperature reaches its maximum at the cutting speed rate of 55 m/min. The results show the effect of temperature on contact processes in the cutting area causing material softening. According to the experiment results, the dependencies of cutting power and temperature on cutting performance were determined under varying cutting parameters (Fig. 6). The material removal rate was calculated by formula Q = ap  Vc  fn. Cutting power was calculated by formula Pc = (Ft  Vc)/6120 where Ft is the tangential cutting force.

1032

A. V. Savilov et al.

Fig. 6 Dependence of the output parameters on the material removal rate when turning VT20 titanium alloy: a cutting power; b cutting temperature

Analysis of the dependences shows that in turning of titanium alloy to a value of Q  28 cm3/min, energy consumption values are almost equal under examined combinations of cutting parameters (Fig. 6a). At Q values varying from 30 to 40 cm3/min, an increase in performance is more preferable due to increasing cutting depth values than cutting speed or feed rates. Analysis of the relationship between temperature and material removal rates under varying cutting parameters shows by increasing a feed rate we can improve the turning performance (see Fig. 6b). Moreover, minimum temperature values were reached under maximum VT20 alloy turning rates.

4 Conclusion The following conclusions may be drawn from the experiments. CoroTurn DCLNL 2020K 12 turning cutter with CNMG 12 04 08-MM 1015 carbide inserts ensure high titanium alloy turning rates. Cutting parameters which minimize energy consumption per removed material unit were identified. A method for improving turning performance without a critical increase in temperature of the machined material was identified. Vibrations accompanying increasing cutting parameters, vibration elimination methods [13], and surface control methods [7] can be avenues for further research. It will contribute to high surface quality even for rough turning. These methods can be used for control of structural and phase changes in materials under high temperatures and cutting forces [14, 15]. The results can be used to improve VT20 titanium alloy cutting data of turning.

Studies on Titanium Alloy Turning Rate Improvement

1033

References 1. YaN Oteniy, Martynenko OV (2015) Studies on cutting features for cylindrical surfaces of machine parts. Int J Appl Fundam Res 9–3:452–456 2. Karaguzela U, Bakkalb M, Budakc M (2017) Mechanical and thermal modeling of orthogonal turn-milling operation. Procedia CIRP 58:287–292 3. Rodygina AE (2010) Development of a finite element model of non-free cutting. Bull ISTU 6 (46):56–60 4. Balaji JH, Krishnaraj V, Yogesvaraj S (2013) Investigation on high speed turning of titanium alloys. Procedia Eng 64:926–935 5. Jagadesh T, Samuel GL (2014) Investigation into cutting forces and surface roughness in micro turning of titanium alloy using coated carbide tool. Procedia Mater Sci 5:2450–2457 6. Chinesta F, Filice L, Micari F et al (2008) Assessment of material models through simple machining tests. Int J Mater Form 1:507–510 7. Krainev DV, Polyanchikova MY, Bondarev AA et al (2017) Influence of the surface layer characteristics on the regularities of the cutting process. In: MATEC web of conferences, vol 129. https://doi.org/10.1051/matecconf/201712901045 8. Belhadi S, Marouki T, Rigal JF et al (2005) Experimental and numerical study of chip formation during straight turning of hardened AISI 4340 steel. J Eng Manuf 219:515–524 9. Nikolaev, AY (2017) Simulation of the plain milling process. In: IOP conference series: materials science and engineering, vol 177. https://doi.org/10.1088/1757-899x/177/1/012080 10. Mavliutov AR, Zlotnikov EG (2018) Optimization of cutting parameters for machining time in turning process. IOP conference series: materials science and engineering, vol 327. https:// doi.org/10.1088/1757-899x/327/4/042069 11. Svinin VM, Prokhorov AYu (2016) Damping of autovibrations of flexible shaft embodied in centers when turning by multiple-tool head with variable tooth pitch. Vektor Nauki Togliatti State Univ 2(36):67–75. https://doi.org/10.18323/2073-5073-2016-2-67-75 12. Svinin VM, Savilov AV, Zarak TV (2018) Self-oscillation regeneration control by irregular tooth pitches. Adv Eng Res 158:399–405. https://doi.org/10.2991/avent-18.2018.77 13. Eynian M, Altintas Y (2009) Chatter stability of general turning operations with process dumping. J Manuf Sci Eng 131(4). https://doi.org/10.1115/1.3159047 14. Sergeev AS, Tikhonova ZS, Uvarova TV (2017) Method for measuring thermo-emf of a “tool-workpiece” natural thermocouple in chip forming machining. In: MATEC web of conferences vol 129. https://doi.org/10.1051/matecconf/201712901044 15. Nikolaeva EP, Mashukov AN (2017) Evaluation of residual stresses in high-pressure valve seat sur-facing. Chem Pet Eng 53(7–8):459–463

Ultrasonic Impact Study on Strain Hardening of Thread Profile Surface Layer V. Golovkin, O. Batishcheva and V. Papshev

Abstract The paper presents the experimental results of various types of ultrasonic vibrations impact on the strain hardening of a thread profile surface layer when cutting the thread. The research results have shown the effectiveness of using a progressive method of cutting threads in tough titanium alloys with the introduction of ultrasonic vibrations into the cutting zone. The analysis of publications on this issue showed that many authors also note the viability and economic efficiency of using ultrasonic technologies in the industry. This paper also presents the results of a comprehensive study of ultrasonic vibrations impact, varying in directions, on the quality formation of the surface layer of a thread, namely the important indicator affecting the performance of thread pieces, that is, strain hardening. The attention is drawn to the fact that the formation of the surface layer on different parts of the thread profile occurs in different ways. Accordingly, the effect of ultrasonic vibrations in different areas manifests itself in various ways, as the research findings have shown. From the obtained regularities, it has been established that the use of tangential ultrasonic vibrations leads to a decrease in strain hardening on the side surface of the thread and the root, while processing with axial vibrations, an increase in the micro-hardness on the side surface occurs, and when cutting with radial vibrations, the side surface and the root of the thread are subjected to hardening. Keywords Strain hardening


Thread
Ultrasonic vibrations

1 Introduction In modern technologies, in order to increase the efficiency of various types of mechanical processing, several physicochemical methods are widely used, in particular, with the ultrasonic vibration’s use. It should be noted that today the industry V. Golovkin (&)
O. Batishcheva
V. Papshev Samara State Technical University, 244, Molodogvardeyskay str., Samara 443100, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_110

1035

1036

V. Golovkin et al.

is in an urgent need for new technologies for the titanium and its alloy’s processing. This is due to its widespread use in various fields, including biomedicine and aerospace field. In this regard, the processing technology of titanium and its alloys based on the use of ultrasonic vibrations [1–7] is promising and economically justified. It should be noted that, unlike other processing methods, ultrasonic treatment leads to a significant improvement in the operational characteristics of the working surfaces that is important for durability and performance of products under extreme operating conditions [8–12]. The use of ultrasonic vibrations for the process of cutting threads deserves special attention. This allows to improve the performance of the process and durability of the tool, as well as to provide the required quality of the surface layer, on which the performance characteristics of thread pieces depend [13–15]. Due to the fact that the thread has a complex profile, different areas of physical and mechanical properties are formed in the surface layer during machining. It is a common practice to distinguish three main zones of a thread, that is, its crest, side surface and root. It is known that the performance of thread pieces is significantly affected by the quality of the surface layer, namely residual voltages, surface roughness and strain hardening [16, 17]. At the same time, it should be noted that when cutting threads with imposing ultrasonic vibrations on the thread-forming tool, there is a significant change in the quality characteristics of the surface layer, the direction of ultrasonic vibrations being a particularly significant factor.

2 Experimental Procedure To identify the patterns of ultrasonic vibrations influence on the strain hardening of the thread profile surface layer, the authors undertook special studies. According to the experimental method, the determination of micro-hardness on the thread profile was carried out according to the scheme shown in Fig. 1.

Fig. 1 Chart of micro-hardness measuring on the thread profile

Ultrasonic Impact Study on Strain Hardening of Thread Profile …

1037

The samples for the study were made as follows. With the help of special ultrasonic devices [1], M6 threading was carried out on the bars of titanium alloy BT16. In this case, threading was realized with the introduction of various ultrasonic vibrations in the cutting zone: tangential, axial and radial. In order to obtain a comparative evaluation, threading without ultrasound was performed.

3 Results As a result of the experimental studies, the following data were obtained. Figure 2 presents the results of the direction impact of ultrasonic vibrations on the micro-hardness of the thread profile crest. The experimental data obtained show that for all types of processing, there is no significant change in micro-hardness at the crest of the threaded profile. This is due to the fact that there is no direct impact of the tool on the thread profile crest. Figure 3 shows the results of experimental studies related to the influence of ultrasonic vibration’s direction on the micro-hardness in the root of the thread. From the dependences illustrated in Fig. 3, it can be seen that the introduction of radial and axial ultrasonic vibrations into the cutting zone leads to an increase in the micro-hardness in the subsurface layer compared to cutting without ultrasound. The impact of tangential ultrasonic vibrations on the instrument leads to a decrease in micro-hardness values. Figure 4 shows the results of experimental studies related to the influence of ultrasonic vibration’s amplitude on the micro-hardness in the root of the thread. The results of an experiment to study the influence of ultrasonic vibration’s direction on the micro-hardness of the thread lateral surface are given in Fig. 5.

Fig. 2 Influence of the ultrasonic vibration’s direction on the micro-hardness of the thread profile crest (according to C-C): 1—threading without ultrasonic vibrations; 2—threading with tangential ultrasonic vibrations; 3—threading with axial ultrasonic vibrations; 4—threading with radial ultrasonic vibrations

1038

V. Golovkin et al.

Fig. 3 Influence of ultrasonic vibration’s direction on the micro-hardness in the root of the thread (along A-A): 1— threading without ultrasonic vibrations; 2—threading with tangential ultrasonic vibrations; 3—threading with axial ultrasonic vibrations; 4 —threading with radial ultrasonic vibrations

Fig. 4 Influence of ultrasonic vibration’s amplitude n on the micro-hardness in the root of the thread (along A-A): 1— threading without ultrasonic vibrations; 2—threading with tangential ultrasonic vibrations; 3—threading with axial ultrasonic vibrations

Fig. 5 Influence of ultrasonic vibration’s direction on the micro-hardness in the middle of the lateral surface of the thread (along the section B-B): 1—threading without ultrasonic vibrations; 2— threading with tangential ultrasonic vibrations; 3— threading with axial ultrasonic vibrations; 4—threading with radial ultrasonic vibrations

Micro-hardness in the experiment was measured in the middle of the thread side surface. The results of an experiment to study the influence of ultrasonic vibration’s amplitude on the micro-hardness of the thread lateral surface are given in Fig. 6. Micro-hardness in the experiment was measured in the middle of the thread side surface. From the dependences experimentally obtained, it can be seen that ultrasonic vibrations in the axial and radial direction contribute to the hardening of the surface

Ultrasonic Impact Study on Strain Hardening of Thread Profile …

1039

Fig. 6 Influence of ultrasonic vibration’s amplitude n on the micro-hardness of the thread side surface (along B-B): 1— threading without ultrasonic vibrations; 2—threading with tangential ultrasonic vibrations; 3—threading with axial ultrasonic vibrations

Fig. 7 Influence of ultrasonic vibration’s direction on micro-hardness along the side surface of the thread (D-D): 1 —threading without ultrasonic vibrations; 2— threading with tangential ultrasonic vibrations; 3— threading with axial ultrasonic vibrations; 4—threading with radial ultrasonic vibrations

layer. Changing the direction of vibrations to tangential vibrations reduces micro-hardness compared to threading without ultrasound. The results of a study of ultrasonic vibration’s direction impact on micro-hardness along the side surface of the thread can be seen in Fig. 7. When measuring micro-hardness along the surface of the thread from crest to root, the distribution of micro-hardness values has a distinctive feature, namely that an increase in the depth and degree of work hardening is observed as it approaches the root of the thread. Perhaps this is due to the more severe conditions of the final profile formation of the thread in its root. The results of a study of the amplitude of ultrasonic vibrations impact on micro-hardness along the side surface of the thread can be seen in Fig. 8. Fig. 8 Influence of ultrasonic vibration’s amplitude n on the micro-hardness of the thread side surface (along D-D): 1— threading without ultrasonic vibrations; 2—threading with tangential ultrasonic vibrations; 3—threading with axial ultrasonic vibrations

1040

V. Golovkin et al.

It should be noted that an increase in the amplitude of ultrasonic vibrations to the values n = 5 lm does not significantly affect the change in micro-hardness.

4 Summary Thus, in all variants of threading when applying radial ultrasonic vibrations to the tool, an increase in micro-hardness in the root and on the side surface of the thread was observed, which is associated with high-frequency shock-cyclic interaction of the tool with the thread surface. When machining with axial ultrasonic vibrations, the lateral surface of the thread profile is subjected to shock-cyclic impact, which also leads to an increase in micro-hardness. In the case of introducing tangential vibrations into the cutting zone, the micro-hardness decreases, as in this instance, the formation of the surface layer occurs at lower cutting forces. In this, the shock-cyclic effect of the thread-forming tool on the thread profile is absent. As a result, we can say that when machining with axial or radial ultrasonic vibrations, the micro-hardness of the thread surface can be increased by 10–20%. This will ultimately improve the performance of thread pieces.

References 1. Golovkin VV, Batishcheva OM, Papshev VA (2018) Enhance the efficiency of the internal threads cutting process with taps by applying ultrasonic vibrations. In: ICMTMTE 2018, MATEC web of conferences, vol 224 2. Golovkin VV, Batishcheva OM, Papshev VA (2017) Enhance the efficiency of the internal threads cutting process with taps by applying ultrasonic vibrations. Izvestia Volgograd State Techn Univ 9(204):85–88 3. Chen P, Liao WB et al (2018) Ultrafast consolidation of bulk nanocrystalline titanium alloy through ultrasonic vibration. Scientific reports, vol 8, Article number 801 4. Gorunov AI, Nyukhlaev OA, Gilmutdinov AK (2018) Investigation of microstructure and properties of low-carbon steel during ultrasonic-assisted laser welding and cladding. Int J Adv Manuf Technol 99(9–12):2467–2479 5. Shao L, Wu S et al (2018) Microstructure and mechanical properties of ultrasonic pulse frequency tungsten inert gas welded Ti-22Al-25Nb (at.%) alloy butt joint. J Mater Process Technol 259:416–423 6. Ward A, Zhang Y, Cordero Z (2018) Junction growth in ultrasonic spot welding and ultrasonic additive manufacturing. Acta Mater 158:393–406 7. Zheng J, Ince A, Tang L (2018) Modeling and simulation of weld residual stresses and ultrasonic impact treatment of welded. Joints Procedia Eng 213:36–47 8. Sharma V, Pandey PM (2018) Experimental investigations and statistical modeling of surface roughness during ultrasonic-assisted turning with self-lubricating cutting inserts. J Process Mech Eng 232(6):709–722

Ultrasonic Impact Study on Strain Hardening of Thread Profile …

1041

9. Li S, Wu Y et al (2017) Improving the grind ability of titanium alloy Ti–6Al–4V with the assistance of ultrasonic vibration and plasma electrolytic oxidation. CIRP Ann—Manuf Technol 66:345–348 10. Abbasi A, Amini S, Emamikhah A (2016) Design and implementation of the ultrasonic cold forging technology process for improving surface mechanical properties of 6XB2C cold-worked alloy steel tool. J Eng Manuf 230:2 11. Fromentin G, Poulachon G et al (2005) Precision and surface integrity of threads obtained by form tapping. CIRP Ann 54(1):519–522 12. Yea C, Xianfeng Z et al (2016) Surface amorphization of NiTi alloy induced by ultrasonic nanocrystal surface modification for improved mechanical properties. J Mech Behav Biomed Mater 53:455–462 13. Agapov SI, Golovkin VV (2010) Increase of efficiency of mechanical processing with the use of ultrasound. Publishing Office of the Samara Scientific Center, Samara 14. Vologin MF, Kalashnikov VV et al (2002) Application of ultrasound and explosion during processing and assembly. Machine Building, Moscow 15. Singh R, Khamba JS (2007) Taguchi technique for modeling material removal rate in ultrasonic machining of titanium. Mater Sci Eng, A 460–461:365–369 16. Singh Khamba JS (2006) Ultrasonic machining of titanium and its alloys: a review. J Mater Process Technol 173:125–135 17. Statnikov E (2004) Physics and mechanism of ultrasonic impact treatment. IIW Document 13:2004–04

Dissipative Structure of Contact Interaction When Cutting Metals V. A. Kim, B. Ya. Mokritsky and A. V. Morozova

Abstract Processing of metals cutting proceeds in the system of non-equilibrium processes including high-speed plastic deformation of the cutoff layer, contact and frictional interaction of the processed material with the asymmetric cutting wedge and its wear. The synergetic algorithm of development of the non-equilibrium process provides the formation of dissipative structures which arise in the system of cutting of metals. The structure and mechanisms of functioning of a dissipative structure of contact and frictional interaction opening new approaches of the management of processes of cutting and quality of machining are considered. The dissipative structure includes the insular and continuous outgrowths covered with the adsorbed and amorphous superficial films, and also the deformation strengthened layer outgrowths. Dissipation is connected with the production of entropy and is carried out due to the work of frictional interaction on a forward surface of the cutting wedge and is defined by friction coefficient size between the sliding shaving and an external surface of outgrowths. The total coefficient of friction is defined by such dissipative structure and spontaneously reaches such sizes at which the density of thermal stream and tension in the zone of frictional interaction extend on the area of primary plastic deformations and minimize deformation processes of shaving formation and work in the shift plane. Varying the modes of cutting and the external technological environment when cutting materials, it is possible to influence actively the dissipative structure for ensuring the necessary quality and productivity of machining.




Keywords Dissipative structure Contact interaction Non-equilibrium process Tool performance





Metal cutting


V. A. Kim
B. Ya. Mokritsky Komsomolsk-on-Amur State University, 27, Lenin Avenue, Komsomolsk-on-Amur 681013, Russia A. V. Morozova (&) Bryansk State Technical University, Boulevard of the 50th Anniversary of October, Bryansk 241035, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_111

1043

1044

V. A. Kim et al.

1 Introduction Dissipative structures are understood as a set of elements and links of a non-equilibrium system that have “artificial intelligence” that convert absorbed internal energy into thermal dissipative flows and minimize the effects of external energy influences [1–3]. Current investigations of dissipative structures in different environments [4–9] are based on classic fundamentals of the theory of their formation and development [10–13]. The approach given is used also in theoretical investigations of the contact interaction in definite environments at the realization of different processes [14–17] and in the analysis of experiment results accompanying them [18, 19]. The properties of dissipative structures are determined by the laws of non-equilibrium thermodynamics, but their structure and functioning mechanism are individual by nature, depending on the nature of the non-equilibrium system, and is not always obvious [1, 20, 21]. In systems of contact-friction interaction, the initial period or the running-in period should be considered as a stage of a dissipative structure formation and a steady-state period as a stage of a formed and functioning dissipative structure [1, 20, 22]. Metal cutting refers to a non-equilibrium process involving high-speed plastic deformation and destruction of the material being processed, friction and wear of the cutting wedge. The disclosure of the structure and mechanism of functioning of dissipative structures during metal cutting allows us to predict the nature of a non-equilibrium process development and reveals new algorithms to control the quality of mechanical processing and the cutting tool performance. The article describes the composition of the dissipative structure of the contact interaction when cutting metals and the mechanisms of its functioning.

2 Research Methodology The studies were carried out during the process of turning of various structural materials (45 steel, 12X18H10T stainless steel and VT22 titanium alloy) with cutting tools made of R6M5 high-speed tool steel in semi-finishing cutting conditions. The morphology of the surface layer of the cutting tool wedge was studied using a S3400-N scanning electron microscope (Hitachi). The shrinkage of the chips was determined by the gravimetric method. The structural and deformation state of the chip formation zone was analyzed by the microstructure of the chip roots. The microstructure itself was recorded with a Nikon 200A metallographic microscope at a magnification of 400, and the resulting digital images were processed by ImagePro.Plus.5.1 (USA).

Dissipative Structure of Contact Interaction When Cutting Metals

1045

3 The Construction of the Dissipative Structure of the Contact-Friction Interaction During Cutting The result of contact-friction interaction and mutual mass transfer of the processed and tool materials during cutting is the formation of build-ups. The continuity of the build-ups and the strength of their adhesion to the cutting wedge surface are determined by their location on the tool working surfaces. Figures 1 and 2 show some types of build-ups on the front surfaces of turning tools made of P6M5 high-speed steel when turning 45 and 12X18H10T steel. Similar types of build-ups were obtained by turning the titanium alloy VT22. The build-ups are divided into continuous and discontinuous. The continuous build-up consists of discontinuous clings and differs in higher strength of adhesion with the

Fig. 1 Build-ups on the front surface of the high-speed cutter made of P6M5 when turning 12X18H10T

Fig. 2 Build-ups on the front surface of the high-speed cutter made of P6M5 when turning 45 steel

1046

V. A. Kim et al.

substrate and degree of plastic deformation. In the structure of build-ups, the probability of the presence of amorphous phases is high. The discontinuous build-ups are characterized by a long-range order associated with their orientation in the direction of the sliding chips. Such orderliness is preserved in the continuous growth morphology. In the pictures of continuous and discontinuous build-ups obtained using a scanning electron microscope, there are bright areas representing microzones with a high electrostatic charge. It can be assumed that these are most likely zones with stable adsorption films and an amorphous phase, with relatively high dielectric properties. Figure 3 schematically shows the construction of the dissipative structure of the contact interaction process when cutting metals, explaining some of the mechanisms of its operation. On the working front surfaces of the cutting wedge 1, discontinuous and continuous build-ups are located. Their strength of adhesion with the substrate depends on their location relative to the cutting edge. The closer the build-up is to the cutting edge, the higher is the strength of its adhesion with the cutting wedge. Build-ups are highly deformed formations predominantly made of the material being processed, having a nanostructured organization, saturated with decomposition products of the technological environment and chemical elements of the tool material [23]. The outer surface of individual discontinuous build-ups is covered with a stable adsorption film 3, and it may be absent on the surface of a continuous and some discontinuous build-ups. Chips 4 slide along the tops of the build-ups. In separate cavities between the troughs of discontinuous build-ups and chips, microvolumes of a lubricating cooling process medium (LCPM) and products of the external technological environment 5 can be found. Under the build-ups, there is a cutting tool subsurface layer with a modified structure 6 changed as a result of

Fig. 3 Construction of the dissipative structure of the contact process when cutting metals: 1—front surface of the cutting wedge; 2—discontinuous build-ups; 3—adsorption film; 4—chips; 5—lubricating cooling process medium; 6—hardened layer

Dissipative Structure of Contact Interaction When Cutting Metals

1047

plastic deformations and thermal effects that can initiate strain hardening and diffusion transfer of elements from deep layers or external technological medium to the surface layers. The dissipative structure is influenced by a flow of high-density thermal and mechanical energy transmitted through sliding chips. Contact interaction proceeds under non-equilibrium conditions, while external dynamic disturbances are determined by the chip formation process, and dynamic reactions to these disturbances are generated by the dissipative structure itself. The process dynamics assume that all speed, power and thermal parameters of the contact interaction system are nonlinear stochastic functions of time and the generalized state coordinate of the system [22, 24, 25]. The heat flux coming from the chip formation zone permeates the contact-friction interaction system and dissipates in the external environment. It affects the mechanical and deformation properties of the cutting tool contact layers, the stability of the adsorption film on the outer surfaces of discontinuous build-ups, the friction coefficient between sliding chips and the cutting tool, as well as the build-up adhesion strength with the cutting wedge front working surface. The system of contact-friction interaction is dynamic, in which various changes are constantly taking place, for example, the removal of discontinuous build-ups, individual sections or an entire continuous build-up. The build-up removal is preceded by a discontinuity of the adsorption film on its outer surface and the appearance of a “dry” metal contact. This leads to the formation of a strong adhesive bond between the build-up and sliding chips. If the strength of such a bond is higher than the strength of adhesion between the build-up and the cutting wedge surface, then the build-up simply removes without destroying the tool’s working surface. If the adhesive bond between the build-up and the tool is sufficiently high, then the destruction occurs either on the build-up itself or at the deep level of the tool surface layer, which represents an elementary wear event.

4 The Functioning of the Dissipative Structure of Contact Interaction The dissipative structure functions in such a way as to minimize the influence of external disturbances, by implementing the principle of least action [1, 24]. The algorithm for reducing external influences includes minimizing the friction coefficient between the build-up and chips, the formation of a strong adhesive bond between the build-up and the cutting wedge surface, the maximum hardening of the tool subsurface layer and the increase in the specific wear work. Let us consider in more detail each of these mechanisms. The stability of the film adsorption depends on the temperature and power conditions of the contact process. With an increase in temperature and contact stresses, the antifriction function of the adsorption film decreases. A stable

1048

V. A. Kim et al.

adsorption film can exist only in the area of discrete contact and in the presence of surface-active additives in LCPM [26, 27]. In the absence of an adsorption film, the outer surface of the build-up can provide a low friction coefficient due to a strongly deformed amorphous phase. Consequently, one of the conditions for the dissipative structure’s normal functioning is to ensure the minimum friction coefficient between the build-up and sliding chips. The friction coefficient during cutting is an integral value that takes into account different modes of friction processes in individual contact areas. a1 f1 ðtÞ þ a2 f2 ðtÞ þ a3 f3 ðtÞ ! min

ð1Þ

where f1 ðtÞ—coefficient of dry friction between the build-up and sliding chips; f2 ðtÞ—friction coefficient between the build-up and chips in the presence of an adsorption film; f3 ðtÞ—friction coefficient between the build-up and chips in the presence of the amorphous phase; a1 , a2 , a3 —weigh coefficients considering area ratio of the contact areas with different modes of external friction. A decrease in the friction coefficient leads to a change in the stress tensor in the chip formation zone, and this changes the stress-strain state in the entire area of primary chip formation deviations [28]. The strength of the adhesive bond is determined by the accumulation of surface defects of the crystal structure, concentrating around itself the elastic internal energy. With the approach of two surfaces and their mutual plastic deformation, a part of this energy is released in the form of a thermal impulse, facilitating the formation of strong metal or molecular bonds between the contacting surfaces. Consequently, the strength of the adhesive bond can be increased by increasing the density of surface defects of a crystal structure. Hardening of the subsurface layer of the cutting wedge and the increase in the specific wear work represent a single process associated with an increase in the surface microstructure chemical potential. Taking into consideration the deformation nature of the contact-friction interaction during cutting, the main mechanism for hardening the cutting wedge surface layer will be strain hardening due to the generation of dislocations. The equation of energy balance of contact-friction interaction in energy flows can be represented as follows: pða1 f1 þ a2 f2 þ a3 f3 Þ t ¼ Q þ aadh Ftrue þ DlVhard þ awear DM

ð2Þ

where p—average normal stress in contact; aadh—adhesion bond formation energy; Ftrue—true contact area; Dl—increment of chemical potential during strain hardening of the tool contact layers; awear —specific wear work; DM—loss of mass per unit contact area of the cutting wedge during wear; Vhard —the volume of the hardened surface layer of the tool; Q—heat flux released during contact-friction interaction. The heat flux Q takes into account only the energy that was released directly in the process of contact-friction interaction in the zone of secondary plastic

Dissipative Structure of Contact Interaction When Cutting Metals

1049

deformation, i.e., does not take into account the heat coming from the chip formation zone. Let us single out the components of Eq. (2) related to the free energy absorbed during contact-friction interaction. U ¼ aadh Ftrue þ awear D þ DlVhard

ð3Þ

Then, the balance equation takes the following form: p
ð a1 f 1 þ a2 f 2 þ a3 f 3 Þ
t ¼ Q þ U

ð4Þ

A dissipative structure, by definition, must provide a maximum of free energy, i.e., dU ¼ 0

ð5Þ

The non-equilibrium system of contact-friction interaction in its development tends to the stable dynamic equilibrium condition, when the normal pressure in the contact, chip sliding speed and temperature in the cutting zone are stabilized. Therefore, in the steady-state period of the process, the cutting system is in a quasistable state. In this case, the Eq. (4) takes the form pt
d ða1 f1 þ a2 f2 þ a3 f3 Þ ¼

d2 S T dt2

ð6Þ

where S—contact process entropy; T—absolute temperature; t—time. From (6), it follows that the production of entropy in the contact steady-state process occurs predominantly in the zone of frictional interaction of chips with surface layers of build-ups on the cutting wedge. All other dissipation mechanisms are activated periodically for a short period of time with peak disturbances. The rate of entropy production will be determined by the integral value of the friction coefficients. The processes of contact-friction interaction have a direct impact on the chip’s formation through the angle of action [29], which is determined by x ¼ arctgða1 f1 þ a2 f2 þ a3 f3 Þ  c

ð7Þ

The influence of the dissipative process on the resultant cutting force can be represented by the following relation: qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi R ¼ Nbef 1  ða1 f1 þ a2 f2 þ a3 f3 Þ2 where Nbef —normal load on the cutting wedge front surface.

ð8Þ

1050

V. A. Kim et al.

Then, the cutting work will be equal to Ap ¼ Nbef t

qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1  ða1 f1 þ a2 f2 þ a3 f3 Þ2
cos½arctgða1 f1 þ a2 f2 þ a3 f3 Þ  c

ð9Þ

The presence of a close connection between the processes occurring in the zones of primary and secondary plastic deformation is confirmed by the relationship between the quantitative microstructural indicators of chip formation and the conditions of contact-friction interaction during metal cutting [28].

5 Conclusions 1. The dissipative structure that forms and functions in the contact process consists of discontinuous and continuous build-ups, an adsorption film on their outer surfaces and a reinforced subsurface layer of the cutting wedge. 2. During the period of steady (quasi-stationary) flow of the cutting process, dissipation in the zone of secondary plastic deformations is realized due to different modes of friction processes between the descending chips and the external surfaces of the build-ups. By influencing the nature of the friction process, the durability of the cutting tool and the quality of the machining can be controlled.

References 1. Ivanova V, Balankin L, Bunin I et al (1994) Synergetic and fractals in materials science. Science, Moscow 2. Kim V (2001) Self-organization in the processes of hardening, friction and wear of the cutting tool. Dal’nauka, Vladivostok 3. Migranov M, Shuster L (2011) Features of thermodynamic processes on the cutting tool contact surfaces. Proc Samara Sci Center Russian Acad Sci 13(4(3)):1126–1129 4. Ali A, Yaqub S, Usman M, Zuhaib KM, Khan AM et al (2018) Motion planning for a planar mechanical system with dissipative forces. Robot Auton Syst 107:129–144 5. Cramer MS, Crickenberger AB (1991) Dissipative structure of shock waves in dense gases. J Fluid Mech. https://doi.org/10.1017/s0022112091001441 6. De Matteis G, Brando G (2018) Comparative analysis of dual steel frames with dissipative metal shear panels. Key Eng Mater 763:735–742 7. Kondoh Y (1993) Eigenfunction for dissipative dynamic operators and the attractor of the dissipative structure. Phys Rev E Stat Phys Plasmas Fluids Relat Interdisc Top 48(4):2975– 2979 8. Malkov VB, Nikolaenko IV, Shveikin GP et al (2018) Formation of dissipative structures in an amorphous film. Dokl Phys Chem 478(2):39–41 9. Wessling B (1993) Dissipative structure formation in colloidal systems. Adv Mater 5(4):300– 305 10. Artigiani R (1987) Revolution and evolution: applying Prigogine’s dissipative structures model. J Soc Biol Struct 10(3):249–264

Dissipative Structure of Contact Interaction When Cutting Metals

1051

11. Condepudi D, Prigogine I (2014) Modern thermodynamics: from heat engines to dissipative structures. Wiley, New York 12. Edelstein BB (1970) Instabilities associated with dissipative structure. J Theor Biol 26 (2):227–241 13. Prigogine I, Lefever R (1968) Symmetry breaking instabilities in dissipative systems. II. J Chem Phys 48:1695 14. Fazeli N, Tedrake R, Rodriguez A (2018) Identifiability analysis of planar rigid-body frictional contact. Robot Res 665–682 15. Jayaraman A (2019) Coarse-grained models for predicting structure and thermodynamics in polymer systems with specific and directional intermolecular interactions. Bulletin of the American Physical Society. Session H42: Dillon Medal Symposium, Boston, Massachusetts 16. Neto AG, de Mattos Pimenta P, Wriggers P (2018) Contact between spheres and general surfaces. Comput Methods Appl Mech Eng 328(1):686–716 17. Palasantzas G, Babamahdi Z, Svetovoy V (2018) Casimir interactions of complex surfaces and materials. Bulletin of the American Physical. APS, abstract id. V07.008 18. Chen X, Deng X, Xu L (2018) A three-dimensional dynamic model for railway vehicle–track interactions. J Comput Nonlinear Dyn. https://doi.org/10.1115/1.4040254 19. Guo J, Ding L, Gao H, Guo T, Liu G (2018) An apparatus to measure wheel-soil interactions on sandy terrains. IEEE/ASME Trans Mechatron 23(1):352–363 20. Fedorov S, Assenova E (2017) Synergetic principle of self-organization during friction. Bull Sci Educ North-West Russia 3(3):1–20 21. Kremneva LV, Snegireva KK, Ershova IV (2014) Method of calculating the energy dissipation coefficient for cutting materials. Vestnik MGTU Stankin 4(31) 22. Kim V, Yakubov F, Skhirtladze A (2017) Mesomezanic contact interaction processes during friction and cutting of metals. TNT, Stary Oskol 23. Kim V, Karimov Sh (2014) Manifestation of physical mesomechanics during contact interaction and wear. Scientific notes of Komsomolsk-on-Amur State Technical University II-1 (18):79–85 24. Kabaldin Yu, Oleynikov A, Burkov A (2003) A synergistic approach to the analysis of dynamic processes in machine tools. Mach Tools 1(2):3–6 25. Zakovorotny V, Tung Fan Din, Bykador V (2014) Self-organization and bifurcation of a dynamic system for metal cutting. Proc Univ Appl Nonequilibr Dyn 22(3):26–39 26. Verkhoturov A, Yakubov F, Kim V, Konevtsov L, Yakubov C (2014) The role of air in contact processes of metal cutting. Sci Notes Komsomolsk-on-Amur State Techn Univ III-1 (19):65–72 27. Yakubov Ch (2008) The hardening effect of the cutting fluid during metal cutting. Simferopol City Printing House, Simferopol 28. Kim V, Otryaskina T, Sarilov M (2014) Structural and quantitative ratios of the process of chip formation. Fundam Res 6:933–936 29. Bobrov V, Granovsky G, Zorev N et al (1967) The development of the science of cutting metals. Mechanical Engineering, Moscow

Cutting Temperature by Polymer-Abrasive End Brushes for Machining Planes D. B. Podashev and Yu. V. Dimov

Abstract Polymer-abrasive end brushes are very effective at finishing the planes on parts from various materials. The article carries out the thermos-physical analysis of the process of treatment with end polymer-abrasive brushes. In this case, the total thermal power of the process Q arises from the conversion of mechanical work into heat and depends on the cutting force and the speed of movement of the heat source. In the presence of heat exchange with the environment and the steady-state processing, the temperature in the cutting zone depends on the intensity of the source, the contact area of the brush with the surface of the workpiece and the thermal properties of the interacting bodies. The regularities of the influence of the treatment modes on the cutting temperature are established. Experimental studies confirmed the adequacy of the developed analytical temperature determination to the actual processing process. When setting processing modes, the cutting temperature must not exceed the melting temperature of the polymer base of the brush villus. Knowing the thermos-physical properties of the polymer-abrasive material and the dimensions of the brush villus, according to the developed mathematical model, it is possible to determine the temperature occurring in the cutting zone. In the development of technological processes, this makes it possible to select processing regimes for performing given finishing operation, excluding the fusion of the polymeric binder of the brush. Keywords End brush deformation Feed





Cutting temperature
Rotational frequency
Brush

1 Introduction Operations’ mechanization and automation of parts finishing are relevant for modern engineering. Polymeric-abrasive brushes are very effective tools for finishing surfaces on parts in aircraft construction, missile construction and other D. B. Podashev (&)
Yu. V. Dimov Irkutsk National Research Technical University, 83 Lermontov str., Irkutsk 664074, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_112

1053

1054

D. B. Podashev and Yu. V. Dimov

branches of engineering. Polymer-abrasive end brushes can effectively be used to prepare surfaces for galvanic and paint coatings, to provide the required roughness after blade machining. However, such a tool with a binder of a polymer is very sensitive to the temperature that occurs during processing. High temperature can cause fouling of the surface layer of villi, which is unacceptable. Therefore, when setting processing modes, it is necessary to take into account the temperature that occurs during cutting.

2 Methods Mathematical modeling of the process of heat formation during the interaction of the brush with the material being processed was carried out using the thermal physics processes of metals’ mechanical treatment [1], the theory of thermal phenomena during grinding [2–4] and mathematical analysis. Experimental confirmation of the analytical studies’ results was carried out using Scotch-Brite ™ end brushes: BD-ZB Bristle P50; BD-ZB Bristle P80; BD-ZB Bristle P120 ∅115 mm of 3 M (Minnesota Mining and Manufacturing Company) with granularity P50, P80 and P120. The brush contains 540 villi, located in 36 rows, and in each row of 15 villi, the length of the villi is 20 mm.

3 Results and Discussion Polymer-abrasive end brushes (PAEB) are effectively used for surface finishing. However, such a tool with a binder of a polymer is very sensitive to the temperature that arises during processing. High temperature can cause fouling of the surface layer of villi, which is unacceptable. Therefore, when setting treatment modes, it is necessary to take into account the temperature that occurs during cutting. A number of studies have been devoted to the study of temperature in the cutting zone when working with abrasive tools, when grinding [1–4], when processing by radial polymer-abrasive brushes—[5–11]. However, there are no studies of temperature during the treatment of planes with polymer-abrasive end brushes. The processed material is a high-strength aluminum alloy V95pvhT2, widely used in aircraft building. The process of heat generation in the treatment of PAEB planes occurs as a result of the interaction of polymer-abrasive villi of the rotating brush with the surface being treated. The total thermal power of the process Q (W) is due to the conversion of mechanical work into heat:

Cutting Temperature by Polymer-Abrasive End Brushes …

Q ¼ Px
vx þ Py
vy þ Pz
vz ;

1055

ð1Þ

where Px, Py and Pz—component forces of cutting along the axes X, Y and Z, N; vx, vy and vz—velocity of the source along the axes X, Y and Z, m/s. The input part of the heat balance, according to [1], during machining can be represented as heat equivalent to the work of deformation and heat equivalent to the work of friction at each of the m sections where it occurs. The consumption part of the heat balance of the processing of the PAEB plane can be written (according to [2]) with the following equation: Q ¼ Qp þ Qb þ Qc þ Ql ; where Qp, Qb, Qc, Ql—the proportion of heat, going, respectively, to the part, the brush, the chips, the cooling liquid. Due to the fact that the study of PAEB treatment was carried out dry, and without cooling, the fraction of heat going to the coolant is excluded from the formula. The calculations carried out by Sypailov [3] establish that the fraction of heat going into the circle when grinding with a conventional abrasive tool is no more than 1%. When treated with brushes, this component is even smaller and it can be neglected. The fraction of heat that flows into the chips according to the data [2] for ordinary flat end grinding is only 3%. Since very fine particles of material are formed when treated with polymer-abrasive brushes, and also regarding the small amount of thermal energy in the chips, the component can be neglected. In the method of sources, the mathematical apparatus is based on a description of the temperature field that arises in an unbounded body under the action of heat introduced by a planar stationary source with an intensity of q (W/m2). q¼

Q ; F

where F is the contact area of the brush with the workpiece being processed, m2. In the presence of heat exchange with the environment and a steady-state process, this temperature field is described by the expression [1]: " rffiffiffiffi# q b Hðx; 1Þ ¼ pffiffiffi exp x ; x 2k b

ð2Þ

where H(x, ∞)—body temperature with coordinates x, ∞, arising from the heat source, °C; k—coefficient of thermal conductivity of the processed body, W/(m °C); x—coefficient of thermal diffusivity of the body in which the source moves, m2/s;

1056

D. B. Podashev and Yu. V. Dimov

b¼

2
a
p ; c
q
F

where a—heat transfer coefficient, W/(m2 °C); p—figure perimeter of the brush contact surface with the workpiece being processed, m; cq—coefficient of volumetric heat capacity (c—specific heat capacity of the mass, q—density), J/m3 (°C); x—distance from the edge of the source of thermal energy to the point at which the temperature is determined, m. The velocity of the heat source in (1) depends on the position of the workpiece end relative to the brush axis (size A in Fig. 1). When the brush starts processing, plane A varies from 51.5 mm to 0, then from 0 to−51.5 mm. Since the velocity depends on the radius on which the villi are located, we will calculate the average radius. For A < R1, the mean velocity is defined as: 9 8 p
n >
ðR þ R1 Þ
sin dav ; > vx ¼ > > > > 30 > > = < p
n
ðR þ R1 Þ; vy ¼ > > 30 > > > > > ; : v ¼ p
n
ðR þ R Þ
cos d > z 1 av 30

Fig. 1 Scheme to determine the weighted average angle between the Z-axis and direction of the velocity

ð3Þ

Cutting Temperature by Polymer-Abrasive End Brushes …

1057

For A  R1, the average velocity is 9 8 p
n >
ðR þ AÞ
sin dav ; > vx ¼ > > > > 30 > > = < p
n
ðR þ AÞ; vy ¼ > > 30 > > > > > ; : v ¼ p
n
ðR þ AÞ
cos d > z av 30

ð4Þ

The weighted average angle dav. between the Z-axis and the velocity v direction of the heat energy source in Eqs. (3) and (4) as the angle between the radius passing through the center of the figure CDE (see Fig. 1) and the X-axis. For A < R1, dav ¼

Fk ; R2  R21

where Fk—area in contact: Fk ¼

R2 R2
ða  sin aÞ  1
ða1  sin a1 Þ; 4 4

where a ¼ 2arccos AR ; a1 ¼ 2arccos RA1 : For A  R1, dav ¼

R2

Fk ;  A2

2

where Fk ¼ R4 ða  sin aÞ  12 ðdav
R2  A2
tgdav Þ: This is accepted that dav  tgdav : The average angle dav, the area Fk and the perimeter of the contact surface are shown in Table 1. The heat transfer coefficient a characterizes the intensity of thermal energy transfer from the source to the workpiece being processed. As noted in [1], this indicator depends on a number of factors: the volume of the body, its configuration, the nature of the contact with another body, the temperature and the properties of the environment, as well as the parameters of the processing process (brush rotation speed and its deformation).

Table 1 Mean values of the angle dav, area and perimeter of the contact surface

Parameter

A (mm) 0

22

30

38

dav, deg. Fk (mm2) p (mm)

45 2607.52 300.752

32.40 1653.043 213.571

21.48 1223.436 182.638

15.28 641.932 145.711

1058

D. B. Podashev and Yu. V. Dimov

Taking into account the variety of factors affecting the value of a, it was decided to determine this coefficient experimentally. The results of the study are given in Table 3. It is established that with an increase in the brush rotational speed n and the feed S, the coefficient of heat transfer increases, and with increasing strain, Dy— decreases. The regression equation for a is obtained from the rotational speed n (min−1), the brush deformation along the Y-axis—Dy (mm) and the feed S (mm/min). a ¼ ðd1
n2 þ d2
n þ d3 Þ
ðd4
Dy2 þ d5
Dy þ d6Þ
ðd7
S2 þ d8
S þ d9 Þ: ð5Þ The values of the coefficients and the free terms of Eq. (5) are given in Table 2. Calculation of the temperature at the contact surface of the end polymer-abrasive brush with the machined plane Ht at distances from the brush end x = 0, x = 5 mm and x = 10 mm was carried out under the following conditions: Initial data: 1. Processed material: high-strength aluminum alloy V95pvhT2: k = 155.03 W/(m °C), cq = 54.892.6 J/(m3 °C), x = 5  10−5 m2/s, a—by the Eq. (5). 2. Cutting force Px, Py and Pz—according to work [12], accepted as the arithmetic mean for three brushes: BD-ZB Bristle P50; BD-ZB Bristle P80; BD-ZB Bristle P120, differing only in the granularity of the material of the villi. The results of calculating the temperature according to the proposed method are given in Table 3. To confirm the adequacy of the proposed theory of the occurrence of temperature in the cutting zone, polymeric-abrasive brushes were used to conduct experimental studies of the temperature at x = 0 using the FLIR Orion SC7200 (series SC7000) thermal imager (Sweden). Figure 2a shows the experimental temperature dependences on the rotational speed of the brush. Table 3 shows the theoretical and experimental values of temperatures in the cutting zone. It is established that the temperature increases with increasing rotation speed, since the speed of the source [according to (1)] directly affects the thermal power of the process Q. In addition, with increasing speed, all the cutting force components Px, Py and Pz also increase in Eq. (1). It is established (Fig. 2b and Table 3) that with increasing deformation of the brushes’ villi, Dy the temperature increases. This is due to the increase in cutting forces Px, Py and Pz, which affect the mechanical power of the process.

Table 2 Values of the coefficients and the free terms in (5) Symbol Value Symbol Value

d1 8.999  10−6 d6 7.07

d2 –4.553  10−3 d7 4.5289  10−5

d3 4.926 d8 –4.0816  10−3

d4 1.7733 d9 0.7652

d5 –6.7066 – –

Cutting Temperature by Polymer-Abrasive End Brushes …

1059

Table 3 Temperature in the cutting zone by the end brush Ht, °C x=0

Parameter

Px, N

n, min−1 500 1000 1250 1600 Dy, mm 0.5 1 1.5 2 S, mm/min 82 130 255 395 A, mm 0 22 30 38

By Dy = 1.5 mm; S = 130 mm/min; A = 0 65.10 23.53 151.75 75.14 80.00 33.00 180.00 117.55 94.95 37.61 190.75 132.36 124.28 43.91 202.02 150.42 By n = 1000 min−1; S = 130 mm/min; A = 0 49.50 26.00 121.00 50.75 63.50 30.00 148.00 74.107 80.00 33.00 180.00 117.55 99.00 35.00 217.00 153.40 By n = 1000 min−1; Dy = 1.5 mm; A = 0 77.558 26.69 178.19 130.54 80.023 33.00 180.00 117.56 84.929 46.85 184.74 83.16 87.827 57.91 190.05 63.02 −1 By n = 1000 min ; Dy = 1.5 mm; S = 130 mm/min 80.00 33.00 180.00 117.55 65.75 19.19 122.02 99.03 53.68 14.14 94.70 91.51 37.93 9.13 64.04 70.96

Pz, N

Py, N

He, °C x=0 75.14 117.55 133.64 150.41 50.75 82.51 117.55 155.87 130.54 117.55 88.21 63.02 117.55 – – –

Fig. 2 Dependence of temperature in the cutting zone from: a rotation speed n at Dy = 1.5 mm, S = 130 mm/min; b deformation of the brush Dy at n = 1000 rpm, S = 130 mm/min; c feed S at n = 1000 rpm, Dy = 1.5 mm for the end brushes: 1—BD-ZB P50, 2—BD-ZB P80, 3—BD-ZB P120

As the supply increases, the temperature decreases (Fig. 2c), which is associated with an increase in the heat transfer coefficient a in accordance with formulas (2), (5) and Table 3. In conclusion, it should be noted that the proposed theory of calculating the temperature in the cutting zone adequately reflects the actual process.

1060

D. B. Podashev and Yu. V. Dimov

When assigning processing modes, it is inadmissible that the cutting temperature is higher than the melting point of the polymer bond of the circle. For example, the melting point of polyamide-6 and nylon is 215 °C, polyurethane—175 °C, polystyrene—220 °C.

4 Conclusion End polymer-abrasive brushes are very effective at finishing the planes on parts from various materials. Knowing the thermo-physical properties of the polymerabrasive material and the dimensions of the brush villi, according to the developed mathematical model, one can determine the temperature arising in the cutting zone. In the development of technological processes, this makes it possible to justifiably select processing regimes for performing current finishing operation, eliminating the fusion of the polymeric binder of the brush.

References 1. Reznikov NA (1981) Thermophysics of metals mechanical processes. Mechanical Engineering, Moscow 2. Svirschev VI, Podbornov IV, Zubairova LH (2011) Thermal balance with flat face planetary grinding. Izhevsk Bull IzhSTU 1(49):11–13 3. Sipailov VA (1978) Thermal processes during grinding and surface quality management. Mechanical Engineering, Moscow 4. Diligensky NV, Kamayev YuP (1969) On the thermal physics of the grinding process. Phys Chem Mater Process 1:43–50 5. Abrashkevich Y, Pelevin LE, Machishin GM (2011) The influence of thermal processes on the performance of a polymer-abrasive brush. Indust Constr Eng Facili: Sci Prod 3:44–47 6. Abrashkevich YD, Pelevin LE, Machishin GM (2014) Modeling the heat transfer process of a polymer-abrasive brush and surface. In: Transport and transport-technological systems: materials of the international scientific and technical conference, p 3–7 7. Machishin GM (2014) Determination of the rational application of a polymer-abrasive tool. Bull Kharkov Natl Autom Highway Univ 65–66:117–122 8. Ustinovich DF, Golub VM (2012) Control of thermal regimes in the treatment of flat surfaces by disk polymer-abrasive brushes. Bull Polotsk State Univ Ser B: Ind Appl Sci 3:90–94 9. Ustinovich DF, Golub VM (2011) Modeling of thermal processes during processing of revolution bodies by polymer-abrasive brushes. Bull Belarus Natl Sci Acad Ser Phys Tech Sci 2:62–68 10. Ustinovich DF, Golub VM (2012) Modeling of thermal processes during processing of flat surfaces by polymer-abrasive brushes. Bulle Belarus Nat Sci Acad Ser Phys and Tech Sci 4:63–68 11. Ustinovich DF, Golub VM (2011) Management of thermal conditions for grinding shafts with brushes. Mech Mach Mech Mater (Minsk) 3(16):38–42 12. Dimov YV, Podashev DB (2018) Cutting forces when working planes with polymer-abrasive end brushes. Messenger IrSTU 22/5 (136):28–45. https://doi.org/10.21285/1814-3520-20185-28-45

Researching the Influence of the Location Tool on the Treatment of Large Shafts Requiring Surface Shaping Y. A. Bondarenko, N. A. Maslennikov and A. A. Mamchenkova

Abstract Rotary kilns for clinker production furnaces have lengths of up to 220 m and sleeve diameters of 6 m. The metal shell is up to 80 mm thick at the dress rings that transmit their mass on oven support rollers. Due to temperature variations (precipitation lining) shells can become deformed and the bent rotational axis of the furnace can resulting in a ring becoming trapped in this zone. In addition to rotating about its axis of rotation there is also movement in the plane of rotation, something that depends on the curvature of the axis of the furnace. This movement leads to the fact that the ring is not always situated on two rollers. Therefore, the side machine tool module, that is mounted on the imaginary axis of rotation of the furnace and is in connection with the ring is not always in contact with the cutting tool. Mathematical modeling of the processing is necessary to find the dependence of the ring.




Keywords Processing Large part Installation Placement





Cutter
Precision
Roughness


1 Introduction For mathematical simulation the handling ring must be found dependence between determining the position of the surface (axis) with respect to the cutter for different angles of rotation, and the surface coordinate values of the ring after cutting (Fig. 1). If we take the amount of movement of point A to be the displacement of the coordinate axes then on the basis of similar codirectional coordinate vectors we have the following representation by vectors x; y; z through x0 ; y0 ; z 0 :

Y. A. Bondarenko
N. A. Maslennikov (&)
A. A. Mamchenkova Belgorod State Technological University Named After V.G. Shukhov, 46, Kostyukov Str., Belgorod 308012, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_113

1061

1062

Y. A. Bondarenko et al.

Fig. 1 Linear displacement of the coordinate system

2 3 0 8
x 1 0 <
x ¼ 1
x þ 0
y þ 0
z 6 7
y ¼ 0
x þ 1
y þ 0
z or 4
y 5 ¼ @ 0 1 :
z ¼ 0
x þ 0
y þ 1
z 0 0
z

10
x0 1 0 B C 0 A@
y0 A 1
z0

ð1Þ


will in all cases be different (Fig. 2): However, vector OO ð2Þ ð3Þ ð4Þ Consequently: OA ¼ xA
x þ yA
y þ zA
z;

O0 A ¼ O0 O þ OA ¼ x0A x0 þ y0A y0 þ z0A z0 ;

ð5Þ

where A denotes the index of the coordinates of A in the old system. Finally, when we move axially along OX we have:

ð6Þ

Researching the Influence of the Location Tool …

1063

Fig. 2 Angular displacement around X axis coordinates

Moving axially along OY gives:

ð7Þ

Moving axially along OZ gives:

ð8Þ

Therefore, when the rotational axis is displaced, and its axes are in a vertical plane, the processed ring is displaced by the cutting tool by amount A. If the roller ring is formed with points of tangency circle radius of the ring at this point, then the displacement axis radius of the circle will increase, that is, this point will not be in a ring on the rollers. Knowing the size of A allows the calculation of gap size d1 and d2 —if there is no gap it is possible to calculate the actual size of the ring [1–3]. When the linear displacement of the coordinates of point A is an angular displacement around its axes—if its displacement relative to the X axis is most pronounced—then relative to axes Z and Y it is not expressed explicitly. This displacement depends on the displacement of the coordinate axes of rotation of the ring along axes OZ and OY, so it is necessary to consider this relative to axis OX:

1064

Y. A. Bondarenko et al.

xA x ¼ xA x0 þ 0y0 þ 0z0

ð9Þ

yA y ¼ 0x0 þ yA cos ay þ yA sin az0

ð10Þ

zA z ¼ 0x0  zA sin ay0 þ zA cos z0

ð11Þ

32 30 1 x0A 1 0 0 xA 4 y0 54 0 cos a  sin a 5@ yA A A z0A 0 sin a cos a zA

ð12Þ

Consequently: 2

Thus, in addition to its linear displacement the point also has an angular one, which occurs if the distance d on one of the rollers is 0 while on the other roller d 6¼ 0; or when the values of d on the rollers have opposite signs, that is, d1 [ 0; d2 \0 or d1 \0; d1 [ 0: Distances d1 and d2 can be calculated mathematically (Fig. 3) [4–6]. Having considered the linear and angular displacement of the rotational axis of ring O0 A; the support rollers are set at an angle of 30°—on the basis of experimental data and mathematical calculations. Therefore, if the radius of the perfect ring axis direction and the offset value are known, it allows one to calculate clearances d1 and d2 between the ring and roller. If these values are compliant with calculations, it is possible to create treatment programs with the help of a computer.

Fig. 3 Calculating the cutter position in the presence of gaps d1 and d2

Researching the Influence of the Location Tool …

1065

Based on the fact that the values of rolling surface defects are small relative to the diameter of the ring, the movement of the treatment surface, with respect to the cutting tool, can be taken as half the sum of the errors: dA ¼ ðd1 þ d2 Þ=2

ð13Þ

When calculating error dA one should take into account the signs of distances d1 and d2 , since they can be either positive or negative. In our case, the center axes offset was above or below axis OY. Therefore, while processing the ring, via the installation of a cutting tool under the ring and between the rollers, it is important that any cutting depth error should not exceed 0.5 mm. The following case (Fig. 4) most often occurs during processing of large parts by the additional machine tool when mounted on the module support with the cutting tool to one side, sitting below the horizontal axis of the part. Extra machine modules mounted on side view of the cutting tool, due to virtue of their design, define a plane in 45 about a horizontal axis of rotation of the items. In this case, if there is a single defect in the detail then we have the situation where d1 [ 0; d2 ¼ 0; when the reel defect on the first roller ring is displaced upwards and towards the second roller, whereby one of the three points (the two rollers and the cutter) form a circle and move circumferentially and the allowance for distortion is not completely removed. As a consequence, the defect is copied and an adjustment must be made to the movement of the tool. The angular position of point B; where a defect is copied, is calculated as follows: the first roller is set at 30° about a vertical axis, and the cutting tool at an angle a to this plane. Consequently, the difference (a  30 ) in repetition angle gives the maximum value of the defect. Knowing the sizes of d; R and the angles of the triangle KOB we can calculate the value of FB [1, 3, 7]. In the case where d1 \0 (Fig. 4), the workpiece is moved towards the cutting tool and is comprised of the defect. If the size d1 is big enough, then cutting tool breakage may occur. Considering that movement is 30° to the vertical plane and

Fig. 4 a Position of the ring for d1 [ 0; d2 ¼ 0: b Position of the ring for d1 \0; d2 ¼ 0

1066

Y. A. Bondarenko et al.

that the tool is angled to it at 45° (with all other factors being known), it is possible to calculate the defect manifest in the cutting tool. Knowing the difference between the angles and the radius of the workpiece it is easy to determine the value of BF for the introduction of corrections to the tool’s movement. Thus, in certain cases (Fig. 4) we can consider the passage of a single defect (d1 6¼ 0Þ: from the first roller it passes through the second with the second video being copied as a mirror image, that is, if d1 \0 the ring is pushed over the tool, when it passes through the second roller, an opposite phenomenon occurs, with the ring farther away from the cutting tool and the tool copying the defect in an inverted form. The same thing happens when d1 [ 0: Consequently, in this case, the installation tool copies the existing defect at an angle of 45° to the vertical plane. To avoid this phenomenon, it is necessary to first determine defect processing and then to remove it by controlling the movement of the cutting tool. When such a defect is removed it is necessary to set the depth of cut based on the equipment and cutting tools strength while it should not be setted more than the half defect size, as shown above. One of the most common positions for the cutting tool is on the horizontal axis of the shaft. Figure 5a shows the behavior of the ring with a single defect when d1 [ 0 and when the tool is located on the horizontal shaft. In this case the ring is displaced towards the second roller and the cutter defect is copied—its value can be calculated since certain values are known: O1 K ¼ R þ d1 , the angle of the rollers at 60°, and the angle of cutter installation at 90°. The horizontal axis, where the single defect is found, is moved against the direction of rotation at an angle a. Knowing the offset center OO1 and angle b one can determine the FB copying defect and its time of occurrence. Figure 5b shows morphogenesis at d1 \0 and at the location of the cutting tool on the horizontal axis of the ring. After contact with the defect, d1 \0 on the roller ring, it “zooms in” on the cutter and the defect is copied. Knowing the center O; the value of OK ¼ R  d1 , and angles a and b , one can calculate the value of BF and its time of occurrence [8, 9].

Fig. 5 a At ring position R: b At ring position d1 \0

Researching the Influence of the Location Tool …

1067

So, in both cases (Fig. 5) defects are copied to the cutting tool. In the first case, copying occurs closer to the defect, and is greater in magnitude than the second case. In both cases, defects falling on the second roller are also copied to the specular reflection and manifest themselves about a horizontal axis being symmetrically opposite in magnitude. So, in these cases, the manifestation of a single defect occurs as a result of processing from one to another. However, the resulting defects are smaller in size. Through mathematical modeling and experimental turning it was found that the greatest magnitude defects are copied if the cutting tool is closer to the base surface. The best mode-cutter arrangement is when it is based between the two rollers. In this case, if the ring (shaft) loses regular geometric shape (Fig. 6), that is, if it has many defects, then the number of defects increases exponentially upon processing. This means that the treated surface will tend to a polyhedron shape, close to a circle [9, 10]. When processing is necessary in order to find the largest imperfection, determining the depth of cut requires a gradual approach to defects, that is, processing them in small quantities. One peculiarity of this processing will be that if defects are of different sizes then the center of the circle obtained will shift towards the area containing the smaller quantity of defects [4, 5, 11]. Figure 7 shows displacement of the center circumferentially during recovery. The new circle is not perfect, therefore, further processing using grinding stones is required to ensure it acquires the right shape.

Fig. 6 Processing rings with a large number of defects

1068

Y. A. Bondarenko et al.

Fig. 7 Displacement of the circle’s center during the recovery process

2 Conclusion We have resolved questions of cutting tool installation area for forming bulky rotating parts and manifestation of defects when installing a tool in different zones of the workpiece. We have also defined the cutting depth starting position when machining workpieces with large numbers of defects.

References 1. Fedorenko TM, Fedorenko MA, Bondarenko YA (2009) Restore the health of the axles of the grinding mills with the use of a portable machine. Eng Technol 3:20–21 2. Fedorenko MA, Bondarenko YA, Fedorenko TM (2008) Machine for machining axles grinding millspatent for utility model. RUS Patent 75339. 07 Feb 2008 3. Bondarenko YA, Fedorenko MA (2003) Method of machining of large shafts without dismantling. Bondarenko Y.A., Fedorenko. Patent for invention RUS Patent 2242346. 05 May 2003 4. Bestuzheva OV, Fedorenko MA, Bondarenko YA (2016) Experimental study of the recovery of a surface of revolution of large parts of industrial equipment. Bull Belgorod State Technol Univ V. G. Shukhov 11:122–127 5. Fedorenko MA, Bondarenko YA, Sanina TM, Markova OV (2015) Auxiliary machine for machining internal surfaces of large cylindrical type. Eng Technol 11:27–28 6. Fedopenko TM, Fedopenko MA, Bondarenko YA (2009) Recovery of CAPP pumping mills with the application of a penable machine. Eng Technol 3:20–21 7. Fedorenko MA, Bondarenko YA, Fedorenko TM (2008) Study of the ensuring of the necessary roughness of the surface of large-size rotating parts by universal machine modules. Bull Belgorod State Technol Univ V.G. Shukhov 2:35–38

Researching the Influence of the Location Tool …

1069

8. Fedorenko MA, Bondarenko JA, Pogonin AA (2018) Impact of oscillations of shafts on machining accuracy using non-stationary machines In the collection. In: IOP conference series: materials science and engineering 11. CEP. In: International conference on mechanical engineering, automation and control systems 2017—Processing Equipment, Mechanical Engineering Processes and Metals Treatment, 042030 9. Bondarenko YA (2004) Special machine for treatment of dacing large shaft, equipped with a tool self installation device. STIN 1:36 10. Bondarenko JA, Fedorenko MA, Pogonin AA (2018) Research of the value of linear distortion of renewable surface of part during rotary processing of bulky items without dismantling unit. In the collection. In: IOP conference series: materials science and engineering 11. CEP. International conference on mechanical engineering, automation and control systems 2017—processing equipment, mechanical engineering processes and metals treatment, 042031 11. Beloborodov SM, Bondarenko YA, Vereshchak AS et al (2012) Technological support of quality and resource in manufacturing, assembling, repair and restoring. The 120th anniversary of the birth of the outstanding aircraft designer, Hero of Socialist Labor, twice the USSR State Prize laureate N. N. Polikarpov is dedicated. Series Engineering: Technology, Equipment, Moscow 12. Fedorenko MA, Bondarenko JA, Sanina TM (2018) Influence of forces acting on side of machine on precision machining of large diameter holes. In the collection. In: IOP conference series: materials science and engineering 11. CEP. International conference on mechanical engineering, automation and control systems 2017—processing equipment, mechanical engineering processes and metals treatment, 042018 13. Ambrosimov SK, Bondarenko YA, Vereshchaka AS et al (2016) Progressive machine building technologies, equipment and tools, vol 7. The Collective Monograph, Moscow 14. Bestuzheva OV, Fedorenko MA, Bondarenko YA (2016) Experimental study of the restoration of the surface of rotation of large-size parts of industrial equipment. Bull Belgorod State Technol Univ V.G Shukhov 11:122–127 15. Fedorenko MA, Bondarenko YA, Sanina TM, Markova OV (2015) A free machine for treatment of large-sized internal surfaces of cylindrical type. Eng Technol 11:27–28 16. Gorbatykh VV, Bondarenko YA (2014) Improvement of efficiency of repair and accuracy of installation of equipment under conditions of operation excluding disassembly of equipment in the collection: youth and knowledge—a guarantee of success. In: Collection of scientific papers of the international scientific and technical conference, pp 93–94 17. Sanina TM, Bondarenko YA, Fedorenko MA (2009) Machine for treatment of internal surfaces TSAPF pomolnyy mills patent for utility model. RUS Patent 89830. 25 Aug 2009 18. Fedopenko MA, Sanina TM, Bondarenko YA, Pogonin AA, Skhirtladze AG (2009) Removed understanding of functional bargeway agents repairs. Recov Mod 11:11–14 19. Bondarenko YA, Fedorenko MA (2003) Method for treating large shafts without their demontages patent for invention. RUS Patent 2242346. 05 May 2003 20. Bondarenko YA, Fedorenko MA (2003) Unruptated restoration of tapp mill pipes. Construct Mater 8:16

Numerical Modeling of the Material Layer Upset Forging with Extrusion Under the Stiffening Rib into the Forging Cavity O. A. Nikitina and T. M. Slobodyanik

Abstract An array of input data is formed and boundary conditions defined for carrying out numerical experiments on the material layer for upset forging into a forging cavity under the stiffening rib. Numerical experimentation is applied to plane upset forging of a blank from an initial body width of ho = 16.13 mm to a finite width of h4 = 12.5 mm in four stages. Calculations are made for each of the four steps with a decrease in body thickness at the first step of 0.63 mm; and a decrease of 1 mm thickness at the second, third, and fourth steps. At the same time, a cavity width under a stiffening rib was accepted by equal finite thickness of forging body, cavity conjugate radius with deformed plane of forging was 15 mm. The stiffening rib finite calculation height was 30 mm. During the research we quantitatively estimated stress changes and the deformation ratio in fibers external to forging. Calculation results were displayed in the form of distortion of finite-elementary grids and isolines pictures that considered sizes in the blank material, and also the stress distribution in externally forged fibers. Keywords Forging Deformation


Stiffening rib
Upset forging
Pressure
Stress


1 Introduction To ensure the profitable activity of various industrial enterprises an increase in the efficiency of equipment is very important. The economic aspects of the rational design of industrial equipment has been considered in several research papers [1–9]. O. A. Nikitina (&) Nosov Magnitogorsk State Technical University, 38, Lenin Ave., Magnitogorsk 455000, Russia e-mail: [email protected] T. M. Slobodyanik National Research Technology University NUST “MISIS”, 4, Leninsky Ave., Moscow 119049, Russia © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_114

1071

1072

O. A. Nikitina and T. M. Slobodyanik

With development of branches of machine construction and with die forging in particular, the production of details with difficult configurations on a thin body and high stiffening ribs have sufficient rigidity and small increased weight. Purpose of development and implementation each of the name of a product is defect-free of forging according to rational technology scheme. Application in forging and stamping production hydrostatic presses with large deformation force changed relation to design process aluminum alloys forging. In forging and stamping production main volume of design efforts are made by semi-empirical methods [10]. For processes of die forging design application of the models is relevant. But nature research are single, owing to imperfection of theory of developments and absence of information in publications. The main current methods used for solving tasks in fluid–plastic environments are approximate methods and methods based on the collateral decision of approximate equilibrium equations and plasticity equations. Methods for solving tasks associated with large plastic deformations are realized by the finite element method. The main difficulties associated with the application of this method are the large number of calculations and the method’s reliability in terms of boundary conditions. Program complex applications for solution of this tasks promoted widely used this method for research processes of a ribbing forging. The non-linear, stress-deformed state of a material is described by an equation in vector form [11]. Anand’s model allows us to consider deformation hardening at an increased deformation ratio:   feg ¼ eth þ eel ;

ð1Þ

 T feg ¼ ex ; ey ; exy is the sum vector of deformation and

ð2Þ

 T  th e ¼ DT ax; ay ; 0 is the vector of temperature deformation

ð3Þ

where

for which DT determines current and input temperature difference; and parameters ax and ay are the coefficients of thermal expansion on axes OX and OY respectively:

1=Ex ½D1 ¼ myx =Ex 0

 el e ¼ ½D1 frg; where

ð4Þ

  frg ¼ rx ; ry ; rxy is the stress vector and

ð5Þ

mxy 1=Ey 0

0 0 is the matrix that is inverse to the matrix of plasticity; 1=Gxy

ð6Þ

Numerical Modeling of the Material Layer Upset …

1073

the components of which are Ex and Ey ; that is, Young’s modulus, and Gxy and vxy that represent the shearing modulus and Poisson’s coefficient on axes OX and OY respectively. Deformations are bound to movements at certain points by: feg ¼ ½Bfug

ð7Þ

where [B] is the matrix of derivation of movement and {u} is the movement vector. The collateral solution of Eqs. (1) and (7) allow one to get a dependence for definition the vector of plastic deformation as:  el  e ¼ ½Bfug  eth

ð8Þ

The solution of this equation demands the preparation of boundary conditions and an array of input data. Plastic forming region 1 (Fig. 1) is limited. Bottom half of a forging 2 is accepted as a fixed one that is deprived of all degrees of mobility. The top half of the tool is seated on the OY axis. Numerical experience is carried out and applied to alloy 1024. Values of s, η, and a are determined by Anand’s method where s is stress sensitivity to the deformation ratio, η is the coefficient of material sensitivity to deformation hardening, and a is the coefficient of influence on material hardening to the distribution of the deformation ratio. Blank forging was divided into 480 elements and deprived of any broadening. The array of input data included the tension modulus of the forging material Ef = 2  105 MN/m2 and of the blank material Eb = 0.746  105 MN/m2, the Poisson coefficient m = 0.3, the forging temperature 320–400 °C, the blank temperature 400 °C, and the friction coefficient between blank and forging materials l = 0.15.

Fig. 1 Scheme of forming input data and boundary conditions: 1—plastic-forming region; 2— bottom half of forging; and 3—top half of forging

1074

O. A. Nikitina and T. M. Slobodyanik

One research numerical method is the non-steady process of monitoring the form change of a metal, based on the application of the finite element method, allowing analysis of this processes for a finite number of body deformation ratio increments. A similar analysis, applicable to plane upset forging of blanks from initial body widths of ho = 16.13 mm to finite widths of h4 = 12.5 mm over four stages, was completed on paper. Calculations were completed in four steps with different upset forging body ratios: a decreased body thickness on the first calculation step of 0.63 mm and body thickness decreases of 1 mm on the second, third, and fourth steps. At the same time, cavity widths under the stiffening rib were accepted according to the equal finite thickness of the forging body—the cavity conjugate radius with a deformed plane of forging being 15 mm. The stiffening rib finite calculation height was 30 mm. During research a quantitative estimate was made of the changes in stress and deformation ratios in the external fibers of a forging. Calculation results were displayed in the form of the distortion of finite-elementary grids and pictures of isolines that considered sizes in the blank material. The first calculation step gives minimum values for ratios of deformation and stress. Increases in these parameters occur with a decrease in body thickness—the location of maxima and minima values remain invariable at each calculation step. For the first step, maximum values of ratio deformation and stress are located on the blank surface and over the rib zone. After the second pass the maximum values of ratio deformation move to an area of interconnection between the materials and tools, beginning to form an area of intense current located over the rib zone on the surface. Intense growth of the current area is followed by increased rib length. Maximum values of stress distribution are similar to maximum values of deformation ratio distribution, as an increase in stress occurs over the rib zone, in the rib itself, and in the area of the zone-forming radius. After the third pass the area of intense current increases and approaches the rib, with maximum values of deformation ratio located in the area of contact friction between materials during forging. Maximum values of stress are located on the surface detail: in an area of contact between the forged deforming surface and radius zone. Upon the last pass the area of intense material current reaches its maximum and is located at the conjugation of the rib and forging body. The maximum values of the deformation ratio and stress ratio arrangement are similar to values encountred during the previous pass. The core of maximum values, of both deformation ratio and stress ratio, is located in the central plane of symmetry of the stiffening rib—visible from isoline distributions. Results of the stress-deformed state (material tension on axis OX and compression on axis OY) can be seen on the back of the material (strength to outflow). The pressure distribution on surfaces produced via deforming forging at different stages of upset forging is presented in Fig. 2. The increased minimum values of pressure from one stage to another are: from first to second stage 22–25%, from second to third stage 13–16%, and from third to fourth stage 17–20%. The increased maximum values are: from first to second stage 22–25%, from second to third stage 32–35%, and from third to fourth stage 7–10%.

Numerical Modeling of the Material Layer Upset …

1075

Fig. 2 Pressure distribution of the deforming forging surface at different stages of upset forging: 1—h1 = 15.5 mm; 2—h2 = 14.5 mm; 3—h3 = 13.5 mm; 4—h4 = 12.5 mm

The distribution of the deformation ratio in external forging fibers at different stages of upset forging is presented in Fig. 3. Stress distributions in external forging fibers at different stages of upset forging are presented in Fig. 4.

Fig. 3 Distribution deformation ratio in external forging fibers at different stages of upset forging: 1—h1 = 15.5 mm, 2—h2 = 14.5 mm, 3—h3 = 13.5 mm, and 4—h4 = 12.5 mm

1076

O. A. Nikitina and T. M. Slobodyanik

Fig. 4 Stress distributions in external forging fibers at different stages of upset forging: 1— h1 = 15.5 mm, 2—h2 = 14.5 mm, 3—h3 = 13.5 mm, and 4—h4 = 12.5 mm

The distributions of deformation ratio (Fig. 3) and stress ratio (Fig. 4) in external forging fibers occurs thus: minimum values of deformation ratio and stress ratio occur over the rib zone. Inflexions of lines occur in the radius zone and the central part of the forging body. Analysis of results from the stress-deformed state at upset forging, flowing into a forging cavity under a stiffening rib, shows that the presented method is suitable for calculations of processes that is associated with plastic forging shaping with developed lateral area. Numerical experimentation with these considered processes will permit the collection of significant results in subsequent research.

References 1. Nikitina OA, Slobodyanik TM (2008) Selection of profitable loan conditions on the basis of annuity payments. Appl Math Econ Tech Studies 2(2):140–142 2. Nikitina OA, Kharitonov AO (2017) Elastic stamp for forgings with reduced allowance. In: “Irreversible processes in nature and technology”: proceedings of the ninth All-Russian conference. MSTU them. N.E. Bauman, p 33 3. Nikitina OA, Slobodyanik TM (2017) Budgeting as an instrument of financial management of a construction organization. “Modern management model: problems and prospects”. In: Materials of the international scientific and practical conference. Under the General Editorship of N.V. Kuznetsova, pp 73–77

Numerical Modeling of the Material Layer Upset …

1077

4. Nikitina OA, Slobodyanik TM, Melikhova YuM (2016) Development of an action plan for restructuring the accounts payable of the repair enterprise. Guide Entrepreneur 31:114–120 5. Slobodyanik TM, Nikitina OA (2017) Investigation of the influence of the construction of conveyors on inertial parameters. In: “Education and its role in shaping the worldview of modern man”: materials of the international scientific and practical conference, pp 137–141 6. Nikitina OA, Litovskaya YV, Savinkova TA, Zinoveva EG, Ponomareva OS (2017) The use of the budget planning mechanism in construction companies: evidence from LLC “STROYTEKHNOLOGIYA”. Espacios 38(33):17–26 7. Nikitina OA (2005) Development of design methodology for stamping of aluminum panels with single-sided fins using the vertical hydraulic presses. Moscow State University of Steel and Alloys, Moscow 8. Nikitina OA, Litovskaya YV, Ponomareva OS (2018) Development of the cost management mechanism for metal products manufacturing based on budgeting method. Acad Strat Manage J 17(5):1–7 9. Gerasimova AA, Radyuk AG, Titlyanov AE (2015) Creation of a diffusional aluminum layer on the narrow walls of continuous-casting molds. Steel Transl 45(3):185–187 10. Solomonov KN, Nikitina OA (2003) New developments in die forging. Steel Transl 33 (3):44–48 11. Gunghin L, Chun L, Sheng S, Tingdong G A system of CAS/CAD system for die forging process. J Jpn Light Metals 989(7):295–300

Centrifugal Rolling of Flexible Shafts for Achieving Best Possible Roughness of the Surface N. V. Vulykh

Abstract The improvement of the surface plastic deforming technology is pertinent in the machine building technology and constitutes the optimization of deforming processes while forming microgeometry of the surface. Shafts and an axis represent the majority of parts manufactured in machine building plants. Today requirements for processing accuracy, surface finish, fatigue strength, and durability of products are becoming more demanding. When a shaft has a small rigidity, the surface layer plays a more important part in ensuring the stability of the form and size of the shaft. The research presents an effective technology of centrifugal rolling of flexible shafts in comparison with traditional techniques of surface plastic deforming, e.g., ball and roller rolling that allows to strengthen the surface layer of a product with the specified force action. The construction of a centrifugal rolling machine is presented as a deforming instrument. Centrifugal processing regimes are set allowing to roll details made of carbon steel with carbon mass content of *0.2 to *0.6% and optimal surface finish according to arithmetic average of the roughness profile (Ra) and mean spacing of profile irregularities (Sm). Keywords Rolling Flexible shaft


Roughness
Surface plastic deforming
Surface of a part


1 Introduction Improving operational characteristics of machine parts that operate under dynamic and contact loading and conditions of wear is essential for the modern machine building industry.

N. V. Vulykh (&) Irkutsk National Research Technical University, 83, Lermontov St, Irkutsk 664074, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_115

1079

1080

N. V. Vulykh

In this regard, the issue of machine parts performance is pertinent. This results in expanding the application of the processes of the strengthening technologies with surface plastic deforming (SPD) being one of them. Under SPD, the surface layer undergoes quality changes: Micro-roughness is levelled, solidity and durability are enhanced, and residual compressive stress is created that improves cyclic strength of parts [1–5]. The researchers often use modeling processes to study the stress-deformed condition of the surface layer [6–12]. One of the disadvantages of the rolling process when rolling flexible shafts is that local deforming does not produce the desired effect because of considerable specific loading and as a result the products lose their geometry in the axis direction. Here, this drawback of the process is eliminated by the parallel deforming impact of three rollers on a part at an angle 120° using centrifugal rolling machine [13]. Thus, the goal of the article is to produce flexible shafts with the processed surface that has a minimum arithmetic average of the roughness profile (Ra) and mean spacing of profile irregularities (Sm).

2 Materials and Methods The rolling process was performed on a screw-cutting lathe 1K62. The samples made of normalized carbon steel: steel 20 (C  0.2%; nominal value  150); steel 35 (C  0.35%; NV  180); steel 45 (C  0.45%; NV  200); and steel 60 (C  0.6%; NV  225), with the length of (l) 300 mm, diameter (d)—16 mm, was chucked in a footstock spindle of the lathe and with the help of a support longitudinal movement is given to the samples. The samples were not rigid (l=d [ 12Þ and had initial roughness Rainit = 1.25; 2.5 and 5.0 µm with an angle at micro-roughness apexes of 60° edgeways, in the rolling plane that lies on the axis of the shaft. The processing of the samples was performed by a rotating rolling machine that is installed in a 3 jaw chuck with 4 different movements (S) and 4 different rates of rotation (n) (Table 1). The change in the rotation rate of the rolling machine helped to change the force put by the polishing roller on the sample. Correlation of the rotation rate and rolling force is presented in [14], and the values of the rolling forces (F) for the corresponding rotation rates of the rolling machine are shown in Table 1. Rollers 3 with the profile radius 5 mm produced from steel SHKH15, hardness HRC 62–64 (Fig. 1) are the deforming instruments in the rolling machine. Surface finish of the samples before and after processing was measured on a profile recorder, profile meter Abris PM-7 by the arithmetic average of the roughness profile (Ra) as the most stable of the surface profile height parameters and the mean spacing of profile irregularities (Sm).

Centrifugal Rolling of Flexible Shafts …

1081

Table 1 Experiment matrix Experiment conditions with Rainit = 5.0 µm F, H 400 600 950 1500 n 250 315 400 500 (rpm) S = 0.07 mm/rev The results of the experiment: steel 20; steel 35; Ra, 0.285; 0.255; 0.319; 0.457; µm 0.319; 0.269; 0.359; 0.437; 0.35; 0.295; 0.325; 0.41; 0.452 0.352 0.29 0.357 Sm, 37; 35; 52; 55; 59; 65; 43; 52; µm 32; 22 48; 39 57; 53 46; 63 S = 0.15 mm/rev F, H 400 600 950 1500 n 250 315 400 500 (rpm) The results of the experiment: steel 20; steel 35; Ra, 0.32; 0.276; 0.359; 0.513; µm 0.35; 0.32; 0.389; 0.45; 0.391; 0.351; 0.391; 0.451; 0.512 0.462 0.423 0.452 Sm, 46; 38; 62; 59; 67; 70; 47; 59; µm 37; 25 54; 43 67; 60 53; 66 Experiment conditions with Rainit = 2.5 µm S = 0.07 mm/rev The results of the experiment: steel 20; steel 35; Ra, 0.256; 0.224; 0.262; 0.415; µm 0.305; 0.259; 0.32; 0.345; 0.305; 0.269; 0.301; 0.357; 0.399 0.298 0.259 0.305 Sm, 33; 30; 45; 50; 49; 62; 39; 50; µm 26; 17 43; 36 51; 52 44; 62 S = 0.15 mm/rev The results of the experiment: steel 20; steel 35; Ra, 0.274; 0.251; 0.301; 0.432; µm 0.331; 0.309; 0.355; 0.385; 0.361; 0.329; 0.375; 0.425; 0.452 0.379 0.331 0.365 Sm, 45; 35; 54; 54; 60; 66; 49; 57; µm 33; 25 51; 44 60; 61 50; 66 Experiment conditions with Rainit = 1.25 µm S = 0.07 mm/rev

400 250

600 315

950 400

1500 500

S = 0.11 mm/rev steel 45; steel 60 0.27; 0.24; 0.32; 0.28; 0.36; 0.302; 0.463 0.384 35; 36; 50; 55; 33; 21 50; 36 S = 0.21 mm/rev 400 600 250 315

0.301; 0.37; 0.34; 0.347 56; 67; 60; 47

0.425; 0.43; 0.419; 0.403 37; 53; 51; 61

950 400

1500 500

steel 45; steel 60 0.35; 0.302; 0.385; 0.351; 0.423; 0.385; 0.588 0.49 55; 42; 69; 63; 40; 30 60; 48

0.39; 0.423; 0.425; 0.453 75; 73; 70; 62

0.55; 0.495; 0.475; 0.492 54; 61; 56; 71

0.245; 0.273; 0.265; 0.279 52; 63; 54; 55

0.395; 0.311; 0.335; 0.331 42; 51; 47; 64

0.32; 0.405; 0.403; 0.355 65; 71; 67; 65

0.465; 0.458; 0.45; 0.388 53; 60; 53; 70

S = 0.11 mm/rev steel 45; steel 60 0.24; 0.202; 0.251; 0.231; 0.291; 0.244; 0.412 0.318 33; 31; 47; 51; 28; 19 47; 39 S = 0.21 mm/rev steel 45; steel 60 0.303; 0.265; 0.366; 0.331; 0.397; 0.355; 0.499 0.401 51; 38; 62; 57; 35; 28 57; 49 S = 0.11 mm/rev

(continued)

1082

N. V. Vulykh

Table 1 (continued) The results of the experiment: steel 20; Ra, 0.224; 0.211; 0.263; µm 0.305; 0.261; 0.334; 0.315; 0.279; 0.288; 0.351 0.302 0.242 Sm, 28; 25; 38; 45; 40; 60; µm 23; 16 41; 36 48; 47 S = 0.15 mm/rev The results of the experiment: steel 20; Ra, 0.275; 0.242; 0.301; µm 0.329; 0.29; 0.363; 0.359; 0.323; 0.373; 0.409 0.373 0.324 Sm, 42; 29; 51; 49; 53; 66; µm 29; 22 49; 43 60; 58

steel 35; steel 45; steel 60 0.377; 0.201; 0.175; 0.357; 0.249; 0.232; 0.357; 0.301; 0.262; 0.303 0.364 0.326 32; 49; 29; 25; 43; 46; 44; 52 25; 17 44; 38 S = 0.21 mm/rev steel 35; steel 45; steel 60 0.399; 0.292; 0.267; 0.399; 0.361; 0.32; 0.412; 0.382; 0.35; 0.392 0.451 0.401 44; 54; 48; 34; 57; 53; 54; 60 33; 26 55; 49

0.241; 0.291; 0.265; 0.271 44; 62; 52; 51

0.347; 0.332; 0.312; 0.342 34; 50; 48; 54

0.326; 0.422; 0.415; 0.358 60; 69; 65; 63

0.445; 0.46; 0.455; 0.419 49; 57; 57; 65

Fig. 1 Process plan of centrifugal rolling: 1—the rolled part, 2—rolling machine loads, 3—deforming rollers, 4—double-arm lever, and 5—the stand of the centrifugal rolling machine

3 Results and Discussion As a result of the conducted experiments for the specified materials, the values of the parameters Ra and Sm were obtained and are presented in Table 1. The values of the parameters Ra and Sm are presented in the corresponding cells of the table successively in accordance with the marks of the samples. The results of the experiment (see Table 1) allowed to obtain the functions Ra ¼ f ðnÞ as well as Sm ¼ f ðnÞ:

Centrifugal Rolling of Flexible Shafts …

3.1

1083

Study of the Arithmetic Average of the Roughness Profile—Ra

In order to save space in the chapter, the diagrams in Fig. 2 study the trends in changing functions for boundary marks of steel 20 and 60 with the initial roughness Ra = 1.25 µm. The results of the research (see Table 1; Fig. 2) show that the change of Rainit from 5.0 to 1.25 µm led to a lower index of the processed surface Ra by (10–25) % except in case of the samples made of steel 35 and 45 where, when some regimes were used, the samples with the initial roughness Rainit 2.5 µm, showed almost the

Fig. 2 Impact of the rotation rate of the rolling machine on the roughness of the processed surface of a part (○—S = 0.07 mm/rev.; □—S = 0.11 mm/rev.; D—S = 0.15 mm/rev.; — S = 0.21 mm/rev.)

1084

N. V. Vulykh

same or smaller roughness of the processed surface (up to 7%) than the samples with the initial roughness Rainit 1.25 µm. In case of the samples made of steel 20, 35, and 45, the optimal processing regimes that minimized the roughness of the processed surface were the following: line feed S = 0.11 mm/rev, the rotation rate of the rolling machine n = 315 rpm. It must be noted that the samples made of steel 20 with the rotation rate of the rolling machine at n = 250 and 315 rpm had the roughness of the processed surface that varied less than in case of other samples made of other studied steel marks. The increase in the carbon mass content in the samples from *0.2 to *0.6% changed the optimum of the rotation rate of the rolling machine from 315 to 400 rpm that allowed for the minimum roughness of the processed surface. This is caused mainly by the increase in the resistance to the deforming of the material of the samples. It is also necessary to note the change in the optimal line feed from 0.11 to 0.07 mm/rev in the samples made of steel 60 (see Fig. 2). The samples made of all the studied materials with feed 0.15 and 0.21 mm/rev had the coarsest roughness of the processed surface. When the samples made of steel 20 at a rotation rate of the rolling machine of n = 250 rpm, the initial roughness was considerably crushed, and at the rotation rate n = 500 rpm, the surface layer peeled as the ductility of the material was exhausted. When the samples made of steel 35 and 45 at the maximum rotation rate of the rolling machine, the following results were obtained: At the rotation rate of n = 250 rpm, the initial roughness of the surface of the samples visually seemed not crushed enough, and at the rotation rate n = 500 rpm, the surface layer peeled. When the samples made of steel 60 at a rotation rate of the rolling machine of n = 250 rpm, the initial roughness was insufficiently crushed, and at the rotation rate n = 500 rpm, the surface layer did not peel. On the basis of the experimental data (see Table 1), a function was developed Ra ¼ f ðCÞ shown in Fig. 3 that represents the impact of the carbon mass content in the samples on the creation of minimum roughness of the surface of rolled samples using optimal processing regimes (S and n), Table 2. Fig. 3 Impact of carbon mass content on achieving minimum roughness of surface of the rolled samples

Centrifugal Rolling of Flexible Shafts … Table 2 Experimental values of the regimes of centrifugal rolling achieving minimum roughness of surface of the rolled samples

1085

Sample material

Ra, µm

Rainit, µm

Rolling regimes S, mm/ n, rev rpm

Steel Steel Steel Steel

0.175 0.231 0.244 0.242

1.25 2.5 2.5 1.25

0.11 0.11 0.11 0.07

20 35 45 60

315 315 315 400

The function (see Fig. 3) shows that the minimum roughness of the surface is achieved when the samples of steel 20 (Ra = 0.175 µm) are processed that results from the greater ductility of the material. The increase in the carbon mass content in the samples up to *0.45% (steel 45) leads to the increase in the roughness of the processed surface up to Ra = 0.244 µm. The further increase in the carbon mass content in the samples up to *0.6% (steel 60) does not result in the increase in roughness that is possibly associated with the shift in the optimal processing regimes (S and n).

3.2

The Study of the Mean Spacing of Profile Irregularities Sm

The diagrams in Fig. 4 show the trends in changing functions for boundary marks of steel 20 and 60 with the initial roughness Ra = 1.25 µm. The diagrams (see Fig. 4; Table 1) show that the change of Rainit from 5.0 to 1.25 µm led to lowering the index of the processed surface Sm down to 30%, except for the samples made of steel 60 that at the rotation rate of the rolling machine of 315 and 400 rpm had the parameter Sm increasing and decreasing by about 10% in comparison with the samples with Rainit 2.5 µm. The increase in the carbon mass content in the samples from *0.2 to *0.6% changed the optimum of the rotation rate of the rolling machine from 315–400 to 500 rpm that ensured the maximum Sm. The increase in the line feed from 0.07 to 0.21 mm/rev led to the increase in the mean spacing of profile irregularities for (15– 35) % with the greatest change of Sm happening in samples made of steel 20 that apparently results from the maximum ductility of the samples material (see Fig. 4). On the basis of the experimental data (see Fig. 4; Table 1), a function was developed Sm ¼ f ðCÞ and is presented in Fig. 5 that allows to show the impact of the carbon mass content in the samples on the creation of maximum mean spacing of profile irregularities of the rolled samples using optimal processing regimes (S and n), Table 3.

1086

N. V. Vulykh

Fig. 4 Impact of the rotation rate of the rolling machine on the mean spacing of profile irregularities (○—S = 0.07 mm/rev.; □—S = 0.11 mm/rev.; D—S = 0.15 mm/rev; —S = 0.21 mm/rev.)

The diagram (see Fig. 5) shows that the maximum mean spacing of profile irregularities (75 µm) appears when the samples made of steel 20 are processed and that is apparently associated with the minimum resistance to the deformation of the samples material and consequently leads to better spreading during the deformation process. The increase in carbon mass content in the samples (up to *0.45%) brought parameter Sm down to 70 µm (by 6.67%). The further increase in the carbon mass content in the samples (up to *0.6%) led to an insignificant increase in parameter Sm (up to 71 µm) that can result from the change in the optimal rotation rate of the rolling machine from 400 to 500 rpm.

Centrifugal Rolling of Flexible Shafts …

1087

Fig. 5 Impact of carbon mass content on forming the maximum mean spacing of profile irregularities

Table 3 Experimental values of the regimes of centrifugal rolling when achieving the maximum mean spacing of profile irregularities

Sample material

Ra, µm

Rainit, µm

Rolling regimes S, mm/ n, rev rpm

Steel Steel Steel Steel

75 73 70 71

5.0 5.0 5.0 5.0

0.21 0.21 0.21 0.21

20 35 45 60

400 400 400 500

4 Conclusion The effective technology that allows to process flexible parts, e.g., bodies of rotation is presented. Optimal regimes are set for finishing and strengthening processing of parts made of carbon steel with the carbon mass content varying from *0.2 to *0.6% that makes it possible to have processed surfaces with quite low roughness evaluated according to parameter Ra and maximum mean spacing of profile irregularities evaluated using parameter Sm.

References 1. Papshev D (1978) Finishing and strengthening of the surface by plastic deforming. Machine Construction, Moscow 2. Balter M (1978) Strengthening machine parts. Machine Construction, Moscow 3. Odintsov L (1987) Strengthening and processing of parts by surface plastic deformation. Machine Construction, Moscow 4. Vulykh NV (2017) Analysis of the stressed condition of rough layer in response to local and axially symmetrical plastic deforming. Irkutsk State Technical University Bulletin, 21(11) (130)

1088

N. V. Vulykh

5. Zaides SA, Kuang, LH, Kyong NK (2018) The quality evaluation of the strengthened surface plastic deforming by rollers with various constructions Irkutsk State Technical University Bulletin 22(1) 6. Vulykh NV (2019) Microprofile model form changing research at axisymmetric deformation with account of scale factor. In: Radionov A, Kravchenko O, Guzeev V, Rozhdestvenskiy Y (eds) Proceedings of the 4th international conference on industrial engineering. ICIE 2018. Lecture Notes in Mechanical Engineering, Springer, Cham, pp 1161–1168. https://doi.org/10. 1007/978-3-319-95630-5_ 7. Zaides S, Vulykh N (2013) Covering Hardening of Low-Stiffness Shafts. Theory, technology. Monograph. Saarbruecken, Germany 8. Vylykh NV (2002) Formation of Microgeometry of Part Hardened Layer at Local and Covering Plastic Surface Deformation. Dissertation, University of Irkutsk 9. Vulykh NV, Bubnov AS, Garmazov IY (2003) Experimental observation of deformed state of part surface layer at covering plastic surface deformation. Paper presented at the Modern Problems of Machine Construction and Transport, University of Ulyanovsk, 8–10 October 2003 10. Vylykh NV (2004) Determining a deformed state of the part peripheral layer at plastic surface deformation. Paper presented at the Advanced Technologies of Material Obtaining and Treatment, University of Irkutsk, 20–21 April 2004 11. Zaides, SA, Gorbunov, AV (2016) Improvement of low-rigidity shafts by centrifugal rolling. Russ Eng Res 36 12. Zaides, SA, Van Khuan N (2017) Influence of parameters of the calibration process on bending stiffness of steel rod. Part 1. Determination of residual stresses in the calibrated rod. In Proceedings of higher educational institutions. Ferrous Metallurgy 13. Zaides SA, Zhuravlev DA, Kurguzov SA. (1997) Device for strengthening the surface of cylindric parts. Patent RF 96,105,784, 27. Aug 1997 14. Vulykh NV, Gorbunov AV (2012) Centrifugal rolling of flexible shafts to achieve minimum roughness and maximum surface loading capacity. Irkutsk State Technical University Bulletin 10(69)

Thread Milling Cutter Flute Production Possibility Research by Using Typical Profiles Grinding Wheels O. V. Malkov and I. A. Pavlyuchenkov

Abstract The research of the technological possibility of producing comb-like thread milling cutters flutes on the basis of their mathematical and experimental modeling and the study of the thread mill prototype production possibility by using a typical profile grinding wheels was conducted. The analysis of grinding wheel types used for manufacture various flutes profiles was carried out, as a result of which the possibility of using a straight and conical profile grinding wheels for forming thread mill flutes was established. Mathematical dependencies determining the flute end section profile made by the grinding wheel have been derived, taking into account the kinematic scheme of the five-coordinate grinding and sharpening CNC machines. The proposed mathematical equations are confirmed by experimental studies in the processing of helical flutes on bulk polyamide blanks with a diameter of 34.5 mm on a universal sharpening machine. The flute profiles end section comparison (experimental and calculated) revealed good convergence. The thread milling cutter CoroMill Plura R217.15-140100AC26 N was selected as a prototype for the study of thread milling cutters of flute producing possibility. It was found that the grinding wheel of type 1A1 is not suitable for making prototype flutes, though it is possible to select positioning parameters so that a divergence of the flutes end section (prototype and design) is in the range of 3 … 63 micrometers, which reveals the prototype thread milling cutter flutes technological production possibility by using grinding wheels with typical profiles. Keywords Thread milling cutter grinding End section profile





Flute profiling
Typical wheel profile
Flute

O. V. Malkov
I. A. Pavlyuchenkov (&) BMSTU, 2-ya Baumanskaya ul., 5/1, Moscow 105005, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_116

1089

1090

O. V. Malkov and I. A. Pavlyuchenkov

1 Introduction A thread is one of the most common machine-building producing parts. There are various ways of a thread production and improvements of production performance properties, some of which are described in publications [1–5]. Thread milling has become more popular in the last ten years, due to the thread milling process being more flexible, universal and well-realized on the CNC machines [6–13]. The current technology of thread mill production utilizes grinding on the CNC tool grinding machines (ANCA, WALTER AG, ROLLOMATIC, etc.) in one operation by several typical profiles grinding wheels. The flute surface shape produced on the tool grinding machines with NC depends generally on shape and dimension of the grinding wheel, tool parameters and grinding wheel positioning [14–22], changing which one can produce a tool with a flute profile as accurate as possible to the required one. The goal of this article is to research thread milling cutter flutes production possibilities based on their mathematical and experimental modeling and research of thread milling cutter flutes production possibilities by using typical profiles of grinding wheels.

2 Methods The solid thread milling cutter CoroMill Plura R217.15-140100AC26 N designed for making internal metric threads with 6 h precision has been chosen for research as a prototype within a collaborative agreement between the chair “Tool engineering and technologies” (BMSTU n/a Bauman) and LLC “Sandvik” training center. The thread milling cutter end section traversing through the top of the tooth is shown in Fig. 1. The result profile was scanned and dimensioned for the further research. Table 1 presents the thread milling cutter main parameters according to the manufacturer’s catalog; Table 2 presents the end section profile parameters measured dimensions. The helical flutes of the solid carbide tool are made typically by using grinding wheel type 1A1 (for the opened end section flute profile) or type 1V1 (for the closed end section flute profiles) according to grinding wheels manufacturers (Norton, Pferd, and Isham). The calculation scheme of solid end milling cutters chip flutes grinding on CNC was developed for thread milling cutter flutes production possibility research (finding the range of using grinding wheel geometrical parameters and parameters of its positioning) shown in Fig. 2. The last one produces the ability to create an end section flute profile mathematical model. The generalized shape of the grinding wheel specified by a parameter up is shown in Fig. 2, but mathematical dependencies allow for calculations to do arbitrary grinding wheel shape.

Thread Milling Cutter Flute Production …

1091

Fig. 1 CoroMill Plura R217.15-140100AC26N end section (1—cross section of the milling cutter, flute end section profile approximated by 3 arcs)

Table 1 CoroMill Plura R217.15-140100AC26N main parameters Thread type

Thread pitch P, mm

Cutting length lp, mm

Cutting diameter dp, mm

Number of cutting tooth, z

Flute helix angle x, °

Internal

1

26

14

5

10

Table 2 CoroMill Plura R217.15-140100AC26N end section parameters Flute profile

H (º)

Hк (º)

HZ (º)

rп (mm)

rв (mm)

rcп, (mm)

d c, (mm)

dp, (mm)

1 2 3 4 5 Average value

71.5 71.8 71.1 72.3 73.1 71.96

43.7 43.5 42.8 44.4 42.5º 43.38

27.9 28.4 28.4 27.9 30.6 28.64

4.53 3.94 5.44 4.93 5.76 4.92

1.58 1.65 1.61 1.54 1.61 1.60

23.89 25.04 32.15 19.61 23.66 24.87

7.56 7.78 7.73 7.37 7.34 7.56

13.99 13.63 13.99 13.74 13.72 13.81

The grinding wheel in rotated and offset position relative to base point O0w on the forming cylindrical helix is shown in Fig. 2. Dr—primary motion, DðS1 Þ —grinding wheel feed motion, DðS2 Þ —grinding wheel circular feed, d—thread milling cutter blank diameter, Di ðRi Þ—diameter (radius) i’s grinding wheel elementary disk, Dmax—grinding wheel maximum diameter, Ci—offset of end of grinding wheel relative to i’s elementary disk, C(max)—wheel thickness, A—offset of grinding wheel axis relative to cylindrical blank axis, N and M—offset of grinding wheel coordinate system Xw, Yw and Zw relative to base point O0w on the generating helix

1092

O. V. Malkov and I. A. Pavlyuchenkov

Fig. 2 A grinding wheel positioning calculation scheme (1—grinding wheel, 2—cylindrical blank, 3—forming helix)

along Xw and Zw axis accordingly, uy—rotation angle of grinding wheel end relative to blank axis, x—inclination angle of generating helix to its axis, wj—current (j) parameter of i’s grinding wheel elementary disk in Xw, Yw, Zw Xw0 , Yw0 and Zw0 — coordinate system offset on M and N relative to Xw, Yw, and Zw; Xw00 , Yw00 and Zw00 — coordinate system offset on uy angle relative to Xw0 , Yw0 and Zw0 , XYZ—blank coordinate system. Affine transformations in space are used to determine grinding wheel coordinates in the blank coordinate system XYZ with respect to offset and

Thread Milling Cutter Flute Production …

1093

rotation of coordinate systems. The equation system describing positions of each grinding wheel peripherical point is shown in Eq. 1: 
 8
    > X ¼ ðRi cos wj þ NÞ sin uy þ ðCi þ M Þ cos uy cos 2dz dtan x þ Ri sin wj þ A sin 2dzdtan x > > 
 <
    Y ¼  ðRi cos wj þ NÞ sin uy þ ðCi þ M Þ cos uy sin 2dz dtan x þ Ri sin wj þ A cos 2dz dtan x >   > > ðC þ M Þ cos uy N cos uy dz : wj ¼ arccos i Ri cos u y

ð1Þ The experiment of grinding cylindrical blank (d = 34.5 mm, length 300 mm) with grinding wheel (Dmax = 151.6 mm, Cmax = 20.5 mm, type 1A1) was done to verify the mathematical model. A rod of bulk polyamide (PA-6-MG, TC 2224-036-00203803-2012) with diameter 40 mm and length 1000 mm was chosen. An angle of helix (x) is 10°; center distance (A) is 84.4 mm. Three flutes were produced on the blank. The research was conducted on the universal grinding machine 3D642E with the use of a tool with helical flutes sharpening device 3E642E. Varied parameters are offset N and M, grinding wheel positioning angle uy (Table 3). The blanks were cut in a plane perpendicular to the axis so that the grinding wheel penetration surface does not fall into the section. The end section was scanned and the outline of the profile was encircled with fine pitch spline in KOMPAS-3D. The comparison of experimental and calculated helical flute end section profiles (M = 0 mm, N = 5 mm, uy = 10°) is shown in Fig. 3 and calculated with the use of reference surfaces drawn in 1 mm increment from the outer diameter. The distance between the points obtained at the intersection of each profile and each reference circle was measured. Experimental research confirmed the presented mathematical dependences by the good convergence of profiles for all values of the variable parameters (Table 3), and gap (up to 0.42 mm) revealed as a result of comparison can be explained by the grinding wheel positioning inaccuracy relative to the workpiece axis, by the workpiece flexibility, by the experimental parameters settings inaccuracy, and by an increase of the inaccuracy of measurement direction coincidence approaching to the core.

Table 3 Varied parameters

Parameter

Flute index 1 2

3

Offset N (mm) Offset M (mm) Positioning angle uy (°)

0 5 10

5 5 30

5 0 10

1094

O. V. Malkov and I. A. Pavlyuchenkov

Fig. 3 Experimental (1) and calculated (2) helical flute end section profiles comparison (M = 0 mm, N = 5 mm, uy = 10°) and the end section of the block polyamide blank (3)

3 Results The experiment results conducted with using a grinding wheel type 1A1 confirmed the mathematical dependences, Eq. 1. Taking into account, the presented mathematical dependences can be used for grinding wheel of any shape. We use the model to select the grinding wheel shape, size, and positioning to analyze the thread milling cutter flute production possibility on the example of the selected prototype CoroMill Plura R217.15-140100AC26 N. The program “Profiling of a solid thread milling cutter flute” was developed at the BMSTU chair “Tool engineering and technologies” based on the obtained dependencies (Eq. 1). It was used to select the grinding wheel parameters and the setting parameters relative to the CoroMill Plura R217.15-140100AC26 N thread milling cutter blank. Preliminary calculations have shown that the grinding wheel type 1A1 is not suitable for thread milling cutter flute production because the calculated value Hк (Fig. 1) turns out more than required, so the selection of grinding wheel type 1V1 parameter, which allows the production of high accuracy prototype flutes, is shown below. Initial data: Norton grinding wheel (type 1V1, Dmax = 125 mm, Cmax = 10 mm, Up = 30°), the parameters of the end section flute profile correspond to those specified in Table 2. The most accurate flute is obtained at M = −12 mm, N = 0 mm, A = 66.06 mm, uy 15.9°. The results of comparing the mill prototype end section flute profile and the calculated profile are presented in Fig. 4. The comparison was carried out using support circles with diameters from 14 to 8 mm with a 1 mm increment starting from the outer diameter. The distances between the intersection points of these circles with the real mill profile and the profile obtained because mathematical modeling is measured. The maximum

Thread Milling Cutter Flute Production …

1095

Fig. 4 Calculated (1) and prototype (CoroMill Plura high feed cutters R217.15-140100AC26 N) (2) helical flute end section comparison

distance was 63 µm, and the minimum deviation was 3 µm. The maximum value in the cavity can be explained by the fact that the measurements along the profiles are significantly greater than the normal distance between them as observed on the outer diameter.

4 Conclusion The results show that the use of the grinding wheel 1V1 with the accepted values of the shape parameters and positioning allows the creation of the desired flute shape with high accuracy and confirms the functionality of the mathematical model for calculating the profile parameters of the projected thread mills end section.

References 1. Evsyukov SA, Nebogov SM, Punin VI, Fedotov IL (2016) Surface plastic deformation of threads in ultrasound treatment. Russ Eng Res 36(8):620–625. https://doi.org/10.3103/ S1068798X16080086 2. Evsyukov S, Nebogov S, Fedotov I (2016) Pipe thread wear-resistant ultrasonic hardening unit. Vibroengineering PROCEDIA 8:142–146. https://www.jvejournals.com/article/17692/ pdf 3. Fedorova LV, Fedorov SK, Ivanova YS, Voronina MV (2017) Increase of wear resistance of the drill pipe thread connection by electromechanical surface hardening 12(18):7485–7489 4. Gao H, Lu S, Yang A, Bao Y (2017) A methodology for helical mill-grinding of tiny internal threads made of hard brittle materials. Int J Adv Manuf Technol 91(1–4):25–37. https://doi. org/10.1007/s00170-016-9727-5

1096

O. V. Malkov and I. A. Pavlyuchenkov

5. Morgunov VA, Nebogov SM, Fedotov IL (2018, 04) Elevation of the Wear Resistance of Threads of Tubing Strings Under the Action of Ultrasound. Metallurgist 61(11–12):1108– 1114. https://doi.org/10.1007/s11015-018-0613-2 6. Fromentin G, Poulachon G (2010, 07) Geometrical analysis of thread milling—part 1: evaluation of tool angles. Int J Adv Manuf Technol 49(1–4):73–80. https://doi.org/10.1007/ s00170-009-2402-3 7. Fromentin G, Poulachon G (2010) Geometrical analysis of thread milling—part 2: calculation of uncut chip thickness. Int J Adv Manuf Technol 49(1–4):81–87. https://doi.org/10.1007/ s00170-009-2401-4 8. Glushko EV (2008, 07) Thread milling by the engagement method. Russ Eng Res 28(7):720– 722. https://doi.org/10.3103/s1068798x08070204 9. Danilenko BD (2015, 01) Cutting conditions for thread mills. Russian Engineering Research 35(1):76–77. https://doi.org/10.3103/s1068798x15010098 10. Kirichek AV, Afonin AN (n.d.) Stress-strain state of the thread-milling tool and blank. Russ Eng Res 27(10):715–718. https://doi.org/10.3103/s1068798x07100152 11. Kosarev VA, Grechishnikov VA, Kosarev DV (2009) Milling internal thread with planetary tool motion. Russ Eng Res 29(11):1177–1179. https://doi.org/10.3103/S1068798X09110227 12. Lee SW, Nestler A (2012) Simulation-aided design of thread milling cutter. Procedia CIRP 1:120–125. https://doi.org/10.1016/j.procir.2012.04.019 13. Mal’kov OV (2013). Precision of the external-thread profile in thread cutting. Russ Eng Res 33(3):172–175 14. Balandin AD, Danilenko BD (2008) Margin in sharpening helical channels of tool. Russ Eng Res 28(8):814–815. https://doi.org/10.3103/S1068798X08080200 15. Balandin AD, Danilenko BD (2013) Producing helical channels on taps by means of face mills. Russ Eng Res 33(6):355–357. https://doi.org/10.3103/S1068798X1306004X 16. Chen Z, Ji W, He G, Liu X, Wang L, Rong Y (2018) Iteration based calculation of position and orientation of grinding wheel for solid cutting tool flute grinding. J Manuf Process 36:209–215. https://doi.org/10.1016/j.jmapro.2018.10.012 17. Ivanov V, Nankov G, Kirov V (1998) CAD orientated mathematical model for determination of profile helical surfaces. Int J Mach Tools Manuf 38(8):1001–1015. https://doi.org/10.1016/ S0890-6955(98)00002-9 18. Kaldor S, Rafael AM, Messinger D (1988) On the CAD of profiles for cutters and helical flutes—geometrical aspects. CIRP Ann 37(1):53–56. https://doi.org/10.1016/S0007-8506(07) 61584-4 19. Karpuschewski B, Jandecka K, Mourek D (2011) Automatic search for wheel position in flute grinding of cutting tools. CIRP Ann 60(1):347–350. https://doi.org/10.1016/j.cirp.2011.03. 113 20. Petukhov YE, Movsesyan AV (2007) Determining the shape of the back surface of disc milling cutter for machining a contoured surface. Russ Eng Res 27(8):519–521. https://doi. org/10.3103/S1068798X07080084 21. Zhao XF, He L, Shi HY (2013) Research on the Mathematical Model of Helical Groove of the End Mill Based on the Grinding Wheel Attitude 589–590:416–420. https://doi.org/10.4028/ www.scientific.net/KEM.589-590.416 22. Zhidyaev AN (2018) Impact of grinding wheel position on flute profile of end mill and cutting process. In IOP Conference Series: Materials Science and Engineering vol 302, no 1. https:// doi.org/10.1088/1757-899x/302/1/012069

Thermo-Emf as Method for Testing Properties of Replaceable Contact Pairs Z. Tikhonova, D. Kraynev and E. Frolov

Abstract In the paper on the basis of the technical literature analysis, reference and regulatory sources and experimental studies, it was found that there is one common circumstance as a reason for the significant discrepancy between the calculated and actual values of cutting speed, cutting force components, and the value of the roughness parameter in the development of turning operation for carbon, structural, and corrosion-resistant steel. It is the probabilistic nature of the formation of contact pairs “carbide tool–steel workpiece”, the absence of mathematical models for real-time accounting of their thermophysical properties, and the orientation of the calculation models on the average value of these properties from the entire permissible scale of variation. From the standpoint of high-speed plastic deformation of machined steels, the authors have proposed some cutting modes for a preliminary test run as a test method for obtaining operational information about the variable thermophysical properties of contact pairs “carbide tool–steel workpiece” directly on the cutting machine.




Keywords Accuracy of calculation Steel hardening Strain rate Thermo-emf CNC systems








Degree of deformation


1 Introduction One of the promising directions in the creation of new generations CNC systems for lathes is to endow them with “technological intelligence”, i.e., ability to solve technological problems of management. Modern CNC systems are “taught” to solve complex geometric control tasks (surface shaping), but they do not use any software in solving technological problems yet.

Z. Tikhonova (&)
D. Kraynev
E. Frolov VSTU, 28, Lenin ave, Volgograd 400005, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_117

1097

1098

Z. Tikhonova et al.

At first glance, the situation seems paradoxical. In spite of the increased capabilities of computing and control technology, modern CNC systems of the PC-NC class with a huge memory capacity cannot solve the problem connected with the reliable section of the main parameters in the cutting process [1–12]. First of all, it concerns the choice of the exact value of the cutting speed, which is designed to provide the specified time of reliable tool operation or exact calculation of the cutting force. Using the magnitude of cutting force, we can determine cutting power, torque, clamping force of the workpiece, and a number of other important parameters in the technological process, such as accuracy machining and quality of the treated surface. The problem lies not in the capabilities of the control technology, but in the adequacy of mathematical models of the parameters during the cutting process. These parameters form the basis of their calculation algorithms programmatically (automatically).

2 Relevance In the work [1, 13], on the basis of the technical literature analysis, reference and regulatory sources and experimental studies, it was found that there is one common circumstance as a reason for the significant discrepancy between the calculated and actual values of cutting speed, cutting force components, and the value of the roughness parameter in the development of turning operation for carbon, structural, and corrosion-resistant steel. It is the probabilistic nature of the formation of contact pairs “carbide tool–steel workpiece”, the absence of mathematical models for real-time accounting of their thermophysical properties, and the orientation of the calculation models on the average value of these properties from the entire permissible scale of variation [2, 11, 14]. For example, in the formulas for determining the cutting speed with a carbide tool, there is an adjustment made for the grade of tool material, but on the condition that its properties within the grade composition are the same. We can see a contradiction between the technical specifications for sintering the carbide blades, which allow for a twofold variation of their cutting properties and the method of calculation. In the formulas for calculating cutting force components, there is no adjustment for the grade of tool material. While calculating the adjustment coefficient for the grade of steels being machined (Cg), the existing techniques use their strength properties rtemp (temporary breaking strength). The value rtemp is obtained either by selective mechanical testing of steel workpiece from the delivery lot, or they take a certain value for each grade from reference books. However, the hardened condition of steel, which it receives during the cutting process and which determines the resistance value to plastic deformation, and, therefore, the magnitude of the cutting force components, does not have an unambiguous connection with the characteristics obtained as a result of mechanical tests at room temperature. This circumstance determines the relevance of this work.

Thermo-Emf as Method for Testing Properties of Replaceable …

1099

3 Problem Statement Thus, based on the above-mentioned reasons for the discrepancy between the calculated and actual parameters of the cutting process, we can formulate the problem: to propose and justify a way to eliminate these shortcomings.

4 Theoretical Part To solve the problem, it is proposed to obtain the data on the thermophysical properties of the contact pairs and the cutting conditions before performing the test run at strictly defined processing modes. It is also necessary to measure the magnitude of the generated thermo-emf of a natural thermocouple, which, due to the physical nature of the thermoelectric phenomena occurrence while cutting, indirectly characterizes these properties [2, 15]. As a methodological basis for the proposed method of obtaining preliminary information about the thermophysical properties of contact pairs directly on the machine while setting it up using the program, it was decided to take the Brinell method for determining steel hardness invented in 1900. Brinell offered to press balls into the surface of the samples of the steels to be tested to evaluate their hardness. The balls were of the fixed diameter and load, and it should have been done in diameter of the incenter, which depends on steel properties. In this method, the hardness of steels (alloys) is averaged since the steels contain phases of different hardness (ferrite, cementite, perlite, austenite, etc.), but it promptly reflects its properties. The value of the steel hardness according to Brinell HB (GOST 9012-59) is used in mathematical models for calculating the parameters of the cutting process when determining the corrections for the properties of the processed steel. The hardness of most carbon and low-alloyed steels is related to their thermal conductivity by an inverse relation: the higher the hardness is, the lower the thermal conductivity is, and the “paradox” is explained from the standpoint of the actual cutting scheme [3, 16]: the cutting force is more important while machining the steel with lower hardness than steel with the higher one. The thermo-emf value of the test cut has a clear link with the thermal conductivity of steel. But, to ensure the calculation accuracy, it is necessary to have some information about the properties of the contact pair. This information is given by the thermo-emf value of the preliminary test cut, the conditions of which are strictly the same for all test contact pairs, as the conditions for pressing the ball into steel samples in the Brinell method. In this case, the choice of cutting modes for a test cut plays an important role. The conditions of the test run mode should have a clear link with the conditions for steel hardening in the applicable operating mode ranges for semi-finishing and finishing turning, and they must also ensure the possibility of accurate measurement of thermo-emf.

1100

Z. Tikhonova et al.

What are these conditions? It is commonly known that carbon and alloy steels are hardened differently in the process of high-speed plastic deformation [3]. There is a level of strain rates, after which the strain hardening stops and thermal softening occurs. According to Reht [4] these are “critical strain rates”, which he proposed to define by the expression:
@s=@e Kn2
e_ jp ¼ 4
p
k
q
Cm
e  ey
; @s=@h s2y
L2

ð1Þ

where k is the thermal conductivity; q is the density; Ct is the specific heat; ey and e are the initial and final shear deformation; Kn is the mechanical equivalent of heat; L is the size of the deformation region;sy is the relationship between the initial yield @s stress and shear; @s @e, @h are partial derivatives (determined experimentally). For carbon steels, Recht [4] defined critical strain rates e_ jp = 103–104 1/s. From Eq. (1), it follows that steels with high thermal conductivity correspond to higher critical strain rates. As for carbon steels, steel 20 has high thermal conductivity. In relation to other steel types, when the carbon content increases, the thermal conductivity decreases. Alloy steels have a lowered thermal conductivity. It is important that the test cut mode provides the presence of a critical strain rate for the whole range of steels being processed. When carrying out the cutting process, the thermal conductivity of the cutting tool will influence the critical strain rate of the steel. This impact is evident through the size of the chip formation zone—the area of steel hardening C2 [3]. Table 1 presents the results of measurement and calculation of the contact characteristics during the cutting process of steel 20 with hard alloy T15K6 and VK8. The calculation of the average strain rate in the chip formation zone was made using the following equation: e_ c ¼

Vpe3
½cos b þ ðsin b  cÞ
tg b ; C2

ð2Þ

The equation was derived analytically, and it is based on the kinematic laws of the chip formation process [4]. The calculation of the strain rate in the zone of contact plastic deformation at the end of the hardening section C2 was made according to the formula: e_ j ¼

Vpe3
tg b hj

ð3Þ

where Vcut is the cutting speed, m/s; hc is the height of the zone of contact plastic deformation at the end of the hardening section which is equal to 0.1 of chip thickness ac. It should be noted that the values e_ j in Table 1 for the hardening areas are not finite and obtained by calculation. The strain rate in the area of plastic contact increases during the movement of the strained volume of the metal as the height of

Section C2 (mm)

Chip shrinkage n

Chip thickness ac (mm)

Shear angle b (degrees)

(a) steel 20—T15К6 (S = 0.3 mm/rev; t = 2 mm; u = 45°; c = 0°; k = 0°) 60 (1) 0.27 3.84 0.9 15 90 (1.5) 0.14 3.56 0.75 15.7 110 (1.83) 0.12 3.15 0.69 17.6 130 (2.1) 0.12 3.10 0.67 17.8 150 (2.5) 0.12 2.94 0.65 18.7 (b) steel 20 —VK8 (S = 0.3 mm/rev; t = 2 mm; u = 45°; c = 0°; k = 0°) 60 (1) no 3.94 – – 75 (1.25) no 3.86 0.27 14.5 90 (1.5) 0.13 3.72 0.26 15.0 105 (1.75) 0.12 3.51 0.26 15.9 120 (2.0) 0.12 3.33 0.25 16.7 135 (2.25) 0.09 3.27 0.25 17.0 150 (2.5) 0.07 3.18 0.24 17.4

Cutting speed V [m/min (m/ s)] 2.96  103 5.62  103 8.4  103 1.06  104 1.3  104 – – 1.54  104 1.91  104 2.4  104 2.7  104 3.2  104

103 103 103 103 103

– – 7.95  103 8.6  103 8.7  103 1.16  104 1.49  104 4.2 6.2 8.8 10.6 11.6 12.0 12.6

Strain rate e_ j , 1/s in the zone of contact plastic deformation

5.17 1.11 1.59 1.82 2.25

    

Strain rate e_ c , 1/s in the chip formation zone

6.0 8.6 9.2 10.3 10.5

Thermo-emf of the test cut E (mV)

Table 1 Contact characteristics of the pair steel 20—T15K6 (a) and the pair steel 20—VK8 (b)

Thermo-Emf as Method for Testing Properties of Replaceable … 1101

1102

Z. Tikhonova et al.

the zone hc decreases. At the end of the plastic contact, hc takes values of 1–2 µm. e_ j reaches the values of 105–106 1/s. The calculation of the resistance to plastic deformation in the chip formation zone and in the zone of contact plastic deformations was performed using the microhardness value measured on the chip roots [3, 8]. The results of the strain rate calculation (Table 1a, b) show that for the cutting conditions studied, the lower limit of the critical strain rate for the pair steel 20—T15K6 starts with cutting speeds of 60 m/min (Table 1a). At a cutting speed of 90 m/min, the strain rates reach 104 1/s. (Table 1a). Using this information, we can come to the conclusion that the level of cutting speed of a test cut as a preliminary test for the hardenability of steels must ensure critical strain rates. Table 1b presents the results of measurement and calculation of the contact characteristics of the pair steel 20—VK8. This pair is made up of materials with the highest thermal conductivity coefficient (steel 20—k = 52 W/m °C; VK8—k = 52 W/m °C). If, with these thermophysical characteristics of the pair, the speed range of 60–130 m/min (1–2.1 m/s) ensures the presence of critical strain rates, it will also provide them for other contact pairs made up of steel and solid alloys grades which are widely used in mechanical engineering. This conclusion is based on the fact that machined steels with high thermal conductivity have an increased area of hardening C2 and an increased height (volume) of the zone of contact plastic deformations hc, which was experimentally confirmed in [3, 10, 17]. At constant cutting speed of the test cut, these contact characteristics, according to formulas (2) and (3), determine the strain rates e_ c and e_ j . For carbon steels, when carbon content increases, thermal conductivity decreases. As it was mentioned above, alloy steels have a lower thermal conductivity than steel 20. Of the entire range of carbide tools used for steel machining, VK alloys have the highest thermal conductivity, which also contributes to C2 and hc. According to the experimental measurements of the contact characteristics for the pair steel 20-VK8 and the calculated values of the strain rates (Table 1b), it may be concluded that the critical strain rates of steel 20 are observed when treated with VK8 alloy during the test cut (S = 0.1 mm/rev, t = 1 mm) starting at a speed of 90 m/min. At speeds above 90 m/min, the onset of critical strain rates is guaranteed. Formula (1) indicates the fact that if we consider the geometry of the tool, we can see that rake angle c affects the magnitude of the strain rate. Is the change of the rake angle c by the value of the thermo-emf of the test cut taken into account? Is the thermo-emf connection with the deformation conditions of the cut steel volumes unambiguous if we consider the range of cutting speeds above 90 m/min? In [3, 18], it is shown that the strain rate e_ c and e_ j in the range of variation of the cutting speed equal to 90–120 m/min and rake angle c from −10° to +10° varies from 103 to 104 1/s, which corresponds to the critical strain rates of steel. In [5, 6], it was experimentally shown that the thermo-emf value of the test cut is arcwisely connected with the change of the rake angle c.

Thermo-Emf as Method for Testing Properties of Replaceable …

1103

It should be noted that the quantitative values of the resistance to plastic deformation, the strain rate, and the degree of deformation depend both on the thermophysical properties of the steels being machined and on the processing modes: feed and speed. With the same processing modes of the test cut of various steels, the level of characteristics that determine their hardened state during the cutting process is different, but the nature of their change is the same.

5 Practical Significance Test run conditions are a kind of testing for a contact pair workpiece—hard alloy. The results of this testing provide more complete and accurate information in comparison with the data about the properties of steel and hard alloys given in references. On the basis of the calculated and experimental data on the effect of cutting speed on the hardenability characteristics of steels during the test cut, it is proposed to use cutting speed of 100 m/min. The depth is assumed to be equal to one millimeter for the reasons of the minimum allowance for finishing (semi-finishing) machining. The feed is assumed to be 0.1 mm/rev. The cutting time during the trial cut equals to 4–5 s. We can observe a stable adhesive bond on the front surface of the tool (cutter) within the plastic contact when machining steels with a carbide tool in the accepted test run modes. The presence of this bond ensures reliable measurement of the thermo-emf of a natural thermocouple [19, 20]. At lower cutting speeds, for example, cutting with build-up forming, thermo-emf signal is unstable, with significant amplitude fluctuations.

6 Conclusion Based on the study results of the informative value of the thermo-emf signal for a natural thermocouple, it is proposed to use its value, measured in the justified mode of preliminary test cut at V = 100 m/min, S = 0.1 mm/rev, t = 1 mm in mathematical relationships for more accurate (automated) determination of the parameters in the cutting process by the NC system itself.

References 1. Plotnikov AL, Sergeyev AS, Uvarova TV (2017) Avtomatizirovannyye sposoby opredeleniya znacheniy parametrov protsessa tokarnoy i frezernoy obrabotki na stankakh s CHPU (Automated ways of determining the parameter values of turning and milling processes on CNC machines). Tonkiye naukoyomkiye tekhnologii (TNT) (Thin science based technologies (TNT)), Volgograd, Staryy Oskol

1104

Z. Tikhonova et al.

2. Plotnikov AL, Sergeyev AS, Zaytseva NG (2013) Ispol’zovaniye informativnoy sposobnosti signala yestestvennoy termopary dlya obespecheniya nadozhnosti avtomatizirovannogo opredeleniya rezhimov lezviynoy obrabotki (Using the informative ability of the signal of a natural thermocouple to ensure the reliability of automated determination of the modes of blade processing). Naukoyomkiye tekhnologii v mashinostroyenii (High Technol Mech Eng) 5:35–40 3. Talantov NV (1992) Phizicheskie osnovy protsessa rezaniya, iznashivaniya i razrusheniya instrumenta (Phisical basics of the tool’s cutting, wear and break down). Mashinostroenie Publ, Moscow 4. Rekht RF (1964) Razrushayushchiy termoplasticheskiy sdvig (Destroying thermoplastic shift). Proc Am Soc Mech Eng 31(2):34–39 5. Plotnikov AL, Sergeyev AS, Zaytseva NG (2015) Povysheniye nadozhnosti upravleniya sherokhovatost’yu obrabotannoy poverkhnosti detaley v SAPR TP tokarnykh i frezernykh operatsiy (Improving the reliability of managing the roughness of the machined surface of parts in a CAD system for turning and milling operations). VolgGTU, ZAO ONIKS, Volgograd, Tol’yatti 6. Krainev DV, Sergeev AS, Tikhonova ZhS, Ngo KC (2017) The reliability improvement of CNC machining centers. In: Proceedings of 2017 20th IEEE international conference on soft computing and measurements, SCM 2017, pp 627–629. https://doi.org/10.1109/CM.2017. 7970670 7. Krainev DV, Polyamchikova MYu, Bondarev AA (2017) Influence of the surface layer characteristics on the regularities of the cutting process. In: MATEC web of conferences. Available via DIALOG. https://www.matec-conferences.org/articles/matecconf/abs/2017/43/ matecconf_icmtmte2017_01045/matecconf_icmtmte2017_01045.html. Accessed 7 Nov 2017 8. Frolov Ye M (2009) Development of a mathematical model for the automated calculation of the cutting speed for thin-turning conditions. Proc VSTU 8:108–110 9. Starkov VK (2009) Physics and optimization of metals. Mechanical engineering, Moscow 10. Zaytseva NG, Ingemansson AR, Kraynev DV et al (2012) Problems of quality management while processing and modeling of the formation process of surface roughness in turning. News VSTU 13(100):15–18 11. Plotnikov AL, Sergeev AS, Tikhonova ZhS (2016) Raschet nyagovykh usilij privodov stanochnych prisposoblenij pri zakreplenii i tochenii korrozionno-stojkikh stalei (Thrust force calculation of machine tool actuators in fastening and turning corrosion resistant steels). Certificate of the state registration of the software program no. 2016618188, 22 Jul 2016 12. Plotnikov AL, Sergeev AS, Tikhonova ZhS (2016) Osobennosti ispolzovaniya signala EDS rezaniya v usloviyakh avtomatizirovannogo stanochnogo proizvodstva (Pecularities of the use of cutting EMF signal in automated machine tool manufacturing). In: Naukoemkie tekhnologii v mashinostroenii. Science based technologies in machine building 6:21–27 13. Plotnikov AL, Krylov EG, Frolov EM (2010) Diagnostics of the state of a multicutter hard-alloy tool on the basis of thermoelectric phenomena in the cutting zone. Russ Eng Res 30 (2):161–165 14. Sergeev AS et al (2017) Method for measuring thermo-EMF of a «tool-workpiece» natural thermocouple in chip forming machining. In: MATEC Web of Conferences. Available via DIALOG. https://www.matec-conferences.org/articles/matecconf/pdf/2017/43/matecconf_ icmtmte2017_01044.pdf. Accessed 7 Nov 2017 15. Sergeev AS, Tikhonova ZhS, Krainev DV (2017) Automated thrust force calculation of machine tool actuators in fastening and turning steels. In: Procedia engineering. International conference on industrial engineering (ICIE 2017), Available via DIALOG. http://www. sciencedirect.com/science/article/pii/S1877705817352943. Accessed 8 Nov 2017 16. Krainev DV, Sergeev AS, Tikhonova ZhS, Ngo KC (2017) The reliability improvement of CNC machining centers due to on-line diagnostics of the cutting process In: XX IEEE international conference on soft computing and measurements (SCM). St. Petersburg, 24–26 May 2017, pp 627–629. https://doi.org/10.1109/scm.2017.7970670

Thermo-Emf as Method for Testing Properties of Replaceable …

1105

17. Polyanchikov YuN, Krainev DV, Norchenko PA et al (2011) Improved cutting of steels by means of preceding plastic deformation. Russ Eng Res 31:82–84. https://doi.org/10.3103/ S1068798X11010187 18. Plotnikov AL, Tikhonova ZhS, Eplov PE et al (2017) Physical principles of the thermo EMF of natural thermocouple for rapid evaluation of properties of the “high-speed tool—steel workpiece” contact pairs. Izvestia VSTU 12(207):79–83 19. Tikhonova, Z, Kraiynev D, Frolov E (2018) Efficiency improvement for assigning of cutting conditions on the basis of the thermo-EMF signal. In: MATEC Web of Conferences. Available via DIALOG. https://www.matec-conferences.org/articles/matecconf/abs/2018/83/ matecconf_icmtmte2018_01067/matecconf_icmtmte2018_01067.html. Accessed 30 Oct 2018 20. Kraynev D, Tikhonova, Z, Bondarev A (2018) On-line diagnostics of the physical and mechanical properties of the surface at the production stage. In: MATEC Web of Conferences. Available via DIALOG. https://www.matec-conferences.org/articles/matecconf/ abs/2018/83/matecconf_icmtmte2018_01079/matecconf_icmtmte2018_01079.html. Accessed 30 Oct 2018

Mathematical Apparatus for Predicting Cutting Tool Life in Turning Process After Prior Plastic Deformation D. Kraynev, A. Bondarev and Z. Tikhonova

Abstract The paper considers the mechanism of action of turning with prior plastic deformation (PPD) on the laws of the cutting process and the durability of the cutting tool. The reasons and physics of this method of processing on the characteristics of the treated surface are analyzed. The theoretical substantiation of the established effect, as well as practical recommendations of application, are offered. It is established that the change of physical and mechanical properties of the processed material after PPD, the change of physical processes in the processing zone, reduces the temperature and force stress of removal of the allowance and the load on the cutting edge. The authors developed a mathematical model that predicts the life of the cutting tool blades in turning after the APD of the surface machined. The regression analysis has revealed the response function effect and allowed evaluating the effect of each factor on it. A formula for calculating the life of the cutting tool blade has been proposed, which expands the applicability of the APD cutting method in the practice of machining.




Keywords Numerical scheme Turning Prior plastic deformation Hard alloy





Wear
Durability
Cutting tool


1 Introduction Structural alloyed steels are widely used for the manufacture of machine parts. The operations of turning processing are applied in technological processes of manufacturing shafts and other rotary bodies demanded much of high surface quality and wear resistance. The processing specificity of structural alloyed steels is the complex loaded state of the cutting edge, which causes high wear of the tool blades [1]. Increasing the machining efficiency of structural alloyed steels is an urgent task.

D. Kraynev (&)
A. Bondarev
Z. Tikhonova Volgograd State Technical University, 28, Lenin Ave, Volgograd 400005, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_118

1107

1108

D. Kraynev et al.

2 Theoretical Relevance The task presented can be solved in the following ways: increasing the service life of the cutting tool, machining efficiency, and quality of the resulting products. Cutting with the prior plastic deformation (PPD) that combines the stages of prior surface plastic deformation and subsequent removal of the stock left for cutting tool machining is a combined processing method that allows obtaining a complex result of an increase in efficiency in these areas. Thus, the change in the physicomechanical properties of the material being PPD processed determines the physical processes conditions in the cutting zone that reduce the loads on the cutting edge [2, 3]. According to the theory of crystal structure, all metals and alloys are polycrystalline; that is, they consist of a set of fused crystallites, which are irregular grains with anisotropic mechanical, physical, and chemical properties. The crystallites in ferritic and pearlitic steel and alloys are characterized by a bcc structure and consequently, there are eight possible slip directions. The hcc crystals of corrosion resistant and high-temperature austenitic steel and alloys are characterized by 12 possible slip directions at normal temperature: four octahedral planes and three directions in each plane. Therefore, the plasticity of these materials is considerably higher [1]. Plastic deformation is accompanied by transition of the material from an equilibrium state, with minimum potential energy, to a non-equilibrium state, with increased potential energy. As a result, not only the geometric shape of the body varies in plastic deformation but also the mechanical and physical properties, which depend on the excess potential energy and its distribution over the volume [4]. Much of the thermal energy associated with plastic deformation is not retained within the deformed element, as shown in [5]; the small proportion that is retained increases the internal energy of the element. Note also that energy accumulation ends at some limiting value corresponding to disintegration of the metal. According to the first law of thermodynamics, the external action on a system may be mechanical or thermal, as noted in [6]. The heat is assumed to be destructive if the internal energy density reaches the critical value in at least one local volume. The deformation in cutting is associated with an increase in the dislocation density. The presence of dislocations leads to elastic distortion of the lattice and increase in its energy. The atoms in the dislocation zone are less stable than the others in the lattice, on account of their higher free energy. Therefore, their displacement requires considerably smaller shear stress than for the simultaneous displacement of all the atoms in the slip plane within an ideal crystal lattice. Hence, the shear in the crystals is not due to simultaneous slip of all the atomic planes but to the successive motion of the dislocations along the shear planes. In both cases, the final result will be the same: displacement of the two parts of the crystal by a distance corresponding to a single atom [1].

Mathematical Apparatus for Predicting Cutting Tool Life …

1109

When the cutting tool begins to act on the cut layer, some of the work that would be done in plastic deformation during chip formation in the case of regular cutting has already been performed by an additional device. During preceding plastic deformation, the crystal lattice of the blank is saturated with deformational energy. The dislocation density increases. The atoms in the dislocation zone are less stable because of their higher free energy. Therefore, their displacement along the shear plane and the corresponding chip formation will require much less shear stress from the tool than in standard cutting. Hence, in the presence of preceding plastic deformation, the tool does not perform all the work, but only part of it. This reduces the cutting force and heat dissipation and as a result decrease tool wear [7–9]. As a result of this effect, the PPD has a beneficial effect on the microprofile of the treated surface. In particular, the arithmetic mean deviation of the RA profile decreases [10, 11]. The research has established that PPD turning structural alloyed steels over the surface to be machined can considerably reduce the wear area on the main rear surface, thereby increasing the cutting process efficiency [12, 13]. In PPD turning of hot-strength relaxation-resistant constructional alloyed steel 30XMA, there was a decrease in the value of the wear area on the front surface by up to 2 times compared with the conventional machining. The reserves of essential improvements in the cutting process efficiency were identified. The same values of the wear area on the front surface were obtained in conventional turning at the same speed and in cutting at high speeds of the priory deformed metal. To study the wear patterns of the cutting tool blades in PPD turning, predict the results of the process, and expand the applicability of the cutting method in the light of the advantages presented, it was necessary to perform the regression analysis of the effect of the main processing parameters of the cutting tool blades wear [14, 15]. The studies were performed for semi-finishing turning (cutting depth t = 1 mm) of structural steel alloy 30XMA (profile iron, delivery condition, and rв = 930 MPa). The tool was represented by replaceable multi-faceted carbide plates T15K6, BK6, TH20 of the W form (a = 0° (ISO 1832-1991)); the conjugation radius of the main and auxiliary cutting edges was r = 0.8 mm. Processing was performed without cooling medium. The cutting speed, thermal conductivity of the tool material, feed rate, and PPD coefficient were selected as the factors that determine the response function value (T). CPPD ¼

hriv t

ð1Þ

where hriv is the depth of the riveted surface layer, mm; t is the depth of cut, mm. The natural and non-dimensional values of the factors accepted in the study are presented in Table 1. The studies revealed the extreme change in the PPD effect of the surface being machined on the efficiency of the turning process [16]. The increase in CPPD to a certain value caused the cutting force decrease to the minimum at a certain ratio between hriv and t (for the conditions described in this article, this ratio was 2). A further increase in CPPD led to a decrease in the process

1110

D. Kraynev et al.

Table 1 Natural values of factors Factors

Natural values

Normalized value (level) Cutting speed (m/min) Thermal conductivity of tool material (W/m C) Speed of longitudinal feed (mm/rev) PPD coefficient (CPPD)

−1 90 11 (TH20) 0.083 0.001

0 135 27 (T15К6) 0.166 1

+1 180 50 (BК6) 0.256 2

efficiency. According to the experiments, the lower CPPD level was 0 (in this case, the cutting process took place with no PPD), but for building the model, it was assumed to be equal to 0.001 in order to enable performing further calculations involving logarithms. Under this assumption, the depth of work hardening would be equal to 0.0005 mm, which was insignificant. Therefore, this convention would not affect the result of the regression analysis. For the research, there were mathematical models of additive (linear) and multiplicative (power and exponential) specifications accepted as the most frequently considered. The task of using mathematical methods for planning experiments was to obtain a mathematical description of the response function in the form of a mathematical model linking this function with variable factors after the experiments conducted. The minimum and sufficient amount of statistical sampling, the simulation was performed on, was determined depending on the model specification and number of factors considered (in this study, there were 4) and was equal to For additive model: kmin ¼ 2n ¼ 24 ¼ 16

ð2Þ

kmin ¼ 3n ¼ 34 ¼ 81

ð3Þ

and for multiplicative models:

where kmin is the number of independent experiments in a full factorial experiment and n is the number of factors. When building a model, the initial formal evaluation for the existence of the dependence of the response function on each of the factors considered was being obtained on the basis of the value of the pair correlation coefficient [17] that indicated the probability of a linear relationship between two statistical samples of equal volumes. Similarly, the relationship between factors was roughly estimated. Based on the calculations performed, the selected factors were established to be pairwise mutually independent and each of them influenced the magnitude of the response function. To evaluate the influence degree of each factor, a multifactorial regression model was constructed. The model building included the following necessary operations performed.

Mathematical Apparatus for Predicting Cutting Tool Life …

1111

The factors’ effects on the nature and magnitude of the response function were possible to evaluate correctly only if the values of all factors were comparable, including the objective function itself. Since the factors under consideration had different scales, the initial data was normalized according to the rules [15, 17] corresponding to the regression models of various specifications. Building a model after normalization involved the following independent specifications and sequential calculations [15, 17], i.e., the calculation of the regression coefficients for the factors selected; calculation of the regression coefficient that determined the influence of random factors; calculation of the normalized values of the response function; calculation of student’s criterion that determined the materiality of the difference between the regression coefficients and zero values; evaluation of statistical significance and determination of the effect of the factors considered (for the linear, power and exponential models, all the factors analyzed turned out to be statistically significant); calculation of the predicted normalized values of the response function, taking into account the statistical significance of the factors; calculation of the restored (in real scale) values of the response function for the normalized series; and adequate assessment of the models constructed by the F-criterion value. For the analysis of the mathematical models constructed, we summarized the most significant simulation results in Table 2. The data analysis (in Table 2) allowed us to draw the following conclusions. The average ratio error (the divergence between the T tool life values calculated by the model and obtained experimentally) of the exponential model was lower than the similar parameters of the power and linear models (9.08% linear; 8.61% power; and 6.49% exponential), which was an advantage. The effect of the random factors on the tool life was expressed by the corresponding regression coefficient and in the exponential and power models in absolute magnitude was equal to −0.03, i.e., insignificant. In the linear model, the influence of random factors had an absolute

Table 2 Comparison of regression models Parameter Factor impact evaluation Random factors Cutting speed Thermal conductivity of the tool material Feed movement PPD coefficient (CPPD) Reliability of an adequate model (%) Error of estimate Standard deviation F-test Average ratio error (%)

Model specification Linear

Power

Exponential

−0.28 −0.40 0.14 0.09 0.35 99.9

−0.03 −0.48 0.08 0.03 0.22 99.9

−0.03 −0.47 0.08 0.04 0.33 99.9

2.014 2.530 16.08

2.014 2.344 8.61

2.014 3.342 6.49

1112

D. Kraynev et al.

value of −0.28, which was quite significant against the background of the factors adopted in the study. Thus, the exponential model facilitated a further prediction of the cutting tool life after PPD turning, as it allowed concluding that the cutting speed, thermal conductivity of the tool material, feed movement, and PPD had the most significant effect on the cutting tool life. The regression coefficients in the exponential model showed that the cutting tool life in PPD turning was influenced by the following factors in ascending order: feed movement, thermal tool conductivity, PPD, and speed. A feed increase caused a growth of the cutting tool service life (the regression coefficient of 0.04 for the exponential model). An increase in thermal conductivity caused an increase in the cutting tool life, in accordance with the ideas traditionally established (the regression coefficient of 0.08 for an exponential model). The PPD could significantly improve the cutting tool life (the regression coefficient of 0.33 for a exponential model). An increase in cutting speed caused a decrease in the blades life of the cutting tool (the regression coefficient of −0.47 for the exponential model), which was consistent with traditional ideas about the effect of cutting speed on the cutting tool life.

3 Practical Relevance The regression mathematical model of the main PPD cutting parameters’ effect on the value of the cutting tool life in turning structural alloy steel 30XMA was implemented in the form of the following relationship: T ¼ 23:65
0:995V
1:0023k
1:3S
1:2CADD

ð4Þ

where V is the cutting speed, m/min; k is the thermal conductivity of the tool material, W/m K; So is the feed, mm/rev; and CPPD is the ratio of the PPD coefficient. This formula enabled calculating the magnitude of durability in conventional turning and PPD turning, i.e., predicting an increase in the efficiency of the cutting process by the method being studied and taking this into account in the technological processing route. The ranges of factors’ variation (for example, cutting speeds of 90–180 m/min) and processing conditions for the practical application of the dependence are described above.

4 Conclusion The prior plastic deformation is a comprehensive way to increase the efficiency of turning structural alloyed steels. A change in the physicomechanical properties of the material being processed after the PPD and in the physical processes in the

Mathematical Apparatus for Predicting Cutting Tool Life …

1113

treatment zone allow reducing the temperature and force intensity of the stock removal and load on the cutting edge. In addition, as a result of the impact of PPD, the stability of the chip formation process increases, the tool life increases, and the roughness of the treated surface improves. A regression analysis was carried out and allowed us to evaluate the effect of the factors on the cutting tool life. A formula for calculating the cutting T tool life was proposed, which expanded the applicability the PPD cutting method in the practice of machining.

References 1. Poduraev VN (1974) Rezanie trudnoobrabatyvaemykh materialov (cutting of hard-to-work materials). Vysshaya Shkola, Moscow 2. Polyanchikov YuN, Kraynev DV et al (2012) Improving medium to finish turning of stainless and heat resisting steels by the use of advanced plastic deformation. ONIKS, Togliatti 3. Polyanchikov YuN, Krainev DV et al (2011) Improved cutting of steels by means of preceding plastic deformation. Russ Eng Res 31(1):82–84 4. Mozberg RK (1991) Materialovedenie (materials science). Vysshaya Shkola, Moscow 5. Lebedev VA, Podel’skii MA (2004) Assessing the efficiency of preceding plastic deformation on thermodynamic principles. Vestn Mashinostr 9:63–67 6. Ignatov SN, Karnov AV (2004) Assessing the efficiency of cutting by means of a dimensionless energy index. STIN 12:23–25 7. Polyanchikov YuN, Krainev DV, Norchenko PA et al (2011) Improved cutting of steels by means of preceding plastic deformation. Russ Eng Res 31:82–84. https://doi.org/10.3103/ S1068798X11010187 8. Bondarev AA, Kraynev DV et al (2014) Increasing of working capacity of the cutting tool during turning with advanced plastic deformation. In: Modern scientific research and their practical application: research bulletin SWorld. https://www.sworld.com.ua/e-journal/j11410. pdf. Accessed 22 Mar 2014 9. Nesterenko PS, Bondarev AA et al (2014) Mathematical model of formation the axial component of the cutting forces at turning structural alloyed steel 3120. In: Modern scientific research and their practical application: research bulletin SWorld. https://www.sworld.com. ua/e-journal/j11410.pdf. Accessed 22 Mar 2014 10. Ingemansson AR, Zaitseva NG et al (2012) Improving the quality of processing and a mathematical model of formation of surface roughness when cutting with outrunning plastic deformation. Bulletin of Ufa State Aviation. Tekhn Univ 4:165–170 11. Norchenko P, Kraynev D, Kozachuknenko I (2016) The problem of forecasting the facial layer quality while cutting steel with advance plastic deformation. In: Radionov AA (ed) Procedia engineering: 2nd international conference on industrial engineering (ICIE-2016), vol 150. Elsevier Publishing, pp 821–826 12. Ingemansson AR (2012) Improving the turning efficiency of hard-to-cut steels of ferritic, martensitic-ferritic, and martensitic classes using prior plastic deformation. Dissertation, Volgograd State Technical University 13. Loladze TN (1958) Prochnost i iznosostoykost rezhushego instrumenta (Strength and wear resistance of the cutting tool) Mashgiz, Moscow 14. Tchigirinsky JuL (2010) Matematicheskie metodi upravlenia processami mehanicheskoy obrabotki (Mathematical methods of the control of the machining processes). Volgograd State Technical University, Volgograd

1114

D. Kraynev et al.

15. Evdokimov YuA et al (1980) Planirovanie i analiz experimentov pri reshenii zadach trenia i iznosa (Planning and analysis of experiments in solving problems of friction and wear). Science, Moscow 16. Bondarev AA, Kraynev DV et al (2014) Improving the performance of the cutting tool in prior plastic deformation turning. In: Modern scientific research and their practical application: research bulletin SWorld. https://www.sworld.com.ua/e-journal/j11410.pdf. Accessed 22 Mar 2014 17. Chigirinsky YuL, Chigirinskaya NV et al (2002) Stohasticheskoe modelirovanie v mashinostroenii (Stochastic modeling in mechanical engineering). VSTU, Volgograd

Numerical Modeling of Heat Transfer and Material Flow During Wire-Based Electron-Beam Additive Manufacturing A. V. Shcherbakov, D. A. Gaponova and R. V. Rodyakina

Abstract The development of a mathematical model provides an analysis of heat transfer and metal flow during wire-based electron-beam additive manufacturing described. The procedure for solving the heat equation for the metal in the solid phase and the Navier–Stokes equations in the liquid phase, based on the use of the finite-difference method and the predictor–corrector procedure, is described. An algorithm for numerical approximation of the motion of the free surface of the melt, using the concept of the volume of fluid (VOF) is described. The original method for calculating the effect of surface tension forces, based on the numerical calculation of the surface curvature index, is proposed. The results of simulating the melting of a wire element localized above a substrate made of 316L steel are described. Experiments showed the predominant role of surface tension force in the formation of deposited layer and also that metal’s flow has a laminar structure. These results were obtained by simulating a short-time beam exposure (t = 0.1 s) with a power of 6 kW. Thus, even when the wire and the substrate are exposed with a more intense beam than often used in practice, the metal transfer is not characterized by the formation of intense vortex flows. This can simplify the solution of the problem of additive manufacturing modeling in the future.







Keywords Additive manufacturing Electron beam Melting Mathematical model Navier–Stokes equations Heat transfer Surface tension









Fluid flow


1 Introduction Additive manufacturing processes are related to new production technologies, and their development and introduction is necessary to realize the concept of digital production in our life. At present, additive technologies are of great interest for the A. V. Shcherbakov
D. A. Gaponova
R. V. Rodyakina (&) National Research University MPEI, 17, Krasnokazarmennaya str., Moscow 111250, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_119

1115

1116

A. V. Shcherbakov et al.

industry, which allow to produce prototypes of products or half-finished products from metallic materials [1]. Using of powders as a filler material imposes certain restrictions on both process itself and technical and economic indicators of production. Spheroidized powders have a high cost, and collecting and processing powder for its reuse takes much time and causes an increase in waste material [2]. The current field of application of this technology is the production of small-sized products from a limited range of materials. The search for alternative solutions to make the technology inexpensive and universal has led to the creation of a whole group of processes in which as a filler material it is used such available material as welding filler wire [3, 4]. Electric arc [5, 6], plasma jet [7], laser [8], and electron beams [9] are used as heat sources for realization of this technology. Wire-based electron-beam technology of additive manufacturing (named electron-beam freeform fabrication, EBF3 [10]) best meets the requirements for structure and quality of formed metal layer, since the whole process is realized in vacuum (10−1–10−5 Pa). At the same time, the equipment has high efficiency (in comparison with laser), and using of electron-beam oscillations allows practically inertia-free control of the position, shape, and dimensions of heating zone to regulate metal transfer mode. At the same time, it is too early to talk about the wide industrial introduction of this technology. At present, control of metal transfer mode, which is influenced by many factors, remains a practically unsolved problem and is primarily associated with inconstancy of heat transfer conditions in “heat source—underlayer” system. The study of the process of heat and mass transfer can be carried out by means of experimental methods, which requires using of expensive thermal imaging equipment [11, 12]. However, even if it is used, researchers receive only indirect information about temperature distribution over the surface of layer and wire. To obtain information about speed distribution and pressure, as well as about temperature distribution over the depth of weld pool, it is necessary to use methods of mathematical modeling.

2 Problem Statement and Numerical Method Heat and mass transfer in the system “electron beam—filler wire—product” should be described in a formulation that takes into account the flow of metal under the action of pressure field, gravity, and forces due to viscosity and surface tension. For this, it is necessary to use the Navier–Stokes equations, which describe the flow of a viscous incompressible fluid, and include, firstly, the continuity equation @vx @vy @vz þ þ ¼ 0; @x @y @z

ð1Þ

where vx, vy, and vz are the projections of speed vector on the x, y, and z axes. Secondly, the system includes the equation of motion, which, when written in projection form, will have the form

Numerical Modeling of Heat Transfer and Material … @tx @t @ty @t

@tx @tx 1 @P x þ tx @t @x þ ty @y þ tz @z ¼  q @x þ l

@tz @t

z þ tx @t @x

þ tx

@ty @x

þ ty

@ty @y

z þ ty @t @y

þ tz

@ty @z

z þ tz @t @z




¼

@ 2 tx 2
@x2 @ t y  q1 @P @y þ l @x2

¼

 q1 @P @z

þl




@ 2 tz @x2

1117

þ

@ 2 tx @y2 @ 2 ty @y2

þ

@ 2 tz @y2

þ



1 dL q rj dx

þ

@ 2 tx @z2 þ  @ 2 ty þ 2 @z

þ

@ 2 tz @z2

1 dL q rj dz

þ



þ

1 dL q rj dy ;

ð2Þ

where P is pressure, Pa; l is kinematic viscosity, m2/s; t is time, s; g is free fall acceleration, m/s2; q is density, kg/m3; r is surface tension coefficient, J/m2; j is curvature index, 1/m; L is dimensionless function of filling the control volume with fluid, 0 < L < 1. At L = 1, the control volume is filled with fluid, and at L = 0, it is empty. This method is known as the volume of fluid (VOF) method, and it has proven itself for solving hydrodynamic problems [13, 14]. In accordance with this method, the flow of fluid between the control volumes is described by means of equation @L @L @L @L þ tx þ ty þ tz ¼ 0: @t @x @y @z

ð3Þ

And closes the system the following equation of heat transfer   @T @T @T @T k @2T @2T @2T Qv þ tx þ ty þ tz ¼ : þ 2 þ 2 þ 2 @t @x @y @z cq @x @y @z cq

ð4Þ

In the last equation, T is temperature, K; c is specific heat capacity, J/(kg K); Qv is volume power of heat release, W/m3. Heat transfer in the solid phase is described by the equation of heat conductivity,   @T k @2T @2T @2T Qv @w ¼  Hm q m ; þ 2 þ 2 þ @t cq @x2 @y @z cq @t

ð5Þ

where Hm is latent heat of fusion of metal, J/kg; wm(T) is the proportion of the melt in two-phase region, calculated in each control volume at each moment of time with using equation wm ðTÞ ¼

8 < 0;

TTs : TL TS

1;

T\TS ; TS \T\TL : T [ TL

ð6Þ

In Eq. (6) TS and TL are, respectively, solidus and liquidus temperatures of the material being remelted, K. To solve the system of Eqs. (1)–(6), the finite-difference method and a fixed grid with cubic cells were used. This grid has “staggered” structure [14, 15]. In such a grid, the nodes, in which speeds are calculated, are displaced along all three axes by half the coordinate step relative to the nodes in which there are calculated pressure P, function of cell’s filling L, and fluid temperature T (Fig. 1).

1118

A. V. Shcherbakov et al.

Fig. 1 Grid cell for a finite-difference approximation of Eqs. (1)– (6). In parentheses are shown the indices of grid nodes numbers

At the initial stage, when melting had not yet begun, a numerical solution of Eq. (5) was made. The grid spacing was taken to be h = 0.1 mm, and the time spacing was 8  10−5 s. At the current stage of research, an explicit difference scheme was used with fixed coordinate and time steps. To calculate the temperature in the current coordinate node at time moment t + 1, we used the following finite-difference approximation of Eq. (5): tþ1 Tðx;y;zÞ



¼

t Tðx;y;zÞ

k Qv Bþ þ Dt cq cq

 ð7Þ

where B¼

t t Tðxt þ 1;y;zÞ  2Tðx;y;zÞ þ Tðx1;y;zÞ

þ

þ

t t t Tðx;y þ 1;zÞ  2Tðx;y;zÞ þ Tðx;y1;zÞ

h2 t t t Tðx;y;z  2Tðx;y;zÞ þ Tðx;y;z1Þ þ 1Þ

h2

ð8Þ

h2

The normal distribution function was used as the Qv function

Qv ¼

8 0; > < 1

pr2 > : eff 0;

Ib Ua exp

   2  ðxxpos ðtÞÞ þ ðyypos ðtÞ2 Þ 2 reff

z\zsur dz ; 1

z ¼ zsur

ð9Þ

z [ zsur

In Eq. (9) zsur is z-coordinate of material’s surface, on which electron beam acts; Ib is beam’s current; Ua is accelerating voltage; xpos, ypos are the current values of electron-beam center coordinates; reff is effective beam radius (assumed to be 1.2 mm).

Numerical Modeling of Heat Transfer and Material …

1119

After reaching the temperature TL in the control volume, metal was considered to be molten, and the system of Eqs. (1)–(4) was numerically solved. To solve Eqs. (1)–(2), the SIMPLER method proposed by Patankar [15] was used. In accordance with this method, at the first time step, speed field is assumed known. In our case, the fluid rest condition was set (all the speeds are zero). Under these conditions, the speed increments were calculated by Eq. (2) without taking into account values of pressure (predictor step). For example, to calculate the speed increments along the x-coordinate the following equation was used: ðt þ 1Þp vxfx;y;zÞ

¼

vtxfx;y;zÞ

! t t 1 Lðx þ 1:y:zÞ  Lðx1;y;zÞ þ Dt A þ lB þ rj ; q 2h

ð10Þ

where A ¼ vtxfx;y;zÞ



vtxfx þ 1;y;zÞ  vtxfx1;y;zÞ 2h

vtxfx;y;z þ 1Þ  vtxfx;y;z1Þ 2h

 vtyfx;y;zÞ


vtxfx;y þ 1;zÞ  vtxfx;y1;zÞ 2h

 vtzfx;y;zÞ

; ð11Þ

B¼

vtxfx þ 1;y;zÞ  2vtxfx;y;zÞ þ vtxfx1;y;zÞ þ

þ

h2 vtzfx;y;z þ 1Þ  2vtzfx;y;zÞ þ vtzfx;y;z1Þ h2

vtyfx;y þ 1;zÞ  2vtyfx;y;zÞ þ vtxfx;y1;zÞ h2

ð12Þ

:

Then, the scalar pressure field in the liquid metal (step-corrector) was calculated using the method described in [15]. The pressure was calculated by the Poisson equation, so as to satisfy the continuity equation [16, 17].   @2P @2P @2P q @vx @vy @vz þ þ þ 2 þ 2 ¼
: @x2 @y @z Dt @x @y @z

ð13Þ

Equation (13) was solved with using simple iteration method by expressing the pressure in the node under consideration (with the subscripts x, y, z) through the pressures in the surrounding nodes and the derivatives of speeds calculated in displaced grid nodes (see Fig. 1). 1 h2 q D; Pðx; y; zÞ ¼ A  6 6Dt where

ð14Þ

1120

A. V. Shcherbakov et al.

A ¼ Pðx þ 1; y; zÞ þ Pðx  1; y; zÞ þ Pðx; y þ 1; zÞ þ Pðx; y  1; zÞ þ Pðx; y; z þ 1Þ þ Pðx; y; z  1Þ; D¼

vxðx þ 1;y;zÞ  vxðx1;y;zÞ 2

2

h

þ

vyðx;y þ 1;zÞ  vxðx;y1;zÞ 2

2

h

þ

ð15Þ

vzðx;y;z þ 1Þ  vzðx;y;z1Þ 2

2

h

:

ð16Þ

After finding the pressure field, the acceleration projections due to its action were calculated and the speed field was corrected. For example, for speed projection on the x-axis, we have the following equation: tþ1 vxðx;y;zÞ

¼

ðt þ 1Þp vxðx;y;zÞ



 1 Pðx þ 1;y;zÞ  Pðx1;y;zÞ þ  Dt: q 2h

ð17Þ

After the final calculation of speed field, one can proceed to the calculation of temperature field in the melt. To do this, the following equation was used: tþ1 Tðx;y;zÞ



¼

t Tðx;y;zÞ

 k Qv B  Dþ þ Dt ; cq cq

ð18Þ

where B¼

t t Tðxt þ 1;y;zÞ  2Tðx;y;zÞ þ Tðx1;y;zÞ

þ D ¼ vtxðx;y;zÞ

h2 t t t Tðx;y;z  2Tðx;y;zÞ þ Tðx;y;z1Þ þ 1Þ h2

t Tðxt þ 1;y;zÞ  Tðx 1 ;y;zÞ 2

2

h

þ vtyðx;y;zÞ

t t t Tðx;y þ 1;zÞ  2Tðx;y;zÞ þ Tðx;y1;zÞ

þ

h2

ð19Þ

;

t t Tðx;y  Tðx;y 1 þ 1;zÞ ;zÞ 2

2

h

þ vtzðx;y;zÞ

t t Tðx;y;z  Tðx;y;z 1 þ 1Þ Þ 2

2

h

:

ð20Þ After calculating vector and scalar fields in the melt boundaries, it is necessary to simulate the motion of free surface, and also to determine the fields of surface tension force.

3 Simulation of Free Surface Motion and Surface Tension Forces The finite-difference approximation of Eq. (3) for the case, if no additional conditions are specified, should have the following form

Numerical Modeling of Heat Transfer and Material …

1121

tþ1 Lðx;y;zÞ ¼ Ltðx;y;zÞ þ Dt
ðA þ CÞ ¼ Ltðx;y;zÞ þ Dt
X;

ð21Þ

where A¼

t Ltðx þ 1;y;zÞ
vtxðx þ 1;y;zÞ  Tðx
vtxðx1;y;zÞ 1 ;y;zÞ 2

2

2

2

h

þ

t Ltðx;y þ 1;zÞ
vtyðx;y þ 1;zÞ  Tðx;y
vtyðx;y1;zÞ 1 ;zÞ 2

2

2

h

2

;

ð22Þ C¼

t t Ltðx;y;z þ 1Þ
vtzðx;y;z þ 1Þ  Tðx;y;z 1
v xðx;y;z1Þ Þ 2

2

2

h

2

:

ð23Þ

However, a direct numerical solution of such a system can lead to the appearance of anomalous results, since this approach does not take into account that function L in each control volume should always take values in the range from 0 to 1. This circumstance led to the creation of various methods for solving Eq. (3), described in the literature [14, 16–20]. In this work, the sign of the speed and magnitude of fluid streams flowing into the control volume, or, on the contrary, flowing from it, was tracked. In the case of overflow of control volume, or, on the contrary, its emptying, a proportional decrease in the stream was made to prevent the output of the function L value out of the interval 0–1. In addition to the above conditions introduced into the algorithm for solving Eq. (14), the calculation of the function L distribution was carried out iteratively with the fragmentation of time step. The maximum fractional step Dtfrac was selected from the condition Dtfrac
max X  0:1:

ð24Þ

The technique of calculating surface tension forces is of particular interest. Calculation of surface tension forces field was made on the basis of original developed algorithm. The normal to the free surface was calculated as the direction of the function gradient L, and the sign and magnitude of the force were calculated from the surface curvature j (see Eq. 2). The curvature index was determined as follows. In the computational grid, a spherical region with a center at the point under consideration (O2) lying on the free surface was distinguished (Fig. 2a). The region has radius R2 (in this model there was used a fixed radius of 0.3 mm). Then, by summing the volumes inside the sphere and occupied by the fluid, the total volume of fluid inside the sphere was calculated (the volume V2 is shown in Fig. 2) and the curvature index was calculated from geometric considerations by means of equation j¼

1 12V2  8pR32 ¼ R1 3pR42

ð25Þ

1122

A. V. Shcherbakov et al.

Fig. 2 To the calculation of curvature index of the surface (a) and acceleration caused by the force of surface tension (b); 1—liquid metal, 2—technological vacuum

Since the expression for determining the acceleration due to the action of surface tension force, given in (2), contains the derivative of function L, this allows to provide the “compression” of free surface due to the action of surface tension force and eliminates its “blurring,” which is characteristic for many numerical algorithms.

4 Results of the Computational Experiment and Their Analysis The model described by expressions (1)–(25) was realized as a program in the Microsoft Visual Studio environment. As the first computational experiment, the problem of simulating the melting of a fixedly set filler wire under the action of an electron beam was chosen. The input parameters used in this model are listed in Table 1. Typically, when surfacing using 1.2 mm wire diameter, beam power does not exceed 1–2 kW; however, during this computational experiment, the exposure time was reduced, and the power was increased to reduce the calculation time. The results of computational experiments are graphically presented in Fig. 3. The initial stage of electron-beam influence on wire feed and substrate material is shown in Fig. 3a. Melting has not yet occurred at this stage (t = 2.4  10−4 s). For a metal in the solid phase, the temperature is displayed by a gradient of gray (the higher the temperature, the brighter the shade). A drop of metal formed on the wire under the action of gravity and surface tension force is shown in Fig. 3b, which corresponds to the point t = 0.0527 s. The color temperature corresponds to metal’s temperature in the liquid pool. White strokes show speed vector of metal. Figure 3c corresponds to the moment of drop’s merging with liquid pool formed on the substrate under electron-beam influence (t = 0.059 s), and Fig. 3d corresponds to the end of calculations carried out at this stage of research (t = 0.108 s). At this

Numerical Modeling of Heat Transfer and Material …

1123

Table 1 Model parameters Parameter

Meaning

Symbol

Wire feed and substrate material Heat conductivity Heat capacity Density Kinematic viscosity Surface tension coefficient Latent heat of fusion Coordinate step Initial (specified) time step Pressure in the vacuum chamber Effective electron-beam radius Electron-beam power Wire feed diameter Substrate size (Length  Width  Height) Beam exposure time

316L steel 30 500 7000 6  10−6 1.8 280,000 1  10−4 8  10−5 0.1 1.2 6000 1.2 12  8  2 0.1

– W/mK J/(kg
К) kg/m3 m2/s J/m2 J/kg m s Pa mm W mm mm s

Fig. 3 Results of computational experiments on melting 316L steel wire

moment, the beam is already turned off, and the crystallization on the peripheral area of the molten pool has begun. Series of computational experiments showed the decisive role of surface tension in the process of deposited layer shape formation. Also series of computational experiments showed that metal’s flow is laminar in nature, without the formation of intense vortex flows.

1124

A. V. Shcherbakov et al.

These results are confirmed by experimental data, namely the fact that during the process of deposition the “soft” mode of metal transfer is usually achieved, in which there are no significant fluctuations in height and width of deposited layers, splashes, and other defects typical for materials processing by means of concentrated energy flows. This is primarily due to the fact that during wire-based electron-beam additive manufacturing, a technological sweep is usually used to distribute the heating source over a large area, which reduces effect of metal vapor pressure on liquid pool surface and transition from turbulent to laminar flow. All of these circumstances confirm the effectiveness of proposed model for studying of more complex tasks, such as surfacing of multilayer structures.

5 Conclusions 1. A model of heat transfer and metal flow during wire-based electron-beam additive manufacturing was developed. It is based on the numerical solution of the Navier–Stokes equations and heat transfer equation, which allows analyzing the movement of melt’s free surface, the speed field, and the temperature field change over time during the process of deposited metal wire with given thermal characteristics. 2. A numerical algorithm is proposed for calculating the surface tension force. It is based on the selection of a spherical region with a center at a point lying on the surface of the melt, which allows to take into account the curvature of the surface and minimize the width of “liquid–vacuum” border. 3. Series of computational experiments were carried out to calculate the metal flow during deposition with the use of 316L steel wire feed. Experiments showed the predominant role of surface tension force in the formation of deposited layer and also that metal’s flow has a laminar structure. These results were obtained by simulating a short-time beam exposure (t = 0.1 s) with a power of 6 kW. Thus, even when the wire and the substrate are exposed with a more intense beam than often used in practice, the metal transfer is not characterized by the formation of intense vortex flows. This can simplify the solution of the problem of additive manufacturing modeling in the future. Acknowledgements This work was carried out in National Research University “Moscow Power Engineering Institute”; it was supported by grant from the Russian Science Foundation (Project 17-79-20015).

Numerical Modeling of Heat Transfer and Material …

1125

References 1. Zlenko MA, Nagaytsev MV, Dovbysh VM (2015) Additive technologies in mechanical engineering. Manual for Engineers. Moscow: State Science Center of Russia Federal State Unitary Enterprise “NAMI” 2. Gibson I, Rosen D, Stucker B (2015) Additive manufacturing technologies. 3D printing, rapid prototyping, and direct digital manufacturing. Springer Science + Business Media, New York 3. Ding D, Pan Z, Cuiuri D, Li H (2015) Wire-feed additive manufacturing of metal components: technologies, developments and future interests. Int J Adv Manuf Technol 81(1–4):465–481. https://doi.org/10.1007/s00170-015-7077-3 4. Ji L, Lu J, Tang S, Wu Q et al (2018) Research on mechanisms and controlling methods of macro defects in TC4 alloy fabricated by wire additive manufacturing. Materials 11:1104. https://doi.org/10.3390/ma110711041213 5. Williams SW, Martina F, Addison AC, Ding J et al (2016) Wire + arc additive manufacturing. Mater Sci Technol 32(7):641–647 6. Feng Y, Bin Z, He J, Wang K (2018) The double-wire feed and plasma arc additive manufacturing process for deposition in Cr-Ni stainless steel. J Mater Process Technol 259:206–215. https://doi.org/10.1016/j.jmatprotec.2018.04.040 7. Jhavar S, Jain N (2014) Development of Micro-plasma wire deposition process for layered manufacturing. In: Katalinic B (ed) DAAAM International scientific book, pp 239–256 8. Ding Y, Akbari M, Kovacevic R (2018) Process planning for laser wire-feed metal additive manufacturing system. Int J Adv Manuf Technol 95(1–4):355–365 9. Fuchs J, Schneider C, Enzinger N (2018) Wire-based additive manufacturing using an electron beam as heat source. Weld World 62(2):267–275 10. Taminger KMB, Hafley RA (2003) Electron beam freeform fabrication: a rapid metal deposition process. In: Proceedings of third annual automotive composites conference. Society of Plastic Engineers, Troy, Michigan, USA, pp 9–10 11. Zalameda JN, Burke ER, Hafley RA, Taminger KM et al (2013) Thermal imaging for assessment of electron-beam freeform fabrication (EBF3) additive manufacturing deposits. Proc SPIE 8705(87050M):8 12. Taminger KM, Domack CS, Zalameda JN, Taminger BL et al (2016) Thermal imaging of the electron beam freeform fabrication process. In: Proceedings of SPIE Commercial + Scientific Sensing and Imaging. Baltimore, MD, United States, Report No. NF1676L-22574 13. Noh WF; Woodward P (1976) SLIC (Simple line interface calculation). In: van de Vooren AI, & Zandbergen PJ (ed) Proceedings of 5th international conference of fluid dynamics. Lecture Notes in Physics 59, pp 330–340 14. Hirt CW, Nichols BD (1981) Volume of fluid (VOF) method for the dynamics of free boundaries. J Comput Phys 39(1):201–225 15. Patankar SV (1980) Numerical Heat transfer and fluid flow. Series in computational methods in mechanics and thermal sciences 16. Kampanis Nikolaos A, Ekaterinaris John A (2006) A staggered grid, high-order accurate method for the incompressible Navier–Stokes equations. J Comput Phys 215(2):589–613 17. Versteeg HK, Malalasekera W (1995) An introduction to computational fluid dynamics: the finite volume method. Longman Scientific & Technical 18. Gopala Vinay R, van Wachem Berend GM (2008) Volume of fluid methods for immiscible-fluid and free-surface flows. Chem Eng J 141(1–3):204–221 19. Rhee SH, Makarov BP, Krishinan H, Vladimir I (2005) Assessment of the volume of fluid method for free-surface wave flow. J Mar Sci Technol 10(4):173–180 20. Hargreaves DM, Morvan HP, Wright NG (2007) Validation of the volume of fluid method for free surface calculation: the broad-crested weir. Eng Appl Comput Fluid Mech 1(2):136–146. https://doi.org/10.1080/19942060.2007.11015188

Control of Weld Bead Position in Additive Manufacturing Process with Using Backscattered Electron Collector Signal A. V. Shcherbakov, D. A. Gaponova and R. V. Rodyakina

Abstract The urgency of solving the problem of determining the deposited layer’s height for controlling the process of additive forming products from metallic materials in a vacuum is substantiated. A scheme of electron-beam additive manufacturing process with wire feed is shown. A method for controlling the distance between deposited bead and electron gun by the signal taken from a collector of backscattered electrons is proposed. The analysis of the signals recorded during the process of deposited metal bead scanning has been carried out, and technical feasibility of this method has been substantiated. The technical solutions to introduce a method for continuous monitoring and control of additive forming process are proposed. A series of computational experiments at using the Monte Carlo method were carried out. They confirmed the relevance of the proposed method for controlling the distance between the electron gun and deposited bead. The calculation of beam energy fraction absorbed by the collector made it possible to substantiate the need for using a signal preamplifier. Based on the analysis of the results of experimental studies conducted on the ELA-15I power complex, the relevance of the proposed method for measuring the working segment and the possibility of using it as part of a process control system was confirmed.




Keywords Additive manufacturing Electron beam system Distance sensor Backscattered electrons








Filler wire
Control

1 Introduction The technology of additive manufacturing with a wire feed, called EBF3 or electron-beam additive manufacturing (EBAM), in the near future may be one of the most promising methods for prototyping products in mechanical engineering [1, 2]. The advantages of this technology are low cost of raw materials (widely used A. V. Shcherbakov
D. A. Gaponova
R. V. Rodyakina (&) National Research University MPEI, 17, Krasnokazarmennaya str., Moscow 111250, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_120

1127

1128

A. V. Shcherbakov et al.

Fig. 1 Scheme of electron-beam additive manufacturing process with wire feed: a the correctly chosen size of working distance, the wire is fed directly into the melt bath; b elongated working distance, the wire melts over the formed bead; c too small working distance, the wire rests against the sublayer; d–f corresponding to cases (a), (b), and (c) photographs of resulting deposited beads; 1—electron gun; 2—deflecting coils; 3—electron beam; 4—wire feeder; 5—wire; 6—liquid bath; 7—formed layer (deposited bead); 8—sublayer

welding wire), high performance [3], and possibility of flexible control of both input energy and feed speed of filler material [4]. At the same time, the implementation of this technology is hampered by problems caused by the complexity of layer-by-layer surfacing process as a control object. To substantiate the relevance of the proposed solution, it is necessary to consider the process diagram shown in Fig. 1. Electron gun 1 forms a beam of electrons 3 with an energy of 20–100 keV and with a power of 1–3 kW [5, 6], which with the help of magnetic deflecting coils block 2 quickly, with a frequency of hundreds Hertz, moves along a closed trajectory to create heating spots with the required size and shape. The roller mechanism 4, rigidly connected with the gun 1, feeds the wire 5 directly into the liquid bath 6. The product is moved relative to the gun, and thus, layer-by-layer welding of weld beads 7 to sublayers 8 is carried out. Figure 1a, corresponds to the mode with correctly selected working distance Z = Zw (distance “gun–weld bead”). In this mode, a stable formation of layers is achieved (Fig. 1d), and the deposition speed is maintained at a constant value. The movement of the formed product relative to the electron gun is carried out by a robotic manipulator [7]. At a constant deposition speed, the value of working segment Z should also be maintained at a constant level. However, determination of magnitude of this velocity is a time-consuming task, usually solved by means of trial and errors.

Control of Weld Bead Position in Additive

1129

On Fig. 1, b, a case when the size of working segment exceeds the required value (Z > Zw) is shown. In this mode, the melting of the wire occurs at a certain distance from the formed bead, which leads, on the one hand, to overheating of filler material with the formation of droplets and spatter, and on the other hand, to a change of speed field in the region of formed bead. As a rule, in such a mode, wider and less high beads with a shape and section that are not constant in length are formed (Fig. 1e). In Fig. 1c, an example of situation when the value of working segment is less than required value (Z < Zw) is given. In this case, there may be a bending of the wire, its displacement from the area of electron beam, and, accordingly, incomplete melting. The shape and dimensions of the bead will also be inconsistent in length, and the formed bead may contain serious defects in the form of fragments of unmelted wire (Fig. 1f). The given examples show the relevance of the construction of sensors and methods for continuous measurement of the position of deposited bead relative to electron gun and filler wire feed mechanism. Moreover, the need to control the power introduced into the product during the deposition process is already obvious [8]. This is due to changes in the conditions of heat transfer in the region offormed layer and is associated both with gradual heating of entire product and with the change in the thermal engineering massiveness of formed layer. The change in power during the surfacing process is likely to entail the need to regulate deposition rate and speed of movement along the vertical axis while maintaining the value of cutting Z constant. Thus, it is necessary to create an electron gun–bead distance stabilization system.

2 Development of Measuring Method It is obvious that the distance measurement should be carried out for an already formed bead, in order to exclude the influence of hydrodynamic processes on measurement accuracy. At the same time, the measurement should be carried out in the immediate vicinity of the area of electron-beam exposure to reduce the inertia of regulation. For this reason, metal’s temperature in the area where it is necessary to measure it, can be quite high. In this regard, the use of contact measurement methods based on the use of resistive, inductive sensors [9], or optical rulers [10] will be difficult due to thermal effects on the sensitive element of sensor. Therefore, it is necessary to give preference to contactless sensors. One of the most promising contactless methods for measuring distances is optical method based on the use of laser distance sensors [11] and laser profile meters [12]. In our case, the process is implemented in a vacuum chamber, in the volume of which an intensive evaporation of metal takes place. Vapors are deposited on the walls of the chamber and optical surfaces of sensors. For example, when installing a video camera with a protective glass, its dusting may occur within a few seconds after beginning of surfacing process. These circumstances compel to look for other methods of measuring distance between electron gun and bead.

1130

A. V. Shcherbakov et al.

Backscattered electron detectors are actively used in electron microscopy [13] and in process control systems for electron-beam welding [14, 15]. When scanning the surface of samples in the mode of low current, one can obtain information on the surface relief [13] and on the location of welded parts joint [15]. As beam current rises to the values at which intense metal evaporation begins, a plasma torch arises, which has a strong effect on the detected signal: The signal contains variable components of a wide frequency spectrum, making it difficult to solve the problem of obtaining information about the relief surface. Such signals are used mainly to control the mode of formation of penetration channel [16]. When analyzing the processes of electrons scattering from the surface of solids, it is usually assumed that backscattered electrons have a cosinusoidal angular distribution [17], and particle flux density decreases as the detector moves away from the material irradiated by electrons. The average angle of particles scattering is also related to the angle of incidence, but since small scattering angles (up to 5°) are used in technological installations, this fact can be disregarded. In view of the above, a scheme for detecting the current of backscattered electrons was proposed, providing information about the distance between electron gun 1 and deposited bead 2 (Fig. 2). There are practically no publications on the use of backscattered electron detectors for such task. The only published patent describes the possibility of measuring the working segment during deposition of beads by triangulation using various sensors, including arrays of detectors of backscattered electrons [18], but technical means and characteristics of this method were not disclosed, which causes the need for research in this area. To implement the proposed method, it is necessary to modulate electron-beam current 3. During periods of time when beam current has a value that ensures Fig. 2 Scheme for detecting the current of backscattered electrons to control the position of deposited bead: 1—electron gun; 2—formed layer (deposited bead); 3— electron beam; 4—filler wire; 5—liquid bath; 6— backscattered electrons; 7—collector; 8—preamp; 9—recorded signal; I—beam movement during surfacing; II—beam movement during scanning

Control of Weld Bead Position in Additive

1131

melting of the wire 4 with the formation of liquid metal bath 5, the beam moves along the trajectory I required given temperature field. Usually beam current in such a mode is 20–40 mA [6]. In other periods, beam current decreases to values at which there is no intensive heating of deposited bead (less than 1 mA), and the beam moves to already formed bead and moves across it along path II. After collapse of plasma torch formed over the liquid bath during melting period, it becomes possible to detect the signal of backscattered electrons 6 that fall on metal collector 7. The signal from collector is amplified in the screened preamplifier unit 8 and amplified signal 9 enters the analog–digital converter installed in the industrial computer.

3 Related of Recorded Signals with Bead Shape and Working Distance To confirm the efficiency of proposed method, theoretical and experimental studies were performed. Figure 3 shows the results of a statistical simulation of electron-beam interaction with an energy of 60 keV with metal bead (iron was chosen as bead’s material) and copper collector. The model of electrons’ interaction with a metal in solid phase was implemented with using the Monte Carlo method and continuous loss approximation [19]. The depth of electron penetration into the materials of deposited beads and collector when electrons are braked with energies of 60 keV and less is of the order of some micrometers, that is, several orders of magnitude less than their overall dimensions. The main task of calculation was to determine the fraction of particle energy absorbed by collector in the total energy of primary beam. If we set the current of primary beam, we can determine the current detected by collector during scanning by changing the position of the beam relative to the bead. Using this model, it was found that the current measured by collector is from tenths of a percent to 1–2% of the total electron-beam current. This confirms the need for a signal amplifier. A family of curves obtained by moving the beam

Fig. 3 Trajectories of backscattered electrons, calculated with using a statistical model: a in the transverse plane; b in the longitudinal plane; 1—deposited bead, 2—collector, 3—primary electron beam

1132

A. V. Shcherbakov et al.

Fig. 4 Dependences of energy fraction absorbed by collector on electron-beam displacement relative to the center of deposited bead, obtained with using a statistical model at different distances between collector and deposited bead: 1—15 mm, 2—20 mm; 3—25 mm, 4—30 mm: Wc is energy absorbed by collector; Wb is the total energy of the beam

across the bead is shown in Fig. 4. Since detector has limited dimensions, moving the target (bead) relative to the collector will lead to a change in the level of useful signal (all other conditions being the same): When removed from collector, the recorded current will decrease. Using the statistical model, curves 1–4 were obtained, corresponding to different positions of deposited bead relative to collector (and electron gun). It should be noted that the dependencies shown in Fig. 4 do not have a clearly marked maximum. The presence of several extremes in these curves is determined by model’s assumptions, namely that deposited bead is represented as consisting of cubic elements with dimensions of 0.2  0.2  0.2 mm, interacting with the edges which will scatter the electrons. To verify the results obtained, experimental studies were carried out on the ELA-15I power complex. An AD-711 micro-based inverter amplifier circuit [20] was used as a preamplifier, and the E-14-140-M L-Card analog-to-digital converter module was used to record the signal. The curves obtained as a result of experiments are shown in Fig. 5. The sampling frequency of ADC module was set at 100 kHz. The frequency of unwrapping electron beam is 100 Hz. Curve 1 is typical for the case, when distance between the collector and the product is equal 15 mm, and curve 2 is typical for the case, when this distance is equal 20 mm. The preamplifier gain is 500. The beam current is 1 mA. The calibration of obtained dependences at the current stage of research was not conducted, and a steel bar with a diameter of 5 mm was placed in vacuum chamber as a physical model of

Control of Weld Bead Position in Additive

1133

Fig. 5 Experimentally obtained dependences of the voltage removed from inverting preamplifier mounted on the collector of backscattered electrons at a beam current of 1 mA: 1—distance between the collector and the rod, simulating welded bead, is equal 15 mm, 2—20 mm

deposited bead. The horizontal displacement of the collector relative to the point of beam’s interaction with the rod was chosen experimentally without any criteria and amounted to about 10 mm. The dependences confirm the adequacy of the results obtained with using this model, as well as the applicability of the proposed method for determining the size of working segment (e.g., by measuring the maximum amplitude of detected signal corresponding to beam’s position at the top of deposited bead). At the same time, it was found that if the distance between collector and the rod is not great, as beam current increases to the values at which melting begins, the sensor may be deformed due to the effect of thermal radiation from the bath. Obviously, to eliminate this situation, it is necessary to increase the distance between sensor and the product and increase the gain of preamplifier.

4 Conclusions 1. The necessity of working segment stabilizing during electron-beam surfacing with filler wire is substantiated, and the characteristic defects of bead formation that arise when working segment size deviates from specified value were shown. It was shown that with an increase of this segment, a regime occurs, accompanied by overheating of filler material, transition to the formation of metal splashes, and drip transfer, and with a decrease—deformation of the wire and its incomplete melting can occur.

1134

A. V. Shcherbakov et al.

2. Technical limitations that prevent the introduction of contact and optical sensors for measuring the distance between the electron gun and the product are described, and relevance of using method of backscattered electrons detecting has been substantiated. 3. A series of computational experiments with using the Monte Carlo method were carried out. They confirmed the relevance of proposed method for controlling the distance between electron gun and deposited bead. The calculation of beam energy fraction absorbed by collector made it possible to substantiate the need for using a signal preamplifier. 4. Based on the analysis of the results of experimental studies conducted on the ELA-15I power complex, the relevance of proposed method for measuring the working segment and the possibility of using it as part of a process control system was confirmed. Acknowledgements This work was carried out in National Research University “Moscow Power Engineering Institute”; it was supported by grant from the Russian Science Foundation (Project 17-79-20015).

References 1. Gibson I, Rosen D, Stucker B (2015) Additive manufacturing technologies. 3D printing, rapid prototyping, and direct digital manufacturing. Springer Science + Business Media, New York 2. Milewski JO (2017) Additive manufacturing of metals: from fundamental technology to rocket nozzles, medical implants, and custom jewelry. Springer, New York 3. Gaponova DA, Shcherbakov AV, Rodyakina RV (2018) Investigation of wire feed control channel in additive manufacturing unit. Lecture notes in mechanical engineering. Springer, Cham, pp 2411–2418. https://doi.org/10.1007/978-3-319-95630-5_260 4. Taminger KMB, Hafley RA (2003) Electron beam freeform fabrication: a rapid metal deposition process. In: Proceedings of third annual automotive composites conference. Society of Plastic Engineer, Troy, Michigan, USA, pp 9–10 5. Fuchs J, Schneider C, Enzinger N (2018) Wire-based additive manufacturing using an electron beam as heat source. Weld World (Le Soudage Dans Le Monde) 62:1–4. https://doi. org/10.1007/s40194-017-0537-7 6. Gudenko AV, Dragunov VK, Sliva AP (2017) Metodika opredelenia rezhimov posloinoy electronno-luchevoy naplavki provoloki dlya additivnykh technologiy (Method for determining modes of layered electron-beam wire surfacing for additive technologies). Vestnik MEI, Moscow 5:8–14. https://doi.org/10.24160/1993-6982-2017-5-8-14 7. Ding D, Pan Z, Cuiuri D, Li H (2015) Wire-feed additive manufacturing of metal components: technologies, developments and future interests. Int J Adv Manuf Technol 81(1– 4):465–481. https://doi.org/10.1007/s00170-015-7077-3 8. Taminger KM, Hafley RA, Zalameda JN, Domack CS et al (2017) Process control of electron beam wire additive manufacturing. US Patent Application Publication 2017/0297140 A1, 19 Oct 2017 9. Lastovirya VN, Cherepakhin AA, Rodyakina RV (2018) Determination of functional relation between parameters of weld penetration shape and measurable output variables in electron-beam welding process. Lecture notes in mechanical engineering. Springer, Cham, pp 2443–2452. https://doi.org/10.1007/978-3-319-95630-5_264

Control of Weld Bead Position in Additive

1135

10. Shcherbakov AV, Rodyakina RV, Gaponova DA (2018) Using of smoothed particle hydrodynamics method for constructing a mathematical model of electron-beam surfacing process. Solid State Phenom 284:523–529. https://doi.org/10.4028/www.scientific.net/SSP. 284.523 11. Shcherbakov AV, Martynov VN, Kharitonov IA, Gaponova DA et al (2018) Control of electron-beam surfacing-process parameters using current signals of the wire and article. Russian Electr Eng 89(4):249–254. https://doi.org/10.3103/S1068371218040120 12. Taminger KM, Hafley RA, Martin RE, Hofmeister WH (2013) Closed-loop process control for an electron beam free-form fabrication and deposition processes. US patent 8,452,073 B2, 28 May 2013 13. Egerton RF (2005) Physical principles of electron microscopy. Springer, New York 14. Petrov P, Georgiev C, Petrov G (1998) Experimental investigation of weld pool formation in electron beam welding. Vacuum 51(3):339–343 15. Adam V, Clauß U, Dobeneck DV, Krüssel T et al (2011) Electron beam welding: the fundamentals of a fascinating technology. pro-beam AG & Co. KGaA, Germany 16. Trushnikov D, Krotova E, Koleva E (2016) Use of a secondary current sensor in plasma during electron-beam welding with focus scanning for process control. J Sens 1:302681. https://doi.org/10.1155/2016/5302681 17. Goldstein J, Newbury DE, Echlin P, Joy DC et al (1992) Scanning electron microscopy and X-ray microanalysis. a text for biologists. In: Materials scientists, and geologists. Springer, New York 18. Stecker S, Wollenhaupt PE (2011) Electron beam layer manufacturing using scanning electron monitored closed loop control. US Patent Application publication US 2011/0114839 A1, 19 May 2011 19. Joy DC (1995) Monte Carlo modeling for electron microscopy and microanalysis. Oxford University Press, Oxford (Oxford Series in Optical and Imaging Sciences) 20. Ya G, Rosen D, Staker B (2016) Additive production technologies. Three-dimensional printing, rapid prototyping and direct digital production. Technosphere, Moscow

Operating Efficiency of Worm Gears Under Ultrasonic Vibration Imposition in the Cutting Region S. I. Agapov, Yu. I. Sidyakin and A. F. Tolstyakov

Abstract The paper studies the effect of the imposition of ultrasonic vibrations in the cutting zone during the process of small-module hobs milling in regard to the wear and performance of worm-gear cutters, depending on the design parameters of the wheel. Cutting gears with a worm-modular cutter is a universal method, these operations are the most prolonged in the technological process of manufacturing the gear; therefore, increasing the durability of the cutting tool is an important scientific and production challenge, the solution of which will increase the productivity of the gear-cutting process. Difficulties arising from the operation of worm-modular cutters provoke one into looking for new ways to improve their performance, because the cost of the gear manufacturing includes the cost of a gear-cutting tool, which is 50–60%. The paper is aimed to explain the nature of the wear process of the teeth of a worm-modular cutter, shows the causes of local wear, and proposes fundamentally new ways of machining. The objective of the paper is to determine the stability patterns of the worm-modular cutters in regard to their geometrical parameters by investigating their performance capacity with due regard to both the module and the number of teeth.










Keywords Ultrasonic vibrations Small-module hob milling Wear Worm-gear cutters operating efficiency

1 Introduction The cutting of gears with worm-gear cutters is a generic method. Gear-cutting operations are the longest in the technological process of wheel manufacturing, this is the reason why increasing the durability of the cutting tool is an important scientific and production task, the solution of which will increase the productivity of the cutting process. S. I. Agapov
Yu. I. Sidyakin (&)
A. F. Tolstyakov Volgograd State Technical University, 28, Lenin Avenue, Volgograd 400005, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_121

1137

1138

S. I. Agapov et al.

Worm-gear cutters, as well as other tools used for cutting precision gears, feature a complicated design and are more labor-intensive in manufacturing; moreover, fine-pitch hobs need a smaller number of regrind passes. Complexities that arise in the manufacture and operation of fine-pitch hobs make it necessary to search for new ways to improve their performance because in the structure of the prime cost of manufacturing worm-gear cutters; the cost of gear-cutting tools approximates 50–60% [1]. In worm-gear cutters, when cutting wheels, the teeth are usually cut with two or three cutting edges. It is known that if thickness of the layer being cut is increased, wear of the faces of all cutting tools becomes greater, so one should expect an increased wear of those cutting edges of the teeth of cutters that cut thicker layers, but this does not happen. On the contrary, the edges of the blades which cut off the thinnest layers wear off hardest of all, and, what is noteworthy, such wear is local. It goes mainly along the posterior and anterior edges of small sections of the outgoing blades adjacent to the corners and not along the entire length [2, 3]. This paper explains the nature of the wear of the teeth of worm-gear cutters, shows the causes of the occurrence of local wears, and proposes fundamentally new ways of processing.

2 Experiment Wear of the teeth of standard worm-gear cutters occurs along the back edge of a small section of the outgoing blade, near the corner, or directly at the exit corner. Therefore, the stiffness patterns of mills, depending on various factors, were determined on the basis of wear on these tooth elements. During the experiments, wear on the posterior faces of the teeth of the cutters was measured at regular intervals with the help of the BMI-1 instrumental microscope and the Brinell magnifier with a scale interval of 0.05 mm. Correctly chosen blunting criterion ensures necessary accuracy and roughness of the machined surface of the gears during the period of durability, minimum costs for operating the cutting tool and its maximum service life. The optimal blunting values were taken for the tool of high-speed steel h = 0.15–0.20 mm at m = 0.5 mm, h = 0.2–0.3 mm at m = 0.9 mm and h = 0.3–0.4 mm at m = 1.2 mm. For the carbide tool, the following values were taken: h = 0.1 mm at m = 0.5 mm, h = 0.1– 0.2 mm at m = 0.8 mm, and h = 0.3–0.5 mm at m = 1.2 mm. In the process of research, straight-toothed cylindrical wheels with external teeth from Zk = 24 to Zk = 100 were cut. The studies were carried out on 40X, OXH1M, and bronze BRAVM steel, followed by plotting the resistance of the cutters to various factors in conventional gear cutting and in cutting with introducing ultrasonic vibration into the cutting zone. To determine the effectiveness of gear milling with the use of ultrasonic testing, in comparison with conventional processing, extensive production tests were carried out at FSUE “PO Barrikady”, Volgograd.

Operating Efficiency of Worm Gears Under Ultrasonic …

1139

Fig. 1 Schematic pattern of the wear of a worm-gear cutter

At present, the operability of worm-gear cutter of the middle and large modules has been studied quite thoroughly [4–9]. Investigation of the influence of technological factors on the durability of fine-pitch hobs was carried out in a somewhat limited manner. On the front face along the cutting blades, small holes, the sizes of which are not the same, occur (Fig. 1). The deepest hole appears at the top point BC. It somewhat retreats from the ingoing side CD and approaches the exit AB, thereby reducing the strength of its blade. Dimensions of the hole near the entrance edge are less than at the apex, only the traces of wear are observed at the outgoing blade almost along its entire length. Near the top of the tooth, there is a sharp increase in wear with the formation of a deep hole. The presence of intense abrasion of this section indicates that the formation of chips is difficult in here. The wear of the back faces along the cutting contour also occurs unevenly. The back edge of the blade point width is weak. The width of the chamfer of wear is almost the same, increasing only at the corners. Wear of the back edge of the incoming blade increases to the vertex edge and at the apex of the tooth makes a jump. The increase in wear is caused by an increase in both the thickness of the shear layer and the cutting speed of the most distant points from the axis of the cutter. In absolute terms, the wear of the back edge of the incoming blade is higher than that of the top, although the latter cuts the thicker layers. This is explained by lower values of the rear corners and a higher intensity of deformation of the cut edge by the lateral edge. With the simultaneous operation of several edges, the transition of non-deformed material into chips takes place in two ways: as a result of shear and as a result of joint deformations of shear and crushing. Corresponding to these types of deformations, two types of chips are formed: monolithic chips and chips with a break at adjacent ends.

1140

S. I. Agapov et al.

The back edge of the outgoing blade is most worn out, despite the fact that the thickness is small. Wear is localized near the top in the form of a triangle, one side of which is approximately equal to the width of the hole on the front face. Very often you can see, along the triangle of wear, a deep scratch, formed as a result of the break of the cutting edge. The back edge of the rest of the active part of the blade wears out poorly. A high degree of deformation of the cut edges by the lateral edges of the teeth of the worm-gear cutter is one of the main reasons for the increased wear of their faces, which is facilitated by a smaller value of the rear corners of the side surfaces. Until now, there was an opinion [10–14] that the outgoing edges of gear-cutting tools (worm cutters) cut layers that are less than the radius of rounding of cutting blades. Under this condition, the cutting edge does not cut, but it brushes the metal into the cutting surface, causes considerable pressure from the side of the latter to the back face, which increases its wear, this can be explained by the following reasons: • In cutting a thin layer, a significant length of the outgoing cutting edge is involved, and increased wear should go along its entire active part. In fact, a small area near the top is subject to wear; • The back edges of the outgoing sides of the profiling teeth must also be strongly worn, also cutting very thin layers; • Referring to Fig. 2, the teeth have the greatest wear, which cut off the layers, exceeding the radius of the roundness of the cutting edges by 0.03–0.05 mm. The intensity of wear of the back edges of the corners largely depends on the deformation of the chip material in the border layers, when the slip path depends on the thickness of the cut layers. With an increase in thickness, the contact area of the top chip with the front edge of the tooth becomes larger, and in order to carry out the gathering of the side chippings, it is necessary to crush a large volume of metal adjacent to the corners. Hence, the pressure considerably increases both along the butt end of the chippings and in the areas of the edges adjacent to the top and the outgoing side of the tooth, with all the negative consequences: the temperature increase and the wear of the cutter. The thickness of the cut layer depends on the Fig. 2 Wear of the back edge of the cutter teeth along the loop: (1)—wear of the side outgoing edge h; (2)—wear of the side ingoing edge h1; (3) —wear of the top point—h3

Operating Efficiency of Worm Gears Under Ultrasonic …

1141

Fig. 3 Gear resistibility versus wheel modulus V = 0.66 m/s; S = 0.5 mm/rev; Zc = 36; h = 0.2 mm: 1—conventional cutting; 2—cutting with ultrasonic vibration

feed to the tooth: with increasing feed, the thickness of the cut layer increases, and the deformation in the boundary layers of the chips is substantially increased, and as a result, intensive wear occurs. Teeth are usually cut with three edges: two lateral and top. Figure 3 shows the wear curves along the rear faces of the teeth of the cutter along the loop, constructed for the most common points B, C, and K. Points B and C lie on the side blades, and point K is in the middle of the top (Fig. 1). The graphs show the uneven wear of the teeth along the loop and the wear of the rear edges of the output edges relative to the input and vertex edges. The wear values were measured after 4 h of operation of the 9XC steel cutter 32 mm in diameter, 0.9 mm module, with the number of chip grooves 12 and the profile angle of 20°. The crowns with Zk = 24 from 40X steel were milled at the following regimes: S = 0.5 mm/rev; V = 0.52 m/s. The first teeth of the cutter starting from 1 to 6 are cut with only one input cutting edge. The cutting was free, the deformations of the cut layers were small, and the wear was relatively small. Beginning with tooth 7, three edges were simultaneously involved in the cutting, the teeth cut the layers of the U shape, as a result of which the deformations of the cut layers sharply increased, especially on the lateral edges, and the wear also increased. Especially, it was great at the exit corner. Maximum wear occurred in tooth No. 22, having the greatest arc of contact with the workpiece being processed. Wear on the back edge of the outer corner was twice as high as the inner one. Wear on the back edge of the top point was small; it was three times smaller than the outer corner. Numerous observations of the wear of the teeth of fine-pitch hobs have shown that the wear of the rear faces proceeds as a result of their wear and tear, and the destruction of the areas of blades adjacent to the corners or the corners themselves never occurs. Wear on the trailing edge is uneven. The posterior side of the vertex

1142

S. I. Agapov et al.

margin is least worn out. Wear slightly increases to the corners. The back edge of the outgoing side wears out more. Wear is of a local nature. In worm gears, starting from the module m = 2 mm, the wear of the rear edges of the outgoing side is always accompanied by a breakthrough of the cutting outgoing blades at the corners, and breakthrough is never observed in fine-pitch hobs. Thus, the nature of wear of the teeth of fine-pitch hobs is fundamentally different from that of the cutters of medium and large modules. The incoming edge of the tooth of the cutter operates in lighter conditions. The sheared layer undergoes shear deformation only. The run-out leading to a blade break is not present. The deviation of the runaway of the chips toward the output blade reduces the shear, increases the contact area of the chips with the front edge, moves the hole toward the outgoing edge, thereby increasing the strength of the cutting blade, so the wear of its rear face proceeds more slowly. Intense deformation of the layer cut by the outgoing edge entails considerable plastic and elastic deformation of the layer below the cutting surface. The result of the latter is a high pressure on the back face, especially at the edge area where the crushing occurs, and increased wear of the back edge. The resulting scratch accelerates wear. As for the wear of the back edge of the incoming blade, it was, as can be seen in Fig. 3, half that of the outgoing one, despite the fact that the entrance edge cut the thicker chips and the arc of its contact with the workpiece was larger. The main reason was the absence of warping deformation. The cut-off layer was subjected to shear only; this was indicated by the parallelism of the wear patterns of the holes at the ingoing and vertex edges. They did not have a cut-off chips; they came together in a single stream. In addition, the shear strain was not so great, because the chips were running down the front edge at a large angle to the incoming blade. The posterior edge of the top edge was weaker than all. Its wear was slower four times than the wear of the back edge of the outgoing blade. This is explained by the fact that the top point of the tooth had three to four times larger posterior corners, and the shear strain of the layer cut by it was two or more times less. Based on the analysis of the nature of wear of worm-gear cutters, the following conclusions can be drawn: 1. The wear of the tooth along the cutting perimeter is not the same. The worst side has the back edge of the outgoing cutting edge adjacent to the corner; the smaller is the back edge of the leading edge, and the smallest is the trailing edge of the tooth’s top point. 2. The back edges of the side edges wear out faster than the vertex edge for the following reasons: (a) They have three to four times less rear corners; (b) The layers cut by them have a deformation of two or more times greater than the layer cut by the vertex.

Operating Efficiency of Worm Gears Under Ultrasonic …

1143

A high degree of plastic deformation of the cut layer is accompanied by considerable elastic and plastic deformation of the layer, which results in increased pressures on the back faces and their intensive wear. 3. The main reason for the increased wear of the back edge of the outgoing blade relative to the ingoing edge when forming the chips is the deflection of the outgoing chip together with the apex toward the outgoing edge under the pressure of the chips from the incoming blade. As a result, the shear strain in the incoming chip is increased. 4. In accordance with the formation of two types of chips: monolithic and with a break in the adjacent ends—there are two types of wear of the teeth of worm cutters: without breaking the cutting edges and with a breakthrough. The second kind is most undesirable. In this case, abrasion of the faces proceeds faster; and wears of these teeth limit the resistance of the cutter. In order to answer whether the cutters will work effectively when cutting small-sized wheels from various materials, the studies were carried out on blanks of the most common steel and bronze grades, which differ significantly in their physico-mechanical properties and chemical composition. The indicated brands on physical and mechanical properties and chemical composition corresponded to GOST 1050-74 and GOST 1628-78. In carrying out the experiments, a 10% solution of sulfofrezol was used as a cooling lubricant. Single-pass machining of fine-pitch hobs faces a number of difficulties. One of the effective methods for increasing the resistance of the cutting tool and the quality of the gears is the introduction of ultrasonic testing with the amplitude of 4 … 6 lm into the cutting zone. To determine the blunting criterion, both the dependence of the surface roughness and the accuracy of fine-pitch hobs on the wear value were studied. The roughness of the surface and the accuracy of the treatment of the gear wheels are decisively affected by the wear on the back surface of the tooth of the worm-gear cutter, since it is the area that forms the treated surface, here the greatest specific loads are observed [15–19]. When the wear is h  0.4 mm, vibrations occur (with conventional cutting), leading to rapid chipping of the cutting edges, a sharp increase in the surface roughness and a decrease in the accuracy of the gear wheel. To remove the chips at the cutting edges of the tool, an oversized allowance is required in the recesses, which reduces the already small number of regrinding of fine-pitch hobs. When the ultrasonic vibration is introduced into the cutting zone, in general, the tool wear is determined by the joint action of three main factors: a change in the stress–strain state in the cutting zone, the appearance of high fatigue stresses in local volumes of the tool material, and an increase in the actual cutting speed. Depending on the cutting conditions and the amplitude of ultrasonic vibrations in the cutting zone, the influence of each of these factors can be different. At small ultrasonic amplitudes, the improvement of cutting conditions (reduction of the volume of plastic deformation, reduction of cutting forces and friction coefficient) is of primary importance, which reduces abrasive and adhesive wear and increases the tool’s efficiency in comparison with conventional cutting.

1144

S. I. Agapov et al.

Preliminary experiments have shown that the maximum tool life is achieved at 4 < x < 6 lm, so all further investigations were carried out in this range of amplitudes. At an amplitude of more than 5–6 lm, the change in the loading conditions and the increase in the true cutting speed, which violate the interatomic bonds caused by irreversible distortions of the crystal lattice during the mass emergence of dislocations on the surface, exert a prevailing influence on tool wear. With further operation of the tool and increasing loading cycles, submicroscopic cracks develop up to the size of microcracks, and fatigue failure of the working part of the tool begins [20–23]. Since the investigation of the chip formation process in the formation of monolithic and discontinuous chips showed that when gearing small-module wheels, the stiffness patterns should be different than for the wheels of medium modules, the influence of the feed modules on wear, and durability of cutters during traditional cutting and introduction into the cutting zone of ultrasonic testing was studied carefully. To carry out the experiments, mill modules 0.5 were taken; 0.9; 1.0 mm, there was a passport for each cutter. The cutters of each module were tested at three feeds. The technique of carrying out the stability tests was adopted classical. After a certain amount of time, wear was measured over the rear faces. According to the measurements, wear curves were plotted, from which the working time of the mill was determined to a certain amount of wear, accepted as a criterion for blunting. Then, the experimental data were processed and the dependence of the resistance on various parameters was determined. The dependence T = f(m) for fine-pitch hobs is fundamentally different than for the medium-modulus ones. If the durability of the wheels of medium modules increases with increasing the module, then, when cutting small-sized wheels, on the contrary, it increases, and quite significantly [3]. Our experiments confirm this (see Fig. 4). When monolithic chipping is formed, the sheared layers undergo only shear deformation, the deformation along the thickness of the cut is the same. When monolithic chipping is formed, the wear of teeth occurs as a result of abrasion of their faces. When the chips are discontinuous, wear is always accompanied by a breakthrough of the cutting blade at the crushing site [6]. Let us explain these patterns in more detail. Let us consider a condition of work of a tooth of a standard mill. The tooth works with three edges: a vertex and two lateral edges. Shavings are formed with a break. The deformation along the width is uneven. The total deformation from shear and crushing at the adjacent inner end of the side chipping is great. With the reduction of the module, the thickness of the cut layer, removed by the tooth of the standard cutter, becomes thinner, which reduces the nose and increases the durability of the cutters. When the depth of in-feed is reduced, the collapsing area also decreases. The breakthrough of the cutting blade will be made to a lesser depth, as a result of the durability and for this reason, will grow, and the application of ultrasonic testing increases the durability of small-milling cutters by 2–2.3 times.

Operating Efficiency of Worm Gears Under Ultrasonic …

1145

Fig. 4 Gear resistibility versus number of teeth: 1—conventional cutting m = 1.2 mm; V = 0.83 m/s; S = 0.5 mm/rev; 2—cutting with ultrasonic vibration m = 1.2 mm; V = 0.83 m/s; S = 0.5 mm/rev; 3—conventional cutting m = 0.5 mm; V = 0.66 m/s; S = 0.5 mm/rev; 4— cutting with ultrasonic vibration m = 0.5 mm; V = 0.66 m/s; S = 0.5 mm/rev

The introduction of ultrasonic vibrations into the cutting zone is advisable to apply for gear milling of small-module wheels, when the cutting forces are relatively small and the thickness of the chips is negligible. Great interest in gear cutting deals with studying the durability of worm-module milling cutters with respect to the number of teeth of the wheel. We conducted experiments on cutting gears with the modules of 0.5 and 1.0 mm and with numbers of teeth Zk = 36, 72 and 100. Monograph [21] represents the graphs that characterize the dependence of the durability of the mills on both the module and the number of teeth of the wheel during the conventional and ultrasonic gear milling of tooth gears with the modules of 0.5 and 1.0 mm and with numbers of teeth Zk = 36, 72 and 100. And this paper gives the values of the operating capacity of the tool with an increased Zk (see Figs. 3 and 4). When monolithic chips are formed, the layers being cut are subjected only to shear deformations, but the shear through thickness deformation is the same. As it can be seen from Fig. 4, the working capacity of the tool increases as soon as ultrasonic vibrations are introduced into the cutting zone, and the higher the value Zk, the greater the effect occurs, which is explained by the following: the higher the number of cuts, the lower of the cut layers [22, 23], which results in a decreased build-up.

1146

S. I. Agapov et al.

Tool durability increases with the increased Zk. If Zk = 36, this value equals to 2.0–2.1 times, and in case Zk = 100, this value equals to 2.2–2.3 times. This pattern is also valid when cutting gears with m = 1.0 mm, which can be explained by the fact that a change in the number of teeth of the wheel affects the thickness of the layer being cut by the cutting edge of the tooth of the cutter. As Zk increases, the cut layer becomes smaller, but the thickness variation has an uneven effect on the resistance in cutting thin layers and thick layers. When thin chips are cut, the effect of thickness on resistance is much less than when cutting thick chips. When cutting the teeth of the wheels of small modules, the cutting edges of the cutter teeth cut the micro-chips, so in this case, a change in thickness cannot lead to a significant change in the durability of the cutters.

3 Conclusion 1. With the increase of the module to m = 1 mm, the tool life increases, the use of ultrasonic testing increases the durability of fine-pitch hobs by 2–2.3 times. 2. As the number of teeth of the cut wheel increases, the tool life increases. With an increase in Zk, the efficacy of ultrasound vibration is increased. 3. Usage of ultrasonic vibration reduces wear of fine-pitch hobs and increases their resistance by 1.8–2.2 times in comparison with conventional treatment.

References 1. Suslov AG, Korsakova IM (2008) Tabular method of assigning the parameters of the roughness of the surfaces of machine parts: reference book. J Eng 4:9–14 2. Baldev R, Rangedran V, Padanichami P (2006) The world of physics and engineering. Use of ultrasound. Tekhnosfera, Moscow, p 576 3. Derli AN (2001) Controlling the process of tooth cutting, Reference book. J Eng 6:8–12 4. Medveditskov SN (1974) High-performance gear cutting using new cutting schemes. VPI, Volgograd, p 93 5. Agapov SI (2003) A nature of the wear of fine-pitch hobs. Tekhnologiya mashinostroyeniya 5:24–26 6. Kalashnikov AS (2002) Gear-shaping machines and cutting modes for toothing. J Eng 11: 19–23 7. Astashev VK (2007) On the nonlinear dynamics of ultrasonic technological processes and systems. Bull Sci Tech Dev 2:12–18 8. Prikhodko VM, Kazantsev VF, Nigmetzyanov RI, Fatyukhin DS (2011) High technology in mechanical engineering with the use of ultrasound. High Technol Eng 1:29–37 9. Agapov SI (1986) Intensification of the process of deployment of difficult-to-work materials by introducing ultrasonic vibrations into the cutting zone. Dissertation, Kuibyshev, p 20 10. Agapov SI (2007) Increasing the tool durability in cutting of fine-grained tooth wheels with ultrasonic vibration imposed on the workpiece. Vestnik mashinostroyeniya 2:46–49 11. Agapov SI (2008) Studying wear resistance of fine-grained tooth wheel hobbing with ultrasonic vibration introduced into a cutting zone. Vestnik mashinostroyeniya 4:66–68

Operating Efficiency of Worm Gears Under Ultrasonic …

1147

12. Brzhovskiy BM, Berkenev NV, Zakharov OV, Trofimov DV (2006) Physical bases, technological processes and equipment for ultrasonic material processing: student book. Saratov, SGTU, p 205 13. Kovalnogov VN (2009) Methodology of modeling and the study of thermal interactions of objects that are in contact during mechanical processing in an ultrasonic area. Dissertation, UlGTU, Ulyanovsk, p 32 14. Brzhovskiy BM, Berkenev NV (2006) Ultrasonic technological processes and equipment in machine building and instrument making: student book. Saratov, SGTU, p 348 15. Kiselyov ES (2003) Intensification of machining processes using the energy of the ultrasonic field: Student book. Ulyanovsk, UlGTU, p 186 16. Kravchenko BA (2002) Physical aspects of the theory of metal cutting. Samar. Gos.tekh.un-t, Samara, p 167 17. Polukhin OV, Tarapanov AS, Kharlamov GA (2000) Development and analysis of the mathematical image of the kinematic scheme of cutting teeth with worm-type tools. Reference book. J Eng 8:11–14 18. Agapov SI, Golovkin VV (2010) Increasing the efficiency of machining by applying ultrasonic vibration: Monograph. Izd-vo SNTS, Samara, p 134 19. Agapov SI, Korpelyanskiy OF, Smolnikov NY (2009) Analyzing the possibility of imposing ultrasonic vibration in the cutting zone. Bull Volgograd State Tech Univ 8:47–50 20. Agapov SI (2009) Intensification of the processing of hard-to-machine materials when ultrasonic vibration is introduced into the cutting zone: Monograph. VolgGTU, Volgograd, p 78 21. Agapov SI (2004) Increasing the operating capacity of small-module worm hobs when introducing ultrasonic oscillation in a cutting zone. Bull Samara State Tech Univ 20:43–48 22. Agapov SI (2010) Determining the optimal amplitudes and directions of ultrasound vibrations in cutting small–module gears. Russ Eng Res 30(2):141–143 23. Agapov SI (2008) Hobbing of small-module gears in the presence of ultrasound. Russ Eng Res 28(4):343–345

Features of Formation of Surface Layer Properties in Multistage Processing of Cr–Ni Steel I. V. Firsov, Ju. L. Tchigirinskiy and N. V. Chigirinskaya

Abstract The paper provides the evidence from the experiment in regard to the formation of physical–mechanical characteristics of the surface layer of products from Cr–Ni alloyed steels X45CrNiMoV2-2 during multistage machining with blade and abrasive tools. The following characteristics of the surface layer with changed physical and mechanical properties (work-hardened, deformed or defected layer) are considered: depth and degree of riveting, microhardness distribution along with the layer depth. The differences in the regularities of the formation of the surface layer characteristics in the processing of low-alloy and chromium-nickel steels due to the peculiarities of the thermal characteristics of the processed material, in particular, low thermal conductivity, are shown and justified. The possibility of operational monitoring of the level of residual stresses by the magnitude of the change in the microhardness of the surface is substantiated. Mathematical models, reflecting the correlation between the change in the microhardness of the surface layer and the value of the degree of hardening, are developed.










Keywords Machining operation Routing Surface layer Degree of hardening Residual stresses




1 Introduction To the present date, methods for mapping parts machining routes of a given quality and accuracy are primarily empirical. The systems used in computer-aided design lack the linkage of technologies to such indicators as cost, performance, and more importantly, the likelihood of achieving the specified criteria of accuracy and quality [1–6] of the machined surface.

I. V. Firsov
Ju. L. Tchigirinskiy (&)
N. V. Chigirinskaya Volgograd State Technical University, 28, Lenin Avenue, Volgograd 400005, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_122

1149

1150

I. V. Firsov et al.

The informational basis of modern technological routing consists of data arrays developed on the basis of generally accepted reference information. The data given in common publications [7–12] are generally devoted to the most common construction materials. The individualization is given in general and refers to non-ferrous metals, cast irons, carbon, and alloyed steels. This classification does not allow obtaining reliable solutions to technological design problems, which are adapted to the conditions of production at specific enterprises. In the design documentation, the concept of “product quality” is usually limited to indicating the degree of accuracy considering the size and one of the height parameters of the surface roughness. The physical–mechanical characteristics of the surface layer, which directly determine the operational properties of products in general engineering, are extremely rarely standardized and are practically not mentioned in the reference documentation of technological routing (Table 1). The given papers estimate the degree of cold work by using various techniques. Paper [12] defines the degree of cold work as a relative increment of microhardness of the machined surface layer (1) to microhardness of the basic material; paper [11] uses Formula (2).

Table 1 Data on depth and degree of hardening from reference books Machining method (machining stage)

Rough milling Finish milling Fine milling Finish slotting Rough planing Finish planing Rough turning Semi-rough turning Finish turning Fine turning Semi-rough grinding Finish grinding Fine grinding Superfinishing Regular lapping

Dalskiy [7]

Suslov [12]

Hardened layer depth (h, mm) Min Max

Hardened layer depth (h, mm) Min Max

Degree of hardening (U, %) Min Max

Kosilova [11] recalculated due to [12] Degree of Hardened hardening layer depth (U, %) (h, mm) Min Max Min Max

0.200 0.100 – – – – 0.200 –

0.500 0.100 – – – – 0.500 –

0.200 0.080 0.050 – 0.250 0.150 0.200 –

0.50 0.15 0.10 – 0.6 0.30 0.45 –

20 10 0 – 20 10 10 –

30 40 20 – 40 30 40 –

– 40 – 60 – – – 20

– 60 – 100 – – – 50

– 0.04 – 0.12 – – – 0.03

– 0.10 – 0.15 – – – 0.15

0.020 – 0.015

0.080 – 0.030

0.050 0.020 0.030

0.20 0.05 0.05

10 0 10

30 20 40

– 40 –

– 80 –

0.02 –

0.06 –

– – – –

– – – –

0.015 0.010 0.005 0.005

0.03 0.02 0.01 0.01

0 0 0 0

30 20 20 20

40 – – –

60 – – –

0.016 – – –

0.035 – – –

Features of Formation of Surface Layer Properties in Multistage …

Hlmach  Hlinit  100%; Hlinit

ð1Þ

Hlmach  100% ¼ U þ 100%; Hlinit

ð2Þ

U¼ U0 ¼

1151

where Hµmach—microhardness of the surface layer after machining; Hµinit—initial microhardness of the material being machined. The ranges of the potential capacities of the methods for machining are combined due to the depth of cold work and in accordance with the principle of defining average values for both the left and right border [1, 5]. The verified data are given in Table 2.

2 Techniques It is essential to note that the data given in reference books [11, 12] and other reference outlets are obtained for non-alloyed and low-alloyed constructional steels [13–18]. In order to receive more precise evidence and to specify their applicability in machining steel X45CrNiMoV2-2, several experiments have been conducted for machining plane surfaces (Table 3). Table 2 Updated reference data about the methods of processing planes on physical and mechanical parameters Machining method (machining pass) Rough milling Finish milling Fine milling Finish slotting Rough planing Finish planing Rough turning Semi-rough turning Finish turning Fine turning Semi-rough grinding Finish grinding Fine grinding Superfinishing Regular lapping

Hardened layer depth (h, mm)

Degree of hardening (U, %)

Min

Max

Av.

Range

Min

Max

Av.

Range

0.200 0.070 0.050 0.120 0.250 0.150 0.200 0.030 0.035 0.020 0.020 0.015 0.010 0.005 0.005

0.500 0.120 0.100 0.150 0.600 0.300 0.480 0.050 0.140 0.060 0.040 0.033 0.020 0.010 0.010

0.350 0.095 0.075 0.185 0.425 0.225 0.340 0.040 0.088 0.040 0.030 0.024 0.015 0.008 0.008

0.300 0.050 0.050 0.030 0.350 0.150 0.280 0.020 0.105 0.040 0.020 0.018 0.010 0.005 0.005

20 10 0 – 20 10 10 – 10 0 10 0 0 0 0

30 40 20 – 40 30 40 – 30 20 40 30 20 20 20

25 25 10 – 30 20 25 – 20 10 25 15 10 10 10

10 30 20 – 20 20 30 – 20 20 30 30 20 20 20

1152

I. V. Firsov et al.

Table 3 Condition of the surface layer after machining steels X45CrNiMoV2-2 No.

Machining method (machining pass)

Microhardness of the surface layer, HV Min Max

Hardened layer depth (h, mm) Min

Max

Degree of hardening (U, %) Min Max

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16

Bandsawing Rough milling Finish milling Fine milling Rough slotting Finish slotting Rough planing Finish planing Rough turning Semi-rough turning Finish turning Fine turning Semi-rough grinding Finish grinding Fine grinding EDM

415 353.6 348.6 352.3 346.1 348.3 355.1 355.3 350.1 353.8 355.8 351.6 346.2 348.1 350 351.5

0.19 0.145 0.085 0.042 0.165 0.101 0.115 0.075 0.102 0.075 0.047 0.035 0.04 0.028 0.026 0.042

0.35 0.175 0.105 0.055 0.215 0.148 0.157 0.111 0.133 0.099 0.062 0.044 0.054 0.038 0.03 0.063

28 21 11 8 36 16 28 14 22 8 6 2 12 5 2 8

450 450.1 407.6 391.1 495.4 425.6 482.7 425.6 444.1 406.6 386.8 372.6 410.6 375 364.5 395.1

47 27 17 10 43 20 36 20 25 15 9 5 17 7 4 12

Initial hardness of the material (core hardness) was equal to 340–360 HV [19]—hardened steel in original condition. Machining was done in the modes recommended by Suslov and Dalskiy [11] for each of the studied types of machining. The depth of cold work was determined due to the results of the measurements of microhardness on the angle lap sections. Residual stresses of the second order were defined by using a DRON-3 X-ray diffractometer (the Bragg–Brentano focusing scheme). Univariable regression analysis was utilized with the purpose of analyzing the experimental data obtained and of building the models [20].

3 Depth and Degree of Hardening on the Surface Layer Analysis of the refined data obtained from reference books, in combination with the evidence from the experiments, showed significant discrepancies in the analyzed parameters.

Features of Formation of Surface Layer Properties in Multistage …

1153

Fig. 1 Hardened layer properties after machining via various technological methods: a depth; b degree of hardening

The tolerance on hardened depth for rough machining methods aggregated from the reference data is displaced upwards with respect to the average values obtained experimentally (Fig. 1a). This is due to the fact that the workpiece features a high hardness (340–360 HV). Experimental averages for finishing and final machining techniques are within the allowable range. As a rule, such a discrepancy between rough and finishing methods can be observed if the material being machined demonstrates significant heterogeneity and uneven distribution of physical and mechanical properties in the cross section. Practically all experimental values characterizing the degree of hardening on the deformed layer fit into the range determined with the help of refined reference data (Fig. 1b). It should be noted that the values obtained are valid for all materials having a similar chemical composition subject to the machining conditions and in accordance with cutting modes. When machining hardened chrome-nickel steel, the higher feed rates, the greater the depth of the deformed layer, and the higher the cutting speed, the shallower the depth [8, 10].

4 Residual Stress in the Surface Layer The hardening of the layer of plastically deformed metal adjacent to the machined surface can also be estimated in regard to the intensity of residual stresses. When machining a metal, the surface layer of the part experiences either tensile or compressive microstresses, which can be investigated by studying the X-ray diffraction pattern. During the turning operation on hardened materials, an increase in the surface layer temperature associated with an increase in cutting speeds can cause metal tempering and a decrease in its specific volume, which leads to a decrease in residual stresses, thus provoking a decrease in compressive and even tensile stresses. It is also known that when machining low ductile hardened materials, an

1154

I. V. Firsov et al.

increase in feed leads to a significant increase in temperature in the cutting zone, which can cause thermal tensile stresses reducing the overall amount of residual compressive stresses in the surface layer [9]. The sign and depth of compressive stresses resulting from the phase transformations of the metal on its surface layer are determined by the completeness of phase transformations and the ratio of the specific volumes of the structural components of the metal layers adjacent to the surface layer. At the same time, the chemical composition of the metal and its ability to structural changes, plasticity, elasticity, thermal conductivity, thermal diffusivity, and other mechanical and physical properties of the material being machined are of great importance for the formation of residual stresses. Table 4 gives the reflections recorded with the help of the diffractometer, according to the qualitative test of the characteristics including the intrinsic broadening of X-ray lines, the crystal size and the elastic strain of the crystal lattice, and reflections which were recorded. The analysis of the data in Table 4 shows that the experimental intensity of the residual stresses formed in the surface layer exceeds the reference data by about

Table 4 Reference and experimental data on residual stresses after the machining of plain surfaces Machining method

Residual stresses, ±rres, MPa Reference data [12]

Bandsawing Rough milling Finish milling Fine milling Rough slotting Finish slotting Rough planing Finish planing Rough turning Semi-rough turning Finish turning Fine turning Semi-rough grinding Finish grinding Fine grinding EDM Superfinishing Regular lapping

Experimental data

Min

Max

Av.

Range

Min

Max

Av.

Range

– 250 200 100 – – 250 200 200 – 150 100 200 300 200 – 100 100

– 300 250 200 – – 350 250 300 – 200 150 400 400 500 – 200 250

– 275 225 150 – – 300 225 250 – 175 125 300 350 350 – 150 175

–

600 490 360 320 530 360 530 325 430 285 280 235 680 420 260 340 – –

710 550 410 360 610 420 630 375 500 335 320 265 760 480 300 390 – –

655 520 385 340 570 390 580 350 465 310 300 250 720 450 280 365 – –

50 60 50 40 80 60 100 50 70 50 40 30 80 60 40 50 – –

50 50 100 – – 100 50 100 – 50 50 200 100 300 – 100 150

Features of Formation of Surface Layer Properties in Multistage …

1155

2 times. This discrepancy is due to the fact that the reference database was aggregated for low-carbon steels, while the experimental studies were carried out to assess complex-alloyed hard-to-work steels in the hardened state. At the same time, the trend of the changes in the average intensity of the residual stresses for the reference and experimental data is the same (Fig. 2a). The residual stresses obtained in the experiment on steel X45CrNiMoV2-2 are compressive. The evidence from studying the semi-rough and finish passes during the grinding operation showed a high fine structure, which evidenced both the presence of small blocks and a rather high level of microstresses common to finishing operations. In case of finishing operations conducted at low speeds and under low pressures, the metal is practically unheated, which proves that these operations cannot be the reason for the formation of residual stresses. The results of the experimental studies (Fig. 2b) prompt suggestions that there are relationships between the parameters determining the physical–mechanical condition of the defective layer, i.e., intensity of residual stresses (r) and a degree of hardening (U). r  10:6
U þ 212:2; r  148:6
U 0:6 ;

R2 ¼ 0:92; R2 ¼ 0:95;

d  9% d  3%

ð3Þ ð4Þ

In the array of the results obtained, we select two groups of technological methods: edge cutting (Fig. 3a) and abrasive machining (Fig. 3b). This separation is due to a significant increase in residual stresses in the surface layer (which is typical for the machining of hardened chromium-nickel steels) during grinding operations in comparison with the finishing operations of edge cutting. The analysis of the models obtained confirms the assumption of a correlation (3, 4) between the value of microscopic (second-order) stresses and the degree of hardening on the surface layer.

Fig. 2 Residual stresses of defect layer: a comparing for the reference and experimental data, b correlation between residual stresses and degree of hardening

1156

I. V. Firsov et al.

Fig. 3 Correlation between residual stresses and a degree of hardening on the surface layer: a during edge cutting machining; b during abrasive machining

5 Conclusions The implementation of the regression models showing the relationship between the intensity of residual stresses and a degree of cold work on the surface layer gives the possibility to control stresses in the surface of the product being machined instead of using labor-intensive and expensive X-ray techniques, which are inapplicable under operating conditions. Data on building models must be obtained under specific manufacturing conditions and for specific constructional materials. Only in this case, rather high fidelity can be guaranteed for the forecasted manufacturing results. Both residual compressive stresses in the surface layer and its microhardness contribute to better operational and wear resistance properties of machines and products. The negative effect of surface hardening is associated with the brittle behavior of parts’ face, a tendency to deformation and cracks formation during subsequent heat treatment, as well as possible increased tool wear in subsequent passes during machining the same workpiece.

References 1. Firsov IV (2015) Computer-aided planning of parts machining in interrupted edge cutting of chromic steels and aluminum alloys. In: Issues on ensuring and increasing product quality and competitive power in engineering and aircraft-engine building (TM-2015). In: Proceedings of the 7th international scientific technical conference, Bryansk, pp 200–201 2. Haas M, Chang T-C (1987) A survey on the usage of computer aided process planning systems in industry. Engineering Research Center on Intelligent Manufacturing Systems, Purdue University, W. Lafayette 3. Houdek F (2012) Improving requirements engineering processes impressions during one decade of improvement at Daimler. In: Lecture notes in computer science, vol 7343, pp 1–2 4. Kanuminy M, Chang T-C (1991) Process planning in a completely automated manufacturing environment. J Des Manuf 1(1):7–15

Features of Formation of Surface Layer Properties in Multistage …

1157

5. Tchigirinsky JL (2014) Ensuring the accuracy and quality of surfaces in multi-stage machining based on the improvement of information and mathematical means of the CAPP designing subsystem. Dissertation, Saratov State Technical University 6. Tchigirinsky JL et al (2016) Structural optimization of technological route using simulation modeling. In: 2015 International conference on mechanical engineering, automation and control systems (MEACS), 1–4 Dec 2015, Tomsk, Russia. https://doi.org/10.1109/meacs. 2015.7414925 7. Dalskiy AM, Bazrov BM, Vasilyev AS, et al (2000) Technological heredity in engineering. In: Dalskiy AM (Eds.), MAI, Moscow 8. Granovskiy GI, Granovskiy VG (1985) Metal cutting: student’s book for engineering and tool manufacturing. Vysshaya Shkola, Moscow 9. Matalin AA (1977) Machining technologies. Mashinostroyeniye, Leningrad 10. Matalin AA (1985) Engineering technologies. Mashinostroyeniye, Leningrad 11. Reference guide for engineers (2001) Book 1 (2). In: Kosilova AG, Meshcheryakov RK (eds), Mashinostroyeniye, Moskow 12. Suslov AG, Dalskiy AM (2002) Scientific basics in engineering technologies. Mashinostroyeniye, Moscow 13. El-Khabeery MM, Fattouh M (1989) Residual stress distribution caused by milling. Int. Journal of Machine Tools and Manufacture 29(3):391–401 14. Firsov IV, Tchigirinsky JL (2016) The condition of the surface layer of the part after rough blade machining of steel X45CrNiMoV2–2. Bull VSTU. Ser Progress Technol Mech Eng 8 (187):55–59 15. Miko E, Nowakowski Ł (2012) Analysis and verification of surface roughness constitution model after machining process. Procedia Eng 39:395–404 16. Tchigirinsky JL, Trong NQ (2015) Phenomenological regularities of formation the treatment quality with milling. In: Issues on ensuring and increasing product quality and competitive power in engineering and aircraft-engine building (TM-2015), Proceedings of the 7th international scientific technical conference, Bryansk. pp. 211–212 17. Tchigirinsky JL, Trong NQ (2017) Influence of technological factors of blade processing on the forming of the defect layer. In: MATEC web of conferences. International conference on modern trends in manufacturing technologies and equipment (ICMTMTE 2017), September 11–15 2017 (vol 129). Sevastopol State University, National University of Science and Technology “MISIS”, Sevastopol, Russia, Polzunov Altai State Technical University, Inlink Ltd. and International Union of Machine Builders. https://doi.org/10.1051/matecconf/ 201712901020 18. Tchigirinsky JL, Firsov IV, Nesterenko PS (2017) Influence of finishing methods of treatment on condition of steel 36NiCrMo16 part surface layer. In: Procedia Engineering. International conference on industrial engineering (ICIE 2017) (vol 206). Peter the Great Saint-Petersburg Polytechnic University, South Ural State University (national research university), Platov South-Russian State Polytechnic University and Far Eastern Federal University. https://doi. org/10.1016/j.proeng.2017.10.627 19. GOST 2999-75 (2018) Metals and alloys. Vickers hardness measurement method. Izdatelstvo standartov, Moscow 20. Chigirinskaya NV, Tchigirinsky JuL, Gorobtsov AS (2015) Experiment planning in technical and economical tasks: student’s book. VolgGTU, Volgograd

Contactless Monitoring of Processed Surface Microrelief at Manufacturing Environment A. P. Gontar, S. V. Mednikov and N. V. Chigirinskaya

Abstract The article considers the possibility of non-contact assessment of the surface condition obtained as a result of machining. Non-contact methods of control allow, firstly, ensuring the safety of the surface after finishing methods of processing and, secondly, reducing the time spent on control operations. The applicability of optical properties of a rough surface to estimate the height of the microrelief is substantiated. The design scheme of the laboratory installation for the study of the possibilities of non-contact assessment of the micro-irregularities’ height parameters was made on the basis of a prototype device. Experimental studies of the proposed method for assessing the quality of the treated surface were conducted. The technique of statistical processing of measurement results and construction of calibration dependence is described. The convergence of the results of measurements of the height of arithmetic average of the microprofile (Ra) with the reference roughness samples for blade processing methods is proved.







Keywords Surface microprofile Assessing microrelief Surface roughness Contactless assessment Specular reflection Diffusion reflection










1 Introduction One of the main properties of the quality of machining products deals with the condition of microrelief of their working area, since the behavior of microgeometry deviations of the working faces defines the fault tolerance and operating life of these products during operation. In particular, the assignment of optimal initial characteristics of the microrelief of rubbing surfaces makes it possible to reduce their wear during their breaking-in by 2–5 times and to shorten the breaking-in time up to several minutes.

A. P. Gontar (&)
S. V. Mednikov
N. V. Chigirinskaya Volgograd State Technical University, 28, Lenin Av., Volgograd 400005, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_123

1159

1160

A. P. Gontar et al.

This research is devoted to solving the problem of creating modern systems of technical quality control and of building methodological basics for investigating the microgeometric deviations of various surfaces of machine parts under production conditions via the construction of an optical–electronic [1, 2] informationmeasuring system (IMS).

2 Methods of Measurement and Evaluation of Surface MicroRelief The solution to this problem puts forward a number of requirements that must be met by modern mobile devices for obtaining measurement information on the microrelief of the surface under study. These measuring tools should be small-sized, implement a contactless method of obtaining measurement information directly in production and in real time, provide digital processing of the information received, and be able to evaluate the microrelief of surfaces having complex shapes, such as the internal cavities of small diameter holes, die faces and molds. In addition, these measuring tools must have metrological characteristics that are invariant with respect to changing production conditions and must provide an energy-saving mode of using these tools. Analysis of the existing (Fig. 1) measuring methods and devices for this purpose has also showed that they cannot fully meet the requirements formulated

Methods of data obtaining and processing about deviations microgeometry of the surface

Contactless

Tactile methods

Contact

Interference method

Light section, shadow projection and raster method

Digital image processing

Indirect

Direct

Replica method

Measuring stylus

Reflectometry

Fig. 1 Methods for controlling surface microrelief in metalworking manufacturing

Contactless Monitoring of Processed Surface Microrelief …

1161

hereinabove. For example, using the stylus method and profilometers and profile testing instruments, it is impossible to quickly measure the roughness of the surface, which features complex shapes [3–5]. Contactless optical methods (Fig. 1)—i.e., interferential method, reflectometry, light section techniques, and shadow method —require the use of specialized microscopes and are utilized only under laboratory conditions, and the roughness measurement on their basis is characterized by large time expenditures. Reflectometric methods and devices require the stability of the incident light on the surface under test; this leads to a significant increase in their dimensions [6–12]. Such methods, as well as the optical tools discussed hereinbefore, cannot be used to assess surface roughness in hard-to-reach areas of industrial products.

3 Technique of Contactless Evaluation Microrelief Parameters The optical roughness properties [3, 13–15] of a surface are functions of its statistical parameters including the standard deviation of the heights of surface microprofile points from the centerline r, the correlation interval a, and the average angle of inclination of the microfaces on the rough surface c. These parameters determine the intensity of the reflected or scattered light and its spatial distribution. In practical use of rough surfaces, it is necessary to know the numerical values of these characteristics. The phenomenon of light reflection from a rough surface is explained as Fresnel [13, 15–24] reflection from microfaces in terms of geometrical optics. It is believed that the distribution of reflected light in space is determined by the orientation of these microfaces in relation to the incident rays. In order to isolate specular reflection in its pure form, it is necessary: (1) to have a rather rough machined surface; (2) to examine a surface with a high reflection coefficient, (3) to use a device for recording indicator functions, with high angular resolution [13]. Specular reflection from a smooth surface is always accompanied by diffuse reflection. The function expressing the dependence of the intensity of the diffuse component on the height of microroughness of the surface is more sensitive [13] than the function for the specular reflection component; therefore, the research is focused on the diffuse component of the reflected flux. The illumination source (Fig. 2) is semiconductor laser 1 with a wavelength of 460 nm, which alternately illuminates [25] sample 2 and reference sample 3 at an angle of 8°. It should be noted that the angle (8°) should not affect the measurement results, since the function of the reflection angle X (h) (2) does not depend on the angle of illumination. Reference sample 3 is a verified sample of the roughness of a milled surface having a roughness of Ra = 0.8 lm and Ra = 0.4 lm. The reflected light passes through lens 5, frosted lenses 6, and falls on the semiconductor photocell 7, the output signal of which is amplified.

1162

A. P. Gontar et al.

Fig. 2 Assessment of surface roughness properties: a measuring schemes [13]; b functional scheme of the device for contactless assessment of surface microprofile height

Light reflected diffusely is recorded by a receiver with an aperture of 8 º by means of photomultiplier 4 within the angles ranged 20–75 [25]. These angles are provided by a sinus plate and a rotary table, on which the photomultiplier is fixed. The output signals of receivers 4 and 7 are fed to the input of the microcontroller, and their ratio is recorded as a function of the reflection angle X (h) (2). The experiment has revealed several factors that significantly affect the measurement results: • Optical encapsulation of the device—extra light sources, which give additional illumination of the photoresistor, significantly reduce the accuracy and stability of measurements. • Coaxiality of the photoresistor and the laser—the errors of the relative position of the light source (laser) and the receiver (photoresistor that records the beam reflected from the surface) reduce the accuracy of measurements. To eliminate these factors, a second scheme has been developed (Fig. 2b). The semiconductor laser and the photoresistor are housed in a metal case. Angle 2h equals 18° due to constructive reasons. Light enters the verified surface roughness sample and, after being reflected diffusely, is detected by the photoresistor. The output signal is fed to the input of the multimeter; the ratio of the output and input signals is recorded as an estimate of the roughness of the test surface.

Contactless Monitoring of Processed Surface Microrelief …

1163

Investigations of the method for the contactless assessment of microprofile height were carried out on the verified surface roughness samples in terms of the following types of treatment: cylindrical milling, face milling, face grinding, and turning. By the example of the verified roughness samples (during face milling), the minimum number of observations was revealed, namely over what time the multimeter could show reliable readings. The stability of the experiments was assessed in regard to the change in the standard deviation of the measurement results progressively as the experiments were repeated. The significance of the change in the average value and the standard deviation was assessed in terms of the values of Student’s and Fisher’s criteria [26]. The hypothesis of an insignificant change in the standard mean-square deviation, with an increase in the number of experiments, was considered as confirmed with the confidence a = 0.95. Table 1 shows an example of calculating the minimum number of tests in an experiment for a verified roughness sample Ra = 0.4 lm with a confidence probability a = 0.95. As in the case of calculation due to the Student criterion, the minimum number of test repetitions cannot be less than three; this is due to the peculiarities of the calculation of the standard deviation. Since already in the seventh line of the calculation table (Table 1), the table values of the Student and Fisher criteria exceed the calculated ones, and we conclude on the stability of the standard deviation, which allows for the evidence that seven repetitions are enough to complete this experiment. Within a 30 s interval, 170 values were obtained, thus proving that reliable research values could be obtained in a split second, which was essential for the rapid measurement of the surface microprofile. Experimental studies carried out on verified roughness samples showed that there was a stable correlation dependence—the correlation coefficient was in the range 0.88–0.95—between the readings of the instrument recording the change in photoresistor resistance and the height of the microrelief, estimated in terms of the value of the arithmetic average of the microprofile Ra. To assess the surface microprofile, a calibration curve was built (Fig. 3), which established a relationship between the instrument readings R (X) and verified surface roughness samples Ra (lm). Relative value of the error is ±10%.

4 Conclusion Production testing of the method for the contactless control of the quality of the treated surface was carried out at the production plant PA “Akvamash” (Volgograd) during the production of large-sized products of the “Plate” type. The quality (Ra = 0.63; 3.2 lm) of plain and cylindrical surfaces obtained as a result of boring, milling, and grinding was evaluated in two ways:

Measurement

2.719 2.694 2.685 2.678 2.679 2.682 2.686 2.688 2.689 2.692

No.

1 2 3 4 5 6 7 8 9 10

2.719 2.707 2.699 2.694 2.691 2.690 2.689 2.689 2.689 2.689

Average

0.000 0.013 0.014 0.016 0.015 0.014 0.013 0.012 0.012 0.011

Standard n−1

Student’s table

0.588 0.380 0.198 0.101 0.036 0.009 0.001 0.027

0.071 0.068 0.067 0.066 0.065 0.065 0.065 0.064

Not calculated

Student’s test

0.616 0.729 0.852 0.923 0.972 0.993 0.999 0.978

Student’s probability

Continued Continued Continued Continued Correct Correct Correct Correct

Student’s result

1.324 1.162 1.053 1.133 1.157 1.142 1.125 1.103

Fisher’s test

Table 1 Calculating the minimum number of value selections due to Student’s and Fisher’s criteria

2 3 3 4 5 6 7 8

m

1 2 4 5 6 7 8 9

k

199.5 19.16 6.591 5.192 4.387 3.866 3.500 3.230

Fisher’s table

0.476 0.507 0.539 0.564 0.575 0.573 0.568 0.561

Fisher’s probability

Correct Correct Correct Correct Correct Correct Correct Correct

Fisher’s result

Continued Continued Continued Continued Correct Correct Correct Correct

General result

1164 A. P. Gontar et al.

Contactless Monitoring of Processed Surface Microrelief …

1165

Fig. 3 Correlation dependence Ra = f(R)

• contactless—as described hereinabove; • conventional—using a profilograph—profilometer. Assessment of the quality of the surface machined via the contactless method was performed directly on the machine, without unfastening and reinstalling the workpiece. The compared methods showed satisfactory convergence of the results. Cases of unreasonable bad quality were not identified. It was established that the time spent on carrying out the quality control of the machining process was significantly reduced. Thus, it is a fair assumption to say that the considered contactless optical method for estimating the height of the microrelief is applicable under production conditions.

References 1. Gizenburg VL, Motulevich GP (1955) Optical properties of metals. Prog Phys Sci (Uspekhi fizicheskikh nauk) 4:469–535 2. Shester AA (1935) Basics of theoretical optics. ONTI, p 19 3. Hariharan P, Oreb BF, Eiju T (1987) Digital phase-shifting interferometry: a simple error-compensating phase calculation algorithm. Appl Opt 26:2504–2506 4. Fleischer Matthias, Windecker Robert, Tiziani Hans J (2000) Fast algorithms for data reduction in modern optical three-dimensional profile measurement systems with MMX technology. Appl Opt 39(8):2019–2023 5. Kino G, Chim SM (1990) Correlation microscope, vol 29

1166

A. P. Gontar et al.

6. Hildebrand BP, Gordon RL, Allen EV (1974) Instrument for measuring the roughness of supersmooth surface. Appl Opt 13(1):177–180 7. Kubo M (1967) Statistical analysis of surface roughness waveforms. Ann CIRP 14(3) 8. Tatrov A et al (2002) Rigorous coupled-wave analysis calculus of submicrometer interference pattern and resolving edge position versus signal-to-noise ratio 41(8):l886–1892 9. Sinclair Michael B, de Boer Maarten P, Corwin Alex D (2005) Long-working-distance incoherent-light interference microscope. Appl Opt 44(36):7714–7721 10. Wyant James C (2007) Improved interferometric optical testing. Opt Photonics News 18 (7):32–37 11. Zhu Y, Gemma T (2001) Method for designing error-compensating phase-calculation algorithms for phase-shifting interferometry. Appl Opt 40:4540–4546 12. Sysoev EV (2010) Nanorelief measurements errors for a white-light interferometer with chromatic aberrations. Key Eng Mater pp 51–55 13. Toporets AS (1968) Specular reflection of light from a rough surface. Opt Spectrosc 1: 159–179 14. Manfred A, Biondi M (1954) Physica 96:534 15. Hybl O, Berger A, Hausler G (2008) Information efficient white-light interferometry. In: DGaO-Proceedings. http://www.dgao-proceedings.de 16. Tchigirinsky JuL, Gontar AP (2016) Possibility of evaluation the geometrical parameters of surface microrelief with contactless method. Bull Volgograd State Tech Univ. Ser Progress Technol Mech Eng 14(193):40–42 17. Pavel Pavlicek, Ondrej Hybl (2008) White-light interferometry on rough surfaces-measurement uncertainty caused by surface roughness. Appl Opt 47(16):2941–2949 18. Morgan CJ (1982) Least squares estimation in phase-measurement interferometry. Opt Lett 368–370 19. Osten W (2005) White-light interferometry with higher accuracy and more speed. Fringe 605–612 20. Jan Niehues, Peter Lehmann, Klaus Bobey (2007) Dual-wavelength vertical scanning low-coherence interferometric microscope. Appl Opt 46(29):7141–7148 21. Pavel Pavlicek, Jan Soubusta (2004) Measurement of the influence of dispersion on white-light interferometry. Appl Opt 43(4):766–770 22. Biegen James F (1989) Calibration requirements for Mirau and Linnik microscope interferometers. Appl Opt 28(11):1972–1974 23. Wyant J (2002) White light interferometry. In: Proceedings of SPIE, vol 4737 24. Andreev V, Indukaev K (2003) The problem of subrayleigh resolution in interference microscopy. J Russ Laser Res 3(24):26–32 25. Beckman P (1965) The scattering of electromagnetic waves from rough surfaces. Pergamon Press, Oxford, p 492 26. Chigirinskaya NV, Tchigirinsky JuL, Gorobtsov AS (2015) Experiment planning in technical and economical tasks: student’s book. VolgGTU, Volgograd

Characteristics, Composition, Mechanisms of Function and Modern Aspects of Implementation of Digital Production Systems in Mechanical Engineering Industry A. R. Ingemansson

Abstract It is pointed out that the implementation of digital production systems (DPS) is the urgent way of increase of efficiency of the mechanical engineering industry. DPS in mechanical engineering are the production systems, which are based on the integration of modern computing technologies, science-intensive equipment and technologies. The increase of efficiency of mechanical engineering by means of implementation of DPS firstly must be realized in the metalworking industry. DPS for metalworking industry consist of technological, measuring, transport and storage equipment. The composition and mechanisms of the function of DPS for metalworking are presented, and the directions of the increase of efficiency of production technologies in metalworking are investigated. It is carried out that the implementation of DPS provides the efficiency increase of machining production by the decrease of time, spend on preparing operations on MC; increase of cutting process effectiveness and reliability; increase of machining operations quality; increase of cutting tools usage effectiveness and increase of machine tools usage effectiveness. The aspects of the implementation of DPS in view of modern level of technological, inspection, transport and storage equipment are analyzed. Keywords Digital production


Metalworking industry
Machining operations

1 Introduction The development of science-intensive solutions for the implementation of digital production systems (DPS) is the urgent way of increase of efficiency of the mechanical engineering industry. DPS in mechanical engineering are the production systems, which are based on the integration of modern computing

A. R. Ingemansson (&) Volgograd State Technical University, 28, Lenin Avenue, Volgograd 400005, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_124

1167

1168

A. R. Ingemansson

technologies, science-intensive equipment and technologies. Implementation of DPS in the industry allows one to increase flexibility, reliability, ecological properties and economic output of production processes [1–6].

2 Relevance DPS functioning is connected with the following aspects: • Digital designing, technological preparation of production process and product lifecycle management (CAD-, CAM-, PLM-systems); • Numerical control (NC-systems) for equipment, industrial sensors and controllers, industrial robotics; • Industrial Internet of Things (IIoT); • Processes with big data arrays; • Informational systems of enterprise management (ERP-, MES-systems). DPS use as subtractive technologies, which are based on shaping principle by removing some volume of material from workpiece, so DPS use and additive technologies, which are based on phased grooving of parts. DPS form the «factories of future» («digital», «smart», «virtual» factories). Creation of such factories is the topical subject of the road-map «Technet» of Russian National Technological Initiative up to 2035. Similar initiatives for the implementation of DPS in industry, in part in mechanical engineering, are developing in other countries. Among them are the following projects: «Industrie 4.0» (Germany), «Advanced Manufacturing Initiative» (USA), «Made in China 2025» (China). The efficiency increase of mechanical engineering production due to implementation of DPS is realized by the following directions: • Decrease of expenses (up to 50%) for production due to the implementation of advanced CAD technologies, technological preparation of production process, enterprise management, science-intensive equipment and technologies; • Decrease of time space between designing and manufacturing of product; • Increase of production division flexibility and the opportunity of unmanned technology; • Increase of responsiveness and objectivity of information about the production process; • The efficiency increase of use and maintenance of technological equipment by the analysis of objective information about its condition; • The efficiency increase or quality management by the analysis of objective information at the stages of starting production, production process and quality control of finished products. The development of science-intensive solutions for the implementation of DPS is the urgent way of increase of efficiency of the mechanical engineering industry.

Characteristics, Composition, Mechanisms of Function …

1169

The benefits from the implementation of DPS should be firstly realized in the metalworking industry.

3 Setting of the Problem The subject of development of science-intensive solutions for the implementation of DPS in the metalworking industry is a complicated and complex task. To solve, it is necessary to determine and design composition and functional mechanism of DPS and the ways of technological processes efficiency increase.

4 The Theoretical Part DPS for metalworking consist of technological, measuring, transport and storage equipment, see Fig. 1 [7].

Fig. 1 Composition of DPS for metalworking industry

1170

A. R. Ingemansson

Technological equipment is represented by NC-machines and machining centers (MC) with contact and less contact devices for control and measuring of working parts and control of tool wear. Measuring equipment is represented by coordinate-measuring machines and the devices for control of workpiece material (portative and stationery analizators of chemical composition and hardness measurement). It is needed to stress the attention on presetters for cutting tools. Storage equipment is represented by automatic storages for tools, pallet-tables and technological equipment. For flexible manufacturing systems (FMS), the transport-loading equipment, including industrial robots, should be engaged in the DPS. It is necessary to investigate the informative abilities of elements of DPS for metalworking. NC-machines and MC are able to generate urgent information about the stage of NC-program, current machining regimes and quantity of machined parts. So-called «monitors of tool condition», used in production practice, are able to supply information about its condition, based on torque-level on the spindle, vibrationacoustic emission signals and by means of laser control of cutting edges [8, 9]. Modern metalworking machines are equipped by accelerometers which allow to estimate wear of parts of equipment on the information of frequency and amplitude of vibration [10–12]. Thermo-sensors, mounted, for example, on spindle, allow to make compensation of temperature deformations of metalworking machine parts [13–16]. Coordinate-measuring machines are able to supply information about dimensional accuracy of produced parts for correction of MC settings and tool settings. Above that, in some FMS (e.g., “Fastems” (Finland) [17]) coordinate-measuring machine is used for the identification of dimensions of workpiece, mounted on pallet-table, before starting of machining to reduce time for measuring of workpiece in cutting zone of MC. Presetters produce information about the dimensions (diameter, length) of instruments [18]. Devices for control of workpiece material are able to supply information for the correction of starting machining (cutting) regimes. Automated storages of tools, pallet-tables and technological equipment are engaged in exchange of information about the presence of needed workpieces, quantity of machined parts, quantity of consumed tools and others [19]. Pallet-tables and tools are able to identify individually, for example, by using RFID-systems, see Fig. 2.

5 Practical Significance Then, it is necessary to investigate the functional mechanism of DPS for metalworking industry [7].

Characteristics, Composition, Mechanisms of Function …

1171

Fig. 2 RFID-label set on the tool

At the stage when the production subdivision recipes workpieces, the system should function by the next way according to the type of workpiece. In case of supplying by stampings, forged workpieces or castings a member of quality-service division puts RFID-label on the workpiece and records needed data on it (order number of part, actual hardness). It is necessary to point out that it is possible to record on RFID-label another data from passport of workpiece, for example, melt order and thermo-treatment order. Information allows to monitor the parameters of object (part of the engine, vehicle, etc.) on all the stages of its creation and, if needed, during the exploitation, that appears as an element of PLM-systems. For workpieces produced by plazma-cutting machines, laser beam-cutting machines, hydro abrasive-cutting machines and cold-stamping operations, mechanical properties are usually set for the production batch, but not for each workpiece. In this case, it is reasonable to put RFID-label on the transport container for a batch of workpieces. At the next stage workpieces transports to storage of metalworking production division or directly to machine-tool. At the storage, the information about order number of part to be produced is read from RFID-label and the data about quantity of supplied workpieces forms. It is necessary to point out that at this stage, it is possible to form linkage with ERP-system of enterprise to provide uninterruptable flow of workpieces with minimal needed stores. At the next stage, assignment for start of production goes. Firstly, the assignment for the start of production goes to the tool-setting division. Tools for machining operations are prepared according to the data from

1172

A. R. Ingemansson

technological documentation. Technological documentation is stored in PC-base and requested by member of tool-setting division on local PC. Presetter devices produce the measurements of tool’s diameter and length. Obtained data record on RFID-label which is put on tool. Then, the tools are sent to MC for tool magazine. MC is equipped with RFID-scanner. Tool parameters (diameter, length) are read and input in NC-system of MC. It is necessary to stress the attention that the time for traditional MC adjustment is reduced by this way. MC operator scans the information about order number of part from RFID-label, which is set on workpiece, on pallet-table or on transport container and obtains NC-program from central technological database. NC-program corrects starting machining regimes according to actual hardness of workpiece, which is recorded on RFID-lable, to provide desired tool life. Definition of workpiece «zero point» could be done with the use of coordinatemeasuring machines at the stage of setting workpiece on pallet-table or with contact devices for control and measuring of working parts on MC, produced, for example, by «Renishaw» (UK), «Blum» (Germany), «Hexagon» (Switzerland). During the machining operation process, so-called «monitors of tool condition», used in production practice, are able to supply information about its condition based on torque-level on the spindle, vibration-acoustic emission signals and by means of laser control of cutting edges. An estimation of tool ware, based on torque-level of spindle, is acceptable for rough and preliminary machining operations. To realize this function exist known products, for example of «Omative» (Israel) [8] and others. For the control of tool’s cutting edges by mean of laser ray are known devices «Renishaw» (UK), «Blum» (Germany), «Hexagon» (Switzerland) and others, which are set in MC working zone. The benefit of tool ware control based on analysis of vibration-acoustic emission signals consists in presence of opportunity to estimate ware on semi-finish and finish machining operations [10–12]. Modern metalworking machines are equipped by accelerometers which allow to estimate wear of parts of equipment on the information of frequency and amplitude of vibration. Thermo-sensors, mounted, for example, on spindle, allow to make compensation of temperature deformation of metalworking machine parts [13–15]. At the next stage, the machined part could be controlled by coordinatemeasuring machine or directly in MC cutting zone with contact measuring probes, which are serially produced, for example, «Hexagon» (Switzerland) [20]. Measuring data are sent to the PC-base for further analysis of enterprise production quality. So, the implementation of DPS provides the efficiency increase of machining production by the following directions: • Decrease of time, spend on preparing operations on MC, due to information from RFID-lables about the tools; • Increase of cutting process effectiveness and reliability due to information from RFID-lables about the actual hardness of workpieces; • Increase of machining operations quality due to the machining process corrections, based on urgent information from measuring systems;

Characteristics, Composition, Mechanisms of Function …

1173

• Increase of cutting tools usage effectiveness due to the cutting process active monitoring and adaptive maintenance; • Increase of machine tools usage effectiveness due to the active monitoring of machine tools parts condition. Above that, implementation of DPS provides the opportunity to increase metalworking enterprise operational effectiveness by means of linkage of production departments with ERP-systems, MES- and APS-systems and PLM-systems.

6 Conclusions The implementation of DPS is the urgent way of increase of efficiency of the mechanical engineering industry and it is the topical subject of the road-map «Technet» of Russian National Technological Initiative up to 2035. The increase of efficiency of mechanical engineering by means of implementation of DPS firstly must be realized in the metalworking industry. DPS for metalworking industry consist of technological, measuring, transport and storage equipment. The aspects of composition and mechanisms of function of DPS and the ways of technological processes efficiency increase in metalworking are analyzed.

References 1. Wahlster W (2013) Industry 4.0: the semantic product memory as a basis for cyber-physical production systems. https://www.ida.lui.se/ida30/program/Wolfgang_Wahlster-IDA3020130924-Industrie_4_0_Active_Semantic_Memories_for_Smart_Factories.pdf. Accessed 8 Nov 2018 2. Bartevyan L (2015) Industry 4.0—summary report. DLG-Expert Rep 5:1–8 3. Wang S (2016) implementing smart factory of Industrie 4.0: An outlook. Int J Distrib Sensor Netw 23–33 4. Ingemansson AR (2016) The relevance of implementation of concept of “industry 4.0” in the modern machine-building industry. Naukoemkie tekhnologii v mashi-nostroenii (Sci Intensive Technol Mech Eng) 7:45–48 5. Ingemansson AR (2016) Modern scientific problem of increasing the efficiency of machining industry by mean of implementation of cyber-physical systems due to the concept of “industry 4.0”. Naukoemkie tekhnologii v mashinostroenii (Sci Intensive Technol Mech Eng) 12:40–44 6. Schwab K, Davis N (2018) Shaping the fourth industrial revolution. Exmo, Moscow 7. Ingemansson AR (2017) The development of composition and mechanisms of function of informational-executive cyber-physical systems in metalworking industry. Naukoemkie tekhnologii v mashinostroenii (Sci Intensive Technol Mech Eng) 11:40–45 8. Systems of adaptive maintenance Omative ACM for CNC metalworking machines: technical specification (2016). 29 p 9. Novak VA (2017) New generation laser inspection system “Blum-Novotest”. RITM of Machine-Building Industry 9:7 10. Zaloga VA (2013) The modern state of cutting instruments diagnostics at milling. Cut Instru Technol Syst 83:118–126

1174

A. R. Ingemansson

11. Kozochkin MP (2012) The tasks of technical diagnostics during the creation and operation with technological equipment. Proceedings of USATU 4:98–104 12. Grigoriev SN et al (2011) The diagnostic of automated industry. In: Grigoriev SN (ed) Mashinostroenie, Moscow 13. Rudzhio H (2016) DMG-Mori condition analyzer—60 sensors control machine-tool condition. DMG-Mori J 2:7 14. Hwacheon (2015) Turning and milling machines: technical catalogue, 293 p 15. Efremova OA (2017) “Industry 4.0” in metalworking. Stankoinstrument 3:78–80 16. Peverill I (2017) A step change in affordable machine health monitoring. Focus 110:15 17. Fastems (2016) Automation of machining industry: technical catalogue, vol 3, 32 p 18. Muller D (2016) Efficiency booster. Z. Mag 2:60–67 19. Zoller (2015) Solutions for tool management: technical catalogue, 40 p 20. Hexagon (2016) Measurement on metallworking machines: technical catalogue, 36 p

Abrasive Machining of Low-Carbon Steels: Ways to Improve the Surface Quality O. A. Kursin, S. B. Fam and N. A. Fedotov

Abstract The chapter discusses issues of improving the efficiency of honing in machining low-carbon complex alloyed steels. The disadvantages of honing “soft” stainless steel 12H18N10T (AISI 321) and frost-resistant steel 09G2S are given. The known methods of increasing the surface hardness such as nitriding, carburization, ni-carbing, hydrogen absorption, and laser alloying of the surface layer have been considered. However, the known chemical methods of increasing the surface hardness have common disadvantages, i.e., reduced corrosion resistance of products that is especially important for the chemical and food industry, as well as thermal deformations of large-sized products arising from the distribution of internal stresses due to heating to temperatures of 700–950 °C. Therefore, these methods are not applicable in some cases. The well-known method of increasing the surface hardness and strength of products made of hardened steels is the plastic deformation, carried out, for example, by rolling or burnishing. For this reason, we propose a new processing method. This is honing with prior plastic deformation.




Keywords Honing Prior plastic deformation steel Abrasive machining





Surface quality
Low-carbon

1 Introduction Low-carbon steels have been more widely applied in various industries. Their weldability is considerably higher in comparison with other steels and makes it possible to create complex engineering structures being used in shipbuilding, mechanical engineering, aircraft manufacturing, automotive industry, and railway transportation. High mechanical resistance to low temperatures allows using pipes made of these steels for hydrocarbons transportation in the north of Russia. High corrosion resistance determines their use in the chemical, food, paper, O. A. Kursin
S. B. Fam
N. A. Fedotov (&) Volgograd State Technical University, 28, Lenin Ave., Volgograd 400005, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_125

1175

1176

O. A. Kursin et al.

pharmaceutical, and textile industries. Compressors and pumps operating in dilute solutions of nitric, acetic, and phosphoric acids and alkalis and salts solutions are made from these steels. Honing is the finishing stage of processing such products. This is the most efficient way to machine the cylinders of compressors and pumps. The main advantages of the honing process are high performance, high dimensional accuracy, and low roughness of the surface machined, practically independent on the accuracy of the machine. The temperature in the treatment area does not exceed 100 °C, and the depth of the defective layer does not exceed 0.006 mm. At present, the manufacture of hones for machining holes with a diameter of from 2.5 to 1000 mm and a length of up to 20,000 mm has been mastered. In comparison with internal grinding, honing has the following advantages: machining holes in complex workpieces (e.g., engine cylinder blocks) is possible; correct geometry of a hole is easier to achieve, since there is no spindle deflection, nor vibration. The shape accuracy achieved in honing is 0.01–0.02 mm for holes with a diameter of 100–120 mm [1, 2]. Characteristic features of honing are a large area of contact between the cutting bars and surface to be machined; minor pressure of 0.1–1.0 MPa; and low cutting speeds of 15–60 m/min that is 40–60 times lower than in grinding. So in the surface layer, there are no physical reasons for microcracks, burn places, or tensile residual stress. Due to the swiveling joint between the honing head and machine spindle, the tool and part are self-aligning, which ensures high precision of the holes to be machined. The swiveling joint between the tool and spindle allows for some discrepancy between the axis of the hole being machined and axis of the tool. This feature of the honing process enables reducing the stock left for machining and idle time for installation, alignment, and fastening of a part.

2 Relevance However, honing “soft” stainless steel 12H18N10T (AISI 321) and frost-resistant steel 09G2S gives a poor surface quality; i.e., lapping along the machining mark (Fig. 1) and scouring are observed.

Fig. 1 Lapping along the machining mark: 1 is the abrasive grain, and 2 is lapping along the machining mark

Abrasive Machining of Low-Carbon Steels …

1177

Various studies [3–10] proved that the characteristics of honed surface microrelief depend on the hardness of the workpiece material (Fig. 2), which indicates the need to increase the surface hardness. We are to consider these known methods for increasing the surface hardness such as nitriding [11], carburization [12], ni-carbing [13], hydrogen absorption [14], and laser alloying of the surface layer [15]. Nitriding [11] (Fig. 3) is a process of saturation of the product’s surface with nitrogen. Such a processing method proposed by Academician N. P. Chizhevsky can improve many characteristics of products made from steel alloys.

Fig. 2 Dependence of arithmetic average surface roughness Ra on surface hardness in steels honing

Fig. 3 Scheme of the experimental installation for nitriding in aqueous ammonia solution during induction heating: 1 is the ammonia solution; 2 is the part; 3 is the air-ejector fan; 4 is the screen; 5 is the cooling water; 6 is the water cooling coil; 7 is the glass vessel; 8 is the support plate; and 9 is the inductor

1178

O. A. Kursin et al.

Fig. 4 Carburization scheme in a solid carburizing compound: 1 is the iron box; 2 is the solid carburizing compound; 3 are the parts; 4 are the control rods; 5 is the cement of the cover; and 6 is the cover

The process occurs at a temperature of 700–1200 °C. Nitriding reduces the toughness of steel, increases its strength, weakens the effect of stress concentrators on fatigue strength of steel, and considerably increases the fatigue strength, especially of thin parts and components operating in certain corrosion media. Nitriding increases the resistance to scouring and metal sticking in processing, especially at elevated temperatures. Nitrided layer has high hardness and wear resistance. The wear resistance of nitrated steel is 1.5–4 times higher than that of hardened high-carbon, cemented, cyanated, and nitro-cemented steels. The disadvantage of the process is the toxicity and high cost of cyanide salts, changes in the size, and buckling of large parts when heated. This process is widely used to process car parts (crankshafts, gears, etc.), dies, molds, and so on. Carburization [12] (Fig. 4) is a chemical saturation process of the product’s surface with carbon. The enrichment of the surface layer of low-carbon and unalloyed steel with carbon to a concentration of eutectoid or proeutectoid structure with some residual amount of austenite and carbides is used to obtain high surface hardness, wear resistance, and fatigue strength. The carburization is carried out in various types of medium at temperatures ranging from 850 to 950 °C. By the medium type, the following methods are distinguished: (1) carburization in a solid carburizing compound, when inorganic (coke) and organic substances (wood, animal bones, etc.) are used with activators being added. Carbon saturation results from a chemical reaction of carbon oxidation. Activators contribute to this reaction. (2) gas carburizing, when specially enriched gases (natural gas, mains gas, etc.) or inert gas (nitrogen) are used. At the same time, small amounts of aliphatic saturated hydrocarbons, alkanes, most often propane, are added to the gas used for cementation. A similar method is increasingly being introduced in thermal processes. In this case, mixtures of high-molecular-weight organic compounds (turpentine, ethyl alcohol, etc.) are introduced into hot retort furnaces that decompose under the action of nickel catalysts.

Abrasive Machining of Low-Carbon Steels …

1179

(3) bath carburizing of two types: in cyanide and cyanide-free baths. However, since the cyanide bath, among other things, contains nitrogen that can be absorbed by steel and may adversely affect the subsequent cutting of the workpiece. In addition, cyanide baths are dangerous to the environment and humans. Therefore, when using liquid carburizing in cyanide baths, it is necessary to observe the safety measures prescribed. In this regard, cyanide-free baths are more appropriate to use. Cyanation and ni-carbing [13]. Simultaneous saturation of the steel surface with carbon and nitrogen is called cyanidation or ni-carbing. The process is carried out either in molten salts containing cyanide group CN or in a gaseous medium. In the first case, the process is called cyanidation; in the second case, it is ni-carbing. Most often, cyanidation is used in the automotive and tractor industry for small parts of medium-carbon steels operating at low specific loading, as well as for high-speed steel cutting tools. The strengthened carbonitride layer has increased hardness, wear resistance, and fatigue strength, which considerably increases the reliability and durability of parts in various operating conditions. Ni-carbing is a more advanced process of chemical heat treatment. Compared with carburization, ni-carbing has a number of essential advantages. When austenite alloying with nitrogen, the a $ c-transformation temperature decreases, which allows saturation at lower temperatures. Hydrogen absorption [14] (Fig. 5) is a process of saturating the surface of a product with hydrogen at room temperature in an electrolytic bath in a low-concentrated solution of sulfuric acid with a low current density.

Fig. 5 Installation scheme for hydrogen absorption: 1 is the workpiece; 2 is the titanium anode; and 3 is the electrolyte

1180

O. A. Kursin et al.

It was experimentally established that under these conditions, the saturation of the surface with hydrogen using electrolysis increases the hardness of low-carbon frost-resistant steels of the 09G2S type and stainless steel 12X18T10 on average by 50–60%. The disadvantage of the hydrogen absorption is the presence of harmful acid waste that must be disposed of. The process also slightly increases the metal shortness after processing. Laser alloying of the surface layer [15] is a process of preliminary processing of the surface treated with coatings that contain alloying elements and further melting them with a laser beam. The method has substantial potential for increasing the hardness. Laser alloying of the surface layer enables a more considerable increase in the surface layer hardness and, consequently, higher surface quality obtained in final abrasive machining. But the currently existing methods of laser alloying of the surface layer have several disadvantages. Alloying the surface in some cases reduces the corrosion resistance of stainless steel products due to the elimination of a part of chromium in the surface solution. In addition, the depth of the layer alloyed this way is not thick and will not always be able to provide the required stock left for subsequent machining, since the melting process of the workpiece material causes large energy consumption and has limitations. The disadvantages of the known methods of increasing the surface hardness are shown in Table 1.

Table 1 List of functions associated with the model nodes of the PPD device for the supply of lubricant-cooling technological means in milling Ways to increase the surface hardness

Disadvantages

Nitriding

– Long process time (up to two days) – For nitriding, expensive alloy steels are used; nitrided parts are 2–3 times more expensive than regular ones – In a solid carburizing compound, the labor intensity is high, and variability of carburizing conditions in carburization is poor – Gas carburizing has high cost – In bath carburizing, longer immersion leads to corrosion of the workpiece – High toxicity of production components – Slightly increased shortness of metal after processing – Toxicity of cyanic salts – Presence of harmful acid wastes that must be disposed of – The process increases the metal shortness after processing – Laser alloying of the surface in some cases reduces the corrosion resistance of stainless steel products due to the elimination of part of chromium in the surface solution – The depth of the layer alloyed in this way is thin and will not always be able to provide the stock left for subsequent machining required

Carburization

Ni-carbing Cyanation Hydrogen absorption Laser alloying of the surface layer

Abrasive Machining of Low-Carbon Steels …

1181

Thus, the known chemical methods of increasing the surface hardness have common disadvantages, i.e., reduced corrosion resistance of products that is especially important for the chemical and food industry, as well as thermal deformations of large-sized products arising from the distribution of internal stresses due to heating to temperatures of 700–950 °C. Therefore, these methods are not applicable in some cases.

3 Purpose and Objective The foregoing makes it possible to formulate the purpose of the work. That is to improve the efficiency of honing in machining low-carbon complex steels. To reach this purpose, the following tasks were set: • Consider the existing methods of increasing the hardness of the surface to be machined before finishing. • Develop a method of honing with prior plastic deformation. • Design and create equipment for the implementation of the honing method with prior plastic deformation. • And develop recommendations for the rational modes of honing steel of austenitic and pearlitic structures with prior plastic deformation.

4 Theoretical Relevance The well-known method [16–24] of increasing the surface hardness and strength of products made of hardened steels is the plastic deformation, carried out, for example, by rolling or burnishing. For this reason, we propose a new processing method. This is honing with prior plastic deformation. The honing method with prior plastic deformation combines two processes, i.e., prior plastic deformation (PPD) and honing itself. Whereby by the time the abrasive tool starts to machine the material of the layer being cut, some of the work intended for plastic deformation in chip formation with conventional honing has been preliminarily performed in the PPD process. The VSTU research team (Kraynev D. V., Ingemansson A. R. and Norchenko P. A.) [17–21] carried out a wide range of cutting studies with prior plastic deformation on steels of pearlitic, austenitic, martensitic–ferritic, and martensitic classes. According to their research studies, the roughness after cutting with prior plastic deformation is less than after conventional cutting. Figure 6 shows a decrease in the surface roughness average Ra (lm) for the surfaces machined with prior plastic deformation, compared with conventional cutting. As Fig. 6 shows, the method under study enables the absolute values of the roughness parameter Ra to decrease.

1182

O. A. Kursin et al.

Fig. 6 Dependence of the roughness of the surface machined on the PPD stress (12H18N10T (AISI 321) steel; T5K10 cutter; t = 1 mm; and So = 0.128 mm/rev). Vc cutting speed is 1 for 90 m/min; 2 for 120 m/min; 3 for 150 m/min; and 4 for 180 m/min

However, they investigated only blade processing with prior plastic deformation. In our work, abrasive processing with prior plastic deformation of low-carbon hard-to-treat steels has been investigated.

5 Practical Relevance In pilot experiments, there were obtained results presented in Figs. 7 and 8. These are profilograms of surfaces processed by honing at different pressure values of prior plastic deformation. Honing was performed under the following modes: T = 1 min was the honing time; Px = 0.4 MPa was the bars unclamping pressure in

Fig. 7 Microprofile of the surface obtained after honing treatment without prior plastic deformation of austenitic stainless steel 12H18N10T (AISI 321) (obtained roughness parameter Ra = 1.3 lm)

Abrasive Machining of Low-Carbon Steels …

1183

Fig. 8 Microprofile of the 12H18N10T (AISI 321) austenitic stainless steel surface obtained by honing after prior plastic deformation. The pressure of the deforming element PPPD = 36 MPa (obtained roughness parameter Ra = 0.65 lm)

honing; Vb–f = 8 m/min was the speed of the back-and-forth motion of the honing head; and Vrot = 50 m/min was the speed of rotational movement of the honing head. The length of the integration path was 4 mm. Increasing the hardness of the surface being processed by prior plastic deformation before honing makes it possible to reduce the roughness parameter Ra up to 50%, increase the surface hardness up to 100%, and enhance the saturation of the metal in the rough layer and the physical and mechanical characteristics of the surface.

6 Conclusions The PPD honing method considered can essentially improve the quality of the surface machined and increase the service life of the product and its competitiveness. Thus, one of the ways to increase the honing efficiency in processing of low-carbon complex alloyed steels is the prior plastic deformation. This method allows avoiding thermal deformation of products and reducing energy costs.

References 1. Kulikov SI (1973) Honing: reference book. Mashinostroenie, Moscow, p 168 2. Babichev AP (2013) Honing: monograph. VGASU, Volgograd, p 245 3. Kao SC (2018) Improving the process of honing the holes of large-sized hydro-pneumatic cylinders by preliminary low-temperature hydrogen treatment of the surfaces to be machined: thesis for a Candidate of Degree in Technical Sciences. Volgograd, 22 p 4. Kremen ZI, Dugin VN, Medvedev VV (1983) The quality of the surface layer in machining abrasive bars. Vestnik Mashinostroeniya 6:73–75

1184

O. A. Kursin et al.

5. Kudoyarov RG (2006) Improving the quality of parts in diamond honing. Mach Tools 5: 35–37 6. Matalin AA (1949) Roughness of the surface of parts in instrument manufacture. Mashgiz, Moscow, p 197 7. Melnikova EP (2003) The influence of technological factors on the surface quality in finishing abrasive machining. Tekhnologiya Mashinostroeniya 3:13–16 8. In: Novikova MP, Orlov PN. Mashinostroenie (1977) In: Handbook of metalist, vol 2 (5 vols). Moscow, 720 p 9. Kaamaikeru GE, Bii J (1994) Patent of Japan JP 6155282. Method of honing. Published in Mar 6, Bulletin 22, Japan 10. Alan Richter (2006) Honing in on perfection. Cutt Tool Eng 8:204–205 11. Lakhtin YuM, Kogan YD (1976) Steel nitrogenization. Mashinostroenie, Moscow, pp 81–88 12. Gulyaev AP (1986) Metallography: textbook for universities, 6th edn. Metallurgy, Moscow, p 544 13. Przhenosil B (1969) Ni-carbing. Mashinostroenie, Moscow, p 212 14. Kursin OA, Plotnikov AL, Kao SCh, Fam SB, Egorov NI, Polyanchikova MYu (2017) Investigation of ways to improve the quality of surfaces of products from low-carbon steels at finishing abrasive machining: monograph. VSTU, Volgograd, p 104 15. Grigoryants AG, Shiganov IN, Misyurov AI (2006) Technological processes of laser processing. Bauman MSTU, Moscow, p 664 16. Poduraev VN, Yaroslavtsev VM, Yaroslavtseva NA (1971) Method of machining with prior plastic deformation. Vestnik Mashinostroeniya 4:64–65 17. Polyanchikov YN, Kraynev DV, Norchenko PA, Ingemansson AR, Amelchenko VV, Razdrogin AV (2010) Positive effect of prior plastic deformation on the formation of surface roughness by cutting. Izvestiya VSTU: inter-university collection of scientific articles. Seri Adv Technol Mech Eng 12(6):41–43 18. Polyanchikov YN, Kraynev DV, Norchenko PA, Gerontidi GV, Ingemansson AR (2009) Obtaining optimal characteristics of the surface layer of parts in cutting by the prior plastic deformation method. Izvestia VSTU: inter-university collection of scientific articles. Ser Adv Technol Mechan Eng 8(5):31–33 19. Polyanchikov YuN, Kraynev DV, Norchenko PA, Ingemansson AR (2010) Improvement of roughness parameters in machining with prior plastic deformation. Vestnik SSTU 1:67–71 20. Yaroslavtsev VM (2010) Cutting with prior plastic deformation: study guide for the course “Perspective technologies of renovation”. Publishing House of Bauman Moscow State Technical University, Moscow, p 46 21. Cheremushnikov NP (1980) Investigation of the plastic deformation process and its instability in metal cutting: abstract of the thesis for a Candidate of Degree in Technical Sciences. Saratov, 22 p 22. El-Hossainy TM (2010) El-Hossainy’s new technique for enhancing the surface of the material. Mater Manuf Process 25:1505–1512 23. Krainev DV, Norchenko PA, Ingemansson AR (2008) Progressive method of cutting and heat-proof steels and alloys. Eur J Nat Hist 4:94 24. Polyanchikov YuN et al (2011) The improvement of stainless steels machining parameters. Int J Appl Fundam Res 2:42–43

Features of Contact Interaction in Cutting High-Alloyed Steels with Carbide Tool A. A. Lipatov, J. L. Tchigirinsky and Hoang Trung Pham

Abstract The paper describes the features of the development of contact interaction on the front and rear (worn area) faces of carbide tools when turning chromium high-alloy steels and analyzes similarities and differences in regard to the regularities of the contact processes during the cutting of austenitic steel, as well as carbon and low-alloy steels of the pearlitic class. The authors revealed differences in the patterns of the contact interaction for high-alloy steels in comparison with the pearlitic steels. It is possible that the identified differences are associated with a different character of the dependence of thermal conductivity on temperature for these groups of machining steels. The studies were carried out during longitudinal turning of steels X10CrNiTi18-10, X14CrNi17-2, and X13CrNiVMoW11-2-2 by using a tool from WC6-Co94 hard alloy. It is shown that at lower cutting speeds, plastic contact is established in the worn area. For steel X10CrNiTi18-10, plastic contact is established in the zone when there is still a pronounced build-up on the front face. The presence of plastic contact in the worn area is experimentally confirmed.













Keywords Turning High-alloy steels Carbide tool Front surface Wear platform Contact interaction Temperature dependence Thermal conductivity










1 Preface The paper describes the features of the development of contact interaction on the front and rear (worn area) faces of carbide tools when turning chromium high-alloy steels and analyzes similarities and differences in regard to the regularities of the contact processes during the cutting of austenitic steel [1–4], as well as carbon and low-alloy steels of pearlitic class [5–7].

A. A. Lipatov
J. L. Tchigirinsky (&)
H. T. Pham Volgograd State Technical University, 28, Lenin Av., Volgograd 400005, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_126

1185

1186

A. A. Lipatov et al.

2 Techniques The studies were carried out during longitudinal turning of steels X10CrNiTi18-10, X14CrNi17-2, and X13CrNiVMoW11-2-2 by using a tool from WC6-Co94 hard alloy [8–11]. Steels X10CrNiTi18-10 and X14CrNi17-2 were machined in the as-delivered state, and blanks from X13CrNiVMoW11-2-2 were previously underwent heat treatment to a hardness of 415 HB. The cutting speed was varied in the range of cutting speeds v = 10–150 m/min with a feed of 0.3 mm/rev and a cutting depth of 1.5 mm. The tool featured the following main geometrical characteristics including rake angle c = 0°, back relief angle a = 10°, main approach angle u = 45º. In the course of the experiments, the chip shrinkage coefficient n was measured. The roots of the chips were obtained using a “falling” cutter. At the same time, on the rear surface, a specific area with width hз  0.2 mm was grinded and worn in during the cutting process.

3 Experiments Figure 1 shows the experimentally obtained dependences of chip shrinkage n on cutting speed v. The appearance of the curves indicates a smooth (to various extents) change in shrinkage n as the velocity v increases. Such a character of n (v) dependences sharply differs from the one peculiar to the treatment of pearlite steels; i.e., when the cutting speed increases, an abrupt (at almost one the same speed) increase in n is observed in a transfer from a build-up to plastic contact [12]. Thus, with an increase in cutting speeds, in regard to steels X14CrNi17-2 and X13CrNiVMoW11-2-2, the contact interaction on the front face should develop similarly to cutting austenitic (and not pearlite) steels.

Fig. 1 Dependencies of chip shrinkage n on cutting speed v when turning steel X10CrNiTi18-10 (1), X14CrNi17-2 (2), and X13CrNiVMoW11-2-2 (3) by using a tool from hard alloy WC6-Co94

Features of Contact Interaction in Cutting High-…

1187

This was evidenced from studying the microsections of the chip roots. When cutting high-alloy steels X14CrNi17-2 and X13CrNiVMoW11-2-2, a transfer from the build-up to the plastic flow of contact layers did not occur directly, but through an intermediate type of interaction, similar to that typical for cutting steel X10CrNiTi18-10—a stagnant zone with a out-of-position build-up [1–3, 12–14]. This is due to the same nature of the dependence of thermal conductivity of these groups of steels on temperature. The thermal conductivity of high-alloy steels is increases with an increase the temperature—unlike carbon, low—and medium-alloy steels of pearlite class, which experience a sharp decrease in thermal conductivity as the temperature grows (in the temperature range 200–800 °C, by 1.5–1.8 times [13, 15]). Therefore, when cutting pearlite steels in the contact zone, there is a positive feedback on temperature—with an increase in speed and temperature, the thermal conductivity of the material being machined drops sharply. Therefore, the heat flow to the chips is also reduced, which in turn leads to an increase in temperature. Therefore, a transfer from a build-up to plastic contact occurs abruptly. The increase in thermal conductivity of high-alloy steels, with speed and temperature growing higher, causes the presence of negative feedback on temperature, intensifying the heat flow from the contact zone to the chips and prolonging the transfer from the build-up to plastic contact in a fairly wide range of speeds [1, 2, 14, 15]. An analysis of the relevant photographs (Fig. 2) gives the following evidence: For X14CrNi17-2 (Fig. 2a), the stagnant zone with an out-of-position build-up is almost indistinguishable from the corresponding type of interaction [1, 2] when cutting X10CrNiTi18-10—the build-up displaced from the cutting edge retards the flow of the contact layers at a distance from the front surface not less than the height of the out-of-position build-up. On the contrary, when cutting steel X13CrNiVMoW11-2-2 (Fig. 2b), the thickness of the stagnant zone located between the build-up remainder on the cutting edge and the out-of-position build-up is much lower than the height of both build-ups. As the rate increases and the build-up on the cutting edge disappears, a sufficiently large out-of-position build-up remains, but the thickness of the stagnant zone becomes even smaller—so that the interaction in this area can be considered as plastic contact (Fig. 2c). This is due to increased heat generation in the chip formation zone due to the high strength of steel X13CrNiVMoW11-2–2 [16–19]. A second difference from machining pearlitic steels (and similarity with austenitic steel) is as follows: The decrease in the thermal conductivity of pearlite steels, as the temperature increases, causes the weakening of the heat flow to the heated chips as compared with a cold workpiece [20, 21]. Therefore, the development of contact interaction in the worn area, with v growing higher, lags behind the corresponding processes on the front surface [12]. On the contrary, the increase in thermal conductivity of high-alloyed chromium steels X14CrNi17-2 and X13CrNiVMoW11-2-2 (as well as X10CrNiTi18-10), with an increase in temperature, leads to an increase in the heat flow to the heated chips (as compared with the colder workpiece).

1188

A. A. Lipatov et al.

Features of Contact Interaction in Cutting High-…

1189

JFig. 2 Microsections of the chip roots (100) formed in turning the following steels:

a X14CrNi17-2, at cutting rate v = 25 m/min; b X13CrNiVMoW11-2-2, at the cutting rate v = 20 m/min; c X13CrNiVMoW11-2-2, at the cutting rate v = 30 m/min

Fig. 3 Microsection of the chip root formed in turning the steel of grade X10CrNiTi18-10, hard alloy WC6-Co94, at the cutting rate v = 20 m/min: a general view (100); b beginning of the contact along the worn area (1000)

1190

A. A. Lipatov et al.

Therefore, at lower cutting speeds, plastic contact is established in the worn area, rather than on the front face (Fig. 2c). At the same time, high-chromium steels are characterized by a less intense increase in thermal conductivity, with an increase in temperature (in the range 200–800 °C—by less than 1.1 times as compared with almost 1.5 times for austenitic steels [13]). Therefore, the advanced development of contact processes in the worn zone is not as clearly expressed as for austenitic steel. For X10CrNiTi18-10, plastic contact is established in the zone when there is still a pronounced build-up on the front face (Fig. 3a). A photograph with a higher zoom (Fig. 3b) confirms the presence of plastic contact (rather than external friction) in the worn area. The high-alloyed chrome steels still experience congestions in the worn zone at the cutting rates, at which a congestion zone with an out-of-position build-up occurs (Fig. 2a, b).

4 Resume The aforementioned results prove that the dependence of the thermal conductivity of the material being machined on temperatures affects the patterns of the development of contact processes.

References 1. Lipatov AA (2013) Features of transferring from build-ups to interactions with plastic contact in austenitic steels machining. Izvestiya VolgGTU 7/110(9):31–34 2. Lipatov AA (2017) Nature of contact interaction in the worn zone of a carbide tool in austenitic steels turning. Izvestiya VolgGTU 5(200):28–30 3. Zeldovich VI (2008) Three mechanisms of austenite formation and structural heredity in iron alloys. In: Works collection of ideas development of academician V.D. Sadovsky. Yekaterinburg, pp 84–98 4. Norchenko PA (2010) Increasing the efficiency of the cutting process of stainless steels of austenitic class with advanced plastic deformation. Dissertation, Volgograd state Technical University 5. Smith L, Celant M (2002) Martensitic stainless steels in context. SMSS 2002, paper No. 017 6. Bhadeshia HKDH (2001) Bainite in steels, 2nd edn. The University Press, Cambridge 7. Brjozovsky BM et al (2012) Improving the efficiency of machining parts on the basis of application tools with the modified work surfaces and optimization of cutting process In: High technologies in mechanical engineering. Science intensive technologies in mechanical engineering, vol 1, pp 3–10 8. Gerasimova NC (2016) Tool materials: tutorial. Moscow State Technical University named N. E. Bauman, Moscow 9. Zubkov NN (2013) Tool materials for the manufacture of cutting tools. Scientific Edition of the Bauman MSTU Named N. E. Bauman, Science and Education, Moscow 10. Sharyn JuS, Tishenina TI (1985) Handbook turner. Environments Ural Book Publishing House, Sverdlovsk

Features of Contact Interaction in Cutting High-…

1191

11. Kozhevnikov DV, Kirsanov SV (2012) Metal cutting. Mashinostroyeniye, Moskow 12. Talantov NV (1992) Physical basics of cutting, tool wear and tool collapsing. Mashinostroyeniye, Moscow 13. Lipatov AA, Tchigirinsky JuL, Kluikov DS (2015) Formation of transitional layer at the processing of high-alloyed steels. World Res Pap 4:7–10 14. Lipatov AA (2008) Instability of chip formation and the wear of a hard-alloy tool in cutting austenitic steel. Russ Eng Res 9:904–905 15. Reznikov AN (1981) Thermophysics of materials machining processes. Mashinostroyeniye, Moscow 16. Schastlivtsev VM (2006) Features of the structure and crystallography of lath martensite in steels. In: Merson DL (ed) Collection of perspective materials: structure and research methods. Tomsk State University, National University of Science and Technology MISIS, Moscow 17. Talantov NV (1992) Physical fundamentals of the cutting process of wear and destruction of the tool. Mechanical Engineering, Moscow 18. Talantov NV (1984) Physical fundamentals of cutting process. In: Physical processes in the cutting metal: collection of articles. Volgograd Politechnical Institute, Volgograd 19. Kimura M et al (2000) Effect of retained austenite on corrosion performance for modified 13% Cr steel pipe. In: Corrosion 2000, Paper No. 00137 20. Pham HT, Tchigirinsky JL (2017) Analysis of mechanism of carbide tool wear and control by wear process In: MATEC web of conferences. International conference on modern trends in manufacturing technologies and equipment (ICMTMTE 2017), vol 129. https://doi.org/10. 1051/matecconf/201712901058 21. Tchigirinsky JL, Trung PH, Lipatov AA (2016) Features of processing of corrosion resistant steels. Mater Sci Forum 870. https://doi.org/10.4028/www.scientific.net/MSF.870.598

Analysis of Influence of Strains of Technological System Elements on Machining Accuracy Under Turning of Non-rigid Shafts Based Between Centers P. S. Nesterenko, J. L. Tchigirinsky and E. N. Nesterenko

Abstract The mechanism of the formation of the machining error caused by the strains of the technological system elements under the turning of non-rigid shafts-based between the centers is described. The basing scheme and distribution scheme of forces acting in the process of turning are presented. The influence of each component of the cutting force on the form and sizing error of the machined surface is considered. A refined mathematical model for calculating this type of error is proposed, the use of which will improve the reliability of technology at the design stage of the technological process. Experimental data and the results of the calculation of the machining error according to the proposed refined theoretical dependencies are compared. The analysis of the data showed that the proposed refined mathematical model has a higher convergence of results with experimental data in comparison with traditional mathematical models adopted in mechanical manufacturing engineering. The resulting refined model can be used in control algorithms for adaptive control systems, which will improve the accuracy of their operation.










Keywords Turning machining Non-rigid shaft Machining accuracy Strains of the technological system elements

1 Introduction The performance characteristics of the product as a whole and its durability depend on the accuracy and quality of manufacture of individual parts and assemblies. However, in the course of processing, under the action of the components of the P. S. Nesterenko (&)
J. L. Tchigirinsky Volgograd State Technical University, 28, Lenin Av, Volgograd 400005, Russia e-mail: [email protected] E. N. Nesterenko Volzhsky Polytechnic Institute, 42a, Engels Str, Volzhsky 404121, Russia © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_127

1193

1194

P. S. Nesterenko et al.

cutting force, elements of the technological system are displaced from the initial (unloaded) state, thereby causing the appearance of machining error [1]. This is especially evident in the turning of non-rigid shafts. Due to the low rigidity of the blanks, the magnitude of this type of error in the turning of them reaches 80–90% of the total machining error [2]. These operations are preliminary, and their result, due to the effect of technological heredity, substantially predetermines the quality of the subsequent finishing treatment. The formulas adopted in manufacturing engineering for calculating the errors of this type have relatively low convergence with experimental data [3], because they are simplified and do not take into account the peculiarities of the course of the treatment process. This circumstance causes the appearance of an error in the technological process at the design stage of the technological process, in particular in assigning allowance for machining for subsequent finishing operations.

2 Elastic Strains of the Technological System Elements Under the Turning of Non-rigid Shafts Based Between the Centers Consider the mechanism of the formation this type of machining error under the turning a smooth non-rigid shaft. This type of mounting is widespread, because provides the highest accuracy and for its implementation requires minimal time investment compared with other bonding technique. Figure 1 shows basing scheme and distribution scheme of forces acting in the process. In the above scheme, the centrifugal inertia force is not taken into account, because the effect of which on the amount of elastic strains under the turning a smooth shaft is insignificant [4]. The shaft-based between centers is represented a beam and freely supported by its ends. The tangential component of the cutting force PZ causes the appearance of bearing pressures RZH and RZR, acting on the head and rear centers, respectively. In addition, the weight of the shaft Q is perceived by the centers and also affects to the magnitude of reactions RZH and RZR. The radial component of the cutting force PY

Fig. 1 Basing scheme and distribution scheme of forces acting in the process of turning

Analysis of Influence of Strains of Technological System …

1195

causes the appearance of bearing pressures RYH and RYR, acting on the center in the horizontal plane perpendicular to the axis of the lathe. In addition, the moment in the horizontal plane, caused by the action of the axial component of the force PX applied at a distance of half the diameter of the workpiece from the centerline, also affects to the magnitude of reactions RYH and RYR. The very force PX is entirely perceived by the head center, thereby unloading the rear center. Therefore, in order to avoid a weakening of the rear center, he is informed about the preliminary tightening of FT equal to the value of PX. Consider separately the effect of strains of each of the elements of the technological system on the accuracy of machined surface. It should be noted that elastic strains of the system, primarily due to the effect of the radial component of the cutting force PY and partly action of the axial component of the cutting force PX. As for the strains under the action of the tangential component of cutting force PZ, their influence on the accuracy of machining is small, and it can be neglected [2].

2.1

Strains of the System Under the Action of Radial Component of the Cutting Force PY

In considering the bending of the shaft axis under the action of the components of cutting forces in turning, the machine units we will assume absolutely rigid. Then, the magnitude of the bending of the shaft y1i under the action of the radial component of the cutting force can be determined according to the energy method: Zx y1i ¼ 0

MPY I
M1I dx þ E
I1

Zl x

MPY II
M1II dx; E
I2

ð1Þ

where • MPYI, MPYII—bending moment from the radial component of the cutting force PY on the treated and treated area, respectively; • M1I, M1II—bending moment from a unit force on the treated and treated area, respectively; • I1, I2—moment of inertia of cross section of the finished and treated stages, respectively. Based on the design scheme shown in Fig. 1, and after the transformations, Expression (1) takes the form: y1i ¼

PY
x3
ðl  xÞ2 PY þ 2 3
E
I1
l E
I2

 3   x2 l  x3 x2  2  l  x2
þ ð l  xÞ
x2 ;
l 3 l ð2Þ

1196

P. S. Nesterenko et al.

However, in practice, for the calculation of the shaft strains instead of Expression (2) is used a simplified calculation formula which does not take into account the changing shaft stiffness due to the change in its cross section during turning. It is wrong. Calculations showed that the difference in determining the amount of deflection by using the simplified formula adopted in manufacturing engineering, with the calculation by using the refined formula (2) reaches 50% compared. Such a ratio often occurs in practice while turning workpieces with a diameter of less than 40 mm with a cutting depth of 0.25 d1. Consider the displacement of centers. In this case, we while considered the shaft absolutely rigid and the caliper is unyielding. Under the action of the force PY in the process of machining, the centers of the lathe are elastically pressed, and after the removal of the load, they return to their initial position (Fig. 2). So, in the initial position of the cutter, the head center is not under load and stay at rest (yH = 0), while the rear center under the action of the force PY will move to the maximum value (yR = max). As the cutter moves to the headstock, displacement of rear center depresses to zero and displacement the head center increasing and reaches its maximum at the extreme left position of the cutter. The curve A1aB1 in Fig. 2 expresses the distorted shaft profile. The magnitude of the error y2i can be found analytically based on the similarity of triangles Aicb and AidBi:  x x y2i ¼ 1 
yRi þ
yHi ; l l

ð3Þ

Since the strain of stocks is directly proportional to the loads, the equation of shaft cross-sectional line can be written in the form:  x2 x 2 y2i ¼ PY
1 
eR þ
eH ; l l

ð4Þ

Fig. 2 Displacement of the centers of the lathe under the action of the radial component of the cutting force PY

Analysis of Influence of Strains of Technological System …

1197

where • ɛH, ɛR—flexibility of the headstock and tailstock, respectively. As for the stain of the support, the operating experience of lathe showed that the flexibility of the support ɛS remains almost constant over the entire length of the treatment, conditional upon it properly assembled and in good condition. Its release under the action of force PY is constant and equal to: y3i ¼ PY
eS ;

2.2

ð5Þ

Strains of the System Under the Action of Axial Component of the Cutting Force PX

Applied at a distance d/2 from the axis of rotation, the axial component of the cutting force PX creates a torque MX, under the influence of which bearing pressure P′R and P′H occur. It acting, respectively, on the tailstock and headstock of the lathe in opposite directions: d P0R ¼ P0H ¼ PX
; l

ð6Þ

Under the action of a bending moment MX, the shaft axis is bent through an angle h, as a result of which longitudinal deformation of the shaft axis y4i occurs, as shown in Fig. 3a. Equation bending moment at the section x for this case has the form: M ¼ P0R
x  MX ¼ PX


d d
x  PX
; 2
l 2

ð7Þ

Fig. 3 a Bending of the shaft axis under the action of moment MX; b Displacement of the line of centers under the action of moment MX

1198

P. S. Nesterenko et al.

Equation of the elastic line for this case has the form: PX
d
x2 PX
d
x  ; 2
E
I1
l 2
E
I1 PX
d
x3 PX
d
x2  ; y4i ¼ yðxÞ ¼ h0
x þ 6
E
I1
l 4
E
I1

hðxÞ ¼ h0 þ

ð8Þ

where h0—initial angle of rotation of the shaft. To determine the integration constant h0, we use the boundary conditions. At the points corresponding to x = 0 and x = l, the diameters of the shaft are not distorted, then: h0 ¼

PX
d
l ; 12
E
I1

ð9Þ

Thus, the magnitude of the strain of the shaft axis, caused by the action of the moment MX, can be determined by the following expression: y4i ¼

  PX
d
l
x x2 x
2
2  3 þ1 ; 12
E
I1 l l

ð10Þ

However, calculations show that the bending of the shaft axis caused by the action of the moment MX can be significant only when turning very thin shafts with a diameter of up to 20 mm and a ratio of length to diameter over 15. In this case, the bending of the axis from the action of the radial component of the force PY turns out to be so large that the bending of the shaft axis caused by the action of the moment MX can be neglected. Since both stocks have unequal flexibility, they are displaced on different values y′R and y′H in opposite directions under the action of the bearing pressure P′R and P′i. As a result, the shaft rotates in a horizontal plane passing through a line of centers, as shown in Fig. 3b. If we assume that the shaft is absolutely rigid, then the offset position of the axis A1B1 will remain unchanged throughout the length of the turning. In this case, the diametrical dimensions of the shaft will change by the amount of double displacement of the axis y5i, which can be determined based on the similarity of triangles AA2B and B1cd:  x x y5i ¼ 1 
y0R 
y0H ; l l

ð11Þ

Since the strain of stocks is directly proportional to the loads and unchanged during processing, Eq. (11) can be written as: y5i ¼ PX


 x i d h x
1 
eR 
eH ; 2
l l l

ð12Þ

Analysis of Influence of Strains of Technological System …

1199

However, according to the calculations, the magnitude of the displacement of the centerline under the action of the moment MX is insignificant in comparison with the total flexibility of the system, which is why this type of error can be neglected. Also, under the action of the force PX, the cutter is pressed, bent and turned; however, these deformations are so small that they practically do not affect the accuracy of the diametrical dimensions of the machined shaft and can be neglected.

3 Experimental Procedure The final formula for calculating the machining error DR caused by the elastic strains of the technological system elements under the turning of non-rigid shafts-based between the centers will take the form: X DR ¼ 2
yji ¼Py 8  3 9  3  x2 > > x
ðl  xÞ2 1 l  x3 x2  2 > 2 2 > > þ  l  x
þ ð l  xÞ
x >
= < l 3
E
I1
l2 E
I2 3 l ;
 x2 > > x 2 > > > > þ 1 
e þ
e þ e : R H S; l l ð13Þ Table 1 presents the experimental data on the study of machining error DR obtained by turning a shaft with a diameter of 24.8 mm and length of 325 mm from steel 45 on a lathe model 16A20F3 CNC with a T15K6 cutting plate with the following cutting conditions: feed −0.25 mm/rev, depth of cut −0.6 mm, and cutting speed −46.3 m per min [5]. Table 1 also presents the data obtained by calculating the machining error by the refined formula (13) and by the traditional formula given in the source [6]. A significant discrepancy between the calculated data obtained in the calculation using the refined formula (13) and the experimental data at the beginning of the turning process may indicate that the turning was carried out without sufficient preliminary pressing of the rear center (sliding of the center). The discrepancy in the data at the end of turning process is due to significant wear of the metal-cutting tool during treatment.

4 Conclusion The analysis of the data showed that the proposed refined mathematical model for calculating the machining error due to the strains of the technological system elements has a higher convergence of results with experimental data in comparison with traditional mathematical models adopted in mechanical manufacturing

1200

P. S. Nesterenko et al.

Table 1 Experimental data and the results of the calculation of machining error for theoretical dependencies Considered section x (mm)

Experimental data

ΔR (mm)

Calculated data (according to the traditional formula [6]) ΔR (mm) Deviation (%)

Calculated data (according to the revised formula (13)) ΔR (mm) Deviation (%)

D (mm) 325 300 275 250 225 200 175 150 125 100 75 50 25 0

23.71 23.72 23.73 23.75 23.77 23.79 23.80 23.82 23.81 23.79 23.77 23.74 23.70 23.67

0.11 0.12 0.13 0.15 0.17 0.19 0.20 0.22 0.21 0.19 0.17 0.14 0.10 0.07

0.090 0.089 0.107 0.133 0.159 0.179 0.189 0.187 0.172 0.149 0.119 0.089 0.068 0.065

0.090 0.089 0.108 0.136 0.166 0.190 0.204 0.204 0.191 0.166 0.132 0.097 0.071 0.065

18.08 25.58 17.80 11.57 6.71 5.99 10.15 15.11 17.69 21.60 29.90 36.05 31.88 7.03

18.08 25.41 16.74 9.01 2.37 0.01 2.98 7.12 9.02 12.86 22.37 30.64 29.42 7.03

engineering. Application of this mathematical model will improve the reliability of technology at the design stage of the technological process, in particular in assigning allowance for machining for subsequent finishing operations. In addition, the resulting refined model can be used in control algorithms for adaptive control systems, which will improve the accuracy of their operation.

References 1. Perelygina TA, Frolova AV (2015) Vliyanie sil, dejstvuyushchih v tekhnologicheskoj sisteme na tochnost’ obrabotki nezhestkih valov 9Influence of the forces acting in the technological system on the accuracy of processing non-rigid shafts). Molodoj uchenyj, Kazan 2. Podporkin VG (1959) Obrabotka nezhyostkih detalej (Processing of non-rigid parts). Mashgiz, Leningrad 3. Balakshin BS (1973) Adaptivnoe upravlenie stankami (Adaptive machine control). Mashinostroenie, Moscow 4. Tchigirinsky JL, Nesterenko PS (2017) K voprosu upravleniya tochnost’yu prodol’nogo profilya detalej tipa nezhyostkij val (To the issue of controlling the accuracy of the production profile of non-rigid shaft parts). IzvestiyaVolgGTU, Volgograd 5. Nabnpkin AU (2012) Avtomatizirovannaya sistema upravleniya formoj nezhyostkih valov pri tokarnoj obrabotke (Automated shape control system for non-rigid shafts during turning). Vestnik SamGTU, Samara 6. Matalin AA (1985) Tekhnologiya mashinostroeniya (Engineering technology). Mashinostroenie, Leningrad

Analysis of Influence of Strains of Technological System …

1201

7. Venkata Rao R (2011) Modeling and optimization of machining processes. In: Advanced modeling and optimization of manufacturing processes, Springer Series in Advanced Manufacturing. Springer, London, pp 55–175 8. Gavrilov VA (2004) Calculation and testing of machines for accuracy: tutorial book. OmGTU, Omsk 9. Cardi AA, Firpi HA, Bement MT et al (2008) Workpiece dynamic analysis and prediction during chatter of turning process. Mech Syst Sig Process 22:1481–1494 10. Palaniappan D (2011) A general solution of equations of equilibrium in linear elasticity. Appl Math Modell, Elsevier 35:5494–5499 11. Świć et al (2014) Accuracy control in the machining of low rigidity shafts. Appl Mech Mater 613:357–367 12. Azimov BM, Sulyukova LF (1996) Mathematic modeling and the optimal control of the small rigidity details processing technological system. Reference book, Engineering magazine 3:45–51 13. Romano F (1996) Deflections of Timoshenko beam with varying cross-section. Int J Mech Sci 38:1017–1035 14. Kashyzadeh KR, Ghorabi MJ (2012) Study of chatter analysis in turning tool and control methods—a review. Int J Emerg Technol Adv Eng 2:1–5 15. Kosilova AG, Meshcheryakov RK (2003) Spravochnik tekhnologa-mashinostroitelya (Handbook of technology-machine builder). Mashinostroenie, Moscow

Peculiarities of Application of Th20 Hard Alloy for Turning Processing of Various Steels with Advanced Plastic Deformation P. A. Norchenko, V. A. Solodkov and S. I. Kormilitsyn

Abstract The problems of management of the mechanical treatment modes and provision of the set indicators of the processed surface quality are being discussed. The concept of the technological operations of machine treatment in the direction of an operational assessment of the cutting tool characteristics, the processed material and the detailed surface layer for the purpose of the efficiency increase is justified. The paper discusses the physical basis for using the TN20 hard alloy for machining steels of various grade (stainless and carbon steels) from the point of view of prospects for improving the efficiency of the cutting process with advanced plastic deformation, as well as demonstrates ways for optimizing cutting conditions. The changing of the development and management concept of the technological operations of machine treatment in the direction of an operational assessment of the cutting tool characteristics, the processed material and the detailed surface layer for the purpose of the efficiency increase is expedient. Keywords Hard alloy conditions


Advanced deformation
Thermal emf
Cutting

1 Introduction It is known that in a substantial part of sources, recommendations on the selection of a cutting part material of the tool [1] as a priority parameter use the mechanical properties of the tool material, such as hardness, plasticity, and impact strength. As a rule, the selection of tools for various operations (rough, finishing, and fine) is expressly linked to the phase composition, particularly, to the content of the bond (cobalt, nickel, molybdenum, etc.). The selection of the cutting part material of the tool for cutting various types of materials to be processed is often made using recommendations obtained experimentally [2]. P. A. Norchenko (&)
V. A. Solodkov
S. I. Kormilitsyn Volgograd State Technical University, 28 Lenin Ave, Volgograd 400005, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_128

1203

1204

P. A. Norchenko et al.

Given the increasing requirements for predicting the parameters of the cutting process, it should be noted that the said sources do not take into account such an important characteristic of the tool material as thermal conductivity. The thermal conductivity of the tool material obviously has a significant effect on the distribution of heat fluxes in the cutting zone, temperature and deformation regularities of the process, as well as the wear pattern of the cutting part of the tool [2]. Established that a significant part of the energy of plastic deformation does not linger in the crash element, it is released in the form of heat, and only a negligible part lingers therein, increasing its internal energy [3]. It was also noted that the accumulation occurs to a certain limiting value, after which the metal is destroyed. As noted in [3], in accordance with the first law of thermodynamics, the external influence may be mechanical or thermal. A body is considered destroyed if the density of internal energy has increased to a critical value in at least one of its local volumes. The specific potential energy intensity of the processed material in the volume affected is defined as follows: Dw ¼ ½u  u0 ¼ HT þ LT ;

ð1Þ

where [u] is the limiting (critical) level of internal energy density; u0 is the initial level of the internal energy density; HT is the specific thermal capacity of melting of a material; LT is the latent heat of melting of a material. It is obvious that the initial level of the internal energy density, u0, is related to the internal energy of the system, U, which suggests a direct relationship between the energy state of the system and the process of destruction, regardless of the type of influence. Cutting heat-resistant and stainless steels and alloys is hampered due to the following features of these materials: high hardening of material in the process of cutting deformation, the low thermal conductivity of the material being processed, ability to retain the original strength and hardness, high abrasion ability, and reduced vibration resistance of cutting. Taking into account the features above the process of cutting stainless and heat-resistant steels and alloys is carried out as follows: first, the working surfaces of the tool come in contact with a relatively soft, non-hardened metal, and plastic deformation, which is accompanied by significant absorption of energy applied from the outside (by the tool), occurs under influence of such working surfaces. In this case, the layer being cut gains much hardening and acquires the properties of work-hardened metal; i.e., it becomes brittle. Plasticity margin is then largely exhausted, and a shear, i.e., destruction, formation of a chip element, occurs. The low thermal conductivity of these materials leads to a sharp decrease in heat removal to the chips and the workpiece, and, consequently, an increase in temperature in the contact area between the cutting part of the tool and the workpiece with the activation of adhesion and diffusion processes. As a result, tool wear and sticking (setting) phenomena, which cause the destruction of the cutting edges, are significantly increased. The intensification of these processes is

Peculiarities of Application of Th20 Hard Alloy for Turning …

1205

facilitated by the increased mechanical characteristics of the material being processed at high temperatures, such as large abrasion capacity of materials, as well as the by the variable effect of these factors due to vibrations [1]. To solve the problem above, it is proposed to use a cutting method with advanced plastic deformation (APD). Cutting with APD provides the improvement of chip formation conditions by a rational change of the physical and mechanical properties of the material of the layer being cut due to its hardening prior to the cutting process. Hardening is carried out using a rolling device that creates the depth and degree of work hardening necessary to obtain maximum efficiency of the subsequent process. During the machining of plastic materials, the separation of the material of the layer being cut from the workpiece is preceded by its intense plastic deformation; i.e., the main part of the cutting work is expended for the plastic deformation of the metal being removed. The essence of cutting with advanced plastic deformation of the material of the layer being cut consists in combining two processes, plastic pre-deformation, and direct cutting process. Whereupon, by the time the cutting tool starts to process the material of the layer being cut, part of the work is expended for plastic deformations in the process of chip formation during normal cutting is already pre-performed by an axillary device. From the physical point of view, in the process of APD, the crystal lattice of the metal being processed is saturated with deformation energy according to the concepts described above and is removed from the equilibrium state. The density of dislocations increases and the atoms located in the dislocation zone have lower stability due to the higher level of free energy compared to other ones. Therefore, from the side of the cutting tool, a much smaller amount of shear stress is required for their displacement along the shear plane and performance of the direct process of removal of metal in the form of chips compared to processing without APD. Consequently, in the process of cutting with plastic pre-deformation, the cutting tool does not perform the entire work, but only part thereof [4, 5]. The purpose of the scientific inquiry is supposed to be the prospect to identify the possible connection of thermal conductivity of materials of the tool/part pair and the type of process operation, thereby expanding the list of input parameters when solving the task of selecting the material of the cutting part of the tool [6, 7]. When machining steels, there are certain threshold velocities, which are the boundary of the transition of the flow chip formation to the cyclic one. These velocities are individual for each material, and for heat-resistant and corrosion-resistant steels and alloys; this phenomenon is already present during traditionally applied modes. It is expressed in the loss of the chip formation zone of thermodynamic stability and the repetition of cycles in time, consisting of two phases, shear deformation in a thin rotating chip formation zone and compression. This destructive phenomenon is expressed in the fluctuation of cutting forces, temperature, velocities of chip movement, which intensifies the wear of the cutting tool [3]. To reduce the detrimental effect of the described process, particularly, on the durability of the cutting tool, it is proposed to use advanced plastic deformation (APD) during cutting. When processing with the help of APD, by the moment of

1206

P. A. Norchenko et al.

the beginning of the contact by the cutting tool to the material of the layer being removed, a part of the work expended for plastic deformation in the process of chip formation during normal cutting is already pre-performed by an axillary device creating the necessary depth and degree of work hardening. In the process of cutting with the help of APD, the cutting tool does not perform the entire work, but only a part thereof, which ensures a reduction in the force and temperature of cutting [2]. It should be noted that initially, the direction of scientific inquiry concerned the area of cutting difficult-to-machine steels and alloys. Particularly, a method was developed to improve their machinability, associated with advanced plastic deformation, the essence of which is described in [8–10]. Thus, when cutting austenitic stainless steels, the finishing and semi-finishing modes of processing are of particular interest, since they are characterized by considerable tool wear and problems in achieving adequate surface layer properties. The most characteristic representative of the grade is heat-resistant steel 12X18H10T (thermal conductivity of 16 W/m K). Based on the available recommendations and production experience, when processing this material, a tool of the VK group is most commonly used; TC tool is used less commonly (thermal conductivity: VK8—52 W/m K, T15K6— 27 W/m K). However, as a result of the research conducted, the conclusion was made that the effect of the application of advanced plastic deformation for alloys with a lower content of tungsten carbide and, consequently, with lower thermal capacity turned out to be considerably more significant. The use of carbide tools with the lowest possible thermal conductivity seemed promising [11, 12].

2 The Physical Prerequisites for the Improvement of the Cutting Process Among the brands used in production, the TN 20 tungsten-free hard alloy was selected (phase composition: TiC—80%, Co + Mo—20%; thermal conductivity— 11 W/m K). The traditional field of application of this tool is carbon steels and alloys, non-hardened tool steels [2]. Analysis of the mechanical properties of this material clearly indicates its applicability only in the field of impact-free loads, which excludes its use in rough operations. A cutting process, in which a pair of tool/workpiece with very low values of thermal conductivity of materials is involved, was organized. [2] indicate the decisive value of the temperature in the cutting zone for the course of the process and thermal conductivity of materials of the tool and the workpiece as the most significant factor, which forms this temperature. The temperature in the cutting zone determines the process of thermal softening and the length of the contact areas, which, in turn, is associated with the formation of stresses and, consequently, cutting forces. It is also indicated that it often forms the distribution of temperature fields in the tool, which is associated with its wear and tear [13].

Peculiarities of Application of Th20 Hard Alloy for Turning …

1207

The most important process affecting the formation of temperature is the distribution and removal of heat from the cutting zone. Due to the very low thermal conductivity of the material being processed and tool material, of the three possible methods of heat exchange, convective one is undoubtedly predominant [2]. And it is reasonable to assume that a significant part of heat will be removed into the environment from the surface of chips, especially when cutting without the use of lubricating process coolant. This was partly confirmed by visual observation of chips (dark color and brittleness). As a result, this determines the small length of the contact areas, the quick onset of the set mode, and the small variability of the size of the cutting plate due to poor heat exchange [14].

3 Experiment and Results Indicates that processing of austenitic steels of the 12X18H10T type is characterized by a different set of types of contact interaction compared to processing of carbon ones [2]. In this regard, the influence of material properties of the tool/ workpiece pair on the magnitude of the transient velocity, which determines the point of change of types of contact interaction, is of the greatest interest. [2] indicates that a decrease in thermal conductivity of the material being processed and tool material leads to a displacement of the transient velocity toward smaller values. In addition, during advanced plastic deformation, the mechanical strength of the material being processed increases, which, according to research results of [2], also leads to a decrease in the value of the transient velocity. The result is a reduction in the transient velocity from the interaction accompanied by the formation of the build-up to the interaction with the existence of plastic and viscous contacts, as well as to the expansion of the range of existence of the latter type of interaction. It becomes possible to expand the range of applied velocities both toward smaller and toward larger values [15]. The influence of the thermal conductivity of the tool material on the tangential and normal components of the cutting force was studied as well. [2] indicates that with a decrease in the thermal conductivity of the tool material, the length of plastic and full contacts decreases, and the value of tangential stresses somewhat decreases resulting in a decrease in the area of the layer being cut and the tangential component. The thermal conductivity of the tool material has a similar effect on the normal component of the cutting force as well [2]. The use of advanced deformation has a qualitatively identical effect on the lengths of contact areas. Thus, the use of this material and this method should lead to a reduction in cutting force. Notes that a decrease in thermal conductivity of the tool material leads to a displacement in the range of ideal cutting velocities toward large values. Pre-deformation leads to both an increase in the mechanical strength and a decrease in the hardenability of the material being processed. These factors affect the wear curves in different directions. However, in this case, a relative change in these factors is such that the effect of a change in strength is substantially less than the

1208

P. A. Norchenko et al.

effect of a change in hardenability. Thus, the use of advanced deformation leads to a displacement in the range of ideal cutting velocities toward large values and to a decrease in intensity of diffusion wear, which prevails in the range of cutting speeds under study [2]. In the course of the research conducted, these assumptions were confirmed. The cutting force determined by the method of Professor Plotnikov [16] was 148 kg for load-free processing and 146 kg for processing after deformation with a force of 200 kg (cutting depth of 1 mm, feed of 0.1 mm/rev, and cutting velocity of 100 m/ min). For comparison, the use of VK6 material leads to an increase in cutting forces to 170 and 169 kg, respectively. Calculation of cutting force through measuring cutting power confirmed the data obtained. The values of 157 and 154 kg were obtained, respectively [17]. The results of measurements of wear showed a significant increase in the resistance of TN20 plates when processing with the help of advanced plastic deformation. The results are presented in Fig. 1. Moreover, in the course of experiments conducted, a very significant (up to 4 times) improvement in the roughness of the processed surface was observed when using advanced plastic deformation. This effect was identified with the use of hard alloys of various groups, including TH20. The results are shown in Fig. 2. After obtaining significant results in the processing of austenitic steels, it was decided to explore the prospects for the use of the TN20 alloy on processed materials “traditional” for this tool, particularly, low-alloy carbon steels, particularly, steel 20 [18]. It should be noted that the operation of the TN20 tool when processing steel 20 without the use of APD is characterized by a rather high resistance and relatively good quality of the surface layer. The use of APD, as in the case of stainless steels, results in a decrease in thermal conductivity of the material being processed, as a result of which the length of

Fig. 1 Dependence of the wear land along the rear face on the cutting path (velocity of 150 m/ min, feed of 0.128 mm/rev, and cutting depth of 1 mm): 1—cutting without APD, cutting with APD under 200 kg load

Peculiarities of Application of Th20 Hard Alloy for Turning …

1209

Fig. 2 Dependence of surface roughness on cutting velocity (feed of 0.128 mm/rev, cutting depth of 1 mm): 1—cutting without APD, 2—cutting with APD under 200 kg load

plastic and full contacts decreases, and the value of tangential stresses somewhat decreases resulting in a decrease in the area of the layer being cut and the tangential component. However, measurements of thermo emf showed that the relative decrease in thermal conductivity is much less than during hardening of stainless steels. The effect of the use of APD is there in this case, but it is not as significant as when cutting stainless steels. Similarly, the use of APD improves the quality of the surface layer. The reason for reducing the roughness may be the stabilization of the processing process when using APD. As is known, the zone of contact interaction of the cutting tool, forming chips, and the processed surface are characterized by a process of high-velocity plastic deformation [19, 20]. This process is described by physical quantities (degree, velocity, resistance to deformation, temperature, etc.) variable in time. For example, the processing of austenitic-grade stainless steels, which include the 12X18H10T steel used in the experiments, is accompanied by the cycling of chip formation, which is manifested in the traditionally used modes. The instability of the cutting process has an adverse effect on the output processing parameters, including surface roughness. Reducing the plasticity of the processed material after APD favorably changes conditions of contact interaction, which results in less roughness of the processed surface. Generally, the economic effect of the use of cutting with the help of APD for low-carbon steels is not so significant, which means that the prospects for the application of this method for these materials are very negligible. However, there is a process effect, which determines the prospects for application in cases when there are no other options for improving the surface quality, and requirements for this quality are very strict.

1210

P. A. Norchenko et al.

4 Conclusions Thus, it is advisable to draw the following conclusions: • When processing stainless steels of austenitic grade, the use of the TN20 hard alloy is justified from the point of view of the ratio of properties of the material being processed and tool material; • The use of the TN20 alloy when cutting with the help of APD results in a reduction in cutting forces and a significant reduction in wear; • When using the TN20 alloy when cutting with the help of APD, the ideal curves are displaced toward higher cutting velocities, which increases the process efficiency; • In some cases, the surface roughness decreases up to four times when cutting with the help of APD; • The use of this method for processing low-carbon steels leads to a slight increase in tool durability and surface quality, but not to as significant as in the case of stainless steels, which significantly limits the prospects for its use for these materials.

References 1. Kosilova AG, Meshcheryakov RK (1985) Reference book for a technologist and mechanical engineer. In: Kosilova AG, Meshcheryakov RK (eds) Mechanical Engineering, Moscow 2. Talantov NV (1992) Physical fundamentals of the process of cutting, wear, and destruction of the tool. Mechanical Engineering, Moscow 3. Chirchiq Branch of VTIITS (1982) Recommendations for the application of Tungsten-free metal and ceramic carbide alloy “Monitikar.” (1982)/Chirchiq-19, Tashkent region 4. Suslov AG (1987) Technological support of a condition parameters of the details surface layer. Mechanical Engineering, Moscow 5. Yaroslavtsev VM (2015) The efficiency of methods of the advancing deformation hardening of material of the cut-off layer by cutting processing. Bulletin of Bauman Moscow State Technical University. Series: Mechanical Engineering 6. Norchenko PA, Popov AI (2014) On the roughness of the surface layer when processing steel 20 with advanced plastic deformation with TH20 cutting plates. Proceedings of the Volgograd State Technical University 7. Norchenko PA, Popov AI (2014) Prospects for the application of the TN20 hard alloy for finishing work of low-carbon structural steels with the help of advanced plastic deformation. Herald of the Volgograd State Technical University 8. Krainev DV, Norchenko PA, Ingemansson AR (2008) Progressive method of cutting stainless and Heaproof steels and alloys. Eur J Nat Hist 9. Podurayev VN (1974) Cutting of Hard-to-Machine Materials. Higher School, Moscow 10. Podurayev VN, Yaroslavtsev VM, Yaroslavtseva NA (1971) Method of cutting using advanced plastic deformation. Herald of Engineering 11. Sergeyev AS, Zaytseva NG, Plotnikov AL (2012) Mathematical model of a surface roughness formation when steel turning on the basis of thermoEMF operational signal. Processing of metals

Peculiarities of Application of Th20 Hard Alloy for Turning …

1211

12. Plotnikov AL, Krylov EG, Frolov EM (2010) Diagnostics of the state of a multicutter hard alloy tool on the basis of thermoelectric phenomena in the cutting zone. Russian Engineering Research 13. Tchigirinsky JL (2010) Mathematical methods of the control of the machining processes: monography. Volgograd State Technical University, Volgograd, Russia 14. Soldatov AA (2012). Undestructive control of plastic deformation by method of measurement of differential thermoEMF. Defectoscopy 15. Firsov AM, Vdovin AV, Perepelkin PV, Timakhovich IV (2012) Interrelation of parameters of acoustic emission with the modes of cutting and roughness of a surface when turning. Processing of metals 16. Plotnikov AL, Taube AO (2003) Management of cutting conditions on CNC Lathe machines. PKK Politekhnik, Volgograd 17. Miroshin IV (2010) Technique of acoustic-emission researches of the inherited parameters of quality of a blanket in the conditions of cutting and superficial plastic deformation. Bulletin of the Kuzbass state technical university 18. Poduraev VN (1974) Cutting of hard-to-process materials. High school, Moscow 19. Polyanchikov YuN, Kraynev DV, Norchenko PA, Ingemansson AR (2012) Improvement of semi-fair and fair turning corrosion-proof and heat resisting steels by use of the advancing plastic deformation: monograph. ONIKS, Tolyatti 20. Ingemansson AR, Zaytseva NG, Kraynev DV, Bondarev AA (2012) Improvement of quality of processing and mathematical model of formation of a roughness of a surface when turning with advancing plastic deformation. USATU Bulletin

Application of Smoothing Rollers in Processes Finishing–Strengthening Treatment of Shafts’ SPD Yu. I. Sidyakin, S. N. Olshtynsky and S. Y. Abakumova

Abstract The paper proposes a method for improving the quality of the surface of shafts in traditional strengthening treatment by surface plastic deformation (SPD) by using additional smoothing rollers having a small curvature of the working profile in axial section. On the basis of the solution of the elastoplastic contact problem, an analytical dependence is obtained for determining the radius of curvature of the profile of smoothing rollers. The dependences of changes in hardness and yield strength over the depth of the hardened surface layer of steel shafts are experimentally established, and regression equations describing the nature of these changes are given. As an example, the curves of hardness distribution in depth of the plastically deformed shaft layer with a diameter of 50 mm from normalized steel 45 subjected to rolling treatment with toroidal rollers by various forces are presented. Also, this paper presents the results of testing the effectiveness of the use of smoothing rollers in strengthening the roll running toroidal rollers with a roller of a diameter of 40 mm for shafts with materials steel 45, 50, and steel 40X. The authors experimentally confirmed the high efficiency of the proposed methodology in the organization of the technological process of finishing and hardening treatment of shafts, which allows one to significantly improve the quality of their surface.




Keywords Elastoplastic deformation Contact deformation Smoothing roller Hardness An indentor Roughness











Rolling
Roller


1 Introduction By now, extensive evidence has been accumulated in regard to the practical application of the methods of surface plastic deformation (FPD) to increase the durability of machine parts, mainly shafts, with the purpose of increasing their Yu.I. Sidyakin (&)
S. N. Olshtynsky
S. Y. Abakumova Volgograd State Technical University, 28, Lenin Avenue, Volgograd 400005, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_129

1213

1214

Yu.I. Sidyakin et al.

cyclic strength. However, it is equally important to ensure high-quality indicators of the surface itself (optimal microrelief with minimal parameters of waviness and roughness), which is not always possible to obtain in the conventional hardening of shafts with toroidal rollers. After this FPD treatment with the use of one, two, or three rolling devices, the surface of the shafts still demonstrates a regular, albeit weakly evident, undulating profile with small microprotrusions following each other evenly with a step, which is a multiple of the advance path of the working tool. To eliminate these irregularities, it is proposed to simultaneously use a smoothing roller with a smaller curvature of the working surface, which will “overlap” and crush at least 5 or 6 individual microprotrusions (in axial section of the shaft) formed in the initial rolling. Obviously, in this case, a significant improvement in the quality of the processed surface should be expected, and therefore, the task of creating an analytical method for determining the required additional smoothing roller profile radius for the implementation of such a process is quite relevant. The solution to this problem is given hereinafter. The calculation of the radius R of the curvature of the profile of a smoothing barrel-shaped roller is based on the following reasons. First, in a two-roller machine, both rollers (strengthening and smoothing) are installed, as a rule, opposite to each other, and therefore, they load the shaft with the same working force F. The situation is the same for a three-roller machine, which is equipped with two reinforcing rollers positioned at an angle of 120° to each other in the rolling plane and one smoothing roller. Secondly, the smoothing roller under this load should not create residual plastic deformation on the shaft surface, and therefore, the nature of the deformation of the surface layer of the hardened shaft can be considered close to elastic. This circumstance is reinforced by the fact that, due to hardening, the material of the surface layer will have enhanced strength characteristics, in particular, HD* hardness (GOST 18835-73) and yield limit rT (or r0;2 ), and moreover, in this layer, a favorable system of residual compressive stresses is formed, which together impede the emergence and development of new plastic deformations under a smoothing roller. To solve this problem, we use the known dependence [1]: h ¼

F  FS ; pDh HD

ð1Þ

which binds the depth h of residual dent (or roller track) in the shaft surface (hardness HD ) with both a force factor ðF  FS Þ and geometric dimensions of the contacting bodies ðDnp Þ, in which the numerator should be zero, since we assume h ¼ 0: Therefore, the condition for the absence of plastic deformation under the smoothing roller is the following in equation:

Application of Smoothing Rollers in Processes …

F  FS ;

1215

ð2Þ

where FS is the critical load at which plastic deformation occurs in the center of the contact of the roller with the hardened shaft. Previously, M.S. Drozd [2], on the basis of the well-known Hertz problem, obtained a solution to a similar contact problem for the case of a static load compression of a spherical indenter of diameter D and half-space, according to which the value of the critical force FS was determined by the following expression: FS ¼

4p
rT a2y ; 3ð1  2l2 Þ

ð3Þ

and the radius of the contour of the elastic contact area at the time of the onset of plastic deformation at its center was calculated from the dependence [1]: rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi p2 ðk1 þ k2 Þ 3 3p
ðk1 þ k2 ÞFS D ¼ rT D; ay ¼ 8 2ð1  2l2 Þ

ð4Þ

  where additionally rT is the yield limit of a half-space material; k1 ¼ 1  l21 =pE1   and k2 ¼ 1  l22 =pE2 are the elastic constants of the materials of the spherical indenter and half-space, respectively; l1 , l2 and E1 , E2 are their Poisson’s coefficients and elastic moduli. In the case of reinforcing rolling of the shafts by using toroidal rollers, at the initial contact of the tool with the part, a plastic imprint is formed on the surface of the latter, with a contour which in most cases differs from the circular one. Its shape in the plane is close to elliptical and has a major (a) and minor (b) semi-axes, while the latter, oriented, as a rule, along the axis of the shaft, plays an important role in assigning the tool feed S [1, 3–5]. According to studies [1], the values of these semi-axes are calculated via the approximate expressions sufficiently accurate for practical use: rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi h þ ay =2 A0 b20 ½KðeÞ  LðeÞ a¼ and ¼ A B0 LðeÞ  b20 KðeÞ

ð5Þ

where h—the residual displacement of the center of the dent (its depth); A and B— the smaller and larger main curvatures of the surfaces of the roller and shaft at the point of contact, respectively; ay —elastic approach of these bodies in the contact when they are compressed by force F, determined by solving the following incomplete cubic equation: sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 ð k1 þ k2 Þ 2 F 2 3 9p : ay ¼
 8 Dh 1 þ 2h=ay

ð6Þ

1216

Yu.I. Sidyakin et al.

Technological support of this process of SPD shafts with a diameter DB and hardness HD (GOST 18835-73) is theoretically and practically well developed, with the main parameters of the processing modes, including the rolling power F, tool geometry (roller diameter Dr in the rolling plane, profile radius r), which provide for obtaining a predetermined intensity of plastic deformation ei;0 , which is close to the maximum uniform deformation ep of the material, in the surface of the shaft, and the required intensity of residual stresses rocm i , which is determined by calculating via the methods of works [1, 6–16]), in the hardened layer. The diameter Dh of the contacting bodies and their main curvatures A and B in the place of pairingis interconnected by the following relations: Dh ¼ ðA
BÞ0:5 ;

ð7Þ

if A ¼ 1=Dp þ 1=DB and B ¼ 1=2r: When using a barrel-shaped smoothing roller, the geometrical parameters of the contour of the contact zone and the relationship between them will drastically change, since it is assumed that the radius R of its profile (with a constant rolling pattern and working load) will significantly exceed the profile radius r of the reinforcing rollers. In this case, the already large semi-axis a0 of the elastic contact will be oriented along the axis of the shaft, and its value must be calculated by the formula [17] sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 3 3p ðk1 þ k2 ÞF
a0 ¼ na ; 4 ðA0 þ B0 Þ

ð8Þ

where by analogy with (7), the pre-selected value of the diameter DC of the smoothing roller in the rolling plane is as follows: A0 ¼ 1=2R and B0 ¼ 1=DC þ 1=DB ;

ð9Þ

and a coefficient na is calculated by the following dependence: sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ffi   B0 3 2
1þ na ¼
DðeÞ; p A0

ð10Þ

qffiffiffiffiffiffiffiffiffiffiffiffiffi where additionally e ¼ 1  b20 —the eccentricity of the contour of an elastic elliptical print with the ratio of semi-axes, b0 =a0 ¼ b0 ; 1 1 DðeÞ ¼ 2
ðKðeÞ  LðeÞÞ ¼
e 2

Zp 0

sin2 /
d/ pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi; 1  e2 sin2 /

ð11Þ

Application of Smoothing Rollers in Processes …

1217

with KðeÞ and LðeÞ—complete elliptic integrals: Zp=2 KðeÞ ¼ 0

d/ pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi and LðeÞ ¼ 1  e2 sin2 /

Zp=2 qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1  e2 sin2 /
d/;

ð12Þ

0

it is more convenient to find the values of these integrals by expansion in series by a parameter n ¼ ð1  b0 Þ=ð1 þ b0 Þ due to their faster convergence. Now, we solve the problem in relation to the critical load FS . In the elastic stage of deformation, the main normal stresses rx ; ry and ry in the center of the elliptical contact area with the large a0 and small b0 semi-axes, when the axis z coincides with the line of the contact load, i.e., when z ¼ 0; are determined by the following expressions [17, 18]:  rx ¼ p0

   2l2 þ b0 1 þ 2l2 b0 ; ry ¼ p0 and rz ¼ p0 ; 1 þ b0 1 þ b0

ð13Þ

where p0—pressure in the contact center, calculated from the known Hertz–Belyaev dependence: 3 F 3 F ¼
: p0 ¼
2 pa0 b0 2 pa20 b0

ð14Þ

The possibility of plastic deformation on the surface of the shaft is determined under the given conditions by the decision of Henki-Mises [18], according to which the intensity of these stresses should reach values close to the yield limit rT of the deformed material, increased due to the cold work from the previous hardening treatment, i.e., when 1 ri ¼ pffiffiffi 2

qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ðrx  ry Þ2 þ ðry  rz Þ2 þ ðrz  rx Þ2 ¼ rT :

ð15Þ

According to dependencies (13) and (14), Eq. (15) is the following qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1  b0 þ b20 3 ð1  2l2 ÞF
¼ rT ;
2 ð1 þ b0 Þ pa20 b0 from which, under condition (2), critical load FS can be calculated: FS ¼

2p a20 b0 ð1 þ b0 ÞrT qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi : 3ð1  2l2 Þ 1  b0 þ b20

ð16Þ

1218

Yu.I. Sidyakin et al.

Here, however, it should be noted that after the SPD treatment, despite some slight decrease in the elastic modulus [19], the value is also not constant across the thickness of the hardened surface layer, due to the variable and attenuated along the depth of propagation (along the axis) intensity of the contact elastic–plastic deformations [1, 6, 7, 16], and therefore, in the calculations, we are most likely to be talking about some average value of the yield strength within the thickness of the layer. In addition, in this layer, as noted hereinabove, a system of residual compressive stresses is formed, which significantly prevent the emergence and development of new plastic deformations under the smoothing roller [11]. All these factors in their concentrated form has a significant effect on the patterns of change in hardness HV(z). Since there is a fairly stable correlation between the material hardness (in HV, HB, or HD units) and its physicomechanical properties (rT and rB ) [1, 2, 6], the nature of the change within the thickness of the hardened layer can be determined from the hardness distribution HV(z). This dependence is in most cases quite satisfactorily described by the equation: HVðzÞ ¼ HV0 þ DHVð1  zÞm ;

ð17Þ

where DHV ¼ HVP  HV0 —increment of hardness due to SPD, HVP and HV0 — Vickers hardness near the surface and shaft core (initial state), respectively, z ¼ z=zs —relative coordinate (zs —depth of cold work), m—exponent depending on the structural state of the material and its tendency to hardening [20, 21]. The interval of its values, established on the basis of statistical processing of numerous experimental data of our research, as well as the results of other works [3–5, 20– 22], is 1.7 … 2.4. Figure 1 shows, as an example, HV (z) curves for shafts with a diameter of 50 mm from normalized steel 45 (HV0186)   subjected to rolling treatment with toroidal rollers Dp ¼ 110 mm; r ¼ 15 mm by various forces. The value in both cases turned out to be close to 2.0. By analogy with Eq. (16), it is possible to describe in a similar way the nature of the change in the yield limit of the deformed material through the layer thickness by the following dependence: rT ðzÞ ¼ rT þ DrT ð1  zÞm ;

ð18Þ

where DrT ¼ rB  rT , since when the condition ei;0  ep is realized, the surface layers of the shaft must achieve the greatest degree of hardening, i.e., the voltage rB . Taking into account the fact that metal layers adjacent to the surface of the roller experience more resistance to the introduction of a smoothing roller into the hardened surface of the shaft, and that the maximum compressive residual stresses acting in the hardened layer fall at a depth of the following order z ¼ 0:10. . .0:25

Application of Smoothing Rollers in Processes …

1219

Fig. 1 Hardness distribution HV10 in depth z of plastically deformed shaft layer diameter DB = 50 mm from steel 45 (HV0 186) after rolling with the roller (Dp = 110 mm, rp = 15 mm): 1—rolling power of 6.9 kN, 2—rolling power of 9.8 kN

[3–5, 9], it is possible to transform Eq. (17) with m = 2 to a simpler form, which is more convenient for practical application: rTðavÞ

rT  3

  rB 1þ2 : rT

ð19Þ

This ratio can be considered as fair, at least, for normalized and improved high-quality medium-carbon and low-alloy steels. We now turn to the definition of the radius of curvature of the smoothing roller. In the process of traditional rolling with the use of two-roller machines, when the smoothing roller works simultaneously with the strengthening roller with constant pressure on both rollers, the force of the rolling process does not change, i.e., F ¼ FS . In this case, the semi-major axis of the contact area under the smoothing roller is determined by the expression (8), and taking into account (9)–(12), we obtain: a0 ¼

p ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 3 3Rðk1 þ k2 ÞF DðeÞ:

ð20Þ

1220

Yu.I. Sidyakin et al.

Further, from the joint solution of Eqs. (16) and (20), in case rT ¼ rTðavÞ , we find R: 2 R¼4

qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi33=2 pffiffiffiffiffiffi 1  b0 þ b20 3F 5
: ðk1 þ k2 ÞDðeÞ 2pb0 ð1 þ b0 ÞrTðavÞ

ð1  2l2 Þ

ð21Þ

This value R should also satisfy the basic transcendental equation of the general contact problem [20], which binds the ratio of the semi-axes b0 of the contact spot with the ratio of the main curvatures A0 =B0 of the surfaces of bodies, which has the following form: A0 b20 DðeÞ ; ¼ B0 KðeÞ  DðeÞ

ð22Þ

hence, taking into account relations (9), we obtain the second expression for R: R¼

KðeÞ  DðeÞ : 2A b20 DðeÞ

ð23Þ

The joint solution of the system of Eqs. (21) and (23) gives the values R and b0 . For these calculations, it is more convenient to use the MathCAD 7 software package with the subprogram root (y (x), x). The effectiveness of applying the smoothing rollers in the SPD processes with geometrical parameters calculated according to the proposed method was experimentally confirmed on shafts with a diameter DB ¼ 40 mmfrom normalized steel 45 (ND2250, ep  0:11) and 50 (ND2530, ep  0.10), as well as improved steel 40X (ND3250, ep  0.075). The test results are shown in (Table 1). Table 1 Results of testing the effectiveness of the use of smoothing rollers in reinforcing rolling of shafts with the use of toroidal rollers (diameter Dp = 40 mm) Roller’s material

F (kN)

r (mm)

S (mm/ rev)

Smoothing roller a0 Kпep Dc R (µm) –

Ra after Strengthening (µm)

Smoothing

Steel 45 4.5 3.3 0.15 60 415 3.51 47 0.25 0.07 Steel 50 5.1 3.8 0.15 50 350 3.44 46 0.25 0.07 Steel 7.4 6.3 0.20 40 225 3.28 33 0.35 0.10 40X Notes 1. Knep ¼ 2a0 =S—coefficient of overlap, determined by the number of microprotrusions of the relief in the axial section of the hardened surface, “crushed” by a smoothing roller in one pass. 2. The roughness parameter Ra was determined as the arithmetic average of 5 … 8 measurements

Application of Smoothing Rollers in Processes …

1221

The parameters of the modes of hardening treatment were determined by the method of [1, 6–8, 11, 16] with the observance of the condition ei;0  ep on the surface, while the hard layer remained unchanged (about 2 mm) for all shafts.

2 Conclusion The use of smoothing rollers in traditional processes of SPD hardening treatment of shafts can significantly improve the surface quality while ensuring sufficiently high values of the fatigue limits of the shafts and preserving a favorable system of residual compressive stresses from the previous reinforcing rolling in their surface layers.

References 1. Drozd MS, Matlin MM, Sidyakin YI (1986) Engineering calculations of elastoplastic contact deformation. Mashinostroyeniye, Moscow, p 224 2. Drozd MS (1965) Determination of the mechanical properties of the metal without destruction. Metallurgiya, Moscow, p 171 3. Braslavsky VM (1975) Technology of rolling of large parts by rollers. Mashinostroyeniye, Moscow, p 160 4. Chepa PA (1981) Technological basis for parts hardening by surface deformation. Nauka i tekhnika, Minsk, p 128 5. Papshev DD (1978) Finishing-hardening treatment by surface plastic deformation. Mashinostroyeniye, Moscow, p 152 6. Sidyakin YI (2001) Improving the efficiency of rollers hardening by treating with rollers or balls. Vestnik mashinostroyeniya 2:43–49 7. Drozd MS, Sidyakin YI (1984) The method of optimization of technological parameters of the mode of hardening rolling with the use of rollers. Vestnik mashinostroyeniya 1:26–28 8. Sidyakin YI, Osipenko AP, Bocharov DA (2007) Improving the technology of finishing and hardening treatment of shafts by surface plastic. Hardening technologies and coverage (Uprochnyayushchiye tekhnologii i pokrytiya) 8:23–26 9. Sidyakin YI, Trunin AV, Abakumova SY (2014) Analytical study of residual stresses in solid shafts after strengthening treatment by surface plastic deformation. Vestnik mashinostroyeniya 6:62–70 10. Sidyakin YI, Trunin AV, Shevtsov AN (2010) Spherical model of contact elastoplastic deformation studying. Izvestiya VolgGTU 12(72):48–52 11. Sidyakin YI, Ivanov SV, Olshtynskiy SN, Shchipetyev DA (2010) Effect of residual stresses on depth of cold work. Izvestiya VolgGTU 12(72):45–48 12. Oteniy YN, Olshtynskiy SN, Olshtynskiy NV (2007) Using kinematics of the points of a part’s surface in plastic deformation with the use of rollers. Izvestiya VolgGTU 4(3):61–62 13. Oteniy YN, Olshtynskiy SN, Olshtynskiy NV, Shchegolev NG (2010) Automation of the evaluation of parameters of the elastoplastic deformation zone when machining adjacent surfaces, assamblying in engineering and tool making (Sborka v mashinostroyenii I priborostroyenii) 4:21–24 14. Olshtynskiy SN, Oteniy YN (2007) Features of the formation of the depth of hardening during machining of parts via surface plastic deformation. Izvestiya VolgGTU 4(3):63–66

1222

Yu.I. Sidyakin et al.

15. Oteniy YN, Olshtynskiy SN, Olshtynskiy NV, Shchegolev NG (2010) Features of determining the rational parameters of deforming elements and technological modes of treatment during surface plastic deformation. Hardening technologies and coverage (Uprochnyayushchiye tekhnologii i pokrytiya) 12:3–6 16. Sidyakin YI, Abakumova SY, Vdovenko AV, Konovalova YG, Strukov VA (2016) Prediction of the intensity of the elastoplastic contact deformation in the surface layer of the shafts with treating by SPD rollers. Izvestiya VolgGTU 14(193):33–37 17. Ponomaryev SD, Biderman VL, Likharev KK, Makushin VM et al (1958, 1959) Calculating hardness in engineering. Mashgiz, books, Moscow 2(3):1118 18. Bezukhov NI (1988) Fundamentals of the theory of elasticity, plasticity, and creep flow. Vysshaya shkola, Moscow, p 512 19. Guryev AV, Nosko IN (1976) The effect of PPD on the change of physico-mechanical properties of the surface layer of metal. KPtI, Kuybyshev, pp 8–13 20. Kudryavtsev IV, Rymynova EV (1961) Increased hardness and fatigue strength as a result of cold-hardening of steels with various structures. TsNIITMASh, Moscow 21. Papshev DD, Stulnikov VK (1969) On the nature of hardness in steel rolling. Mekhanika, Kuybyshev, pp 313–314 22. Yankov NI (1977) Hardeninig the surface layer of machine parts up to the maximum hardness. Priborostroyeniye, Minsk, pp 71–74

Preparing Automated of Software Complex for Technological Processes with Imposition of Electric Field O. V. Skrygin, V. P. Smolentsev and E. A. Saltanaeva

Abstract The chapter examines the processes of creating control systems for electrical processing methods that can be based on the principle of contact-free movement of the tool electrode’s working surface in relation to the processed one. The possibilities of tool wear and poorly formalized change in the inter-electrode gap are taken into consideration. For this purpose, it is proposed to apply the statistical methods of process investigation, using the results as boundary conditions in modelling shifts along the controlling coordinate, within the limits laid in CNC machine tools for electro-erosion, electrochemical and combined processing of metal materials, including the most common equipment with an unprofiled wire electrode. The paper presents the examples of using the developed software products for electro-erosion and combined processing of typical complex profiled parts, specific for small-scale flexible manufacturing in modern science-based machine construction. The area of perspective use in various branches of mechanical engineering of the technique of design of control programs for the combined methods of processing with imposing of an electric field offered in the article is opened. Keywords Combined processing Control systems


Electrode tool
Non-profiled electrode


1 Introduction In the process of developing control programs for CNC metal processing tools [1, 2], the true position of the tool cutting edge corresponds to the point of its contact with a part. The wear and pressing of the tool by cutting force are also taken into account. Such approach proved to be unsuitable for the development of control programs for operations using electrical processing methods since, in this case, O. V. Skrygin
V. P. Smolentsev (&)
E. A. Saltanaeva Voronezh State Technical University, 14, Moscow Pr, Voronezh 394026, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_130

1223

1224

O. V. Skrygin et al.

it can involve the variable wear of electrodes (e.g. it cannot occur in electrochemical dimensional processing), and the force of tool electrode pressure upon a part is negligible (e.g. in combined processing). Besides, in some methods, the process does not involve the contact of an electrode with the processing area, and it occurs when the tool is removed from the part. Here, the variable clearance must be considered in calculating the tool shifts since its value is commensurate with part tolerance, or can significantly exceed its value, which impacts the accuracy of the part. The attempts of using tool path converters, adopted in mechanical processing [2] for contact-free formation of products, have shown that the deviation of the part contour from the working part of the tool electrode can significantly exceed the permissible tolerance limit, and, in case of electrical processing methods, the process fades out as the clearance grows larger. The mathematical software developed in recent years (mainly, by foreign companies) largely refers to mechanical processing at the level of reference models, covering a small proportion of products, which requires considerable time for transferring the original parts to machine tools with computerized control systems, and provides almost no automated technological preparation of equipment control programs for electrical processing methods. Therefore, the problem of creating the methodology of designing software for electro-erosion, electrochemical and combined equipment is urgent for machine construction. It becomes especially acute with launching the production of such machine tools in Russia. The means of technological preparation, obtained from abroad, take little account of technical and economic requirements of sectoral machine construction (in particular, of aerospace industry). Thus, it is required to carry out the research on developing the methodology of designing control programs, with an account of the specificity of shaping processes in some cases, with no metal contact between a tool (electrode) and a workpiece. It is also necessary to take account of the long-term experience of using such research for metal-cutting systems, some of which are described in [3, 4]. To design the programs for combined processing, the materials of [5, 6] can be used.

2 The Features of Shape Formation of Parts When Using Electrophysical and Electrochemical Processing Methods The principal feature of electrical processing methods is the removal of material from a workpiece without direct contact of the tool with the part (i.e. electrospark, pulse or electrochemical processing), or with non-shape-forming mechanical impact upon the item processed (i.e. electrical contact, combined method), based on some non-mechanical process (thermal, chemical, magnetic, nuclear, etc.). Hence, it is possible to use these methods for manufacturing the parts without force action

Preparing Automated of Software Complex for Technological …

1225

(or with insignificant pressure) of the tool, with minimal displacement of elements of the technological system (as a rule, one or two control programs for tool movement are sufficient for processing by these methods). It helps to process non-rigid part of modern machinery with limited access of electrodes to the processing zone. Figure 1 shows the example of obtaining hard machining deep holes of small diameter. The mechanical manufacturing of such a part will require the creation of an expensive jig-type device of the conductor type to direct the drill (especially, when it touches the workpiece), the repeated controlled movement of the cutting tool along several coordinates, deburring and a number of other labour-intensive operations, which entail additional operation costs. Besides, due to change in the outer diameter of the part (within tolerance), the safety of small-diameter drills (less than 2 mm) is not ensured because of the gaps occurring between the conductor and the outer surface of the part. When using the electro-erosion or erosion-chemical piercing of holes like those, shown in Fig. 1, it is necessary to specify (or, in some cases, introduce) the adaptation of electrical (and, sometimes, hydraulic) process modes. Such programming does not cause any problems even for equipment having no CNC units; however, in this case, the complexity of production preparation significantly increases. To automate the preparation of control programs, it is required to establish the adaptive links to take account of the actual tool electrode wear in electro-erosion

Fig. 1 Housing with holes, pierced at an angle to the cylindrical surface by the electro-erosion method using the developed control programs

1226

O. V. Skrygin et al.

processing, but, in most cases, this can be avoided by calibrating the obtained profile with an unworn electrode part, which must be also considered in evaluating the depth of tool advance. Here, the motion is limited by the portion of the inner hole, opposite to the working electrode end (Fig. 1). The methodology of designing control programs for mechanical processing methods, including the use of edge and abrasive tools, is fairy fully described in [7, 8]. Many provisions of this methodology are valid for electrical processing methods; however, one must consider the dynamics of variation in inter-electrode gaps and uneven wear of the tool electrode, if any. As concerns the processing zone heating, this factor is almost absent in the process under consideration; therefore, in a general view, the model takes the following form: 8 d2 YðtÞ dYðtÞ dX dY dZ > > > m þ cYðtÞ ¼ FðX; ; Y; ; Z; ; pÞ; þh > 2 > dt dt dt dt dt > > > 2 > > d ZðtÞ dZðtÞ dX dY dZ > > M þ CZðtÞ ¼ FðX; ; Y; ; Z; ; pÞ; þH > 2 > dt dt dt dt dt > > > > A Z > < pðiÞ ðAÞ ¼ pi;0 þ pi wPi ðA  sÞNðsÞds; i ¼ 1; 2. . .s; > > > 0 > > > t > Z > > > > > AðtÞ ¼ NðtÞdt; > > > > > 0 > > : NðtÞ ¼ Vp jF2 ðtÞj: ð1Þ     where m, M are diagonal matrices of 3  3 in size; c ¼ cs;k , C ¼ Cs;k are positive symmetric matrices of changing the position and geometry in the subsystem of the tool electrode and the workpiece of 3  3 in size, which are unchanged with reference to coordinates of the tool movement  and  when the equilibrium point of the system is displaced; h ¼ hs;k , H ¼ Hs;k are positive symmetric matrices of dissipation for the subsystem of the tool electrode and the workpiece of 3  3 in size, also unchanged with reference to coordinates of motion of the tool’s working part and when the equilibrium point of the system is displaced; wPi ðA  sÞ ¼ exp½ T1P ðA  sÞ are kernels of integral operators; Vp is i

the speed of tool advance, which can be considered constant after process stabilization; p ¼ fp1 ; p2 ; . . .ps g is the dynamic characteristic of the tool electrode wear; N is the number of manipulations with the tool (switching, changes, mode parameters), necessary to obtain a high-quality part; F is the function; A is total processing part; T0 is the total processing time in one mode; t is the actual moment of time; and w0 is the mode processing parameter.

Preparing Automated of Software Complex for Technological …

1227

3 The Efficiency Assessment of Using the System of CNC Machine Tools with Wire Electrodes Figure 2 represents the lodgement of 40X steel, on which it is required to obtain the profile with various tilt angles of lateral sides, which needs the establishment of the controlling coordinate, the action of which is limited by angular displacement (up to +−15° in modern electro-erosion equipment). Figure 3 shows the control program, created in accordance with the proposed methodology, with explanations for users %“24713”—the start of the program, marked with the %. The text in  the comment (the program code) N5 G20—the assignment of coordinates in microns (the coordinates are displayed without a point) N10 G92 X8000 Y60000—setting the zero point of the program N15 G42 G01 X8000 Y58000 G101 X8000 Y58000 G50—Activation of the radial tool adjustment on the right, the simultaneous access to the contour without adjustment (G50) along the right (G01) and the left plane (G101) N20 G01 X6000 Y58000 G101 X6000 Y58000—Linear motion with activated mode of adjustment along two planes simultaneously N25 G01 X-18875 Y58000 G101 X-18875 Y58000 N30 G02 X-19000 Y58125 I0 J125 G102 X-19000 Y58125 I0 J125—The clockwise circular motion along two planes: (G02)—lower plane, (G102)—upper plane

Fig. 2 Lodgement. a—Drawing of the part; b—General view

1228

O. V. Skrygin et al.

Fig. 3 Formation of the edge rounding radius for outer (R1 ) and inner и (R2 ) transitional sections with an angle a (the workpiece material is steel, the inter-electrode gap is 0.4 mm): a—the diagram of the transitional section; b—the dynamics of edge shape formation

N35 G01 X-19000 Y72000 G101 X-19000 Y72000 N40 G01 X-36297 Y72000 G101 X-36297 Y72000 N45 G01 X-42460 Y56745 G101 X-42460 Y56745 N50 G02 X-19155 Y-26522 I-42340 J-56745 G102 X-19155 Y-26522 I-42340 J-56745 N55 G01 X-41762 Y-82475 G101 X-33512 Y-62056 N60 G03 X-41848 Y-83375 I1470 J-594 G103 X-33598 Y-62955 I1470 J-594 N65 G03 X-41418 Y-84165 I1472 J289 G103 X-33168 Y-63745 I1472 J289 N70 G01 X-35329 Y-90044 G101 X-27079 Y-69625 N75 G01 X-27981 Y-82127 G101 X-19731 Y-61708 N80 G02 X-26632 Y-81161 I3044 J-2825 G102 X-18382 Y-60742 I3044 J-2825 N85 G02 X-25011 Y-80815 I1620 J-3626 G102 X-16761 Y-60396 I1620 J-3626 N90 G01 X6000 Y-80815 G101 X6000 Y-60396 N95 G01 X8000 Y-80815 G101 X8000 Y-60396 N100 G40 G01 X8000 Y-82815 G101 X8000 Y-62396 G50—The cancellation of radial tool adjustment, the retreat from the contour without 2-plane adjustment %—the symbol of the program end The program was created for the machine tool ARTA 450 PRO with a non-profiled electrode when splitting the material with the use of the following technological modes:

Preparing Automated of Software Complex for Technological …

Voltage Current Wire tension Wire rewinding speed Frequency of pulse generation Operating pulse duration Wire feed for plunging

1229

250 V 0.8 A 20 N 4 m/min 40 kHz 0.38 nks 1, 5 …2 mm/min

The use of this method helps to automate the process of preparing the controlling programs and reduce labour intensity from several hours to 10–15 min.

4 The Control of Channel Geometry Change by an Electrochemical Method In case of electrochemical dimensional processing, the tool is not worn out; therefore, with reference to this parameter, adaptive signals and a control channel are not required. The special case consists in manufacturing of parts by unbound pellets using abrasive (i.e. hydroabrasive workpiece splitting), or smooth pellets (i.e. electrochemical processing with a crumbling tool), in which the creation of multi-coordinate control programs makes it possible to introduce [5, 6] a new scientific area, i.e. remote shaping of parts of any profile from metal blanks. From [5], it is known that, when controlling the removal of metal with reference to processing time, edge rounding of the type, shown in Fig. 3, occurs in the transitional sections. For right angles (a ¼ 90 in Fig. 3), the rounding radius during processing (usually, no more than several minutes) can reach 1 mm. Bearing in mind the regularities of the rounding radius change, we can control the obtainment of the desired radius with reference to time, i.e. obtain a completely new technological result, namely the piercing of circular openings with a tool electrode of the rectangular (e.g. square) cross section. This is relevant for rocket engine filters, in which the diameter of openings is less than 1 mm, and the formation of such surfaces by square electrodes produces a great economic effect since they are quite simply formed using the electro-erosion machine tools with a non-profiled electrode, by dividing the workpiece into dimensional sections by part of its length [9].

5 The Application of the Method of Path Converters for Creating Control Programs In [2], the creation of control programs is based on the principle of analogue modelling using electrical systems (for example, via equivalent resistance). This parameter is part of all performance rates for processing methods with the imposition of an electric field.

1230

O. V. Skrygin et al.

In [2], it is proposed to use the Lissajous curves, i.e. closed-loop movements of the point in two mutually perpendicular directions for plotting the trajectory of the tool shift (in our case, an electrode), which makes it possible to describe the coordinates in any system of shift modelling, with account of the clearance between the tool and the processing zone. The analogue of such a system is automatic control of an electron or light beam in splitting workpieces made from sheet materials. The expression of Lissajous curves via electric equivalents [2] allows for an approximation of algebraic cycloid curves by amplitude-frequency-phase parameters. Then, it is possible to use the system of one-parameter equations, applied for Lissajous curves, to describe the paths of reciprocal (contact and non-contact) motion of the tool and the workpiece. For this purpose, it will be necessary to set the frequency coefficients of relevant facilities. According to Razi [2], such dependencies provide the possibility to obtain linear, circular and elliptic trajectories. The present-day equipment for program planning makes it possible, with a limited number of one-parameter estimating equations, to set the trajectories for a large number of sections with high accuracy, through the transition by command programming via frequency coefficients, which are sufficiently stable for electric methods of processing. The alteration of initial phases allows us to describe not only the geometrical images but also the analytical properties of objects (particularly, the performance rates, the wear-related variations in shape formation zones on the part and the electrode and the fluctuations of the inter-electrode gap value). The variation of frequency terms in one-parameter equation systems enables us to consider the time transformations, which is quite significant, for instance, in non-contact electrochemical processing by fixed electrodes. Such a process is used to fine-tune the geometry of adjacent channels of products, where there is a need to produce a coupling from piecewise smooth sections, with a profile deviating from the estimated one by not more than 20–30 lm. In this case, the shape of the curve sections is weakly predictable. To achieve high accuracy, it is required to know the patterns of inter-electrode gap variation, which are stochastic and difficult to calculate. For this purpose, experimental data are used which allow setting the limits of parameter changes in adaptive process control. In [2], it is proposed to describe the trajectories of the tool’s motion by a system of one-parameter equations that give a fairly full picture of their geometric and analytical properties. In a rectangular coordinate system ðx; yÞ, the controlled coordinates ðx; yÞ in a two-coordinate system used in the electro-erosion and combined splitting of current-carrying materials xb , yb are calculated by dependences xb ¼ A0

m X

ð1Þ

ð1Þ

ð2Þ

ð2Þ

½ak cosðk ð1Þ a þ a01 Þ bk sinðk ð2Þ a þ a02 Þ

i¼1

yb ¼ B0

m X i¼1

ð2Þ ½ak cosðk ð3Þ a þ a03 Þ bk sinðk ð4Þ a þ a04 Þ

Preparing Automated of Software Complex for Technological …

1231 ðnÞ

ðnÞ

where A0 ; B0 is the distance in the reference frame from the origin; ak , bk are the amplitudes of harmonics of the first order; k ðnÞ are the frequencies of harmonic base terms; a is the independent current parameter ð0 a p2Þ; a0 is the initial parameter value; and m is the number of independent sections.

6 Conclusion The described specific nature of creating the control system of part processing by electro-erosion and electrochemical CNC machine tools shows that the purpose-oriented use of standard software products, intended for mechanical processing includes only some procedures, primarily, those, in which the absence of direct contact of the tool with the workpiece does not significantly affect the shape formation [10]. The methods of path converters allow us to consider not only the geometrical parameters but also the analytical factors, determined by the modes of an electric field and hydrodynamic properties, impacting the technological parameters of electric processing methods [11, 12].

References 1. Brailov IG, Chasovskikh AI (1998) The modelling of shape formation by CNC machine tools. The Publishing House of VSTU, Voronezh 2. Razi AA (1984) Path converters. Mashinostroenie, Leningrad 3. Ivakhnenko AG, Kutz VV (2013) The pre-project studies of metal-cutting systems: monograph. South-West State University, Кursk 4. Safonov SV (2015) The criterial system of planning and using the technological processes for enhancing the performance characteristics of the surface layer. The Bulletin of Voronezh State Technical University 11(3):4–10 5. Smolentsev EV (2005) The planning of electrical and combined processing methods. Mashinostroenie, Moscow 6. Kuzovkin AV (2001) The combined processing by an unbound electrode. Publishing House of Voronezh State University, Voronezh 7. Zakovorotny VL, Flek MB (1997) The definition of invariant diversity of trajectories of shape-forming movements. The designing of technological machines: the collection of scientific works, MSTU “Stankin”, Moscow, issue 4 8. Flek MB (1998) The account of evolutionary transformations in shape-forming movement control. Designing technological machines: the collection of scientific works, MSTU “Stankin”, Moscow, issue 9 9. Smolentsev VP, Goncharov EV (2012) Enhancing the accuracy of hydroabrasive processing and quality of the surface layer in the separation zone of viscious materials. Strengthening technologies and coatings 4:45–48 10. Suslov AG (2012) High technologies in mechanical engineering. Mashinostroenie, Moscow 11. Suslov AG (2000) Engineering. Engineering, T- Sh-3, Moscow 12. Vasiliev AS, Kutina AA (2018), Handbook of mechanical engineer, vol 2. Innovative mechanical engineering, Moscow

Technology of Combined Chemical–Mechanical Processing V. P. Smolentsev, V. V. Ivanov and E. V. Panichev

Abstract The combined process for chemical–mechanical processing of durable mechanical engineering products has been designed. In describing the process and for creating the model, we set forth a single criterion and substantiate the scientific hypothesis on a generalized method of controlling the effects upon the flow of chemical change through energy potential of process components themselves or by means of external supply of mechanical energy. It is shown that the supplied external mechanical energy can compensate the costs of chemical reactions, accelerate the combined processes, obtain a new state of coatings, resulting from chemical transformations, and intensify the transformations as to achieve high technological parameters (namely, durability), providing the operational properties, required by a customer and a developer of new equipment. The article describes the ways of controlling external and internal impacts upon the combined process of coating deposition and gives the examples of their use. It is shown that the most common cases of implementing the developed technologies will suffice the energy of pulsed external strikes of hard pellets, similar to processes of vibroimpact hardening, mastered by enterprises and scientific institutions of the country.







Keywords Technology Modeling Criteria estimation Chemical-mechanical coatings Resource





Combined methods


1 Introduction The wave processes associated with the vibroimpact method of hardening treatment make it possible [1–5] to enhance the technological parameters, obtained by consistent use of chemical and mechanical impacts. However, the fullest use of the V. P. Smolentsev (&)
E. V. Panichev Voronezh State Technical University, 14, Moscow Ave, Voronezh 394026, Russia e-mail: [email protected] V. V. Ivanov Don State Technical University, 1, Gagarin Sq., Rostov-on-Don 344010, Russia © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_131

1233

1234

V. P. Smolentsev et al.

potential of these impacts is only possible with their combination in an integrated way of treatment, described in [6–9]. The results of this research helped to create the mechanism of increasing technological durability of products, which is demanded in industry and relevant for present-day domestic and foreign machine construction [10, 11].

2 The Mechanism of Combined Chemical–Mechanical Processing The supplied external mechanical energy can compensate the costs of chemical reactions, accelerate the combined processes, obtain a new state of coatings, resulting from chemical transformations, and intensify the transformations as to achieve high technological parameters (namely durability), providing the operational properties, required by a customer and a developer of new equipment. The energy approach provides the creation of a generalized model, which serves as a basis for the methodology of designing technological processes of depositing combined high-resource coatings and developing the processing methods for standard parts. In the general form, the model can be represented as: Ua [ UMg  Uxg

ð1Þ

where Ua is energy, required for a chemical reaction to flow; UMg is mechanical energy, supplied from outside; Uxg is additional energy, supplied to accelerate the chemical reaction (the «+» sign), or expended (the «−» sign) to compensate for losses during the flow of the chemical reaction. The criterion: UMg  UMmax

ð2Þ

can serve as the boundary condition of the model, in which UMmax is the maximum energy potential of the mechanical component of the process. The analysis of the criteria (1) and (2) demonstrates the necessity for setting the options of chemical impact, whereby the internal energy will have the highest rate, providing the flow of an intensive chemical reaction to achieve the required technological parameter. Then, the external mechanical energy can be provided by its regulated supply to the processing zone, for example, by a wave process. In this case, the energy loss (Uxg ) in the Formula (1) can be reduced that will increase the share of energy (UMg ) for intensification of chemical reactions and result in significant energy saving of the combined process.

Technology of Combined Chemical–Mechanical Processing

1235

The above criteria (1) and (2) allow us to analyze the predicted capabilities of known and newly created kinds of combined, highly durable chemical–mechanical impacts, and to substantiate the most effective ways of typical coating for investigating such methods.

3 The Systematization of Technological Processes for Obtaining Coatings in Various Mechanisms of Their Formation The technology of chemical–mechanical formation of coatings with vibrowave supply of pulse energy differs from other coating methods, because the deposition of coatings occurs under the simultaneous flow of chemical reactions and vibrowave shock action of hard pellets. The contact of operating medium (pellets) with the workpiece surface results in activation of chemical processes, flowing in the surface layer during coating formation, accompanied by modification in the geometric and physical–mechanical characteristics of the material. Based on the conducted analysis and research in the field of vibrational mechanical–chemical coating deposition, a classification has been developed, depending on the energy capacity of the system (i.e., the nature and intensity of processes occurring between the base and the coating): 1. At the molecular level, the coating material is resistant to deformation, possessing chemical and thermal stability. The pattern of the formation of such coatings is the impact of operating medium in maximum processing modes. The metal balls of various diameters are used as operating media. 2. At the electronic and ionic level, the base and the coating material interact at the energy level, making a firm connection. The pattern of the formation of such coatings is the joint impact of two energy types, mechanical and chemical ones, simultaneously responsible for the process of coating formation. The modes of equipment operation are moderate ones, using various materials as operating media (e.g., porcelain, glass, ceramic, and metal balls). The coating is formed on the surface of the metal. 3. At the chemical level, the base metal chemically interacts with the process solution, forming a coating. The basis of the formation of such coatings is the chemical component of the process. The time of coating formation must be several times less than the speed of metal removal. In this case, the mechanical component contributes to activation of the process. The technological modes of equipment operation are lenient, in which the oscillation range is 1–3 mm, and the frequency is 16–30 Hz. The materials of operating media are polyethylene, plastic, and glass of various shapes. The growth of the coating occurs depthwise in the metal. Figure 1 represents the characteristics of the energy level of the most common coatings.

1236

V. P. Smolentsev et al.

Fig. 1 Assessment of the energy level of vibrational chemical-mechanical coating

The energy level of vibrational chemical–mechanical coating VCMC (Fig. 1) is determined by the contribution of energy, caused by entropy growth, to the process of surface layer formation (the first group of impacts), the increase in the energy of elastic–plastic distortions of the crystal lattice resulting from mechanical impact of indentors (the second group), and the change in internal energy of the surface layer, modified as a result of chemical interaction between the contacting media (the third group). The role of each type of impacts in the kinetics of formation of the classification groups of chemical–mechanical coatings is described by energy and standard models, from which it can be seen that the deformation component, monotonically growing with time, and makes the main contribution to the process of VCMC formation of the first group. The creation of VCMC of the second group is the result of deformation and chemical components. In forming the VCMC of the third group, the leading role belongs to the chemical component of the model. The key quality parameter (Fig. 1) of a durable coating, deposited upon the metal surface, is adhesion which characterizes the strength of bonding between two materials and the emergence of a bond between surface layers of two heterogeneous substances, brought into contact. If we equate the mole energy, determining the conditions for creating the modified local microvolume at the

Technology of Combined Chemical–Mechanical Processing

1237

«coating—substrate» interface, to the average energy of the bonding unit, ensuring its adhesion, the energy model becomes the basis for the calculation and analytical model of the technological system of VCMC, which helps to obtain durable coating on the surface of the material, required by operating conditions of products. The process of creating highly durable coatings is a complex set of mechanical, physical, and chemical phenomena, significantly impacting the state of the surface layer of the processed part and the formation of coatings. When depositing the VCMC, the nature of mechanical and physical–mechanical phenomena is determined by the processing technology, the type of coating, and the physical–chemical properties of the coating material and the kinetic regularities of its formation. The main parameters of mechanical impact of the combined process are the speed and force of pellets of the operating medium, striking the surface of parts, and contact pressure in the collision zone. To determine the possibility of depositing the VCMC, the paper defines the value of specified parameters. The mechanical component of the chemical–mechanical processing energy can be represented as: UM ¼ A
Pmax
su

ð3Þ

where Pmax —is the maximum collision force of collision, N. According to [5] Pmax

pffiffiffiffiffiffiffiffiffiffiffiffiffi m1
x20
x
Kp 1  n20  ¼ pffiffiffiffiffiffiffiffiffiffiffiffiffi en0
x0
su sin x0
su 1  n20

ð4Þ

where su is the collision time, sec; Kp is the coefficient of operating medium loosening; m1 —the reduced mass of a pellet of the operating medium, g; x is the vibrating mass transfer, mm. Here, the angular frequency x0 is calculated as: rffiffiffiffiffiffiffiffiffiffiffiffi m1
n x0 ¼ K2

ð5Þ

where K2 is the coefficient of resistance of vibrating operating medium. su ¼

p pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi x0 m1
n
C

ð6Þ

where C is the coefficient, taking into consideration the operating mass activity (part of the kinetic energy of a strike, used to transfer the operating medium). C¼

m2 m1 þ m2

ð7Þ

1238

V. P. Smolentsev et al.

Fig. 2 Classification of VCMC

where m2 is the mass of the part. K0 n0 ¼ pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 m1
n
C

ð8Þ

where K0 is the coefficient of resistance to the transfer of vibrating operating medium. The distinctive feature of VCMC, as compared with other methods of coating deposition, is the fact that their deposition is performed with the use of combined technologies, under the mechanical vibrational striking pulse action of pellets of the operating medium. The dynamic impact provides the activation of chemical and physical–chemical processes, occurring in the surface layer, which is accompanied by the change of its geometrical and physical–mechanical characteristics. VCMC can be divided into three types (Fig. 2). From the classification, presented in Fig. 2, it follows that some coatings are formed by mechanical rapprochement of the coating material with the metal surface and are kept by Van der Waals forces. Others are formed in conditions of a chemical reaction and, in the presence of electrostatic forces in the boundary layer, the process is complicated with the formation of a double electric layer and hydration. However, for all VCMC, regardless of the complexity of chemical processes, the mechanical energy always remains the main activating force. On the basis of the proposed model, the analysis of literary sources and experimental research, the physical models of single-component coating formation have been developed.

4 Conclusion The article describes the technologies of combined deposition of chemical–mechanical coatings, which made it possible to deposit surface layers to obtain the products with the required level of claimed durability.

Technology of Combined Chemical–Mechanical Processing

1239

We have developed the mechanism of energy control of the combined process, involving the influence upon chemical and mechanical components of external and internal impacts, which allowed us to make full use of the advantages of incoming impacts in a single process. On the basis of the energy approach, the classification of chemical–mechanical coatings has been worked out, substantiating the main fields of research for typical, high-resource products, which gave the opportunity to specify the area of highly effective implementation of new coatings.

References 1. Dyachek IM, Volobuev NК, Ermilov AS et al (1969) The deposition of a solid film of molybdenum disulphide upon the surface of rolling bearing parts using ultrasound. Plastic lubricants and solid lubricating coatings. The proceedings of the All-Russian Scientific and Research Institute of Oil Processing. Khimiya, Moscow 2. Efremov IK, Bolotov NL (1989) The features of oxide layer formation on aluminium during microarc oxidation in an alternating electric field. Phys Chem Mech Mater 25(3) 3. Zhuravleva LA, Chernetskaya NB, Kolodyazhnaya LG et al (2009) Zinc coating deposition in a two-phase medium during vibrational processing. East Eur J Adv Technol 2/1:13–17. Kharkov 4. Zhuravleva LA, Kolodyazhny PV (2009) The technology of zinc coating formation on parts made from carbon steels in conditions of vibrational processing. Hardening Technol Coat 3:24–28 5. Kirichek AV, Afonin AN (2009) The progressive technologies of hardening and shape formation by roll threading. High-Tech Technol Mach Constr 9:3–8 6. Ivanov VV (2007) Vibrational mechanical-chemical methods of coating deposition (oxidation). The Publishing Centre of Don State Technical University, Rostov-on-Don 7. Ivanov VV (2010) Vibrational mechanical-chemical methods of coating deposition (zinc coating). The Publishing Centre of Don State Technical University 8. Smolentsev EV (2005) The designing of electrical and combined processing methods. Mashinostroenie, Moscow, p 511 9. Smolentsev VP (ed) (1983) The electrophysical and electrochemical methods of material processing, vol 2, V.1. Vyschaya Shkola, Moscow, p 247 10. Ivanov VV, Babichev AP (2014) Vibrational mechanical-chemical coatings. LAP LAMBERT Academic Publishing is a trademark of: Omni Scriptum GmbH & Co. KG. Heinrich-Böcking-Str. 6–8, 66121. Saarbrucken, Germany, p 119 11. Kopylov YR (1990) The process dynamics and the technology of vibroimpact hardening of parts of complex geometry. Doctoral thesis, Technical Science, Voronezh, p 387

Technology of Combined Treatment of Engine Cooling Elements V. P. Smolensev, A. V. Shchednov and J. S. Smolenseva

Abstract The chapter considers ways of creating heat exchangers with cooling elements, in particular, vortex generators in the form of local roughness on heat exchanging surfaces that distract boundary layer of fluid, gaseous or multiphase coolant. Vortex generators are widely used in engine engineering because they allow significantly increasing the intensity of heat exchange between a cooling object and a bordering wall of a cooling channel with some increase in its hydraulic resistance. The examples of using vortex generators are thermally stressed elements of modern engines (jet aviation engines, liquid propellant engines, etc.). In these engines using such elements allow significantly improve the product performance indicators. It requires the creation of effective technological processes of intensification heat exchange. Intensification comes at the expense of increasing the area of heat exchange and using vortex generators on surfaces of walls of cooling channels in heat exchangers. The examples of such technological processes are electrical processing methods. The combined method is the most promising of these methods. The article offers a new way of the generation heat exchanging section and vortex generators in cooling channels of heat exchangers with local indentations. This way will be illustrated by a shaped channel that widely used, for example, in liquid propellant engines. For this purpose electric-erosive, electrochemical treatment or, in the future, combined electric-erosive-chemical treatment. This makes it possible to manufacture vortex generators not only on the surface of a “fire” wall of the channel as at present but also on the lateral surfaces of ribs that provides additional intensification of cooling.




Keywords Cooling elements Vortex generators Channels Electrochemical treatment





Technological processes


V. P. Smolensev (&)
A. V. Shchednov
J. S. Smolenseva Voronezh State Technical University, 14, Moscow Av., Voronezh 394026, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_132

1241

1242

V. P. Smolensev et al.

1 Introduction Cooling efficiency of the elements of the hot zone of heat engines (combustion chamber, reactive nozzles, afterburning turbofans, injectors) determines the service life and operational reliability of the product and its competitiveness on the local and global market. The designers of new generations of technique offer for this purpose both the known and the new design solutions that need for implementation creation of original technological methods on the level of inventions. Some of these methods are considered in the article [1–4].

2 Local Cooling of Thermally Stressed Area When injectors are working, high temperature is formed around a channel of flowing combustible fluid. It causes injectors to overheat and failure. In [1] technological method of creation, artificial roughness on the front end around the flame of burning fuel is offered. Figure 1 below illustrates the scheme of creation such roughness in the shape of spherical indentations. These indentations obtained in the process of electric-erosive treatment. In the first phase (Fig. 1a), working part of tool 2 for its electrical discharge treatment is processed by electropulse method with electrode 1 (anode) in draft

Fig. 1 Method of making local section for intensification of cooling a—making of electrode tool; b—creation of artificial roughness on detail

Technology of Combined Treatment of Engine Cooling Elements

1243

mode according to a circuit with reverse polarity. Erosion-resisted graphite compositions can be applied as a material for electrode 1. These compositions provide generation on a tool 2 indentations 6 500–600 µm depth. In doing so, the working surface of electrode 1 can be flat (5 on Fig. 1a). This process takes place in dielectric liquid 4 in the energy of impulses above 3–4 J impulses produced by a generator 3. In this operating mode, defective layer with a web of cracks is formed in indentations on a tool 2. That makes impossible its using as a detail of a product. Electrode 1 is replaced by a blank of detail 7 (where artificial roughness 8 must be obtained) after generation indentations 6 (Fig. 1b) on a cooling detail of a product. Generator 3 is switched to straight polarity (blank of detail 7—anode), the mode is changed to finishing electrical discharge, and then, roughness 8 is formed as smoothly related spherical projections and indentations. Result of this is an increase in cooling surface area. Using this on injectors, for example, helps lower the temperature in the cut of a nozzle to 100–150 K and significantly increase the service life of such details. In [3] (Fig. 2), new method of group electrochemical treatment of local cooling areas is offered. Also, there is offered a device for using this method. This device provides an opportunity accelerated by one order of magnitude and more making local cooling areas on surfaces of various shapes through simultaneous creation of indentations with random geometry. Figure 2 illustrates the opportunity of making annular and helical grooves, including internal surfaces of tubular heat exchangers with big length, where access of the cutting tool in the treatment area is difficult or impossible. Making of grooves 1 (Fig. 2) on surfaces of heat exchangers 2 is performed by electrochemical or combined treatment, where the cathode is electrode tool 3. It located on the internal surface of a heat exchanger through flexible elements 4. When grooves are treated electrolyte flows on two channels 5 and 6, which are separated by a membrane 7. This method uses a pumping effect created in channels by impulses of electrolyte

Fig. 2 Scheme of group making of grooves for increasing cooling area

1244

V. P. Smolensev et al.

pressure in a channel 6. It causes fluctuation of a membrane and an electrode tool. Fluctuation is limited by a flexible element. Electrolyte with a variable speed flow in a grooves treatment area through a system of openings in electrode tool and membrane. Electrolyte removes products of treatment and speeds up the process of anodic dissolution in grooves, and expands the technological possibility of increasing depth of grooves. When electrolyte pressure is permanent and has magnitude 0.2–0.23 MPa and impulse of pressure has magnitude 0.5–0.6 MPa. It was able to obtain grooves with a smooth contour up to 0.5 mm deep. These grooves are twice deeper than previously made and allow increase intensity of cooling up to 15%.

3 Cooling of Surfaces by the Method of Making Vortex Generators Cooling details of engines where liquid and gas–liquid flow through cooling area require making of vortex generators. These vortex generators should have a shape, which exclude locking of the cooling channel by gas phase, which is formed, for example, when liquid hydrogen is used. In [5–7], different shapes of vortex generators are presented. These vortex generators are made as local projections and indentations along with the movement of cooling flow. Shape of vortex generators have a significant influence on fluid dynamics of a cooling flow in the channel and determine the intensity of cooling and possibility of removing locking of the cooling channel. Figure 1 below illustrates high-speed footage of flow separation and turbulating of laminar flow on projections with rectangular and smooth shape [8] in a flow section of the channel. Visual comparison of two pictures shows that a rectangular projection (Fig. 3a) creates a vortex area which is higher than projection. Vortex structures also arise above projections and in angular areas, which significantly change fluid dynamics of a flow and can cause locking of the cooling channel, especially if it has a small size. Figure 3b shows that thickness of a vortex area behind smooth shaped projection is similar to its height. Because of a smooth shape of projection, «angular» vortexes are less advanced. That is why there is no locking of the cooling channel and heat exchange remains on a required level. In [2, 4, 9] are presented methods of electrochemical dimensional and combined treatments which created on the level of inventions. Using these methods allow to make vortex generators with a shape which is best achievable when electrical methods of treatment are used. Figure 4 below illustrates cooling channels with vortex generators with different sizes.

Technology of Combined Treatment of Engine Cooling Elements

1245

Fig. 3 Flow separation on vortex generators with shape of rectangular projection (a) and smooth indentation (b)

Fig. 4 Vortex generators, which are placed in bottom area and lateral surfaces of cooling channels. a—Location of cooling channels in a cooling jacket of combustion chamber of a missile engine, b—cross section of a channel, c—shape and sizes of a vortex generator, and d—cross section of a vortex generator

Vortex generators which provide a nonstop movement of coolant in channels have a height (h) 0.2–0.3 mm, length (L) about 1 mm, and width from 0.8 to 0.9 of the width of the channel (it depends on sizes of a channel, type of a coolant, and speed of a flow). Distance (H2) between vortex generators is defined by developers and depending on strength of a chamber under vibration effect. It also depends on terms of coolant flow and a time of a nonstop movement of coolant in a channel with vortex generators.

1246

V. P. Smolensev et al.

Fig. 5 Scheme of group making of vortex generators

In [2] is offered a new method and device, which are used for simultaneous making vortex generators with different shape and depth by a multifunctional template. Figure 5 illustrates the scheme of this. Location of indentations which are forming vortex generators on ribbed surfaces of a heat exchanger (for example, a combustion chamber or a reactive nozzle) is determined by a multifunctional template 1 (Fig. 5) which is made of a film dielectric material with cross-cutting openings with required shape and size. Openings are located on bottom 2 and lateral 3 parts of an external of a heat exchanger. Electrode tool 4 (cathode) moves a template into the groove and fixes it there for the time, needed for making an indentation with required depth on the bottom and lateral parts of a groove. Intensity of making vortex generators and their sizes depends on a gap between electrodes. That is why in the developed method and device system of a template tension is included. This system controls a size of a gap through changing the thickness of a film. Using of this invention allows made vortex generators with the shape of smooth elliptical indentations length 3 mm, gap 6 mm and depth 0.5 mm on a lateral part of a groove, and 0.3 mm on a bottom part of a groove. It was made on models of products in grooves with width 1.5 mm, depth 2 mm, and gap 3 mm, and lavsan film with thickness 0.12 mm was used for making a template. Tension of a film was changed from 580 to 620 N. Time of making all vortex generators was less than 3 min. It greatly decreases the labor input of technological processes. In heat exchangers after making these indentations were created smooth channels. These channels could not be locked by a coolant.

4 Conclusion Heat exchangers and vortex generators that were made by methods described in [1–3] were tested on models [4]. Workability of new heat exchangers was confirmed by computer simulation and calculation by final quantities method [10].

Technology of Combined Treatment of Engine Cooling Elements

1247

Comparative tests were held for channels without vortex generators, for channels with vortex generators with a rectangular shape on one of walls, for channels with vortex generators with smooth angular areas and top sides. Tests were held after electrical discharge, electrochemical, and combined production. Also, tests were held for vortex generators [2] with a drop shape (shape of a drop that was divided into two by its axis with a ratio of a diameter to its length equal d/l = 1/2.5) on one of the walls of product’s groove. Using new methods and devices allows achieving the largest area of a heat exchange without locking of channels. Researches that were described in this article open possibilities of increasing service life and operational reliability of heat exchangers create the basis for a solution to the problem of locking cooling channels by the gas phase when gas– liquid coolants are applied. It also allows to increase productivity and quality of manufacturing products, increase performance indicators of heat exchangers that are applied in aircraft, missile and automobile engine design, nuclear, food, and other knowledge-based branches of industry.

References 1. Smolencev VP, Koptev IT, Kuznecov IYu, Titov AV, Osekov AN (2012) Method of making local cooling area of a thermally stressed detail. RF patent 2464137, 20.10.2012 2. Smolencev VP, Sharov YuV, Smolencev EV, Mozgalin VL (2016) Electrochemical method of making indentations that are forming vortex generators on ribs and bottom part of cooling channels of thermally stressed machines and a device for its creation. RF patent 2573465, 20.01.2016 3. Koptev IT, Smolencev VP, Korovin AA, Kuznecov IYu, Osekov AN (2012) Method of an electrochemical treatment of local areas and a device for its using. RF patent 2470749, 27.12.2012 4. Smolencev VP, Koptev IT, Korovin AA, Klimova GN (2012) Simulator that could be used for optimal technological mode in process of making vortex generators in cooling channels by electrochemical method. RF patent 119663, 27.08.2012 5. Migaj VK (1989) Modelling of energy heat exchangers. Saint-Petersburg 6. Gortyshov YF, Olimpiev VV, Bajgaliev BB (2004) Heat hydraulic calculation and design of machinery with intensive heat exchange. Kazan 7. Vorobej VV, Loginov VE (2001) Technology of making of liquid propellant engines. Moscow 8. Gahun GG (1989) Construction and design of liquid propellant engines. Moscow 9. Smolencev VP, Korovin AA (2011) Making of vortex generators in cooling channels of liquid propellant engines. Basic Appl Probl Tech Technol 5(289):81–90 10. Alyamovskij AA (2010) Engineering calculations in solidworks simulation. Moscow

Temperature in Intermittent Cutting V. A. Solodkov, S. I. Kormilitsin and P. A. Norchenko

Abstract It is customary to assume that intermittent cutting temperature is much lower than steady-state temperature. However, according to the findings of experimental studies and the supporting theoretical calculations, the intermittent cutting temperature at the end of the cutting stroke (with a sufficiently long length) approximates to the temperature of steady-state cutting. This is due to a higher rate of plastic deformation in the chip formation zone and in the contact zone during intermittent cutting and, as a result, due to a higher intensity of heat output. It has been established that the cause of a higher plastic strain rate is a smaller amount of chip shrinkage and, as a consequence, a higher rate of its movement along the front face of the tool. The simplest, most accessible method for estimating temperatures, which provides for higher accuracy of calculations, is the calculation method by A. N. Reznikov. This technique is based on the most common thermo-physical method of solving the differential equation of thermal conductivity—the method of heat sources—in accordance with which the layout of sources and heat sinks during steady-state and intermittent cutting are adopted. To verify the results of the calculation, the cutting temperature was determined by a well-known and widely used method—the method of a natural thermocouple.





Keywords Intermittent cutting Cutting temperature Plastic deformation Strain rate Strain ratio





Contact straining


1 Introduction The theoretical and experimental studies of temperatures in intermittent cutting, which have been carried out so far, testify to their significantly lower values than those in steady-state cutting [1, 2]. This implies that during the main characteristics of the process of intermittent cutting feature the same values as those of steady-state V. A. Solodkov (&)
S. I. Kormilitsin
P. A. Norchenko Volgograd State Technical University, 28, Lenin Ave., Volgograd 400005, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_133

1249

1250

V. A. Solodkov et al.

cutting in the same machining conditions. However, according to the earlier studies [3], the characteristics of the intermittent cutting process are variable along the length of a single cut and significantly differ from the characteristics of steady-state cutting under the same machining conditions. In the studies comparing intermittent and steady-state cutting at the achieved level of thermo-EMF, the cutting process was not always carried out under the same conditions (e.g., in paper [1] with different thickness of the layer being cut off). When temperatures are measured by using different types of thermocouples, it is impossible to measure them on the contact surfaces (closer than 30 lm from the tool surface). Therefore, in this study, the temperature characteristics of intermittent and steady-state cutting processes were compared based on the results of measurements using the natural thermocouple method, which is the most reliable. The experimental data are confirmed by theoretical calculations performed by generally accepted methods.

2 Reasons for the Increased Intensity of Heat Generation in Intermittent Cutting The variability of the main characteristics of the intermittent cutting process influences the degree and rate of plastic straining in the contact zone and in the chip formation zone. As is known [4], these are the conditions and the nature of plastic straining that in the first instance determine the intensity of heat release and heat sink, and, consequently, the level of temperatures in the cutting zones. Due to the variability of the shear angle, shrinkage of chips, and cutting forces along the length of the cut, it is possible to assume that the characteristics of plastic straining are also variable and differ from those occurring in steady-state cutting. Figure 1 (steel 45-BK8; V = 60 m/min; S = 0.34 mm/tooth) shows the degree and speed of straining in the chip formation zone calculated on the basis of experimental data. By comparison, the same characteristics are given for steady-state cutting under similar conditions. The calculation was made according to standard formulas [5]. Since chip shrinkage is an approximate characteristic of the degree of straining for the material being machined, relative shear is taken as a characteristic of the degree of strain: e ¼ ctgb þ tgðb  cÞ;

ð1Þ

where b—shear angle; c—rake angle. The rate of strain in the zone of chip formation is calculated by using the following formula: e_ ¼

v
½cos b þ tgðb  cÞ
sin b ; Cc

where v—cutting rate; Cc—width of the zone of chip formation.

ð2Þ

Temperature in Intermittent Cutting

1251

Fig. 1 Degrees and rates of strain in the zone of chip formation

Average strain rates in the contact zone along the front surface are subject to the following dependence: e_ ¼

D
n ; v

ð3Þ

where D—thickness of the contact; n—chip shrinkage. The aforereferenced formula demonstrates that rate of contact deformations depends not only on the thickness of the contact zone, but also on the speed of movement of the chip along the front surface, defined by the ratio v/n. In other words, the smaller the shrinkage of the chips (the greater the shear angle), the higher the speed of the chip movement (at a given cutting speed) and, in general, the higher the speed of contact plastic deformations. In this case, it is essential to take into account the thickness of the zone in which the strains occur. Figure 2 shows the dependence of the rate of contact deformations along the length of the cut, calculated on the basis of experimental findings. It is seen that the strain rates with the maximum values achieved immediately after the penetration and the establishment of adhesive interaction fall along the length of the cut, approaching those values taking place at steady-state cutting. Figure 2 also shows the averaged value of the strain rate at steady-state cutting. According to the calculations, in order to achieve the same strain rate common to the rate during intermittent cutting, the cutting speed during steady-state cutting should be higher (in this case about 80 m/min). It can be assumed that the intensity of heat generation during intermittent cutting is higher than that during steady-state cutting under similar machining conditions [6]. Indeed, as noted in [7], “… the

1252

V. A. Solodkov et al.

Fig. 2 Rates of strain in the contact zone

work of friction is directly proportional to the cutting speed and inversely proportional to the shrinkage of the chips. Since shrinkage of the chips decreases simultaneously with an increase in cutting speed, both parameters influence in the same direction, causing an increase in heat generation.” In addition, an increase in the strain rate (as evidenced by a smaller shrinkage of the chips, and consequently, a higher speed of its movement) causes a significant increase in the yield strength of the material being machined, which also increases the power of the heat-evolution source [8].

3 Temperature in Intermittent Cutting The simplest, most accessible method for estimating temperatures, which provides higher accuracy for calculations, is the calculation method by A. N. Reznikov [2]. The intensity of heat generation in the chip formation zone is assumed to be uniformly distributed in the shear plane. Therefore, for a zero rake angle, it is calculated by using the following formula (all quantities are substituted into the following formulas in the units of measurement adopted in [2]): qn ¼

3:9
v
sin b
ðN
n  F Þ ; b
l
n

ð4Þ

where a, b—thickness and width of the layer being cut; N; F—normal and tangential components of the cutting force on the front surface;

Temperature in Intermittent Cutting

1253

On the front surface, the distribution of the heat-forming source according to the combined law is taken: qn ðx=lÞ ¼ qn ¼

5:85
v
ðN þ FÞ ; b
l
n

ð5Þ

if 0 < (x/l) < 0.5 qn ðx=lÞ ¼ qn
exp
½6
ðx=1  0:5;

ð6Þ

if 0.5 < (x/l) < 1.0; where l—length of full chip contact with the front surface; x—point coordinate. To calculate the temperature of steady-state cutting, the final formula obtained in [2] is used. With the introduction of the correction factor, this formula can also be used to determine the intermittent cutting temperature. The calculation program makes it possible to calculate the temperature characteristics of both intermittent and steady-state cutting under the same conditions in a parallel manner and determines the intensity of heat generation in the chip formation zone and on the front face, the cutting temperature, and the temperature in the chip formation zone. Experimental acquisition of the initial data was carried out with the same cutting conditions (Sz = 0.2 mm/tooth; t = 1.0 mm) and with the same tool geometry (c = 0°; a = 10°; u = 45°; r = 0), which are taken into account when calculating. When obtaining experimental initial data, special attention is paid to the creation of identical conditions for intermittent and steady-state cutting. All characteristics are obtained by cutting the same carbide plate, fixed in one holder without reinstalling. Since the characteristics of the intermittent cutting process are variable over the length of the cut, their average values over the cut length of 200 mm are taken for calculation. To accomplish this, the characteristics of the intermittent cutting process at various points from the beginning of the penetration are taken. The data obtained are the basis for the calculation of averaged values. Table 1 shows the initial data and the results of the calculation of the temperature characteristics of intermittent cutting (P) and gives the reference values for steady-state cutting (V). According to the given results, the temperatures of intermittent and steady-state cutting under the same conditions are rather close (the difference in all cases does not exceed 30–40 °C). In the zone of chip formation (at a significantly lower temperature level), the difference in temperatures is within one and the same order. The intensity values of heat generation in the chip formation zone are close, and since with a lower degree of straining, the strain rate during intermittent cutting is higher in this zone. Another important result is the fact of significantly higher intensity of heat generation on the front face under conditions of intermittent cutting. It is the increased intensity of heat generation that is apparently the main reason for the proximity of temperatures in steady-state cutting and in intermittent cutting (at the end of the cut). Another reason for the high-contact temperature under

U P U P U P

150

90

120

Type of cutting

Cutting rate (m/ min)

0.88 0.61 1.01 0.69 1.16 0.84

Length of contact (mm)

0.073 0.051 0.089 0.059 0.112 0.070

Height of contact zone (mm)

2.51 2.14 2.79 2.30 3.17 2.40

Chip shrinkage

57.3 52.6 61.0 55.1 66.5 58.6

41.8 35.4 43.8 37.9 47.7 42.1

Cutting force on the front face (kg) Tangential Horizontal

42,870 44,140 28,150 29,530 18,710 20,200

Intensity of heat release in the chip formation zone (cal/ sm2∙s)

Table 1 Comparative temperature characteristics of intermittent and steady-state cutting

269 257 245 242 217 210

Temperature in the chip formation zone (°C) 776 751 685 654 612 591

Cutting temperature (°C)

11,076 16,620 7411 11,112 4807 7665

Intensity of heat release on the front face (cal/ sm2∙s)

1254 V. A. Solodkov et al.

Temperature in Intermittent Cutting

1255

intermittent cutting conditions is obviously the worst conditions for heat flow— firstly, due to the shorter contact length, and secondly, due to the shift of the heat source to the cutting edge. In order to assess the dynamics of changes in the intensity of heat release along the length of the cut, on the basis of experimental data, the relative intensity of heat release is calculated in comparison with the intensity during steady-state cutting (Fig. 3). The figures show that throughout the entire length of the cut (with the exception of the penetration area), the intensity of heat generation during intermittent cutting (qint) is higher than that during steady-state cutting (qsteady). Moreover, at the beginning of the cut, when the strain rates are maximal (Fig. 2), the intensity of heat release can be more than twice as high.

4 Experimental Determination of Interrupted Cutting Temperature To verify the results of the calculation, the cutting temperature was determined by a well-known and widely used method—the method of a natural thermocouple. As it is commonly known [9], the method can be used to determine the average integral cutting temperature. The recommendations of paper [10] were also taken into account to conduct the research. The peculiarities of measuring thermo-EMF in intermittent cutting dealt with the closing of the formed chips on the workpiece being machined; this was caused by the phenomenon of intensive curling of chips in the vertical plane, which was typical of intermittent cutting. In order to prevent

Fig. 3 Ratio between the intensity of heat release during intermittent and steady-state cutting

1256

V. A. Solodkov et al.

distortion of the thermo-EMF signal due to this reason, the surface of the workpiece was covered with a dielectric plate. Experiments (measurement of thermo-EMF) during intermittent and steady-state cutting were carried out under strictly the same conditions using the same carbide plate. The thermo-EMF values were converted to temperature values using the refined calibration dependencies [11]. It should be primarily noted that there was an increase observed for the thermo-EMF and the temperature of intermittent cutting at a constant thickness of the cut along the cut length. This refuted the data [12–15] according to which the level of thermo-EMF established during cutting does not change with further cutting. By comparison, Fig. 4 shows the level of thermo-EMF during steady-state cutting and a typical graph of thermo-EMF during intermittent cutting (steel 45-T5К10; V = 150 m/min; S = 0.2 mm/tooth). The temperatures that correspond to these levels were also referred to in the course of the experiment. Intermittent cutting was performed with a single-tooth face-mill with a diameter of Ǿ160 mm according to a symmetrical milling scheme with a 90-mm-wide milling surface. An analysis of the results obtained allowed us to draw the following conclusions. First of all, it should be noted that both thermo-EMF and the temperature of interrupted cutting along the length of the cut showed a real increase, and the level of thermo-EMF during intermittent and steady-state cutting showed their proximity. In most cases, the level of thermo-EMF in steady-state cutting only slightly exceeded the thermo-EMF in intermittent cutting. Table 2 shows the results of cutting thermo-EMF measurement and the corresponding temperatures (for intermittent cutting, the thermo-EMF values and the corresponding temperatures at the beginning and at the end of the cutting pass are

Fig. 4 Relative temperatures during intermittent and steady-state cutting

Temperature in Intermittent Cutting

1257

Table 2 Comparing experimental and theoretical temperatures during intermittent and steady-state cutting Cutting rate (m/ min)

Thermo-EMF of cutting (мV)

Cutting temperature (°C)

Predicted temperature of cutting (°C)

Steady-state cutting

Steady-state cutting

Steady-state cutting

Intermittent cutting

Intermittent cutting Start of cut

End of cut

Intermittent cutting Start of cut

End of cut

150

13.6

12.4

13.4

811

724

783

781

761

120

13.0

11.6

12.8

724

631

695

697

663

90

12.2

11.0

12.0

652

539

617

633

589

given). The previously calculated temperatures are given as the reference values (Table 1). There is a good match between the data obtained experimentally and the calculated data. Cases of high thermo-EMF values during intermittent cutting are of greater interest than those during steady-state cutting [16, 17]. This phenomenon occurred in two cases—first, when a workpiece with a large pass length (up to 500 mm) was subjected to intermittent cutting. Analyzing the causes of this phenomenon and taking into account the fact of both a higher intensity of heat generation and a shorter contact length (worse than the flow condition) during intermittent cutting [3, 18–20], it can be assumed that at some point in time the temperature may exceed the steady-state value (Fig. 5). In the second case, the temperature can be exceeded at the very moment of tool penetration. However, in order to detect an increase in the steady-state value in

Fig. 5 Changes in temperatures during penetration with a zero (a) and full cut thickness (b)

1258

V. A. Solodkov et al.

terms of temperatures during penetration, the tool was to penetrate throughout the full thickness of the cutting layer. We could observe the following phenomenon—when a tool is inserted, the thermo-EMF level increases sharply and, at the first moment, exceeds the value then established (Fig. 5b). This phenomenon serves as an indirect confirmation of a higher intensity of heat release during intermittent cutting. By comparison, the same cutter was used under the same machining conditions to build a graph of thermo-EMF during a normal penetration with zero cut thickness (Fig. 5a). It can be seen that the gradually increasing thickness of the layer being cut obscures the burst of thermo-EMF at the initial moment, and its growth has a monotonic character.

References 1. Kasimov LN, Akbarov M, Rakhman-Zade AZ (1977) Researching the influence of some factors of intermittent turning of viscous steels on the cutting temperature. Optimization of cutting of extra strong and hot-strength alloys 2. Reznikov AN (1981) Thermophysics of materials machining processes. Mechanical Engineering, Moscow 3. Solodkov VA (2012) Increasing the efficiency of machining via cutting. Izdatelskiy dom “Spektr”. Moscow 4. Bobrov FB, Granovskiy GI, Zorev NN (1967) Developing the science of metal cutting. Mechanical Engineering, Moscow 5. Zorev NN (1956) Issues on the mechanics of metal cutting. Mechanical Engineering, Moscow 6. Solodkov VA (2013) Temperature features of intermittent cutting and the impact on them of solid lubricants. Bulletin of Volgograd State Technical University (Series of “Progressive technologies in mechanical engineering”), Volgograd 7. Victor H, Kamme H (1977) Standzeit und Verfanrensoptimierung beim Messer-kopffrasen. Zeitschrift fur industrielle Fertigung 8. Murarka P, Hinduja S, Barrow G (1981) Influence of strain, strain-sate and temperature on the flow stress in the primary deformation zone in metal cutting. Int J Mach Tool Des Res 9. Fokin OV (1964) Measuring the temperatures in the area of the contact of the cutter with both the chips and the product. Bulletin of engineering 10. Platonov AV (1989) Intensifying the turning section of pipes via preliminary plasma heating of the workpiece. Dissertation, Leningrad 11. Voevodin GA (1985) Patterns of the cutting process of steels with different microstructures and ways to reduce the intensity of wear of carbide tools. Dissertation, Saratov 12. Bardnikov LN (1976) Prevention of brittle fracture of the cutting tool due to thermal stress. Bulletin of engineering 13. Talantov NV (1992) The physical basis of the process of cutting, wear and fracture of the tool. Mechanical Engineering, Moscow 14. Podurayev VN (1974) Cutting of hard-to-machine materials. Higher School, Moscow 15. Zorev NN (1963) About the interaction processes in the zone of chip formation and in the contact zone of the front surface of the tool. Bulletin of engineering 16. Solodkov VA, Tibirkova MA (2010) The influence of the conditions of access to the performance carbide tool for interrupted cuts. Bulletin of Volgograd State Technical University (Series of “Progressive technologies in mechanical engineering”), Volgograd

Temperature in Intermittent Cutting

1259

17. Solodkov VA (2014) Shavingformation and contact interaction in the output. Bulletin of Volgograd State Technical University (Series of “Progressive technologies in mechanical engineering”), Volgograd 18. Chandrasekaran H, Nagarajan R (1977) Tranzient strains and chip formation during tool entry. Pros Int Cont Prod Eng 19. Pekelharing AJ (1978) The exit failure in interrupted cutting. CIRP Ann 20. Pekelharing AJ (1980) Cutting tool damage in interrupted cutting. Wear

Increasing Calorific Value of Biogas by Steam Explosion Activation of Renewable Raw Materials D. B. Prosvirnikov, A. R. Sadrtdinov and Z. G. Sattarova

Abstract Anaerobic microbial transformation of organic substrates, resulting in the production of various types of biofuels, currently occupies a leading position among the studies on the search and use of alternative energy sources. Cellulose-containing materials are among the most promising substrates, but for their full use there are a number of unresolved issues relating to the completeness of their utilization, improving the quality and volume of biogas production, as well as maintaining stability and increasing the functional activity of microbial communities. The article shows that the activation of plant materials by steam explosion can increase the caloric content and energy intensity of biogas produced from both straw and wood raw materials. Quantitative values of the biogas energy of combustion directly depend on the content of lignin and hemicelluloses in the feedstock, wherein the tendency to increase the calorific value of gas with an increase in steam explosion temperature is observed for all types of plant raw materials. The quantitative values of the biogas energy of combustion should directly depend on the content of lignin and hemicelluloses in the feedstock. Keywords Cellulose


Lignin
Biogas
Steam explosion
Activation

1 Introduction Reducing stocks of traditional fuels (oil, coal, natural gas, oil shale, and wood), methods of their extraction, transportation and use, leading to global ecosystem disturbances and steadily deteriorating ecology, identified current trends in the D. B. Prosvirnikov (&) Kazan State Power Engineering University, 51, Krasnoselskaya Street, Kazan 420066, Russia e-mail: [email protected] A. R. Sadrtdinov
Z. G. Sattarova Kazan National Research Technological University, 68, K. Marx Street, Kazan 420015, Russia © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_134

1261

1262

D. B. Prosvirnikov et al.

development of biotechnological processes for the production and use of alternative renewable energy. At the same time, renewable energy currently accounts for about 14% of the world’s consumption [1–5]. In addition to natural resources such as sunlight, wind, rain, tides, and geo-thermal sources, bioenergy can play a significant role among alternative energy sources, which implies using fuel derived from microbiological transformation of a number of organic substrates of biological origin: specially grown plants, products and waste of forest use and wood processing, waste of agriculture and livestock, and various types of organic household and industrial wastes [6–8]. The advantage of technologies based on natural processes and mechanisms for the conversion of organic substances using enzymes or microbial cultures is that waste and by-products of processes can also serve as additional sources of raw materials, which allows you to create completely waste-free technologies [9, 10]. One of the most energy-efficient ways to produce biofuels is biogas production (biomethane) by anaerobic microbial processing of organic substrates of various origins. Biogas consists of mainly methane (55–75% CH4) and carbon dioxide (25– 45% CO2) with trace amounts of nitrogen, hydrogen, hydrogen sulfide, and oxygen [11–13]. In addition, depending on the raw material used, a small amount of ammonia, alkanes, aromatic and halogen aromatic hydrocarbons, and oxidized compounds, such as sulfur dioxide, are found among the products. Biogas belongs to environmentally friendly fuels, although its combustion with pollutants can be accompanied by toxic emissions. Biogas with a methane content of at least 60% is most often burned, and the heat generated is used to heat the premises and various technological purposes, and however, to obtain a clean fuel that is identical in composition to natural gas, it needs additional cleaning. When separating the original gas mixture at specialized gas treatment plants and membrane modules, the concentration of methane rises to 95% and higher, with the result that biomethane can be used to generate electricity, as well as in internal combustion engines [14–18]. Depending on the methane content, the calorific value of biogas is 4700– 6000 kcal/m3. Biogas has several advantages over other types of alternative fuels. Biomethane produces significantly less harmful emissions than gasoline or diesel. The energy contained in methane is about three times more than that of hydrogen fuel, and the conditions for the storage and transportation of methane are also more efficient. In addition, in the formation of biogas from biomass, there is no need for special cultivation of agricultural plants, as it is done when obtaining biodiesel and bioethanol [19–24]. Basically, biogas is obtained by decomposing organic waste, such as cattle manure, plant waste from the food and agricultural industries. Despite the fact that anaerobic wastewater treatment has long been widely used, the application of this technology on an industrial scale in the processing of solid organic waste without dumping it in landfills is less successful. In a number of countries, a significant amount of biogas is obtained by processing so-called energy plants grown exclusively for the purpose of further processing into biofuels.

Increasing Calorific Value of Biogas by Steam Explosion …

1263

2 Plant Material for Biogas Production Various types of organic substrates can be used to obtain biogas; however, the main raw material for its production at industrial enterprises remains manure of farm animals and bird droppings, as well as the organic part of household and industrial waste in the form of wastewater. Studies are being conducted to obtain biogas from difficult substrates such as peat and coal. It has been shown that it is much more profitable, both from an economic and an ecological point of view, to ferment not “clean” waste, but to supplement it with the foundations, for example, from “energy” plants, which include specially cultivated grass crops (sugarcane, corn, millet, sunflower, miscanthus, rape, and some others), as well as woody undergrowth. The plant wastes from wood processing, farming, and livestock can also significantly increase the yield of biogas, and they also do not require the cost of growing, harvesting, and processing them, as is in the case with “energy” plants. The main polysaccharide component of higher plants is cellulose, the content of which may be about 35–50% of the dry mass of plant fibers. The second most abundant plant polysaccharide is hemicellulose, which is 25–35% of lignocellulosic biomass and is represented mainly by xylanes substituted (mostly in hardwood), glucans, mannans, arabinans, or galactans (the content of glucomannans is elevated in softwood) [25–29]. Different physiological groups of microorganisms have the ability to anaerobically decompose cellulose. The most complete information about this is presented in the reviews. Since cellulose fibers are tightly bound to other polymers, such as hemicellulose and lignin, this makes cellulose-containing materials extremely resistant to destruction. The hydrolysis of cellulose by bacteria is usually carried out slowly, but the microbial consortium of the farding bag of ruminant is able to hydrolyze 60–65% of cellulose in 48 h [30–32]. The ability to decompose cellulose is found in aerobes, representatives of the genera Acidothermus, Bacillus, Caldibacillus, Cellulomonas, Cellvibrio, Cytophaga, Dyella, Erwinia, Microbacterium, Micromonospora, Pseudomonas, Pseudoxanthomonas, Sporocytophaga, Rhodothermus, Streptomyces, and Thermobifida. Among the anaerobic microorganisms, cellulolytics are found in the representatives of the genera Acetivibrio, Anaerocellum, Bacteroides, Butyrivibrio, Caldicellulosiruptor, Clostridium, Desulfurococcus, Enterococcus, Eubacterium, Fibrobacter, Halocella, Ruminococcus, Spirochaeta, and Thermotoga [33–35]. It is known that the energy contained in organic raw materials can be converted to 85% in methane (calculated on the basis of cellulose conversion), which is the main component of biogas [36]. The speed and completeness of biodegradation of cellulosic material are influenced by the presence of lignin. In various types of cellulose-containing substrates, its content may be (%): 18–35—wood, 30–40—nutshell, 10–30—grass, 0–15—various types of papers, 15—wheat straw, 0—leaves, 18–30—newspaper, and 2.7–5.7 —solid cattle manure [34, 35]. Despite the fact that there is a direct relationship between an increase in methane concentration in the culture medium with an increase in the content of cellulose and

1264

D. B. Prosvirnikov et al.

hemicellulose in the substrate, the presence of lignin can significantly inhibit the formation of biogas. The stage of hydrolysis of cellulose is generally a rate-limiting step in the anaerobic cleavage of substrates containing cellulose, which is mainly due to the presence of lignin. The multicomponent and heterogeneous lignocellulosic biomass, as well as the crystalline structure of cellulose, significantly complicate the process of enzymatic hydrolysis of this raw material [36]. In this connection, various and not always cheap methods of pretreatment of cellulose-containing raw materials are often used, which may include mechanical crushing and grinding, pyrolysis at temperatures above 300 °C, and the use of gamma and microwave frequencies. Hot steam treatment which is the high-temperature autohydrolysis and hydrothermolysis—often with the addition of extra agents in the form of inorganic acids, carbon dioxide vapors, and ammonia— is used in the woodworking industry [34]. Chemical methods for the pretreatment of cellulose-containing materials include ozonolysis, alkaline hydrolysis, and hydrolysis with concentrated and dilute acids, as well as oxidation with hydrogen oxide in a humid atmosphere and the so-called organosolv—a process in which, along with inorganic acids, methanol, ethanol, acetone, ethylene glycol, also tetrahydrofurfural, and some other organic solvents are used. The greatest practical interest for the pretreatment of plant materials for biodegradation is the process of steam explosion, which was used in this study.

3 Materials and Methods Branches of deciduous plants and straw were used as the raw material. The raw materials were subjected to steam explosion at the plant, the scheme of which is shown in Fig. 1. The installation consists of two reactors—the first for direct steam explosion treatment, the second is for studying the process of steaming plant materials. In this work, the first reactor was mainly used. The experimental setup allows to obtain both liquid products containing sugars, oligosaccharides, and organic acids, and gaseous products containing furfural. Solid products are lignocellulosic mass in which physical and chemical changes take place under the action of pressure and temperature: The hemicellulose content is reduced, the lignin is converted into a destructive form, and the amorphous part of the cellulose is partially hydrolyzed. Thus, this raw material is an activated lignocellulosic pulp, with a large specific surface for penetration of cellulolytic enzymes. Steam explosion processing of plant materials was carried out at temperatures of 160, 190, and 220 °C in an environment of saturated water steam for 5–7 min. The microstructure of the obtained samples was studied using a LEXT4000 confocal laser scanning 3D microscope—images of the surface of dispersed capillary-porous lignocellulosic material were obtained in the focusing plane with magnification up to 300 lm. After the explosive treatment, the raw materials were subjected to biological treatment using the Cellulomonas Bacteria at a temperature of up to 37 °C in laboratory methane fermentation apparatus for

Increasing Calorific Value of Biogas by Steam Explosion …

1265

Fig. 1 Experimental plant for steam explosion activation of plant biomass

45–50 days under anaerobic conditions. The methane content in the composition of the exhaust gases was determined by a gas analyzer GAMMA-100. Gas samples with maximum methane yields were quantitatively collected to determine the heat of combustion on a DSC823e differential scanning calorimeter.

4 Results and Discussion The morphological properties of capillary-porous materials activated by steam explosion treatment were evaluated from images of the surface of dispersed capillary-porous lignocellulosic material (Fig. 2). As can be seen from the images, with an increase in the temperature of the steam explosion treatment, an increase in the specific surface of the material occurs. This is primarily due to chemical transformations that occur during steam hydrolysis. Thus, the effect of temperature during steam hydrolysis contributes to the destruction of hemicelluloses to mono- and oligosaccharides, as well as the partial

1266

D. B. Prosvirnikov et al.

Fig. 2 Photomicrographs of dispersed capillary-porous lignocellulosic material (straw)

destruction of lignin to low-molecular-weight fragments. At the same time, cellulose, as a macromolecule, is practically not subjected to destruction. However, close-packed crystalline chains of cellulose alternating with amorphous regions undergo hydrolytic degradation in unstable non-crystalline regions. All these points in the cell wall contribute to an increase in the internal specific surface. In the macro-picture after steam explosion treatment, the separation of the plant material fibers from each other is observed due to the pressure drop resulting from the depressurization of the reactor. Since the moisture in the material is in a superheated state, with a decrease in pressure above it, it boils sharply with a large amount of energy released, which exerts a tensile force on the fibers, thereby separating them from each other. Figure 3 shows the results of a study of the content of methane in biogas, formed as a result of biotreatment of plant materials activated at different temperatures. The results showed that the lowest yield of methane in the composition of biogas is observed for raw materials that are not subjected to steam explosion, the highest yield is observed for raw materials activated at 190 °C, and the biodegradation of the material treated at 220 °C does not lead to a significant decrease in methane output after 50 h of processing. The low methane yield is due to the fact that cellulolytic enzymes with great difficulty enter the cell wall of the material, respectively; the low rate of biological conversion of cellulose-containing raw materials leads to a lower methane yield. The inaccessibility is due to the high content of the original lignin and hemicelluloses, which physically prevent the penetration of the enzyme. The increased

Increasing Calorific Value of Biogas by Steam Explosion …

1267

Fig. 3 Methane yield in biogas in %, depending on the temperature of the steam explosion activation of plant raw materials

processing temperature, and hence the higher specific surface area obtained, on the contrary, promotes the penetration of enzymes deep into the cell wall, but, as studies have shown, up to a certain processing temperature. Thus, the maximum yield of methane in biogas was in raw materials activated at a temperature of 190 ° C, and with an increase in the temperature of the steam explosion treatment up to 220 °C, it leads to a decrease in the yield of methane. This is explained by the fact that at temperatures above 190 °C, low-molecular-weight fractions of lignin undergo additional changes in the form of plastification and condensation of fragments with the formation of more stable and insoluble forms, which, similar to the original lignin, prevent the penetration of enzymes into the cell wall. Figure 4 shows the results of determining the heat of combustion of biogas containing methane, obtained from the activated material at different temperatures of steam explosion. Thus, the maximum energy intensity is observed in biogas obtained from vegetable raw materials activated at a temperature of 190 °C, which is obviously related to the content of methane in this gas. Fig. 4 Calorific value of biogas depending on the temperature of the steam explosion activation of plant materials

1268

D. B. Prosvirnikov et al.

These results are similar for both straw and wood. In the case of wood, the calorific value of gas is 10–15% lower, since the initial lignin content in it is 30–40% higher than in straw, and also as a result of the different morphological structure of the cell wall.

5 Conclusions Thus, it is shown that the activation of plant materials by steam explosion can increase the calorific value and energy intensity of biogas produced from both straw and wood raw materials. Obviously, the tendency to increase the calorific value of gas with an increase in steam explosion temperature will be observed for all types of plant raw materials, but the quantitative values of the biogas energy of combustion should directly depend on the content of lignin and hemicelluloses in the feedstock, which impedes the penetration of enzymes. Acknowledgements The results presented in the work were obtained with the support of the Scholarship of the President of the Russian Federation (SP-4422.2018.1).

References 1. Mahapatra S, Kumar S, Dasappa S (2016) Gasification of wood particles in a co-current packed bed: experiments and model analysis. Fuel Process Technol 145:76–89 2. Tuntsev DV et al (2016) The mathematical model of fast pyrolysis of wood waste. In: Proceedings of 2015 international conference on mechanical engineering, automation and control systems, MEACS 2015. Article no 7414929 3. Dasappa S, Paul PJ, Mukunda HS, Shrinivasa U (1994) The gasification of wood-char spheres in CO2 N2 mixtures: analysis and experiments. Chem Eng Sci 49(2):223–232 4. Bui T, Loof R, Bhattacharya SC (1994) Multi-stage reactor for thermal gasification of wood. Energy 19(4):397–404 5. Prosvirnikov DB et al (2017) Modelling heat and mass transfer processes in capillary-porous materials at their grinding by pressure release. In: Proceedings of 2017 international conference on industrial engineering, applications and manufacturing, ICIEAM 2017. Article no 8076443. https://doi.org/10.1109/icieam.2017.8076443 6. Gusev VG, Fomin AA, Saldaev VA (2018) Mathematical model of cut layer at intensive profile milling of workpieces. In: International conference on industrial engineering, p 517– 526 7. Fomin AA (2017) Limiting product surface and its use in profile milling design operations. Solid State Phenom 265:672–678. https://doi.org/10.4028/www.scientific.net/SSP.265.672 8. Prosvirnikov DB et al (2017) IOP Conf Ser Mater Sci Eng 221:012010. https://doi.org/10. 1088/1755-1315/221/1/012010 9. Gusev VG, Fomin AA (2017) Multidimensional model of surface waviness treated by shaping cutter. Procedia Eng 206:286–292 10. Timerbaev NF et al (2017) Software systems application for shafts strength analysis in mechanical engineering. Procedia Eng 206:1376–1381. https://doi.org/10.1016/j.proeng. 2017.10.648

Increasing Calorific Value of Biogas by Steam Explosion …

1269

11. Thijs Defraeye, Blocken Bert, Carmeliet Jan (2013) Influence of uncertainty in heat–moisture transport properties on convective drying of porous materials by numerical modelling. Chem Eng Res Des 91(1):36–42 12. Bouzid Majda et al (2011) In-pore tensile stress by drying-induced capillary bridges inside porous materials. J Colloid Interface Sci 355(2):494–502 13. Fomin AA et al (2018) Geometrical errors of surfaces milled with convex and concave profile tools. Solid State Phenom 284:281–288. https://doi.org/10.4028/www.scientific.net/SSP.284. 281 14. Stepanov VV et al (2017) Composite material for railroad tie. Solid State Phenom 265:587– 591 15. Sadrtdinov AR et al (2016) IOP Conf Ser Mater Sci Eng 124:012092. https://doi.org/10.1088/ 1757-899x/124/1/012092 16. Lashkov VA et al (2016) IOP Conf Ser Mater Sci Eng 124:012111. https://doi.org/10.1088/ 1757-899x/124/1/012111 17. Anisimova IV, Gortyshov YF, Ignat’ev VN (2016) Russ Aeronaut 59:414. https://doi.org/10. 3103/s1068799816030193 18. Timerbaev NF et al (2017) Application of software solutions for modeling and analysis of parameters of belt drive in engineering. IOP Conf Ser Earth Environ Sci 87(8):082047. https:// doi.org/10.1088/1755-1315/87/8/082047 19. Popov IA et al (2015) Cooling systems for electronic devices based on the ribbed heat pipe. Russ Aeronaut (Iz VUZ) 58(3):309–314 20. Safin R, Barcik S, Tuntsev D, Safin R, Hismatov R (2016) A mathematical model of thermal decomposition of wood in conditions of fluidized bed. Acta Facultatis Xylologiae Zvolen res Publica Slovaca 58(2):141–148 21. Timerbaev NF, Ziatdinova DF, Safin RG, Sadrtdinov AR (2017) Gas purification system modeling in fatty acids removing from soapstock. In: Proceedings of 2017 international conference on industrial engineering, applications and manufacturing, ICIEAM 2017. Article no 8076418. https://doi.org/10.1109/icieam.2017.8076418 22. Hu G, Li G, Zheng Y, Zhang Z, Xu Y (2015) Euler-Lagrange modeling of wood chip gasification in a small-scale gasifier. J Energy Inst 88(3):314–322 23. Tuntsev DV et al (2018) Physical and chemical properties of activated lignocellulose and its areas of application. Solid State Phenom 284:779–784. https://doi.org/10.4028/www. scientific.net/SSP.284.779 24. Saravanakumar A, Haridasan TM, Reed TB (2010) Flaming pyrolysis model of the fixed bed cross draft long-stick wood gasifier. Fuel Process Technol 91(6):669–675. https://doi.org/10. 1016/j.fuproc.2010.01.016 25. Fomin AA, Gusev VG, Sadrtdinov AR (2018) Assurance of accuracy of longitudinal section of profile surfaces milled at high feeds. In: International conference on industrial engineering, p 527–536 26. Janajreh I, Al Shrah M (2013) Numerical and experimental investigation of downdraft gasification of wood chips. Energy Convers Manag 65:783–792 27. Tuntsev DV, Filippova FM, Khismatov RG, Timerbaev NF (2014) Pyrolyzates: products of plant biomass fast pyrolysis. Russ J Appl Chem 87(9):1367–1370 28. Pakdel H, Roy C (1991) Hydrocarbon content of liquid products and tar from pyrolysis and gasification of wood. Energy Fuels 5(3):427–436 29. Storodubtseva TN, Aksomitny AA, Sadrtdinov AR (2018) Thermal insulation properties of wood polymeric sand composite. Solid State Phenom 284:986–992. https://doi.org/10.4028/ www.scientific.net/SSP.284.986 30. Palmer ER (1984) Gasification of wood for methanol production energy in agriculture 3:363–375 31. Mazarov IY, Timerbaev NF, Sadrtdinov AR (2018) Cogeneration power plant for processing biomass with the application of solid oxide fuel cells. In: International multi-conference on industrial engineering and modern technologies (FarEastCon), Vladivostok, Russia, p 1–4. https://doi.org/10.1109/fareastcon.2018.8602699

1270

D. B. Prosvirnikov et al.

32. Prosvirnikov DB et al (2017) Modeling of delignification process of activated wood and equipment for its implementation. IOP Conf Ser Mater Sci Eng 221(1):012009. https://doi. org/10.1088/1755-1315/221/1/012009 33. Popov IA, Shchelchkov AV, Gortyshov YF et al (2017) High Temp 55(4):524. https://doi. org/10.1134/s0018151x17030208 34. Sychevskii VA (2015) Drying of colloidal capillary-porous materials. J Heat Mass Transf 85:740–749 35. Drapalyuk MV et al (2016) IOP Conf Ser Mater Sci Eng 142:012090. https://doi.org/10.1088/ 1757-899x/142/1/012090 36. Prosvirnikov DB, Safin RG, Zakirov SR (2018) Microcrystalline cellulose based on cellulose containing raw material modified by steam explosion treatment. Solid State Phenom 284:773– 778 37. Storodubtseva TN, Aksomitny AA, Saldaev VA (2018) The study of soundproofing properties of wood polymer-sand composite. Solid State Phenom 284:993–998. https://doi. org/10.4028/www.scientific.net/SSP.284.99

Application of Statistical Modeling Methods to Assess Decontamination Effect of Electromagnetic Field on Raw Materials for Food Industry N. N. Ovchinnikova, G. S. Kochetkova and T. A. Tolmacheva

Abstract The article considers a modern method to decontaminate dried fruits from the molds of storage by the electromagnetic field of microwave radiation. The research is based on the active planning of the technological experiment. The experiments were carried out three times. In order to show a dependence on the contamination of dried fruits on time and heating rate, we construct regression equations. Also, we represent by diagram an influence of the electromagnetic field of microwave radiation on various fungus of dried fruits. Effective modes of decontamination were found. This research was conducted within the framework of the “Industry 4.0” concept with the use of modern technologies and equipment. The results of the research can be introduced into the food production in order to improve the efficiency and achieve a new product quality.







Keywords Regression equations Decontamination effect Microwave method of decontamination Two-factor experiment by the Kono plan




1 Introduction The breakthrough in technology provides a start of new industrial revolution, or “Industry 4.0”, all over the world. In the near future, large-scale changes in technology will lead to the situation, when economic growth is stimulated by not capital and natural resources, but innovation and creative thinking. In contrast to the previous industrial revolutions, the industrial revolution 4.0 develops not linearly, but exponentially [1]. “Industry 4.0” is permeated with the spirit of innovation and is focused on individual solutions, and also has a great potential for automation. On the one hand, N. N. Ovchinnikova (&)
G. S. Kochetkova South Ural State University, 76, Lenin Avenue, Chelyabinsk 454080, Russia e-mail: [email protected] T. A. Tolmacheva FSAEI HE “RSAU-MTAA”, 49, Timiryazevskaya Str., Moscow 127550, Russia © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_135

1271

1272

N. N. Ovchinnikova et al.

it is necessary to create solutions to automate production, which has the functionality corresponding to “Industry 4.0”. On the other hand, new products and technologies are necessary [2, 3]. In this paper, we consider a modern method to decontaminate dried fruits from mold fungi of storage by the electromagnetic field of microwave radiation. In modern bakery and confectionery production, the use of high-quality environmentally friendly raw materials, modern technologies, and equipment is especially topical. These conditions ensure maximum preservation of vitamins and nutrients, as well as organoleptic advantages of products, and an increase in shelf life [4, 5]. High quality of confectionery products is achieved by used raw materials having a variety of chemical composition and properties: flour, sugar, starch syrup, fruits and berries, cocoa beans, oil-containing kernels of nuts [6]. Analysis of the quality of the used raw materials shows that in order to regulate the properties and qualitative indices of food products made from vegetable raw materials, and it is necessary to apply scientifically grounded methods to impact, as well as to and select additional raw materials [7]. In the production of confectionery products, both flour and sugar, dried fruits are used as an additional raw material. Dried fruits contain a lot of useful minerals (iron, magnesium, calcium, and phosphorus), organic acids, and pectins. The basic vitamins of fresh fruits disappear at the time of drying, although vitamins A, C, PP, and some vitamins of group B remain. Dried fruits help to remove radionuclides, heavy metals and many other unhealthy substances from the body Preservation of fruits and products of their processing is reduced to regulation of life processes, which form the basis of the spoilage. The biological processes taking place in the raw materials, as well as the vital activity of the microorganisms, lead to the spoilage of fruits. In order to suppress the vital activity of spoilage agents and preserve raw materials, we can change environmental conditions and act on raw materials or on microorganisms by physical, biological, and chemical factors.

2 Review of Existing Methods for Processing Raw Materials in the Food Industry There are various methods of preservation for processing raw materials having vegetable origin. These methods are subdivided into physical, chemical, physicochemical, biological, and biochemical [8]. Physical methods are the most common and include heat processing (pasteurization, sterilization, cooling, and freezing); sterilization by ultrasound; high-frequency current processing; ultraviolet, red, and blue spectra of laser radiation processing; ionizing radiation processing; mechanical sterilization [9].

Application of Statistical Modeling Methods to Assess …

1273

Chemical methods are based on the addition of preservatives, antiseptics, and antibiotics to products of fruit and berry processing. Physicochemical methods include drying process of fruits and vegetables, as well as preservation by sugar [10, 11]. Biochemical methods are based on the use of natural preservatives produced by microorganisms or accumulated in plant cells (antibiotics, phytoncides). Biological methods to preserve fruit and berry products are mainly based on the use of the biosynthetic potential of microorganisms. Chemical methods include processing by sulfurous anhydride, sugar, alcohol, food surfactants, as well as sulfur fumigation. Sulfur dioxide (SO2), sodium benzoate (C6H5COONa), and sorbic acid (CH3(CH)4COOH) are permissible to use in order to preserve fruits and berries, fruit purees, fruit and berry juices. Sodium benzoate meets all the requirements for antiseptics. However, sodium benzoate cannot be removed from the product before food intake. In recent years, canning production successfully uses sorbic acid or its potassium salt, which are harmless to humans and have a preservative effect in small concentrations (0.05–0.1%). Sugar processing allows to produce comestibles, which are semi-finished products made from fresh fruits or berries with a pronounced aroma. Such fruits or berries are preserved by sugar in two ways: hot and cold. Alcohol processing is as follows. Put fresh, sorted fruits and berries in alcohol– sugar solution, and then bottle obtained product using an airtight corking. The aim of processing by food surfactants is to accelerate the drying process of stone fruits and grapes. For stone fruits, good results are obtained by processing fruit surface with 6–7% oleic acid emulsion for 2–3 min at normal temperature. In this case, the possibility that the fruits crack or lose coloring substances is excluded; because, there is no heating. Despite that the chemical methods to preserve fruit and berries are relatively widespread, the physical methods are preferable. Indeed, in the food preserved by physical methods, there are no foreign substances, which can impact on human health. We, especially, caution against to use chemicals, which are not permitted by public health authorities, such as salicylic acid and its derivatives (e.g., aspirin). The simplest, cheapest, and least laborious method to preserve fresh fruits and berries is drying. At the same time, drying is a complex thermophysical and technological process. During the drying process at elevated temperatures, a complexity of chemical composition of raw materials and low stability of its components cause enough deep physicochemical, structural, and biochemical changes. The combination of these changes usually leads to a change in the initial organoleptic properties and nutritional value of the product. The nature and depth of these changes depend on the composition and initial properties of the raw material, on the processing raw materials before drying, on the methods and mode of drying, and on the amount of moisture removed from the products.

1274

N. N. Ovchinnikova et al.

In order to maximize the preservation of valuable natural properties of raw materials and to obtain a product, which is steady under storage in different conditions, the process and mode of drying should be justified. In order to extend the shelf life of products, it is necessary to reduce the contamination of the products with microbes. To this end, there are many methods based on physical, biological, and chemical impacts [12–14]. The microwave method of decontamination is based on the influence of two fields: electromagnetic and thermal. This method is characterized by selective heating, i.e., the more humid components of dried apricots are heated faster [15]. Under the action of the microwave field, water molecules (dipoles) perform vibrational and rotational motions, orienting with the field frequency along electric lines of the field. The motion of molecules is the thermal energy. The more water in a given volume and the more molecules participate in this movement, the more thermal energy liberates. Therefore, whole volume of dried fruits is warmed, and the most humid areas receive more energy. Hence, moisture is removed from the dried fruits, and at the same time, the moisture level is equalized throughout whole volume. At present, the use of microwave field energy is one of the special variants for heat processing of various food products [16, 17]. Microwave heating has several advantages over traditional methods of heat processing: • high speed of heating and uniform heating due to “volumetric” heat supply; • preservation of vitamins and other essential nutrients of the food product; • possibility of a soft heat processing mode, to supply heat by pulses, i.e., step heating; • in order to create a given temperature unevenness during heat processing of food products, it is enough to choose a shape of working elements of the microwave generator or to use screens that control the transmission of microwaves to the product; • high economical efficiency of the process (indeed, there is no contact with the heat carrier, and the heat is generated in the product itself, therefore, heat loss for heating of equipment and external environment is minimized; in addition, the consumption of electricity by microwave generators is much less than the consumption by electrolytes and other heating devices); • improvement of work conditions by reducing the liberation of gaseous substances, steam, and heat into the environment [18].

Application of Statistical Modeling Methods to Assess …

1275

3 Decontaminating Effect of the Microwave Electromagnetic Field on Dried Fruits and Evaluation of the Effect by Statistical Methods Laboratory and laboratory-production experiments were carried out during the research of the microwave energy impact on the microflora of dried fruits. The impact of the microwave field energy on the causative agents of mold fungi having genus Penicillium, Mucor, and the total microbial number was studied in laboratory conditions. An application of the method to decontaminate by the microwave field energy in the technological process was studied in laboratory-production conditions. The research was based on the method of active planning of the technological experiment. The method allows to obtain a regression equation that connects modes of impact on dried fruits and indicators of decontamination efficiency and quality improvement [19, 20]. Dried apricots having extended storage were studied during the research of reaction of the saprophyte microflora to the microwave field energy impact. There were 11 variants in a scheme of the experiment, including two control ones: a control, i.e., to wash dried fruit with sterile water, and a standard control, i.e., a standard according to the existing technology (to boil fruits in sugar syrup). The experiments were carried out three times. The permissible number of mold fungi in dried fruits is 100 CFU/g (Sanitary regulations and codes 2.3.2.1078-2001). An initial level of seeding of mold fungi having genus Penicillium in dried apricots was determined by the control and standard control. According to the results of the control, a contamination with these mold fungi was 2170 CFU/g while the standard control showed that the contamination was 10 CFU/g. The obtained results were tabulated (Table 1). Analysis of the results shows that the effect of complete decontamination of dried apricot (variant №1, №5, №6, №7) takes place in the following conditions: an increase in heating temperature is 85–100 °C, heating time is 60–90 s, and heating rate ranges from minimum to maximum. Note that organoleptic and taste qualities in these variants either deteriorate or correspond to control level. In the case of maximum decontamination modes (heating time is 90 s, heating rate is 0.8 °C/s, heating temperature is 100 °C), there is a complete destruction of the fungus contamination, but the dried apricots are caramelized and the taste of burned berries appears. In the case of minimum decontamination modes (heating time is 30 s, heating rate is 0.4 °C/s, heating temperature is 40 °C), the contamination halves. Organoleptic and taste qualities correspond to the control level. If the heating temperature is 65 °C, then the contamination halves. At the same time, if the heating temperature is 70 °C, then the contamination decreases 20 times. In both these cases, a taste becomes sour-sweet, and a color corresponds to the control level. A further increase in heating temperature up to 75–80 °C leads to

1276

N. N. Ovchinnikova et al.

Table 1 Impact of microwave energy on the number of mold fungi of genus Penicillium on dried apricots N

Modes

Contamination CFU/g; 1*10−2 I II III Average

Time (s)

Rate, v (°C/s)

Temperature, t (°C)

1

90

0.8

100

0

0

0

0

2

30

0.8

80

1

0

0

0.3

3

90

0.4

70

2

2

2

2

4

30

0.4

40

11

11

12

11

5

60

0.8

90

0

0

0

0

6

60

0.4

70

0

0

0

0

7

90

0.6

85

0

0

0

0

8

30

0.6

65

12

13

11

12

9

60

0.6

75

1

0

1

10

Control (to wash dried fruit with sterile water)

18

22

25

11

Standard control (to boil fruits in sugar syrup)

0

0

0.3

0.7

22

0.1

Organoleptic and taste qualities

Darkened, caramelized, taste of burned berries Burned, caramelized, tart taste Colorless, sour-sweet taste Taste and color correspond to the control level Darkened, caramelized, honey taste Color is lighter than control color, taste of fresh berries Color corresponds to the control level, sour-sweet taste of ripe berries Color corresponds to the control level, sour-sweet taste of ripe berries Color corresponds to the control level, sweet honey taste Yellow-brown color, sweet taste of dried apricots Yellow-brown color, very sweet taste of boiled dried apricots

a decrease in the number of mold fungi of genus Penicillium up to standards of maximum permissible concentration. If heating time is 60 s, heating rate is 0.6 °C/s, and heating temperature is 75 °C, then the contamination is 67 CFU/g. If heating time is 30 s, heating rate is 0.8 °C/s, and heating temperature is 80 °C, then the contamination is 33 CFU/g. Note that if heating temperature is 80 °C, then the

Application of Statistical Modeling Methods to Assess …

1277

organoleptic and taste qualities deteriorate (burned fruits have a tart taste), and if heating temperature is 75 °C, then the color corresponds to the control level, and the taste improves and becomes sweet and honey. It follows from the table that not all considered variants for processing by microwave energy are acceptable to decontaminate dried apricots. Indeed, in some of the variants, the decontamination is accompanied by the deterioration of organoleptic and taste qualities. Effective modes for processing of dried apricots by microwave field method are variants, where heating time is 60 s, heating rate is 0.6 °C/s, and heating temperature is 75 °C. Decontamination of dried apricots from the mold fungi of genus Penicillium was carried out by the plan of the two-factor experiment according to the Kono plan. Therefore, analysis of the obtained data allowed to make a regression equation in order to show the dependence of contamination of dried apricots on heating time and heating rate. The regression equation in the coded variables has the form: Y1 ¼ B0 þ B1 x1 þ B2 x2 þ B12 x1 x2 þ B11 x21 þ B22 x22 ; where Y1 is a contamination of dried apricots with mold fungi of genus Penicillium, x1 is a heating time, x2 is a heating rate. In our case, B0 ¼ 1:5; B1 ¼ 1:6; B2 ¼ 4:2; B12 ¼ 2:2; B11 ¼ 1:9; B22 ¼ 4:0. Therefore, we obtain the following regression equation: Y1 ¼ 1:5  1:6x1  4:2x2  1:9x21 þ 4:0x22 þ 2:2x1 x2

ð1Þ

We use regression Eq. (1) to construct the response surface (Fig. 1) and the graphical dependence (Fig. 2), which shows the impact of microwave field energy on the mold fungi of genus Penicillium on dried apricots. The obtained response surface (Fig. 1) and the graphical dependence (Fig. 2) show that partial decontamination of dried apricots from mold fungi of genus Penicillium takes place starting with minimum parameter values of the processing by microwave field, i.e., heating time is 30 s, and heating rate is 0.4 °C/s. If heating time is 30–90 s, heating rate is 0.6–0.7 °C/s, and heating temperature is 70–75 °C, then the dried apricots are decontaminated to the level of maximum permissible concentration (100 CFU/g) while the control level of contamination is 2170 CFU/g. If the parameters are increased, i.e., heating time is 30–60 s and heating rate is 0.7–0.8 °C/s, then there is a slight increase in mold fungi contamination. Complete decontamination takes place when heating time is 75–90 s and heating rate is 0.56–0.62 °C/s. Analysis of tabular data on the decontamination of dried apricots from mold fungi of genus Penicillium and the graphical dependence show that the most effective modes to decontaminate dried apricots from these molds have the following parameter values: heating time is 60 s, heating rate is 0.6 °C/s, and heating temperature is 75 °C. These parameter values ensure decontamination

1278

N. N. Ovchinnikova et al.

Fig. 1 Impact of the microwave field energy on mold fungi of genus Penicillium on dried apricots

Fig. 2 Impact of microwave field energy on mold fungi of genus Penicillium on dried apricots

Application of Statistical Modeling Methods to Assess …

1279

of dried apricots, decrease in contamination up to 67 CFU/g, and improvement of taste qualities. Namely, dried apricots acquire sweet honey taste and remain brightly colored. Contamination with mold fungi of genus Mucor on dried apricots in the control and standard control was 1770 and 70 CFU/g, respectively. Complete decontamination of dried apricots from mold fungi of genus Mucor by the microwave field energy takes place when heating time is 60–90 s, heating rate is 0.6–0.8 °C/s, and heating temperature is 75–100 °C, see Table 2. If an impact mode of the microwave field is maximum (variant №1), i.e., heating time is 90 s, heating rate is 0.8 °C/s, and heating temperature is 100 °C, then there is the complete decontamination of dried apricots from the mold fungi spores, while dried apricots become darkened and caramelized, and acquire taste of burned berries. Note that there is a stimulation of growth and development of mold fungi in the variants with minimal parameter values, i.e., heating time is 30 s, heating rate is 0.4–0.6 °C/s, and heating temperature is 40–60 °C. In this case, contamination of dried apricots is 1830 CFU/g, i.e., higher than in the control level. Organoleptic indicators of the quality of dried apricots remain at the control level. Complete decontamination of dried apricots from the mold fungi spores takes place in variant №7, where heating time is maximum, i.e., 90 s, heating rate is mean, i.e., 0.6 °C/s, and heating temperature is 85 °C. In this case, dried apricots acquire sour-sweet taste while color of fruits corresponds to the control level. Consider modes having the following parameter values: heating time is 60 s, heating rate is 0.4 °C/s, and heating temperature is 70 °C, see, variant №6. There is the decontamination of dried apricots from the mold fungi of genus Mucor by the microwave energy up to 300 CFU/g, i.e., the number of colonies decreases in comparison with the control level on average six times. In the case of modes, where heating rate is maximum, i.e., 0.8 °C/s, while heating time is 30–60 s, and heating temperature is 80–90 °C, there is the decontamination from the mold fungi of genus Mucor up to 33 CFU/g. Complete decontamination of dried apricots from the mold fungi of genus Mucor takes place in variant with mean parameter values: heating time is 60 s, heating rate is 0.6 °C/s, and heating temperature is 75 °C. In this case, organoleptic and taste qualities of dried apricots are improved in comparison with the control level and are as good as the standard control level (Table 2). Tabular results on the decontamination of dried apricots from the mold fungi of genus Mucor show that the most acceptable modes having the following parameter values: heating time is 60 s, heating rate is 0.6 °C/s, and heating temperature is 75 °C. Decontamination of dried apricots from the mold fungi of genus Mucor was carried out by the plan of the two-factor experiment according to the Kono plan. Therefore, analysis of the obtained data allowed to make a regression equation in order to show the dependence of contamination of dried apricots on heating time and heating rate.

1280

N. N. Ovchinnikova et al.

Table 2 Impact of microwave energy on the number of mold fungi of genus Mucor on dried apricots N

Modes Time (s)

Contamination CFU/g; 1*10−2 Time Rate, Time (s) v (°C/s) (s)

Rate, v (°C/s)

1

90

0.8

100

0

0

0

0

2

30

0.8

80

0

1

0

0.33

3

90

0.4

70

2

0

0

0.7

4

30

0.4

40

20

18

17

18.3

5

60

0.8

90

0

0

0

0

6

60

0.4

70

3

3

3

3

7

90

0.6

85

0

0

0

0

8

30

0.6

65

19

18

18

18.3

9

60

0.6

75

0

0

0

0

10

Control (to wash dried fruit with sterile water)

18

16

19

17.7

11

Standard control (to boil fruits in sugar syrup)

0.7

0.7

0.7

0.7

Organoleptic and taste qualities Darkened, caramelized, taste of burned berries Burned, caramelized, tart taste Colorless, sour-sweet taste Taste and color correspond to the control level Darkened, caramelized, honey taste Color is lighter than control color, taste of fresh berries Color corresponds to the control level, sour-sweet taste of ripe berries Color corresponds to the control level, sour-sweet taste of ripe berries Glossy, color corresponds to the control level, sweet honey taste Yellow-brown color, sweet taste of dried apricots Glossy, yellow-brown color, very sweet taste

The regression equation in the coded variables has the form: Y2 ¼ B0 þ B1 x1 þ B2 x2 þ B12 x1 x2 þ B11 x21 þ B22 x22 ; where Y2 is a contamination of dried apricots with mold fungi of genus Mucor, x1 is a heating time, x2 is a heating rate. In our case, B0 ¼ 2:6; B1 ¼ 3:5; B2 ¼ 6:2; B12 ¼ 4:3; B11 ¼ 2:4; B22 ¼ 5:3. Therefore, we obtain the following regression equation:

Application of Statistical Modeling Methods to Assess …

1281

Fig. 3 Impact of the microwave field energy on mold fungi of genus Mucor on dried apricots

Y2 ¼ 2:6  3:5x1  6:2x2  2:4x21 þ 5:3x22 þ 4:3x1 x2

ð2Þ

We use regression Eq. (2) to construct the response surface (Fig. 3) and the graphical dependence (Fig. 4), which shows the impact of microwave field energy on the mold fungi of genus Mucor on dried apricots. Analysis of the obtained response surface and the graphical dependence shows that partial decontamination of dried apricots from mold fungi of genus Mucor takes place starting with minimum parameter of the processing by microwave field. If heating time is maximum, then there is intensive decontamination. Complete decontamination of dried apricots from the mold fungi of genus Mucor takes place, when heating time is 90 s and heating rate is 0.8 °C/s. If heating time is 30–75 s, then either there is a minimal permissible concentration, or complete decontamination of dried apricots from the mold fungi of genus Mucor. More effective modes have the following parameter values: heating time is 60 s, heating rate is 0.6 °C/s, and heating temperature is 80 °C. Results of the experiments given in the graph (Fig. 3) show that the increase in heating time and heating rate leads to the gradual decrease in the contamination of dried apricots with the mold fungi of genus Mucor.

1282

N. N. Ovchinnikova et al.

Fig. 4 Impact of microwave field energy on mold fungi of genus Mucor on dried apricots

4 Conclusion Analysis of the published methods used in the technological process for processing raw materials shows that modern chemical and biological methods to destruct microorganisms do not provide sufficient biological efficiency. In addition, most of the reagents used for decontamination create environmental problems, since processed dried fruits are consumed in food. The method of microwave processing is the most acceptable. The method is effective both to protect against a complex of pathogenic microorganisms having different etiology and to preserve the organoleptic and taste properties of dried fruits. As a result of the research, we establish parameters of the decontamination of dried fruits by microwave field energy, i.e., heating time and heating rate. Method of active experiment planning (Kono plan 2n−1) allows to estimate the quantitative characteristics of decontamination of dried fruits from unhealthy microflora by microwave using minimal volume of experimental study. The results of the research show that the complete decontamination of dried fruits from the mold fungi of genera Mucor and Penicillium takes place, if the modes have the following parameter values: heating time is 60 s and heating rate is 0.6 °C/s. Therefore, the studied data on the use of microwave energy show the effectiveness of this physical method for the processing of dried fruits. Mechanisms of the impact of microwave energy on food need further study with the help of technologies that will become possible in the age of “Industry 4.0”.

Application of Statistical Modeling Methods to Assess …

1283

References 1. Schwab K (2018) Fourth industrial revolution. Publishing House E, Moscow, p 252 2. Khan MK, Ahmad K, Hassan S, Ahmad N, Xu (2018) Effect of novel technologies on polyphenols during food processing. Innov Food Sci Emerg Technol 45:361–381 3. Orlov VV, Alferov AS (2006) Prospects of microwave processing of liquid food products. In: Processes and devices of food production, vol 2, no 2. Research Institute of ITMO, pp 52–54 4. Vorobiev VV (2005) Scientific and practical basis to create efficient technologies for the production of high-quality products from hydrobionts using the electromagnetic field of microwave. Dissertation, Moscow, p 398 5. Ushakova NF, Kopysova TS, Kasatkin VV, Kudryashova AG (2013) Experience of microwave heating application for food production. Food Ind 10:30–32 6. Arslanov Sh (2010) Impact of electrophysical influences on the technological process of baking of bread. Bread Prod 11:56–57 7. Yusupova GG, Yu I, Zdanovich, Cherkasova EI (2005) Application of microwave energy to ensure safety and improve the quality of products of vegetable origin. In: Storage and processing of agricultural raw materials, vol 7, pp 27–29 8. Skokan LE, Zharikova GG (2006) Microbiology of main types of raw materials and semi-finished products in the confectionery production. Moscow, p 148 9. Koh DA, Tilsina NN (2009) Impact of the freezing process on fruit of small-fruited apples. In: Modern problems of equipment and technology of food production: materials XII international. Scientific-practical conference, Barnaul, pp 150–154 10. Altukhov IV, Ochirov VD (2009) Analysis of methods for drying of food products. Bull Irkutsk State Agric Acad 36:16–21 11. Antipov ST, Shirshov EA, Kazartsev DA (2002) Impact of electromagnetic field intensity on the process of dielectric drying of coriander seeds. In: Storage and processing of agricultural raw materials, vol 9, pp 50–51 12. Rogov IA (1988) Electrophysical methods for processing of food products. Moscow, p 272 13. Zuglionok VN (2004) Substantiation of the technological process and effective modes of microwave decontamination of grain in the production of grain bread. Dissertation for the degree of Candidate of Technical Sciences, Krasnoyask, p 122 14. Xu F, Chen Z, Huang M, Li C, Zhou W (2017) Effect of intermittent microwave drying on biophysical characteristics of rice. J Food Process Eng 40(6):e1259 15. Yusupova GG, Kosovan AP, Yusupov RK (2010) Microbiological control of food products from grain. Moscow, p 422 16. Chizoba Ekezie FG, Sun DW, Han Z, Cheng JH (2017) Microwave-assisted food processing technologies for improving product quality and process efficiency: a review of recent developments. Trends Food Sci Technol 67:58–69 17. Guo Q, Sun DW, Cheng JH, Han Z (2017) Microwave processing techniques and its recent applications in the food industry. Trends Food Sci Technol 67:236–247 18. Rushchyts AA, Shcherbakova EI (2014) Analysis of methods for drying of food products. Bull South Ural State Univ Ser Food Biotechnol 2(1):9–15 19. Pen RV (1982) Statistical methods to simulate and optimize processes of pulp and paper production. Krasnoyarsk, p 168 20. Kretov IT, Shashkin AI, Shakhov SV (2003) Modeling of the vacuum-sublimation drying of food products by the microwave field. Proc High Educ Inst Food Technol 5–6:65–68

Energy Consumption Modeling of Machining Processes V. Salnikov and Yu. Frantsuzova

Abstract Energy consumption management is an effective tool for adapting the production system under the conditions of changing the operating mode. Rational energy consumption improves not only the energy performance of the enterprise but also the ecological and general economic parameters. Energy consumption assessment of each technological operation and for each workplace requires all production facilities (equipment, materials, operations, etc.) to be clearly identified. This study represents machining operations as the transformation of workpiece’s surface into the desired surface of the part. This is important in order to evaluate its effectiveness in terms of energy conversion. This approach serves as the basis for mathematical modeling of machining operations’ energy consumption, which considers the type, parameters, and time of exposure. This information is obtained from the control program text at the preparatory stage by means of syntactic and lexical analysis. This chapter presents a planned schedule of the machine tool’s energy consumption during the processing of a given surface. A processing experiment, performed on real equipment, helped to obtain a graph of actual energy consumption. Finally, the paper discusses efficiency strategies that can be used at some hierarchical levels of an industrial enterprise.




Keywords Machining Energy efficiency production Mathematical modeling





Energy consumption
Industrial

1 Introduction Production is a complex system, which depends on many factors [1]. Even the most well-organized systems cannot be effective all the time. Therefore, in order to get the maximum efficiency level, the structure of the system should be flexible and dynamic and should be provided with some regulatory mechanisms [2]. V. Salnikov (&)
Yu. Frantsuzova Tula State University, 92, Lenin Ave., Tula 300012, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_136

1285

1286

V. Salnikov and Yu. Frantsuzova

Energy consumption management is an effective tool for adapting the production system under the conditions of changing the operating mode. Rational energy consumption improves not only the energy performance of the enterprise but also the ecological and general economic parameters [3–7]. Enterprise managers often underestimate the amount of savings they can get due to energy efficiency programs. They believe that there is a certain degree of technical and financial risk connected with their implementation. These programs have a lower priority than traditional commercial offers. In addition, despite the relatively low energy resources’ costs, it is hard to convince the management of the need for complex energy efficiency projects. It should be added that energy prices are constantly changing, depending on a large number of factors (economic situation in the country, currency exchange rate, market position of the energy supplier, etc.), and as a result, their share in the product costs is also changing. These arguments emphasize the urgency of improving the energy efficiency of industrial enterprises.

2 Mathematical Modeling Energy consumption assessment of each technological operation and for each workplace requires all production facilities (equipment, materials, operations, etc.) to be clearly identified [8]. As a rule, machine-building production systems have a complex multi-level hierarchical structure [1, 9]. The lower level is determined by the impact on the processing object. The upper level is determined by the structural unit for which energy consumption is estimated. Each level is characterized by a certain energy loss [10]. Still there are no generally accepted methods of energy consumption regulation, unlike the regulation of other resources [1]. This situation is explained by the lack of mutual understanding between energy services, acting as an energy supplier, and technological services acting as the consumers, in terms of the energy consumption efficiency [1, 11]. Even the most detailed regulation method, based on the calculation of energy consumption per unit of process equipment, does not imply considering the actual state of production capacity and the types of operations performed [12, 13]. These parameters serve as the basis for operational management and long-term cost planning, including energy consumption, according to technical criteria: energy intensity, energy consumption, etc. [14–18]. It is known that the main task of any method of machine parts’ processing is the formation of surfaces of the required functional purpose with given operational properties [19]. The machining operation, in particular, is inextricably linked to the destruction, that is, to the formation of new surfaces on the part and the cutting layer of the workpiece [20].

Energy Consumption Modeling of Machining Processes

1287

Previously, a mathematical model of industrial equipment’s energy consumption was proposed, basing on the representation of the energy characteristics of the technological impact as a function of the surfaces formed. According to this model, machining operations consist of the transformation of the workpiece’s surface into the desired surface of the part. This enables the evaluation of their energy conversion effectiveness [2, 21, 22]. Nðb; V; SÞ ¼ AbS þ BbV þ ESV;

ð1Þ

where S [mm] is the feed rate due to the operation of the ith axis drive; V [mm/s]— cutting velocity; b—stock allowance; bSi [mm2]—section area formed during the processing of chips; bV [mm2/s]—chips’ lateral surface formation rate; SV [mm2/s] —directly formed surface; and A [J/s mm2], B [J/mm2], E [J/mm2]—energy coefficients of the energy consumption model. A, B, and E coefficients are obtained by approximating the known experimental dependence [23] of the cutting power Nex ¼ C
bx
Sy
V n
ðr=75Þh ;

ð2Þ

where C, x, y, n, h are table coefficients and measures of empirical dependence, respectively; r is the ultimate tensile strength of the material being processed. When evaluating operations’ energy consumption, it is important to consider not only the type and parameters of the impact but also its temporal characteristics. Therefore, energy consumption functions for working and idle strokes can be represented as follows: Wws ðDt; b; S; VÞ ¼ DtNðb; S; VÞ ¼ ðtEw  tSw ÞNðb; S; VÞ;

ð3Þ

Wis ðDt; b; S; VÞ ¼ DtNð0; S; VÞ ¼ ðtEi  tSi ÞNð0; S; VÞ;

ð4Þ

where Dt—transition time between successive reference points of the tool path; tSw , tEw are, respectively, the starting and the ending time of the working stroke; and tSi , tEi —are, respectively, the starting and the ending time of the idle stroke. During machining, the total energy consumption for processing a complex surface consists of working and idle strokes’ consumption in accordance with the formula: Wðt; b; S; VÞ ¼

l X i¼1

Wws i ðDti ; bi ; Si ; Vi Þ þ

m X

Wis j ðDtj ; 0; Sj ; Vj Þ;

ð5Þ

j¼1

where Wws i , Wis j —is the energy consumption at the ith working stroke and the jth idle one, respectively; l, m—the number of working and idle strokes of the operating tool, respectively.

1288

V. Salnikov and Yu. Frantsuzova

Another component of energy consumption of a particular technology is determined by the availability of preparatory final and additional time for each part operation, which can be calculated from the following dependency DWadd ¼

lNgiven v ; q

ð6Þ

where l—is the given capacity coefficient (Ngiven) of a particular machine drive; m = Tadd/T—the amount of additional time in the actual time of the machine operation; q—the total number of part operations performed on a particular machine. Thus, the total energy consumption for a particular part operation on a particular machine, and, consequently, the energy intensity of the operation can be defined as Wproc ¼

p Z X

Wðt; b; Si ; VÞdt;

ð7Þ

i¼1

where p is the number of machine’s operating tools involved in the part operation. The proposed mathematical model for estimating the operation energy consumption considers the type and parameters of operations, as well as their duration. Temporal characteristics can be obtained from the text of the control program (CP) at the preproduction stage. The text of the CP contains the tool path, processing modes, and other supporting information. Tool path R is a set of reference points p1, p2 … pn. Its graphic representation is shown in Fig. 1, where 1 is a cutting tool; 2—ith point of the path; 3—motion path; and 4—formed surface. Each point pi is characterized by coordinates X, Y, and Z. This enables calculation of coordinate increments along the axes oX, oY, and oZ for each elementary part of the tool path and helps to determine the tool drives involved in this motion.

Fig. 1 Tool path

Energy Consumption Modeling of Machining Processes

1289

DXi ¼ p½Xi þ 1  p½Xi ;

ð8Þ

DYi ¼ p½Yi þ 1  p½Yi ;

ð9Þ

DZi ¼ p½Zi þ 1  p½Zi ;

ð10Þ

where DXi , DYi , DZi —the coordinate increments of the motion along the axes oX, oY, oZ on ith path area, respectively (oX, oY, and oZ axes drives are involved). Let us add the following parameters to each point: cutting velocity V, feed rate S. The value of the latter can be obtained by comparing the coordinates of the tool position and the geometry of the workpiece being machined. A graphic image of the formation of surfaces during milling is presented in Fig. 2, where 1 is a cutting tool; 2—ith point of the path; 3—tool path; 4—formed surface; and 5—workpiece. The necessary information for energy consumption modeling can be found in the control program text. At the preproduction stage, it can be used to calculate and optimize the temporal and energy characteristics at each point of the tool path. The absolute value of horizontal motion on the ith linear area of the tool path can be determined by the following formula: Di ¼

qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ðDXi Þ2 þ ðDYi Þ2 :

ð11Þ

The direction of tool motion can be determined using already known values Di , DXi , or DYi as follows: a ¼ arccosðDXi =Di Þ

or

a ¼ arcsinðDYi =Di Þ

ð12Þ

To estimate the drives’ operating time along coordinate axes oX and oY on ith area, we use the following formula:

Fig. 2 Graphic representation of surface formation during milling

TiX ¼ DXi =ðSi
cosðaÞÞ;

ð13Þ

TiY ¼ DYi =ðSi
sinðaÞÞ;

ð14Þ

1290

V. Salnikov and Yu. Frantsuzova

The tool path is obtained at the preproduction stage using the syntactic and lexical analysis of the control program text. Basing on the information about the path and current geometry of the workpiece, we calculate the stock allowance at each point of the path and determine the tool drives involved in the movement and the duration of their work. This data is sufficient for estimating the energy consumption in the path section. The total energy consumption for the machine control program execution will be equal to the sum of the energy consumption on separate sections of the path. Figure 3 shows the algorithm described. The proposed approach offers an opportunity to improve the accuracy of energy planning for machining operations in industrial enterprises. It takes into account the type, parameters, and temporal characteristics of the operations. At the production stage, energy consumption is monitored, and the technological process is controlled

Fig. 3 Algorithm for estimating the energy consumption at the preproduction stage

start 6

1 3D CP workpiece model of i-th part’s processing

Calculating stock allowance at the jth point of the path 2

7 Calculating energy consumption of j-th point of the path

3D model-based workpiece’s geometry obtaining 3

8

4

9

Obtaining tool path

Obtaining cutting velocity and feed rate at each point of the path

Energy consumption for ith part production

5 end

Energy Consumption Modeling of Machining Processes

1291

to minimize energy loss. The deviation of the actual energy consumption from the planned graph indicates that there was a disruption in the system’s normal production process. This deviation may be caused by a violation of the technology, i.e., by the inconsistency of technological parameters with specified values, or abnormal situations in equipment and power systems, i.e., failure of power transmission networks (energy loss during transportation from supplier to consumer); changes in energy consumption and specific return from the use of production capacity (wear and tear and, as a consequence, an increase in energy consumption), etc. [9, 22].

3 Simulation Modeling Figure 4 shows a fragment of the tool path (green curve). The ratio of the diameter of the tool and the inner and outer diameters of the machined surface allows us to determine the contact angle of the tool and the workpiece, and, consequently, the uneven distribution of the cutting forces. When machining the inside diameter, or the pocket, the contact angle is greater, and the cutting force is greater as well. When machining the protrusion, the contact angle is smaller, as well as the cutting force. Since the parameters of the cutting mode along the path remain unchanged, the calculated value of power consumption repeats the graph of the cutting force change (see blue curve in Fig. 4). This curve reflects the ideal law of energy consumption by a metal cutting machine when processing a given part (see a green curve in Fig. 4).

4 The Performance of the Experiment The studies were carried out using a part with die-rolled section surface. A typical element of this section is shown in Fig. 4 (green curve). Ten workpieces were processed on the machine HAAS UMC 750SS. The control program was created using the Autodesk Powermill CAM system. Figure 5 shows the curve of the actual energy consumption, constructed according to the data produced by the CNC system in the diagnostic mode

Fig. 4 Result of energy consumption assessment according to the control program text

1292

V. Salnikov and Yu. Frantsuzova

Fig. 5 Actual and planned energy consumption during the processing of a given surface

(red curve). A comparison of the planned (blue curve obtained by means of the developed module) and the actual energy consumption shows that the actual curve has a constant component of 4.5 kW, which is 10% larger than the calculated curve. This can be explained, for example, by the influence of the idle stroke power of the machine drives involved in the forming process, or the influence of friction in the tracks, etc. The flat top of the curve when processing protrusions indicates a decrease in power consumption because of almost zero value of the contour tool feed. In summary, we can conclude that the control program designed using the Powermill CAM system is not optimal in terms of energy consumption processing conditions (change in cutting depth) along the shape-forming path. Consequently, there is an urgent task to stabilize the cutting mode along the entire tool path in order to increase the energy consumption efficiency. In addition, it was shown that in order to solve the problem, all the necessary data (workpiece form, parts, tool parameters, tool path, etc.) should be known before the processing starts. This enables the optimization (stabilization) of operation parameters at the preproduction stage. The analysis of the tool path obtained using the developed energy analysis module and of the workpiece’s geometry showed the “irrational” idle tool strokes (Fig. 6). It can be concluded that they significantly increase the energy consumption of this control program since a large amount of energy is spent on maintaining the spindle head and the work of the tool drives. Therefore, the identification and optimization of these “irrational” idle tool strokes are essential for the overall improvement of energy consumption efficiency.

Energy Consumption Modeling of Machining Processes

1293

Fig. 6 Irrational idle tool strokes

5 Results and Discussion The identified reasons for the deviations of the actual energy consumption from the calculated values allow us to give a number of recommendations aimed to improve the planning and efficiency of energy consumption. For example, it is useful to include the dependence of power consumption on efficiency into the design model, and during the processing of a cavity, it is advisable to reduce the contour feed. Thus, summarizing all the above, we would like to highlight the following ways to improve energy efficiency on some hierarchical levels of machine-building industrial enterprises, which are the aim of this project: Operation level: – selection of optimal operation parameters b, V, S (taking into account the actual state of the equipment) in terms of energy efficiency (minimum required amount of energy at the operation area); Technological process level: – selection of the optimal path for processing of the workpiece’s surfaces (elimination of “irrational idle strokes,” reducing the proportion of additional time in preparatory final and additional time periods in the total machine operating time); – selection of the optimal path for part processing, taking into account a large number of possible sizes of equipment and the level of wear.

1294

V. Salnikov and Yu. Frantsuzova

At the highest hierarchical level, one of the main barriers to improving energy efficiency and, consequently, increasing the competitiveness of products is managers’ conservatism. Industrial enterprises often underestimate the potential savings from energy-saving programs.

References 1. Salnikov VV, Ivutin AN (2017) The role of production information for the service of the main energy production system. Izv Tula State Univ Tech Sci 8(1):165–170 2. Salnikov VS (2003) Technological basis for energy efficiency of production systems. Tula 3. Gludkin OP, Gorbunov NM, Gurov AI (1999) Total quality management. In: Gludkina (ed.) Moscow 4. North K (2001) Quality and efficiency management: module program. In: Prokopenko I (ed.) Moscow 5. Venikov VA (1990) Energy saving technology. Power supply of the national economy. Moscow 6. Smith CB (1981) Energy management principles. Elmsford, New York 7. Thuman A (1984) Handbook for energy audit. Atlanta, Georgia 8. Erzin OA, Salnikov VV (2014) One aspect efficiency evaluation of technological systems. Izv Tula State Univ Tech Sci 11(2):594–603 9. Salnikov VV, Ivutin AN, Frantsuzova YV (2018) Computer support use of energy resources in the production system. Izv Tula State Univ Tech Sci 6:106–114 10. Salnikov V, Ivutin AN (2016) Increasing energy efficiency in production systems. In: Collection of theses of the scientific-practical conference “innovative high-tech information technologies”. Materials, Tula 11. Salnikov VV, Ivutin AN (2017) One of the aspects of collecting technological information for assessing the efficiency of production energy consumption. In: Collection of theses of the congress of young scientists, Electronic edn. ITMO University, Saint-Petersburg 12. Kireeva NV (2012) The methodology of the rationing of the power consumption at the industrial enterprises: problems and development directions. “Kazan Science”, pp 30–33 13. Salnikov VV (2016) Methodology of technological operations’ energy consumption rationing. In: Intellectual and information systems. Materials, Tula 14. Semenova EI, Vlchkevicha LI (1987) Automation of discrete production. Moscow 15. Alifanov AY (1993) The basic principles of electrophysical, electromechanical and combined processing methods. Russ Eng Res 41–45 16. Pukhovsky ES (1989) Technological bases of flexible automated production. Kiev 17. Rastrigin LA (1980) Modern principles of complex objects management. Moscow 18. Tishchenko N (1986) Introduction to the design of control systems. Moscow 19. Bobrov VF (1975) Fundamentals of metal cutting theory. Moscow 20. Poduraev VN (1985) Technology of physico-chemical methods of machining. Moscow 21. Salnikov VV (2017) Mathematical model of industrial equipment’s energy consumption In: Intellectual and information systems: materials of the international scientific and technical conference, Tula 22. Salnikov VV (2018) Computer based support for efficient use of energy in manufacturing. In: 7th Mediterranean conference on embedded computing (MECO). https://doi.org/10.1109/ meco.2018.8405994 23. Malov AN (1973) Handbook of metallurgy, vol 4 (ed: Malov AN)

Design and Calculation Method of Composite Housings for New Generation Magnetorheological Devices K. V. Naigert and V. A. Tselischev

Abstract By designing of special purpose drive systems, the construction materials must possess the functional properties, for example, transmission of magnetic and electromagnetic radiation or magnetic shielding being important properties. The control of characteristics in the magnetorheological (MR) systems occurs due to exposure to external electromagnetic fields on the working environment. It is obvious that in the MR drive systems, structural elements for installing of electromagnetic control units must have transmission properties of electromagnetic waves and metal is not applicable to these structural elements. Using of polymer composites is a promising solution to the problem. Particularly, high strength properties have fiber polymer composites, which are able to withstand mechanical stress that exceeds allowable stresses for preserving integrity of steel elements. The fiber polymer composites can withstand the high tensile and compressive stresses, but it is only by stresses, which are attached in axial direction of composite fiber, and the ability to resist loading in the radial or tangential direction significantly concedes to values of allowable axial stresses. MR devices by exploitation under multi-directional dynamic loads need the use of composite materials with more isotropic material properties (strength properties). In the research, this problem was solved and the authors made a new fiber polymer composite material, which has significantly less anisotropy in strength properties. The calculation method of strength properties of this new fiber polymer composite material is proposed. Keywords Fiber polymer composites Strength properties


Magnetorheological devices


K. V. Naigert (&) South Ural State University, 76, Lenin Avenue, Chelyabinsk 454080, Russia e-mail: [email protected] V. A. Tselischev Ufa State Aviation Technical University, 12, K. Marx Street, Ufa 450008, Russia © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_137

1295

1296

K. V. Naigert and V. A. Tselischev

1 Introduction The problem in the implementation of technical projects often is the absence of materials with desirable physical properties. This is especially widespread in industries, which are associated with rocket, aircraft and precision engineering; they have specific requirements for the strength properties of materials, durability and ability to resistance to different types of strains and stresses. Therefore, the housing elements of aircraft and spacecraft structures are usually manufactured with composite materials. The working principle of MR devices and the installation of electromagnetic control units on the housings need manufacturing of housings from magnetic conductive materials and do not allow application of metal material for housings, because the metal causes the magnetic shielding effect; it is associated with magnetic flux closure on the metal surfaces. The difficulties by processing of polymer composite materials after the consolidation of the polyester compounds are obvious, and technological additional processing of fiber polymer composites leads to weakening of composite structures and to appearance of internal stresses by machining processes [1–11].

2 Actuality of the Research The use of new materials and technologies in the manufacture of elements of MR devices can increase their effectiveness and variants of design that allows expanding application field of MR devices. Consequently, the development of integration concept of composite materials in the MR systems and perfection of design of MR devices with composite elements is relevant.

3 Statement of the Problem The existed composite materials do not meet all the requirements of industry, and fiber composite materials have anisotropic structures, which need special calculation methods that are adapted to the features of composite structure. The original composite structure requires theoretical basis for calculation of their physical characteristics. So, it is important to realize the improvement of composite structures of MR device housings and development of calculation methods for their design.

Design and Calculation Method of Composite Housings for New …

1297

4 Constructive Solution to the Problem and Numerical Model of Multilayer Composite Housing The design of composite housing is published in patent application (the decision to grant a patent RU 2018130967) [12]. Composite housing for MR damper has nonmetallic shell and composite liner with aramid frame, which are manufactured from whole aramid fiber. The aramid frame is reinforced by whole carbon fiber. Along the circumference of cylindrical aramid frame, the circular array of carbon fiber reinforcing turns is formed; the carbon turns are interwoven with aramid fiber and between each other. The use of nonmetallic shells and composite liners allows significantly to increase the strength characteristics of housings and simplify the process of their production. Whole fiber aramid frame can be produced in the form of cylinder or sleeve and reinforced to increase strength with whole carbon fiber. Aramid and carbon fibers have high tensile strength, but it is just correct under condition of preserving fiber integrity and application of loads in the axial direction, along the fiber. The specific strength of carbon fiber is inferior to specific strength of aramid fiber. Therefore, it is rational to manufacture the frames of composite liner from aramid fiber. The implementation of composite liners with fiber structure of woven aramid fibers and reinforcement with whole carbon fibers contribute to increasing the resistance of housing to multi-directional dynamic loads. Structure of new polymer composite is formed with whole fibers that increase the strength characteristics of finished products, and it allows them to withstand significant multi-directional dynamic loads. The option of physical properties’ modeling of composite structure is considered. The MR damper housing has shell of rotation geometry, axial symmetry and orthotropic fiber composite layer. The characteristics’ calculation of multilayer structure of MR damper housing will be realized on the base of availability of three-layer structure—fiber composite layer (composite liner), which is bound by compound and placed between two polymer shells. Carbon fibers of composite material have pronounced orientation in the structure. We describe the elastic characteristics of every layer of composite housing. Composite layer is formed by carbon and aramid fibers. Taking into account that the reinforcing fibers and stresses in the housing have different directions, the fibers of composite layer of selected fragment are angled to the direction of loading. Therefore, we introduce two coordinate systems: orthogonal coordinate system with axes x, y, z, which coincide with the direction of loading in the composite layer and coordinate system with axes 1, 2, 3. The axis 1 is positioned in direction of reinforcing fiber of composite layer and is angled to the axis x by angle u. Stresses and strains in the fiber composite layer are written as [13–15]:

1298

K. V. Naigert and V. A. Tselischev

rx ¼ r1 cos2 ðuÞ þ r2 sin2 ðuÞ  s12 sinð2uÞ ry ¼ r1 sin2 ðuÞ þ r2 cos2 ðuÞ  s12 sinð2uÞ sxy ¼ ðr1  r2 Þ sinðuÞ cosðuÞ þ s12 cosð2uÞ sxz ¼ s13 cosðuÞ  s23 sinðuÞ syz ¼ s23 cosðuÞ  s13 sinðuÞ e1 ¼ ex cos2 ðuÞ þ ey sin2 ðuÞ þ cxy sinðuÞ cosðuÞ

ð1Þ

e2 ¼ ex sin2 ðuÞ þ ey cos2 ðuÞ  cxy sinðuÞ cosðuÞ
 c12 ¼ ey  ex sinð2uÞ þ cxy cosð2uÞ c13 ¼ cxz cosðuÞ þ cyz sinðuÞ c23 ¼ cyz cosðuÞ  cxz sinðuÞ Aramid structure has a complex woven structure, which is formed by whole fiber, and it allows considering of orthotropic structure in several directions as symmetrical. In the case of symmetric reinforcement, the stresses and strains will take the form:
   r 1 ¼ E 1 e1 þ l21 e2
   r 2 ¼ E 2 e2 þ l12 e1  s 12 ¼ G12 e12  2  2  r x ¼ r1 cos ðuÞ þ r2 sin ðuÞ  s12 sinð2uÞ  2  2  r y ¼ r1 sin ðuÞ þ r2 cos ðuÞ  s12 sinð2uÞ
    s xy ¼  r1  r2 sinðuÞ cosðuÞ þ s12 cosð2uÞ

ð2Þ

2 2 e 1 ¼ ex cos ðuÞ þ ey sin ðuÞ  sinðuÞ cosðuÞ 2 2 e 2 ¼ ex sin ðuÞ þ ey cos ðuÞ  sinðuÞ cosðuÞ
 c 12 ¼  ey  ex sinð2uÞ þ exy cosð2uÞ

Matrix of compliance for composite material and the relation between stress and strain in the matrix form are [16]: l21 =E2 1=E1 l12 =E1 1=E2 l13 =E1 l23 =E2 ½D ¼ 0 0 0 0 0 0 ¼ ½D
frg þ fe0 g

l31 =E3 l32 =E3 1=E3 0 0 0

0 0 0 1=G23 0 0

0 0 0 0 1=G13 0

0 0 0 ; fe g 0 0 1=G12

ð3Þ

Design and Calculation Method of Composite Housings for New …

1299

l12 ; l13 ; l21 ; l23 ; l31 ; l32 —lateral strain coefficient, E1 ; E2 ; E3 —elastic modulus, G23 ; G12 ; G13 —shear modulus. Elastic characteristics in direction along the reinforcing fiber of composite material are established experimentally. Determining of elastic characteristics of fiber carbon and aramid composites is realized for layers. The elastic modulus of fragment of fiber composite layer can be calculated by: El ¼ Efc Vfc þ Efa Vfa þ Ec Vc

ð4Þ

Efc —elastic modulus of carbon fiber, Vfc —volume fraction of carbon fiber, Efa — elastic modulus of aramid fiber, Vfa —volume fraction of aramid fiber, Ec —elastic modulus of compound, Vc —volume fraction of compound. Equilibrium differential equations of composite housing fragment are [17–20]: @rxx @ryx a @rzx @rxy @ryy a @rzy @rxz @ryz a @rzz þ þ ¼ 0; þ þ ¼ 0; þ þ ¼0 h @h @h @@n @f @n @f @n @f ð5Þ Dimensionless coordinates n ¼ ax ; f ¼ ay ; - ¼ hz express the stresses as: ryz  rðh=aÞ; rzx  rðh=aÞ; rzz  rðh=aÞ2

ð6Þ

r—relative value of strength and h—thickness of fragment. Taking into account that size of composite housing in the dimension along the axis z is significantly smaller than sizes in the dimensions along two other axes x and y, so the displacements are expressed as: ux ðx; y; zÞ ¼

m X k¼0

ak ðx; yÞzk ; uy ðx; y; zÞ ¼

m X k¼0

bk ðx; yÞzk ; uz ðx; y; zÞ ¼

n X

ck ðx; yÞzk

k¼0

ð7Þ m þ 1; n þ 1—number of terms in the expansion and ak ; bk ; ck —coefficients for decomposition of function arguments x and y. The described power series allow determining of displacement values in relation to argument z. It is obvious, by the working pressure, the multilayer housing of the MR damper experiences transverse shear and tensile/compressive stresses. The tangent displacements are distributed over the cubic parabola of argument z, and normal displacements get along the quadratic parabola: ux ¼ a0 þ a1 z þ a2 z2 þ a3 z3 ; uy ¼ b0 þ b1 z þ b2 z2 þ b3 z3 ; uz ¼ c0 þ c1 z þ c2 z2 ð8Þ

1300

K. V. Naigert and V. A. Tselischev

As the main structural element of housing is composite liner, its specific strain energy is: ð3Þ

ð3Þ

ð 3Þ

W ð3Þ ¼ W1 þ W2 þ W3

ð9Þ

Specific strain energies are given as: ð3Þ

  1  ð3Þ ð3Þ 1 ð3Þ ð3Þ rxx exx þrðyy3Þ eðyy3Þ þrðxy3Þ cðxy3Þ ; W2 ¼ rðxz3Þ cðxz3Þ þrðyz3Þ cðyz3Þ ; W3 2 2 1 ¼ rðzz3Þ eðzz3Þ 2

W1 ¼

ð10Þ

ð3Þ

W1 —specific strain energy of composite liner in the considered plane of housing ð3Þ

ð3Þ

fragment, W2 —specific strain energy of transverse shear stresses, W3 —specific strain energy of transverse tensile and compressive stresses. Elastic characteristics of polymeric material of shell are presented as characteristic module: E  ExðiÞ  EyðiÞ  GðxyiÞ

ð11Þ

Elastic characteristics of composite liner are: Exð3Þ  Eyð3Þ  Gðxy3Þ  b1 E; Gðxz3Þ  Gðyz3Þ  b2 E; Ezð3Þ  b3 E

ð12Þ

b1 ; b2 ; b3 —ratio order of composite liner modules to the characteristic module. The order of major stresses and strains of shell are based on the Kirchhoff–Love model: r  rðxxiÞ  rðyyiÞ  rðxyiÞ ; e ¼ eðxxiÞ  eðyyiÞ  cðxyiÞ 

r E

ð13Þ

Equilibrium differential equations of shell fragment of multilayer housing are written: @rðxyiÞ @rðyyiÞ @rðzyiÞ @rðiÞ @rðyziÞ @rðzziÞ @rðxxiÞ @rðyxiÞ @rðzxiÞ þ þ ¼ 0; þ þ ¼ 0; xz þ þ ¼0 @x @y @z @x @y @z @x @y @z ð14Þ The main strength element of multilayer housing structure is the composite liner, so let us say that: @ @ 1 @ 1 h   ;  ; rðiÞ  rðyyiÞ  rðxyiÞ  gr; rðzziÞ  g2 r; g ¼ @x @y l @z h xx l

ð15Þ

Design and Calculation Method of Composite Housings for New …

1301

l—fragment length. Stress–strain states of multilayer fragment are: rðxz3Þ  rðyz3Þ  rðxziÞ  rðyziÞ  gr ¼ gEe; rðzz3Þ  rðzziÞ  g2 r ge g2 e ¼ g2 Ee; cðxz3Þ  cðyz3Þ  ; eðzz3Þ  b2 b3

ð16Þ

Major strains and specific strain energy of fragment can be rewritten as:

W ð3Þ

e  eðxx3Þ  eðyy3Þ  cðxy3Þ ; rðxx3Þ  rðyy3Þ  rðxy3Þ  b1 Ee   g2 g4 ð3Þ ð3Þ ð3Þ ¼ W1 þ W2 þ W3  Ee2 b1 þ þ b2 b3

ð17Þ

5 Calculation Results Main interest in the design of housings for MR devices has fiber composite liners. As it was already noted, reinforcing element of composite housing is fiber composite liner, which is formed by carbon fiber and aramid fiber matrix, and the space between fibers is filled with compound. Therefore, it is necessary to determine the strength characteristics of the composite liner by design and calculation of composite housings. This composite liner is manufactured as multi-directional reinforcing structure and bonded by compound. The optimization of stiffness and strength properties of composites are possible to realize by varying the volume fractions of reinforcing fibers and compound or by changing the composition of compound. Taking into account that the choice of compound depends on the expected loading conditions, polymer shell materials, operating temperatures and many other designs and operational features, it is rational to consider the impact of total volume fraction of components. Results of numerical experiment in the MATLAB represented stiffness and strength of composite material in influence of total volume fraction of reinforcing fibers (carbon and aramid fibers) and angle between the directions of carbon and aramid fibers (the dependencies 1 are for angle between carbon and aramid fibers 0°, and the dependencies 2 are for angle between carbon and aramid fibers 45°), Fig. 1. Simulation is performed for composite liner section with equal shares of volume fractions of reinforcing carbon and aramid fibers and for polyester resin compound. The obtained simulation results illustrate acceptable stiffness and strength characteristics of composite housing. By sufficient high volume content of reinforcing fibers and rational choice of reinforcement angle, the strength of composite material exceeds the strength of many metal alloys; it gives the low weight of product and corrosion resistance. This indicates the possibility and feasibility of using of composite housing in the MR systems.

1302

K. V. Naigert and V. A. Tselischev

Fig. 1 a Dependence of stiffness/elastic properties E of composite liner on total volume fraction of reinforcing fibers Vf . b Dependence of strength properties r of composite liner on total volume fraction of reinforcing fibers Vf

6 Novelty, Practical and Scientific Significance of Research Created design of composite housing for new generation MR devices has nonmetallic shell and composite liner; aramid frame of composite liner is manufactured from whole aramid fiber and reinforced by whole carbon fiber [12]. The proposed calculation method of multilayer composite housing is applicable to optimize the elastic characteristics and to design process of original MR damper housing.

7 Conclusions The multilayer housings with fiber composite liners have good mechanical properties. Therefore, it is advisable to perform the polymeric housings with composite liners. Multilayer structure allows production of precision surface treatments without reducing their strength. Proposed design achieves technical results such as increase in the strength and resource boost of nonmetallic housings of MR dampers (supports) and gives possibility to installation of electromagnetic control units on the device housings.

Design and Calculation Method of Composite Housings for New …

1303

References 1. Burchenkov VN et al (2000) Magnitozhidkostnoye ustroystvo dlya gasheniya kolebaniy (MR device for vibration damping). RU patent 2,145,394, 10 Feb 2000 2. Korchagin AB et al (2012) Reguliruyemyy magnitoreologicheskiy pnevmaticheskiy amortizator (Adjustable magnetorheological pneumatic damper). RU patent 2,449,188, 27 Apr 2012 3. Gusev EP, Plotnikov AM, Voevodov SYu (2003) Magnitoreologicheskiy amortizator (MR shock absorber). RU patent 2,232,316, 27 Oct 2003 4. Kudryakov YuB et al (1998) Magnitoreologicheskiy vibrogasitel’ (MR vibration damper). RU patent 2,106,551, 10 Mar 1998 5. Yamanin IA et al (2009) Dinamicheskiy gasitel’ (Dynamic absorber). RU patent 2,354,867, 10 May 2009 6. Gordeev BA et al (2015) Magnitoreologicheskiy amortizator (MR damper). RU patent 2,561,610, 27 Aug 2015 7. Harris CE et al (2002) Emerging materials for revolutionary aerospace vehicle structures and propulsion systems. SAMPE J 38(6):33–43 8. Strong AB (2006) Plastics: materials and processing, 3rd edn. Prentice-Hall Inc, Upper Saddle River, NJ 9. Warren CD (2001) Carbon fiber in future vehicles. SAMPE J 37(2) 10. Driver D (2000) Towards 2000—the composite engine. In: Meeting of Australian Aeronautical Society 11. Dubois D, Fargier H (2014) Capacity refinements and their application to qualitative decision evaluation. ftp://ftp.irit.fr/IRIT/ADRIA/PapersFargier/dubois-fargier-xkru09.pdf. Accessed 28 Nov 2018 12. Naigert KV, Tutynin VT (2018) Kompozitnyy korpus dlya magnitoreologicheskogo dempfera (Composite housing for magnetorheological damper). RU patent application 2018,130,967, the decision to grant a patent 6 Nov 2018 13. Beer F et al (2009) Mechanics of materials. McGraw-Hill Companies, New York 14. Harmsen SC, Rogers AM (1986) Inferences about the local stress field from focal mechanisms: applications to earthquakes in the southern Great Basin of Nevada. Bull Seism Soc Am 76:1560–1572 15. Grabisch M (2004) The Moebius transform on symmetric ordered structures and its application to capacities on finite sets. Discrete Math 287:17–34 16. Landau LD, Lifshits EM (1986) Theory of elasticity, 3rd edn. Heinemann, Oxford 17. Dudarev SL, Ma PW (2018) Elastic fields, dipole tensors, and interaction between self-interstitial atom defects in bcc transition metals. Phys Rev Mater 2:033602 18. Dederichs PH, Schroeder K (1978) Anisotropic diffusion in stress fields. Phys Rev B 17:2524 19. Skvortsov YuV (2013) Mekhanika kompozitsionnykh materialov. In: Mechanics of composite materials, Samara 20. Alexandrov S (2018) Elastic/plastic disks under plane stress conditions. Springer, New York, USA

Methodology and Constructive Implementation of Active Vibration Protection of Large-Scale Structures K. V. Naigert and V. A. Tselischev

Abstract It is advisable to implement active vibration protection by structures; those include one spatial dimension with the value, which exceeds the values of other two spatial dimensions. The proposed vibration protection method is based on the oscillation generation that leads to growing frequencies of large-scale structures (LSS). The rise of frequency shifts the natural frequency of LSS oscillations in relation to the disturbing frequency; at the same time, there is an increase in the frequency and decrease in the amplitude that allow getting away from the resonance. Another vibration protection method is oscillation generation, which coincides with damped LSS frequencies but is in the antiphase with them. The main task on the development of active vibration protection is exclusion from designs mechanical oscillators and moving mechanical elements, solution of the problem is the use of the ferromagnetic fluid volume as an oscillator. Constructive implementation of such active vibration protection device with a ferrofluid (FF) oscillator is considered in this research paper. The principle of operation of the device is based on the oscillation generation and on the change of the electromagnetic component value of pressure in the ferromagnetic fluid by pulsed electromagnetic fields and on imparting oscillatory trajectories to magnetic particles in the vortex electromagnetic fields too. Keywords Active vibration protection Ferrofluid oscillator


Viscoelastic damper


1 Introduction To maintain their integrity, LSS require vibration protection, because of large-amplitude oscillations, can cause the formation of defects in the material, strength reduction, and structure displacement. In the case of LSS oscillations, the K. V. Naigert (&) South Ural State University, 76, Lenin Avenue, Chelyabinsk 454080, Russia e-mail: [email protected] V. A. Tselischev Ufa State Aviation Technical University, 12, K. Marx Street, Ufa 450008, Russia © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_138

1305

1306

K. V. Naigert and V. A. Tselischev

resonance is especially dangerous. Currently, operated active vibration protection systems for LSS have mechanical oscillators and pressure sources containing moving mechanical elements and have long transient processes, which reduce performance. Common designs of active vibration protection devices generate the oscillations by mechanical oscillators that use the electromechanical or hydromechanical converters; it depends on the application of electromechanical drives or pressure impulse control systems for working chamber of support [1–9].

2 Actuality of the Research Existing active vibration protection systems have some disadvantages; many vibration protection systems include the moving mass of a mechanical oscillator in their designs that is inertial, which reduces transmission speed of control signal and decreases vibration protection device reliability. So, the development of vibration protection equipment and structures without the mechanical oscillators is an urgent task of research.

3 Statement of the Problem Active vibration protection methods are not new but standard designs of vibration protection devices have a lot of disadvantages and require working process improvement. Therefore, research objectives are improving the active vibration protection method, its constructive implementation, and development of the calculation method for new FF active vibration protection device construction.

4 Constructive Solution of the Problem, Numerical Model of Active Vibration Protection Process, and Method of Active Vibration Protection of LSS Active vibration protection device includes FF oscillator, the design is published in a patent application (the decision to grant a patent RU 2018124391) [10]. In order to increase the efficiency of eliminating resonant vibrations, the ferromagnetic fluid volume is used and placed in the external electromagnetic fields, which are generated by a vortex toroidal inductor. The oscillations are created at pulsed changes of electromagnetic pressure component and by an oscillatory motion of magnetic particles in the vortex electromagnetic fields. Electromagnetic pressure component is part of total pressure in the ferromagnetic fluid volume. The frequency increases and leads to shifting the natural frequency of LSS oscillations in relation to the disturbing frequency. Oscillations increase in the frequency and reduce the amplitude. By vibration

Methodology and Constructive Implementation of Active Vibration …

1307

protection, the FF oscillator uses the damping frequencies, which coincide with damped LSS frequencies but are in the antiphase with them. The control of operating parameters of FF active vibration protection device for LSS is carried out by means of frequency phase and current–voltage characteristics’ changes of an electrical signal. Direct current in the ring coil leads to change in the viscosity offerromagnetic fluid and makes it possible to operate the device as a rheological vibration damper. The use of the combination of FF oscillator and viscoelastic damper significantly improves the device performance, extends the frequency range, and allows reducing the impact of shock loads on the structures. Therefore, the combination of ring coil and toroidal inductor, which generates the control vortex electromagnetic field, is useful. The modeling of the viscosity characteristics of ferromagnetic fluids eliminates the resonance and shock loads by operating in the viscoelastic damping mode. Elimination of resonant vibration in LSS was founded on the approximation of oscillations’ frequency of disturbing forces to lower tones of LSS natural frequencies. Disturbing forces are created by FF oscillator. Vortex electromagnetic fields cause the rotation and movement of magnetic particles; it happens due to a change in the kinetic energy of magnetic particles and in the orientation of magnetic particles by external electromagnetic fields that give to particles radial or axial component of velocity. It creates mechanical vibrations. Located in the vortex electromagnetic fields, magnetic particles acquire an oscillatory motion trajectory (translational and rotational return trajectory). Vortex electromagnetic fields are induced by the differential coil with multilayer windings; these fields interact and form trajectories of the magnetic particles. Toroidal inductor creates the vortex variable control electromagnetic fields; the control is realized by the electrical signal generator and is based on the regulated changes in frequency phase and current–voltage characteristics of control electrical signal. The required frequency phase and current–voltage characteristics of the control electrical signal are initiated by the controller as feedback responses to signal of sensor. The active vibration protection process is considered to lead by the example of pipeline, which is fixed on the edges. It is important to describe the process of pipeline oscillations. Pipeline oscillation and shape of oscillation equations are based on Kirchhoff model (in the case of bending) [11]: @4w @2w @2w  ðT  Pi Fi Þ 2 þ ðqF þ qi Fi Þ 2 ¼ 0 @x4 @x @t @ 4 w T  Pi Fi @ 2 w ðqF þ qi Fi Þx2
2 w¼0  @x4 @x EJ EJ pffiffiffiðqF þ qi Fi Þx2 @w x w ðT  Pi Fi ÞL2 ;k¼L 4 ¼ 0ðx ¼ 0; LÞ; n ¼ ; z ¼ ; p ¼ w ¼ 0; 2EJ ! EJ @x rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi!L L rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi qffiffiffiffiffiffiffiffiffiffiffiffiffiffi qffiffiffiffiffiffiffiffiffiffiffiffiffiffi z ¼ A cos n p2 þ k4  p þ B sin n p2 þ k4  p EJ

rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi! qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi! 4 2 þ Cch n p þ p þ k þ Dsh n p þ p2 þ k4

ð1Þ

1308

K. V. Naigert and V. A. Tselischev

E—elastic modulus of pipeline, q—material density of pipeline, qi —fluid density, Pi —pressure in the pipeline, T—tensile force in the pipeline, w—pipeline deflection, x—coordinate, t—time, J—axial moment of inertia, F—cross-sectional area of pipeline, Fi —effective cross-sectional area, x—frequency, L—length of pipeline, z—displacement. We write motion equations of FF active vibration protection device [12–14]: m€z þ g_z þ cðtÞz ¼ Fd sin xt €z þ 2h_z þ x2 z ¼ Fr sin xt Fd €z þ 2h_z þ x2 z ¼ sin xt m z ¼ Fm sinðxt þ u0 Þ h pi vos ¼ Fm x cosðxt þ u0 Þ ¼ Fm x sin ðxt þ u0 Þ þ 2 2 2 aos ¼ Fm x sinðxt þ u0 Þ ¼ Fm x sin½ðxt þ u0 Þ þ p

ð2Þ

m—mass of volume, c—coefficient of proportionality, which characterizes the restoring force, z_ —oscillatory speed, €z—acceleration of oscillating mass, g—coefficient of fluid friction, Fd —disturbing force, Fr —restoring force, h—reduced viscosity drag coefficient, Fm —amplitude of disturbing force, u0 —initial phase, vos —speed of oscillator displacement, aos —acceleration of oscillator displacement. It is observed that the resonance phenomena occur in the case of amplitude increase of the oscillations by external disturbing force. Therefore, it is necessary to describe the amplitude of ferromagnetic fluid volume oscillations in the FF active vibration protection device. The oscillation frequency of ferromagnetic fluid volume depends on the pressure changes in the working chamber of support; it is set by the frequency of inclusion of an electromagnetic control unit. Proposed method is based on the generation of impulse perturbations by FF oscillator. The determination of optimal frequency parameters helps to avoid the resonance; it allows the joint solution of equations of the pipeline and FF support oscillations. Resonant frequency of oscillatory process is defined by the maximum allowable pipeline deflection. Based on the obtained values of resonant frequency of pipeline oscillation, the required frequency of FF oscillator is solved. The main condition for the lack of resonance is described as [15–19]: Fm ðtÞ

ln Fm ðt þ Tm Þ x0 v d  pffiffiffi ; d ¼ ¼ Tm Tm 2

ð3Þ

x0 —undamped natural frequency of oscillation, Tm —period, d—attenuation coefficient, v—logarithmic decrement of attenuation. The proposed design of FF support allows the complex method regulation, namely, it is the modeling a number of parameters which have an influence on the oscillatory process. These parameters are: viscosity of FF, frequency of FF oscillator, and restoring force, which is formed

Methodology and Constructive Implementation of Active Vibration …

1309

by increasing pressure in the working chamber of FF support. Increasing the pressure is the result of changes in the electromagnetic component value of total pressure. The maximum allowable displacement of FF support is calculated on the basis of the maximum allowable pipeline deflection by undamped natural frequency of oscillation. The values of resonant amplitude and resonant frequency can be determined by expressions: Fmax:m Fmax:m F  ¼ max:m zr ¼ qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ¼  2 sM ð H Þ g gr þ 14
1 þ ðsss HM ðH ÞÞ=J xd 2dmxd g x20  4m 2 vffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi  2 u u ðH Þ qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi gr þ 14
1 þ ðsssM t s HM ðH ÞÞ=J ¼ x20  2d2 xr ¼ x20  2m2

ð4Þ

xd —frequency by attenuation, g—viscosity of ferromagnetic fluid by control electromagnetic field, gr —viscosity of ferromagnetic fluid without control electromagnetic field, M—magnetization, H—magnetic field strength, zr —resonant amplitude, xr —resonant frequency. Modulation of required displacement values of FF support is achieved by varying parameters: ferromagnetic fluid viscosity [20], frequency of FF oscillator, and restoring force; this is obtained by regulating the parameters of control electromagnetic field. Pressure in the working chamber of FF support can be solved as: Fr Fd SMRC ¼ cjzj ¼ pf þ pem ¼ SMRC0 m 1 ZH   0 0 ¼ pf þ @const  qgz þ M H dH A

PMRC ¼

ð5Þ

0

pf —initial pressure in the working chamber. The way to avoid the resonance could be selected on the basis of values of critical amplitude–frequency characteristics for oscillatory process. It is preferable for large values of maximum allowable pipeline deflection, the method of impulse oscillations, which lead to increase in the frequency of LSS oscillations but reduce the amplitude. By frequency control, the equation of resulting oscillations effect takes the form:     ztot ¼ zp þ zMRS ¼ Zm sin xp t þ sinðxMRS tÞ  x  x  xp þ xMRS p MRS ¼ 2Zm cos t sin t 2 2

ð6Þ

Zm —amplitude of both oscillations, zp —displacement of LSS, zMRS —displacement of FF support, xp —oscillation frequency of LSS, xMRS —oscillation frequency of FF support. In the cases of small values of maximum allowable pipeline deflection, the use of phase control method is rational and is based on the generation of oscillations, which coincide with damped LSS frequencies but are in the antiphase

1310

K. V. Naigert and V. A. Tselischev

with them. The displacement at the resulting oscillations by phase control is calculated by: ztot ¼ zp þ zMRS ¼ ZmP sinðxt þ uoP Þ þ ZmMRS sinðxt þ uoMRS Þ ¼ Zmtot sinðxt þ uotot Þ

ð7Þ

ZmP —oscillation amplitude of LSS, ZmMRS —oscillation amplitude of FF support, uoP —initial phase of oscillation of LSS, uoMRS —initial phase of oscillation of FF support. The amplitude and initial phase of the resulting oscillations are written as: Zmtot ¼

qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 þ Z2 ZmP mMRS þ 2ZmP ZmMRS cosðuoP  uoMRS Þ

uotot ¼ arctg

ZmP sinðuoP Þ þ ZmMRS sinðuoMRS Þ ZmP cosðuoP Þ þ ZmMRS cosðuoMRS Þ

ð8Þ

5 Calculation Results By numerical experiments, the analytical dependencies are prepared. Results represent the influence on increase in the frequency, which leads to frequency shift of LSS oscillations in relation to the disturbing frequency and to vibration amplitude reduction. These MATLAB simulation results are presented in Figs. 1, 2, and 3. The pictures illustrate the dynamics of vibration protection process of pipeline (dynamics of changes in the pipeline displacement values z; mm in the time t; ms, s). Numerical simulation of high-frequency vibration of pipeline represents the typical case of pipeline loading, for example, by high seismic activity (seismic events). A wide band of oscillation amplitude (pipeline displacement) is shown in

Fig. 1 The vibration of pipeline

Methodology and Constructive Implementation of Active Vibration …

1311

Fig. 2 The vibration of pipeline by active vibration protection mode

Fig. 3 The vibration of pipeline by active vibration protection and viscoelastic damping modes

Fig. 1. High-frequency oscillations of FF oscillator can form the significantly narrow band of oscillation amplitude and lead the values of fundamental oscillations in the band of oscillation amplitude close to zero, Fig. 2. It saves the structural integrity and minimizes damages of vibration exposure. By simultaneously using the active vibration protection and viscoelastic damping modes can realize the rapid attenuation of vibration impulse and reduce to zero pipeline axis deviation, Fig. 3. The implementation of active vibration protection on the lower frequencies is useful for the regulation of deposition degree of paraffinic fractions in the oil; it preserves the homogeneity of dispersed environment and prevents the separation of fractions of solutions and suspensions. The viscoelastic damping makes it possible to effective suppression of vibration and shock loads and provides static stability of LSS. The design of FF support has the adaptive viscoelastic properties, which are

1312

K. V. Naigert and V. A. Tselischev

changed in the real time by controlled electromagnetic field. Simulation results illustrate the high efficiency of proposed method of active vibration protection of LSS, significant narrowing of band of fundamental amplitude, and efficiency of viscoelastic damping mode, which allows additional vibration damping of ferromagnetic environment at short time of transient process and leads the oscillation amplitude values to zero. This proves the appropriateness of combination of active vibration protection and viscoelastic damping modes.

6 Novelty, Practical, and Scientific Significance of Research The proposed design of active vibration protection device with FF oscillator uses the ferromagnetic fluid volume in order to increase the efficiency of the resonant vibration removal, viscoelastic damping mode, and vortex control electromagnetic fields [10]. The developed new method of calculation of active vibration protection device characteristics is applicable to design of active vibration protection device for LSS.

7 Conclusions The use of presented active vibration protection device allows expanding of frequency range of vibration disturbing forces, providing the LSS protection by resonant vibration mode and optimizing the active vibration protection device design. Active vibration protection device does not require the mechanical parts by realization of changes in the energy of ferromagnetic particles and formation of radial or axial velocity components of magnetic particles in the control electromagnetic fields. It reduces the inertia of the system, decreases the transient time, and increases the reliability. Combination of FF oscillator and viscoelastic damper significantly improves the performance of active vibration protection device, extends the frequency range, allows reducing of impact on shock loads, and maintains integrity of LSS. New technical solutions are based on the proposed active vibration protection method, which was implemented in the construction and tested by numerical experiment. The results of research form the way of improvement of methodology of active vibration protection of LSS.

References 1. Burchenkov VN et al (2000) Magnitozhidkostnoye ustroystvo dlya gasheniya kolebaniy (Magnetorheological device for vibration damping). RU patent 2,145,394, 10 Feb 2000

Methodology and Constructive Implementation of Active Vibration …

1313

2. Korchagin AB et al (2012) Reguliruyemyy magnitoreologicheskiy pnevmaticheskiy amortizator (Adjustable magnetorheological pneumatic damper). RU patent 2,449,188, 27 Apr 2012 3. Gusev EP, Plotnikov AM, Voevodov SYu (2003) Magnitoreologicheskiy amortizator (Magnetorheological shock absorber). RU patent 2,232,316, 27 Oct 2003 4. Kudryakov YuB et al (1998) Magnitoreologicheskiy vibrogasitel’ (Magnetorheological vibration damper). RU patent 2,106,551, 10 Mar 1998 5. Yamanin IA et al (2009) Dinamicheskiy gasitel’ (Dynamic absorber). RU patent 2,354,867, 10 May 2009 6. Mikhailov VP et al (2012) Magnitoreologicheskaya pozitsioniruyushchaya i vibroizoliruyushchaya sistema (Magneticoreological positioning and vibrational insulation system). RU patent 2,443,911, 27 Feb 2012 7. Gordeev BA et al (2015) Magnitoreologicheskiy amortizator (Magnetorheological damper). RU patent 2,561,610, 27 Aug 2015 8. Deshmukh SS, McKinley GH (2012) Fluid-filled cellular solids for controlled. US patent 8,091,692, 10 Jan 2012 9. Vlasov AV (2011) Sposob dempfirovaniya kolebaniy podvizhnoy sistemy i ustroystvo dlya yego osushchestvleniya (Vibration damping method for mobile system and device for its implementation). RU patent 2,426,922, 20 Aug 2011 10. Naigert KV, Tutynin VT (2018) Magnitoreologicheskoye ustroystvo aktivnoy zashchity dlinnomernoy konstruktsii ot rezonansnoy vibratsii (MR device of active protection of large scale structures from resonant vibration). RU patent application 2018,124,391, the decision to grant a patent 21 Nov 2018 11. Khakimov AG (2014) Opredeleniye osevogo usiliya, plotnosti zhidkosti v truboprovode po sobstvennym chastotam kolebaniy (Determination of axial effort, fluid density in pipeline by own frequency vibrations). IVTN-2014: 09.09.2014, pp 1–7. http://www.ivtn.ru/2014/pdf/ d14_08.pdf. Accessed 28 Nov 2018 12. Inman Daniel J (2001) Engineering vibration. Prentice Hall 13. Thompson WT (1996) Theory of vibrations. Nelson Thornes Ltd 14. Tongue B (2001) Principles of vibration. Oxford University Press 15. Fan H, Gao F (2014) Asymptotic stability of solutions to elastic systems with structural damping. Electron J Diff Equat 245:1–9 16. Fan H, Li Y (2014) Analyticity and exponential stability of semigroups for the elastic systems with structural damping in Banach spaces. J Math Anal Appl 410:316–322 17. Balescu R (1975) Equilibrium and nonequilibrium statistical mechanics. Wiley 18. Suli E, Mayers D (2003) An introduction to numerical analysis. Cambridge University Press 19. Ciccotti G, Hoover W (1986) Molecular-dynamics simulation of statistical-mechanical systems: Varenna on Lake Como, Villa Monastero, 23 July–2 August 1985. In: Proceedings of the international school of physics Enrico Fermi. North-Holland, pp 43–65 20. Takeketi S, Tikazumi S (1993) Magnitnyye zhidkosti (Magnetic fluids). Mir, Moscow

Substantiation of Parameters of Machine with Volumetric Hydraulic Drive for Formation of Wells in Ground A. V. Gorin, N. V. Tokmakov and I. S. Kyznetsov

Abstract The article discusses the prerequisites for the substantiation of the parameters of the machine with a volumetric hydraulic drive. A rheological model of the soil is presented in the modernized Maxwell model. A mathematical model of the operation mode of the machine with a volumetric hydraulic drive for the formation of wells in the ground is presented. A structural diagram of the combined drive of the working body of the machine is presented. For the experimental confirmation of the obtained analytical dependencies and results, a full-size test stand has been developed. As a result of the analysis of the data obtained on the test bench and field tests, the graphical dependence of the power consumption of drilling through on a single machine impact is presented. A dimensionless coefficient of power impact on the face was proposed, allowing to evaluate the possible use of a machine with a volumetric hydraulic drive for the formation of wells in the soil.




Keywords Hydraulic drive Experimental stand Rheological model Soil Well








Mathematical model


1 Introduction Currently, there is a steady trend of increasing volumes of urban and industrial construction, reconstruction of the existing part of cities, and industrial enterprises [1, 2]. This trend will continue in the nearest and long run. The integral process of this perspective is the construction of new underground engineering communicaA. V. Gorin (&)
N. V. Tokmakov Orel State University named after I.S. Turgenev, 95, Komsomolskaya St., Orel 302026, Russia e-mail: [email protected] I. S. Kyznetsov Orel State Agrarian University named after N.V. Parahin, 69, Komsomolskaya St., Orel 302020, Russia © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_139

1315

1316

A. V. Gorin et al.

tions [3, 4] for various purposes: sewage, plumbing, electrical and heating networks, gas pipelines, etc. The construction of underground pipelines by the open (trench) method in the conditions of the city [5, 6] is connected with certain difficulties: necessity of disassembly and then restoration of road surfaces; traffic disruption; pollution. The closed (trenchless) method of construction is devoid of many of the listed drawbacks [7, 8], including both the construction of pipelines under the roads and the construction of extensive underground the collectors for engineering communications. The analysis of the works published on the research topic showed that 70% of underground pipelines have a diameter of up to 300 mm only in city conditions. Nowadays, there is a variety of machines for trenchless construction of communications; their effectiveness largely depends on the correct choice of their schemes, development of designs, and calculation parameters. However, scientifically based methods for assessing the choice of these parameters have not yet been fully completed and need to be adjusted and refined. Therefore, the substantiation and selection of the most effective parameters of machines that make use of modern hydraulic pressure mechanisms with great force (1500 kN and more) in combination with shock and vibration effects on the face with the possibility of their regulation in automatic mode is an actual problem for trenchless laying of wells in the soil. Existing information on the combined method of the trenchless well laying up to 300 mm in diameter makes it possible to characterize it with a much lower energy intensity of the process, mainly due to a significant increase in the rate of drilling through.

2 Simulation of the Machine Operation Process The dynamic theory of any technological machine designed to transform materials is built on the knowledge of the mechanical model of the object of influence with which the working body of this machine interacts [9, 10]. The external environment with which the investigated combined machine interacts is the soil. The description of the physicomechanical properties of the soil is based on the rheological models [11–13]. The closest model that describes the rheological properties of the soil in our case is the Maxwell model (Fig. 1). The consistent connection of elements according to the third law of Newton means that the same forces act on both components of the model (shear stresses), and the deformations of the elastic and viscous elements add up. As a result of studying and analyzing the work of predecessors [14–16], our own research and our understanding of the processes occurring during the puncture and punching a well in the soils with the use of shocks and vibrations, the conclusion is drawn about the expediency of using the Maxwell model in a modified (modernized) representation, specifically:

Substantiation of Parameters of Machine with Volumetric …

1317

Fig. 1 Rheological model of soil

– initially, viscous flow of the soil occurs under static exposure, while the material is compacted by filling the pores with a substance, its density increases; – when a certain (maximum force) is reached, the soil becomes more dense and (with approximation) can be considered elastic; at this moment of time, the action of the shock impulse is added to acting static force (Fig. 2). The structural scheme of the combined working body of the machine (Fig. 3) consists of a combined volumetric hydraulic drive, which includes a pressure hydraulic cylinder (HC); hydraulic hammer (HH); pressure switch (PS); distributors (P1, P2); hydro accumulator (A); pump (H); oil tank (T). When compiling a mathematical model, we assume that the machine is driven by a volumetric hydraulic pump, which supplies working fluid to a pressure line with constant performance, the flow process is considered to be isothermal. In this case, the dynamic model of the “machine-soil” system is reduced to a single-mass system with reduced rigidity and weight of the machine and reduced rigidity of the soil; we neglect the impact phase as it is 100 or more times less than other phases. In drawing up the design scheme of the working body of the machine (Fig. 4), the following notation is used: soil reaction (RC); the reduced mass of the working body and the soil (M); reduced machine stiffness (C); shock pulse (QS); fluid speed reduced to a hydraulic cylinder (V0).

1318

A. V. Gorin et al.

Fig. 2 Diagrams: a soil resistance forces; b shock impulses

Fig. 3 Structural diagram of the combined drive of the working body of the machine

Fig. 4 Design scheme of the working body of the machine

The differential equation of motion of the working body is written in the form of a relation (1). € ¼ ðV0 t  XÞC  R0  C1 X: MX

ð1Þ

The elastic properties of hydraulic transmission are characterized by its reduced to a pressure hydraulic cylinder, stiffness C (Fig. 3), 1 1 1 1 1 ¼ þ þ þ ; C CP CA CF CHC

ð2Þ

Substantiation of Parameters of Machine with Volumetric …

1319

where CP—the stiffness of elastic pipelines; CA—the stiffness of the pressure accumulator; CF—the fluid stiffness; CHC—the stiffness of the pressure cylinder. The mass of the working body M includes the mass of the pipe, pressure hydraulic cylinders, impact mechanism, and tool. The groundmass participating in the movement with the tool is taken into account using the coefficient: lGR ¼

mGR ; mPR

ð3Þ

Then the total mass of the working body and the ground will be equal to: M ¼ mPR þ mGR ¼ mPR ð1 þ lGR Þ:

ð4Þ

A simplified view of the soil resistance force chart is presented in Fig. 3, a. We accept that the resistance force RC changes according to the law: RC ¼ R0 þ C1 X;

ð5Þ

where C1—the soil stiffness; R0—the initial value of resistance force. The impact pulses of QS are transmitted to the instrument periodically at a time when the resistance force reaches the maximum value Rm (Fig. 2, б), while the working body instantly receives the speed VS in the direction of motion. The mathematical model of this mode of movement of the machine has the following form: ðVS  C0 V0 Þ sin kt; k X_ ¼ C0 V0 þ ðVS þ C0 V0 Þ cos kt;
 ðVS  C0 V0 Þ sin kt ; R ¼ R0 þ C V0 tð1  C0 Þ  k X ¼ C0 V0 t þ

ð6Þ

One of the main advantages of this model is the possibility of theoretical determination not only of the speed and movement of the tool but also of the force acting on the face at any time. The productivity of the machine (the rate of drilling through) in this operation mode is presented with the form: VPR ¼ V0 

2pV02 ð1  C0 Þ2 pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi : 2pV0 ð1  C0 Þ þ VS ðVS  2C0 V0 Þ

ð7Þ

The computational experiment to assess the influence of the working and geometrical parameters of the hydraulic drive of the combined machine was carried out using a program developed in the engineering applications environment MatLab and Excel.

1320

A. V. Gorin et al.

3 Experimental Studies The experimental studies were carried out in two stages: conducting research on an experimental full-size stand; the performance of field tests in production conditions. The experimental stand (Fig. 5) consists of a mechanical 1, a power 2, and a control and measuring system 3. The mechanical part has a hydraulic impulse drive with a free drain 1, mounted on a mobile platform 4 (Fig. 6), which moves along the base 3 with a power cylinder 2, abutting against the supporting wall 6. The loads generated by the pulse drive and the power cylinder are perceived by the damper 5. The test bench power system is a KPG-1421 hydraulic pump station, specially designed for this test bench. It consists of two pumps. The first is connected to a pulse drive; the second feeds the power cylinder. The control and measuring system consists of a number of sensors registering the specified parameters. The signals from the sensors are received by the YT-4 amplifier (7), the amplifier transmits it to the converter 10, which is powered from the power supply unit 9, or to the computer. The field tests under production conditions were performed at the landfill of Orel city. The drilling through was carried out by pipes 80, 160, 230, and 320 mm with the use of a static–dynamic machine SDM-1500 see Fig. 7. The length of drilling through ranged from 15 to 28 m at a depth of 1.5 to 2 m from the day surface. The tests were comparative, i.e., the other things being equal (soil properties, pipe diameter, laying depth, etc.), axial forces were determined during the static penetration and the impact on the pipe.

Fig. 5 Test stand: 1—stand, 2—hydraulic station КPG-1421, 3—control and measuring equipment

Substantiation of Parameters of Machine with Volumetric …

1321

Fig. 6 The scheme of the experimental stand and the connection of the control and recording equipment

Fig. 7 Static–dynamic machine SDM-1500 under production conditions: 1, 2—pressure and impact mechanisms; 3—pipe; 4—instrument; 5—the impact mechanism body; 6—frame; 7—retaining wall; 8—drive of pressure and impact mechanisms; 9—drive of pulse mechanism; 10—drive tracking control system; 11—ADC board; 12—computer

1322

A. V. Gorin et al.

Fig. 8 The dependence of the energy of drilling through HD from the energy of a single blow A of the machine

The study of the effect of operating parameters on the drilling through force and well-laying performance was carried out during the first full-scale test. The dependence of the energy intensity of drilling through HD with a dynamic load application from a single impact energy A (Fig. 8) shows that the static force FST has a pronounced extremum, which corresponds to the optimal impact energy A. So, with FST = 800 kN, the optimum energy of a single strike is A = 860 J, with FST = 600, 400, and 200 kN, respectively A = 670, 600, and 400 J. In this case, the minimum value of energy consumption corresponds to the mode at FST = 600 kN, A = 670 J and has a value 1.3 kW h/m3. The performance of the well drilling through largely depends on the combination of the geometrical and operating parameters of the machine. Expedient to use the following dimensionless parameter coefficient of the force impact, which includes the geometric and operating parameters of the machine: P PGR Si KF ¼ ; ð8Þ FR where RPGR—the total pressure of the front and side faces; S—the total surface area of the element to be laid; FR—the total force of the pressure and impact mechanism. The comparative analysis of theoretical and experimental data showed satisfactory results. The discrepancy composes 13–16%.

4 Conclusion The rheological model of the soil presented in the article in the form of a modified Maxwell model makes it possible to carry out theoretical calculations quite easily and accurately. The test bench based on a pulsed hydraulic drive provides experimental results in a wide range. As a result of the field tests, the dependence of the energy of penetration of the well on the energy of a single impact of a pulsed hydraulic drive was obtained. Taking into account the influence of the working and geometrical parameters of the machine with a hydraulic drive for the formation of wells in the soil, a new power factor is proposed.

Substantiation of Parameters of Machine with Volumetric …

1323

Acknowledgements This work was supported by the Ministry of Education and Science of the Russian Federation under the project No 9.2952.2017/4.6.

References 1. Balandinsky E, Vasiliev V, Minaev V, Ladyzhensky V (1991) Trenchless laying of engineering communications. Subsoil, Moscow 2. Vasiliev N (1994) Closed piping. Subsoil, Moscow 3. Gorin A, Yeshutkin D, Gorina M (2015) Volumetric hydraulic drive of the combined machine for formation of wells in soils. State University—USPC, Orel 4. Gorbunov V, Yeshutkin D, Piven G (1992) Hydraulic and pneumatic drilling hammers. Institute of Mining, Novosibirsk 5. Ushakov L, Kotylev Y, Kravchenko V (2000) Hydraulic impact machines. Engineering, Moscow 6. Bashta T (2005) Hydraulics, hydraulic machines and hydraulic drives. Engineering, Moscow 7. Gorin A, Yeshutkin D, Gorina M (2015) The use of hydraulic impact machines for the formation of wells in the soil. State University—USPC, Orel 8. Kushul M, Shlyakhtin A (1988) Theory of vibration immersion of cylindrical rod in elastic-plastic medium. Subsoil, Moscow 9. Rakishev B, Sherstyuk B, Yastrebow E (1998) Drilling of special wells in frozen rocks. Subsoil, Moscow 10. Lavrov G, Saratov T (1978) Mechanization of construction of trunk pipeline crossings under roads and railways. Subsoil, Moscow 11. Gorin A, Eshutkin D, Zhuravleva A (2011) Modeling of actuator static and dynamic machines for the trenchless construction of pipelines. Fundam Appl Probl Eng Technol 3(287):20–26 12. Zelenin A (1968) Foundations of destruction of soils by mechanical means. Engineering, Moscow 13. Protasov Y (1985) The theoretical foundations of mechanical destruction of rocks. Subsoil, Moscow 14. Kershenbaum N, Minaev V (1984) Drilling of horizontal and vertical wells by impact method. Subsoil, Moscow 15. Holstein M (1983) Mechanical properties of soils. Stroiizdat, Moscow 16. Yeshutkin D, Zhuravleva A, Gorin A (2011) The rigidity of the pressure line elements of hydraulic impact machines. Bull TulGU Ser Actual Probl Mech 7:58–63

Research of Pneumodrive with Energy Recovery into Additional Volume A. N. Sirotenko, S. A. Partko and S. A. Voinash

Abstract The authors consider the issue of energy saving in pneumatic drives of the technological equipment. The principal circuit of the energy saving pneumatic drive is provided. The energy consumption decreases at the retardation of the output element of the pneumatic drive with the help of reverse pressure, blocking the pneumatic cylinder chambers and the recuperation of the energy of the air, compressed in the output chamber, in an additional volume, with its subsequent use for the pneumatic drive output element reverse motion. The authors present a mathematical model of dynamic processes, describing the acceleration and retardation of the pneumatic drive output element. Changing the initial parameters of the additional volume provides an opportunity for the efficient control of energy and speed-drive characteristics. The authors compared the known pneumatic drive retardation methods with the suggested retardation method by means of the energy recuperation with pre-set initial parameters.




Keywords Pneumatic drive Recovery Braking Speed performance





Backpressure
Additional volume


1 Introduction Because of the known advantages, pneumatic drives are actively applied for automation and mechanization of auxiliary and main technological equipment operations [1]. Such operations include a test [2], weighting and pre-packing,

A. N. Sirotenko
S. A. Partko (&) Don State Technical University, 1, Gagarin Square, Rostov-on-Don 344000, Russia e-mail: [email protected] S. A. Voinash Rubtsovsk Industrial Institute (Branch) of Federal State Budgetary Educational Institution of Higher Education Polzunov, Altai State Technical University, 2/6, Traktornaya Str., Rubtsovsk 658207, Russia © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_140

1325

1326

A. N. Sirotenko et al.

packing and insulating [3, 4] and auxiliary technological operations [5]. These peculiarities preconditioned an active pneumatic fixing of actuated motions of the weighing and packing equipment, including packing automatic devices (ADN, ADNK, ADNK LR, AFB) [3, 4]. There are various known methods to increase energy and speed characteristics of a pneumatic drive. Some of them include decreasing energy consumption by means of reverse pressure retardation, which is relevant and of scientific and engineering interest [1, 6–8]. The key feature of this method consists in the presence of some significant volumes of compressed air inside the chamber of a pneumatic cylinder at its acceleration which prevents obtaining accurate data on pressure change and the rate of the pneumatic drive output elements with the help of simplified calculation methods. These parameters are input data to calculate retardation parameters [9].

2 Purpose of the Paper To find out a mathematical dependence between the initial parameters of additional volume, parameters of the air, compressed in the pneumatic cylinder chambers, and the braking distance of its output element.

3 Article Main Part The principal circuit of a pneumatic drive is given in Fig. 1. The response rate of a pneumatically fixed rotary platform of a packing and filling machine hinders a shock-free shutdown, and kinetic energy needs to be dissipated with a ribbon brake. For these conditions, it is reasonable to use retardation with a reverse pressure, i.e., by means of creating an overpressure in the pneumatic drive exhaust chamber. For the purpose of retardation energy recuperation, an additional volume was connected to the pneumatic cylinder output chamber. During retardation, the operating medium is compressed inside a brake chamber of the pneumatic drive and the additional volume. After the equipment stop, the accumulated energy is used for other actuated motions or for the pneumatic engine reversing [4, 5, 8]. The actuated body fixation can be arranged ‘on the rebound’ of a stem after its insignificant ‘overtravel’ due to spring stop [3] or the application of built-in [6] or external hydromechanical devices [6, 10–16]. A more active retardation is provided by a complete overlap of a pump main and the pneumatic drive exhaust, i.e., retardation proceeds with a constant mass of air inside the pneumatic drive chambers [6]. Changing the additional volume initial parameters (both the pressure and the value of a geometrical volume) in such a way one has an opportunity to control the braking effort at the drive piston, the pneumatic drive fast action and energy consumption [8, 17]. The initial parameters of such additional volume are selected so as to ensure the amount of the compressed air, accumulated at retardation, is enough

Research of Pneumodrive with Energy Recovery …

1327

Fig. 1 Pneumatic drive of the weighting and packing machine: (a) general view before modernization; (b) general view after modernization; (c) principal circuit of the cost-effective pneumatic drive with energy recuperation in the additional volume

Table 1 Sequence of the pneumatic drive component connection No

Operation name

Sensors, electrical magnets, power bodies (information on actuation) S1 S2 DCV1 DCV2 DCV3 DCV4 PC

1 2 3 4 5

Initial position Turret rotation Turret slowdown Reversing motion Stop

+ – – – +

– – + – –

– – – – –

– + – – –

– + + + –

+ – – – –

! ! ←

for a full return of the pneumatic drive stem into its initial position. The sequence of the pneumatic drive component connection is presented in Table 1. A mathematical description of thermodynamic processes in the suggested pneumatic drive was made on the following assumptions [9, 18]: pressure in the mains was taken as constant; the processes in the suggested pneumatic drive chambers were assumed to be quasistationary; the process of flow and filling of the pneumatic drive chambers with compressed air is considered as being polytropic with a variable polytropic parameter; compressed air was studied as ideal gas.

1328

A. N. Sirotenko et al.

The process of flow and filling of the pneumatic cylinder chambers is considered as a polytropic one with a variable polytrope [6, 9]. The control valve spool switchover is considered to be instantaneous. The design circuit of a pneumatic drive under research is given in Fig. 2. Compressed air is fed to the main (parameter with the ‘m’ index) from an infinitely large volume. The main is communicated with a channel having the section area f1 and the flow coefficient l1, with the pneumatic drive decreasing space (parameters with the index ‘p’). The operating medium, displaced from the pneumatic drive exhaust chamber (parameters with the index ‘b’), passes through the channel with the section area f2, the flow coefficient l2. The compressed air is exhausted into atmosphere in the indefinite volume (atmosphere, parameters with the index ‘a’). Designate the length of ‘passive’ volumes of the operating and exhaust chambers as x01 and x02, correspondingly—friction force. Under the pressure action in the jetting chamber, the piston, overcoming the friction force (P0), starts moving with the rate dx/dt, displacing the mass «m» towards the stem. The piston acceleration starts. The compressed gas exhaust lasts until the value of a current displacement coordinate (x) is equal to the control valve spool switchover for retardation (xтп). After that the control valve spool shuts off the channels (f1 = 0 and f2 = 0) and simultaneously connects and exhaust chamber of the pneumatic engine with the additional volume, pre-set initial parameters (parameters with the index ‘ak’). Further motion of the pneumatic drive piston takes place under the reverse action of the air compressed

Fig. 2 Design scheme of pneumatic actuator with compressed air energy recovery into an additional volume

Research of Pneumodrive with Energy Recovery …

1329

inside the exhaust chamber and the additional volume. The pressure inside the exhaust chamber will keep growing while the pressure in the jetting chamber— decreasing. Thermodynamic processes in the pneumatic drive chambers at the acceleration are described by a known equation system [9] while the retardation process takes place with a constant mass of the air in the chambers. The system of equations, describing thermodynamic processes in the pneumatic drive chambers at retardation, has the following form: ⎧ 2 ⎪m d x = p ⋅ F − P − p ⋅ F n p п ш 0 2 ⎪ ⎪ n ⎛ x01 + xтп ⎞ ⎪ ⎪ p р = ⎜ x + x ⎟ ⋅ p pпт ⎝ 02 ⎠ ⎪ n ⎪ 1 1 ⎛ ⎛ ⎞⎞ ⎪ s + x02 − x )( pшт ) k + h ⎜ ( pak ) n ⎟ ⎟ ( ⎜ ⎪ ⎝ ⎠⎟ ⎨p = ⎜ ⎜ ⎟ + + − s x h x ( ) 02 ⎪ ⎜ ⎟ ⎪ ⎝ ⎠ ⎪ n −1 ⎪ n ⎪T = ⎛ p pтп ⎞ ⋅ T ⎜ ⎟ р рпт ⎪ ⎜ pp ⎟ ⎝ ⎠ ⎪ ⎪ Tb ⋅ Tak )) ⎪T = p ⋅ (( s + x02 + h − x ) ( s x x ) T + − ak ⋅ pb + h ⋅ Tb ⋅ pak 02 ⎩

(1.1) (1.2)

(1.3)

ð1Þ

(1.4) (1.5)

where mn—reduced mass of the pneumatic engine and working body rotating parts, kg; Fn, Fш—pneumatic engine piston and stem net areas, correspondingly, m2; pa, pp, pb—absolute pressures in the atmosphere, operating and exhaust chambers of the pneumatic drive, Pa; paк—absolute initial pressure in the additional volume, Pa; p—absolute current pressure in the additional volume, connected to the exhaust chamber of the pneumatic engine, Pa; x—current displacement of the pneumatic engine piston, m; s—piston maximum operating stroke, m; x01, x02—‘passive volume’ values reduced to the operating areas of the piston and stem parts of the pneumatic engine piston, m; t—displacement time, c; P0—friction force, H; n— polytropic parameter, for Eqs. (1.2) and (1.4) n = 1+(pa/pm (k − 1)/(pa/pm)), for Eqs. (1.3) and (1.5) n = 1+(pa/pm(k − 1)/(pa/pb)); Tp, Tb,—absolute air pressure in the operating and exhaust chambers of the pneumatic engine, correspondingly, К; Taк, T—absolute air temperature in the additional volume before and after the connection to the exhaust chamber of the pneumatic engine, correspondingly, К; Tpтn—initial temperature value in the pneumatic engine jetting chamber, К; pbmn, ppтn—initial pressures at retardation in the exhaust and jetting chambers, correspondingly, Pa; xтn—coordinate of the control valve switchover for retardation, m; h—geometrical volume of the additional volume reduced to the pneumatic cylinder stem area, m.

1330

A. N. Sirotenko et al.

Equations (1.2) and (1.3) describe the changes in the compressed air pressure in the piston and stem chambers of the pneumatic chamber while Eqs. (1.4) and (1.5) —temperature changes. A simultaneous solution of the known [9] and represented equation system (1) allows describing the pneumatic drive dynamic processes at all piston motion stages. For a simultaneous solution of the equations, the authors used the Runge– Kutta method of numerical integration with an automatic stroke selection [17–20]. The mathematical model validity is confirmed by the experiments [21] and by Fisher’s criterion [22]. To confirm the efficiency of the suggested method, the authors applied the mathematical model to conduct a comparative simulation experiment. The initial parameters of a pneumatic drive for all retardation methods were taken as similar ones except for the initial parameters of the additional volume and the coordinate of the switchover to retardation. The pneumatic drive had the following initial parameters: piston net area of the pneumatic engine— 85.4  10−5 m2; stem net area of the pneumatic engine—42  10−5 m2, passive volume of the jetting chamber, reduced to a piston net area—0.2 m; ‘passive’ volume of the pneumatic engine exhaust chamber reduced to a stem chamber net area—0.18 m; maximum piston stroke of the pneumatic engine—1 m; ambient working environment temperature—340 К; operating medium temperature in the main—340 К; absolute pressure of the operating medium in the main— 7.3  105 Pa; absolute pressure of the ambient working environment—105 Pa; reduced mass of the transferred cargo—34 kg; operating stroke length—0.934 m. The obtained dynamic dependences are presented in Fig. 3. At the graphs, the characteristics of each method are presented by a certain line type, designated by numbers (Fig. 3d): 1—throttling method; 2—reverse pressure method with the recuperation in the additional volume 0.4 m3 and the initial pressure of 0.1  105 Pa; 3—reverse pressure method with the recuperation in the additional volume of 0.4 m3 with the initial pressure of 7.3  105 Pa; 4—reverse pressure method with the recuperation into the additional volume of 0.1 m3 and the initial pressure 0.1  105 Pa; 5—reverse pressure method with the recuperation in the additional volume of 0.1 m3 and the initial pressure of 7.3  105 Pa; 6—reverse pressure method with the recuperation in a ‘passive’ volume of the pneumatic engine. The research results are implemented at the modernization of the pneumatic drive of the rotation platform of a weighing and packing machine (Fig. 1a, b). The replacement of the pneumatic drive table rotation (Fig. 1a) by a pneumatic drive with energy recuperation (Fig. 1b) allowed controlling the table stop by changing the additional volume parameters. At the same time, with the production performance preservation, the design of such pneumatic drive has simplified while the consumable rate decreased from 0.46 m3/h to 0.13 m3/h.

Research of Pneumodrive with Energy Recovery …

1331

Fig. 3 Dynamic characteristics of the pneumatic drive retardation methods: (a) dependence of the output element motion on the time; (b) dependence of the output element rate on time; (c) dependence of pressure in the pneumatic engine jetting chamber on time; (d) dependence of pressure in the brake chamber on time

4 Conclusion The analysis of the obtained outcomes demonstrated that the pneumatic drive with a reverse pressure retardation in a ‘passive’ volume is less cost-efficient comparing with the suggested pneumatic drive with energy recuperation in the additional volume: by 1.12 times for the additional volume parameters 42  10−6 m3 and the initial pressure inside it of 7.3  105 Pa; by 1,47 times—for the additional volume value of 168  10−6 m3 and the initial pressure inside it of 105 Pa. The pneumatic drive fast response with the reverse pressure retardation in a ‘passive’ volume is by 2 and 6% higher than the suggested pneumatic drive model for the indicated

1332

A. N. Sirotenko et al.

parameters of additional volume. The least energy-intensive method is retardation with the recuperation in the additional volume with the parameters 40  10−6 m3 and initial pressure of 105 Pa.

References 1. Yusop MYM (2006) Energy saving for pneumatic actuation using dynamic model prediction. Dissertation, school of engineering, Cardiff University, 212p 2. Blagojevic Vladislav, Šešlija Dragan, Stojiljkovic Miodrag (2011) Cost effectiveness of restoring energy in execution part of pneumatic system. J Sci Ind Res 70:170–176 3. Diachenko AD, Udovkin AI, Sirotenko AN (2006) Improving characteristics of pneumatic drive of unit for liquid product pre-packing and packing. Improving processes and technical means in agroindustrial complex: 4. Collection of scientific papers. Ministry of Agriculture and Food of the Russian Federation. Azov and Black Sea Region Academy of Agricultural Mechanization, Zernograd, p 79–81 4. Sirotenko AN, Partko SA (2015) Improvement of characteristics of a pneumodrive of table turn of packaging and filling machine. In: 8th international scientific-practical conference materials: condition and prospects of development of agricultural machine-building, Within the framework of the 18th International agricultural industry. Sp. “Interagromash 2015”, Rostov-on-Don, p 179–182 5. Diachenko AG, SavostinaTP, Kolpakov MV (2018) Development of structure of feeding machine for cylindrical parts. In: Innovative materials and technologies: collection of articles on outcomes of international research-to-practice conference, Sterlitamak, p 11–13 6. Filipov IB (1987) Deceleration devices of pneumatic actuator. Machine-building, Leningrad 7. Krytikov G, Strizhak M, Strizhak V (2017) The synthesis of structure and parameters of energy efficient pneumatic actuator. Eastern-Eur J Enterp Technol 1(7):38–44 8. Sirotenko AN, Partko SA (2014) Energy-saving pneumatic actuator of technological equipment. Materials of VI Intern. scientific-practical. In: Conferrence: innovative technologies in machine-building and metallurgy, Rostov-on-Don, p 173–178 9. Hertz EV (1985) Dynamics of pneumatic systems of machines. Machine-building, Moskow 10. Anh Dao The, Sidorenko VS, Dymochkin DD (2014) Dynamics of position fast robot with pneumatic drive of brake unit. Dyn Vibroacoustics Mach Conf 3:176–183 11. Sirotenko AN, Chernavsky VA (1998) Mathematical model of valve hydraulic absorber. Hydraulic pneumatic systems of technological and mobile machines: inter-university collection of scientific papers, Rostov-on-Don, p 28–32 12. Dao TA, Sidorenko VS, Dymochkin DD (2015) Study on positioning accuracy of automated pneumatic drive with an outer brake. Vestnik Don State Tech Univ 15(4):46–53 13. Grishchenko VI, Sidorenko VS, Poleshkin MS (2009) Pneumatic position drive. Patent RUS for invention 2450174 14. Grishchenko VI, Kilina MS, Chernavskiy VA (2012) Positioning dynamics of drive gears with hydroabsorber. Vestnik Don State Tech Univ 12(4):16–21 15. Poleshkin MS, Al-Kudakh AM, Grishchenko VI, Sidorenko VS (2008) Identification of operation processes in multifunctional braking device. In: Hydraulic machines, hydraulic drives and hydraulic pneumatic automation: proceedings of reports of the 12th international science and technology conference, Moscow, p 54–55

Research of Pneumodrive with Energy Recovery …

1333

16. Grishchenko VI, Al-Kudakh AM, Poleshkin MS, Sidorenko VS (2008) Structural and parametric control of position pneumatic hydraulic mechanical device. In: Hydraulic machines, hydraulic drives and hydraulic pneumatic automation: proceedings of reports of the 12th international science and technology conference of students and post-graduate students. Ministry of Education and Science of the Russian Federation, Moscow Power Engineering Institute (Technical University), Bauman Moscow State Technical University, Moscow, p 22–23 17. Sirotenko AN, Partko SA (2017) Decrease in power inputs in pneumodrive weighing-and-packing machine. Int J Appl Eng Res 12(14):4599–4603 18. Sirotenko AN, Partko SA (2017) Mathematical model of dynamic processes of pneumatic drive during braking by reverse pressure, with recovery of energy into additional volume. Sci Rev 1:67–74 19. Sirotenko AN, Partko SA (2018) Calculation of dynamic parameters of pneumatic hydraulic drive with energy recovery. Certificate of registration of the computer program, RUS 2018613130 RF 20. Sirotenko AN, Partko SA (2018) Calculation of dynamic characteristics of pneumatic-mechanical drive. Certificate of registration of the computer program, RUS 2018663925 RF 21. Udovkin AI, Sirotenko AN (2009) Experimental check of validity of mathematical model of cost-efficient pneumatic drive of the unit “Alur-1500”, Improving processes and technical means in agroindustrial complex: collection of scientific Papers– Zernograd/ FGOU VPO ACHGAA 8:50–52 22. Sirotenko AN, Partko SA, Saed BA (2017) Dependence of energy-speed characteristics of pneumatic drive on initial parameters of additional volume under counterpressure braking. Vestnik Don State Tech Univ 17(4):69–76

Model of Airflow Process Through Throttling Sections of Automated Deadweight Absolute Pressure Measurement System A. Markov

Abstract The aim of the work is to develop a mathematical flow of air with constant pressure drops through the throttling areas of the non-compacted piston of the automated cargo-piston absolute pressure measurement system. Research methods include the theory of automatic control and simulation of systems, as well as the basic laws and regulations of gas dynamics. The results of theoretical studies of the processes of airflow with constant pressure drops through the throttling areas of the non-compacted piston are presented in a mathematical model, the main parameters of gas-dynamic processes occurring in a closed volume, in which the absolute air pressure is set. The connection between the value of the pressure drop and the airflow through the throttling areas of the non-compacted piston is established. The proposed mathematical model allows conducting theoretical studies and computer experiments, as a result of which the optimal values of the constant pressure drop can be selected, allowing to provide the necessary dynamic and precision characteristics of the automated cargo-piston system of absolute pressure measurement. The mathematical model is brought to the calculated level and can be used to solve the problems of designing automated cargo-piston systems of absolute pressure measurement as a precision tool for quality control.










Keywords Pressure measurement Pressure setting Automated control Quality control Unpacked piston Pressure sensor







1 Introduction Aviation technology development, increasing requirements for metrological characteristics of monitoring air-speed flight parameters devices, automated information, and measurement systems use require modernization of quality control A. Markov (&) Baltic State Technical University “VOENMEH” Named After D. F. Ustinov, 1, Krasnoarmeyski, Saint Petersburg 190005, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_141

1335

1336

A. Markov

processes of absolute pressure sensors [1, 2]. Requirements for air pressure measurement equipment are changing [3–5]. In connection with this, automated testing complexes are being created based on different physical principles. However, only automated testing and calibration complexes based on a deadweight pressure gauge are able to implement the fundamental pressure measurement principle, allowing the absolute pressure of a gas to be reproduced with an accuracy of 20 Pa in the range from 0.7 to 100 kPa and 0.01% in the range from 100 up to 285 kPa. The automated deadweight absolute pressure measurement system (ADPMS) meets these requirements. The ADPMS sets the air pressure in a closed volume connected to a pressure sensor, which is controlled by the control program. ADPMS is built on the basis of an automatic pressure control device (APCD), using the principle of a deadweight pressure gauge. The process of quality control consists of comparing the output signals of the controlled sensor and ADPMS [6–8]. For the practical implementation of ADPMS, it is necessary to conduct theoretical and experimental studies on the airflow through throttling elements and in closed volumes, as well as it is necessary to develop a mathematical model of airflow through the annular gaps of the piston pair. Thus, the purpose of the article is to develop a mathematical model for calculating the air consumption through the throttling sections of the unpacked piston of APCD, built on the basis of the deadweight pressure gauge.

2 Design Features of Automatic Pressure Control Device Dynamic and accuracy characteristics of ADPMS are determined by the structure of the APCD. The APCD should contain an automatic pressure control system that sets and stabilizes the absolute air pressure in a closed volume, determined by the weight of the load on the load platform of the unpacked piston; differential unpacked piston with lever balance [9, 10], which allows setting both absolute and excessive pressure; a mechanism based on an unpacked piston is used as a regulator and a sensing element. Thus, a design feature of APCD is implementing the principle of direct control, in which the unpacked piston is the measuring and actuating element of the automatic pressure control system. The nature of the transient process of setting the pressure at the selected design parameters of the elements of the pressure control circuit is determined by the volume of the receiver and the damping properties of the lubrication fluid filling the gaps between the piston and the cylinder. Selection of these factors ensures the setting of pressure in the working volume according to law, which is close to aperiodic. Dynamic characteristics are mainly determined by dimensions of the closed volume, where the absolute air pressure is set; damping properties of the lubricant filling the gap between the piston and the cylinder in the unpacked piston; inertia of moving parts of the APCD. Optimization of these factors, as a rule, allows ensuring the transient process of setting the air pressure in a closed volume close to aperiodic. However, with a large range of

Model of Airflow Process Through Throttling Sections …

1337

setting pressures, the difference in air consumption rates in the throttling section of the unpacked piston during pumping and air draining reaches a large value, which in many respects prevents the transient process from being aperiodic. To solve this problem, it is proposed to use special devices in the design of APCD—pneumatic regulators of constant pressure differential. They provide airflow through the throttling sections of the unpacked piston with constant pressure drops over the entire range of specified air pressures. The structure of the pneumatic regulator must contain resilient sensor (diaphragm) loaded by a driver spring, dividing the regulator into controlling and working cavities and having the possibility of displacement from its equilibrium position; input channel; and valve device.

3 Consumption Characteristics of the Unpacked Piston During the Airflow with Continuous Pressure Differentials Designing APCD is impossible without a mathematical model of the process of airflow with continuous pressure differentials in a closed volume (receiver). When developing a mathematical model, simplifying assumptions were made: The change in air density due to pressure changes was not taken into account for a receiver in which air flows at low-pressure differentials; the flow characteristics of the unpacked piston at unsteady operating modes were considered the same as under static conditions (condition of quasi-stationary flow); the pressure did not change from one point of the receiver to another; i.e., average pressure for the receiver was taken; airflow analysis was carried out for limiting conditions (the process of changing the air condition is isothermal or adiabatic); linearized characteristics of throttling sections of the unpacked piston were used. The task of regulating (setting and stabilizing) the pressure px in a closed volume Vx is solved with the help of an unpacked piston UP with throttling sections D1 and D2 (Fig. 1). With a sufficiently large range of pressure change, p2 < px < p1 (p1 is supply pressure, p2 is discharge pressure, Uctrl is the control signal set by the weights placed on the load platform of the unpacked piston) the regulating process gets more difficult since pressure differentials d1 = p1 − px and d2 = px − p2 at the throttling sections D1 and D2 at the ends of the range become too large. The use of pressure differentials regulators P1 and P2 which provide constant pressure differentials d1 and d2 in throttling sections D1 and D2 relatively to px regardless of its size in the design of regulators improves pressure control process [11]. Consumption characteristics of the throttling sections of the unpacked piston were determined considering the flow regime to be turbulent. The flow of gas through an opening (channel) in a partition separating infinitely large cavities with different pressures was accepted as a calculation model of throttling sections of the

1338

A. Markov

Fig. 1 Structural model of pneumatic part of the automatic pressure control device

unpacked piston. This made it possible not to take into account the velocity of the gas flow in the inlet and outlet channels. In the case of a turbulent regime, different relations are obtained for subcritical and supercritical flows. In the particular case of isothermal outflow of air through an opening in a thin side, the following approximated formulas [12–14] can be used to calculate the air consumption: • for subcritical mode

Gsub ¼ k

pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 00 p00 ðp0  p00 Þ; p [ 0:5p0

ð1Þ

• for supercritical mode Gsup ¼ 0:5kp0 ; p00  0:5p0 rffiffiffiffiffiffi 2g k ¼ mF ; RT

ð2Þ

where m—consumption coefficient, which considers the discrepancy between the actual consumption rate and the consumption rate calculated using these formulas at m = 1, F—open flow area of the unpacked piston’s channel, R—gas constant, T— absolute temperature, p′—pressure in front of the side, p″—pressure behind the side. The ratio p2/p1 is a criterion for the existence of a particular flow regime. If compressed air is used as an actuating medium, and this ratio is greater than 0.5, then the flow is considered subcritical. If the ratio is p2/p1  0.5, then the flow is supercritical. We will calculate the air consumption through the throttling areas of the unpacked piston, taking into account the notation introduced in Fig. 1, a when filling (G011 ) and emptying (G012 ) the working volume:

Model of Airflow Process Through Throttling Sections …

G011 G012

 pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi k px ðp1  px Þ; px [ 0:5p1 ¼ 0:5kp1 ; px  0:5p1  pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi k p2 ðpx  p2 Þ; p2 [ 0:5px ¼ 0:5kpx ; p2  0:5px

1339

ð3Þ ð4Þ

When pneumatic regulators are used in APCD which maintain constant pressure differentials d1 = p1 − px and d2 = px − p2 (hereinafter d = d1 = d2). In accordance with this, the weight air consumption rate during filling (G0011 ) and emptying (G0012 ) the working volume is calculated using the following formulas:  pffiffiffiffiffiffiffi k px d ; px [ d 0:5kðpx  dÞ; px  d

ð5Þ

 pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi k ðpx  dÞd ; px [ 2d 0:5kpx ; px  2d

ð6Þ

G0011 ¼ G0012 ¼

We will use (3)–(6) to make graphs of consumption characteristics at p2 = 0, p1 = 8d, d = d1 = d2, k = k1 = k2. Figure 2 a shows the characteristics G011 ¼ G011 (px) and G012 ¼ G012 (px) for consumption rate of filling and emptying working volume in the absence of a pneumatic regulator of continuous pressure differential. Figure 2b shows the characteristics G0011 ¼ G0011 (px) and G0012 ¼ G0012 (px) for the consumption rate of filling and emptying working volume in the presence of a pneumatic regulator of continuous pressure differential.

Fig. 2 (a) Flow characteristics of the throttle in the absence of a pneumatic regulator (b) Flow characteristics of the throttle in the presence of a pneumatic regulator. 1—unconsolidated differential piston; 2—lower cylinder; 3—upper cylinder; 4—load plate; 5—goods; 6—engine; 7, 8—sleeve; 9—pedestrian device; 10, 11—lubricator; 12, 13—oil channels; 14—recess; 15, 17— undercut; 16, 18, 19, 20, 22—air channels; 21—camera feedback; 23—the working volume; 24, 25—pneumotachometry constant differential pressure

1340

A. Markov

4 Functional Model of the Automatic Pressure Control Device Figure 3 shows the functional model of the APCD. Deadweight measuring and regulating column of the device consist of a combined lower unit—cylinder 2, upper unit—cylinder 3, and the unpacked differential piston 1 rotating in them with a load platform 4. The piston is also assembled from units; it consists of three coaxial, hinged-related direct pistons which create a measuring chamber A and abutment pressure chamber B. The assembled piston design is applied to simplify its manufacturing technology. Loads 5 are put on the load platform. These loads are calibrated in proportion to the given area of the piston. This loading method allows setting different discrete pressure values in the arbitrary sequence. To reduce mechanical friction between the piston and the cylinders, the piston is driven to rotate by the kinematic coupling of the drive motor 6 through a gear reducer, bushings 7 and 8, the driving device 9 with the load platform 4 rigidly bound to the piston. Driving device 9 provides free movement of the piston along

Fig. 3 Functional scheme of the automatic pressure control device

Model of Airflow Process Through Throttling Sections …

1341

the vertical axis. The friction in the driving device is so small that it can be not considered when assessing the error in the operation of the device. To balance the mass of the piston with the load platform, a balancing lever device is used with the placement of rater loads on it. In addition, the balancing mechanism is designed to provide a zero reference. To eliminate “dry” friction lubrication is injected into the annular gap between the piston and cylinders 2 and 3 of the column using: oiler 10, in which the oil is under atmospheric pressure; oiler 11 with oil under supply pressure; systems of oil channels 12 with an external oil line, providing lubrication in the upper and lower cylinders of the column under atmospheric pressure. The automatic control system is located in the lower part of the piston pair and includes: (1) Control throttle, consisting of: • Grooves 14 on the piston; • Two recesses 15 and 17 on the inner cylinder of the column, which, through channels 16 and 18, are connected to the vacuum pump Pvs (vacuum source) and to the source of pressure Gsrc (power source), respectively; • Input channel 20. (2) Feedback chamber A with pneumatic valve and channel 22; (3) Receiver 23, to which controlled pressure sensors are connected. The universality of the APCD and the ability to use it as a setter of both excess and absolute pressure is ensured by the presence of a piston chamber of the abutment pressure B. When setting excess pressure, the output channel 19 of the abutment pressure chamber is connected to the atmosphere. To ensure absolute pressure setting the over-piston chamber of the abutment pressure is connected to a separate fore-vacuum pump through the channel 19, and thus, the circuit is prepared for the absolute pressure setting. The operation of the automatic pressure control system proceeds as follows. When applying a load 5 to the load platform 4 of the piston 1, the piston is lowered onto the lower stop. In this position, the pressure from the compressor enters the feedback chamber 21 through the channel 18, recess 17, groove 14, and the pipeline 20 and then through the pipeline 22 enters the working volume 23. As the pressure increases in the working volume 23, the supporting force in the feedback chamber 21 increases, and at the moment when the pressure force in this chamber becomes approximately equal to the weight of the loads, the piston will begin to rise upward to the equilibrium position and block the recess 17, which is connected with a pressure source by channel 18. If at the moment of arrival of the piston to equilibrium, the pressure in volumes 21 and 23 is equal to the pressure specified calibrated load, and the speed of the piston is zero, then the control process will end. If the pressure in these volumes is higher than the pressure specified calibrated load, then the piston will pass the equilibrium position, release pressure through the opened valve 15 and return to the equilibrium position.

1342

A. Markov

Pneumatic controllers 24 and 25 are used in APCD design to set constant pressure differentials of high and low (vacuum) pneumatic power on the grooves respectively, 17 and 15 of piston pair cylinders, which together with the groove 14 on the piston form the regulator of the automatic pressure control system.

5 Dynamics Equation of the Automatic Pressure Regulation System Operation principle of the automatic pressure control system (APCS) is illustrated in the functional diagram (Fig. 3). Equation of translational motion of the piston can be written as follows: mx

d2 x X ¼ Fi ; dt2

ð7Þ

where mx —reduced mass of moving parts of APCD; x—piston displacement; P Fi —sum of all forces applied to the piston in the equilibrium position of the system. The value of the reduced mass can be determined by the formula: mx ¼ m0 þ

Qx ; g

ð8Þ

where m0 —reduced mass without mass of loads; Qx —loads weight; g—gravitational acceleration. The sum of all forces applied to the piston in the equilibrium position will be equal to: X

Fi ¼ Qx þ q þ P0
S02 þ Pb
S01  Nx  Py
S002  Pb
S001  Fff ;

ð9Þ

where q—piston weight; P0 —absolute pressure in chamber B; Pb —barometric pressure; Py —absolute pressure in chamber A; S01 —effective area of the upper end of the piston; S001 —effective area of the lower end of the piston; S02 —effective area of the differential part of the piston in chamber B; S002 —effective area of the differential part of the piston in chamber A; Nx —the reaction of the lever counterweight; Fff — fluid friction force. Fluid friction force can be calculated by equation: Fff ¼ m where m—viscosity coefficient.

dx ; dt

ð10Þ

Model of Airflow Process Through Throttling Sections …

1343

The reaction of the lever at small angles of deflection can be written as: Nx ¼ N þ Cx;

ð11Þ

where N—counterweight reaction d at x = 0; C—gravitational spring rate of the lever. In the selected equilibrium position, when adjusting the APCD, it is possible to ensure equality: q ¼ N þ Cx0 :

ð12Þ

q  Nx ¼ C ðx0  xÞ:

ð13Þ

Then

When developing the manufacturing technology of piston APCD, it is necessary to ensure that the following conditions are met: 

S01 ¼ S001 ¼ S1 S02 ¼ S002 ¼ S2

ð14Þ

When APCD is operating in absolute pressure setting mode, air is being evacuated from chamber B, i.e. P0 ¼ 0. Then, taking into account, the fulfillment of (11)–(14) and under condition 8 < Qx ¼ Qxi ¼ const m ¼ mxi ¼ const ð15Þ : x P0 ¼ 0 Equation (7) will take the following form: mxi

d2 x dx þ m þ Cx ¼ Qxi  Py S2 þ Cx0 2 dt dt

ð16Þ

When X ¼ Xi ¼ const and Py ¼ Pyi ¼ const, we get the statics equation: Pyi S2 ¼ Qxi  C ðxi  x0 Þ

ð17Þ

The change and stabilization of pressure at the level defined by (17) are due to the change in consumption throttling sections of pipelines 18 and 16. We will use the following designations: G1 and G2 —weighting air consumption per second from channel 18 to groove 14 and from grove 14 to channel 16; G3 — weighting air consumption per second through throttle 22. Equation of pressure change Px and Py can be obtained by considering the conditions of filling (emptying) volumes Vx and Vy through throttles of variable cross section of the regulating element.

1344

A. Markov

We will use the gas law for the volume V: PV ¼ q11 RT;

ð18Þ

where P—absolute pressure; V—volume; q11 —weighting gas amount in volume; R—universal gas constant; T—absolute temperature. We will differentiate both parts of (18) by time: dq11 V dP
: ¼ RT dt dt

ð19Þ

Assuming that processes of air outflow through the throttles are quasi-stationary, which is permissible with relatively slow movement of the inertial mass of the piston, we can write: dq11 X ¼ G; dt

ð20Þ

P where G —the algebraic sum of the weighting consumption through the throttles into a given volume. Substituting (20) into (19) for chamber A, we obtain the equation of change: Vy dPy
¼ G1  G2  G3 : RT dt

ð21Þ

where G1, G2, and G3—weighting air consumption through throttles 25, 24, and 23, respectively. Similarly, we obtain the equation of pressure change for the volume of the chamber Vx , where the absolute air pressure is set and where the controlled pressure sensors are connected: Vx dPx
¼ G3 : RT dt

ð22Þ

In compiling (21) and (22), the pneumatic resistance of the pipeline section which connects groove 14 to chamber A was neglected, assuming that the volume Vy also includes the volume of groove 14 and the pipeline volume. We also assume that Vy ¼ const, i.e., piston displacement does not significantly affect the value of Vy . These assumptions should be ensured by the design of the APCD. The resistance of the pipeline section which connects chambers A and 23 is considered, since at small values of Vy and large values of Vx significant outpace in pressure setting in volume A can occur even with a small pneumatic resistance of the throttle 22 in case of the adopted scheme of connecting the volumes. When deriving the above equations, the process of pressure changing in volumes Vx and Vy was assumed to be isothermal. It can be shown that with small relative changes in pressure DP=P in the volume V for a real polytropic process with

Model of Airflow Process Through Throttling Sections …

1345

V polytropic coefficient 1 < n < 1.4, coefficients s ¼ RT change in (21) and (22), specifically:

s¼

s : n

ð23Þ

Consumption characteristics of throttles G1 , G2 , G3 may be related to piston displacement and pressure magnitudes on inputs and outputs of throttles, i.e. 8   < G1 ¼ G1 P1 ; Py ; X  G ¼ G2 Py ; P2; X : 2 G3 ¼ G3 Py ; Px

ð24Þ

Then the dynamics of the automatic pressure control system can be described by the following differential equations system: 8 2 > m d x þ m ddxt þ Cx ¼ Qxi  Py S2 þ Cx0 > < xi dt2       dP sy dty ¼ G1 P1 ; Py ; x  G2 Py ; P2 ; x  G3 Py ; Px > >   : dPx sx dt ¼ G3 Py ; Px

ð25Þ

Air consumption through the throttling sections of the unpacked piston during the flow with continuous pressure differentials is calculated using (5) and (6). We will transform (25) to a convenient for modeling in MATLAB environment form, while linearizing airflow models with continuous differentials in the neighborhood of the operating point (Fig. 4). The “unit step” signal was inputted to the model; modeling results are shown in Fig. 5. Changing parameters of the model depending on the design characteristics of APCD and pneumatic controllers, it is possible to investigate the dynamics of setting the absolute air pressure. Wherein, the type of the transient will not change; i.e., it will be aperiodic.

Fig. 4 Linearized model of automatic pressure control device

1346

A. Markov

Fig. 5 Graph of the transient of air pressure setting

6 Conclusion The mathematical model for calculating the air consumption proposed in the article considers the use of pneumatic regulators of continuous pressure differentials in the design of APCD. Analysis of this mathematical model makes it possible to determine that with equal pressure differentials (d = d1 = d2), the weighting air consumption per second during filling is somewhat different from air consumption during emptying the working volume. However, for certain values of d1 and d2, it is possible to achieve a minimum difference in consumption in the operating range. Wherein, as a computer experiment, it is possible to simulate the values of continuous differentials d1 and d2 and to ensure that consumption characteristics are almost identical during pressure regulating, what increases the accuracy and expands specified pressures range. The proposed dependences do not contradict the results obtained by leading experts in this and related fields of knowledge [15–17]. In addition to the simplicity of automating the process of the quality control of pressure sensors, the advantage of the proposed design of an automated deadweight system for the quality control of absolute pressure sensors is that the fundamental principle of pressure measurement is implemented, which ensures sufficiently accurate and reliable measurements. The proposed mathematical model for calculating consumption through the throttling section of the unpacked piston of APCD when air flows with continuous pressure differentials is convenient for calculations and can be used for practical problems solving when designing real systems that can implement modern methods of quality control of pressure sensors.

Model of Airflow Process Through Throttling Sections …

1347

References 1. Markov AV (2014) Problems and ways of absolute pressure sensors quality control systems modernization. Sci J Age Qual 4:30–32 2. Markov AV (2015) Concept of measuring instruments of absolute pressure sensors quality control systems. Sci J Age Qual 1:34–35 3. Pushkov SG, Lovitsky LL, Korsun ON (2018) Aero-synamic errors of the aircraft static pressure measurement systems in sliding modes. Meas Tech 2:37–42 4. Pushkov SG, Gorshkova OYu, Korsun ON (2013) Mathematic models of in-flight measurements of speed and angle of attack in aircraft landing modes. Mechatronics Automotization Control 18:66–70 5. Korsun ON, Nikolaev SV, Pushkov SG (2016) Algorithm for estimating systematic errors in airspeed measurements, angles of attack and slip in flight tests. In: Proceedings of the Russian academy of sciences. The theory of control systems vol 3, pp 118–129 6. Loparev VK, Markov AV, Stepanyan NM, Dryuk VA (2003) Structure of automatic verification stand of air pressure measuring instruments. Information technologies on transport. Collection of research papers, Politechnics, Saint-Petersburg, p 220‒222 7. Markov AV (2006) Problems of metrological assurance for measuring instruments. In: 61th scientific technical conference abstracts dedicated to day of radio, Saint Petersburg Electrotechnical University “LETI”, p 226‒228 8. Loparev VK, Markov AV, Spiridonov EI, Stepanyan NM (2002) Organization of frequency pressure sensor with ratio of error of verifiable and reference instruments. Methods of applied mathematics in transport systems: issue 6, collection of research papers, Saint Petersburg State University of Water Communications, Saint-Petersburg, p 137‒139 9. Markov AV (2018) The concept of precision automated control systems for pressure sensors as a means of metrological support of aircrafts. Questions of defense technology. Series 16: Tech Means Countering Terrorism. 9–10 (123‒124):150‒154 10. Markov AV (2018) The concept of precision automated deadweight piston quality control systems for pressure sensors. Qual Innovation Edu 4:89–93 11. Markov AV (2010) Software-controlled pressure setting system. Patent of the Russian Federation for useful model 107869 12. Markov AV (2017) Development of automated pressure sensors. International scientific journal “Mathematical modeling”, YEAR I ISSUE 1/2017 ISS 2535‒0978, Sofia, Bulgaria, p 31‒32 13. Mirskaya VA, Nazarevich DA, Ibavov NV (2017) Method of pressure measurement on an experimental installation for studying the complex of thermophysical properties of liquids and gases. Meas Tech 9:33–36 14. Mirskaya VA, Ibavov NV, Nazarevich DA (2016) Automated experimental installation for investigating the complex of thermophysical properties of liquids and gases. Thermophys High Temp 54(2):237–242 15. Siraya TN (2018) Methods of data processing for measurements and metrological models. Meas Tech 1:9–14 16. Korovina OA (2018) Assessment of the risks of the manufacturer and the customer when monitoring the errors of measuring devices at one or several points. Measuring Tech 5:14–17 17. Danilevich SB (2015) Reliability of the results of multiparameter measurement control. Control Commun Saf Syst 4:171–179

Investigation of Radial Gas Bearings with Longitudinal Micro-Grooves of Various Transverse Profiles I. V. Vishtak, V. A. Fedotov and A. N. Solomon

Abstract For a radial gas bearing with two lines of compressed gas supply into the working gap with blind micro-grooves of a constant and variable depth of the following profiles: triangular or stepped, with the increasing or decreasing depth of the micro-groove in the direction of movement of the compressed gas—radial lifting force, gas consumption and performance ranges were founded. In the study of the effect of blind micro-grooves of different transverse profiles (seven options) on the characteristics of radial bearings the, authors used the methods of cyclic sweep (grooves of a triangular transverse profile) and the method of splines (stepped grooves). The calculations with optimal design parameters (maximum ratio of the radial lifting force to the gas flow rates) showed that the radial bearing with longitudinal micro-grooves of various transverse profiles, the depth of which decreases in the direction of gas flow, has a significantly greater radial lifting force (by 54% at PH = 5; k = 2; æ = 0.25; a = 0.676) compared with a suspension with the grooves of a constant depth and efficient in the range of practical use of supports, unlike the suspensions with micro-grooves with a minimum depth in the area of compressed gas supply.




Keywords Radial bearing Depth micro-grooves Lifting force Gas consumption





Triangular
Stepped profile


1 Introduction Gas bearings are divided into classes depending on the type of throttling of the gas flow: with an external throttle (slots, small diameter holes, porous bushings) [1–8], with an internal throttle (gas pressure regulation depending on the load structural I. V. Vishtak (&)
V. A. Fedotov Vinnitsa National Technical University, 95, Khmelnitskoe Shosse, Vinnytsia 21021, Ukraine e-mail: [email protected] A. N. Solomon Vinnitsa National Agrarian University, Vinnytsia, Ukraine © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_142

1349

1350

I. V. Vishtak et al.

elements inside the working gap) [9–11], a combination of external and internal throttling of the compressed gas flow [1, 3, 5–7, 10, 11]. Bearings with an external throttle have better characteristics and are more economical compared to suspensions with internal compensation, especially at close to one unit ratios of the support length to its diameter. But they are prone to loss of their characteristics during operation: the clogging of chains of holes (diameters from 0.1 to 0, 6 mm), cracks (10 … 30 microns wide); high sensitivity to changes in the characteristics of compressed gas (the presence of solid fractions, moisture, etc.); technologically difficult to manufacture. The bearings with internal compensation are technological in production (longitudinal micro-grooves on the surface of the gas bearing shaft in spindle units are made in the form of fillets or flats) [11, 12], do not change their characteristics depending on the humidity of the compressed gas. In the Adams bearing [9], the pressure inside the working gap is controlled by a deaf circular stepped working gap. Hirs [9, 11] proposed to improve the characteristics of such a bearing by replacing the circular stepped shape of the shaft with longitudinal micro-grooves of constant depth. The advantage of longitudinal micro-grooves is explained by the fact that when the shaft is displaced under external load, the flow of compressed gas [11, 13] from the overpressure zone (minimum working clearance) to the under pressure zone (maximum operating clearance) decreases.

2 Literature Review and Defining the Problem Thus, in order to improve the characteristics of a bearing with longitudinal micro-grooves, it is necessary to reduce the circular flow of gas by reducing the ratio of the hydraulic resistance to the movement of gas in the axial direction of the circular. This effect is achieved by performing longitudinal micro-grooves of variable depth with different transverse groove profiles [11, 14–16]. In a radial bearing (Fig. 1), gas (a mixture of gases, air) under pressure PH, from an external source (compressor), flows through two lines directly into the working gap, flows through zones with longitudinal micro-grooves of various transverse profiles and zones without grooves (isotropic) and flows into the environment. Basic notation: c—working gap between the shaft and the bearing housing at their coaxial position; e ¼ e=c—is the relative radial eccentricity; k ¼ l0 =R0 —is the relative length of the support; n ¼ z=l0 —is the dimensionless axial coordinate; r; r1 —depth of longitudinal micro-grooves; m ¼ c=ðc þ rÞ, m1 ¼ cðc þ r1 Þ—are the parameter for changing the gap as a result of applying micro-grooves; b ¼ r=ðc þ r1 Þ—is the micro-grooves parameter of variable depth; c ¼ m þ b, c1 ¼ 1  c—with increasing depth of micro-grooves in the direction of gas flow (Fig. 1d, e); c ¼ 1 and c1 ¼ ð1  m  bÞ—in case of decreasing the depth of micro-grooves in the direction of the gas flow (Fig. 1b, c) —working gap on the projections of the longitudinal micro-grooves; hК = chq—working gap in micro-grooves; hm ¼ 1  e cos /, hq ¼ 1  em cos /—dimensionless working

Investigation of Radial Gas Bearings with Longitudinal …

1351

Fig. 1 Radial gas bearing with two gas supply lines and longitudinal micro-grooves of various profiles: (a)—constant depth of micro-grooves; (b) and (c)—the depth of the micro-grooves decreases in the direction of the gas flow; (d) and (e)—the depth of the micro-grooves increases in the direction of the gas flow; (f)—a form with a greater depth of stepped micro-grooves at the entrance to the working gap; (g)—stepped micro-grooves with a smaller depth at the entrance to the working area

1352

I. V. Vishtak et al.

gaps, respectively, on the ledges and micro-grooves; /1 ; /2 —are the central corners of the micro-grooves and protrusions, respectively; æ ¼ /1 =ð/1 þ /2 Þ—is the relative width of the micro-grooves; a1 ¼ l=l0 —is the relative length of the step; a ¼ l1 =l0 —is the relative length of micro-grooves; pa —is the ambient pressure; pH —is the gas pressure from an external source of compressed gas (at the entrance to the working gap of the bearing); PH ¼ pH =pa —is the dimensionless pressure of the gas supplied to the working gap; p ¼ pðn; /Þ—is the gas pressure in the working clearances of the bearing; P ¼ p=pa —dimensionless pressure (Fig. 2); u ¼ P2 —is the square of dimensionless pressure; F ¼ ppa R20 F  —radial lifting force of the bearing; F  —dimensionless radial lifting force of the bearing; pkp2 c3

Q ¼ 12la Q —gas consumption for bearing operation; Q —dimensionless gas flow; k—is the ratio of gas density to pressure at a given temperature in the gas layer; l—is the dynamic coefficient of viscosity of the gas. The characteristics of radial bearings with micro-grooves of constant depth (Fig. 1a) are sufficiently studied in the static mode with their optimal parameters [10, 11] in the range of linear dependence of the radial lifting force F depending on the relative radial eccentricity within 0:5  e  0:5 that is F  ¼ Ke
e. If the longitudinal micro-grooves have a variable depth (Fig. 1b–e) with r1 ¼ 0, then the differential equation of pressure distribution in the working area of the bearing with micro-grooves has the form [11, 16, 17]: @ 2 u1 3 @u1  ðmb2 cos /  c1 v0 Þ @n @n2 b1 hq
 3 2 3 2 m kh @ u1 3b4 e @u1 þ hb3 þ sin / ¼ 0; hq @/ @/2 b2 b23 ð1  vÞ2

ð1Þ

where v ¼ /2 =ð/1 þ /2 Þ, v0 ¼ v=ð1  vÞ, h ¼ hm =hq , hq ¼ c  me cos / þ c1 n, hm ¼ 1  me cos /, b 1 ¼ v0 þ m 3 h3 , b 2 ¼ v0 þ m 2 h2 , b 3 ¼ 1 þ v0 m 3 h3 , 4 4 b4 ¼ 1 þ v0 m h . When v ¼ 1, m ¼ 1, c ¼ 1, c1 ¼ 0, the basic equation for smooth sections of the gas layer is obtained from Eq. (1).
2  @ 2 u2 3 @u 3 @u2 2 @ u2 þ k sin / ¼ 0:  cos / þ e @n hm @/ @n2 hm @/2

Fig. 2 Load distribution scheme

ð2Þ

Investigation of Radial Gas Bearings with Longitudinal …

1353

The solution of differential Eqs. (1) and (2) is carried out by an approximate method using cubic polynomials [11, 16–18]. u1 ¼ A1 þ B1 n þ C1 n2 þ D1 n3 :

Work zone 1ðFig:1beÞ

ð3Þ

u 2 ¼ A 2 þ B 2 n þ C 2 n2 þ D 2 n3 :

Work zone 2 ðFig: 1 beÞ

ð4Þ

where Ai ; Bi ; Ci ; Di (i = 1, 2) are unknown coordinate functions u. The dimensionless squares of pressure ui (i = 1, 2) in the working clearances of the bearings, the dimensionless radial force F  and the dimensionless gas flow rate Q are found by the cyclic sweep method [19, 20]. 0 Fe ¼ k@

Za

Zp dn

0

pffiffiffiffiffi u1 cos / d/ þ

Z1

Zp dn

a

0

1 pffiffiffiffiffi u2 cos / d/A

0

0 2p 1 Z Z2p 1 @u @u 2 @ h3 1 d/ þ Q ¼ d/A h3m m pk @n @n 0

0

Taking into account the study of the effect of the inclination of the longitudinal micro-grooves (Fig. 1a–e) on the characteristics and performance of gas bearings, similar results are achieved by applying an appropriate number of steps of different heights in the zone profiled with micro-grooves. For micro-grooves with a decrease (Fig. 1f) or increase (Fig. 1g) of their depth in the direction of gas flow, the stepped micro-grooves have a maximum depth (Fig. 1f) at the entrance to the working gap or a minimum depth (Fig. 1g) (throttle inlet). The constructive advantage of stepped micro-grooves lies in the fairly reasonable method of calculating such bearings [10, 11, 17]. A decrease in the length of the zone with micro-grooves increases the accuracy of calculations. The squares of gas pressure u1 ; u2 ; u3 in three sections of the working gap of a radial bearing with stepped micro-grooves (Fig. 1f, g) are found according to the procedure [10, 11, 17, 21].   u1 ¼ a01 þ a02 n þ a11 ehkn þ a12 ehkn cos /;   u2 ¼ b01 þ b02 n þ b11 eh1 kn þ b12 eh1 kn cos /;   u3 ¼ d01 þ d02 n þ d11 ekn þ d12 ekn cos /: where h2 ¼ ððv þ m3 ð1vÞmÞðvm3 ð1vÞÞ ; h21 ¼ 3

m3

1 : ððv þ m31 ð1vÞÞð1v þ vm31 Þ The coefficients aij ; bij ; dij —depend on the gas pressure and the structural dimensions of the bearing and micro-grooves. Dimensionless lifting force F  of a radial gas bearing with stepped longitudinal micro-grooves:

1354

I. V. Vishtak et al.

0 F  ¼ k@

Za1

Zp dn

0

pffiffiffiffiffiffi U1 cos /d/ þ

dn a1

0

Zp

Za1

pffiffiffiffiffiffi U2 cos /d/ þ

0

Zp

Z1 dn a

1 pffiffiffiffiffiffi U3 cos /d/A

0

When calculating and optimizing the design parameters of radial bearings and their practical use, the dimensionless gas flow rates Q are at e ¼ 0 [10, 11].   s P2H  1 : Q ¼ kð a2 m 3  s ð 1  aÞ Þ 

3 Objective Investigation of the characteristics and performance of gas bearings with micro-grooves by changing the shape of the grooves (stepped shape).

4 Materials and Research Results If the depth of the micro-grooves of the bearing increases in the direction of gas flow (Fig. 1d, e) at o1 ¼ 0, then its dimensionless radial rigidity is much less (1.75 times at PH ¼ 5; k ¼ 2; æ ¼ 0:25; a ¼ 0:676; b ¼ 0:3) in comparison with a bearing with a back tilt of micro-grooves (Fig. 1b, c). A radial bearing with a maximum depth of micro-grooves on the gas pressure line (Fig. 1b, c) has extremes of dimensionless radial rigidity in terms of the depth of micro-groove b and the parameter of the gap smoothness m. According to the criteria [11] for a micro-grooved bearing, shown in Fig. 1b, c in [22], no static instability zones were detected at 0  b  1 and 0:2  m  1. A radial bearing in which the depth of the longitudinal micro-grooves increases in the direction of gas flow (Fig. 1d, e) loses its static stability at b  0:25 and m  0:35. The bearing (Fig. 1c) has a dimensionless radial lifting force 15% less compared to a micro-grooved bearing shown in Fig. 1, b. Calculations of the radial bearing with stepped micro-grooves (Fig. 1f, g,) showed that if the depth of the longitudinal micro-grooves at the entrance to the working gap (Fig. 1g) is less (at 0:4  a1  a), then the dimensionless radial lifting the force begins to grow only when m  0:35; if the depth of the longitudinal micro-grooves in the direction of gas flow along the working gap decreases (Fig. 1f) [23], then the bearing will be operational even with m  0:17 ðr  4:9cÞ; the bearing with stepped micro-grooves (Fig. 1g) has an instability zone to m ¼ 0:35.

Investigation of Radial Gas Bearings with Longitudinal …

1355

When increasing only the relative length of the step a1, without changing the depth of the longitudinal micro-grooves, the dimensionless radial lift force of the bearing increases; for the profile of micro-grooves in which the depth of the step decreases in the direction of gas flow (Fig. 1f), the working gap is shorter than the longitudinal micro-grooves (Fig. 1g), and then for a1  0:45 the gas bearing is disabled (the dimensionless radial lifting force is zero, the zone of instability). Radial gas bearings with stepped longitudinal micro-grooves (Fig. 1f) with optimal design parameters (maximum ratio K  =Q ) have a large dimensionless radial lifting force (approximately 51.5% at PH ¼ 5; k ¼ 2; æ ¼ 0:25; a ¼ 0:676) compared with a bearing, in which micro-grooves of constant depth and design parameters are also optimal.

5 Conclusions A radial sub-bearing with two lines of compressed gas supply with longitudinal micro-grooves, the depth of which decreases linearly (Fig. 1b, c) in the direction of gas flow, has a radial lifting force (54% at PH ¼ 5; k ¼ 2; æ ¼ 0:25; a ¼ 0:676) compared with a micro-grooved bearing of constant depth (Fig. 1a) and 75% compared with a bearing with a minimum micro-grooved depth (Fig. 1d, e) in the compressed feed zone gas. Bearings with stepped longitudinal micro-grooves with a greater depth at the entrance to the working gap (Fig. 1f), unlike longitudinal micro-grooves with a smaller depth at the entrance (Fig. 1e), are efficient in the range of practical use of bearings and have a large dimensionless radial lifting force and more economical as compared to bearings in which micro-grooves of constant depth. The decrease in radial lift (approximately 4.67% at PH ¼ 5; k ¼ 2; æ ¼ 0:25; a ¼ 0:676) of a bearing with stepped micro-grooves (Fig. 1f) compared with micro-grooves (Fig. 1b) is explained by increasing the accuracy of approximate calculations, since the length of the micro-grooved section (Fig. 1b) is, and in stepped areas—the areas are approximated by cubic polynomials at 0:4  a1  0:669 and 0  a1  0:4. With optimal design parameters, the consumption of compressed gas for the operation of bearings with longitudinal micro-grooves of various transverse at e ¼ 0 (Fig. 1) is almost the same.

References 1. Gressem NS, Pauell JU (1966) Gas lubricated bearings 2. Konstantinesku VN (1968) Gas Lubricated, Mechanical engineering 3. Gross UA (1969) Overview of work in the field of gas bearings with external supercharging for the period since 1959, Friction and lubrication problems, pp 180–185

1356

I. V. Vishtak et al.

4. Dee CW, Shires GL (1971) Current state of development of bearings with feed gaps. Friction and lubrication problems, pp 1–9 5. Sheinberg SA, Ged VP, Shisheev VL (1969) Slide bearings with gas lubrication. Mechanical engineering 6. Pinegin SV, Tabachnikov YB, Sipenkov IE (1982) Static and dynamic characteristics of gas-static bearings, Science 7. Snopov AI, Luchin GA, Danilcheko VF et al (1977) Calculation of radial gas-static bearings of nuclear power turbomachines. PTM 108. 129. 101: 84 8. Adams CR, Dworski J, Shoemaker EM (1961) Externally pressurized Step journal bearings. Trans. ASME, D:125–134 9. Hirs GG (1968) Design bearing bearings with longitudinal grooves and external lubrication injection. Friction and lubrication problems, pp 324–331 10. Emelyinov AV, Fedotov VA, Priyitelchuk VA (1977) Characteristics of radial gas-static supports with double throttling of the gas stream. Mach Sci, 97–104 11. Fedotov VA, Fedotova IV (2010) Gazovye podvesy shpindelnyh uzlov (Gas bearings of spindle unit). Vinnitsa National Technical University, Vinnitsa 12. Fedotov VA, Shevchuk AI (1985) Calculation of gas-static bearings with longitudinal grooves in the form of grooves and fillets. Mech Eng Bull, 21–23 13. Fedotov VA, Shevchenko AV, Molchanov AV (1980) Influence of circumferential gas flows in flow channels of a cylindrical suspension on its characteristics. Mech Eng, 67–71 14. Fedotov VA (1989) On improving the characteristics of gas suspensions with longitudinal grooves. Gas lubrication in machines and devices, All-Union Coordination Meeting, Moscow, pp 45–46 15. Stepanchuk VI, Fedotov VA (1994) Linear static analysis of gas-static tapered bearings profiled with longitudinal grooves. Bulletin of Vinnitsa Polytechnic Institute, pp 57–61 16. Fedotov VA, Savulyak VV (2003) Influence of longitudinal grooves on the characteristics of radial bearings. Problems of Tribology, pp 62–66 17. Vishtak IV, Fedotov VA (2017) Areas of research to improve the characteristics of radial gas bearings with internal compensation. In: Proceedings of the international scientific conference science of the XXI century: problems and prospects of researches, Warsaw, Poland, 17 Aug 2017 18. Emelyinov AV, Shevchuk AI (1983) The method of corrective splines and its application to the theory of gas bearings. Research and application of sliding bearings with gas lubrication. All-Union Coordination Meeting, Vinnitsa, pp 47–48 19. Samarskiy AA (1977) Difference schemes, Science 20. Samarskiy AA, Lazarev RD, Makarov VL (1987) Theory of difference schemes for differential equations with generalized solutions. Graduate School 21. Vishtak IV, Fedotov VA (2016) Influence of stepped longitudinal grooves on the characteristics of radial gas-static bearings. Bulletin of Vinnitsa Polytechnic Institute, pp 110–117 22. Vishtak IV, Fedotov VA (2015) Gas static bearing of the electric machine. UA Patent 102617 U F16c 32/06 HO2K 5/00, 10 Nov 2015 23. Vishtak IV, Fedotov VA (2015) The electric machine on the gas-static bearing. UA Patent 102619 U F16c 32/06 HO2K 5/00, 10 Nov 2015

Flexible Composite Diversion Water Conduits of Small HPPs for Recreational Facilities in the Republic of North Ossetia D. V. Kasharin, S. A. Kalmikov and O. A. Surzhko

Abstract This article is devoted to the substantiation of the application of flexible diversion multi-shell water channels from composite materials for mobile microand small hydropower plants for the recreational zones of the Republic of North Ossetia (Alania). A comparison is made of steel, polyethylene, and fiberglass pipelines of various diameters currently used to manufacture diversion conduits according to weight, cost, and strength characteristics and hydraulic conditions of work and empirical dependencies used later to assess the effectiveness of their use in various sections of the designed conduit are defined. New technical solutions for creating flexible pipelines for the implementation of the derivation are proposed. A comparative analysis of existing technical solutions of shell structures from flexible composite materials as well as calculated positions for determining their optimal parameters while ensuring operational reliability is given. New technical solutions for the flexible diversion water conduit, its advantages, as well as the issues of construction and ensuring reliability in construction under conditions of geological phenomena, were reviewed.










Keywords Diversion conduit Flexible Mobile Hydro Mount Composite material Cable systems Cost











Shell
Support


1 Introduction One of the factors adversely affecting the technical and economic development of the Republic of North Ossetia (Alania), as well as the use of its recreational potential, is the shortage of generating capacities. At the same time, the Republic has one of the highest specific hydropower potential in the North Caucasus, reaching 1800 MW h/km2 and capable of covering the power generation deficit D. V. Kasharin (&)
S. A. Kalmikov
O. A. Surzhko Platov South-Russian State Polytechnic University (NPI), 132, St. Prosvescheniya, Rostov Region, Novocherkassk 346428, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_143

1357

1358

D. V. Kasharin et al.

with about 2000 million kWh. Also, the existing hydroelectric power plants are mainly represented by large stations with the capacity of 854 MW, in the second place are the average of 41, 8 MW and the small ones providing a capacity of 30 MW. Given the significant centralization of the power system with unproductive losses in the transmission of electricity reaching 30%, the main attention should be paid to distributed power generation facilities. The most promising are the conditions for the development of recreational facilities, taking into account the provision of the least negative impact, the diversion micro and small HPPs. One of the main costs and the source of the negative impact is the construction of water intake and water supply structures (diversion conduits). The use of new technologies will ensure a faster payback and increased use in recreational areas by lowering the environmental impact. Consider the conditions for the construction of micro and small hydropower plants in the recreational areas of North Ossetia (Alania). Temperature fluctuation is within a small range from −18 to +40. Radiation balance is 360 MJ/m2. Intense rainfall and solid precipitation can lead to landslides and avalanches, and seismicity reaches 9 points [1, 2].

2 Technological Scheme of Diversion Micro and Small HPPS The most reliable under such conditions will be the use for the diversion of tunnel sections. However, they will be economically unjustified, since underground pipeline laying in rocky and semi-local soils requires significant capital expenditures and is often unacceptable due to the need for additional road infrastructure and heavy construction equipment in conservation areas. The external laying of pipelines is conducted using steel, polyethylene, and fiberglass pipelines to ensure the reliability that is necessary when erecting engineering protection structures and overpasses. The delivery and laying of these pipelines require the use of heavy construction equipment, the work on the relief layout and the installation of various types of supports, compensators to prevent temperature effects and displacements, as well as a device for damping hydraulic impact. We shall consider the main influencing factors for the diversion water conduit (Fig. 1). Characteristics of the water body, depending on its regulation and uniformity of supply, alluvial regime, etc. In case of above-ground laying, it is required to protect the pipeline with heat-insulating materials to prevent freezing of the transported substance at negative air temperatures and heating of the pipe walls under the influence of solar radiation and high air temperatures. The biggest costs for heat-insulating materials are incurred on steel pipelines with the thermal conductivity of the material at 58 in the

Flexible Composite Diversion Water Conduits of Small HPPs …

1359

Impacts on water conduit Hydrometeorological - Characteristics of the water body - Precipitation -Temperature mode - Wind load - Solar radiation - Flooding of the territory

Geological characteristics - Characteristics of the soil - Seismic effects - Dangerous geological phenomena

Technological impact - Static and dynamic pressure - Abrasive wear - hydraulic shock - Other man-made impacts

Fig. 1 Influencing factors for a surface-laid diversion water conduit

second place on the cost of polyethylene 0.36–0.43, and in the third—glass fiber 0.25–0.33 [3–16]. Where, for the steel pipelines, the most important problem is anticorrosive protection, for the polyethylene and fiberglass pipes, solar radiation and ultraviolet rays are subjected to photo-degradation polyethylene. Therefore, benzophene derivatives and soot are used as stabilizers. As can be seen from the above-mentioned types of impacts, water-conducting structures made of materials having the greatest weight and modulus of elasticity, thermal conductivity, and coefficient of hydraulic resistance have the least advantages, which relates to steel pipelines. Consider the comparative characteristics between existing pipelines and traditional materials: cold-rolled steel (Fig. 2a), polyethylene and fiberglass (Fig. 2b, c). The dependence of the cost of steel, plastic and fiberglass pipelines on the diameter and strength characteristics is shown in Fig. 2. For steel pipes, cost dependence on diameter and weight at a pressure of 3.23 MPa. Ss ¼ a0 þ a1
D þ a2
P þ a11
D2 þ a12
P
D

ð1Þ

Fig. 2 a Dependence of cost and weight indicators cold-rolled steel, pressure 3.2 MPa; b dependence of weight indicators on material, pressure ranges and pipeline diameters; c dependence of weight indicators on material, pressure ranges, and pipeline diameters: 1— polyethylene; 2—fiberglass

1360

D. V. Kasharin et al.

Dependence of polyethylene and fiberglass cost on pressure and diameter (Table 1): Spol ¼ a0 þ a1
Pp þ a2
D þ a11
P2p þ a22
D2 þ a12
Pp
D

ð2Þ

Sf ¼ a0 þ a1
Pp þ a2
D þ a11
D2 þ a12
Pp
D

ð3Þ

Dependence of polyethylene and fiberglass weight on pressure, diameter: Pmpol ¼ a0 þ a1
Pp þ a2
D þ a11
P2p þ a22
D2 þ a12
Pp
D

ð4Þ

Pmf ¼ a0 þ a1
Pp þ a2
D þ a11
P2p þ a22
D2 þ a12
Pp
D

ð5Þ

1. Polyethylene hydraulic resistance dependence on diameter (Fig. 3):

Table 1 Coefficient of regression Eq. (1)–(6) a1

a2

a11

a12

Fiberglass dependence of weight on pressure and diameter, Sпэ −29.01 0.013 7.535 9.8  10−5 Polyethylene dependence of weight on pressure and diameter, Scп 88.756 −0.012 −31.481 1.495  10−4 Polyethylene dependence of hydraulic pressure on diameter, k 3.18  10−9 – – −7.334  10−6 The dependence of the cost of diameter and weight for steel pipe pressure 3.2, −214.519 899.107 0.412 3.941 Polyethylene dependence of cost on pressure and diameter, Pm 1.605  104 −6.257 −5.661  103 0.026 Fiberglass dependence of cost on pressure and diameter, Pm 1.182  104 −62.075 −1.197  104 0.023 Polyethylene dependence of hydraulic resistance on diameter, k 9.529  10−10 – – −3.843  10−6

a22 −2.32  10−3 0.05 – Scт (1) −2.719 8.930 51.086 –

Fig. 3 Hydraulic characteristics of the various pipelines are shown in a; b hydraulic resistance dependence on diameter

Flexible Composite Diversion Water Conduits of Small HPPs …

k ¼ a0 þ a1
D þ a2
D 2

1361

ð6Þ

3 New Technical Solutions for Flexible Derivational Conduits When creating new technical solutions for the flexible part of the derivational material, the problem of reducing the cost and weight of the diversionary conduit is solved, and its installation is simplified in difficult terrain without significant planning and alignment along the route. The advantage is the use of local building materials. In view of the new technical solutions, a diversion water pipeline has been developed, including flexible shells made of composite materials. In the course of the experimental studies, the authors determined the working conditions of the flexible diversion water conduit, the following variants of the diversion water conduit are suggested, depending on the conditions of laying of the water conduit (slope of the track, friction on the surface: • Single-shell without a base, used for slopes along the path of less than 10° with a friction coefficient of at least 0.6 • Single-shell with a partition without a base for slopes along the route from 10° to 15° not less than 0.6; • Bi-shell without a base, in which only the upper part of the cladding is secured • Bi-shell with an outer protective cover providing external protection of the waterway. Accounting of elastoplastic deformations allows getting additional weight reduction. Despite some decrease in the capacity to 15%, flexible shells have an order of magnitude less weight and the number of supports. The reduction in capacity is due to the form of the diversion water conduit.

4 Determination of the Initial Shape of the Cross Sections of a Flexible Diversion Water Conduit) The main difference between flexible diversion water conduit and rigid ones is the interrelation of their form with internal and external loads, and boundary conditions. Therefore, the calculation of closed water-retaining shells is made taking into account the change in their shape, depending on: hydrodynamic impact of the flow on the water-filled shell of the LW; conditions of attachment; and characteristics of the shell base [18–22].

1362

D. V. Kasharin et al.

Fig. 4 Design scheme of a water-filled shell

In determining the form and NDS of a flexible diversion water conduit, we introduce the following assumptions: the shell is mounted on a rigid, horizontal base, inextensible, weightless, and filled with an incompressible fluid. Consider the design scheme of a soft water-filled shell (Fig. 4) [2, 17]. We introduce the following notation: W, B—width of the shell and the cladding of the shell to the base; pbot —pressure at the base of the shell; qg—specific weight of water; L—shell perimeter; the relative conformance of the shell to the base and the width can be determined by the following formulae: b=L; w ¼ W=L. To determine the value of the relative fit of the shell, we use the Meok Kim Eq. b [5, 6]:
 w ffi 0; 5p0bot k2 1 þ k 2 p=32  ½k4 ð3p  10Þ=64 þ b

ð7Þ

where k—elliptic integral module; relative pressure at the shell base: p0bot ¼ pbot =ðqgLÞ

ð8Þ

To build the upper and lower parts of the shell, we use the flexible thread equation. When k1  1 the shape of the lower thread will be an elastic of the second kind, and when k1 > 1—elastic of the first kind. For the lower elastic, the parametric equations have the form [2, 3, 7]:

Flexible Composite Diversion Water Conduits of Small HPPs …

1363



qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi y ¼ h 1  1  k12 sin2 u

ð9Þ

  x ¼ h Eðu; kÞ  ð1  k 2 =2ÞFðu; k1 Þ

ð10Þ

where x y—coordinates of the lower thread point, m; k—distance of the most distant lower thread point to the backfill surface, m; Fðu; k1 Þ and Eðu; k1 Þ—elliptic integrals of the first and second kind, respectively; k1 —elliptic integral module
2; N
1 —upper thread tension, kN/m; u—angle (modulus of elasticity) k12 ¼ 4N=h u ¼ a1 =2;a1 —angle between the tangent to the thread and the axis of abscissae, degrees. Values x1 , y1 , a1 , and u at supporting points: 

y1a

qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ¼ h 1  1  k12 sin2 ua1

  x1a ¼ h1 Eðua1 ; k1 Þ  ð1  k12 =2ÞFðua1 ; k1 Þ

ð11Þ ð12Þ

where ua1 ¼ a1a =2. The outline of the lower thread using (11), (12), the parameter u has values within 0  u  ua1 . When k1  1 the shape of the lower thread will be an elastic of the second kind, and at k1 [ 1—an elastic of the first kind, given that for the upper thread, we take the coordinate system x2 O2 y2 , and the parametric equations of the upper elastic have the form [2, 7]: y 2 ¼ h2

qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 



1  k22 sin2 u = 1  k22  1




1  k22 =2 ½Fðp=2; k2 Þ  Fðu; k2 Þ qffiffiffiffiffiffiffiffiffiffiffiffiffi  ½Eðp=2; k2 Þ  Eðu; k2 Þg= 1  k22

x 2 ¼ h2

ð13Þ

ð14Þ

where x2 and y2 —coordinates of the upper thread point, m; h2 —distance of the most distant upper thread point, m; k2 —elliptic integral module (modulus of
2 =h22 þ 4N
2; N
2 —upper thread tension, kN/m; a2 —angle elasticity) k22 ¼ 4N between the tangent to the thread and the axis x2 , degrees; u2a ¼ ðp  a2a Þ=2. Anchor point coordinates: y2a ¼ h2

qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 

1  k22 sin2 ua2 =1  k22  1

ð15Þ

1364

D. V. Kasharin et al.

x2a




¼ h2 1  k22 =2 ½E1 ðp=2; k2 Þ  E 1 ðua2 ; k2 Þ qffiffiffiffiffiffiffiffiffiffiffiffiffi  ½E 2 ðp=2; k2 Þ  E 2 ðua2 ; k2 Þg= 1  k22

ð16Þ

In the general case, when the cross-sectional contour at the reference point has a fracture ua2 6¼ ua1 , we use the (7)–(15) to determine the initial shell shape used for 2D modeling in the Mechanical APDL Ansys module, taking into account deformations and large displacements. To reduce the twisting effects of the flexible DV with a single-shell structure that is being acted upon by transverse circulations, as well as provide sufficient stability and reliability of operation, multi-shell structures are used (Fig. 5a). The water-filled base allows laying a flexible diversion water conduit along a route with a difficult terrain without preparation (Fig. 5), while its smooth outline allows minimizing pressure losses. Here are some characteristics of flexible conduits made of polymeric materials. In the case of the application of material with the 2PVCRX30/EW double-layer reinforcement, the maximum withstandable design head is 150 m (Table 2). Two alternative materials—Mirasol STS 850 (PAD-25) and Unisol 900 are suitable for providing a head up to 70 m in strength characteristics. The weight characteristics are given in Table 2. To study the head losses, hydraulic tests were carried out on a straight section of a single-shell and double-shell diversion water conduit with a diameter of 910 mm. When compared with the head loss, determined by the readings of manometers, the capacity of the water conduit is reduced to 12%. Compared to polyethylene and fiberglass pipes, head losses are increased due to changes in the initial sections of the flexible section. However, even at a pressure of 10 m, the cross section of a single-shell diversion water conduit becomes close to a circular cross section, and the reduction in throughput does not exceed 2%.

Fig. 5 Flexible diversion water conduit on a water-filled foundation: a multiple-shell diversion conduit; b water-filled base partially filled (with the maximum pressure valve); c pressure gauges for determining the pressure in the water channel and the water-filled base:: 1—filling shell; 2— inner shell of the water conduit; 3—outer shell of the water conduit

Flexible Composite Diversion Water Conduits of Small HPPs …

1365

Table 2 Standard sections of a flexible diversion water conduit made on 60 m Name of material

Diameter, mm

Single-shell Mirasol STS 850(PAD-25) 910 Unisol 900 2PVCRX30/EW Double-shell with an outer coating (Fig. 4) Mirasol STS 850(PAD-25) 1220 Unisol 900 2PVCRX30/EW

Section weight, kg 145 154 420 360 370 1321

Fig. 6 Polymer supports and connections of a flexible section of a diversion water conduit: a— intermediate support of a flexible diversion water conduit; b and c—connection between the sections of single- and double-shell water conduit

We have developed and patented new technical solutions for fixing the water conduit made of polymer materials without a waterproof shell with significant slopes, as well as overpasses on the cable suspension, allowing to ensure the laying of the water conduit in conditions of a complex relief through obstacles when it is vertically and horizontally fixed due to the guys of braces. (Fig. 6a). A special feature of constructive solutions is the use of perforators for the installation of chemical anchors, sufficient for fastening flexible sections of a diversion water conduit. Calculation of anchor supports is carried out on the action of axial, transverse, and vertical forces applied to them from the left and right side. The anchor supports installed at the fracture points of the pipeline axis perceive the centrifugal force arising at the turn of the flow and two moments, on the upper edge of the uniformly distributed load in the lower (zero) span (moment on the anchor support), and also on the lower face of the support from uniformly distributed load. At pressures of more than 150–300 m, the distribution of water conduits from polyethylene and fiberglass materials is considered, higher than steel pipelines on the lower sections.

1366

D. V. Kasharin et al.

5 Conclusion 1. Analysis of the natural and climatic conditions of the hydropower and recreational potential of the Republic of North Ossetia (Alania) based on which the use of micro and small diversion hydroelectric power stations is justified. 2. The performed comparisons of technical characteristics revealed the preeminent use of fiberglass pipelines at pressures of up to 320 m as compared to polyethylene and steel pipes. 3. The technical solutions of flexible sections of diversion water conduits are considered, the conditions for their location are identified, and the throughput, which is insignificantly reduced compared with polyethylene and glass fiber pipelines, is estimated, which makes it possible to use them on the upland sections (up to 150 m of the height difference) of the diversion water conduit. 4. The developed technical solutions of supports and overpasses allow the installation of a diversion water conduit without the use of heavy construction equipment. The results of the work were obtained with the support of the Ministry of Education and Science of the Russian Federation within the framework of the state task for conducting research work, Code No. 13.1236.2017/4.6.

References 1. Konovalova OE, Ivanova EA (2013) Small hydro: the problems, difficulties and ways of their solution. Transactions Kola Sci Centre Russ Acad Sci 5:64–75 2. Kasharin DV (2007) Mobile hydropower structure of composite materials. Reclamation Water Manage 4:45–46 3. Caleyo F, Alfonso L, Alcántara J, Hallen JM (2008) On the estimation of failure rates of multiple pipeline systems. J Press Vessel Technol 130 4. Matveev EA, Frolov AB (2008) Final deflections of circular cylindrical shells before loss of stability under the action of uniform external pressure. Izvestiya MSTU MAMI 2(6):152–157 5. Rowe RK, Rimal S, Sangam H (2009) Ageing of HDPE geomembrane exposed to air, water and leachate at different temperatures. Geotext Geomembr 27:137–151 6. Zhang M, Fu S, Song LJ et al (2018) Hydrodynamics of flexible pipe with staggered buoyancy elements undergoing vortex-induced vibrations. J Offshore Mech Arctic Eng 140 (6). https://doi.org/10.1115/1.4040509 7. Leventeli1 Y, Yılmazer I, Ertunc A (2013) Turkey pressure pipeline system to produce hydropower:ecemis-turkey. In: International conference on renewable energy research and applications (ICRERA), pp 46–51. https://doi.org/10.1109/icrera.2013.6749724 8. Kasharin DV (2016) Numerical simulation of the structural elements of a mobile micro-hydroelectric power plant of derivative type. Ser Adv Intell Syst Comput 423:51–61. https://doi.org/10.1007/978-3-319-27644-1_6 9. Leroy JM, Estrier P (2001) Calculation of stresses and slips in helical layers of dynamically bent flexible pipes. Oil Gas Sci Technol Rev IFP 56(6):545–554 10. Gavryushin SS (2014) Numerical modeling of processes of nonlinear deformation of thin elastic shells. Mat Model Number Meth 1:115–130. https://doi.org/10.18698/2309-36842014-1-115130

Flexible Composite Diversion Water Conduits of Small HPPs …

1367

11. Fedorov VM (2011) Constructive and technological measures for the effective operation of the water supply network of irrigation systems. Sci J KubSAU 66(02):1–10 12. Bazhenov VA, Nightingale NA, Krivenko OP et al (2014) Modeling of nonlinear deformation and loss of stability of elastic inhomogeneous shells. Build Mech Eng Constr Struct 5:14–33 13. Bourguet R, Triantafyllou MS (2015) Vortex-induced vibrations of a flexible cylinder at large inclination angle. Philos Trans A Math Phys Eng Sci 28:373(2033). https://doi.org/10.1098/ rsta.2014.0108 14. Rogalevich VV, Timashev SA (2013) An approximate method for calculating flexible sloping shells of constant and variable thickness. Acad Herald Uralniiproject Raasn 1:71–75 15. Savenkov VN (2011) Calculation of rational parameters of multilayer polymer-metal flexible pipes, Zbirnik Naukovych prac. DonIZT 25:175–179 16. Chen ZS, Kim WJ (2010) Numerical investigation of vortex shedding and vortex-induced vibration for flexible riser models. Inter J Nav Archit Oc Engng, pp 112–118. https://doi.org/ 10.3744/jnaoe.2010.2.2.112 17. Haberyan KM (1956) Rational forms of pipes, tanks and pressure slab/Gosstroiizdat: 206 18. Sousa JRM et al (2010) An experimental and numerical study on the axial compression response of flexible pipes. J Offshore Mech ad Arct Eng 134(3). https://doi.org/10.1115/ omae2010-20856 19. Souza AP (2002) Flexible Pipe’s Collapse under External Pressure, D.Sc. thesis, COPPE, Federal University of Rio de Janeiro, Rio de Janeiro, Brazil. (In Portuguese) 20. Rogers CDF, Chapman DN, Wan F et al (2002) Laboratory testing of pipe splitting operations. Tunn Undergr Space Technol 17:99–113. https://doi.org/10.1016/S0886-7798(01) 00061-X 21. Stolyarov NN (1996) Elastic-plastic deformation and optimization of flexible shells and variable-rigidity plates. Bulletin of the Samara State Technical University. Series Physics and mathematics, pp 63–78 22. Ramberg S (2006) The effects of yaw and finite length upon the vortex wakes of stationary and vibrating circular cylinders. J Fluid Mech 128:81–1071983. https://doi.org/10.1017/ S002211208300039

Experimental Research on Reducing Hydraulic Resistance When Transporting High-Viscosity Fluids by Pipeline L. Ilina, N. Goncharov and A. Shagarova

Abstract The paper dwells upon reducing the hydraulic resistance of a pipeline by generating a wall-adjacent gas layer when transporting a high-viscosity medium. The research team has designed an experimental setup that contains a device for generating two-layer annular flow. The stability of such two-layer annular flow has been studied by simulating a viscous fluid (motor oil) and a wall-adjacent gas layer. Experiments show using a two-layer annular flow effectively reduces the hydraulic resistance when transporting a high-viscosity medium by a round pipe. Using two-layer annular flows in hydraulic systems and industrial pipelines improves the throughput and reduces the energy costs of transporting high-viscosity fluids by reducing pressure losses that occur when overcoming the friction forces arising during such transport. The experimental data thus show that a wall-adjacent gas layer can improve the pipeline throughput by up to *50% when transporting a high-viscosity Newtonian fluid through a rough-surfaced steel pipe. Keywords Hydraulic resistance Annular flow


Wall-adjacent layer
High-viscosity fluid


1 Introduction Reducing hydraulic resistance will help spend less energy to transport viscous and highly viscous media by industrial and transportation pipelines, making it a challenge worth addressing when designing new industrial pipelines and hydraulic systems, or retrofitting the existing ones.

L. Ilina (&)
A. Shagarova Volgograd State Technical University, 28, Lenin Avenue, 400005 Volgograd, Russia e-mail: [email protected] N. Goncharov JSC Kaustik, 40 Let VLKSM Street, Building 57, 400097 Volgograd, Russia © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_144

1369

1370

L. Ilina et al.

Review of the hydraulic resistance reduction methods shows that two-layer flow of viscous media in an annulus of low-viscosity fluid or gas is a promising technique [1–7]. Various designs have been made to generate stable two-layer annular flow in pipelines [8–12]. The theory behind the flow of fluids in pipelines with low-viscosity wall-adjacent layers has been described in many sources; numerous researchers have proposed flow models for viscous fluids with various rheological characteristics in an annulus of low-viscosity fluid or gas [13–22]. The existing mathematical two-layer annular flow model of high viscosity Newtonian fluid in an annulus of low-viscosity fluid (gas) shows that the lowviscosity fluid must flow at 10 to 25% the flow rate of the primary high-viscosity fluid to generate stable two-layer annular flow [16]. This results in a significant reduction in hydraulic resistance and energy spent to transport high-viscosity fluids. To prove the developed two-layer annular flow model, the research team has run a series of experiments to study the stability of such flow in a vertical pipeline [16]. Experimental data (annular gas layer, horizontal pipeline, and high-viscosity fluid [17, 23]) show a considerable increase in the viscous-fluid flow rate while spending less gas to generate a wall-adjacent layer. The experimental setup used in [17, 23] uses a plastic pipe, while the compressor used therein enables no constant gas flow rate to generate wall-adjacent layers. Notably, hydrodynamics assumes that roughness has only minor effects on the pressure losses in straight gas and plastic pipelines [24], i.e., the pressure losses will not be significant for viscous fluids. Therefore, it is advisable to research the stability of viscous-fluid flow in an annulus of low-viscosity fluid or gas in a straight rough-surfaced pipeline, as such research will enable optimizing the primary viscous fluid or gas flow to generate stable annular flow.

2 Experimental Setup with a Two-Layer Annular Flow Generator An experimental setup has been made to test the stability of two-layer annular flow. The setup is shown in Fig. 1 and comprises the primary-fluid feed tank 1 mounted on the stand 2 and connected to the receiving tank 5 in the reservoir 6 via the spacer 12 of the thin-walled steel pipe 4, which is 1000 mm long and has an ID of 16 mm. Gas is supplied to the wall-adjacent layer by the compressor 7. Compressor equipment includes the pressure gauge 9, the air-feed tap 8, the tap-equipped pipe 10, the gas meter 11, and the wall-adjacent annular layer generator 12; this equipment is connected to the pipe 4 to generate a wall-adjacent annular gas layer. The pipe 4 is mounted horizontally to the reservoir 6 and is held in place by the holder 13.

Experimental Research on Reducing Hydraulic Resistance …

1371

Fig. 1 Experimental setup layout. 1 is the feed tank; 2 is the stand; 3 and 10 are taps; 4 is the pipe; 5 is the receiving tank; 6 is the reservoir; 7 is the compressor; 8 is the tap; 9 is the pressure gauge; 11 is the gas meter; 12 is the wall-adjacent annular layer generator; and 13 is the holder

Fig. 2 Wall-adjacent annular layer generator. 1 is the casing; 2 is a bushing; 3 is a sleeve; 4 is a socket; 5 is the pipe; 6 is a lining; and 7 is a screw

To generate wall-adjacent gas layer, a generator has been made, see Fig. 2. The generator has a casing with an annular cavity inside, which is connected to the gas supply system to generate annular wall-adjacent layer; the other end is connected to the pipe via an annular gap between the outer surface of the sleeve and the inner surface of the pipe.

3 Research Objects and Their Properties The primary fluid used herein is mineral motor oil for gasoline cars, see Table 1 for characteristics. Wall-adjacent annular gas layer is generated by means of air, see Table 2 for characteristics.

1372 Table 1 Lukoil Standard 10w-40 oil

Table 2 Ambient air

L. Ilina et al. Indicator

Value

Kinematic viscosity, m2/s, at 20 °C Dinamic viscosity, Pa∙s, at 20 °C Mass fraction of mechanical impurities, % Mass fraction of water, % Congelation temperature, °C Density, kg/m3, at 20 °C

0.594  10−3 0.526 0.012 traces 35 886

Indicator

Value

Dinamic viscosity, Pa∙s, at 20 °C Density, kg/m3, at 20 °C

0.018  10−3 1.184

4 Experimental Technique Experiments use the primary-fluid (motor-oil) flow as a function of the secondarymedium (gas) flow, where the secondary flow generates a wall-adjacent annular layer. A standard glass thermometer is used to find the temperature of oil and air; then a standard glass capillary viscosity type BПЖ-1 and an areometer type AHT-1 are used to measure the kinematic viscosity and the density of motor oil. Experimental viscosity and density testing techniques are well-known, whence their description is omitted. The viscosity and density of the air used to generate the wall-adjacent layer are taken from reference tables [4] for known temperatures. The primary-fluid flow Qo is found by measuring the volume and flow time by a measuring tank and a stopwatch. The gas (air) flow Qa is found by measuring the gas supplied over a unit of time by means of an CГMБ rotor-type gas meter and a stopwatch. Below is the experimental procedure to find the primary-fluid (motor and vegetable oil) flow Qo as a function of the secondary-medium (air) flow Qa, where the auxiliary medium is used to generate the wall-adjacent layer: • pour the primary fluid into the feed tank 1 and pump air into the compressor 7 tank while monitoring its pressure by the gauge 9; • start the stopwatch and open the tap 3 so that gravity could force the primary fluid to flow from the feed tank 1 to the pipe 4 via the channel in the sleeve of the spacer 12; • open the tap 10 and adjust it to a specific flow; supply compressed air via the tube 9 to the annular cavity in the casing of the spacer 12, then to the pipe 4 via the annular gap so as to generate a wall-adjacent gas layer of a thickness specific to the feed flow; • drain the primary fluid to the receiving tank 5;

Experimental Research on Reducing Hydraulic Resistance …

1373

• stop the stopwatch in some time, close the taps 3 and 10, and stop supplying the primary fluid and the secondary fluid (gas); • measure the primary-and secondary-fluid volumes flown during the experiment to find their volumetric flows Qo, Qa; • the air tap 8 to adjust the secondary-medium (gas) flow and repeat the experiment.

5 Experimental Data Analysis Figure 3 presents the results of the experiments carried out to find the primary-fluid (motor-oil) flow Qo as a function of the secondary-medium (air) flow Qa, where the secondary medium is used to generate the wall-adjacent layer: Generating a wall-adjacent fluid (gas) layer while the primary-fluid (motor oil) flows through a pipe does reduce the hydraulic resistance of the pipe and improves the primary-fluid flow rate; primary-fluid flow Qo greatly correlates with the secondary-medium (gas) flow Qa. Increasing the gas flow Qa increases the primary-fluid flow Qo to a certain threshold, after which the latter is becoming reduced as the secondary-medium ousts the primary fluid, while the flow transforms from laminar to turbulent, which entails a higher resistance to the primary-fluid transport. Compare the experimental data obtained herein to those presented in [17, 23]. For clarity, see Fig. 4.

Fig. 3 Primary-fluid (motor-oil) flow Qo as a function of the gas flow Qa. Motor oil: t = 20 °C, µ = 0.526 Pa∙s, q = 886 kg/m3A. ir: t = 20 °C, µ = 0.018 Pa s, and q = 1.184 kg/m3

1374

L. Ilina et al.

Fig. 4 Primary-fluid (motor-oil) flow Qo as a function of the air flow Qa: a comparison. 1 for the experimental data per [17, 23]. 2 for the experimental data obtained herein

The functions 1 and 2 shown above mean that the experimental findings may be deemed correct, as the results deviate proportionally, which is due to the pipe insides being different in roughness. When transporting the primary fluid (motor oil) without a wall-adjacent gas layer, the function 1 shows a 30% increase in the motor-oil flow at the onset against the experimental data (2), which is also due to the pipeline surfaces being smooth. Reproducibility is confirmed by the array of data obtained experimentally by the proposed procedure. Considerable reduction in the gas flow in the experiment (2) is due to stabilizing the annular wall-adjacent layer by placing a compressor to adjust the air feed. The experimental data thus show that a wall-adjacent gas layer can improve the pipeline throughput by up to *50% when transporting a high-viscosity Newtonian fluid through a rough-surfaced steel pipe. Comparing the viscous-fluid flow with a wall-adjacent gas layer in a smooth or rough pipeline makes it clear this approach improves the throughput regardless of the pipeline type.

References 1. Chernikin VI (1949) Gidrotransport nefteproduktov po trubam (Hydraulic Transport of Petroleum Products by Pipelines). Proceedings of Gubkin Moscow Petroleum Institute, iss. 9 2. Charles ME, Redberge PJ (1967) The reduction of pressure gradients in oil pipelines by the addition of water. Cana J Chem Eng

Experimental Research on Reducing Hydraulic Resistance …

1375

3. Mastobayev BN, Shammazov AM, Movsumzade EM (2002) Khimicheskiye sredstva i tekhnologii v truboprovodnom transporte nefti (Chemical Agents and Technologies for Oil Transport by Pipelines). Khimiya, Moscow, p 296 4. Aliyev RA et al (1988) Truboprovodny transport nefti i gaza (Pipeline Transport of Oil and Gas): university textbook, Nedra, 368 p 5. Skripnikov YuV (1988) Vliyaniye malovyazkogo pristennogo sloya na ustoychivost techeniya v krugloy trube (How Low-Viscosity Wall-Adjacent Layer Affects Round-Pipe Flow Stability). In: Improving Trunkline Operation and Control Systems (Sovershenstvovaniye sistem upravleniya i ekspluatatsii magistralnogo transporta), Ufa, VNIISPTneft 6. Tugunov PI, Nechval MV, Novosyolov VF, Akhatov SHN (2005) Ekspluatatsiya magistralnykh truboprovodov (Trunkline Operations). Bashkir Publ., Ufa-25, 84 p 7. Ilina LA, Golovanchikov AB, Ilin AV, Vasilyeva YeV (2014) Snizheniye gidravlicheskogo soprotivleniya pri dvukhsloynom koltsevom techenii vysokovyazkoy zhidkosti v truboprovode (Hydraulic-Resistance Reduction by Two-Layer Annular Flow of High-Viscosity Fluids in Pipelines). Izvestia VSTU. Series: Rheology, Processes, and Devices in Chemistry. Iss. 7: Inter-University Proceedings, VSTU 1 (128):114–117 8. Patent 84925 Russia, IPC F 15 D 1/06, F 17 D 1/20. Device to Reduce Hydraulic Losses in Pipelines (Ustroystvo dlya umensheniya gidravlicheskikh poter v truboprovode)/ A. B. Golovanchikov, Ye. N. Konopaltseva, A. V. Ilin, L. A. Ilina, M. G. Chernobrovkina; VSTU. – 2009 9. Golovanchikov AB, Ilina LA, Ilin AV e al (2008) Ustroystvo dlya umensheniya gidravlicheskikh poter v truboprovode (Device to Reduce Hydraulic Losses in Pipelines). Patent 2334134 Russia, IPC F 15 D 1/06, F 17 D 1/20, VSTU 10. Golovanchikov AB, Ilina LA, Ilin AV e al (2008) Sposob peremeshcheniya vysokovyazkikh zhidkostey po truboprovodu (Method for Pipeline Transport of High-Viscosity Fluids). Patent 2334161 Russia, IPC F 17 D 1/14, F 15 D 1/02, VSTU 11. Golovanchikov AB, Ilina LA, Ilin AV e al (2006) Ustroystvo dlya umensheniya gidravlicheskikh poter v truboprovode (Device to Reduce Hydraulic Losses in Pipelines). Patent 2285198 Russia, IPC F 17 D 1/20, F 15 D 1/06. VSTU 12. Golovanchikov AB, Ilina LA, Ilin AV e al (2004) Ustroystvo dlya umensheniya gidravlicheskikh poter v truboprovode (Device to Reduce Hydraulic Losses in Pipelines). Patent 2241868 Russsia, IPC 7 F 15 D 1/06, F 17 D 1/20. VSTU 13. Golovanchikov AB, Ilina LA, Ilin AV et al (2006) Transportirovka nefti i nefteproduktov s gazovym pogranichnym sloyem (Boundary Gas Layer in Oil and Petroleum-Product Transport). Tekhnologii nefti i gaza 4:10–14 14. Golovanchikov AB, Ilin AV, Ilina LA (2012) Model dvukhsloynogo koltsevogo techeniya vyazkoy zhidkosti s proizvolnoy reologicheskoy krivoy s malovyazkim pristennym sloyem v trube (Two-Layer Annular-Flow Model for a Viscous Fluid with an Arbitrary Rheological Curve and a Low-Viscosity Wall-Adjacent Layer). Izvestia VSTU. Series: Rheology, Processes, and Devices in Chemistry. Iss. 5: Inter-University Proceedings 1:12–14 15. Golovanchikov AB, Ilin AV, Ilina LA (2004) Techeniye v trube nenyutonovskoy zhidkosti s malovyazkim pogranichnym sloyem (Non-Newtonian Fluid: Pipe Flow with a Low-Viscosity Boundary Layer). Izvestia VSTU. Series: Conceptual Design in Education, Engineering, and Technology: Inter-University Proceedings 1(5):19–21 16. Golovanchikov AB, Ilin AV, Ilina LA (2007) Teoreticheskiye osnovy techeniya zhidkostey v truboprovode s malovyazkim pogranichnym sloyem (Theory Behind Fluid Flow in Pipelines with Low-Viscosity Boundary Layers). VSTU, Volgograd, p 108 17. Golovanchikov AB, YeV Vasilyeva, Ilina LA (2017) Granichnaya ustoychivost geterofaznykh zhidkostey na makro- i mikrourovnyakh (Marginal Stability of Heterophase Lluids at Macro- and Micro-Levels). VSTU, Volgograd, p 131 18. Archibong-Eso A, Shi J, Baba YD, Aliyu AM, Raji YO, Yeung H (2018) High viscous oil– water two–phase flow: experiments & numerical simulations. Heat Mass Trans J

1376

L. Ilina et al.

19. Andrade THF, Crivelaro KCO, Neto SRF, Lima AGB (2013) Isothermal and non-isothermal water and oil two-phase flow (core-flow) in curved pipes. Int J Multiphysics, 7(2) 20. Cazarez-Candia O, Piedra-González S (2017) Modeling of heavy oil-water core-annular upward flow in vertical pipes using the two-fluid model. J Pet Sci Eng 150:146 21. Gupta R, Turangan CK, Manica R (2016) Oil-water core-annular flow in vertical pipes: a CFD study. Can J Chem Eng 94 22. Gadelha AJF, de Farias Neto SR, Swarnakar R, de Lima AGB (2013) Thermo-hydrodynamics of core-annular flow of water, heavy oil and air using CFX. Adv Chem Eng Sci 3:37 23. Ilin AV, Kareva GN, YeV Vasilyeva, Ilina LA, Golovanchikov AB (2013) Experimental Research of Petroleum-Product Flow Stability in Pipes with Wall-Adjacent Fluid (Gas) Layers (Eksperimentalnye issledovaniya ustoychivosti techeniya nefteproduktov v trube s pristennym zhidkim (gazovym) sloyem). Tekhnologii nefti i gaza 5:36–41 24. Pavlov KF, Romankov PG, Noskov AA (2006) Primery i zadachi po kursu protsessov i apparatov khimicheskoy tekhnologii (examples and problems for the course in chemical processes and devices). 13th edn, reprint, Alians, Moscow, 575p

Modeling Three-Dimensional Liquid Flows in Computer-Controlled Vibrojet Mixer Using FlowVision Yu. S. Sergeev, S. V. Sergeev and G. E. Karpov

Abstract The chapter shows that to study the three-dimensional flows arising from the complex effect of vibration on Newtonian fluids, it is best to use a topological method of computer modeling. For reliability and clarity of the numerical modeling results obtained, FlowVision software and SUSU supercomputer resources were used. The calculation results showed the superiority of the new controlled vibrojet method over the traditional process of mixing liquid media. The calculations of flows rates were performed. The conditions for the formation of internal submerged toroidal flows and counter-swirling jets inside these flows are revealed. The simulation results made it possible to determine the control parameters of the process, to describe the functionality and technological capabilities of the computercontrolled vibrojet mixer. Its design is based on a fundamentally new method of controlled vibration mixing of multicomponent mixtures and solutions used in the chemical, pharmaceutical, food, engineering, and mining industries. As a result, practical technology has received a tool for digital processing of the mixing process.










Keywords Modeling FlowVision Mixing automation Liquid multicomponent systems Digital twin mixing




1 Introduction The problems of the quality of liquid multicomponent mixtures and the productivity of production of them are quite acute in chemical [1], pharmaceutical [1], and food industries [2]. No less acutely these problems are indicated in mechanical engineering [3, 4] and mining [5] in the manufacture of process fluids. The useful life of such mixtures is mainly limited by the time of their fractionation and the degree of saturation with impurities during their operation. Traditional methods of mixing, such as circulation, mechanical, shock pulse, acoustic, homogenization, or Yu. S. Sergeev
S. V. Sergeev (&)
G. E. Karpov South Ural State University, 16, Str. Turgeneva, Zlatoust 456209, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_145

1377

1378

Yu. S. Sergeev et al.

bubbling, do not give the desired result [3]. This is due to the fact that, firstly, the rotating flow, even if it rapid, has a low velocity gradient due to the solidity of movement in the vortex, and secondly, intensive processing of the mixture occurs only in a narrow area that is close to the mixer working body surface. And as you move away from it, the intensity of the process decreases, forming stagnant zones [6]. Therefore, any optimization of the modes or minor constructive refinements of traditional mixers as a rule do not give a really tangible increase in the performance of the process and the quality of the product. When solving such tasks in production, technologists and designers usually rely on the opinion of experts, who are not always available in the company. This means that nothing steals development resources more than archaic technologies. The mixing process is essentially the dispersion of liquid and solid media [7]. In the preparation of mixtures and solutions, the layers of liquid and amorphous bodies, that adjacent to the working body of the mixer—the activator, carried away by the force of friction under the effect of centrifugal force are thrown out from it, forming flows [8]. At the same time, they collide and break up into droplets, or amorphous dispersed particles if they encounter obstacles in their way that prevent them from following this force. Such obstacles are the components of the solutions and mixtures themselves, which are outside the active zone. Moreover, the more such flows will be [9] inside the medium, the more frequent the collisions of particles and drops will happen. And this means that these components will be deformed and destroyed more quickly, so the mixing process will be more intensive. It is easiest to provide a multiplicity of particle flows by applying a complex controlled vibration effect on the medium [10]. By increasing the number of interactions of opposing flows, it is possible increase the turbulization of the mixture, i.e., perform the same work in a shorter period of time. This means the performance of the process can be improved.

2 Main Body A significant technological leap in quality assurance of multicomponent liquid systems and at the same time in increasing the performance of their production can be achieved using a new method of vibrojet mixing of them [11]. The efficiency of the process can be significantly improved by eliminating stagnant zones [12, 13] due to increased turbulization and volumetric circulation of flows. This, in turn, is accomplished by transmitting of mutual rotation with the angular velocity x and radial–axial vibrations with the angular velocity xrot from the working body 1 of the mixer 2 (see Fig. 1a) to a liquid medium. Such an approach will make it possible to increase the number of opposing submerged jets (toroidal flows) 3 and the number of their interactions (see Fig. 1b). For deeper turbulization inside the toroidal rotating flows of the mixture (see Fig. 1a, c), along with x and xrot, the working body 1 with perforated disks 4 is additionally given controlled pulsations ±e of the angular velocity xrot. Only with

Modeling Three-Dimensional Liquid Flows in Computer …

1379

Fig. 1 Functional diagram of the vibrojet mixer: a scheme of setup of the complex oscillatory process; b scheme of formation of toroidal flows of the mixture; c scheme of formation of swirling mixture flows

such a complex vibration effect on the liquid medium do the swirling opposing jets 5 (see Fig. 1c) form inside the toroid-shaped streams, amplifying the vibrojet effect [14–16]. Thus, the complex movements of the working body generate more complicated forms of the trajectories of moving flows. They, in turn, more often overlap and interact with each other, which significantly increases the length of the mixing path [11]. All this leads to turbulization increase. In accordance with the new method [11], the generated fluid flows are described by the modernized Navier–Stokes equations system that takes into account all forced vibration moves. In this case will dominate such control parameters as angular velocity x, amplitudes a1 and a2 of radial–axial oscillations, and magnitude of ripples of the angular velocity xrot± e that have an impact on the values of the components of medium flows velocities Vr, Vz, Vt. Moreover, these movement parameters control the step of the spiral trajectories of the oncoming submerged jets, and hence the intensity of their interaction. In particular, the control of the values a1 and a2 (see Fig. 1b, c) in the area of each perforated disk is provided by varying the amplitude a of circular oscillations of the working body 1 according to [17]. a¼

Dxrot ; 2x

ð1Þ

1380

Yu. S. Sergeev et al.

Fig. 2 Functional diagram of the vibrojet mixer drive

and the angular velocity x Pa x¼ þ 2lmxrot

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi P2a j þ ; 2 2 2 4l m xrot m

ð2Þ

where D is the working body diameter, Pa is the working body pressing axial force, l is the displacement of the working body, m is the effective working body mass, and j is the working body stem stiffness (if joint is hinged then j = 0). It was possible to achieve this functional superiority by using the switched reluctance vibrodrive (SRD) in the computer-controlled mixer (see Fig. 2) [14, 16]. The pulsations of xrot± e are controlled by the operation of the electromechanical converter (EC) itself. The switched reluctance motor (SRM) used for this, in turn, is controlled by the logic of the control system [14] in accordance with the required law of the angular acceleration ±e. It is important to note that the dynamic model of SRD [16] takes into account both the mechanical inertia of the drive and the load moment, which is significantly dependent on the viscosity of the mixture. Regulation of the parameters of the angular velocity xrot pulsations is carried out from the control computer via changing the angle of switching on h0 and the angle of switching off he the winding. At the output of EC, the average value of the angular velocity xrot per revolution is controlled via the rotor position sensor (RPS) and the functional converter (FC). The proportional speed controller (SC) generates the voltage USC that is applied to the input of the control unit (CU). The control unit generates control signals Xc to the commutator taking into account set pulsation parameters. The semiconductor commutator (ESC) forms the phase voltages Upi for the converter in accordance with the control signals and the rotor position fr. In turn, EC transmits mechanical movements to the mixer working body (MWB). It is important to note that any adopted technical solution imposes certain limitations, so the design of EC and the choice of its control method must be carried out in a complex, for a specific task. The circuit uses positional switching with the PWM voltage regulation, allowing to set the required laws of oscillations. To control the specified technological parameters, the automated system for monitoring the generated oscillations of working body (AMS) is also provided in the control system of the mixer–disperser (Fig. 3) [18].

Modeling Three-Dimensional Liquid Flows in Computer …

1381

Fig. 3 Functional diagram of the vibrojet mixer drive

The drive stability is ensured by collecting information from proximity sensors, its automatic processing, and system operation adjustment through the integrated electrical feedback [18]. For the practical use of the proposed mixer with computer control, it was necessary to conduct extensive research into the possibilities of controlling the mixing process depending on the physicomechanical properties of the components of the mixture. Therefore, the aim of the work is the modeling such conditions of vibrojet mixing, under which the maximum interaction of multidirectional submerged jets of condensed media is ensured. To achieve this goal, the trajectories of motion of the submerged jets of the mixture were investigated. Since these are fast processes, it is advisable to use a topological modeling method [19–22] that has been widely used recently to study the formation of the structure of non-Newtonian fluids. For the reliability and clarity of the obtained results of modeling experiments, the calculations used the FlowVision program, which used a method based on conservative schemes for calculating non-stationary partial differential equations. Compared with non-conservative schemes, they provide solutions that exactly satisfy the conservation laws (in particular, the continuity equation) [11, 21–24]. To solve problems in the FlowVision package, first a device model was created with specified design parameters using an external program—a geometric preprocessor. As such a preprocessor, the SolidWorks package was used, which became widespread in modern scientific and engineering practice [25]. For a more complete and visual representation of the processes of vibrojet mixing [11] of liquid media, the SUSU supercomputer resources [26, 27] and the FlowVision software were used to calculate the flows rates and the optimal temporal characteristics at varying amplitude–phase–frequency parameters of the vibration drives [14, 17]. Some of the results of calculations are presented in Fig. 4 —comparative analysis (with visualization) of the parameters of the traditional mixing process without superimposed radial–axial vibrations (Fig. 4a) and vibration mixing (Fig. 4b, c). From Fig. 4, it is easy to see that the process of vibrating is much more intensive (see on the scale of rate and density of flows); i.e., the number of internal submerged jets significantly increases and, accordingly, the length of their mixing. At the same time, there are practically no stagnation zones in the mixer.

1382

Yu. S. Sergeev et al.

Fig. 4 Results of the simulation of mixing processes: a traditional way (no oscillations); b with the imposition of radial–axial oscillations; c with the imposition of radial–axial oscillations and angular velocity ripples

In the course of computer simulation, for each concrete mixture, not only the required technological parameters and time intervals were identified, but also the optimal quantities and geometry of perforated disks, the geometry of the holes, and their influence on the performance of the mixing process. The zones of increased turbulence in the total volume of the mixture and possible ways of its further increase were determined by calculation. Some results of the calculation are shown in Fig. 5. A tangible economic effect is expected from the introduction of a new method of mixing liquids, as there are a number of disadvantages in traditional methods of mixing, which do not allow to effectively manage the mixing process. Therefore, replacing the known designs of mixers with a mixer, which will work in a new way of mixing, will increase the efficiency and productivity of the process. And with certain skills of the technologist, it becomes possible to preliminarily model the process with visualization of it and verification of modes, without resorting to carrying out expensive field experiments. Next, the mixer is set to automatically maintain the designed modes, which may even be variable in the mixing process.

Modeling Three-Dimensional Liquid Flows in Computer …

1383

Fig. 5 Zones of turbulent modes: a toroidal flows; b spiral flows (increased turbulence)

3 Conclusions (1) The results of computer simulation have shown that the use of a fundamentally new method of exciting oscillations, implemented in a rotary inertial vibration drive equipped with SRD, allows to obtain a number of the physical effects. Their simultaneous implementation allows to improve the equipment for mixing the components of mixtures and solutions, and, consequently, to intensify the technological process of preparation of multicomponent liquid systems. Therefore, when setting up the installation, it is necessary, depending on the physical properties of the process medium, to select such frequency and amplitude of oscillations of the mixer rotor with perforated disks at which high product quality and maximum performance will be combined. (2) The obtained simulation results allow digital twin through which succeeds not only to calculate and visualize physical processes when mixing the components of a liquid medium, to select and maintain the necessary modes for their implementation, but also to control the intensity of the mixing process in automatic mode. Acknowledgements South Ural State University is grateful for financial support of the Ministry of Education and Science of the Russian Federation (grant No 9.7960.2017/BP).

References 1. Hessel V, Noël T (2012) Micro process technology, 2. Processing. Ullmann’s Encyclopedia of Industrial Chemistry. https://doi.org/10.1002/14356007.b16_b37.pub2 2. Xia B, Sun D-W (2002) Applications of computational fluid dynamics (cfd) in the food industry: a review. Comput Electron Agric 34(1):5–24. https://doi.org/10.1016/S0168-1699 (01)00177-6

1384

Yu. S. Sergeev et al.

3. Latyshev VN (1975) Improving the efficiency of cutting fluid. Mechanical Engineering, Moscow 4. Khudobin LV, Babichev AP, Bulyzhev EM (2006) Lubricating and cooling technological tools and their use in machining. Mechanical Engineering, Moscow 5. Hoelscher KP, De Stefano G, Riley M, Young S (2012) Application of nanotechnology in drilling fluids. In: Paper presented at the SPE International Oilfield Nanotechnology Conference and Exhibition, Noordwijk, The Netherlands 6. Matsunaga D, Imai Y, Yamaguchi T, Ishikawa T (2016) Rheology of a dense suspension of spherical capsules under simple shear flow. J Fluid Mech 786:110–127. https://doi.org/10. 1017/jfm.2015.666 7. Dodd MS, Ferrante A (2016) On the interaction of Taylor length scale size droplets and isotropic turbulence. J Fluid Mech 806:356–412. https://doi.org/10.1017/jfm.2016.550 8. Schlichting G (1974) Theory of the boundary layer. Publishing House “Science”, Moscow, p 712 9. Yatsun SF, Mishchenko VY, Mishchenko EV (2009) The impact of vibration effects on the extraction process in the food industry. News of universities. Food Technol 4:70–72 10. Blekhman II (1994) Vibration mechanics. Science, Moscow 11. Sergeev YS, Sergeev SV, Zakirov RG, Nekrutov VG, Gordeev EN, Irshin AV et al (2013) Method of mixing liquid. Russian Federation Patent 2543204, 27 Feb 2015 12. Kulkarni PM, Morris JF (2008) Suspension properties at finite Reynolds number from simulated shear flow. Phys Fluids 20(4):040602. https://doi.org/10.1063/1.2911017 13. Sergeev YS, Sergeev SV, Maltsev PS (2015) Intensification of hydrodynamic processes in the preparation and regeneration of technological multicomponent mixtures Science SUSU: Materials of the 67th scientific conference. Publishing Center SUSU, Chelyabinsk, pp 1762– 1766 14. Sandalov VM, Sergeev YS (2012) Dynamic model of switched reluctance vibratory drive. Russian Electrical Engineering 83(8):432–435. https://doi.org/10.3103/S1068371212080111 15. Sergeev YS, Sandalov VM, Karpov GE (2017) Modeling switched reluctance vibratory drive. Bulletin of SUSU. Series: Energy 17(4):90–98. https://doi.org/10.14529/power170410 16. Sergeev YS, Sandalov VM, Karpov GE (2018) ‘Modeling of switched reluctance electric vibration drive with adaptive control’ 2018 international russian automation conference (RusAutoCon), 9–16 Sept 2018, pp 1–4 17. Lakirev SG, Khilkevich YM, Sergeev SV (1988) The method of excitation of circular oscillations and device for its implementation. USSR Patent 1664412, 1991 18. Sergeev YS, Sergeev SV, D’yakonov AA, Kononistov AV, Karpov GE, Mikryukov AA (2018) Automated monitoring system for self-synchronizing vibrational drives. Russian Engineering Research 38(2):86–90. https://doi.org/10.3103/s1068798x18020168 19. Rudman M (1998) Volume-tracking methods for interfacial flow calculations. Int J Numer Meth Fluids 24(7):671–691. https://doi.org/10.1002/(SICI)1097-0363(19970415)24:7% 3c671:AID-FLD508%3e3.0.CO;2-9 20. Stickel JJ, Powell RL (2005) Fluid mechanics and Rheology of dense suspensions. Annu Rev Fluid Mech 37(1):129–149. https://doi.org/10.1146/annurev.fluid.36.050802.122132 21. Rosti ME, De Vita F, Brandt L (2018) Numerical simulations of emulsions in shear flows. Acta Mechanica. https://doi.org/10.1007/s00707-018-2265-5 22. Izbassarov D, Rosti ME, Ardekani MN, Sarabian M, Hormozi S, Brandt L et al (2018) Computational modeling of multiphase viscoelastic and elastoviscoplastic flows. Int J Numer Meth Fluids 88(12):521–543. https://doi.org/10.1002/fld.4678 23. Kim J, Moin P (1985) Application of a fractional-step method to incompressible Navier-Stokes equations. J Comput Phys 59(2):308–323. https://doi.org/10.1016/0021-9991 (85)90148-2 24. Pastrana D, Cajas JC, Lehmkuhl O, Rodríguez I, Houzeaux G (2018) Large-eddy simulations of the vortex-induced vibration of a low mass ratio two-degree-of-freedom circular cylinder at subcritical Reynolds numbers. Comput Fluids 173:118–132. https://doi.org/10.1016/j. compfluid.2018.03.016

Modeling Three-Dimensional Liquid Flows in Computer …

1385

25. Alizad Banaei A, Loiseau J-C, Lashgari I, Brandt L (2017) Numerical simulations of elastic capsules with nucleus in shear flow. Eur J Comput Mech 26(1–2):131–153. https://doi.org/10. 1080/17797179.2017.1294828 26. Kostenetskiy P (2016) SUSU supercomputer resources 27. Rekachinsky AI, Chulkevich RA, Kostenetskiy PS (2018) Modeling parallel processing of databases on the central processor Intel Xeon Phi KNL’ 2018. In: 41st international convention on information and communication technology, Electronics and Microelectronics (MIPRO), 21–25 May 2018, pp 1605–1610

Improving Efficiency of Boiler in Case of Coal Hydrotransport K. V. Osintsev, M. M. Dudkin and Iu. S. Prikhodko

Abstract There are different schemes of utilization of heat of combustion products discharged from the boilers of CHP. The efficiency of heat recovery depends on the type of combustible organic fuel. This problem is the most relevant when using solid fuel dust and coal-water slurries. Gas recirculation boilers may be used. For the secondary use of incomplete combustion products, you can use boilers designed to burn industrial gases and coal-water fuel. A variant of an experimental system for physical modeling of coal-water fuel movement through a pipeline has been proposed. The plant improves the reliability of the existing systems of inspection and monitoring of the flame burning in the furnace of the boiler unit by simulating real processes and determining the air excess factor. The use of non-standard ways to improve the efficiency of thermal schemes of CHP and boiler rooms is possible with the use of alternative types of fuel, which include the considered coal-water slurries and coal-water fuel. Keywords Heat utilization


Exhausted gas
Alternative fuel

1 Introduction An economical method of substituting gas for recycling full-combustion products from the gas turbine electrical-generating superstructure can be realized at a power plant equipped with gas-burning boilers Fig. 1a [1–4]. Slotted burners 12 are installed in the bottom part of the furnaces 1 of these boilers. Natural gas and hot air from air heaters 5 are supplied to the slotted burners. A flame is formed in the furnace 1 during the combustion of natural gas. The torch heats the steam–water medium in the pipes of economizers 4 and superheaters 3. Combustion products (flue gases) are discharged through air heaters 5 by smoke exhausters 6 to the chimney 7. To reduce the concentration of nitrogen oxides in the K. V. Osintsev (&)
M. M. Dudkin
Iu. S. Prikhodko South Ural State University, 76, Lenin Avenue, Chelyabinsk 454080, Russia e-mail: [email protected] © Springer Nature Switzerland AG 2020 A. A. Radionov et al. (eds.), Proceedings of the 5th International Conference on Industrial Engineering (ICIE 2019), Lecture Notes in Mechanical Engineering, https://doi.org/10.1007/978-3-030-22063-1_146

1387

1388

K. V. Osintsev et al.

Fig. 1 a Diagram of industrial gas-fired boiler with hearth slotted burner and system of recirculation and additional input combustion products: 1—furnace; 2—exhaust duct with convection heating surfaces; 3—steam superheater; 4—evaporators; 5—air heater; 6—smoke exhausters; 7—smoke duct; 8—connective boiler flue; 9—exhaust duct recirculating flue gases; 10 —boiler fans; 11—air channel; 12—gas slotted hearth burner; 13—exhaust duct dumping combustion products from gas turboelectric power generator; 14—duct of additive exhaust smoke from gas turbine generator; 15—channel of additional cold air to the furnace; b Diagram of an industrial pulverized coal dust boiler with recirculation systems and additional input of complete combustion products: 1—furnace; 2—smoke duct with convection surface; 3—economizer; 4— water wall tubes; 5—steam superheater; 6—smoke exhausters; 7—connective boiler flue; 8— smoke exhausters; 9—chimney; 10—flue gas recirculation fans of the boiler; 11—exhaust duct recirculating duct; 12—disintegrating mill; 13—pulverized fuel pipe; 14—basic blower of boiler; 15—air channel; 16—coal dust burner; 17—cold air channel addition to furnace; 18—exhaust duct dumping combustion products from gas turboelectric power generator; 19—smoke duct transport combustion products from gas turboelectric power generator to furnace; 20, 21—bunker and coal feeder; 22—installation of ash collection

combustion products from the furnace 1 and the boiler, and increasing the temperature of the flue gases above the dew point temperature, the slotted burners 12 are supplied with flue gases together with air. Flue gases are collected behind the smoke exhausters 6 in front of the chimneys 7 on the suction side of the boiler blowing engine 10. The smoke–air mixture is heated in the air heaters 5 before being supplied to the burners 12. For replacing part of the flue gases circulating in the boiler by the combustion products from the exhaust duct 13 from gas turboelectric power generator, it is necessary to cut off the supply of gases to the suction of the blowing engines 10. This is achieved by closing the valves on exhaust duct recirculating flue gases of the boiler 9 and opening valve on the pressure flues of the combustion products additive 14 [5–8]. In the exhaust duct 14 from the blowing engines 10 cold air is supplied at an excessively high temperature level. The effect of integrated heat and power generation technology is realized by reducing the temperature level of combustion products from gas turbine plants and heat losses DqCHPP = q′CHPP − q ″CHPP. The consumption characteristics of smoke exhausters 6 are retained, when the recirculation gases replacing with the gas turbine power plants, therefore DEie  0 [9–11]. Reduction in energy consumption for recycling by blowing

Improving Efficiency of Boiler in Case of Coal Hydrotransport

1389

engines 10 generates a gain to the boiler complex DEg. It is necessary to perform a design check of the chimneys and perform systems for emergency discharge of combustion products to the atmosphere after the gas turbine power plants from the flue 13 for the version Fig. 1b. The combustion in boilers of highly reactive coal dust uses technology to suppress his igniting activity using gas recirculation Fig. 1b [12–14]. The flow of inert flue gases with oxygen concentrations