Advances in Materials and Technologies

Table of contents :
Advances in Materials and Technologies
Preface
Table of Contents
Chapter 1: Materials for Electronics
Optical and Electrical Characterization of Polymer Dispersed Nematic Liquid Crystals
Suppress Short Channel Effects on Split Channel-Cylindrical GAA TFET Using Buried Oxide Layer
Spray Pyrolyzed Praseodymium Doped SnO2 Thin Film with Fast Response to LPG: Analysis Based on Microstructural Features
Solvent Effects on the UV-Visible Absorption and Emission of Tris[4-Diethylamino)Phenyl]amine
Chapter 2: Materials Processing
Analysis of the Influence of Build Plate Pre-Treatment and Process Parameters on the Bonding of Additively Manufactured Parts of TiAl6V4 to the Build Plate
Laser Metal Deposition (LMD) Toolpaths with Adaptive Capability for Complex Repairs and Coating Geometries
Innovative Method of Embedding Optical Fiber inside Titanium Alloy Utilizing Friction Stir Forming
Towards a Model-Based Approach for the Optimization of the Mechanical and Economical Properties of Laser-Based Plastic-Metal Joints
Center Line Globular Structure of Strip Cast Using High Speed Twin Roll Caster
Investigation of End Milling Process by Machine Tool Equipped with an Idling-Stop Function in a Feed Driving System
Finite Element Modeling of Upper Ball Joint in a Two-Step Hot Forging Process
Casting of Wire Using Twin Wheel Caster with Melt Holding Nozzle
Chapter 3: Technologies of Welding Production
Optimization of Resistance Spot Welding Parameters in a Shop Floor Environment to Achieve Desired Spot Size in Low Carbon Steel Sheet
Optimization of Process Variables for Prediction of Penetration Depth of HSLA Steel Welds Using Response Surface Methodology
Microstructural and Mechanical Properties of Dissimilar Aluminium Alloys by Friction Stir Welding
Analysis of Thermal Fields, Weld Strength and Microstructural Studies of Friction Stir Dissimilar Weldments of AA6082 and AA7075
Mechanical and Metallurgical Characteristics of Rotary Friction Welded Low Carbon Steel Plate/Rod Joints
Chapter 4: Building Materials
Pervious Concrete Utilized as a Base Slab of Pavement Edge Drains
Experimental Study on Multi-Segmented Circular or Non-Circular Pipes to Determine the Pipe Stiffness
Coal Ash and Rice Husk Ash Binder for Manufacturing Hollow Blocks
Flexural Behaviors of Modified Recycled Aggregate Concrete Reinforced with Pultruded GFRP
Keyword Index
Author Index

Citation preview

Advances in Materials and Technologies

Edited by Dr. Ramya Muthusamy Dr. Thangaprakash Sengodan Prof. Vladimir Khovaylo Nguyen Quang Liem Prof. Takahiro Ohashi

Advances in Materials and Technologies

Special topic volume with invited peer-reviewed papers only

Edited by

Dr. Ramya Muthusamy, Dr. Thangaprakash Sengodan, Prof. Vladimir Khovaylo, Nguyen Quang Liem and Prof. Takahiro Ohashi

Copyright  2022 Trans Tech Publications Ltd, Switzerland All rights reserved. No part of the contents of this publication may be reproduced or transmitted in any form or by any means without the written permission of the publisher. Trans Tech Publications Ltd Seestrasse 24c CH-8806 Baech Switzerland https://www.scientific.net

Volume 934 of Key Engineering Materials ISSN print 1013-9826 ISSN web 1662-9795

Full text available online at https://www.scientific.net

Distributed worldwide by Trans Tech Publications Ltd Seestrasse 24c CH-8806 Baech Switzerland Phone: +41 (44) 922 10 22 e-mail: [email protected]

Preface The presented edition is devoted to investigating modern materials science, technologies of materials processing and welding processes problems. The first chapter is related to the research on electrical and optical properties of semiconductors and polymeric materials used in the modern electronic and optoelectronic industry. The properties of special thin film are also discussed. The following two chapters research structural materials processing technologies: laser deposition, casting, milling, etc., and modern welding technologies. Research on some modern types of concrete and waste-based binder are presented in the last chapter. This publication will be helpful to many specialists whose activities related to machinery and construction.

Table of Contents Preface

Chapter 1: Materials for Electronics Optical and Electrical Characterization of Polymer Dispersed Nematic Liquid Crystals S. Mani, M. Pradhan, P. Rai, S. Khosla and P. Sarawade Suppress Short Channel Effects on Split Channel-Cylindrical GAA TFET Using Buried Oxide Layer P. Dhake, J. Ghosh, M. Joshi, R. Mathew and A. Beohar Spray Pyrolyzed Praseodymium Doped SnO2 Thin Film with Fast Response to LPG: Analysis Based on Microstructural Features S. Deepa, P.K. Krishnan and B. Thomas Solvent Effects on the UV-Visible Absorption and Emission of Tris[4Diethylamino)Phenyl]amine S. Singh, A. Singh Nain and A. Kumar

3 15 23 37

Chapter 2: Materials Processing Analysis of the Influence of Build Plate Pre-Treatment and Process Parameters on the Bonding of Additively Manufactured Parts of TiAl6V4 to the Build Plate C. Bay, A. Mahr, A. Hofmann, C. Wienert and F. Döpper Laser Metal Deposition (LMD) Toolpaths with Adaptive Capability for Complex Repairs and Coating Geometries I. Ortiz, P. Alvarez and M.A. Montealegre Innovative Method of Embedding Optical Fiber inside Titanium Alloy Utilizing Friction Stir Forming H.M. Tabatabaei, T. Ohashi and T. Nishihara Towards a Model-Based Approach for the Optimization of the Mechanical and Economical Properties of Laser-Based Plastic-Metal Joints J.M. Berges, K. van der Straeten, G. Jacobs and J. Berroth Center Line Globular Structure of Strip Cast Using High Speed Twin Roll Caster T. Haga Investigation of End Milling Process by Machine Tool Equipped with an Idling-Stop Function in a Feed Driving System Y. Mizuguchi, T. Hirogaki and E. Aoyama Finite Element Modeling of Upper Ball Joint in a Two-Step Hot Forging Process N. Siripath, S. Suranuntchai and S. Sucharitpwatskul Casting of Wire Using Twin Wheel Caster with Melt Holding Nozzle T. Haga, N. Okuda, H. Watari and S. Nishida

49 59 67 75 81 87 95 103

Chapter 3: Technologies of Welding Production Optimization of Resistance Spot Welding Parameters in a Shop Floor Environment to Achieve Desired Spot Size in Low Carbon Steel Sheet J. Bagali, N.V. Nanjundaradhya and R.S. Sharma Optimization of Process Variables for Prediction of Penetration Depth of HSLA Steel Welds Using Response Surface Methodology D. Pathak, D. Kumar, R.P. Singh and V. Balu Microstructural and Mechanical Properties of Dissimilar Aluminium Alloys by Friction Stir Welding K. Sekar and P. Vasanthakumar

111 119 129

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Analysis of Thermal Fields, Weld Strength and Microstructural Studies of Friction Stir Dissimilar Weldments of AA6082 and AA7075 M.V.A. Ramakrishna and K. Srinivas Mechanical and Metallurgical Characteristics of Rotary Friction Welded Low Carbon Steel Plate/Rod Joints T. Dhamotharakannan, P. Sivaraj, M. Seeman and V. Balasubramanian

139 153

Chapter 4: Building Materials Pervious Concrete Utilized as a Base Slab of Pavement Edge Drains J. Endawati and A. Febriansya Experimental Study on Multi-Segmented Circular or Non-Circular Pipes to Determine the Pipe Stiffness D. Chelot and P. Upadhyaya Coal Ash and Rice Husk Ash Binder for Manufacturing Hollow Blocks Z.S. Culilang Flexural Behaviors of Modified Recycled Aggregate Concrete Reinforced with Pultruded GFRP S. Ismail and M. Ramli

163 171 179 189

CHAPTER 1: Materials for Electronics

Key Engineering Materials ISSN: 1662-9795, Vol. 934, pp 3-13 doi:10.4028/p-5x10ni © 2022 Trans Tech Publications Ltd, Switzerland

Submitted: 2022-01-05 Revised: 2022-06-09 Accepted: 2022-06-13 Online: 2022-11-28

Optical and Electrical Characterization of Polymer Dispersed Nematic Liquid Crystals Santosh Mani1,2,a*, Madhavi Pradhan2,3,b, Pushpendra Rai1,c, Samriti Khosla4,d, Pradip Sarawade2,e K.J. Somaiya College of Engineering, Somaiya Vidyavihar University, Mumbai, India

1

Department of Physics, Univesity of Mumbai, Santacruz (E), Mumbai, India

2

Department of Physics, Gurunanak Khalsa College, Matunga, Mumbai, India

3

Department of Physics, DAV University, Jalandhar, India

4

[email protected], [email protected], [email protected], c [email protected], [email protected], [email protected] a*

Keywords: Polymer, Nematic Liquid Crystal, Polarizing Optical Microscopy, Fabry Perot Scattering Studies, Dielectric

Abstract. Polymer dispersed liquid crystals are composite functional materials having a variety of application ranging from display to smart window. These materials consist of liquid crystal in which micron size droplets of polymer is dispersed. In the present study the effect of different concentration of polymer 2-ethyl hexyl acrylate on optical and electrical properties of nematic liquid crystal 4-cynophenyl 4-n-hexyl benzoate were investigated by various techniques. The investigation of textures at different temperatures was performed by polarizing optical microscopy for the determination of phase transition temperature. The fabry perot scattering studies using low powered laser beam was used for the confirmation of phase transition temperature. Various textures were found according to the orientation of liquid crystal and polymer. The optical and electrical properties of pure liquid crystal were found to enhance after dispersing polymer. Our investigation suggest that after dispersing polymer into nematic liquid crystal, the material shows more stability, less flicking and sticking of image for display applications. Introduction The nematic liquid crystal (NLC) is one of the important phase commonly used for flat panel display and photonic devices due its simple structure. This phase exist for large temperature range and have uniaxial macroscopic properties for homogeneous NLC. It is found that the plane of polarized light becomes perpendicular when it passes though NLC without applying electric field. This is also useful for magnetic resonance applications because the molecules when dissolved in NLC provide an extremely determined spectrum. The analysis of this spectrum provides information about chemical shift, bond angle, bond length, relaxation process, dipole interaction, and spin coupling. In this phase, the molecules have no positional order but they tend to align along the direction of the director. However the molecules do not maintain long range positional order with respect to each in this phase. Due to the anisotropic behaviour, they can be reoriented by the external stimuli like heat, electric field and magnetic field [1-5]. However, the small ions present in NLC act as barrier when driven by electric field. These ions act as impurity and responsible for instability which lead to many drawbacks like slow response, flicking and sticking of image etc. Therefore for the proper application of NLC, the effect of impurities should be minimized [6-8]. This problem can be overcome by dispersing polymer or nanoparticles to NLCs. When a polymer is dispersed into NLC then it becomes more stable and gives fast response to external stimuli. The polymer dispersed liquid crystal (PDLC) has many advantages in comparison to pure liquid crystals like in twisted NLC display, the liquid crystal is placed between two polarizers kept perpendicular to each other. The PDLC can work without polarizers as no surface alignment treatment is required for their production. For these reasons, the

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PDLC have potential applications in electro-optical displays. In addition to this the physical properties are also enhanced after dispersing polymer into NLC. The proper selection of polymer, its concentration and the compatibility plays a very important role on its optical and electrical properties. When a suitable polymer of proper concentration is dispersed in NLC, then phase separation occurs along a line in the space of temperature-composition. If the process of polymerization is done in the homogeneous state, then increase in the temperature of the phase separation line is found due to growth of the polymer matrix. The first requirement to prepare good quality PDLC is matching of refractive index and second the NLC should not be dissoluble in the given polymer [9-15]. The investigation of various physical properties of PDLC and their possible potential applications were performed by many researcher groups. Mikhail et al. studied the electrically controllable polarizers using polymer dispersed nematic liquid crystal [16], E. P. Pozhidaev et al. studied the electrically controlled light scattering using PDLC in the visible and near-infrared ranges[17], Zemin He et al. fabricated viewing-angle-switching film based on PDLC and investigated its optical properties[18], Zeng Liang et al. prepared ZnO nanoparticles doped PDLC from a mixture of UV curable liquid polymer and investigated morphological and electro-optical properties [19], Haonan Lin et al. investigated the morphologies, electro-optical and mechanical properties of methacrylate monomers on PDLC [20], We have investigated optical and thermal properties of PDLC [21]. Many other researchers investigated the effect of polymer on NLC but as per our literature survey the effect of different concentration of polymer on optical and electrical properties of NLC is not yet investigated [22-47]. The main objective of the present work is to study the optical and electrical properties of NLC dispersed with varying concentration of polymer. The detailed morphology of NLC and PDLC samples were studied at different temperatures. The phase transition temperature, dielectric constant and conductivity due to dielectric loss of PDLC were investigated and compared with NLC. Materials and Method A pure NLC namely 4-cynophenyl 4-n-hexyl benzoate (Solid → Nematic → Isotropic) purchased from Sigma-Aldrich and used as a base material. Then the NLC was dispersed with 2-Ethyl Hexyl Acrylate (2-EHA) which is a functional monomer. Higher concentration of this monomer makes polymer network sturdier which is useful for reducing response time and suppressing hysteresis. It increases operating voltage substantially. The pure NLC, 4-cynophenyl 4-n-hexyl benzoate is designated as NLC. The two different proportions of 2-EHA by weight were dispersed to NLC and composites were synthesized by chemical method in the laboratory. The composite of concentration A (80% NLC and 20% 2-EHA) and BB (60% NLC+ 40% 2-EHA) was characterised by various characterization techniques. The study of optical textures and investigation of phase transition temperature were performed by Carl Zeiss polarizing optical microscope (POM) attached with adjustable hot stage. The phase transition temperatures were confirmed by fabry perot scattering studies using low powered laser beam. The dielectric constant and conductivity due to dielectric loss of PDLC were also investigated for potential display applications. Result and Discussion Polarizing Optical Microscopy The polarizing optical microscopy (POM) is used for the identification of mesophases from the optical textures. However, this technique is also essential to evaluate the physical properties of PDLC in certain phases and over particular temperature range. The identification of mesophase through POM usually involves the magnified view of a thin sample of mesogenic material sandwiched between a glass microscope slide and glass cover slip.

Key Engineering Materials Vol. 934

(a) Smectic A at 340C

(c) Nematic marbled texture at 430C

5

(b) SE Speeulitic region at 40.20C

(d) Melting at 45.60C

Figure 1: Textures of NLC The textures of NLC had shown in Fig. 1 exhibit Smectic A and nematic phases. Their isotropic temperature is 47.50C, when cooling down from isotropic liquid a nematic texture appears at 43.50C. Upon further cooling the nematic textures turn into fan like texture at 38.50C and at 34.50C Smectic A phase is appeared. The material exhibit mesophases of nematic phase.

(a) Broken focal conic at 30.50C

(c) Nematic marbled at 47.20C

(b) SE with sperulitic at 430C

(d) Spherulitic at 520C

Figure 2: Textures of AA

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The texture of concentration of AA is shown in figure 2. The isotropic temperature is obtained at 750C, the nematic texture is observed at 47.20C and fan like texture is appeared at 30.50C.

(a) Broken focal conic texture at 290C

(c) Nematic with oil droplet at 480C

(b) Cholesteric with oil streaks at 630C

(d) Nematic schlieren at 550C

Figure 3: Textures of BB The texture of concentration of BB is shown in Fig. 3. Isotropic temperature for composite BB is obtained at 850C, the nematic texture is observed at 480C and fan like texture appeared at 29.20C for cooling cycle. It also shows broken focal conic texture. The layers are arranged in Dupin cyclides. The characteristic ellipses are visible in the broken polygonal texture and branches of hyperbola appeared in broken fan shaped texture. Fabry-Perot Scattering Studies The fabry perot scattering studies (FPSS) of all samples under investigation were performed using laser of power 2mW. The phase transition temperatures (PTTs) for all the samples at various heating and cooling cycles were investigated. The optical technique of measuring the diameter of fabry-perot rings obtained from fabry-perot interferometer coupled with a spectrometer has been used to determine the PTTs of the sample. The diameter of fabry-perot rings were recorded for all the samples as a function of temperature and plotted as shown in the following graphs. The NLC is heated electrically from room temperature upto 1400C and graph of temperature against diameter is shown in Fig. 4. The change in the diameter indicates transition temperature, which is shown by peaks in the graph. Diameters of five different fringes are shown indicating transition at 380C, 420C, 440C, and 620C. Composite AA is heated from room temperature to 1200C and graph of temperature against diameter is shown in Fig. 5. The changes in the intensities were found at some transition temperatures. Changes in the diameter at transition were found at 470C, 510C, 540C, 620C, 700C, and 820C.

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350 D i a m e t e r

300

Series1

250

Series2

200

Series3

150

Series4

100

Series5

50 0

30

40

50

60

70

80

Temperature (0C)

Figure 4: Graph of temperature vs. diameter for NLC 300 D i a m e t e r

Series3

250

Series4 Series5

200

Series6

150 100 50 0

40

50

60

70

80 Temperature

90

100

110

120

(0C)

Figure 5: Graph of temperature vs. diameter for AA

D i a m e t e r

300

Series2

250

Series3 Series4

200

Series5

150

Series6

100 50 0

40

50

60

70 Temperature

80

90

100

110

(0C)

Figure 6: Graph of temperature vs. diameter for BB The sample of concentration BB is heated upto 120oC and graph of temperature against diameter is shown in Fig.6. As temperature increases intensity of fringe pattern decreases and transition were observed at 410C, 47 0C, 550C, 640C, 780C and 920C.The table 1 below shows the comparison of PTTs of different concentrations obtained by POM and FPSS.

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Table 1: Comparison of PTTs Techniques POM FPSS

PTTs of NLC (0C) 37.8, 42.3, 47.5, 62 38, 42, 44, 48, 62

PTTs of AA (0C) 47.2, 51.1, 62.3, 70.1, 74.9 47, 51, 54, 62, 70, 75, 82

PTTs of BB (0C) 41.1, 47.5, 55.1, 63.9, 78 41, 47, 55, 64, 78, 92

Electrical Properties The dielectric constant were measured by LCR impedence analyser for different frequencies by applying AC voltage at constant amplitude. Influence of polymer on dielectric parameters such as capacitance, dielectric loss impedence has been recorded in the frequency range of 20 Hz to 20 MHz. The dielectric relaxation measurement at different frequency gives information about the dynamics of polar groups and modular motion. The relative dielectric consant is obtained from the measuremnet of the parallel capacitance. The Cole-Cole graph of NLC and PDLC composite of concentration AA and BB is shown in the Fig. 7 below. 1.0

(a)

0.8

ε"

0.6 0.4 0.2 0.0 0.016

1.0

(b)

ε'

1.5

0.004 0.000

2.0 0.006

2.5

(c)

0.004

0.008

ε"

ε"

0.012

0.5

0.002

0.000 -0.024

-0.016

ε'

-0.008

0.000

0.008

-0.027

-0.024

-0.021

ε'

-0.018

-0.015

Figure 7: Cole-Cole Plot for NLC, concentration AA and BB The Cole-Cole plot for pure NLC shown in Fig. 7 (a) gives values for ∈| ranging between 0.5 to 2.5 and maximum dielectric loss that is ∈|| shows value of 0.9. The graph shown in Fig.7(b) for composite AA, shows negative anisotropy having real value of dielectric constant, ranging from -0.024 to +0.005. The maximum values of imaginary part of dielectric constant have value 0.014. Composite AA shows anisotropy ranging from negative value to positive value. The graph in Fig. 7(c) shows Cole-Cole plot of composite BB. It shows negative anisotropy values from -0.027 to -0.015 and maximum value of dielectric loss is 0.006. Conductivity due to dielectric losses The conductivity (σ) for all the samples under investigations were calculated using following formula, σ = ω εo ε' tan δ, where ω is frequency, εo permittivity of free space, tan δ is dilectric loss. The conductivity graph of pure NLC, concentration AA and BB is shown in Fig. 7 (a), 7 (b) and 7(c) respectively in the same frequency range. The conductivities were found to increase for the concentration BB. The graph shows same nature for all the three samples but indicate change in the value of conductivity in lower frequency range for composite BB.

Key Engineering Materials Vol. 934

8.0x10-2 6.0x10-2

9

(a)

4.0x10-2

Conductivity(mho)

2.0x10-2 0.0 4.5x10-4 3.6x10-4 2.7x10-4 1.8x10-4 9.0x10-5 0.0

(b)

1.2x10-3

(c)

9.0x10-4 6.0x10-4 3.0x10-4 0.0

0

200 400 600 800 1000

Frequency (Hz)

Figure 7: Conductivity of pure NLC and concentration AA and BB Conclusions The PDLC of different concentrations were prepared by dispersing monomer 2-EHA into NLC 4-cynophenyl 4-n-hexyl benzoate by chemical method. The textures obtained using POM for several cycles of heating and cooling shows various mesophases. The PTTs of all samples were investigated by POM and the same was confirmed by FPSS. The PTTs were found to increase after dispersing polymer into NLC. The electrical investigation shows fast switching for PDLC which makes them useful in the field of display. The negative anisotropy of PDLCs shows that the lateral polar substitutent induces a dipole moment perpendicular to the principle molecular axis. The different off axis polar group cyano and fluoro have been employed to enlarge perpendicular dipole moment to avoid image flickering high resistivity which can be used for obtaining a high voltage – holding – ratio for active matrix PDLC displays. Acknowledgment We wish to express our sincere thanks to Prof. (Dr.) V. N. Rajasekharan Pillai, Vice-Chancellor of Somaiya Vidyavihar University and Dr. Shubha Pandit, Principal, K. J. Somaiya College of Engineering, for encouragement and support towards this research work.

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References [1] De Gennes, Pierre-Gilles, and Jacques Prost. The physics of liquid crystals. No. 83. Oxford university press, 1993. [2]

S. Chandrasekhar, Liquid Crystals, 2nd ed. Cambridge University Press, Cambridge, 1994.

[3]

Demus, D., et al. "Handbook of liquid crystals set." 2008.

[4] Collings, Peter J. Liquid crystals: nature's delicate phase of matter. Princeton University Press, 2002. [5] Mani, Santosh, et al. "Morphological and Thermal Behaviour of Monomer Dispersed Liquid Crystal." Proceedings of Fourth International Conference on Inventive Material Science Applications. Springer, Singapore, 2022. https://doi.org/10.1007/978-981-16-4321-7_58 [6] P. J, Jessy, and Nainesh Patel. "Highly improved dielectric behaviour of ferronematic nanocomposite for display application." Liquid Crystals 46.5 (2019): 772-786. https://doi.org/10.1080/02678292.2018.1528642 [7] Draude, Adam Paul, and Ingo Dierking. "Thermotropic liquid crystals with low-dimensional carbon allotropes." Nano Express (2021). https://doi.org/10.1088/2632-959X/abdf2d [8] Jayoti, Divya, Praveen Malik, and Arshdeep Singh. "Analysis of morphological behaviour and electro-optical properties of silica nanoparticles doped polymer dispersed liquid crystal composites." Journal of Molecular Liquids 225 (2017): 456-461. https://doi.org/10.1016/j.molliq. 2016.11.100 [9] White, Timothy J., et al. "Contribution of monomer functionality and additives to polymerization kinetics and liquid crystal phase separation in acrylate‐based polymer‐dispersed liquid crystals (PDLCs)." Liquid Crystals 34.12 (2007): 1377-1385. https://doi.org/10.1080/ 02678290701663936 [10] Sun, Yujian, et al. "A study on the electro-optical properties of thiol-ene polymer dispersed cholesteric liquid crystal (PDChLC) films." Molecules 22.2 (2017): 317. https://doi.org/10.3390/molecules22020317. [11] Manda, Ramesh, et al. "Effect of monomer concentration and functionality on electrooptical properties of polymer-stabilised optically isotropic liquid crystals." Liquid Crystals 45.5 (2018): 736-745. https://doi.org/10.1080/02678292.2017.1380239 [12] Bashtyk, Y., et al. "Primary converters for optical sensors of physical values based on polymer dispersed cholesteric liquid crystal." Molecular Crystals and Liquid Crystals 642.1 (2017): 41-https://doi.org/10.1080/15421406.2016.1254509 [13] Mani, Santosh A., et al. "Investigations of optical and thermal response of polymer dispersed binary liquid crystals." Molecular Crystals and Liquid Crystals 646.1 (2017): 183-193. https://doi.org/10.1080/15421406.2017.1287478 [14] Jain, Anuja Katariya, and R. Deshmukh. "An overview of polymer-dispersed liquid crystals composite films and their applications." Liq. Cryst. Disp. Technol (2020): 1-68 https://doi.org/10.5772/intechopen.91889 [15] Sharma, Vandna, and Pankaj Kumar. "Electro-optically oriented Kerr and orientational phase study of normal mode polymer dispersed liquid crystals–Effect of dispersion of nanoparticles." Journal of Molecular Liquids (2021): 118030. https://doi.org/10.1016/j.molliq.2021.118030 [16] Krakhalev, Mikhail N., et al. "Polymer dispersed nematic liquid crystal films with conical boundary conditions for electrically controllable polarizers." Optical Materials 89 (2019): 1-4. https://doi.org/10.1016/j.optmat.2019.01.004

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[17] Pozhidaev, E. P., et al. "Polymer dispersed liquid crystals with electrically controlled light scattering in the visible and near-infrared ranges." Optical Materials Express 10.12 (2020): 30303040.https://doi.org/10.1364/OME.410163 [18] He, Zemin, et al. "Viewing-angle-switching film based on polymer dispersed liquid crystals for smart anti-peeping liquid crystal display." Liquid Crystals (2021): 1-7. https://doi.org/10.1080/02678292.2021.1944357 [19] Liang, Zeng, et al. "Influence of ZnO NPs on morphological and electro-optical properties of polymer-dispersed liquid crystals." Liquid Crystals (2021): 1-10. https://doi.org/10.1080/02678292.2021.1898055. [20] Lin, Haonan, et al. "Effects of the methacrylate monomers with different end groups on the morphologies, electro-optical and mechanical properties of polymer dispersed liquid crystals composite films." Liquid Crystals 48.5 (2021): 722-734. https://doi.org/10.1080/02678292.2020.1815091 [21] Mani, Santosh, et al. "The influence of polymer on optical and thermal properties of nematic liquid crystals." Journal of Physics: Conference Series. Vol. 2070. No. 1. IOP Publishing, 2021. https://doi.org/10.1088/1742-6596/2070/1/012055 [22] Ziqian He, Kun Yin, and Shin-Tson Wu, Passive polymer-dispersed liquid crystal enabled multi-focal plane displays Optics Express, (2020), Vol. 28, Issue 10, 15294-15299, https://doi.org/10.1364/OE.392489 [23] Saeed, Mohsin H.; Zhang, Shuaifeng; Cao, Yaping; Zhou, Le; Hu, Junmei; Muhammad, Imran; Xiao, Jiumei; Zhang, Lanying; Yang, Huai, Recent Advances in The Polymer Dispersed Liquid Crystal Composite and Its Applications (2020), Molecules 25, no. 23: 5510. https://doi.org/10.3390/molecules25235510 [24] Zhao, Xiaoshuai Li, Kemeng Wang, He Huai, Hongmei Ma & Yubao Sun, Effect of the introduction of mono-functional monomer on the electro-optic properties of reverse-mode polymer stabilised cholesteric liquid crystal, (2020) Liquid Crystals https://doi.org/10.1080/02678292.2020.1849835 [25] Li Jinqian, Yuzhen Zhao, Hong Gao, Dong Wang, Zongcheng Miao, Hui Cao, Zhou Yang & Wanli He (2021) Polymer dispersed liquid crystals doped with CeO2 nanoparticles for the smart window, Liquid Crystals, https://doi.org/10.1080/02678292.2021.1942573. [26] Mishra, Krishnakant G., et al. "Comparative study of nanoparticles doped in liquid crystal polymer system." Journal of Molecular Liquids 224 (2016): 668-671. https://doi.org/10.1016/j.molliq.2016.10.075 [27] Shumeng Guo, Xiao Liang, Huimin Zhang, Wenbo Shen, Chunxin Li, Xiao Wang, Cuihong Zhang, Lanying Zhang & Huai Yang,An electrically light-transmittance-controllable film witha low-driving voltage from a coexistent system of polymer-dispersed and polymer-stabilised cholesteric liquid crystals, (2018), Liquid Crystals, 45:12, 1854-1860, https://doi.org/10.1080/02678292.2018.1501820 [28] Li, Chen-Yue; Wang, Xiao; Liang, Xiao; Sun, Jian; Li, Chun-Xin; Zhang, Shuai-Feng; Zhang, Lan-Ying; Zhang, Hai-Quan; Yang, Huai. Electro-Optical Properties of a Polymer Dispersed and Stabilized Cholesteric Liquid Crystals System Constructed by a Stepwise UVInitiated Radical/Cationic Polymerization (2019) Crystals 9, no. 6: 282. https://doi.org/10.3390/cryst9060282 [29] Daniela Ailincai, Daniela Pamfil, Luminita Marin,Multiple, Bio-responsive polymer dispersed liquid crystal composites for sensing applications, Journal of Molecular Liquids (2018), Volume 272, 572-582, https://doi.org/10.1016/j.molliq.2018.09.125.

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[30] Jong-Min Baek, Seung-Won Oh, Sang-Hyeok Kim, Tae-Hoon Yoon, Fabrication of an initiallyfocal- conic cholesteric liquid crystal cell without polymer stabilization, Displays, (2018) Volume 52, 55-58, https://doi.org/10.1016/j.displa.2017.10.002. [31] Gardymova, Anna P.; Krakhalev, Mikhail N.; Zyryanov, Victor Y.; Gruzdenko, Alexandra A.; Alekseev, Andrey A.; Rudyak, Vladimir Y. Polymer Dispersed Cholesteric Liquid Crystals with a Toroidal Director Configuration under an Electric Field (2021), Polymers 13, no. 5:732. https://doi.org/10.3390/polym13050732 [32] Mishra, Krishnakant K., Sheshmani K. Dubey, and Santosh A. Mani. "Optical characterization of inorganic nanoparticles doped in polymer dispersed liquid crystal." Molecular Crystals and Liquid Crystals 647.1 (2017): 244-252. https://doi.org/10.1080/15421406.2017. 1289603 [33] Sun, Yujian; Gao, Yanzi; Zhou, Le; Huang, Jianhua; Fang, Hua; Ma, Haipeng; Zhang, Yi; Yang, Jie; Song, Ping; Zhang, Cuihong; Zhang, Lanying; Li, Fasheng; Zhao, Yuzhen; Li, Kexuan. A Study on the Electro-Optical Properties of Thiol-Ene Polymer Dispersed Cholesteric Liquid Crystal (PDChLC) Films (2017). Molecules 22, no. 2: 317. https://doi.org/10.3390/molecules22020317. [34] Mishra, Krishnakant K, S. K. Dubey, Santosh A. Mani, Madhavi S. Pradhan. Comparative study of nanoparticles doped in Liquid Crystal Polymer System. Journal of Molecular Liquids, (2016), 224: 668-671. http://dx.doi.org/10.1016/j.molliq.2016.10.075 [35] Y. Bashtyk, O. Bojko, A. Fechan, P. Grzyb & P. Turyk, Primary converters for optical sensors of physical values based on polymer dispersed cholesteric liquid crystal, (2017) Molecular Crystals and Liquid Crystals, 642:1, 41-46, https://doi.org/10.1080/15421406.2016.1254509. [36] Hongqi Gao, Shuaifeng Zhang, Mohsin Hassan Saeed, Gang Chen, Haonan Lin, Junyi Huang, Lanying Zhang, Qian Wang & Hui Cao, Study on the morphologies and electro-optical properties of cyano-phenyl-ester liquid crystals/polymer composite films prepared by stepwise polymerisation, (2020), Liquid Crystals, 47:10, 1497-1506, https://doi.org/10.1080/02678292.2020.1737976. [37] Sharma, Vandna, Pankaj Kumar, and Kuldeep Kumar Raina. "Simultaneous effects of external stimuli on preparation and performance parameters of normally transparent reverse mode polymer-dispersed liquid crystals—a review." Journal of Materials Science 56.34 (2021): 1879518836. https://doi.org/10.1007/s10853-021-06489-7 [38] Mhatre, Manoj M., Anuja Katariya-Jain, and Rajendra R. Deshmukh. "Enhancing morphological, electro-optical and dielectric properties of polymer-dispersed liquid crystal by doping of disperse Orange 25 dye in LC E7." Liquid Crystals (2021): 1-14. https://doi.org/10.1080/02678292.2021.2007548 [39] De Filpo, Giovanni, et al. "Order parameter and electro-optical properties in polymerdispersed liquid crystals." Liquid Crystals 48.8 (2021): 1206-1214. https://doi.org/10.1080/02678292.2020.1851416 [40] Kumari, Asha, et al. "Pseudopeptidic Polymer Microsphere-Filled Liquid Crystals as HighPerformance Light-Scattering Switches." ACS Applied Polymer Materials (2021). https://doi.org/10.1021/acsapm.1c00945 [41] Meng, Xiangshen, et al. "Polymer dispersed liquid crystals doped with low concentration γFe2O3 nanoparticles." Liquid Crystals (2021): 1-15. https://doi.org/10.1080/02678292.2020.1852620 [42] Ramanitra, H., et al. "Application of polymer dispersed liquid crystal (PDLC) nematic: optical-fiber variable attenuator." Mol. Cryst. Liq. Cryst. 404.1 (2003): 57-73. https://doi.org/10.1080/15421400390249952

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[43] Higgins, Daniel A., Jeffrey E. Hall, and Aifang Xie. "Optical microscopy studies of dynamics within individual polymer-dispersed liquid crystal droplets." Accounts of chemical research 38.2 (2005): 137-145. https://doi.org/10.1021/ar040106p [44] Serbutoviez, C., et al. "Polymerization-induced phase separation. 2. Morphology of polymer-dispersed liquid crystal thin films." Macromolecules 29.24 (1996): 7690-7698. https://doi.org/10.1021/ma960293+ [45] Drzaic, Paul S. "Polymer dispersed nematic liquid crystal for large area displays and light valves." Journal of applied physics 60.6 (1986): 2142-2148. https://doi.org/10.1063/1.337167 [46] Ondris‐Crawford, Renate, et al. "Microscope textures of nematic droplets in polymer dispersed liquid crystals." Journal of applied physics 69.9 (1991): 6380-6386. https://doi.org/10.1063/1.348840 [47] Coates, David. "Polymer-dispersed liquid crystals." Journal of Materials Chemistry 5.12 (1995): 2063-2072. https://doi.org/10.1039/JM9950502063

Key Engineering Materials ISSN: 1662-9795, Vol. 934, pp 15-22 doi:10.4028/p-i221xc © 2022 Trans Tech Publications Ltd, Switzerland

Submitted: 2022-02-02 Revised: 2022-06-07 Accepted: 2022-06-08 Online: 2022-11-28

Suppress Short Channel Effects on Split Channel-Cylindrical GAA TFET Using Buried Oxide Layer Pratiksha Dhake1,a*, Jyotirmoy Ghosh1,b, Mayuresh Joshi1,c, Ribu Mathew2,d and Ankur Beohar2,e MTech scholar, School of Electrical & Electronics Engineering (SEEE) VIT Bhopal University Bhopal, India

1

Assistant Professor, School of Electrical & Electronics Engineering (SEEE) VIT Bhopal University Bhopal, India

2

[email protected], [email protected] ,c [email protected] d [email protected] , [email protected]

a

Keywords TFET, Uniform Doping (UD), Buried oxide layer(BOx), InP based TFET

Abstract The idea of a buried oxide layer (BOx) in a split channel Gate All Around-Tunnel Field Effect Transistor (GAA-TFET) is investigated in this paper. This work examined the impact of buried oxide layer on the device's performance. With the BOx layer present and channel length of 20nm, the channel area of the TFET device investigated in this study is divided equally on the same side. The doping concentration has been transferred to the split channel on the drain side. The device’s performance is examined using numerical simulation utilizing simulation software of CAD devices. The final results which incorporate the buried oxide layer is compared to the uniform split channel GAA-TFET. The parameters like ON current (Ion),OFF current (Ioff), subthreshold swing (SS) and electric-field (E) intensity are observed and compared with silicon (Si) based GAA-TFET and Indium phosphide (InP) based GAA-TFET. It is found that InP based GAA-TFET with buried oxide layer is more advanced device design than the others with Ion and Ioff of 3.02 x 10-05 A/um and 2.09 x 10-22 A/um, respectively. Introduction The benefits of continuous scaling of field-effect transistor (FET) devices have been constrained by the limits imposed by short channel effects (SCEs). SCEs have hindered the scaling of traditional MOSFETs beyond 100 nm length, and considerable research into alternate devices is underway. As a result, device performance has suffered significantly. The gate-all-around (GAA) TFET device are being used to circumvent the limitations of SCEs. [1-3]. GAA-TFET provide a number of advantages over traditional MOSFETs, including a lower OFF-state current, a reduced sensitivity to SCEs, and due to low SS, it has higher running speed [4 -7]. Beside these shortcomings, regular TFETs perform poorly in the DC and RF domains, and physically doped TFETs exhibit random dopant oscillations (RDFs) [8-9] Various approaches have been described to overcome SCEs. The splitting technique is one of the techniques. In Junction Field Effect Transistors (J-FETs), Sturm-Rogon et al. employed the concept of split channel to increase the FET's radio frequency (RF) performance [10]. Similarly, Beohar et al. studied the influence of a cylindrical GAA-TFET with a Germanium (Ge) source area on analog or radio-frequency (RF) properties. They developed the drain underlapping approach and used a high-constant dielectric material across the drain region, which minimized fringing field effects and decreased leakage current (Ioff) [11-12].Moreover, several TFET designs have been proposed to increase ON current, such as DG TFET [13], hetero structures [14], TFET of strained silicon [15] and so on. M.Joshi investigated the concept of channel splitting in GAA-TFET of 20nm channel length. The subthreshold swing is likewise shown to be decreasing at 33.14 mV/decade. 0.092V was discovered to be the threshold voltage[16] This work presents a uniform doped GAA-TFET based on InP with a split channel with the incorporation of the buried oxide (BOx) layer and SiO2 oxide layer (InP-SiO2-UD-BOx). The effect of a BOx layer by dividing channel and evenly doping on the performance of GAA-TFETs is

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numerically investigated using device simulation tools. The goal is to achieve low off current (IOFF) and raise on current (ION), therefore reducing the device's leakage current. At the same technology node, the results were compared to a uniform split channel GAA-TFET based on InP with HfO2 oxide layer (InP-HfO2-UD) device and uniform doped split channel GAA-TFET based on Si with HfO2 oxide layer (Si-HfO2-UD) device. Simulation Process The proposed device is a cylindrical gate all around TFET with uniform doping and buried oxide layer. The device has uniform doping in both the source region as well as drain region, with doping concentration moved to the drain side. In addition to this, the oxide layer under gate is of high k dielectric and oxide layer to right and left of drain and source respectively is of low k dielectric. The device features that were employed to model the InP-SiO2-UD-BOx device are as follows (Table1) Table 1. Device features employed to InP-SiO2-UD-BOx S.No Characteristics

Value

1.

Source.doping concentration (p-type)

1.0 x 1020cm-3

2.

Drain.doping concentration (n-type)

5 x 1016cm-3

3.

Oxide thickness(tox)

1nm

4.

Channel.length (Lchannel)

20nm

5.

Radii

1nm

6.

Gate work function

4.63eV

7.

Extension.length.of source.(Lextss).and.drain (Lextdd)

40nm

The source and drain regions have low k dielectric oxide (SiO2) material whereas high k oxide (HfO2) material is incorporated over channel region. TCAD Synopsys® has been used to model and simulate the device. Figures 1 (a) and 2 (a) depict the proposed device's 3D design and cross-sectional view, accordingly.

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(a)

(b) Fig.1. (a) 3-Dimensional view of Cylindrical-Gate All Around (Cyl.-GAA) with buried oxide(BOx) layer , (b) 3-D view of buried oxide layer in proposed structure

(a)

(b) Fig.2. (a) A cross-sectional view of Cyl. GAA with BOx layer and uniform doped TFET (b) A cross-sectional view of InP-SiO2-UD-BOx device InP is composed of indium and phosphorus, therefore categorized as binary semiconductor.As InP has higher electron velocity compared to other semiconductors, it is utilized in high-power and high-

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frequency devices. Some transistors like High Electron Mobility Transistors (HEMT) or Heterostructure Bipolar Transistors (HBT) use InP microelectronics. Energy bandgap of InP at 300K is 1.27eV. Moreover the oxides at source and drain side are of low-k dielectric (SiO2) material that helps in reducing the fringing field effects and the oxide under gate is of high-k dielectric (HfO2) , which substantially decreases leakage current. Results and Discussion The impact of different parameters on the performance of InP-SiO2-UD-BOx GAA-TFET is detailed in this section. Impact of buried oxide layer & Channel Splitting on the Ion/Ioff Figure 3 shows the relationship between InP-SiO2-UD-BOx GAA-TFET, Si-HfO2-UD GAATFET & InP-HfO2-UD GAA-TFET. The degradation in off current has been observed high in the case of InP-SiO2-UD-BOx GAA-TFET because of the asymmetric doping across the drain-side which is more and incorporates the buried oxide layer . The on-current ( Ion) of all devices are 3.02 x 10-05 A/um, 1.5 x 10-05 A/um and 5.85 x 10-06 A/um, respectively as the order mentioned above. InP-SiO2-UD-BOx GAA-TFET, on the other hand, exhibits a modest improvement in both off and on current. In InP-SiO2-UD-BOx GAA-TFET, on the other hand, on employing the variable doping profile in the drain area, the depletion barrier width between the channel and drain junction rises, resulting in an improvement in OFF-current. In InP-SiO2-UD-BOx GAA-TFET, the Ion/Ioff ratio reaches to 1.44 x 10-17 and the primary concept behind a buried oxide layer is to minimize parasitic junction capacitance. Furthermore, the lower the parasitic capacitance, the faster the device will operate. There are less undesired leakage pathways that are far from the gate because of the BOx layer. As a result, power consumption is reduced. As a result, when compared Si-HfO2-UD GAATFET & InP-HfO2-UD GAA-TFET with proposed design, the leakage current of InP-SiO2-UD-BOx GAA-TFET, the proposed device is found to be least.

Fig.3. Comparision of Ids vs Vgs transfer characteristics The concentration of doping in the source is higher than in the drain. As a result, the barrier width between drain and channel will be greater than the width between source and channel regions. As a result, the OFF current is lowered. The electric field intensity in the region of source-channel is higher than in the region of drain-channel, as shown in Figure 4. ,which in turn lowers the OFF current.

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Impact of buried oxide layer & shift in doping on the Electric Field

Fig.4. Comparision of Electric Field for the mentioned devices Figure 4 value shows the effect of shift of doping & impact of BOx layer on the electric field of InPSiO2-UD-BOx GAA-TFET, Si-HfO2-UD GAA-TFET & InP-SiO2-UD GAA-TFET. The electricfield-strengths vary slightly. The electric-field intensity (E) in InP-SiO2-UD-BOx GAA-TFET is higher, as shown by the results. The flow of charge from source to drain is what determines the value of E (electric-field-intensity) and we can observe the drain region near the channel-2 has more electric field. As a result, the electric-field intensity in InP-SiO2-UD-BOx GAA-TFET is greater than SiHfO2-UD GAA-TFET & InP-SiO2-UD GAA-TFET. Further, due to channel splitting, the distance between the conduction and valence bands narrows, allowing for a higher ON current (Ion )as seen in fig. 5.

Fig.5. Energy band characteristic of InP-SiO2-UD-BOx GAA-TFET The subthreshold swing(SS) and threshold voltage(Vth) of the device, both are affected by the electric-field intensity. Band to Band tunneling at the channel-source area causes the first peak in the electric field. The high-k dielectric minimizes the sub-threshold swing by reducing the influence of fringing fields between the source and drain channel. In Si-HfO2-UD GAA-TFET & InP-HfO2-UD GAA-TFET, the subthreshold swings are 33.14 mV/decade & 23 mV /decade, while for the InPSiO2-UD-BOx GAA-TFET, it’s value is 14.9 mV/decade.As a result, the threshold voltage of the InP-SiO2-UD-BOx (proposed) device is reduced.

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Fig.6. Comparison of different devices subthreshold swing Table 2 also shows the overall device comparison and exhibits the effectiveness of the InP-SiO2-UDBOx device which is proposed, over the other experimental and simulated devices. The device which is proposed and the device which is manufactured is compared and tabulated in Table 3. Table 2. Comparison of different parameters of different devices Device Si-HfO2-UD InP-HfO2-UD InP-SiO2-UD-BOx

Ion (A/um) 1.5 x 10-05 5.85 x 10-06 3.02 x 10-05

Ioff (A/um) 1.71 x 10-19 4.84 x 10-22 2.09 x 10-22

SS (mV/decade) 33.14 23 14.9

Ion/Ioff 8.77 x 1013 1.20 x 1016 1.44 x 1017

Table 3. Comparison with manufactured devices Device Vertical nanowire Si TFET [17] Si tunnel transistor [18] Heterojunction TFET[19] Si TFET [20] Cylindrical GAAFET [7] Proposed device

Ion (A/µm) 10 × 10-09 1.2 × 10-05 1 × 10-06 1.5 × 10-06 1.64 × 10-06 3.02 × 10-05

Ioff (A/µm) 1.0 × 10-10 1.7 × 10-12 1.0 × 10-14 4.1 × 10-15 3.48 × 10-18 2.09 × 10-22

Ion/Ioff 10 × 103 7.0 × 107 1.0 × 105 3.6 × 105 4.71 × 1012 1.44 × 1017

SS (mV/decade) 75 46 22 36 29 14.9

Conclusion The design of a buried oxide split-channel GAA TFET device for analog/RF applications is described. To compare the performance of the proposed device to that of a uniform channel GAATFET, simulations were conducted using CAD device simulation software. The suggested device exhibits higher performance metrics, according to the results. When compared to the source-side channel, the doping concentration in the drain-side channel is more, resulting in better off-current performance. The off current (2.09 x 10-22 A/µm) has been reduced. The suggested device's subthreshold swing is 14.9 mV/decade with better electric field intensity due to shift of doping concentration and buried oxide layer. In addition, the Ion current is also improved of proposed device, InP-SiO2-UD-BOx GAA-TFET due to splitting the channel i.e. 3.02 × 10-05 (A/µm) .

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Acknowledgment The authors would like to thank Nanoscale Devices, VLSI Circuit and System Design Research Group, IIT Indore, and VIT-Bhopal University for technical assistance. References [1] A Beohar, S.K. Vishvakarma , “Performance enhancement of asymmetrical underlap 3Dcylindrical GAA-TFET with low spacer width”, IET Micro & Nano Letters, 2016, Vol. 11, Iss. 8, pp. 443-445 [2] A. C. Seabaugh and Q. Zhang, “Low-voltage tunnel transistors for beyond CMOS logic,” Proc. IEEE, vol. 98, no. 12, pp. 2095–2110, Dec. 2010. [3] S. O. Koswatta, M. S. Lundstrom, and D. E. Nikonov, “Performance comparison between p-i-n tunneling transistors and conventional MOSFETs,” IEEE Trans. Electron Devices, Vol. 56, No. 3, pp. 456– 465, Mar. 2009. [4] Q. Zhang, W. Zhao, A. Seabaugh, Low-subthreshold-swing tunnel transistors. IEEE Electron Device Lett. 27(4), 297–300 (2006) [5] S.Tiwari, A.Dutt, M.Joshi, P.Nigam, R.Mathew and A. Beohar: “In-silico Investigation of Cyl. Gate all Around (GAA) Tunnel Field Effect Transistor (TFET) Biosensor”, IOP: Materials Science and Engineering (ICMSMT2021),Coimbatore April 2021. [6] A.Dutt, S.Tiwari, M.Joshi, P.Nigam, R.Mathew and A. Beohar: “On-chip Analysis of Etched Drain based Cyl. GAA TFET with Elevated Density Strip”, IOP:Materials Science and Engineering (ICMSMT 2021), Coimbatore,April 2021, [7] Tiwari Sanjana, Arya Dutt, Mayuresh Joshi, Prakhar Nigam, Ribu Mathew, and Ankur Beohar. "An investigation of a suppressed-drain cylindrical gate-all-around retrograde-doped heterospacer steep-density-film tunneling field-effect transistor." Journal of Computational Electronics 20, no. 5 (2021): 1702-1710 [8] A. Asenov, “Simulation of statistical variability in nano MOSFETs,” in VLSI Symp. Tech. Dig., Jun. 2007, pp. 86–87. [9] N. Damrongplasit, C. Shin, S.H. Kim, R.A. Vega, T.J.K. Liu, Study of random dopant fluctuation effects in germanium-source tunnel FETs. IEEE Trans. Electron Devices 58(10), 3541–3548 (2011) [10] L. Sturm-Rogon, K. Neumeier, and C. Kutter “Low-Noise Si-JFETs Enhanced by Split channel Concept”, IEEE Transactions on Electron Devices, Volume-67, Issue-11, Nov-2020. [11] A. Beohar, N. Yadav, A. P. Shah, S. K. Vishwakarma, “Analog/RF characteristics of a 3D Cyl. Underlap GAA-TFET based on a Ge-Source using fringing field engineering for low power applications”, Journal of Computational Electronics, Springer, Volume-17, Dec-2018. [12] A. Beohar, N. Yadav, and S. K. Vishvakarma, “Analysis of Trap Assisted Tunneling in Asymmetrical Underlap 3D-Cylindrical GAA-TFET based on Hetero-Spacer Engineering for Improved Device Reliability,” IET Micro & Nano Letters, vol. 12, no. 12, Dec. 2017, pp. 982-986 (SCI, Impact factor: 0.97). DOI: 10.1049/mnl.2017.0311, ISSN: 1750-0443 [13] K. Boukart. W. Riess, and A. M. Ionescu, “Double-gate tunnel FET with high-K gate dielectric,” IEEE Trans. Electron Devices, vol. 54 no. 7, pp. 1725-1733, Jul.2007. [14] A. S. Verhulst, W.G. Vandenberghe, K. Maex, S. De Gendt, M. M. Heyns and G. Groeseneken,” Complementary silicon-based hetrostructures tunnel-FET with high tunnel rates,” IEEE Electron Device Lett., vol. 29, no. 12, pp. 1398-1401, Dec. 2008. [15] K. Boucart, W. Riess and A. M. Ionescu, “Lateral strain profile as key technology booster for all silicon Tunnel FETs,” IEEE Electron Device Lett., vol. 30, no. 6, pp. 656-658, Jun. 2009.

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[16] M. Joshi, A. Dutt, S. Tiwari, P. Nigam, A. Beohar and R. Mathew, "Impact of Channel Splitting on Gate All Around Tunnel Field Effect Transistor (GAA-TFET)," 2021 Devices for Integrated Circuit (DevIC), 2021, pp. 1-5, doi: 10.1109/DevIC50843.2021.9455931. [17] Vandooren A, Leonelli D, Rooyackers R et al.“Impact of process and geometrical parameters on the electrical characteristics of vertical nanowire silicon n-TFETs”, Solid-State Electronics, Volume 72, 2012. [18] Jeon K “Si tunnel transistors with a novel silicided source and 46mV/dec swing” 2010 Symposium on VLSI Technology, Honolulu, 2010, pp. 121-122 [19] Walke A & Vandooren A, Rooyackers, et al. (2014)” Fabrication and Analysis of a Si/Si0.55Ge0.45 Heterojunction Line Tunnel FET” IEEE Transactions on Electron Devices. [20] Huang Q et al., “A novel Si tunnel FET with 36mV/dec subthreshold slope based on junction depleted-modulation through striped gate configuration” 2012 International Electron Devices Meeting, San Francisco, CA, 2012, pp. 8.5.1-8.5.4

Key Engineering Materials ISSN: 1662-9795, Vol. 934, pp 23-36 doi:10.4028/p-shh9bt © 2022 Trans Tech Publications Ltd, Switzerland

Submitted: 2022-03-09 Revised: 2022-07-20 Accepted: 2022-07-21 Online: 2022-11-28

Spray Pyrolyzed Praseodymium Doped SnO2 Thin Film with Fast Response to LPG: Analysis Based on Microstructural Features Deepa S.1,a*, Prasannakumari K.2,b and Boben Thomas1,c Research centre in Physics, Mar Athanasius College (Autonomous), Kothamangalam College P.O, Kothamangalam, Kerala 686 666, India

1

Mar Athanasius College of Engineering, Kothamangalam College P.O, Kothamangalam, Kerala 686 666, India

2

[email protected], [email protected], [email protected]

a

Keywords: SnO2, spray pyrolysis, LPG sensor, thin film, oxygen. Praseodymium.

Abstract. Nebulizer aided spray deposition technique is employed for obtaining Praseodymium (Pr) modified (0.1 to 6 wt. %) tin oxide thin film in nano configuration at a deposition temperature of 320 ℃. Investigation of the sample proclaims preferential orientation along (110), (301), and (310) planes, and the developed films are tested against 500 ppm of LPG at different operating temperatures. Creditable sensor response of 99 % with quick response time and recovery time of 9 s and 11 s respectively is obtained for 1wt.% Pr doped SnO2 film at an operating temperature of 350°C, which is appreciable compared to the pristine SnO2 film. The sensor response reduces at lower operating temperatures. The sensing mechanism of 1 wt.% Pr doped SnO2 thin film is explained through microstructural, morphological as well as optical studies. The presence of oxygen vacancies and traps evidenced from Raman and photoluminescence analysis have a crucial role in sensor response. Introduction Tin Oxide (SnO2) based gas sensors offer several advantages in comparison to other gas sensing materials [1]. Altering the structure and grain size by using a suitable dopant can improve many features of SnO2 nanostructured thin film. In this context, rare earth metals are appropriate candidates as they are enhancing catalytic activity, directly involved in oxidation, hence can modify the microstructural features including morphology as well as surface defects which are decisive factors in gas sensing action [2,3,4]. Praseodymium (Pr) with atomic number 59, ionic radius 101.3 pm and electronic configurtion [Xe] 4f3 6s2 has oxidation states of 3 and 4 and it has remarkable catalytic activity [5]. The structural and surface study of praseodymium-doped SnO2 nanoparticles synthesized via the polymeric precursor method (6) is outlined by Aragón et al. Research work from this lab has already reported the gas sensing features of rare-earth doped SnO2 thin film (7,8). However, to the best of our knowledge, the influence of microstructural features of Pr doped SnO2 including defectrelated oxygen vacancies as well as preferential orientation on the sensing mechanism has not been addressed in detail. In the present work, undoped and Pr-doped SnO2 thin films are synthesized by spray pyrolysis process. Microstructural investigations based on X-Ray diffraction (XRD), Field Emission Scanning Electron Micrograph (FESEM), Transmission electron Micrograph (TEM), Atomic Force Micrograph (AFM), and optical studies including UV-Visible Absorbance Spectra, Photoluminescence, Raman Spectroscopy of thin films has been conducted. The sensor based on as prepared films are used for the LPG detection. Remarkable response to LPG accompanied with quick response and recovery time is stamped for1 wt.% Pr-doped film. A humble attempt is made to correlate the improved sensing mechanism with microstructural modifications induced via Pr doping.

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Experimental Details The pristine and Pr-modified SnO2 thin films are prepared by the pyrolytic decomposition of an aqueous solution of SnCl4. 2H2O (99.99%) with 20 ml of isopropanol and distilled water at 320 °C. Doped films are achieved by incorporating Praseodymium (III) oxide Pr2O3 in the spray solution. Highly stable, well coated and uniform films with the compatible structure are obtained. The structural analysis is done by computer-controlled Rigaku MiniFlex 600 X-Ray diffractometer with Cu-Kα radiation (λ = 1.5418Å) as an X-Ray source at 20kV and 20mA in the scanning angle 2θ from 20° to 70°. Field Emission Scanning Electron Microscopy (FESEM) by Quanta SEG 250 is utilized to disclose the morphology and dimension of the samples. UV-Vis absorbance spectra of the samples are recorded by Perkin Elmer Lambda 650 spectrometer in the wavelength range 280 to 800 nm. The results were correlated with the sensing mechanism by establishing the effects of microstructural modifications. The gas sensing measurement of prepared films is carried out by an indigenously built static gas sensing system which consists of temperature controlled airtight glass chamber with four probe leads attached. Resistance measurement and gas sensing response of the samples are done using Keithley 2400 source meter at different temperatures by heating the film inside the chamber with atmospheric air as ambient and then in the ambiance of LPG. The sensor response to a specific gas concentration is defined as S (%) =

(Ra −Rg ) Ra

× 100

(1)

where Rg: resistance of the sensor in test gas and Ra: resistance of the sensor in air. Result and Discussions Structural analysis. Recorded X-ray diffractogram of tin oxide film of varied Pr doping concentration from 0 to 6 wt. % is depicted in Fig. 1. Polycrystalline nature with tetragonal rutile structure of SnO2 is confirmed from the spectrum analysis (JCPDS card No:21-1250). For Pr doping up to 3 wt.%, the diffraction pattern corresponding to any other phase such as Pr2O3 cannot be seen in the XRD pattern, but at 6 wt.%, an additional phase corresponding to SnO orthorhombic phase can be observed. Probably doping Pr3+ in SnO2 crystallite structure vanquish its growth to large crystallite which is indicated by brooding of peaks with the rising amount of dopant. The calculated crystallite size for the prepared samples varies from 9 to 28 nm. For moderate doping concentrations, it is observed that the crystallite size in Pr doped sample increases. It is because of the fact that the dopant atom occupies interstitial positions [9]. Along with the increase in crystallite size, there is a significant reduction in the lattice strain. A small amount of doping concentration supports the progress of cassiterite phases along with preferential orientation in the planes (110), (101), (200), (310) and (301) [10]. From the most prominent peaks of XRD, the lattice constants (a=b and c) values evaluated from using the equation, 1

𝑑𝑑2

=

(ℎ2 +𝑘𝑘 2 ) 𝑎𝑎2

+

𝑙𝑙2

𝑐𝑐 2

(2)

Debye-Scherrer formula and the tangent formula [11], the crystallite size (D) and lattice strain (ε) values of the SnO2 film are determined. 𝐷𝐷 =

0.95 𝜆𝜆

𝛽𝛽𝛽𝛽𝛽𝛽𝛽𝛽𝛽𝛽 𝛽𝛽

𝜖𝜖 = 4 tan 𝜃𝜃

(3) (4)

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Fig.1 XRD pattern of undoped and Pr-doped SnO2 thin films where λ is the wavelength of X-Ray equal to 1.5418Å, θ is the Bragg diffraction angle and 𝛽𝛽 (radians) is the full width half maximum of the peak. The estimated values of stacking fault and the dislocation density [4] show a dip in their values compared to the pristine sample. Table 1 depict calculated values. Table 1. Crystallite size, microstrain, stacking fault, and dislocation density of prepared SnO2 thin films with different Pr concentration Sample code

Sensor Response [%]

undoped

81.2

Crystallite Size Microstrain [nm] XRD W-H XRD W-H Analysis Analysis 9.5 9.93 0.024 0.001

Stacking Fault

Dislocation Density [nm-2]

0.015

8.31E16

0.1 1 2

95.7 99.4 68.4

19.29 13.88 21.88

24.41 10.16 21.54

0.007 0.010 0.007

0.003 0.019 0.005

0.004 0.007 0.004

2.71E15 1.96E16 2.04E15

3 6

85.8 74.4

13.51 28.39

6.368 28.17

0.011 0.006

0.028 0.001

0.007 0.006

2.85E16 1.21E15

The variation of lattice constant ratio with cell volume against Pr doping concentration is shown in Fig. 2. For moderate Pr3+ doping up to 1 wt.%, lattice ratio is reduced linearly, and then it keeps a steady high value which is near to the bulk estimate. Earlier reports have shown that tin interstitials and oxygen interstitials or doubly charged oxygen vacancies increase the size of the lattice, while VO0 induces a reduction in lattice size [12]. A hike in cell volume is observed up to 1 wt.% Pr doping, and then it subsided to the bulk value on higher doping concentration. This behavior asserts that a small amount of Pr doping can induce local defects which may play a crucial role in presiding over the

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sensing phenomenan. From the figure, it is clear that the sample prepared with 1 wt.% Pr incorporation has the maximum distortion in its lattice. When large ionic radii (101.3 pm) Pr3+ ion replaces small ionic radii (70 pm) Sn4+ ion int the host SnO2 structure trigger more lattice distortion which bring out greater oxygen vacancies for charge compensation [13].

Fig. 2 Lattice ratio a/c along with cell volume for the Pr doped samples along with the pristine one

Fig. 3 Texture coefficient of planes for Pr-doped SnO2 thin films along with the pristine one Fig. 3 shows the Texture coefficient (TC) of prominent planes for different Pr doping concentrations [10]. In contrast to undoped samples, the TC of (200), (310), and (301) planes are significantly improved in 1 wt.% Pr-doped sample. Electrical Characterization-DC Conductivity. The V-I curve of pristine and Pr-doped SnO2 films are shown in Fig. 4 (a) and (b) respectively. A roughly linear relationship is noted from -20 V to +20 V. Resistance variations in the air may be attributed to the difference in internal potential barriers arising from defects. The slope of ln 𝞼𝞼ac versus 1/T (Fig. 5 (a) and (b)) gives the value of activation energy. It is 0.525 eV for the undoped sample, whereas 0.157 eV for 1 wt.% Pr doped sample in the temperature regime 573- 623 K, and the results are found to be consistent with the findings of Sagadevan et al [14]. The mobility of charge carriers may be enhanced upon thermal activation, which may result in hopping to cross the barrier potential and enhance current. The activation energy needed for this process is less than 1 wt.% Pr doped sample compared to the pristine one. This variation in energy may be attributed to the tunneling of charge carriers through the grain boundary.

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(a)

27

(b)

Fig. 4 V-I Characteristics at different temperatures of (a) undoped and (b) 1 wt.% Pr doped SnO2 samples

(a) (b) Fig. 5 DC conductivity variation with temperature of (a) undoped (b) 1wt.% Pr doped SnO2 samples Sensor Response to LPG. Gas sensing can be improved with a thin layer having a large surface area, as it is a surfacphenomenonna. Usually, oxygen gets adsorbed on the thin film surface, which extracts conduction electrons resulting in an increased potential barrier [15]. When the sample surface is exposed to a reducing gas like LPG, the adsorbed oxygen reacts with the reducing gas and the trapped electrons are released, resulting in an increase in conductivity [16,17] The mechanism can be explained as follows 𝑚𝑚 2

O2 + [vacant site] + e– → (Om–)site

X + (Om–)site → (XOm)site + e–

(5) (6)

The prepared samples are tested for 500 ppm of reducing gas LPG at different operating temperatures. The sensor response of the samples at three operating temperatures is shown in fig.6. All the samples show good LPG response at 350℃ and 1 wt.% Pr doped sample performs better at all operating temperatures. On increasing the dopant concentration the sensor response shows a decline at all temperatures.

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Fig.6 Sensor response as a function of Pr-doping concentration at various operating temperatures The response behaviour of Pr doped samples along with pristine sample at 350℃ is plotted in Fig. 7

Fig. 7 Sensor response at 350 ℃ of Pr-doped SnO2 thin films along with the pristine one For a deep analysis of the sensor performance at 350 ℃, time vs sensor response curve has been drawn for all samples (Fig. 7). 1 wt.% Pr-doped film shows the best LPG response (500ppm) of 98.84 % with a good response time of 9 s. The sensor response decreases for all other doping concentrations. For light Pr-doping, the chemisorbed oxygen at the SnO2 surface increases and which results in more sensing sites responding [18]. Heavy doping influences gas sensing negatively, which may be due to the excessive disorder on the surface as well as the state density introduced in the film [19]. The sensor response is greatly correlated with the growth orientation. 1 wt.% Pr-doped film shows the lowest response time with good sensitivity and invariably exhibits a characteristic higher texture coefficient for the (200) plane in combination with the (310) and (301) plane.

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Fig. 8 The resistance variation with time of 1 wt.% Pr doped SnO2 thin film on exposure to LPG The variation in resistance occurring in the sample on exposure to LPG is depicted in Fig. 8. The resistance value drops very sharply on exposure to LPG, keeps a low profile for a few minutes and on evacuation, it almost regains its original value. Fig. 9 illustrates a comparison of the sensor response with time at various operating temperatures of undoped and Pr doped SnO2 thin film. It is well plausible from the figure that, though the sensor performance of all the doped films is comparable at higher operating temperatures, it is not so good at a lower operating temperature. At 350 °C, the response time of 1 wt.% Pr doped sample is 9s towards its exposure to 500 ppm of LPG, while it takes approximately 11s to recover. The faster oxidation of the gas in the sample surface may be the reason for its quick response.

Fig.9 Transient sensor response of 1 wt,% Pr doped SnO2 thin film along with the pristine one

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At all operating temperatures, Pr doped samples show better response and recovery compared to the pristine ones. Accordingly, the main achievement in rare earth doping in SnO2is the quick gas response and recovery time, together with enhanced response magnitude. To investigate further into the good response behavior of 1 wt.% Pr doped sample, some studies have been performed. Micro-structural characterization-High Resolution Transmission Electron Microscopy (HRTEM). Transmission Electron Microscopy (TEM) is a method used to determine particle size and distribution. The images of Pr doped SnO2 thin film are shown in Fig 10. Polygonal nanoparticles of size about 18.57 nm with different sizes can be seen in the TEM image. The d spacing corresponds to the (110) plane. The crystal structure with random orientation is evident from the Selected Area Electron Diffraction (SAED). The direct variation of d spacing and reciprocal lattice spacing in the SAED is well indexed to the tetragonal unit cell of the rutile phase. The presence of (110), (101), and (200) planes is clear from the SAED ring pattern.

Fig. 10 (a) TEM, (b) HRTEM, and (c) SAED pattern of 1 wt.%Pr doped SnO2 thin film To compare the microstructural modification brought about by Pr doping, the bright field images of undoped SnO2 are taken and are shown in Fig. 11. It is evident that the film is composed of agglomerated nanometric particles of enlarged size in the range of 14-26 nm. The prominent planes are along (110) and (211), as seen in Fig.11(c). For most of the grains, the average size is less than 30 nm and it has greater importance when developing gas sensor materials [20,21].

Fig. 11 (a) TEM, (b) HRTEM and (c) SAED pattern of undoped SnO2 thin film FESEM. The morphology and grain size of the prepared films have been studied using Field Emission Scanning Electron Microscopy(FESEM). The images of pristine and 1wt.% Pr doped films are shown in Fig. 12 (a) and (b) respectively.

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(a) (b) Fig.12 FESEM image of (a) undoped and (b) 1wt.% Pr-doped SnO2 In 1 wt.% Pr doped SnO2 thin film, the particles are homogeneous with fused together polygonal shaped particles and hence the average grain size found to increase. Thus they are composed of welldefined highly compact grains with huddled nanoparticles. Also, voids in between the grains and large surface irregularities are observed. The average particle size is 130 nm. The incorporation of Pr is accompanied by a change in surface morphology. The voids become more prominent Fig. 12(b). This may be pointed out as a reason for the enhanced sensitivity of this sample. For investigating the morphological change introduced via Pr doping, we examined the FESEM of the undoped sample prepared at 320 °C, which is shown in Fig. 12(a). The film shows a regular and homogeneous distribution of polygonal grains which are closely packed and jumbled up with one another with minimum pores. Undoped SnO2 thin film has an average grain size of 80 nm. The grain size increases with the addition of dopants as discussed in XRD analysis. Atomic Force Microscope (AFM). The surface feature helps to realize surface roughness and deposition growth of thin films. The surface morphology and topology of SnO2 film samples are explored using an atomic force microscope(AFM). The images were obtained for undoped, and 1 wt.% Pr doped SnO2 film are shown in Fig. 13 (a), and (b) respectively. The 3- dimensional and 2dimensional topography gives evidence of uniform thin films with polygonal bottomed granular microstructure. The diameter of these grains depends on the subsequent doping amount. The surface appears rough at the submicron scale. The height of root means square (RMS) roughness dafm is in the order of 19.78 nm and 26.21 nm, for undoped and 1 wt.% Pr doped SnO2 film. The increased value of surface roughness for 1 wt.% Pr doped thin film may be a key factor for its fast response towards LPG.

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(a) (b) Fig.13 3D and 2D topographic view of (a) undoped (b) 1 wt.% Pr doped SnO2 thin film Optical Properties and Band Gap Variation. The band gap energy of the pristine SnO2samples prepared at deposition temperatures of 320 °C has been estimated as 3.899 eV. Initially, the crystallite size increases with deposition temperature and then decreases at higher temperatures. Further, there seems an inverse relation between the crystallite size and band gap. Fig. 14 shows the Tauc Plot of Pr doped SnO2 thin film along with undoped films. To analyze the band gap engineering effect by Pr incorporation in SnO2 films, photon energy (hν) vs (αhν)2 graphs are plotted. (Fig. 14)

Fig.14 Tauc plot for (a) undoped and (b) 1 wt.% Pr doped SnO2 thin films

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α =α0 (hν-Eg)n where n=1/2, for direct transition.

33

(5)

The estimated values of band gaps are 3.9, 3.87, 3.94, 3.93, 3.91 and 3.95 eV for undoped, 0.1 wt.% ,1 wt.%, 2 wt.% 3 wt.% and 6 wt,% Pr-doped SnO2 samples respectively. For undoped SnO2 film, the observed band gap is 3.9 eV which is greater than its bulk value of 3.6 eV. The band gap energy decreases from 3.9 eV to 3.87 eV for 0.1 wt. % Pr-doped SnO2 samples and thereafter it exhibits an increase with the increase in doping concentration. Moderate doping concentration may increase the number of allowed states near the conduction band due to the defects in nano-grained SnO2 thin films which in turn causes the decrease in bandgap [22]. Raman Analysis. The nature of oxygen vacancies imparted by the rare earth ions in the SnO2 host lattice, from the perspective of enhanced gas sensing can be studied from the analysis of Raman spectra. Raman lines are sensitively dependent on the defects and structure pertaining to the sample.

Fig. 15 Raman spectra of undoped and Pr doped SnO2 thin films at room temperature Fig. 15 shows the Raman spectra taken at room temperature of undoped, 1 wt.% Pr doped and 6 wt.% Pr doped SnO2 thin films. Apart from the major peaks at 484, 635, and 776 cm−1 (Eg, A1g, and B2g modes) of SnO2 [23], there are minor peaks at 91,121, 192, 227, 566, 699, and 776. cm−1. The expected high energy shift of absorption edge for nanocrystalline materials is confirmed by UV measurement. Weak Raman bands at 566 cm−1and 699 cm−1can are assigned as (TO) and (LO) modes, both of which are IR active. These abnormal Raman bands are characteristic of nanostructured SnO2 [24]. The most intense peak at 635 cm−1is observed for 1 and 3 wt.% Pr samples [25], which is strengthened as the Pr doping amount increases. Aragon F.H.et al (2010)] and Thangadurai et al (2005) [26,27] have demonstrated that stressed SnO2 exhibits Raman peaks at higher values in comparison with ambient conditions. Therefore, it can be ascertained that apart from nonstoichiometry, residual stresses also play a role for doped the appearance of Raman peaks at comparatively higher wave numbers as in the case of 3 wt.% Pr doped sample. The peak corresponding to Eg mode gets shrunk in 1 wt.% Pr doped sample, giving evidence of increased oxygen vacancies. The spectra of 3 wt.% Pr doped sample has a tapered structure and almost all peaks are narrowed here. This may be due to a large amount of Pr incorporated into the sample Photoluminescence Analysis. The presence of impurities, crystalline quality, and the exciton fine structures in a material can be determined using Photoluminescence Technique [28,29]. The PL

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analysis of the undoped and doped SnO2 thin films is carried out to study the nature of oxygen vacancies. The primary focus is given on the defects like small polarons and oxygen vacancies. The defects present in the bandgap such as oxygen vacancies, dangling bonds, and interstitial tin atoms give rise to the luminescence peaks of the energy states.

. Fig. 16 PL spectra of undoped, 1 wt.% and 3 wt.% Pr doped sample for the excitations of (a) 320 nm and (b) 370 nm Fig. 16 (a) and (b) denote the PL spectra of pristine and Pr-doped SnO2 thin films. For the excitation, 320 nm, for doped samples, two major peaks are observed namely 385nm and 465nm. The intensity of the peaks decreases for 1 wt.% Pr and 3wt.% Pr doped film. There is a significant shift in the peak positions The PL spectra taken with excitation 370 nm show peaks at 505, 609, 657, and 682 nm. The intensity of the peaks increases for 1 wt.% Pr doped samples, but it decreases for 3 wt.% Pr doped sample. When excited with 370 nm, all samples exhibit a peak at yellow emission around 609 nm. This emission peak (2.04 eV) is less than the band gap width of 3.899eV of bulk SnO2 as obtained from UV-visible spectroscopy. Thus the emission peak at 609 nm may be originating from the luminescence centers created from the dangling bonds or interstitials in the SnO2 nanostructure film. [30]. The interaction of high-density oxygen vacancies with interfacial tin creates trapped states within the band gap, and this contributes to the high PL intensity at room temperature [31,32]. The increased intensity of the peak corresponds to 505 nm compared to 609 nm for1 wt.% Pr doped sample gives evidence for the excess of in-plane oxygen vacancies compared to bridging oxygen vacancies. Conclusions SnO2 thin films doped with 0.1wt. % to 6 wt. % of Pr are prepared via spray pyrolysis and their LPG sensing property is analyzed. The tetragonal rutile structure of the prepared films with an average crystallite size of 10 -30 nm is confirmed by the X-Ray diffraction pattern. Pr doped samples exhibit faster response along with good recovery time. The FESEM images of the Pr doped sample shows a significant modification in morphology. The polycrystalline nature of the film is evidenced from TEM analysis and the crystalline sizes obtained from TEM and XRD results are well matched each other.DC conductivity properties tell less activation energy for 1 wt,% Pr doped sample which substantiates its enhanced sensing property. The mechanism behind the sensor response is clear from optical studies of thin films. The prominent change in optical band gap with doping concentration shows the formation of defect states via Pr doping. The role of defects and oxygen vacancies are well understood from Raman and PL analysis. The conduction process via oxygen ion vacancy motion positively influences gas sensing. All the studies suggest that these films are of potential scope for gas sensing applications.

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Acknowledgments The author (DS) is grateful for the financial assistance from RUSA through the award letter No. 009/2019-20 RUSA- MRP(S) MAC and the Faculty Improvement Fellowship from the UGC (SWRO/FIP 12th Plan/ KLMG 038 TF-06), and the author (PK) acknowledges the financial assistance from KSCSTE (822/DIR/2014-15/KSCSTE dated 09.02.2015) The authors are thankful to Dr. P Balraju, CIT Coimbatore, for FESEM analyses and to Dr. Anu of SPAP, MG University for TEM characterization. References [1]

J. Huang, and Q.Wan, Gas sensors based on semiconducting metal oxide one-dimensional nanostructures. Sensors, 9 (2009) 9903-9924.

[2]

Yulia Borchert, Patrick Sonstrom, Michaela Wilhelm, Holger Borchert, and Marcus Ba1umer, Nanostructured Praseodymium Oxide: Preparation, Structure, and Catalytic Properties, J. Phys. Chem. C, 112 (2008) 3054-3063.

[3]

M. A. Pinheiro, T. F. Pineiz, E. A. de Morais, L. V. Scalvi, M. J. Saeki, and A. A.Cavalheiro, Schottky emission in nanoscopically crystallized Ce-doped SnO2 thin films deposited by sol– gel-dip-coating. Thin Solid Films, 517 (2008) 976-981.

[4]

B.Thomas, S. Deepa, & K. P. Kumari, Influence of surface defects and preferential orientation in nanostructured Ce-doped SnO2 thin films by nebulizer spray deposition for lowering the LPG sensing temperature to 150° C. Ionics, 25(2019) 809-826.

[5]

W. Q. Li, S. Y. Ma, Y. F. Li, X. B. Li, C. Y. Wang, X. H. Yang, ... and X. L. Xu, Preparation of Pr-doped SnO2 hollow nanofibers by electrospinning method and their gas sensing properties. Journal of alloys and compounds, 605(2014) 80-88.

[6]

F. H. Aragón, I. Gonzalez, J. A. Coaquira, P. Hidalgo, H. F. Brito, J. D. Ardisson, ... & P. C. Morais, Structural and surface study of praseodymium-doped SnO2 nanoparticles prepared by the polymeric precursor method. The Journal of Physical Chemistry C, 119 (2015) 8711-8717.

[7]

S. Deepa, K. PrasannaKumari, & B. Thomas, Influence of lattice strain and dislocations on the LPG sensing performance of praseodymium doped SnO2 nanostructured thin films. IJRASET, 5 (2017) 1054-1059.

[8]

S. Deepa, B.Thomas, & K.PrasannaKumari, Influence of surface oxygen vacancies on the LPG sensing response and the gas selectivity of Nd-doped SnO2 nanoparticulate thin films. Journal of Materials Science: Materials in Electronics, 30 (2019) 16579-16595

[9]

Y. M. Chiang, D. P. Birnie, and W. D. Kingery, Physical Ceramics: Principles for Ceramic Science and Engineering. J. Wiley, New York, (1997) 357-358.

[10] Boben Thomas, Benoy Skariah, Spray deposited Mg-doped SnO2 thin film LPG sensor: XPS and EDX analysis in relation to deposition temperature and doping, 625 (2015) 231–240. [11] B. D. Cullity, Elements of X-Ray diffraction, Addision-Wesley, New York, 1978. [12] A. K. Singh, A. Janotti, M. Scheffler and C. G. Van de Walle, Sources of electrical conductivity in SnO2.Physical Review Letters, 101 (2008) 055502. [13] S., Deepa, K. P. Kumari, & B. Thomas, Contribution of oxygen-vacancy defect-types in enhanced CO2 sensing of nanoparticulate Zn-doped SnO2 films. Ceramics International, 43 (2017) 17128-17141. [14] S. Sagadevan, and C. Arunseshan, Dielectric properties of cadmium selenide (CdSe) nanoparticles synthesized by solvothermal method. Applied Nanoscience, 4 (2014)179-184

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[15] J. Liu, X. Huang, G. Ye, W. Liu, Z. Jiao, W. Chao, Z. Zhou and Z. Yu, H2S detection sensing characteristic of CuO/SnO2 sensor.Sensors, 3 (2003) 110-118. [16] N. B. Sonawane, K. V. Gurav, R. R. Ahire, J. H. Kim and B. R. Sankapal, CdS nanowires with PbS nanoparticles surface coating as room temperature liquefied petroleum gas sensor. Sensors and Actuators A: Physical, 216 (2014) 78-83. [17] T. Nakahara, H. Koda; Chemical Sensor Technology, Vol. 3 (Kodansha and Elsevier, Amsterdam, (1992). [18] N. Barsan, U. W. Eimar, Conduction model of metal oxide gas sensors, Journal of Electro ceramics 7 (2001) 143-167. [19] R.K Mishra, P.P Sahay, Zn-doped and undoped SnO2 nano particles: A comparative structural, optical and LPG sensing properties study, Materials Research Bulletin 47 (2012) 4112-4118. [20] A. Srivastava, S. T. Lakshmikumar, A. K. Srivastava and K. Jain, Gas sensing properties of nanocrystalline SnO2 prepared in solvent media using a microwave assisted technique. Sensors and Actuators B: Chemical, 126 (2007) 583-587. [21] D. Haridas, A. Chowdhuri, K. Sreenivas, and V. Gupta, Enhanced room temperature response of SnO2 thin film sensor loaded with Pt catalyst clusters under UV radiation for LPG. Sensors and Actuators B: Chemical,153 (2011)152-157. [22] R.K. Mishra, P.P Sahay, Materials Research Bulletin 47 (2012) 4112-4118. [23] M. N. Rumyantseva, A. M. Gaskov, N. Rosman, T. Pagnier and J. R. Morante, Raman surface vibration modes in nanocrystalline SnO2: correlation with gas sensor performances. Chemistry of materials, 17 (2005) 893-901. [24] J.G. Traylor, H.G. Smith, R.M. Nicklow, and M.K. Wilkinson, Lattice dynamics of rutile. Physical Review B, 3 (1971) 3457-3472. [25] M. Batzill, and U. Diebold, The surface and materials science of tin oxide.Progress in surface science, 79 (2005) 47-154. [26] F. H. Aragón, J.A.H. Coaquira, P. Hidalgo, S.L.M., Brito, D. Gouvêa, and R. H. R. Castro, Structural and magnetic properties of pure and nickel doped SnO2 nanoparticles. Journal of Physics: Condensed Matter, 22 (2010) 496003. [27] P. Thangadurai, A. C. Bose, S. Ramasamy, Kesavamoorthy, and T. R. Ravindran, High Pressure effects on electrical resistivity and dielectric properties of nanocrystalline SnO2.Journal of Physics and Chemistry of Solids, 66 (2005) 1621-1627. [28] W.Z. Wang, C.K. Xu, G.H. Wang, Y.K., Liu and C.L. Zheng, Preparation of Smooth Single Crystal Mn3O4 Nanowires. Advanced Materials, 14 (2002) 837-840. [29] S. Luo, P. K. Chu, W. Liu, M. Zhang and C. Lin, Origin of low-temperature photoluminescence from SnO2 nanowires fabricated by thermal evaporation and annealed in different ambients. Applied Physics Letters, 88 (2006)183112. [30] Nguyen VnHieu, Le Thi Ngoc Loan, Nguyen DucKhoang, Nguyen Tuan Minh, Do Thanh Viet, Do Cong Minh, Tran Trung and Nguyen Duc Chien. A facile thermal evaporation route for large-area synthesis of tin oxide nano wires: Characterisations and their use for liquid petroleum gas sensor. Current Applied Physics, 10 (2010) 636-641. [31] A. Kar, M. A. Stroscio, M. Dutta, J. Kumari, and M. Meyyappan, Observation of ultraviolet emission and effect of surface states on the luminescence from tin oxide nanowires. Applied Physics Letters, 94 (2009) 101905. [32] H.T. Chen, S.J. Xiong, X.L. Wu, J. Zhu, J.C. Shen and P.K. Chu, Tin oxide nanoribbons with vacancy structures in luminescence-sensitive oxygen sensing. Nano letters, 9 (2009) 1926-1931.

Key Engineering Materials ISSN: 1662-9795, Vol. 934, pp 37-46 doi:10.4028/p-eft062 © 2022 Trans Tech Publications Ltd, Switzerland

Submitted: 2022-04-02 Revised: 2022-07-18 Accepted: 2022-07-20 Online: 2022-11-28

Solvent Effects on the UV-Visible Absorption & Emission of Tris[4-Diethylamino)Phenyl]Amine Sardul Singh Dhayal1,a,*, Abhimanyu Nain1,b, Amit Kumar2,c Department of Electronics & Communication Engineering, Guru Jambheshwar University of Science & Technology, Hisar, Haryana-125001, India

1

School of Engineering & Technology, Central University of Haryana, Jant-Pali, Mahendergarh, Haryana-123031, India

2

a,*

[email protected], [email protected], [email protected]

Keywords: Effect of solvents, Optical Properties, Organic Semiconductor, UV-Vis spectroscopy, Photoluminescence

Semiconductors,

Molecular

Abstract. Tris[4-(diethylamino)phenyl]amine (TDAPA) is an organic molecular semiconductor generally used to enhance the charge transport of the devices for some time now. TDAPA is dissolved in various Polar and Non-polar solvents like DMF, Acetone, Acetonitrile, Ethanol, Methanol, Toluene and Chloroform. Absorption spectrum of solution is recorded using UV-Vis spectroscopy and absorption peak for different solvents were observed in UV and Near-UV region. PL study and Pl Excitation study is also carried out for these solvents. Results for absorption and photoluminescence show some interesting phenomenon of Stokes’ shift. The colour coordinates for respective emission are represented by CIE 1931. The study is successfully carried out for better understanding of effect of these solvents on the optical properties of TDAPA. 1. Introduction The increased interest of researchers in field of OLED and OFET due to their wide application in optoelectronic devices [1,2]. Devices based on these materials are low weight and flexible cost effective devices which enhance their use in foldable devices as they provide large surface [3,4]. Numerous organic materials have lately been shown to have carrier mobilities equivalent to hydrogenated amorphous silicon in experiments [5–7], with improved stability of OFET [8]. An OFET's fundamental component is an organic semiconductor. It controls both charge carrier transportation and charge carrier injection [9]. The OFETs are widely used in applications like sensors [10,11] and electronic paper like display [12,13]. It is generally established that charge-carrier transport is limited by hopping between molecules in the disordered region, and that molecular ordering must provide sufficient overlap of the p-orbitals of conjugated organic molecules to facilitate efficient charge migration between neighbouring molecules [14–16]. Organic molecules such as TDAPA are used as a dopant in nanostructure films to boost carrier concentration for organic optoelectronic devices such OFET [17–19]. Small molecule fluorophores are used in wide range of applications such as biophysics, molecular biology, medicine, and material sciences [20–24]. Such materials are widely adopted for the optoelectronic devices such as OFET and OLED, and to study their behaviour in different solvents is essential to understand its structural and electrical behaviour. The study of the effects of a solvent on a solute's structure and spectroscopic behaviour is critical for the advancement of solution chemistry [25–28]. The important parameters to study are optical bandgap and Stokes’ shift. The energy difference between the PLE maxima and the lower-energy (red-shifted) emission is Stokes shift. In present study the effect of different polar and non-polar solvents is studied. Absorption spectrum of TDAPA in various solvents and its corresponding PL emission is observed. The Stokes’ shift is observed in different solvents and CIE colour coordinates suggest the emission colour for solvents.

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2. Experimental Work Tris[4-(diethylamino)phenyl]amine (TDAPA) was procured from Sigma Aldrich whereas N,NDimethylformamide (DMF), Acetone, Acetonitrile, Ethanol, Methanol, Toluene and Chloroform all solvents were procured from EMPARTA (99.8%, ACS). The glassware was washed with deionised water and later oven-dried. 2.3 mg of TDAPA powder is dissolved in 5 mL different solvents. The solution is further sonicated for 30 minutes to dissolve the TDAPA powder in solvents. Figure 1 (a) shows TDAPA in different solvents under normal light, and Figure 1 (b) shows solvents under UV light (374 nm). The solvents are marked as Toluene (a), Chloroform (b), Methanol (c), Ethanol (d), Acetonitrile (e), DMF (f) and Acetone (g).

Figure 1: TDAPA in different solvents in natural light (a) and under UV light (b). 3. Result and Discussion a) Absorption Spectroscopy The effect of different solvents, both Polar [Aprotic (DMF, Acetone and Acetonitrile) and Protic (Ethanol and Methanol)] and Non-Polar [Toluene and Chloroform] was explored using UV-Vis spectroscopy. The absorption spectrum of TDAPA in DMF and Chloroform was studied earlier by Tanis et al. [19]. In this study, the absorption spectrum of TDAPA in DMF and Chloroform is given in Figure 2, where for DMF absorption peak is observed at 263.7 nm and 320 nm and for Chloroform 281.8 nm, 318 nm and 404 nm, which suggest that in DMF solvent the absorption peak is in NUV (Near UV) and in Chloroform solvent peaks are in both NUV and Visible region. The optical bandgap (Eg) calculated using Tauc's equation for DMF is 3.35 eV, and it is 2.71 eV for chloroform. The lower bandgap for chloroform is due to the higher molarity of the TDAPAChloroform solution, earlier explained in literature[29].

Figure 2: Absorption spectrum of TDAPA in DMF (a) and Chloroform (b).

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The absorption spectrum of TDAPA in Acetone gives an absorption peak at 324 nm, whereas Acetonitrile gives an absorption peak at 259 nm and 316 nm, as shown in Figure 3 (a, b). In nonpolar solvent Toluene, the absorption spectrum was observed at 326 nm, as shown in Figure 3 (c).The corresponding Egvalues in Acetone, Acetonitrile and Toluene are 3.36 eV, 3.4 eV and 3.27 eV, respectively.

Figure 3: Absorption spectrum of TDAPA in Acetone (a) and Acetonitrile (b) and Toluene (c). The absorption spectra of TDAPA in Protic solvents (Ethanol and Methanol) are shown in Figure 4 (a, b). The absorption peak is observed at 210 nm, 260 nm, and 310 nm for TDAPA in Methanol, whereas absorption peaks for Ethanol were at 207 nm, 254 nm and 316 nm. The Eg values are 3.38 eV and 3.42 eV, respectively.

Figure 4: Absorption spectrum of TDAPA in Methanol (a) and Ethanol (b).

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b) Photoluminescence Studies The Photoluminescence (PL) studies of TDAPA in different solvents were studied to explore its potential use in organic optoelectronic devices. For Toluene, the PL emission spectra were recorded in the excitation range 350 nm to 410 nm, and emission intensity rise till excitation wavelength 395 nm where emission peak is observed to be 440 nm or at 2.82 eV, and then PL intensity decreased. No significant shift in emission peak was observed; hence no EDF (excitation dependent fluorescence) was observed in Toluene. The Photoluminescence excitation (PLE) emission spectra were recorded for Toluene, and an excitation peak was observed at 392 nm. For Chloroform, the emission spectrum is taken from 440 nm to 505 nm, and the emission peak was also shifted to a larger wavelength to 535 nm or to lower energy, i.e. 2.31 eV. With increasing the excitation wavelength, a new peak was observed around 500 nm, which is believed to be due to EDF; a similar effect was observed in PLE where excitation peak wavelength shifts from 479 nm to 490 nm for emission peak 530 nm to 560 nm. To study the effect of TDAPA in Protic (polar) solvent, TDAPA is dissolved in Methanol and Ethanol. PL emission spectra were recorded in the range of the excitation wavelength 345 nm to 410 nm, and maximum emission is observed at 464 nm or 2.67 eV and 450 nm or 2.74 eV at excitation wavelength 370 nm for both and in both Protic solvents, no shift in emission peak is observed. The PLE emission peak for Methanol is observed around 364 nm for emission wavelength 470 nm, and maximum excitation is observed for emission wavelength 470 nm. For Ethanol, PLE peak is observed at 370 nm, and no shift is observed. PL emission spectra for Aprotic solvents are shown in Figure 5 (e, f, g), for Acetonitrile emission is taken for excitation wavelength 340 nm to 400 nm in the interval of 5 nm and PL peak emission is observed at 468 nm for excitation wavelength 355 nm or 2.64 eV and PLE emission peak is observed at 355 nm. For DMF, PL emission is taken in the range 350 nm to 420 nm, and PL maximum peak is observed at 447 nm for excitation wavelength 395 nm or 2.76 eV. PLE emission is observed at 387 nm for an emission peak 445 nm. For Acetone, PL emission is observed for 355 nm to 415 nm, and maximum emission is observed at 443 nm for excitation wavelength 390 nm or 2.79 eV. Uniform emission was observed for each solvent, and no significant shift in emission peak was observed in any solvent.

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Figure 5: PL emission spectra of TDAPA Toluene (a), Chloroform (b), Methanol (c), Ethanol (d), Acetonitrile (e), DMF (f) and Acetone (g) and inset show PLE emission spectra of respective solvent.

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The Normalised PL (NPL) emission vs wavelength for various solvents is shown in Figure 6 (a), where emission wavelength is shifted for various solvents and in Figure 6 (b). It has been discovered that when the polarity of the solvent increases, the emission peak wavelength increases as well [30]. NPL vs energy shows that in various solvents, peaks shift from 2.82 eV to 2.31 eV, giving different emission colours. The Colour coordinates are represented in CIE 1931 shown in Figure 6 (c), and its X-Y coordinates are given in Table 1. The stokes' shift for different solvents was calculated as shown in Table 1, and the Overlap of PLE and PL with calculated stokes shift is shown in Figure 7.The Stokes' shift is the difference between PLE and PL maxima in the wavenumber scale [31]. For non-polar solvents (Toluene and Chloroform), a smaller Stokes' shift is observed, while for Polar solvents significant shift is observed. For Methanol, a large Stokes' shift of 5996.48cm-1 is observed. An intra-molecular charge-transfer excitation of an electron from the HOMO to the LUMO is responsible for the observed high Stokes shifts.[32]. The maximum Stokes’ shift observed is 6801.49 cm-1 which is for Acetonitrile, an Aprotic solvent. Similar significant Stokes' shifts were observed for polar solvents [33]. Table 1: Optical bandgap, emission peak, and CIE coordinate for TDAPA in various solvents. S. No. A B C D E F G

Solvent Toluene Chloroform Methanol Ethanol Acetonitrile DMF Acetone

PL Emission peak 440 nm 535 nm 464 nm 450 nm 468 nm 447 nm 443 nm

PLE Emission peak 392 nm 479 nm 362 nm 370 nm 355 nm 386 nm 392 nm

Stokes’ shift (cm-1) 2782.93 2254.85 5996.48 4804.80 6801.49 3535.37 2936.84

CIE Coordinates X Y 0.15564 0.06001 0.35607 0.61532 0.13636 0.13866 0.14903 0.08818 0.14335 0.21529 0.14792 0.06844 0.15388 0.07423

Figure 6: PL emission spectra of TDAPA Normalised PL Intensity vs wavelength(a), Normalised PL intensity vs Energy (b) in various solvents and CIE 1931 showing colour coordinates for different solvents.

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Figure 7: Normalised PL and normalised PLE spectrum of TDAPA in different solvents and their Stokes’ shift in wavelength.

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4. Conclusion This study provides the better understanding of effect of basic solvents (as discussed above) on the optical properties of TDAPA. The absorption peak is observed in UV and Near-UV region for all studied solvents except Chloroform which has absorption peak in visible region. The optical bandgap of TDAPA in Chloroform is 2.71 eV and for other solvents it lays in range 3.27 eV to 3.36 eV. The PL and PLE emission spectrum gives the lowest emission peak for Toluene i.e. 440 nm (in wavelength) or 2.81 eV (energy) while for Chloroform emission peak is observed at 535 nm or 2.31 eV. Very small Stokes’ shift was observed in non-polar solvents, while the significant Stokes’ shift was observed in Aprotic solvents and maximum Stokes’ shift is observed for Acetonitrile (113 nm). The colour coordinates were represented by CIE 1931. These results indicate that TDAPA has vast possibilities for next generation optoelectronic devices. References [1]

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CHAPTER 2: Materials Processing

Key Engineering Materials ISSN: 1662-9795, Vol. 934, pp 49-58 doi:10.4028/p-yf8y2m © 2022 Trans Tech Publications Ltd, Switzerland

Submitted: 2022-07-27 Accepted: 2022-09-06 Online: 2022-11-28

Analysis of the Influence of Build Plate Pre-Treatment and Process Parameters on the Bonding of Additively Manufactured Parts of TiAl6V4 to the Build Plate BAY Christian1,a*, MAHR Alexander2,b, HOFMANN Andreas3,c, WIENERT Christian3,d, DÖPPER Frank3,e University of Bayreuth – Research Center for Additive Innovations, Universitaetsstrasse 30, 95447 Bayreuth, Germany

1

Fraunhofer-Institute for Production Technology and Automation – Project Group Process Innovation, Universitaetsstrasse 9, 95447 Bayreuth, Germany

2

University of Bayreuth – Chair for Manufacturing and Remanufacturing Technology, Universitaetsstrasse 30, 95447 Bayreuth, Germany

3

[email protected], [email protected], c [email protected], [email protected], e [email protected]

*, a

Keywords: Additive Manufacturing, TiAl6V4, pre-treatment, build plate bonding

Abstract. In this paper, the influence of the surface roughness of build plates as well as the process parameters laser power, scan speed and exposure curing time on the bonding of additively manufactured components made of TiAl6V4 to the build plates is analyzed. These analyses are carried out with build plates made of Titanium Grade 2 and AlMgSi0.5. The analyses show that higher surface roughness leads to lower bending strength and thus poorer bonding of components on the build plate. In addition, it is shown that the bending strength normalized to the bonding surface decreases at high laser power, especially at high scanning speeds. Furthermore, multiple exposure results in lower flexural strength. Introduction Additive manufacturing has increasingly established itself as an industrially applied manufacturing technology, particularly in aerospace technology, medical technology and general mechanical engineering [1]. Powder bed fusion of metals with laser beam (PBF-LB/M) is the most widely used additive manufacturing process for metallic series components in industry [1]. In PBF-LB/M, a layer of metal powder is applied to the build plate and selectively exposed and melted by thermal energy introduced by a laser beam. Subsequent cooling and solidification of the molten material creates a strong bond with the underlying layer. For the further industrialization of PBF-LB/M, the increase of economic efficiency as one of the target criteria in the triangle of effects of production according to Westkämper and Hayes & Wheelwright represents a critical success factor [1,2,3,4]. One approach to increase the economic efficiency of the PBF-LB/M is the use of low-cost materials for the build plate especially for high-cost powder materials like TiAl6V4. Minimizing the effort of machining the build plate in view of reuse is another approach to increase the economic efficiency of PBF-LB/M. Therefore, the objective of this paper is to analyze the bonding of additively manufactured components made of TiAl6V4 by varying the build plate material, the build plate pre-treatment as well as the process parameters laser power, scan speed and exposure frequency. This analysis is carried out by measuring the flexural strength and dimensional stability of test specimens as well as metallographic analysis of the transition zone between test specimen and build plate. Regarding the state of the art, mixed compounds are compounds between dissimilar materials. Dissimilar means that the metallurgical and mechanical-technological as well as chemical and physical properties of the materials differ [5,6]. In terms of metallurgy, the suitability for fusion welding of different alloys depends essentially on the phases occurring in the transition zone and their

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behavior during cooling [5,6,7,8]. If solid solution formation occurs, the two materials are homogeneously welded. If there is no solubility or limited solubility, intermetallic phases are formed, which usually lead to embrittlement of the interface and to poorer properties between the materials to be welded [5,6,7,8]. To estimate the solubility of the various material components, the so-called “Darken-Gurry diagram” based on the so-called “Hume-Rothery rule” can be applied [9,10]. According to this, the difference in the atomic radii of the elements of the solution partners for solid solution formation must not exceed ±15 %—metallurgical favorable is a deviation below ±8 %. In relation to titanium, the elements aluminum, tantalum and niobium are at a deviation of the atomic radius lower than ±8 %. [6,11,12] PBF-LB/M is a laser welding process in which a layer of metal powder is applied to the build plate by a coater and selectively exposed and melted by thermal energy introduced by a laser beam [5,7]. Subsequent cooling and solidification of the molten material creates a strong bond with the underlying layer. The laser power PL is a process parameter that describes how much energy is provided by the laser per unit of time [15,16]. The speed at which the focal point of the laser beam travels along the scan vectors over the powder bed is referred to as the scan speed vS [16]. The energy volume density describes the energy introduced per material volume unit, cf. equation (1) [16,17]. Here, hS is the scan distance and DS is the layer thickness [15]. Ev =

PL vS

× hS × DS

(1)

If the introduced energy per unit material volume is too high, e.g. due to high laser powers or low scanning speeds, there is a risk of the formation of a so-called "keyhole formation" [18]. In this case, a deep welding effect occurs due to the large energy input, whereby pores can appear within the component. In addition, high laser power and scanning speed at the same time can result in the formation of melt droplets ("balling effect"), which in particular reduces the relative component density. On the other hand, too low an energy volume density can lead to bonding defects due to insufficient melting of the powder. Local absorption and reflection properties of the powder particles and the melt further reduce the energy introduced into the powder bed. [15,16,19] In PBF-LB/M, high energy input and cooling rates as well as multiple melting of the powder material are characteristic [20,21]. The resulting temperature-gradient mechanism can lead to thermal-induced residual stresses, which can negatively affect the mechanical properties and lead to distortion [16,22]. In addition to the bonding of intermetallic phases, these residual stresses occur increasingly in the transition zone between material combinations [23]. According to Zhang et al and Yaqoob Mohsin Baqer et al, a welded joint between two components made of different materials (titanium and steel or aluminum and titanium) can be produced by adjusting the process parameters [8,23]. Repeated exposure of the layer is one way to reduce residual stresses [21,22]. According to Baqer et al, for Al/Mg and Mg/Ti compounds, coatings or interlayers, thus increasing the bond strength, can reduce the formation of intermetallic phases. Schaub et al showed that by reducing the temperature gradient between a substrate plate and an additively manufactured component made of TiAl6V4, the bond strength in the transition zone could be increased [20]. Experiments Procedure. In this work, the influence of the pre-treatment of the built plates (test T1) and the process parameters (test T2) on the flexural strength and dimensional stability of test specimens made of TiAl6V4 was investigated. These tests were carried out with both Titanium Grade 2 and AlMgSi0.5 build plate materials (see Table 1). In addition, the transition zone between test specimen and build plate was analyzed metallographically for both build plate materials.

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Table 1: Design of Experiments Build plate material Titanium Grade 2

AlMgSi0,5

T2

T1

not sandblasted Variation of pretreatment of build plate

sandblasted (ds = 30 mm)

Variation of process parameters

Laser power PL Scan speed vS Exposure frequency EL

T1: Analysis of flexural strength

sandblasted (ds = 150 mm) T2.1: Analysis of dimensional stability T2.2: Analysis of flexural strength

–

Metallographic analysis of the transition zone

Test specimen geometry. For the experimental investigations, test specimens for determining flexural strength and dimensional stability were manufactured on the PBF-LB/M machine “Orlas Creator RA” from Coherent. The PBF-LB/M machine has a fiber laser (laser wavelength 1,070 nm). The laser focus diameter is 40 µm, the layer thickness DS = 25 µm and the scan line distance hS = 5 µm. TiAl6V4 with a measured particle size distribution of D10 = 19.2 µm, D50 = 32.9 µm and D90 = 42.5 µm was used as the material for all test specimens. In addition to the actual bending test specimen for the introduction of the bending force, the test specimens consist of a base in the transition area to the build plate, for which the process parameters are varied in the tests. In the tests carried out, the base corresponds to a height of three exposed layers. The test specimen was optimized so that the flexural strength and dimensional stability could be determined with the test setup for both build plate materials without any additional geometry change. Fig. 1 shows the technical drawing of the test specimen and the position of the load application.

Fig. 1: Technical drawing of the flexural test specimen used (a) and schematic representation of the force application F during bending test (b)

Process parameters and pre-treatment of build plate. The upper part of the specimen, the flexural test specimen, was manufactured with standard parameters (PL = 106.75 W, vS = 600 m/s) for both build plate materials. These standard parameters were determined as part of a parameter study to optimize the achievable relative component density. The parameters of the lower part of the test specimen, the base, were varied in test T1 for both build plate materials and in test T2 for the build plate material Titanium Grade 2, see Table 2.

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Table 2: Test parameters for the test specimen base for tests T1(*) as well as T1 and T2

PL

vS EL

Analysis parameter

Titanium Grade 2

AlMgSi0,5

Flexural strength & dimensional stability

100 W; 125 W*; 150 W; 175 W; 200 W*; 225 W; 250 W

50 W*; 100 W*

Flexural strength

400 mm/s*, 800 mm/s*

400 mm/s*; 800 mm/s*

Dimensional stability

400 mm/s; 600 mm/s; 800 mm/s

–

Flexural strength & dimensional stability 1*; 2*, 5*

1*; 2*

Test T1: Analysis of influence of pre-treatment of build plate. To investigate the influence of the surface condition of the build plate on the bond strength, machined build plates made of both the investigated materials were each divided into four quadrants. The first quadrant of the build plates was sandblasted for 30 seconds at a distance ds = 30 mm from the blasting gun. The second quadrant was sandblasted for 30 seconds at a distance ds = 150 mm. The last two quadrants were nonsandblasted. Table 3 shows microscopy images and the surface properties of the build plates made of Titanium Grade 2 and AlMgSi0.5. Table 3: Surface properties of Titanium Grade 2 and AlMgSi0.5 build plates before and after pre-treatment (T1)

AlMgSi0.5

Titanium Grade 2

Parameter/Specification

not sandblasted

sandblasted (ds = 30 mm)

sandblasted (ds = 150 mm)

0.69 µm 4.75 µm

1.36 µm 10.94 µm

1.67 µm 12.65 µm

0.67 µm 4.02 µm

2.43 µm 27.74 µm

2.20 µm 16.69 µm

Microscopy image Centerline roughness Ra Average peak-to-valley height Rz Microscopy image Centerline roughness Ra Average peak-to-valley height Rz

Test T2: Analysis of influence of process parameter. In addition to a variation of the exposure parameters laser power PL and scan speed vS, the influence of the exposure frequency EL (cf. Table 2) on the bonding of additively manufactured component to the build plate was investigated. Experiment Setup. The bending strength was tested in an experimental setup developed for this application, which is based on the operation of a bending testing machine. In order to ensure reproducibility and comparability between the test specimens, the force was applied at a constant drawing speed of 4 m/min by means of an electric winch “PA100D” from Trading EU. The maximum force at the time of failure on the test specimen is recorded by means of a peak value measurement with the force-measuring device “FK 250” from Sauter. The force gauge was designed to detect maximum forces of 250 N with a resolution of 0.1 N.

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Fig. 1: Schematic structure of the test rig (a) and clamped build plate during the bending test (b)

To prevent the force gauge from being overloaded by excessive forces, it is suspended in a factoral pulley block with a ratio of 1:3 to realize maximum forces of up to 750 N and a drawing speed of 1.33 m/min. A fixture developed for this application was used for clamping the build plate with the test specimens, see Fig. 2b. For the optical analysis of the fracture surfaces of the specimens (dimensional accuracy), the 3D profilometer “VR-5000 One-Shot 3D” from Keyence is used. The dimensional accuracy of the test specimens was analyzed on basis of the cross sections of the fracture surfaces (deviation analysis). For the metallographic analysis (topography contrast images with scanning electron microscope SEM “1540EsB Cross Beam” from Carl Zeiss) of the transition zone between the additively manufactured component and the build plate, a metallographic test specimen made of TiAl6V4 has been manufactured on the Titanium Grade 2 build plate (20 mm by 20 mm and a height of 1.1 mm). Due to the expected lower bond strength on a non-metallographic test specimen compared to a metallographic test specimen of the same type, a cylindrical metallographic test specimen with a diameter of 12 mm and a height of 1.1 mm has been fabricated on the AlMgSi0.5 test specimen (Titanium Grade 2 test specimen with PL = 50 W, vs = 800 mm/s and AlMgSi0.5 test specimen with PL = 100 W, vs = 600 mm/s). Results For the experimental analyses, 216 test specimens were manufactured and analyzed with regard to their flexural strength and dimensional stability. Analyzing of pre-treatment of build plate (T1). To evaluate the influence of the surface roughness of the Titanium Grade 2 build plate on the flexural strength, the mean value of the flexural strengths as a function of the energy volume density is shown in Fig. 3.

Fig. 2: Flexural strength in relation to pre-treatment of build plate of Titanium Grade 2 (a) and AlMgSi0.5 (b) and energy volume density (Test T1)

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The T1 tests with the Titanium Grade 2 build plate material show that the specimens on the nonsandblasted build plates have the highest flexural strengths. The flexural strengths of the test specimens on the build plate sandblasted at a distance ds = 150 mm are lower on average than those of the quadrants sandblasted at a distance ds = 30 mm. It is assumed that the increased surface roughness results in a notch effect, which reduces the flexural strength of the specimens on the build plate [24]. When considering the mean values of the flexural strengths of the parameter combinations as a function of the energy volume density, it can be seen that the flexural strength of the specimens on the sandblasted quadrants is at least 62 % lower than the flexural strength of the specimens on the non-sandblasted quadrant. Regardless of the surface roughness, multiple exposure does not lead to a significant improvement of the flexural strength. Analyzing of process parameters—geometric analysis of the dimensional stability (T2.1). Fig. 4 shows the results of the T2.1 tests on the dimensional stability of the specimens on the Titanium Grade 2 build plate

Fig. 3: Influence of process parameter (laser power, scan speed and exposure frequency) on the dimensional stability for build plate of Titanium Grade 2 (Test T2.1)

All measured actual diameters are equal to or larger than the nominal diameter of 1.2 mm. The more often the exposure, the greater this measurement deviation. In addition, at constant scanning speed with increasing laser power, a slight tendency towards increasing dimensional deviation can be observed. Fig. 5 shows the actual diameters marked a), b) and c) in Fig. 4 at the level of the transition zone from test specimen to build plate. The actual diameter on the build plate increases from 1.20 mm (single exposure) to 1.29 mm (+7.5 %; double exposure) to 1.56 mm (+30 %; fivefold exposure).

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Fig. 4: Microscopy and 3d images of Titanium Grade 2 build plate of with a) low, b) medium, and c) large deviation of actual to target diameter

Analyzing of process parameters—mechanical analysis of the flexural strength (T2.2). In test T2.2, the influence of PL, vs and EL on the flexural strength was investigated. In Fig. 6, the flexural strength was normalized to the specific cross-sectional area of the respective specimen (cf. test T2.1).

Fig. 5: Influence of PL and EL for scan speed 400 mm/s (a) and 800 mm/s (b) on normalized flexural strength of test specimens made of Titanium Grade 2 (Test T2.2)

Test T2.2 with the build plate material Titanium Grade 2 shows that the normalized flexural strength initially increases with higher laser power, irrespective of the exposure frequency. From a laser power of 200 W and 225 W, respectively, the normalized flexural strength drops with a single exposure. With increasing exposure frequency, this behavior occurs at lower laser powers and largely. With a reduction of the flexural strength to 19 % of the maximum value, this trend is particularly noticeable with a five-fold exposure. In each case, this decrease is due to favorable spatter and oxide formation, resulting in a decrease in mechanical properties due to adhesion and embrittlement. This trend is more evident at the higher scanning speed of 800 mm/s, see Fig. 6b. Accordingly, the position of the optimum of the bending strength is influenced not only by the scanning speed but also by the laser power and the exposure frequency. At a scanning speed of 400 mm/s and a single exposure, the optimum lies at a laser power of 200 W (global optimum; 2,544 N/mm²), at a double exposure at 175 W (2,430 N/mm²) and at a five-fold exposure at 150 W (2,232 N/mm²). Metallographic analysis. Fig. 7 shows the transition zone between the metallographic test specimen made of TiAl6V4 and the build plate of Titanium Grade 2 in the corresponding SEM images. Due to the only partially visible transition zone, it can be concluded that the alloying elements of the test specimen are well mixed with the alloying elements of the identical build plate. No detachment

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of the metallographic specimen from the build plate was observed. Furthermore, the melt paths of the first layer as well as the lamellar basket weave structure typical for TiAl6V4 can be seen in Fig. 7 b) [25].

Fig. 6: SEM image of the transition zone of the metallography test specimen made of TiAl6V4 on the Titanium Grade 2 build plate at 100-x (left) and 1,000-x (right) magnification

When looking at the topography contrast images of the metallographic specimen manufactured on the AlMgSi0.5 build plate, it is noticeable that it has detached from the build plate at the edge, see Fig. 7a). Furthermore, both Fig. 7a) and Fig. 7b) show a clear demarcation between the specimen and the build plate. This indicates a low degree of mixing between the alloying constituents of the specimen and the build plate and explains the reduced flexural strength of the specimens made of TiAl6V4 compared to the AlMgSi0.5 build plate.

Fig 7: SEM image of the transition zone of the metallography test specimen made of TiAl6V4 on the AlMgSi0.5 build plate at 100-x (left) and 1,000-x (right) magnification

It can be assumed that physical effects, such as mechanical hooking, mainly pronounce the flexural strength. This can be related to the strong oxide layers on the surface of the build plate as well as the low solubility and low diffusion between the TiAl6V4 specimen and the AlMgSi0.5 build plate, Summary In the analyses presented here, it was shown both for a build plate made of Titanium Grade 2 and of AlMgSi0.5 that an initial higher surface roughness leads to lower flexural strength and thus to poorer bonding of components to the build plate. Furthermore, it was shown that at high laser power, the bending strength normalized to the surface decreases. This was particularly evident at high scanning speeds. In further analyses, it was demonstrated that multiple exposures (two and five times) result in a lower normalized flexural strength of the test specimens and thus a poorer bond of the test specimens to the build plate.

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References [1] T. Wohlers, R.I. Campbell, O. Diegel, N. Mostow, J. Kowen: Wohlers Report 2021: 3D Printing and Additive Manufacturing State of the Industry, Wohlers Associates, Fort Collins, Colo., 2021. [2] E. Witt, C. Anton: Additive Fertigung: Entwicklungen, Möglichkeiten und Herausforderungen: Stellungnahme, Halle (Saale), 2020. [3] A. Weckenmann, G. Akkasoglu, T. Werner: Quality management—history and trends. In: The TQM Journal; ISSN: 1754-2731. Vol. 27 No. 3, pp. 281–293. https://doi.org/10.1108/TQM-112013-0125. [4] R.H. Hayes, S.C. Wheelwright: Restoring our competitive edge: Competing through manufacturing, Wiley, New York, N.Y., 1984. [5] H.J. Fahrenwaldt, V. Schuler: Praxiswissen Schweißtechnik: Werkstoffe, Prozesse, Fertigung, 3rd ed., Vieweg + Teubner, Wiesbaden, 2009. [6] C. Pohle: Schweißen von Werkstoffkombinationen: Metallkundliche und fertigungstechnische Grundlagen sowie Ausführungsbeispiele, Verl. für Schweißen und Verwandte Verfahren DVSVerl., Düsseldorf, 1999. [7] S. Katayama (Ed.): Handbook of laser welding technologies, Woodhead Publ, Oxford, 2013. [8] Yaqoob Mohsin Baqer, S. Ramesh, F. Yusof, S. Manladan: Challenges and advances in laser welding of dissimilar light alloys: Al/Mg, Al/Ti, and Mg/Ti alloys, (2018). [9] J.-F. Nie: Physical Metallurgy of Light Alloys, in: D.E. Laughlin, K. Hono (Eds.), Physical Metallurgy, 5th ed., Elsevier Science, Burlington, 2014, pp. 2009–2156. [10] Y. Mae: What the Darken–Gurry Plot Means About the Solubility of Elements in Metals, Metall Mater Trans A 47 (2016) 6498–6506. https://doi.org/10.1007/s11661-016-3730-1. [11] PubChem, Periodic Table of Elements, 2021, https://pubchem.ncbi.nlm.nih.gov/periodic-table/, accessed 23 January 2022. [12] K.A. Gschneidner, M. Verkade: Electronic and crystal structures, size (ECS2) model for predicting binary solid solutions, Progress in Materials Science 49 (2004) 411–428. https://doi.org/10.1016/S0079-6425(03)00026-4. [13] MatWeb: Material Property Data of Aluminum, Al, 2022, http://www.matweb.com/search/ datasheet.aspx?bassnum=AMEAL00&ckck=1, accessed 23 January 2022. [14] MatWeb: Material Property Data of Titanium, Ti, 2022, http://www.matweb.com/search/ DataSheet.aspx?MatGUID=66a15d609a3f4c829cb6ad08f0dafc01, accessed 23 January 2022. [15] J.-P. Kruth, P. Mercelis, J. van Vaerenbergh, L. Froyen, M. Rombouts: Binding mechanisms in selective laser sintering and selective laser melting, Rapid Prototyping Journal 11 (2005) 26–36. https://doi.org/10.1108/13552540510573365. [16] V. Seyda: Werkstoff- und Prozessverhalten von Metallpulvern in der laseradditiven Fertigung. Dissertation. [17] Umberto Scipioni Bertoli, Alexandra Wolfer, M. Matthews, J. Delplanque, J. Schoenung: On the limitations of Volumetric Energy Density as a design parameter for Selective Laser Melting, (2017). [18] W.E. King, H.D. Barth, V.M. Castillo, G.F. Gallegos, J.W. Gibbs, D.E. Hahn, C. Kamath, A.M. Rubenchik: Observation of keyhole-mode laser melting in laser powder-bed fusion additive manufacturing, Journal of Materials Processing Technology 214 (2014) 2915–2925. https://doi.org/10.1016/j.jmatprotec.2014.06.005.

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[19] N.K. Tolochko, Y.V. Khlopkov, S.E. Mozzharov, M.B. Ignatiev, T. Laoui, V.I. Titov: Absorptance of powder materials suitable for laser sintering, Rapid Prototyping Journal 6 (2000) 155–161. https://doi.org/10.1108/13552540010337029. [20] A. Schaub, B. Ahuja, M. Karg, M. Schmidt, M. Merklein: Fabrication and Characterization of Laser Beam Melted Ti-6Al-4V Geometries on Sheet Metal, Proceedings / DDMC 2014, Fraunhofer Direct Digital Manufacturing Conference, March 12 - 13, 2014, Berlin (2014). [21] H. Ali, H. Ghadbeigi, K. Mumtaz: Effect of scanning strategies on residual stress and mechanical properties of Selective Laser Melted Ti6Al4V, Materials Science and Engineering: A 712 (2018) 175–187. https://doi.org/10.1016/j.msea.2017.11.103. [22] K. Osakada, M. Shiomi: Flexible manufacturing of metallic products by selective laser melting of powder, International Journal of Machine Tools and Manufacture 46 (2006) 1188–1193. https://doi.org/10.1016/j.ijmachtools.2006.01.024. [23] Y. Zhang, J. Zhou, D. Sun, X. Gu: Nd:YAG laser welding of dissimilar metals of titanium alloy to stainless steel without filler metal based on a hybrid connection mechanism, Journal of Materials Research and Technology 9 (2020) 1662–1672. https://doi.org/10.1016/j.jmrt.2019. 12.001. [24] E. Siebel, M. Gaier: Untersuchungen über den Einfluss der Oberflächenbeschaffenheit auf die Dauerschwingfestigkeit metallischer Bauteile, VDI-Zeitschrift (1956) 1715–1723. [25] M. Peters (Ed.): Titan und Titanlegierungen, 3rd ed., Wiley-VCH, Weinheim, 2002.

Key Engineering Materials ISSN: 1662-9795, Vol. 934, pp 59-66 doi:10.4028/p-54tx42 © 2022 Trans Tech Publications Ltd, Switzerland

Submitted: 2022-08-02 Accepted: 2022-09-06 Online: 2022-11-28

Laser Metal Deposition (LMD) Toolpaths with Adaptive Capability for Complex Repairs and Coating Geometries Igor Ortiz1,a, Piera Alvarez1,b and M. Angeles Montealegre2,c Ikergune A.I.E. Inzu Group, San Antolin, 3, 20870 Elgoibar, Guipúzcoa, Spain

1

Talens Systems. Inzu Group, Polígono Albitxuri, 20, 20870 Elgoibar, Guipúzcoa, Spain.

2

[email protected], [email protected], [email protected]

a

Keywords: Toolpaths, CAD/CAM, Laser Metal Deposition (LMD), Coatings, Additive Manufacturing.

Abstract. Inside Direct Energy Deposition (DED) processes is the Laser Metal Deposition (LMD) technology. Industries that can implement this technology approach are from automotive to energy sectors where critical parts suffer due to operation cycles, weather or hazardous environments etc. LMD process can be applied for coating, repair and build near net shape geometries. One of the main problems of LMD applied in the coating or in the repair is the dealing with these different types of geometries, to achieve an adherent and homogeneous coating. The current calculation of toolpaths in LMD software is based on a mathematical algorithm that relies on subtractive processes such as machining. The main drawback of using this type of toolpaths is that in this case they do not take into account the overlap between adjacent machining toolpaths. While for machining this parameter is not relevant, in the LMD process, the overlap between two contiguous laser tracks is a critical point to have an adequate process with the required quality. Talens Systems has developed a new Software, Azala software is able to calculate these strategy toolpaths for advanced repairs and coatings in any type of geometry. Beside taking into account the overlapped between contiguous laser clads, the calculated toolpaths have integrated the main laser process in LMD (laser power, process speed, powder flow). The objective of this work is to validate the Azala software developed using a piece with complex geometry on a laboratory scale. The developed software brings the possibility to automate repair and coating, where the LMD process provides a value-added opportunity to reduce production costs due to the repair of value-added components. Introduction Inside Direct Energy Deposition (DED) processes, the Laser Metal Deposition (LMD) is a technology that metallic powder or wire can be used with a laser power source. In LMD the metallic powder is melted on metallic surfaces using a high power laser to repair damaged surfaces or for upgrading surface properties. The LMD process can produce coatings with no porosity, finer surface finishes, more consistent layer thicknesses, and more precise clad placement, than conventional welding process. Inherently brings a low heat input to the process, resulting in fine microstructures, small heat affected zones, and low distortion. Moreover, it helps to reduce processing time, laser cladding may restore parts to a near net shape and could be used other materials for improving mechanical properties. This technology can develop three main processes, coating, part repair and build near net shape 3D geometries [1]. Thanks to laser cladding can repair damaged areas or build directly CAD geometries, these capabilities avoids the design and development of complex tooling and fixtures [2],[3]. This technology is capable of repair and build high added value parts like turbine blades and rocket nozzles [4],[5]. The LMD technology is capable of tailoring the microstructure of the deposited part to meet the requirements of the repair, avoiding the introduction of cracks or microstructure disorder [6]. The toolpaths calculation mathematical algorithm is done based in subtractive processes (machining, grinding, drilling etc.), the problem of using this type of toolpaths is that the overlap between the toolpaths, doesn’t always match with the requirement of LMD process [7]. Without constant

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toolpaths the DED toolpaths are not suitable for the process, due to next issues appears in the material porous, lack of powder, material accumulation in the tracks. LMD process can produce coatings with minimal heat affected area with the substrate, this allows a more consistent layer thickness and more precise clad placement and quality than conventional welding process. Repair of added value parts are done with Laser Metal Deposition, that allows extremely tight control over the size of the fusion zone and the amount of heat introduced into the component [8]. The result of this is that the deposited part with process meets the requirements of the repair, avoiding cracks formation or the disorder on the microstructure [9]. Besides the LMD process can build complex three dimensional features and parts, like aerospace parts and rocket propulsion systems [10]. The LMD process is highly different from the machining process, due to the variables that are involved in the process [11]. Variables like laser power, process speed, powder flow and flow of shielding gas are key variables for configuring a stable laser cladding process [12],[13] while for machining, the main variables involved in the process are the feeds and speeds of the machining process and the type of the cutting tool. On the other hand commercial software’s, currently lack in the capability to configure Laser Metal Deposition process variables, due to fragmented software, forces to use different modules for preparing and calculating LMD processes [14]. Expert knowledge is required to work with this software, standards for the industrialization of processes are lacking [15]. The objective of this work is to validate the Azala software developed by Talens System, a software which has integrate the main process parameters of LMD. A core architecture with specific LMD algorithms, have been developed for creating complex toolpaths and process parameters for coating complex surfaces [16]. The proposed software in this work, Azala software, allows the calculation of different strategies that are focused to coat or repair medium and complex surface geometries. These strategies are calculated to have a regular overlap between continuous laser coatings to maintain process quality. Furthermore the software has the capability to calculate laser hardening toolpaths with process parameters. Methodology Process know-how and process variables have been programed in a customized CAM environment for LMD selected processes, for solving complex variables and toolpaths in laser cladding processes with variable surface geometries. Previously CAD software is used for adding a colour to surfaces that have been selected for coating process. The areas can be detected by the software and the system calculates the toolpaths in previously detected area, with attached process data inside the toolpath. This task reduces the robot or CNC programming due to the part programming is done with and off line methodology [17]. By using this software the operator is able to create automated coating laser strategies for any surface type, additionally the toolpaths integrate the laser process parameters. There are three main process workflows that can be seen in Fig. 1. The Azala software has been developed with laboratory experiments and is based on mesh parametrization mathematical algorithms. The programming process workflow is divided in three main steps. CAD planning, Toolpaths and NC Code. Inside the first main step CAD planning, there are two phases related with CAD model preparation. The first phase is related for preparing the coating areas using any CAD modelling software, inside the CAD model the colour of the coating areas is identified with selected colour. The second phase inside the first main step, is related with loading into Azala software’s previously coloured CAD model surfaces. Second main step the Toolpaths calculation, is related with toolpaths and three phases are identified, the first phase is numbered with the third number and is for creating the zero workplane for part location in Azala software. The fourth phase is related with reference edges selection and extraction for toolpath calculation. The final phase which is numbered with number five, is for selecting boundary zones that are previously coloured in the CAD modelling software.

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The final main step is for creating the NC Code, inside this step are three key phases. The first phase is numbered with the number six and is for executing the toolpath calculation, with previously selected edge, the coloured boundary CAD surfaces and the zero workplane. With this data Azala software calculates the toolpaths with a regular overlap for not disturbing the coating process. The last two phases, seven and eight are related with the numerical control code (NC) for creating machine tool (CNC) or robot cells programs. The seventh phase is for translating the Azala strategies to the machine tool numeral control code. Finally in the eighth phase, the previously created NC programs can be loaded to a CNC or robot.

Fig. 1: Example of the workflow for surface coating toolpaths

Software Testing, This paragraph describes the steps for software testing process, is divided in CAD and part geometries, that is the description of the geometry types that are going to be coated. On the other hand the next paragraph is the toolpaths calculation that is going to be used for surface coating. When the CAD model and the part that is going to be tested is going to be prepared, the next step is to test the toolpaths with the CAD model. Besides for preparing the experimental testing a part is going to be designed and manufactured, the developed software coating toolpaths are going to be calculated and generated in the laboratory machine tool. The calculated toolpaths are going to include the LMD process parameters, finally the program is going to be validated in a three axis machine tool with a Siemens 840D control. CAD and part geometries, For testing the toolpaths in Azala software a part has been designed with different surfaces and angled features. The part has different angled surfaces in the horizontal plane, on the other hand it has an angled surface wall. These aggressive geometry changes are designed, for studding the coating in the whole surface. The combination of these geometries, help to validate the mathematical calculation of the coating toolpaths. These geometrical changes are useful to analyse the type of strategy, the overlap and the process parameters that are going to be integrated. Besides the part has complex features, like variable surfaces that match with horizontal surface the angles between surfaces cannot be measured by 3D CAD drawings or with probing in a coordinate measuring machine. The height of the planar surface is 35mm, on the other hand the top height of the part is 65mm, with a 120 degrees angle between them Fig. 2. The geometries that can be measured are for example, angles between the planar surface and the wall surface that have an angle of 120 degrees. Between these surfaces is a radius of 5mm. On the other hand the coating is going to be studied is the whole green surface marked in Fig. 2. In this area the surface has been machined with a good surface roughness Ra 1,6. This surface roughness is needed for archiving a good LMD coating.

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Fig. 2: CAD drawing of the part that is going to be validated. Plan view left and section view right.

First of all a design of a CAD model has been done, after validating the model a physical prototype has been manufactured, for experimenting the Laser Metal Deposition coating process. For optimizing the behaviour of clad tracks, the surface of the part has been machined to a surface quality roughness Ra 1,6. This roughness quality is necessary, for the laser cladding process powder bonding to the substrate Fig. 3.

Fig. 3: Manufactured test part positioned for testing the coating process

Toolpaths calculation, in this section the mathematical procedures that have been used for developing the toolpaths strategies are going to be tested. The process helps to analyse and identify the quality of the overlap, the toolpath type and the parameters that are configured for the process. The combination of the digital strategies validation with the physical part, helps to verify the quality of the coated geometry. Furthermore the toolpaths calculation with Azala software, automatically extracts the edges for toolpath calculation, in previously green coloured surface area Fig. 4. On the other hand the LMD toolpaths have the data points introduced, like laser ON and OFF commands, geometry control points that are connected to the CAD geometry, this options create adaptive strategies that are linked to the CAD model. This adjusts the cladding process with the part geometry Fig. 4.

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Fig. 4: CAD model with edge selection for calculating equidistant LMD toolpaths in variable surface model

After reviewing the strategy in Azala’s graphics system, the strategies are checked. When the checking is approved a NC program is generated, finally the generated NC program can be loaded in the laboratory machine tool for laser cladding coating. Moreover, each toolpath point has linked process variables for automating the laser cladding process and reduces manual intervention. Besides, the strategy overlap and process parameters have been experimented and validated in laboratory machine tool [18]. Finally after using Azala digital toolpaths checking and validation options, the NC programs have been performed with two layer laser cladding process and tilting orientation, for best adaptation to the part surface, the part orientation is done as horizontal as possible with the nozzle [19]. The realized coating is homogenous across the variable surface and the angled wall geometry. Finally the toolpaths have a constant overlap with integrated process parameters (laser power, speed, laser on and off, powder flow rate etc.) that can be seen in the Fig. 5.

Fig. 5: CAD data of surface views with coating toolpaths. The red lines represent when the laser is ON and blue lines represent laser OFF that are reposition movements.

Software and process validation, in this paragraph is explained the equipment that has been used for LMD experiments, using a high-power 6KW fiber laser from IPG Photonics. The fiber laser source is a YLS-6000 (λ = 1070 nm), with a 5 mm diameter. The optical head was assembled in a three axis system with a Siemens Sinumerik 840D control. The metallic powder particles were introduced in a GTV powder feeder and delivered to the work area through a coaxial nozzle. On the other hand a Fraunhofer IWS coaxial nozzle Powerline 12 model was used for laboratory tests. The carrier gas has been used to deliver the metallic powder from the powder feeder to the melt pool Fig. 6.

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Fig. 6: Coaxial nozzle from Fraunhofer IWS

In a first task single tracks were characterized, this is done previously to manufacture the solid part by LMD. Next process variables have been analysed like, laser power, powder flow and process speeds. In a second phase base material parts have been machined with a surface quality roughness Ra 1.6. A manufactured prototype has been machined that can be seen in Fig. 7, the purpose in this study is to validate the toolpaths calculated with Azala software for process automation and geometry adaptation [20]. Furthermore the strategies have integrated laboratory know how, software performance can be studied with the coating part tests in variable geometry type of surfaces. Finally in laboratory machine tool the calculated coating toolpaths are validated. In conclusion Azala software has a friendly user interface, for coating and LMD process strategies that are adapted to complex part geometries and have integrated process parameters.

Fig. 7: Coating example part in variable surfaces

Conclusions In this research, the Azala software has been validated with the coating of a piece designed with complex geometry. It has been possible to validate the capacity of the Azala software to develop trajectories that adapt to any type of surface geometry. In addition, the strategies of the cladding laser coating process have been developed and contrasted, obtaining a homogeneous and quality coating. In laboratory experiments, next points have been validated: 1. Adaptive toolpaths that are adapted to any geometry type. Complex and variable surface geometries coating experiments. 2. Laser cladding toolpath automation, due to integrated process parameters (laser power, process speed, powder flow…). 3.

Surface selection capability to the areas that are going to be coated.

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Finally, Azala software integrates a friendly user interface, which automates the process workflow and generates automated laser cladding toolpaths. This capability opens new methodologies for automating the coating of different parts like mold and dies, valves, blades etc. Additionally, the toolpaths and process parameters researched in Azala software have been developed and validated for both LMD and hardening processes. Future work A future work is to automate LMD toolpaths using reverse engineering techniques and integrate the with the physical part geometry that has to be coated, repaired or built. Another research area is in line quality assurance system linked with the toolpath parameters. Even though in this article the software is described for cladding processes, it has the capability of laser heat treatment for hardening and softening. Moreover there is the capability, to vary the laser spot shape and energy density. Acknowledgment This work has been co-funded by the Basque Government under Hazitek program ZE-2019/00002 (project Addit4all). Conflict of interest statement No potential conflict of interest was reported by the authors. References [1]

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P. Alvarez, M. Montealegre, J. Pulido-Jiménez, and J. Arrizubieta, “Analysis of the Process Parameter Influence in Laser Cladding of 316L Stainless Steel,” J. Manuf. Mater. Process., vol. 2, no. 3, p. 55, Aug. 2018, doi: 10.3390/jmmp2030055.

[10] P. R. Gradl, “Advancement of Metal Additive Manufacturing Techniques and Materials for Rocket Propulsion Applications Paul Gradl National Aeronautics and Space Administration ( NASA ),” no. August, 2020, doi: 10.13140/RG.2.2.23927.65444. [11] D. Boisselier, S. Sankaré, and T. Engel, “Improvement of the laser direct metal deposition process in 5-axis configuration,” Phys. Procedia, vol. 56, no. C, pp. 239–249, 2014, doi: 10.1016/j.phpro.2014.08.168. [12] S. Saqiba, R. J. Urbanica, and K. Aggarwal, “Analysis of laser cladding bead morphology for developing additive manufacturing travel paths,” Procedia CIRP, vol. 17, pp. 824–829, 2014, doi: 10.1016/j.procir.2014.01.098. [13] S. Zhu, W. Chen, L. Ding, X. Zhan, and Q. Chen, “A mathematical model of laser cladding repair,” Int. J. Adv. Manuf. Technol., vol. 103, no. 9–12, pp. 3265–3278, Aug. 2019, doi: 10.1007/s00170-019-03588-3. [14] A. Bayesteh, J. Ko, F. Ahmad, and M. B. G. Jun, “Improvement of Computer-Aided Manufacturing (CAM) Software for Laser Machining,” J. Korean Soc. Manuf. Technol. Eng., vol. 24, no. 4, pp. 374–385, 2015, doi: 10.7735/ksmte.2015.24.4.374. [15] C. Bremer, “Adaptive metal deposition and data management for automated overhaul of complex turbine components,” 26th Int. Congr. Appl. Lasers Electro-Optics, ICALEO 2007 Congr. Proc., vol. 105, 2007, doi: 10.2351/1.5060988. [16] D. Montoya-Zapata et al., “Generation of 2.5D Deposition Strategies for LMD-based AdditiveManufacturing,” Procedia Comput. Sci., vol. 180, pp. 280–289, 2021, doi: 10.1016/j.procs.2021.01.165. [17] N. Wittig, “Digitalization: Laser Metal Deposition — The Future of Spare Parts and Repairs for Industrial Steam Turbines,” pp. 1–7, 2018, doi: 10.1115/gt2018-75066. [18] P. Alvarez, M. Á. Montealegre, F. Cordovilla, Á. García-Beltrán, I. Angulo, and J. L. Ocaña, “Direct Generation of High-Aspect-Ratio Structures of AISI 316L by Laser-Assisted Powder Deposition,” Materials (Basel)., vol. 13, no. 24, p. 5670, Dec. 2020, doi: 10.3390/ma13245670. [19] J. Marashi, “Remanufacture of hot forging tools and dies using laser metal deposition with powder and a hard-facing alloy,” pp. 189–203, 2019, doi: 10.1007/s13243-018-0063-9. [20] P. R. Gradl, C. Protz, and T. Wammen, “Additive manufacturing development and hot-fire testing of liquid rocket channel wall nozzles using blown powder directed energy deposition inconel 625 and jbk-75 alloys,” AIAA Propuls. Energy Forum Expo. 2019, no. August, 2019, doi: 10.2514/6.2019-4362.

Key Engineering Materials ISSN: 1662-9795, Vol. 934, pp 67-73 doi:10.4028/p-683ks0 © 2022 Trans Tech Publications Ltd, Switzerland

Submitted: 2022-07-27 Accepted: 2022-09-06 Online: 2022-11-28

Innovative Method of Embedding Optical Fiber inside Titanium Alloy Utilizing Friction Stir Forming Hamed MOFIDI TABATABAEI1,a, Takahiro OHASHI1,b and Tadashi NISHIHARA1,c School of science and engineering, Kokushikan University, Setagaya 4-28-1, Setagaya Ward, 154-8515, Japan

1

[email protected], [email protected], [email protected]

a

Keywords: Friction stir welding, friction stir forming, mechanical joining, functional materials, titanium, optical fiber sensor.

Abstract. The development of FBG (Fiber Bragg Grating) sensors is essential for intelligent parts of airplanes to achieve ultra-lightweight structures and intelligent aviation control. Awaiting solution is embedding technology of optical fiber sensors into the base material of the parts; however, it has been challenging to embed fibers into high melting temperature point alloys like titanium-based materials without having any defects. Present research fulfills the mentioned demands effectively by utilizing Friction Stir Forming (FSF). Precisely, optical fiber has been placed into a guide slit inside the titanium sheet. Then, FSF was applied to the surface of the titanium. As a result, titanium plasticizes and flows into the guide slit. This mechanically interlocks the optical fiber inside the titanium alloy. This solid-state stirring process can bury sensors inside the titanium to protect the sensor from harsh environments. Embedded sensors will detect the strain and temperature of the host titanium in real-time. Moreover, this research can have further considerable effects on the areas like aerospace and artificial intelligence systems that demand real-time condition monitoring systems. Introduction The development of FBG (Fiber Bragg Grating) sensors is indispensable for reducing the weight of smart aircraft parts and realizing intelligent air traffic control. For the development of FBG sensors, a technique for embedding an optical fiber sensor in the base material of a component is desired. Still, it is difficult to embed fibers in a refractory alloy such as a titanium-based material without defects. Airlines worldwide spent about $ 50 billion on aircraft maintenance, repair, and overhaul in 2015 and are expected to increase to $ 625 billion in 2020 [1]. As the aircraft system becomes more complex and high-performance, the need for lighter weight and lower fuel consumption continues to increase. In turn, the quality, reliability, and predictive power of the health conditioning system of the aircraft are improved. The FBG sensor is a desirable fiber optic sensor that can simultaneously measure different parameters such as deformation, strain, pressure, and temperature through the deviation of the reflected wavelength of light and is a very accurate measurement. It can be expected to be applied to monitoring the internal state of structures. Most of the research on FBG sensors is to embed optical fibers inside composite materials such as carbon fiber reinforced composite materials used in aircraft frames, windmill blades, and high-performance automobiles. Very few studies on embedding optical fibers inside high melting materials such as titanium have been carried out. While refractory metals such as titanium can protect the sensor from harsh environments, there are strict conditions during the embedding process, such as casting and hot pressing. Ti-4.5Al-3V-2Mo-2Fe (trade name SP-700) is a β-rich superplastic titanium alloy with a microstructure developed for aerospace [2] and has excellent workability and superplastic properties. Therefore, it is suitable for forming complex aircraft and aerospace parts. Traditional methods based on foil gauges require redundancy to deal with failures caused by electrical contacts, so regular health monitoring cannot be performed inside the metal. The optical fiber sensor (FOS) is metal coated to prevent damage from the process. In addition, the material and porosity of the metal protective coating affect the adhesion of the embedded fibers, causing slippage and coat peeling to protect the

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FOS when trying to solve it using deposition techniques. The problem remains that a large metal coating (2 mm) is required [3]. Ultrasonic welding is known as an alternative method; however, the application of the sensor is limited due to the low temperature of the process [4]. The vacuum brazing technique gives some flexibility to the operation using FOS, but it is complicated and expensive because it requires a vacuum chamber [5]. Selective laser melting (SLM) is also a promising method for embedding FOS, and 3D printers for metals are expensive and not readily available. Recently, friction stir forming (FSF) has been attracting attention in the field of dissimilar material joining as a new metal forming method using the principle of friction stir welding (FSW) [6, 7]. As shown in figure 1, the shape of the die transfers to the workpiece by placing the work material on the mold and applying friction stir processing to the back surface of the work material. Research on FSF is being conducted in our laboratory [8-16]. This study proposes a novel method for joining dissimilar materials that mechanically joins titanium alloy and optical fiber by using friction stir forming, which effectively meets the mentioned above requirements. This study aims to experimentally investigate the development of a composite material with a new function in which an optical fiber is embedded inside a titanium alloy without impairing its characteristics using FSF. It can also be proposed to be introduced as a primary material for intelligent structures. It is expected that information can be transmitted as a sensor for detecting the internal state (temperature and strain) of host titanium. The scope of the application of optical fibers is not limited to aircraft applications. Still, it is expected to expand to applications in aerospace and artificial intelligence systems that require rotating machinery, vehicle structures, and real-time condition monitoring systems.

Fig. 1 Schematic Drawing of FSF [7]. Experimental Procedure Friction stir forming experiments. As shown in figure 2, a guide slit is machined in the titanium alloy plate, an optical fiber is placed there, and FSF is performed on the surface of the titanium alloy. The plastic flow of the material occurs in the slit, and the optical fiber is mechanically bonded inside the titanium alloy. As a result, fiber sensor embedded in titanium can be expected to operate reliably in scorching and harsh environments. The life of the FSF tool is one of the most critical issues when performing the friction stir process for high-strength alloys such as titanium alloys. In this experiment, the FSF experiment was conducted by selecting the Tungsten tool as the most suitable material for this experiment. The dimension of the tool is shown in figure 3.

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Fig. 2 Schematic of the experimental procedure for the proposed method.

Fig. 3 Dimension of FSF tool (Ds= 8 mm, Dq= 4 mm, Dp= 2 mm, hp= 1.2 mm, α= -5 deg, β= 45 deg) Formation of interlock by FSF. Figure 4 shows the actual test material after the FSF. In most cases, when the center of the probe was placed directly above the slit and FSF was performed, the optical fiber became brittle and broke. FSF was performed while offsetting toward the Advancing side.

Fig. 4 Appearance of friction stir formed titanium showing surface and back side condition for travel speeds of (a) 5 mm/min (b) 100 mm/min (c) 200 mm/min (d) 400 mm/min (e) 800 mm/min (rotation speed is set to 1240 rpm) The results of previous studies have revealed that the material flows toward the retreating side of the process compared to the advancing side during friction stir processing; therefore, the center of the tool is located on the advancing side by 2 mm [16]. Comparing the surface after FSF with the fixed rotation speed and changing the travel speed, it can be seen that the color of the surface has changed, which is thought to be due to the change in heat input depending on the travel speed. Also, as confirmed from the photograph of the back side, the heat input decreased as the travel speed increased, and the frictional heat was not transmitted to the back of the workpiece, especially at speeds of 200 and 400 mm/min. However, according to the results of cross-sectional structure observation, it was

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confirmed that the material was sufficiently filled in the slit even with the parameter of high travel speeds. It can also be confirmed that when the travel speed is low at 50 mm/min and high speed of 800 mm/min, burrs are likely to occur, and the surface is comparatively rough. Results and Discussion Microstructure observation by scanning electron microscope (SEM). Figure 5 (a) shows a macro photograph of the cross section after FSF. It can be confirmed that the titanium is softened by the FSF and flows into the slit. As shown in Fig. 5 (b), it can be confirmed that a cavity is generated due to the insufficient flow of the titanium alloy, which is due to the inadequate heat input.

Fig. 5 SEM photos of the mechanical interlocked optical fiber after FSF showing (a) cross-section perpendicular to the direction of processing, (b) interlocked fiber within titanium, and (c) map analysis showing that titanium has not filled the slit. Figure 6 shows the results of cross-sectional microstructure observation after the FSF. The grains are refined within the stir zone. Microstructures were observed with a scanning electron microscope, and the average grain size of 2 µm in the not affected zone (Fig. 6 (b)) was refined to 0.8 µm in the stirring zone after the process (Fig. 6 (c)). Figure 6 (d) shows the microstructures of the material that flowed into the slit. In (d), the grains are more refined; therefore, it is highly possible that the thermos-mechanically affected zone (TMAZ, (Fig. 5 (b)) zone has flowed into the groove.

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Fig. 6 SEM photo showing (a) un-affected zone, (b) Thermo-mechanically affected zone, (c) Stir zone, and (d) Flowed material Figure 7 shows the results of cross-sectional macrostructure observation under the condition that the travel speed has been reduced. It can be confirmed that the volume of the material to be softened increases, and the cavity part decreases due to the increase of heat input during the process. Figure 8 (a) is a cross-sectional photograph of increased rotation speed to increase the heat input. It can be confirmed that the flow of the material fills the slit. However, it was confirmed that the optical fiber was damaged due to increasing the shoulder diameter to increase the heat input further. (Fig. 8 (b)) This can be due to the extreme heat input, making the optical fiber brittle.

Optical fiber Cavity Titanium alloy

Fig. 7 Cross-section after FSF showing interlocked fiber within titanium (Shoulder diameter: 10 mm, rotation speed 1750 rpm, travel speed: 25 mm/min, plunge depth: 0.8 mm)

(a)

(b)

500 µm

500 µm

Fig. 8 Cross-section after FSF (a) Shoulder diameter of 10 mm, (b) Shoulder diameter of 15 mm. (rotation speed: 1750 rpm, travel speed: 25 mm/min，plunge depth: 0.8 mm)

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Hardness tests. As shown in figure 9, it can be confirmed that the hardness value increases within the stir zone. This is because of the grain refinement under the probe part of the tool due to stirring. In addition, the hardness value around the optical fiber also increased, which is thought to be due to the TMAZ (thermo-mechanically affected zone) flowing inside the slit.

Fig. 9 Hardness distribution after FSF (The solid blue line shows the hardness of the base metal) Acknowledgment This research was supported by the Amada Foundation 2019 Encouragement Research Grant (Young Researcher Frame AF-2019044-C2). Summary The present study proposed a novel mechanical joining method between the optical fiber and a titanium alloy by using friction stir forming. The possibility of creating a new functional composite alloy and developing innovative materials was investigated experimentally. It was confirmed that embedding fiber inside titanium alloy could be possible by controlling the process parameters of FSF. It was also confirmed that material flow is highly dependent on the process parameters; there were cases where the material flow was insufficient due to low heat input. There were cases where the optical fiber was damaged due to the extreme amount of heat input. Hardness values increased within the stir zone due to the grain refinement under the probe part of the tool due to stirring. References [1] The Ohio State University, Rajendra Singh, 614.292.9044. [2]A. Ogawa, M. Niikura, C. Ouchi, K. Minikawa, and M. Yamada, "Development and Applications of Titanium Alloy SP-700 with High Formability," Journal of Testing and Evaluation 24, no. 2 (1996): 100-109. [3] X. Li, "Embedded Sensors in Layered Manufacturing," Ph.D., Mechanical Engineering, Stanford University, 2001. [4] Y. Li, W. Liu, Y. Feng, and H. Zhang, "Ultrasonic embedding of nickel-coated fiber Bragg grating in aluminum and associated sensing characteristics," Opt. Fiber Technol., vol. 18, no. 1, pp. 7–13, 2012. [5] S. Sandlin, T. Kosonen, A. Hokkanen, L. Heikinheimo, Use of brazing technique for manufacturing of high-temperature fiber optical temperature and displacement transducer, Materials Science and Technology, Vol. 23, Iss. 10, 2007. [6] W.M. Thomas et al. "Friction Stir Welding", International Patent Application No. PCT/GB92/02203.

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[7] T. Nishihara, "Development of Friction Stir Forming", Mater. Sci. Forum, 426-432 (2003) 2971-2978. [8] H. M. Tabatabaei et al., "Friction Stir Forming for Mechanical Interlocking of Ultra-Thin Stainless Steel Strands and Aluminum Alloys", Defect and Diffusion Forum, 382, pp.114-119, 2018. [9] Hamed Mofidi Tabatabaei, Takahiro Ohashi, Tadashi Nishihara, Effect of Friction Stir Forming Parameters and Heat Treatment on Mechanical Properties of Fibre-Reinforced Aluminium Alloy, Key Engineering Materials, 918: 57-66, April 2022. [10] Hamed Mofidi Tabatabaei, Keita Kobayashi, Takahiro Ohashi, Tadashi Nishihara, Effect of Friction Stir Forming Process Parameters on Mechanical Interlock between A5083Alloy and Ultra-Thin Stainless Steel Strands, Key Engineering Materials 858:27-32, August 2020. [11] Hamed Mofidi Tabatabaei, Tadashi Nishihara, Friction Stir Forming for Mechanical Interlocking of Ultra-Thin Stainless Steel Strands and Aluminum Alloys, Defect and Diffusion Forum 382:114-119 January 2018 [12] Hamed Mofidi Tabatabaei, Tadashi Nishihara, Effect of Friction Stir Welding on Mechanical properties of Zn-22Al Superplastic Alloy, International Journal of Engineering Research in Mechanical and Civil Engineering, 2(12) 63-68, December 2017. [13] Hamed Mofidi Tabatabaei, Tadashi Nishihara, Friction stir forming for mechanical interlocking of insulated copper wire and Zn-22Al superplastic alloy, Weld World, 61:47–55, December 2016. [14] Hamed Mofidi Tabatabaei, Takahiro Hara, Tadashi Nishihara, Production of a Superplastic Vibration-Damping Steel Sheet Composite Using Friction Stir Forming, Materials Science Forum 838-839:574-580, January 2016. [15] Takahiro Ohashi, Tadashi Nishihara, Hamed Mofidi Tabatabaei, Mechanical Joining Utilizing Friction Stir Forming, Materials Science Forum, 1016:1058-1064, January 2021. [16] Takahiro Ohashi, Hamed Mofidi Tabatabaei, Tadashi Nishihara, Cylindrical extrusions on A5083 aluminum alloy plate fabricated by friction stir forming, AIP Conference Proceedings, 1896 (1): 080002, October 2017.

Key Engineering Materials ISSN: 1662-9795, Vol. 934, pp 75-80 doi:10.4028/p-wcv8f2 © 2022 Trans Tech Publications Ltd, Switzerland

Submitted: 2022-07-27 Accepted: 2022-09-06 Online: 2022-11-28

Towards a Model-Based Approach for the Optimization of the Mechanical and Economical Properties of Laser-Based Plastic-Metal Joints Julius Moritz Berges1,a, Kira van der Straeten2,b, Georg Jacobs1,c and Jörg Berroth1,d Institute for Machine Elements and Systems Engineering, RWTH Aachen University, Eilfschornsteinstr. 18, 52062 Aachen, Germany

1

Fraunhofer Institute for Laser Technology ILT, Steinbachstr. 15, 52074 Aachen, Germany

2

[email protected], [email protected], c [email protected], [email protected]

a

Keywords: optimization, engineering, simulation

laser

manufacturing,

plastic-metal

joint,

model-based

systems

Abstract. The application of laser-structured metal surfaces to combine plastics and metals is a promising option to enable low-cost lightweight and resource-efficient multi-material joints. The mechanical properties (especially strength), as well as production time and costs, depend on the microstructure of the metal surface (e.g. the number, distance and shape of cavities). Thus, in order to design optimal joints, the properties from the mechanical, as well as production and cost domain, must be considered simultaneously during product development. Therefore, in this paper, a modelbased optimization approach is presented with the goal of identifying an optimum between the strength of the joint and laser manufacturing costs. Parameterized models for the strength estimation, as well as calculation of laser manufacturing time and costs, are developed. The models are linked in an optimization workflow and the optimum positioning and shape of the cavities on the joining zone is determined using a genetic optimization algorithm. The results show that, compared to a benchmark, the manufacturing costs can be reduced by 82 % for the same strength using the proposed model-based optimization approach. Introduction Multi-material structures of thermoplastics and metal are ideally suited to create resource-efficient lightweight structures [1]. However, in order to enable their series application, suitable production processes are required to join the two materials, with their different material properties [2]. In recent years, laser-based direct thermal joining has emerged as a promising technology for joining thermoplastics and metals. With this joining technology, fast process times and low costs can be achieved without the need for additional elements as with screws, or complex process chains as with adhesive bonding [3–5]. The production of laser-based plastic-metal joints is performed in two steps as shown in Figure 1. In the first step, the metal is structured with a laser in order to produce cavities on the metal surface. The high intensity of the laser induces a combination of local melting and sublimation of the metal during the scan. Each laser run changes the shape of the cavity (e.g. depth, undercut, etc.). Thus, this process is usually repeated several times to form cavities with distinctive undercuts in the metal surface (see Fig. 1 step 1). Examples of cavity shapes for two to seven runs can be found in [6]. Theoretically, any geometric pattern can be created on the metal surface. However, due to the simple kinematics, parallel linear structures as shown in Fig. 1 are frequently used. In the second step, the thermoplastic is joined to the metal. Therefore, the plastic is melted and joined with the metal under pressure and temperature using e.g. injection molding [7]. The plastic melts into the cavities and, after solidification, a hybrid joint is formed which takes loads through mechanical interlocking and, depending on the material, specific adhesion.

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1. Microstructuring of the metal surface

Laser Beam

2. Joining with temperature and pressure

Mechanical interlocking and specific adhesion Cavity

Metal

Plastic

Metal

Fig. 1: Two-step manufacturing process of laser-based plastic-metal hybrid joints The present research focuses mainly on maximizing the mechanical properties (especially strength) by determining optimal settings for the distance or shape of the cavities. In this research, the properties (distance, shape of the cavities) are uniform over the entire joining zone and it is not analyzed whether cavities are required on the entire surface. However, not only the strength but also the laser production time and therefore costs strongly depend on the number and shape of the cavities. For example, the laser production time increases when more cavities have to be manufactured or more laser runs are required to produce the desired cavity shape. Thus, to enable optimum joining zones, both mechanical properties and laser production costs must be considered concurrently in the product development of laser-based plastic-metal joints. In this paper, it is investigated whether an optimum between the strength and cost can be determined by the optimal positioning of the cavities in the joining zone. In order to enable the application of optimization algorithms, that usually require a large number of data points, the presented approach is performed based on simulation models. Hence, a finite element micro-model for strength calculation and a novel cost model are used and linked in an optimization workflow. Then, the positioning of cavities on the joining zone is optimized with a genetic algorithm. Finally, the optimized configuration is compared to the classical design as a benchmark, in which the cavities with equal shape and distance are placed over the whole joining zone. Models and Optimization Approach Mechanical Model. The optimization approach is performed using a 2D finite element micro model of a single lap joint specimen (Steel 1.4301, Plastic PP/GF40) as presented in [6] and shown in Fig. 2a. The total length of the specimen is 50 mm, the width is 25 mm and the overlapped length is 12.5 mm. For the plastic, a plastic isotropic material behavior and ductile damage model is applied in the finite element model. The metal is modeled linearly elastic due to the typically small deformations. The material parameters are taken from the materials supplier datasheet. As a non-polar plastic is used, there is only friction (µ = 0.2 [8]), but no specific adhesion, between the plastic and the metal. For the optimization, the joining zone is partitioned into 41 discrete areas with a width of 300 µm (b1 – b41, see Figure 2b). This width is determined in experiments and matches the typical spacing of cavities without causing thermal interaction during laser production, which results in a disturbed forming of undercuts. In each of these areas, it can be located either no cavity or a cavity with dimensions resulting from five laser runs (see Fig. 2c). So, in total up to 41 cavities are possible. In the finite element model, the maximum reaction force RF on the plastic side in the X-direction is used as the target value. The reaction force is related to the joining surface (12.5 mm x 25 mm) and thus the max. strength σmax is calculated.

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(a) Full Model of the specimen

F

Plastic

(b) Discretization

(c) Possible cavity options for each area

12.5 mm Metal

b1 b2 b3

0n

77

b39 b40 b41

5n

300 µm

Fig. 2: Full Model of the single lap joint specimen (a); Discretization of the joining zone (b); possible options for each discrete area, which can be used in the optimization (c) Cost Model. In the following, a model for the calculation of laser manufacturing costs is presented. It should be noted that the following investigations focus solely on the laser production costs, since the material costs and plastic processing costs remain constant independently of the local microstructure, because the total dimensions of the specimen and the joining zone do not change. Laser manufacturing costs CL are calculated based on the laser manufacturing time TL with the machine costs per hour CM for the laser system (equation 1). The laser manufacturing time is calculated via the distance that needs to be structured and the scanning speed vs. For the linear structures discussed in this paper, the total distance can be calculated by the number of cavities n and their respective number of runs Ni with their distance d (300 µm) and the width of component w (25 mm) (see equation 2). Due to the acceleration and deceleration delays of the galvanometer laser scanning system, an additional time delay TD is considered as shown in [9]. The machine costs per hour are calculated by the capital cost of the laser machine CC and the writeoff time tw, the relative maintenance costs per year Cm cost and the building costs approximated by the machine footprint F and the building rent per square meter and year CF (see equation 3). These costs are related to the running hours of the machine per year tr. Furthermore, the energy costs CE are considered. Since the considered laser production machines are typically prototype systems, the values for the laser cost calculation are estimated on the basis of experience (see Table 1). 𝐶𝐶𝐿𝐿 = 𝑇𝑇𝐿𝐿 ∙ 𝐶𝐶𝑀𝑀

𝑇𝑇𝐿𝐿 =

∑𝑛𝑛 𝑖𝑖=1 𝑤𝑤∙𝑁𝑁𝑖𝑖 +𝑑𝑑

𝐶𝐶𝑀𝑀 =

𝑣𝑣𝑠𝑠

(1)

𝐶𝐶𝑐𝑐 ∙(1+𝐶𝐶𝑚𝑚 )+𝐹𝐹∙𝐶𝐶𝐹𝐹 𝑡𝑡𝑤𝑤

𝑡𝑡𝑟𝑟

(2)

+ 𝑇𝑇𝐷𝐷

(3)

+ 𝐶𝐶𝐸𝐸

Table 1: Values for the laser cost calculation Parameter Capital costs CC Write-off time tw Relative maintenance costs Cm Footprint F Building costs CF Running hours tr Energy costs CE

Value 100,000 € 8 years 20 % p.a. 5 m² 3.5 €/m² p.M. 2000 h p.a. 0.405 €/h

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Optimization Setup. The presented finite element model and the cost model are combined in SIMULIA isight (Dassault Systems) and implemented in an optimization workflow. The setup, execution and analysis of the finite element model for the calculation of the strength as well as the cost calculation based on the configuration of the joining zone (position and shape of the cavities) are fully automated using python scripting. In the given optimization problem, the goal is to find a global optimum between the max. strength σmax and the laser production cost CL. Since local optimization methods, such as gradient techniques, are not appropriate for this task, global optimization with an evolutionary algorithm is used [10]. Evolutionary algorithms are widely used for solving engineering problems and are characterized by their high robustness in complex problems and their ability to handle discrete variables. Genetic optimization (GA), a special type of evolutionary algorithms, is based on natural selection in biology, in which only the fittest individuals are selected and then reproduce. The Algorithm NSGA-II is one of the most popular GAs and is applied due to the robustness and the usually fast convergence. The detailed procedure of the algorithm can be found in e.g. [11]. For the given optimization algorithm, variables are assigned to each of the 41 areas from b1 – b41. Any of these variables can have the discrete values of zero or five, representing no cavity or a cavity with five laser runs. Since the dimensions of the joining area are fixed through the constant number of discrete areas, there is no additional geometric constraint required. The maximum strength σmax and the laser production cost CL are defined as optimization target values and are equally weighted. In order to limit the optimization time, the population size is set to 12 and the number of generations is set to 30, resulting in a total number of 360 individuals. A random population is generated at the beginning of the optimization procedure. The settings of the NSGA-II algorithm parameters are given in Table 2. Table 2: Settings of NSGA-II optimization algorithm Option Population Size Number of Generations Crossover Probability Crossover Distribution Index Mutation Distribution Index Initialization Mode

Value 12 30 0.9 10 20 Random

Results and Discussion Figure 3 shows the results of the genetic optimization. Laser production costs and strength are shown as relative values in relation to the benchmark (equal distance of 300 µm, 41 craters, 5 number of runs). The benchmark represents the upper limit for the strength and costs since the maximum number of cavities on the joining zone is used there, leading to the maximum possible strength. For the costs, values below 1 and for the strength values above 1 indicate an improvement compared to the benchmark. In Fig. 3a the relative values are shown over the number of individuals on the x-axis. Over the 360 individuals, the costs and strength tend to decrease during the optimization. It can be seen that the costs of all individuals are significantly lower than the benchmark, while the strength tends towards the value of the benchmark. In Fig. 3b, the relative values are plotted against each other. A correlation between the strength and cost can be observed so that a higher strength is associated with higher costs. Pareto optima can be clearly identified and are shown in blue in Fig. 3b. An optimum at 18% cost and 102% strength can be identified compared to the benchmark (marked in red). By optimizing the positioning of the cavities, the laser production costs can thus be significantly reduced while maintaining similar strength.

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(a)

1.05

0.3

0.85

0.2

0.8 0.75 0.7

0.15

Relative max. load

0.65

50

1 0.95

Pareto frontier

0.9 0.85

Optimum design 18 % costs 102 % max. load

0.8

Relative costs 0

Relative strength

0.25

0.9

Relative costs

Relative strength

Relative values Optimum design Pareto designs

1.05

0.95

0.6

(b)

1.1

1

79

150 100 200 250 Number of individuals

300

350

0.1

0.75

0.1

0.125

Fig. 3: Optimization results

0.15

0.175

0.2

0.225

0.25

Relative costs

Rel. number of cavities

Figure 4 shows the relative number of cavities in comparison to the benchmark (Fig. 4 top) and their allocation in the discrete areas 1 to 41 on the joining zone (Fig. 4 bottom) over the number of individuals. The number of cavities is significantly lower than the benchmark and decreases with the number of individuals. In the optimization process, a concentration of cavities can be identified at the ends of the joining zone. In the middle region, substantially fewer cavities are placed. Considering the well-known stress distribution of single lap joint specimens, it is evident that the stress at the ends is significantly higher than in the middle [3]. Therefore, it is plausible that the greatest demand for cavities is located there, and that the cavities can be reduced correspondingly in the lower stressed middle area. In particular, for the optimized design, almost no cavities are placed in the middle area. 0.7 0.6

Rel. number of cavitites

0.5 0.4

Moving average

0.3

Optimum

0.2 0.1 0 40

Cavity

Area on joining zone

36

No cavity Optimum

32 28 24 20 16 12 8 4 0

0

50

100

150

200

250

300

350

400

Number of individuals

Fig. 4: Number of the cavities and their position on the joining zone during the genetic optimization Figure 5 shows the results of the 2D finite element model for the optimized design in (a) and the benchmark in (b). In both options, the high stress of the cavities at the ends can be observed. In the optimized option, the cavities are almost equally stressed, whereas, in the benchmark, the cavities in the middle region are low stressed. This behavior is a possible reason for the potential to significantly reduce the cavities, especially in the middle region.

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Von Mises Stress (MPa) 3.0 x 102 2.5 x 102 2.0 x 102 1.5 x 102 1.0 x 102 5.0 x 101 0.0 x 100

(a)

(b)

Fig. 5: Optimized design is shown in (a); the benchmark with equal distance is shown in (b) Conclusion and Outlook In this paper, a multi-domain model-based approach for the optimization of the mechanical and economic properties of laser-based plastic-metal hybrid joints is presented. A finite element model for the estimation of the strength and a cost model were implemented in an optimization workflow and a genetic algorithm was applied. The results show that an optimum in terms of strength and costs can be achieved through the optimized positioning of cavities on the joining zone. The laser manufacturing costs are reduced by 82 % compared to the classical design as a benchmark for the same strength properties. The approach is demonstrated using a simple 2D component and solely for parallel linear structures. It is expected that the extension of the degrees of freedom to 3D components and other laser structuring paths (e.g. crossed) holds additional potential. In future, the calculation time of the finite element model must be further reduced to calculate more generations in a shorter period of time (e.g. RVE-homogenization). Furthermore, additional criteria, such as temperature dependence or dynamic load scenarios, shall be implemented in the optimization. Acknowledgements This research has been carried out as part of the project “TailoredJoints”, funding code 03XP0277F which is funded by the German Federal Ministry of Education and Research (BMBF). References [1] B. Bader et al., “Multi material design: A current overview of the used potential in automotive industries,” in Technologies for economical and functional lightweight design, 2019, pp. 3–13. [2] J. F. d. Santos and S. T. Amancio Filho, Eds., Joining of polymer-metal hybrid structures: Principles and applications. Hoboken, NJ: John Wiley & Sons Inc, 2018. [3] G. Habenicht, Kleben: Grundlagen, Technologien, Anwendungen, 6th ed. Berlin, Heidelberg: Springer Verlag, 2009. [4] D. Spancken et al., “Laserstrukturierung von Metalloberflächen für Hybridverbindungen,” Lightweight Design, vol. 11, no. 4, pp. 16–23, 2018, doi: 10.1007/s35725-018-0031-1. [5] A. Klotzbach et al., “Thermal direct joining of metal to fiber reinforced thermoplastic components,” Journal of Laser Applications, vol. 29, no. 2, p. 22421, 2017, doi: 10.2351/1.4983243. [6] J. M. Berges et al., “Model-Based Estimation of the Strength of Laser-Based Plastic-Metal Joints Using Finite Element Microstructure Models and Regression Models,” Materials (Basel, Switzerland), vol. 14, no. 17, p. 5004, 2021, doi: 10.3390/ma14175004. [7] J. Gebauer et al., “Laser structured surfaces for metal-plastic hybrid joined by injection molding,” Journal of Laser Applications, vol. 30, no. 3, p. 32021, 2018. [8] H. Wittel et al., Roloff/Matek Maschinenelemente. Wiesbaden: Springer Fachmedien Wiesbaden, 2013. Accessed: May 11 2021. [9] ScanlabScanlab, Product Brochure: RTC5 control board. [Online]. Available: www.scanlab.de (accessed: Oct. 10 2021). [10] G. Venter, “Review of Optimization Techniques,” in Encyclopedia of Aerospace Engineering, R. Blockley and W. Shyy, Eds., Chichester, UK: John Wiley & Sons, Ltd, 2010. [11] M. Gen et al., Genetic Algorithms and Engineering Design. Hoboken, NJ, USA: John Wiley & Sons, Inc, 1996.

Key Engineering Materials ISSN: 1662-9795, Vol. 934, pp 81-86 doi:10.4028/p-1s2i4z © 2022 Trans Tech Publications Ltd, Switzerland

Submitted: 2022-07-28 Accepted: 2022-09-06 Online: 2022-11-28

Center Line Globular Structure of Strip Cast Using High Speed Twin Roll Caster Toshio Haga Department of Mechanical Engineering Osaka Institute of Technology, 5-16-1 Omiya Asahi-ku Osaka city, 535-8585 Japan [email protected] Keywords: High speed twin roll caster, globular structure, semisolid, dividing of dendrite.

Abstract. Formation of a centerline globular structure in an aluminum alloy strip cast using a vertical type high-speed twin-roll caster was investigated. The theory that the crystals form at tip of the nozzle and are conveyed to the roll gap was investigated by roll casting using a cooling slope and clad-strip-casting in this study. In high-speed roll casting, the center of the strip thickness is semisolid, and the temperature gradient in the thickness direction is small. The latent heat of the semisolid metal heats the dendrite structure, which is divided by melting. The dendrite structure was divided as in thixocasting, which is the mechanism of formation of a globular crystal form. The zone of the semisolid metal becomes narrow as the roll load increases or the pouring temperature of the molten metal decreases. As result, the zone containing globular crystal structures becomes thinner. Introduction Globular crystals exist at the center of cast strips in the thickness direction regardless of the type of aluminum alloy. These crystals form in the center area of aluminum strips cast by high-speed and low load. It has been reported that for horizontal-type high-speed twin-roll casters, globular crystals form at the position where the molten metal contacts the roll, in other words, at the tip of the nozzle and starting point of the strip formation, and then they are conveyed to the roll gap and gather at the center area of the strip [1]. The conveyance of crystals from the tip of the nozzle was investigated using two types of casters. One is a vertical-type twin-roll caster (VHSTRC), without and with a cooling slope, and the other is an unequal-diameter twin-roll caster equipped with a scraper (UDTRCs) for casting clad strips. In this study, experiments were conducted to clarify the origin of the globular structure, and attention was paid to dividing of the secondary arms of the dendrite structure. For aluminum alloy, the roll load is very small in the high-speed twin-roll caster compared with that in the conventional twin-roll caster. The center area in the thickness direction of the strip may be semisolid at the roll bite or just ahead of the roll bite. The secondary arms of the dendrite may be divided as in thixocasting. The divided secondary arms become globular crystals. Latent heat of the semisolid area is released when solidification completes. This heat re-melts the dendrite. If a semisolid metal affects the formation of globular crystals, this means that the thickness of the semisolid area affects the globular crystals. The strips were cast under different conditions, including pouring temperature of the molten metal, roll load, and roll speed. These experiments were conducted to investigate the origin of the globular structure at the center area of the strip. Experiment and Experimental Conditions Effect of Casting Temperature. The VHSTRC shown in Fig. 1(a) was used [2,3]. The rolls were made of copper, and their diameter and width were 300 mm and 100 mm, respectively. Parting material was not used on the roll. The roll speed and load were 60 m/min and 0.56 kN/mm, respectively. The roll speed was much higher and the roll load was much smaller than that for a conventional twin-roll caster for aluminum alloy [4,5]. A cooling slope was attached to the VHSTRC as shown in Fig. 1(b) to carry out semisolid roll casting with a low solid fraction [6]. The cooling slope was made from mild steel 5 mm thick and cooled by an internal water flow. Boron nitride was sprayed on the cooling slope as the parting material. The Al alloy was 6022, and its chemical

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composition is shown in Table 1. The globular crystals formed on the cooling slope and flowed into the molten metal pool made by the side dam and back dam plates. The globular crystal position in the cast strip was investigated. It is presumed that the semisolid metal with a low solid fraction increased the number of globular crystals. If the globular crystals solidified on the roll at the back dam plate, the globular crystals that crystalized on the cooling slope and at the tip of the back dam plate moved to the center area in the thickness direction. As a result, the layer of globular crystals became thicker. Clad Strip. An UDTRCs, shown in Fig. 1(c), was used to cast a clad strip [7]. The diameters of the large and small rolls were 1000 mm and 300 mm, respectively. The width of both rolls was 100 mm, and the roll speed was 30 m/min. The base strip was pure Al, and the overlay strip was Al-10%Mg. The chemical compositions of the pure Al and Al-10%Mg alloy are shown in Table 1. Scraper Crucible Overlay strip Back dam plate Side dam plate

Cooling slope

(a)

Base strip

(b)

(c)

Fig. 1 Schematic illustration of twin-roll casters. (a) Vertical-type high-speed twin-roll caster (VHSTRC), (b) VHSTRC equipped with a cooling slope, and (c) Unequal-diameter twin-roll caster (UDTRCs) equipped with a scraper for clad strip casting.

The liquidus line for Al 6022 is 655°C. The pouring temperature for the VHSTRC in Fig. 1(a) was 660 and 705°C. The pouring temperature of the VHSTRC in Fig. 1(b) was 660°C. For the UDTRCs in Fig. 1(c), the pouring temperatures of the base strip (1070) and cladding (A-10%Mg) were 680 and 640°C, respectively. Alloy 6022 Al-10%Mg 99.7%Al

Table 1 Chemical compositions of aluminum alloys used in this study Si Fe Cu Mn Mg Ti 1.02 0.15 0.01 0.07 0.56 0.02 0.02 0.07 0 0 9.9 0.01 0.05 0.20 0 0 0 0

Al Bal. Bal. Bal.

Results Effect of Casting Temperature. Cross sections of as-cast strips cast from semisolid and molten metals at various temperatures are shown in Fig. 2. Figure 2(a) shows a cross section of the strip cast from the semisolid slurry. In semisolid casting, the temperature of the slurry was 650 to 652°C. The slurry had a low solid fraction, giving it the fluidity of molten metal. Figure 2(b) shows the cross section of the strip cast from molten metal at 660C. Figure 2(c) shows the cross section of the strip cast from molten metal at 705°C. When casting using the cooling slope, the globular crystals formed on the cooling slope, and they existed not only in the center of the thickness direction but also near the surface, as shown by arrow in Fig. 2(a). The large and small globular crystals are mixed as shown in the enlarged view near the center of the thickness direction in Fig. 2(a). The large globular crystal shown by the arrow was formed on the cooling slope.

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Z

50 μm

500 μm

A Z

(a) Semisolid cast strip using the VHSTRC in Fig. 1(b). The left side is a cross section of the strip. The right side is an enlarged view of the center of the cross section.

C

A

B (b) Strip cast from molten metal at 660°C using the VHSTRC in Fig. 1(a). The left side is a cross section of the strip. The right side is an enlarged view of the center of the cross section.

500 μm

50 μm

Z (c) Strip cast from molten metal at 705°C using the VHSTRC in Fig. 1(a). The left side is a cross section of the strip. The right side is an enlarged view of the center of the cross section. 500 μm

50 μm

Fig. 2 Cross sections of as-cast strips and enlarged views of the center of each strip

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In Fig. 2(b), large and small crystals and a dendrite structure existed in the center area. The dendrite structures are shown by arrows marked “A” in the enlarged view in Fig. 2(b). The dendrites shown by arrows A separated from the solidification layer as the grain boundaries were re-melted by the latent heat released from the semisolid metal at the center. Arrow B shows the secondary arms dividing from the dendrite structure, as in thixocasting [8,9]. The secondary arms did not completely divide from these dendrites. Arrow C marks growing crystals. In Fig. 2(c), the zone of the globular crystal shown by arrow Z was wider than that in Figs. 2(a) and 2(b). The sizes of the globular crystals were more uniform than in Fig. 2(a) and 2(b). The thickness of the zone of globular crystals indicated by arrow Z decreased and the number of the fine globular crystal decreased as the pouring temperature of the molten metal decreased. Cross sections of the strips cast at a roll load of 1.12k N/mm, and roll speeds of 30 and 90 m/min, are shown in Fig. 3. The roll load for the strips shown in Fig. 3 was greater than that for the strips shown in Fig. 2. The thickness of the globular crystal zone shown by arrow Z was narrower in Fig. 3(a) than that for the same zone in Fig. 2. There were very fine globular crystals (arrow A), dividing secondary dendrite arms (arrows B) and growing crystal (arrow C) in Fig. 3(a). The black areas in the center zone in Fig. 3(b) are thought to be eutectic. Fig. 3(b) shows very fine globular crystals (arrow A) and dendrites separated from the solidification layer (arrow B).

A

B

C

Z A

B

500 μm

100 μm

500 μm

100 μm

(a) The left side is a cross section of the strip. The right side is an enlarged view of the center. The roll load and speed were 1.12 kN/mm and 30 m/min, respectively. Z: globular crystal zone; A: small globular primary crystal; B: inner dividing dendrite arms; C: growing crystal.

(b) The left side is a cross section of the strip. The right side is an enlarged view of the center. The roll load and speed were 1.12 kN/mm and 90 m/min. A: small globular primary crystal B: dendrites separated from the solidification layer, dark areas that are thought to be eutectic. Fig. 3 Cross section of as-cast strip cast using the VHSTRC shown in Fig. 1(a). Ovelay strip Al-10%Mg

Base strip 99.7%Al 1mm (a) Al-10%Mg

1mm (b) Clad strip with Al-4.8%Mg and 99.7%Al

Fig. 4 Cross section of single and clad strips cast using the caster shown in Figs. 1(a) and (c), respectively.

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Cross sections of a strip of Al-4.8%Mg cast using the VHSTRC in Fig. 1(a) and a clad strip with Al-10%Mg and 99.7%Al cast using the UDTRCs in Fig. 1(c) are shown in Fig. 4. The Al-10%Mg strip contained a band of globular crystals, which did not appear at the interface between the 99.7%Al base strip and Al-10%Mg overlay strip. Discussion Unal [1] showed that globular crystals at the center of the strip were formed at the tip of the back dam plate and decreased in size as they flowed toward the roll bite. If this is true, the quantity of globular crystals increases when a semisolid slurry or low-temperature molten metal is poured, as the quantity of re-melted globular crystals formed on the roll at the tip of the back dam plate decreases and the number of newly formed crystals increases. The results of our experiment indicated the opposite. When molten metal at higher temperature was poured, the amount and temperature of the semisolid metal at the roll gap both increased. The latent heat of the semisolid metal divides the secondary arm of the dendrite structure near the semisolid metal. The quantity of divided secondary arms increased. The quantity of semisolid metal at the roll bite decreased as the roll speed decreased and the roll load increased. The almost semisolid metal usually solidifies at the roll bite. The position of the sump of solidifying material becomes closer to the roll bite as the roll speed increases. When the roll load is not so large and the roll speed is high enough, the center area in the thickness direction of the strip does not completely solidify and semisolid metal remains. In Fig. 3, when the roll speed was 30 m/min, the quantity of globular crystals was small. There were very small globular crystals, which had not accumulated more material, and growing and dividing dendrites were seen in Fig. 3(a). In Fig. 3(b), there were more globular crystals and dividing dendrites than in Fig. 3(a). The number of dividing dendrites in Fig. 3(b) is fewer than that in Fig. 3(a). The roll speed for Fig. 3(b) was three times faster than that for Fig. 3(a). Therefore, the dividing time of the secondary arms was shorter in Fig. 3(b) than that in Fig. 3(a). As a result, the quantity of the divided secondary arms decreased. The roll speed for Fig. 3(b) was three times faster than that for Fig. 3(a). Therefore, the growth time was shorter in Fig. 3(b) than that in Fig. 3(a). As a result, the dendrite structure separated from the solidification layer and many dividing dendrites existed. The black part in Fig. 3(b) might represent eutectic solidification that occurred after the strip was released from the roll pressure, as shown in Fig. 5(b). In contrast, no black area is seen in Fig. 3(a), because the solidification time was sufficiently long, as shown in Fig. 5(b). The sump in Fig. 5(a) is shallower than that in Fig. 5(b). Position of sump

Roll bite

(a)

(b)

Fig. 5 Schematic illustration showing the effect of roll speed on solidification between the rolls, primarily caused by repositioning of the melt sump.

In Fig. 4(a), there is a zone of globular crystal in the centerline. Solidification started at both roll surfaces, and this zone grew and finally solidified. The thickening speed of the front of the solidification layer became very slow, indicating that the temperature gradient was very small in this zone. This zone remained in a semisolid condition with a solid fraction low enough for re-melting and

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re-locating the dendrite divided secondary arms. As a result, globular crystals were formed as in thixocasting [8-10]. If the globular crystals were crystalized at the roll surface where the roll first contacts the molten metal, they would exist at the interface between the 99.7%Al strip and Al-4.8%Mg cladding in Fig. 4(b). However, this was not the case, indicating that the globular crystals formed elsewhere. In Fig. 4(b), the temperature of the 99.7%Al strip, when the molten Al-4.8%Mg was poured, was higher than the solidus temperature of Al-4.8%Mg. The solidification of the Al-4.8%Mg started at the roll surface and moved to the 99.7%Al strip. A temperature gradient existed in the Al-4.8%Mg from the 99.7%Al to the roll surface. As a result, the semisolid metal did not exist long enough to produce globular crystals. In other words, the solidification speed of the solidification front was high enough to prevent formation of a semisolid layer. In this way, the origin of the globular crystals in the center of the thickness direction can be explained as the dividing of the secondary arms of the dendrite structure when the centerline was still semisolid. Conclusion The effect of the casting conditions on the formation of globular crystals in the center of an aluminum alloy strip cast using a high-speed twin-roll caster was investigated to find the origin of the globular crystals. In this study, the pouring temperature of the molten metal, including semisolid casting, and the roll speed were varied. Casting of clad strips using a twin-roll caster was conducted to investigate the origin of the globular crystals. From the results of this study, it can be concluded that the globular crystals in the centerline of the aluminum alloy strip were produced by the secondary dendrite arms dividing. This occurred when enough semisolid metal existed in the strip during casting. References [1] Alcoa. Inc., U.S. Patent 6,672,368 B2. (2004) [2] T.Haga, K.Takahashi, M.Ikawa, H.Watari, A vertical type twin roll caster for aluminum alloy strip, Study on a high-speed twin roll caster for aluminum alloys, Journal of Materials Processing and Technology, 143(2003)895-900 [3] T.Haga, K.Takahashi, M.Ikaw, H.Watari, Twin roll casting of aluminum alloy strips, Journal of Materials Processing and Technology, 153(2004)42-47. [4] A.I.Nussbaum, Three-state-of-the-art Thin –gage high-speed roll caster for aluminum alloy sheet products PartⅢ, Light Met Age, 55(1997)34-39. [5] O.Daaland, A.B.Espedal, M.L.Nedreberg, I.Alvestad, Thin gage twin-roll casting, process capabilities and product quality, Light Met. (1997)745-752. [6] T.Haga, K.Takahashi, H.Watari, Semisolid strip casting using a vertical type twin roll caste, Materials Science Forum, 426(2003)477-482. [7] T.Haga, Casting of Clad Strip Consisting of Al-Sn Alloy and Pure Aluminum, Materials Science Forum Volume 1007(2020)23-28. [8] T.Haga, S.Suzuki, Casting of aluminum alloy ingot for thixoforming using a cooling slope, Journal of Materials Processing and Technology, 118(2001)169-172. [9] T.Haga, P. Kapranos, Billetless simple thixoforming process, Journal of Materials Processing and Technology, 130(2002)581-586. [10] T.Haga, P. Kapranos, Thixoforming of laminate made from semisolid cast strips, Journal of Materials Processing and Technology, 157(2004)508-512.

Key Engineering Materials ISSN: 1662-9795, Vol. 934, pp 87-94 doi:10.4028/p-98un31 © 2022 Trans Tech Publications Ltd, Switzerland

Submitted: 2022-08-09 Accepted: 2022-09-06 Online: 2022-11-28

Investigation of End Milling Process by Machine Tool Equipped with an Idling-Stop Function in a Feed Driving System Yuto MIZUGUCHI1, a, Toshiki HIROGAKI2, b, Eichi AOYAMA3, c 1

Mechanical Engineering, Doshisha Univ., Kyoto Japan

2

Mechanical Engineering, Doshisha Univ., Kyoto Japan

3

Mechanical Engineering, Doshisha Univ., Kyoto Japan

a

[email protected], [email protected], c

[email protected]

Keywords: Sustainable Manufacturing Technologies, Machine Tools, Feed Driving System, Power Consumption, End Mill Process

Abstract. In recent years, as global environmental problems have become more serious, the concept of sustainable development, such as the 3Rs, has become increasingly important. Against this backdrop, machinery and equipment are becoming smaller, lighter, and more sophisticated and multifunctional, and highly integrated products increasingly require the machining and assembly of minute mechanical parts. However, the fact that small machine tools generate relatively large amounts of servo standby power during operation calls for proposals for machining methods that reduce power consumption. In this paper, we have developed a novel 5-axes controlled machine tool equipped with an idling-stop function in a feed driving system to reduce the power consumption of machining processes. A problem has been emerged that start-up power increased as the velocity increased. In the present report, we therefore investigate the influence of various programing paths on the power consumption using an idling-stop function when the end milling process is operated in the X-Y plane. As a result, it can be seen that a suitable programing route is occurs for each feed velocity to reduce the power consumption with a developed machine tool. Introduction In recent years, the concept of sustainable development has been gaining importance due to the increasing seriousness of global environmental problems. Sustainable development is defined as "development that meets the needs of today's generation without endangering the ability of future generations to meet their needs. In other words, sustainability means meeting the needs of the present without jeopardizing the potential to meet the needs of future generations, and to this end, it is required to be sustainable in terms of the three aspects of environment, economy, and society [1], [2]. In the 3R activities, it is generally considered that each of the following is effective in reducing the environmental load: Reduce, Reuse, Recycle, In the 3R activities, it is generally believed that each of the following is effective in reducing the environmental load in the order of Reduce, Reuse, and Recycle, and it is said that the most effective way is to reduce the size of the product itself and to reduce the consumption of resources and energy during the manufacture and operation of the product [3], [4].Against this background, machinery and equipment, including those related to information and communication, are becoming smaller, lighter, and more sophisticated and multifunctional, and highly integrated products are increasingly required to process and assemble small mechanical parts. However, the dimensions of the machines that make them have hardly changed, and the processing and assembly of small parts lacks rationality [5]. In addition, despite the increase in environmental awareness, there are many reports on products and technologies that consider the environmental impact of consumer goods, such as home appliances and automobiles [6], [7]. However, there seem

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to be few research reports on the consideration of the environmental impact of the products [8].Even if there are examples of research reports on production goods, most of them are about predicting energy consumption and improving the efficiency of ancillary systems of machine tools [9], [10]. Therefore, by focusing on the concept of "using a small machine for a small object" based on the 3Rs and using a machine tool that is appropriate for the size of the workpiece, it is possible to not only reduce the environmental impact during product operation, but also to reduce the environmental impact during the product manufacturing stage in the current situation where a significant portion of the environmental impact is generated by the manufacturing department. In addition to reducing the environmental impact during product operation, it is also possible to reduce the environmental impact at the product manufacturing stage and realize environmentally friendly product manufacturing [11]. However, the miniaturization of machine tools has many advantages, such as the small inertia effect on the operation and the ease of speeding up [12], and the fact that small machine tools have relatively high servo standby power consumption during use. The fact that small machine tools have relatively high servo standby power consumption when in use has led to the need to propose machining methods that reduce power consumption [13]. Since the standby power differs depending on the control axis, it is thought that the environmental load at the product manufacturing stage can be reduced by selecting the optimal axis in terms of energy consumption. We have developed a prototype 5-axis controlled machine tool, which is equipped with an idling-stop function (IS) on the X-axis of the translation axis to reduce standby power consumption of the machine tool. By using this IS function, the power consumption during machining is reduced, and the efficiency of energy consumption can increase in some cases. In other words, in tool path patterns that machine geometrically identical shapes, differences in energy consumption occur because of differences in the feed velocity and path of the feed drive system and other motions of each axis [14]. Therefore, it is expected that among several possible tool paths for machining the same shape, one specific tool path exists that consumes the least amount of energy. Furthermore, it is thought that the effect of the IS function can affect the energy consumed in each machining path. However, no systematic study of this has been conducted. In this study, we focused on the plane machining of an end mill and measured the energy consumption in various tool paths to visualize and discuss the power consumption characteristics of each path and the effect of the IS function. To predict the energy consumption without actually operating the tool, simulations were conducted based on the energy consumption at each stage of the path operation. The validity of the simulations was verified by comparing the estimated values with the experimental values. Experimental Equipment

X Z

Y Fig. 1 Outline of developed D5MT equipped with a idling-stop function in X axis feed drive system.

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The prototype 5-axis controlled machine tool (spindle size: BT5, maximum speed: 20,000 rpm) used in this study is an MM400 Lite Multitask (hereinafter referred to as D5MT) manufactured by Iwama Corporation. The D5MT has three translational axes (X, Y, and Z axes) and two swivel axes (A and B axes), as depicted in Fig. 1. The X-axis is equipped with an IS that turns off the servo lock mechanism (the encoder is always on), and the IS state can be recreated when the X-axis fails. IS state is also controlled by M code command in an NC program. Experimental Methods The power consumption of the feed drive system of each axis was measured using the ARTIS monitoring system (GEM-TP, sampling interval of 3 ms) of MARPOSS. It calculates the power

4 mm

4

I.S. ON/OFF

Command

Actual

Com

(a) Linear interpolation. I.S. ON/OFF

2m m

Command

2m m

Actual

Co

(b) Arc interpolation. 4 mm

4

I.S. ON/OFF

45 deg.

Actual Command (c) Diagonal interpolation Fig. 2 Tool path patterns for X- and Y-axis motion.

Co

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consumption required for each axis to move 1 mm. The power consumption of the X and Y axes were measured during constant velocity feed operations at six different velocities ranging from 1000 to 6000 mm/min. The measurements were taken by moving each axis separately. Only the power required for operation was measured. Standby power was excluded. The three tool paths shown in Fig. 2 were operated 10 times and at 15 different velocities from 200 to 3000 mm/min. In each machining path shown in Fig. 2, the interpolation distance was set to 4 mm, assuming mainly interpolation motion at the path edge. Figs. 2(a), (b), and (c) show the linear, arc, and diagonal interpolation motion, respectively. The left side of each figure shows the commanded operation, and the right side figure shows the actual operation. The error between the commanded value and the error when using the IS in the figure on the right was measured. Experimental Results and Discussion

Power consumption Q1mm [-]

Setting evaluation indicators. Fig. 3 shows the results of calculating the operating power consumed per 1 mm at each feed velocity in the X and Y axes. This figure shows that the larger the velocity, the larger the power consumption, which is thought to be due to the increase in the dynamic friction force in the drive system during high-speed operation. In addition, the power consumption of the Y-axis movement was larger than that of the X-axis movement. This is thought to be due to the fact that the Y-axis table is heavier than the X-axis drive unit, thereby consuming more power for the movement.

Velocity F [mm/min] Fig. 3 Moving power consumption. Next, from the results in Fig. 3, the operating power according to the velocity was quantified using the least-squares method, as shown in Eq. 1. Here, Q1mm [1/mm] is the operating power, and a and b are the regression coefficients obtained by the least-squares method. 𝑄1mm = a × 𝐹 + b.

(1)

Fig. 4 shows the start-up power generated when the idling stop is released at each velocity and the interpolation motion. From the results of each interpolation motion shown in Fig. 4, the start-up power increased as the velocity increased for all motion paths. The degree of variation also increased as the velocity and speed increased. The starting power was generally highest for linear interpolation, followed by diagonal interpolation, and then circular interpolation.

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Starting power Q [-]

Starting power Q [-]

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Velocity F [mm/min]

Velocity F [mm/min] (b) Arc

Starting power Q [-]

(a) Linear

Velocity F [mm/min] (c) Diagonal Fig. 4 Starting power consumption. From this result, we quantified the startup power QON [-] using the least-squares method, as shown in

Eq. 1. Using these results, the total power consumption Qpath [-] was defined as shown in Eq. 2: 𝑄path = ∫ 𝑄1mm (𝐹) 𝑑𝐿 + ∑(𝑄ON × 𝑁) + ∑(𝑄i × 𝑇).

(2)

where L [mm] is the travel distance, N is the number of activations, Qi [1/s] is the standby power, and T [s] is the standby time. Comparison of experimental and estimated values. The machining paths simulated in this study are shown in Fig. 5. Each path was assumed to operate on a square plane of 40 × 40 mm. The black section is the area where the IS cannot be reproduced, and the gray section is the area where IS can be reproduced. In the gray section, both patterns of using IS and the normal state are assumed, and experiments and speculations are conducted. Path 1 is a straight path (4-mm spacing) in the Xaxis direction and Path 2 a straight path in the Y-axis direction. Paths 3 and 4 are circular interpolation motions of the path ends of paths 1 and 2, respectively. Path 5 is a diagonal tool path with a normal distance of 4 mm between each path. Path 6 is a spiral path with a distance of 4 mm between paths.

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Path Moving distance L [mm] X-axis movement L X [mm]

1 480 440

2 480 40

3 503 460

4 503 40

5 476 314

6 480 220

Y-axis movement L Y [mm]

40

440

40

460

325

260

480 X-axis 480 travel 500 500LX [mm], 640 Y-axis 480 X + Ltravel Y [mm] Table 1 shows theL tool distance L [mm], distance travel distance Rate of I.S. [%] 8.33 91.7 0 87.5 9.60 54.2 LY [mm], percentage of IS usable area, and number of IS switches for each path. Number of I.S. ON 10 10 0 10 7 10

Table 1 Detail of each tool path patterns. Path Moving distance L [mm] X-axis movement L X [mm]

1 480 440

2 480 40

3 503 460

4 503 40

5 476 314

6 480 220

Y-axis movement L Y [mm]

40

440

40

460

325

260

L X + L Y [mm] Rate of I.S. [%] Number of I.S. ON

480 8.33 10

480 91.7 10

500 0 0

500 87.5 10

640 9.60 7

480 54.2 10

Table 1 Detail of each tool path patterns. A comparison of the inferred results derived using Eq. 2 (based on Fig. 5) and the experimental results is shown in Fig. 6. Fig. 6(a) shows the results of normal operation, and Fig. 6(b) shows the results using the idling stop. The bars show the experimental values, and the grid shows the estimated values. From the experimental values in Fig. 6(a), it can be observed that the smaller the velocity, the larger the energy consumption. This is thought to be because the operation time increased as the velocity decreased, thereby increasing the ratio of standby power. In addition, because the Y-axis table weighs more than the X-axis drive unit, it is reasonable to assume that the energy consumption of paths 1 and 3 is smaller than that of paths 2 and 4. Subsequently, by comparing the experimental and estimated values of the total power consumption in Fig. 6(a), we can see that the evaluation index used here enables the evaluation of the difference in power consumption between routes. Additionally, the experimental values are generally higher than the estimated values. This is thought to be due to the fact that the present evaluation index does not consider the power consumption of the drive system during acceleration and deceleration. The IS use results in Fig. 6(b) suggest that the power consumption can be evaluated but also that the relationship between the experimental and estimated values is not constant. Comparing the experimental values in Fig. 6(a) and Fig. 6(b), it can be seen that the use of IS significantly reduces the energy consumption in paths 2, 4, and 6. This is because

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the ratio of the area where IS can be used is larger in paths 2, 4, and 6 than in paths 1, 3, and 5, demonstrating the effect of the IS. Also, it can be seen that the velocity affects the deference between Fig. 6(a) and Fig. 6(b).

Power consumption Q [-]

loc Ve ity F m [m in] /m el oc ity

Power consumption Q [-]

V

(a) Without I.S.

F

[m

m

/m

in ]

(b) With I.S.

Fig. 6 Comparison of experiment and simulation data.

Summary The IS of the proposed feed axes servo system was examined for end milling by directly measuring the servo motor using a highly responsive power consumption measurement system. It was confirmed that a condition exists where the energy efficiency of IS increased as the velocity decreased, resulting in a larger IS usage ratio. By comparing the estimated values from the set evaluation indexes with the measured values, we managed to evaluate the size of the power consumption depending on the tool path.

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References [1] [2] [3] [4] [5] [6] [7] [8] [9] [10] [11] [12] [13] [14]

Yoshiki Shimomura, Proceedings of the Japan Society for Precision Engineering Autumn Symposium, 10-13, (2006) (in Japanese). H. Koizumi, H. Yahagi, In Search of Sustainability-Learning from Overseas Cities, 2, (2005) (in Japanese). S. Kondo, M. Soma, Y. Umeda, Eco Design Japan Symposium, 153-157, (2002) (in Japanese). Bocken, N.M.P., Allwood, J.M., Willey, A.R., King, J.M.H., Development of an eco-ideation tool to identify stepwise greenhouse gas emissions reduction options for consumer goods, Journal of Cleaner Production, 19(12):1279-1287, (2011). Yuichi Okazaki, Research and Study on Practical Application of MEMS and Microfactories(1), 28, (2004) (in Japanese). Brent, K., Roni, N., Measurement and communication of greenhouse gas emissions from U.S. food consumption via carbon calculators, Ecological Economics, 69(1),186-196, (2009). Yokota, H., Study on methods to reduce exhaust gases from heavy-duty diesel vehicles in use, Journal of Japan society for Atmospheric Environment, 42(1), 1-15, (2007) (in Japanese). Liu, F., Xie, J., Liu, S., A method for predicting the energy consumption of the main driving system of a machine tool in amachining process, Jornal of Clearner Production, 105, 171-177, (2015). Hu, S., Liu, F., He, Y., Hu, T., An on-line approach for energy efficiency monitoring of machine tools, Journal of Cleaner Production 27, 133, (2012). Mori, M., Fujishima, M., Inamasu, Y. and Oda Y., A study on energy efficiency improvement for machine tools, CIRP Annals - Manufacturing Technology, 60(1), 144-148 (2011) (in Japanese). Kurita, T., Hattori, M., Development of new-concept desk top size machine tool, International Journal of Machine Tools and Manufacture, 45(7-8), 959-965, (2005) (in Japanese). Vijayaraghavan, A., Dornfeld, D., Automated energy monitoring of machine tools, CIRP Annals - Manufacturing Technology, 59(1), 21-24, (2010). Hirogaki, T., Aoyama, E., Ogawa, K., Niiyama, T., Suzuki, M. and Iwama, M., Environmental Impact of Desktop-Sized Five-Axis CNC Machine Tool Estimated with LCA, Journal of Environment and Engineering, 6(2), 242-252, (2011) (in Japanese). Akio Hayashi, “Tool Path Evaluation based on Electric Power Consumption of Feed Drive Systems in NC Machine Tool”, Doctoral thesis (Kobe Univ.) (2014) (in Japanese).

Key Engineering Materials ISSN: 1662-9795, Vol. 934, pp 95-102 doi:10.4028/p-iziu0l © 2022 Trans Tech Publications Ltd, Switzerland

Submitted: 2022-08-03 Accepted: 2022-09-06 Online: 2022-11-28

Finite Element Modeling of Upper Ball Joint in a Two-Step Hot Forging Process Nattarawee Siripath1,a, Sedthawatt Sucharitpwatskul2,b and Surasak Suranuntchai3,c* Department of Tool and Materials Engineering, Faculty of Engineering, King Mongkut’s University of Technology, Thonburi, Thailand. 2 National Science and Technology Development Agency (NSTDA), Thailand Science Park Phahonyothin Road, Khlong Nueng, Khlong Luang, Pathum Thani, Thailand. 1, 3

[email protected], [email protected], [email protected]

a

Keywords: Finite Element Modeling, Hot forging process, SNCM8, Upper Ball joint, Qform

Abstract. This paper presented an analysis of the two-step of hot forging process are carried out to manufacture Upper Ball joint, which are including roughing and finishing operation. The part was made from SNCM8 alloy steel. The simulation has been done with application of QForm V10.1.6 software. The constitutive model based on Zener-Hollomon parameter was applied. As a result of simulations, metal flow lines, plastic strain, temperature distribution and effective stress for forgings were obtained. Finite element simulation by QForm V10.1.6. software is suggested as a valuable tool for predicting the hot deformation behavior of material during multi-stage of hot forging process which utilized to enhance the manufacturing process. In addition, the forming load and thickness during the forging process were analyzed. It was found that the deviation of forming load between simulation and experiment was raised to 10.71% and the maximum error of flank thickness was 5.137%. within specification. Therefore, the workpieces of the quality required by specification are obtained. Introduction Metal forming refers to the process in which the given material, usually simple geometry, is transformed into desired shape and size by plastic deformation. The stresses that involve in the process are exceed yield strength of material. Thus, the deformation and its new shape is maintained when the force is removed. Forging is the process in which metals are plastically deformed between two dies using localized compressive forces. Hot forging is widely used in metal industry all over the world. The most common applications for hot forged products are mainly found in the automotive, agricultural, aerospace and construction configurations where strength and durability are required. Hot forging is performed with the workpiece heated above the recrystallization temperature. The hot forging processes is a time consuming with high cost because of the trial−and−error methods which is strongly depends on the designer’s experience [1]. The FE software such as MSC. Marc, ANSYS, QForm, Simufact, etc. is a computer-aided engineering (CAE) tool used for modelling and simulating the engineering problems which commonly used to contribute towards the development and optimization of the metal forming process, leading to enhance the quality of forged product, and reduce both of production cost and time [2]. Many attempts have been discussed the application of simulation results to determine the optimal geometry and dimension of the initial workpiece-billet for control of flash waste, as well as to investigate the various aspects of metal forming. For example, a previous study, Vazquez and Altan [3] proposed the preform design for flashless forging of a connecting rod and introduced a new tooling concept for complex forged parts. Suranuntchai [4] and Wangchaichune et al. [5] reduced the billet size which achieve a final product without any defect as well by using FEA for simulation a Yoke Spline and KVBM gear, respectively, and the effect of several parameters, such as effective stress, plastic strain and temperature were forecasted to enhanced the productivity. Lv et al. [6] successfully presented the simulation of an entire forging process of a gas turbine compressor blade for predicting load/time diagram. During the multi-

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operation process of manufacturing a Hub, Hawryluk et al. [7] used numerical simulation to determine the material flow and properness of impression filling, as well as temperature gradients and plastic deformations in the forging. Sanrutsadakorn et al. [8] simulated an industrial Spacer Wheel part to investigate the forging load and temperature. Jantepa et al. [9] studied the use of FEM as a tool to design the optimal die gap and forming load. The objective of this work is to simulate the two-step hot forging process to manufacture Upper Ball joint, made of SNCM8 alloy steel, in order to determine its final shape and forming load, as well as monitor the temperature, effective stress, and plastic strain distribution by using commercial software QForm V10.1.6. The constitutive equation based on Zener-Hollomon parameter of SNCM8 was employed in this simulation which obtained from experiment. Materials and Experimental Procedure The SNCM8 commercial alloy steel was used in this work, which the chemical composition is shown in Table 1. The flow curves were obtained from hot compression tests under several temperatures (850, 950, 1050 and 1150 °C) and various strain rates (0.01, 0.1, 1 and 10 s-1). The cylindrical specimens with Ø5 x 10 millimeters were prepared. Table 1 The chemical composition (wt %) of the investigated SNCM8 alloy steel. C Si Mn P S Cr Mo Ni V Al W 0.387 0.273 0.695 0.025 0.170 0.768 0.156 1.871 0.010 0.015 0.080 Each specimen was heated up to the given temperature with the heating rate of 10 °C/s and held for 60 seconds for obtaining homogenous temperature distribution. The specimens were upset to a height reduction of 60%, which almost dynamic recrystallization (DRX) could occur completely under this deformation degree. Glass powder was applied on the contacting surface between specimens and dies for reducing the friction. The details of thermo-mechanical processing regime took in this work are displayed schematically in Fig. 1a. The flow curves under different temperatures and at strain rates of 1 s-1 is shown in Fig 1b. (a)

(b)

Fig. 1 (a) Schematic diagram of hot compression test and (b) True stress-strain curves at a strain rate of 1 s-1 under various temperature for SNCM8. Constitutive Model The constitutive model of Zener-Hollomon parameters for hot deformation is gaining wide acceptance that used to characterized the material flow behavior in the simulation. The deformation parameters were fitted using Arrhenius equation and Zener–Hollomon parameter (Z) based on experimental data. The apparent activation energy Q and the constants A, α and n of SNCM8 alloy steel behavior under hot deformation has been investigated by Jantepa et al. [10]. Thus, the constitutive model can be written as follows.

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𝜎𝜎 = 0.011 𝑠𝑠𝑠𝑠𝑠𝑠ℎ

−1

Industrial Process

385.6 ×103

1

��1.019×1014 ∙ 𝜀𝜀̇ ∙ 𝑒𝑒𝑒𝑒𝑒𝑒 �

𝑅𝑅𝑅𝑅

��

1 4.18

97

(1)

�

The material used to manufacture the Upper Ball joint was SNCM8 alloy, and the billet size was Ø48 x 160 millimeters for the hot forging simulation model. The billet was heated up to 1100 °C in the furnace, followed by two forgings in the mechanical press i.e., roughing and finishing forging. After the finishing forging, the flash was trimmed and the workpiece was cooled to room temperature. The material of the upper and lower dies was SKD61 and the dies temperature was set at 250°C. The rougher surface was deeper than the finished surface by about 1 millimeter. The die gaps between the upper and lower dies were 3 millimeters for roughing and finishing forging. The water-based-graphite lubricants was sprayed on the die surface to minimize friction, cooled down the tool surface, prevented sticking and galling of the workpiece to the dies, and reduced erosion of the die surface. The 1350-ton mechanical press with a total stroke of 240 millimeters and a stroke per minute of 85 was used to forge the Upper Ball joint. FE Simulation of Upper Ball joint The 3D-CAD model of the entire hot forging process for Upper Ball joint which consist of the upper die, lower die and the billet is shown in Fig. 2a. The initial 3D-CAD model of dies and billets were transferred to finite element analysis (FEA) software, which was carried out using the latest version of QForm V10.1.6. The tetrahedral elements of the FE model were created. (Fig. 2b). The cylindrical billet, with the diameter of 48 millimeters and a height of 160 millimeters, initially was 2083 nodes, 2612 elements on surface and 8543 volumetric elements. In the simulation, remeshing was required during the numerical solution when large plastic deformation takes place of the element. The workpiece included 83,047 nodes, 87,326 surface elements and 379,285 volumetric elements after the roughing step and 81,011 nodes, 80,062 surface elements and 377,539 volumetric elements after the finishing forging step. The simulation conditions were set according to the actual hot forging process of Upper Ball joint. The cylindrical billet was assigned as a deformable body while the upper and lower dies were assigned as rigid body. The mechanical press was used and its condition was the strokes per minute of 85, stroke length of 240 millimeters and crank radius to conrod length ratio of 0.17. The contact interfaces between the billet and dies was set as shear type, and friction factor was taken to be 0.15 due to the water-based-graphite lubrication. The heat transfer coefficient between the workpiece material and dies material have been assumed 11 N/s·mm·°C. The ambient values used for the heat transfer coefficients were suggested by QForm V10.1.6. library.

(a)

(b)

Fig. 2 (a) 3D-CAD and (b) 3D-FE models of hot forging process for Upper Ball joint.

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Results and Discussions Accuracy of FE software The investigated industrial part was an Upper Ball joint, which was manufactured by two main forging steps i.e., roughing and finishing forging. The main task of roughing forging is mass distribution and to prepare forged part to be shape close enough to the final shape. In the finishing forging step, the forged part is deformed into final shape with correct geometry and dimension. The flash is removed later by a flash trimming operation. The comparison between experimental and simulated workpiece after forging processes is shown in Fig. 3. It was found that the simulation of hot forging process for Upper Ball joint and the experiment were similar. In the simulation, the final shape of Upper Ball joint was completed and no obvious defect is found. The experimental part thickness at region A in roughing and finishing forging step was 11.00−11.50 millimeters and 10.50−11.00 millimeters, respectively, which referred to the specification drawing of shop floor (Fig. 4c). The experimental thickness values were compared with simulated thickness are shown in Table. 2. The average error of thickness at region A was 5.137%, which showed good agreement with simulation results. Thus, the workpieces with the quality as specification required is obtained. The final shape of workpiece is shown in Fig. 4.

(a)

(b)

(c)

Fig. 3 The comparison between actual and simulated workpiece after forging processes (a) roughing, (b) finishing forging and (c) final workpiece after flash trimming.

(a)

(b)

(c)

Region A

Fig. 4 The final shape of (a) simulated and (b) actual workpiece. And (c) the interested region A. The experimental forming load was 1078 and 980 ton in the roughing and finishing forging steps, respectively, which measured directly on the mechanical press. The forming load used in the forging process depicted an average error about 6.376% since the simulated processing condition was a bit different from actual processing condition.

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Table 2 The comparison of thickness at region A after each forging steps and forming load by simulation in QForm V10.1.6. with experiment. Thickness at region A (milimeters) Load forming (ton) Forging steps Experiment Simulation %Error Experiment Simulation %Error Roughing 13.0 12.02 7.538 1078 1056 2.041 Finishing 10.6 10.89 2.736 980 875 10.71 The prediction of metal flow lines During the forging process, the flow lines were distributed as a chain or strip shape along the elongation direction. The forging flow lines could induce anisotropy in the mechanical properties [11]. The mechanical properties of the product were influenced by the flow lines. The flow lines should be arranged in parallel to the working surface; on the other hand, if there were disordered through the flow lines, the mechanical properties will be affected [12]. The distribution of flow lines after roughing and finishing forging operation are presented in Fig. 5a and Fig. 5b, respectively. The formation of flow lines is correlated to the stage of roughing and finishing forging process. At roughing process, the billet is formed thinner to be as near to the final shape as possible. The billet is compressed in the radial direction and the large metal flowed both axially and radially outwards. The distribution of flow lines is along the workpiece due to the larger deformation in this stage. The metal flow in flash region is in the axial direction. The flow lines are densely distribution and demonstrated the continuity and ordered distribution in the cross-section (section A-A). The flow lines in these regions are maintained in the following stage because of the restriction of upper and lower dies. After the roughing process, the workpiece is subjected to the finishing process. At finishing process, the forged workpiece is formed to the final shape with required sizes and geometry. In Fig. 5, the flow lines of the finishing process are large in radial direction in the middle of workpiece. The simulation results showed that there is no breaking and discontinuity of the flow lines in the cross-section area. Hence, there is no forging defects were observed in the two-step hot forging process.

A

A

A A

SECTION A-A (a)

SECTION A-A (b)

Fig. 5 The distribution of flow lines (a) roughing and (b) finishing forging. Plastic deformation distribution The plastic deformation distribution after each forging steps is displayed in Fig. 6. It can be seen that the largest plastic deformation was localized in the flash nearby the concavity which appear in both roughing and finishing forging process which the maximum plastic strains were obtained as 13.85 and 14.12, respectively. As we can see in Fig. 6b, In the finishing forging process, the plastic deformation of the flash area next to the flank edge was increasing. The plastic deformation caused by the heat which generated from the transformation of deformation work and the friction.

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(a)

(b)

Fig. 6 The distribution of plastic deformation (a) roughing and (b) finishing forging steps. Temperature distribution In this section, the initial temperature was 1100°C which was obtained based on the actual condition in shop floor. The temperature distributions within the workpiece after each forging step is shown in Fig. 7. It can be found that the high temperature on the surface and within the workpiece was caused by the conversion of plastic deformation into heat. The temperature within the workpiece (section BB) is greater than the temperature on the surface. Since the long time contact between die and workpiece, the temperature of the workpiece would be drop due to the heat transfer between workpiece and dies. At the end of each operation, the temperature of flash area is higher than other area because of the large deformation and the die friction. B

B

B

B

SECTION B-B (a)

SECTION B-B (b)

Fig. 7 The temperature distribution after (a) roughing and (b) finishing forging steps. The effective stress distribution The amount of forming load undergone is defined by effective stress. Fig. 8 displayed the effective stress distribution of the Upper Ball joint after each forging step. The effective stress is mainly concentrated on flank edge of workpiece nearby the flash region which is the contact areas between workpiece and fillet of the die cavity and the effective stress is gradually decreases on the body of the workpiece. The effective stress in finishing forging step is higher than the effective stress in roughing forging step due to the forming load that applied in the operation.

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(b)

Fig. 8 The distribution of the effective stress (a) roughing and (b) finishing forging steps. Conclusion The simulation of hot forging process for Upper Ball joint which made of SNCM8 alloy steel has been done using commercial QForm V10.1.6 software. Two steps of hot forging process are carried out to manufacture Upper Ball joint, which are including roughing and finishing steps. In the simulation, the constitution model based on Zener-Hollomon parameter is employed and the initial and boundary condition are set based on the actual process. Through analyzed the effect of two-step of hot forging process on temperature, effective stress distribution, plastic strain distribution can be predicted which utilized to enhance the manufacturing process. 1. The deviation of forming load between simulation and experiment was reached to 10.71% and workpiece thickness in region A demonstrated an average error about 5.137%. The accuracy of FEA results depends on the material model and its constants, as well as the input processing conditions in simulation software. 2. The formation of flow lines is correlated to the stage of roughing and finishing forging process. At both roughing and finishing process, the flow line was distributed properly along the Upper Ball joint forging and there is no breaking and discontinuity of the flow lines in the cross-section area. As a result, there is no forging defects were observed in the two-step hot forging process. 3. Plastic deformation is caused by the heat generated by a large transformation of deformation and friction. The largest plastic deformation is in the flash near the concavity, which occurs during both roughing and finishing forging. 4. The temperature within the workpiece is greater than the temperature on the surface of the workpiece. The temperature of the workpiece would be decreased owing to the heat transfer between workpiece and dies. 5. The effective stress is mainly concentrated on flank edge of workpiece nearby the flash region which is the interface between forgings and fillet of the die cavity and the effective stress is gradually decreases on the body of the workpiece. The effective stress depends on to the forming load that applied in the operation. Finite element simulation by QForm V10.1.6. software is suggested as an efficient tool for predicting the hot deformation behavior of material during multi-stage of hot forging process. Future scope of this study may extend about the microstructure evolution for hot forging process using the commercial software QForm V10.1.6. Acknowledgement The authors gratefully acknowledged the financial support by the Petchra Pra Jom Klao Ph.D. Research Scholarship (KMUTT-NSTDA) from King Mongkut’s University of Technology Thonburi. In addition, sincere thanks must be given to S.B.-CERA Co., Ltd. for providing the technique information and industrial part support.

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References [1] [2] [3] [4] [5] [6] [7] [8] [9] [10] [11] [12]

T. Altan, G. Ngaile and G. Shen: Cold and hot forging: Fundamentals and applications, Materials Park, OH, ASM International (2004). J. O. Obiko, F. M. Mwema and M. O. Bodunrin: SN Appl. Sci. Vol. 1 Issue 9 (2019). V. Vazquez and T. Altan: J. Mater. Process. Technol. Vol. 98 (2000), pp. 81-89 S. Suranuntchai: Mater. Sci. Forum Vol. 982 (2020) pp. 106-111 S. Wangchaichune and S. Suranuntachai: Appl. Mech. Mater. Vol. 875 (2018) pp. 30-35 C. Lv, L. Zhang, Z. Mu, Q. Tai and Q. Zheng: J. Mater. Process. Technol. Vol. 198 (2008) pp. 463-470 M. Hawryluk, M. Rychlik, M. Więcław and P. Jabłoński: J. manuf. mater. process. Vol. 5 Issue 2 (2021), pp. 1-13 A. Sanrutsadakorn, V. Uthaisangsuk, S. Suranuntchai and B. Thossatheppitak: Adv. Mater. Res. Vol. 410 (2011), pp. 263-266 N. Jantepa and S. Suranuntchai: Process. Journal of Physics: Conference Series Vol. 1510 (2020), pp 1-3 N. Jantepa and S. Suranuntchai: World J. Mech. Vol. 11 (2021), pp. 17-33 H. Jiang, Y. Wu, X. Gong, D. Shan and Y. Zong: Int. J. Adv. Manuf. Technol. Vol. 106 (2020), pp. 753–764 Y. Zhang, D. Shan and F. Xu: J. Mater. Process. Technol. Vol. 209 (2009), pp. 745-753

Key Engineering Materials ISSN: 1662-9795, Vol. 934, pp 103-108 doi:10.4028/p-r8pk4x © 2022 Trans Tech Publications Ltd, Switzerland

Submitted: 2022-07-27 Accepted: 2022-09-06 Online: 2022-11-28

Casting of Wire Using Twin Wheel Caster with Melt Holding Nozzle Toshio Haga1,a*, Naotsugu Okuda2,b, Hisaki Watari3,c and Shinichi Nishida4,d 1

Osaka Institute of Technology 5-16-1 Omiya Asahi-ku Osaka city, Japan

2

Student of Osaka Institute of Technology 5-16-1 Omiya Asahi-ku Osaka city, Japan 3

Tokyo Denki University, Ishizaka Hatoyama town Saitama 350-0394, Japan 4

Gunma University, 29-1 Honcho Ota city Gunma 373-0057, Japan

a

[email protected], [email protected], [email protected], [email protected]

Keywords: twin wheel caster, wire, nozzle, groove wheel, puddle holding

Abstract. A twin-wheel caster with a hole nozzle was developed to cast thin aluminum alloy wire. The lower large wheel had a trapezoid groove with a cross-sectional area of 22.5 mm2, and the upper small wheel was flat. Molten-metal flow onto the wheel was controlled by the gap between the wheel and the nozzle, and the ejection angle of the molten metal. When the gap was not appropriate, bulges formed on the free solidified surface of the molten metal on the wheel. The bulges were flattened by setting the gap to the correct size of 3 to 4 mm. The appropriate angle was found to be 60°. A thin wire of 6061 aluminum alloy could be cast continuously at a speed of 5 m/min. The cross-sectional area of the nozzle hole was 3 mm2. The cross-sectional area of the as-cast wire was 42.7 mm2. Introduction The demand for aluminum alloy for automotive wire harnesses is increasing because aluminum alloys are lighter and more economical than copper alloys. In particular, lightweight materials are in demand for carbon dioxide emissions reduction. Most aluminum alloy wire is made using the Properzi continuous casting and rolling process [1-8]. The cross-sectional area of the first rods cast using this process is usually 3600 mm2. Cast rods are rolled to a diameter of 9 mm by using 20 three-roll mills in succession. For casting wire, a more compact process is required to reduce the carbon dioxide emissions associated with manufacturing. The present study aims to cast thin wires whose cross-sectional area is smaller than 50 mm2. When the cross-sectional shape of the wire is changed to a circle, the diameter becomes smaller than 8 mm. This diameter is equivalent to the diameter of the wire fabricated by the Properzi continuous casting and rolling process. The process proposed here eliminates the rolling required for size reduction. In this two-wheel process, molten metal is poured through a nozzle into a trapezoid groove machined into the lower wheel, and the upper wheel is flat [9,10]. In this study, we investigated the effect of the gap between the hole nozzle and the lower wheel, the ejection angle of the molten metal against the lower wheel surface and holding of molten metal, during casting of sound wire. A wire with a cross-sectional area of 42.7 mm2 was cast at a wheel speed of 5 m/min. When the cross-sectional shape of the wire was changed to a circle, the diameter was 7.4 mm. Experimental Conditions Twin wheel caster. A schematic illustration of a twin-wheel caster for casting thin aluminum wire is shown in Fig. 1. The upper and lower wheels are made of copper to rapidly solidify the aluminum alloy. Rapid solidification is useful for increasing the casting speed and cooling rate of the cast wire. Two types of lower wheel with different groove sizes were used. The diameter of the upper wheel was 200 mm. The diameter of the lower wheels was 600 mm. The width of the upper and lower wheels was 10 mm. These wheels were not equipped with interior water cooling since the mass of the aluminum alloy was small, for example 500 g, because this caster was a bench-scale laboratory machine. The wheels were each rotated by their own motor.

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The nozzle was inserted into a tundish. The nozzle was made of an alumina-based heat insulator sheet. Photographs of the tundish and the nozzle are shown in Fig. 2. Ejection angle

Nozzle

Rip angle

Upper Wheel

Upper wheel

Upper wheel

Tandish

Gap Solidification length

Lower wheel

Lower wheel

Lower wheel Fig. 1 Left: Schematic illustration identifying the components and showing their geometric arrangement. Right: Two photograhic views of the twin-wheel caster.

Clip Nozzle Hole Tandish

Nozzle

Fig. 2 Photographs of tundish, nozzle and nozzle hole.

Casting Conditions. A 6061 aluminum alloy was used. A wheel speed of 2 to 10 m/min was chosen based on the results of a preliminary experiment. The pouring temperature of the molten metal was 720°C. The molten metal-ejection gas pressure was 5 to 100 kPa. The shapes of the upper and the lower wheels are shown in Fig. 3.

10

φ600

5

φ600

φ125, φ200

2

3 6

2

10

10

2

7

(b) (c) (a) Fig. 3 Schematic illustrations of the (a) upper wheel and (b), (c) lower wheels. Units: mm.

Result and Discussion When the cross-sectional area of the nozzle hole was less than 2 mm2, the molten metal was not ejected at a steady rate. Therefore, a hole with a cross-sectional area of 3 mm2 was used. The

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experimental setup is shown in Fig. 4(a), and the experimental conditions are described below. The speed was 5 m/min, the ejection angle of the molten metal was 70°, the ejection gas pressure was 100 kP, the solidification length was 90 mm, and the cross-sectional area of the nozzle hole was 3 mm2. The as-cast wire is shown in Fig. 4(b). Continuous wire casting was achieved, but ripple projections were formed. The generation of these projections was not restrained by the ejection gas pressure. When the gas pressure was less than 80 kPa, molten metal was ejected at an unsteady rate. When the gas pressure was greater than 100 kPa, for example 200 kPa, the projection formation did not improve, and the mass of the molten metal was too great. The plunger was not suppressed by the wheel speed. When the wheel speed was slower than 2 m/min, the projections became excessive. When the wheel speed was higher 10 m/min, solidification was insufficient. Top view Tundish

Upper wheel

Side view

Projection

Nozzle

Bottom view

Lower wheel

5 mm

Molten metal (a) Casting of a wire

(b) Surfaces of as-cast wire Fig. 4 Cating of a wire and surfaces of as-cast wire. The speed was 5 m/min, ejection angle was 70 º, ejection gas pressure was 100kP, and solidification length was 90 mm

The effect of the solidification length on the projections was investigated. The ejection angle was made small from 70 to 40°. The solidification length was varied between 60 and 90 mm. The casting setup and as-cast wire are shown in Fig. 5. The projection formation was a little improved, but a burr was continuously formed. This means that the quantity of molten metal exceeded the solidification rate when the aluminum alloy was formed by the upper wheel. The width of the burr was not uniform. This means that the projection was not completely improved or molten metal was not ejected at a constant rate. The casting conditions in Figs. 4(a) and 5(a) correspond to the schematic illustrations in Figs. 6(a) and (b), respectively. Tundish

Top view

Nozzle

Side view Upper wheel Molten metal Lower wheel

(a) Casting of a wire

Bottom view

5 mm

Burr (b) Surfaces of as-cast wire

Fig. 5 Cating of a wire and surfaces of as-cast wire. The speed was 5 m/min, ejection angle was 40 º, ejection gas pressure was 100kP, and solidification length was 60 mm.

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Molten metal

Molten metal

Upper wheel

Upper wheel

Puddle 40º

70º

60 mm

90 mm Lower wheel

Lower wheel (b)

(a)

Fig. 6 Schematic illustration shouing castiong situations in (a) Fig.4(a) and (b) Fig. 5(a).

Elimination of the projection, steady flow of molten metal on the lower wheel, and sufficient solidification of the aluminum alloy to prevent burring are essential for casting round wire. The active holding of molten metal flow was attempted to eliminate the projection and achieve a steady flow of molten metal on the lower wheel. Holding of the molten metal was conducted by the tip of the nozzle as shown in Fig. 7. The casting setup is shown in Fig. 8. The wheel speed was 3.5 m/min, the solidification length was 110 mm, and the ejection gas pressure was 10 kPa. When the molten metal was held up by the nozzle, a smaller ejection pressure was suitable. Nozzle

60º

Molten metal 2.5

3.5

Lower wheel

Fig. 7 Schematic illustration showong position of anozzle against a lower wheel.

Fig. 8 Casting of a wire under the condition of Fig. 7 and with a lower wheel of a Fig.3 (b).

The cross sections and surfaces of as-cast wires are shown in Figs. 9(a) and (b), respectively. The cross section of the as-cast wire consisted of two areas “A” and “B,” as shown in Fig. 9(a). The A area is where the molten metal was solidified by the lower wheel, and the B area is where metal was solidified by the upper wheel. The fluctuation of the width of the burr shown in Fig. 9(b) was smaller than that in Fig. 5. The ejection gas pressure decreased from 100 to 10 kPa. This means that the amount of molten metal ejected from the nozzle became constant in spite of the lower ejection gas pressure. This was the effect of holding the molten metal within the tip of the nozzle. The burr became thicker by increasing the solid fraction when the aluminum alloy was formed using the upper wheel. This was the effect of decreased wheel speed and increased solidification length.

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Top view

Interface A

Side view

Bottom view 5 mm

Out line of a lower wheel

(b) Surfaces (a) Cross section Fig. 9 (a) Cross section and (b) surfaces of an as-cast wire cast with the setup of Fig. 8. “A” and “B” in (a) show the areas solidified by the lower wheel and upper wheel, respectively.

The wire had poor contact with the lower wheel, as shown by the cross section in Fig. 9. It was difficult to reduce the ejection gas pressure for the molten metal as the molten metal flow became unsteady at gas pressures less than 10 kPa. The amount of molten metal could not be reduced to make the wire thinner. When the wheel speed increased, the wire became thin. However, the solidification rate decreased, and the burr formation also increased. Therefore, the wheel speed cannot be increased. The wheel with a larger channel was used to increase the contact area between the molten metal and the lower wheel shown in Fig. 3(c). The casting setup shown in Fig. 10 allowed the wire to be cast at a steady rate. The cross section and surfaces of the as-cast wire are shown in Figs. 11(a) and (b), respectively. The cross-sectional area of the as-cast wire shown in Fig. 10(a) was 42.7 mm2. The cross section of the as-cast wire consisted of two areas “A” and “B,” as shown in Fig. 11(a). The A area is where the molten metal was solidified by the lower wheel, and the B area is where the metal was solidified by the upper wheel. The surfaces of the as-cast wire shown in Fig. 11(b) show that the width of the wire became constant. This shows that the amount of ejected molten metal and the solidification rate were constant.

Fig. 10 Casting of a wire under the condition of Fig. 7 and with the lower wheel of a Fig.3 (c).

B Top view

A

Side view

Interface

Bottom view

Out line of a lower wheel (a) Cross section

5 mm

(b) surfaces

Fig. 11 11 (a) Cross section and (b) surfaces of an as-cast wire cast with the setup of Fig. 10. “A” and “B” in (a) show the areas solidified by the lower wheel and upper wheel, respectively.

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Conclusion In this study, a twin-wheel caster with a melt holding nozzle was developed and investigated. The molten metal was ejected onto a grooved lower wheel from a hole nozzle with a cross-sectional area of 3 mm2. A steady flow of molten metal onto the lower wheel was attained by using the nozzle tip to hold the molten metal. The cross section of the as-cast thin wire became uniform by this method. A wire with a cross-sectional area of 42.7 mm2 could be cast at 5 m/min. When the cross-sectional shape of the wire was changed to a circle, the diameter was 7.4 mm. Thus, the objective to cast thin wires with a cross-sectional area smaller than 50 mm2 was attained. References [1] Giuseppe E. Marcantoni, FROM MOLTEN METAL TO 3.2 mm WIRE FOR MECHANICAL APPLICATIONS, Light Metals 2012 Edited by: Carlos E. Suarez TMS (2012). [2] Y. Otsuka, T. Nishikawa, Y. Yoshimoto, Y. Akasofu, Development of Aluminum Wire for Low-Voltage Automotive Wiring Harnesses, SAE Int. J. Passeng. Cars - Electron. Electr. Syst. 5(2012)486-491. [3] M. G. Kim, "Continuous Casting and Rolling for Aluminum Alloy Wire and Rod", Materials Science Forum, 638-642(2010)255-260. [4] M. Jabłoński, Effect of Iron Addition to Aluminium on the Structure and Properties of Wires Used for Electrical Purposes, Materials Science Forum, 690(2011)459-462. [5] K. Buxmann, E. Gold, Solidification Conditions and Microstructure in Continuously Cast Aluminum JOURNAL OF METALS April 1982, Volume 34(1982)28–34. [6] N. Cheung, N. Santos, J. Quaresma, G. Dulikravich, A. Garcia, Interfacial heat transfer coefficients and solidification of an aluminum alloy in a rotary continuous caster, Int. J. Heat & Mass Transfer,52(2009)451-459. [7] D. Lewis, THE PRODUCTION OF NON-FERROUS METAL SLAB AND BAR BY CONTINUOUS-CASTING AND ROLLING METHODS, Metallurgical Reviews, 6(1961) 143- 192. [8] J. Birat, Continuous casting for tomorrow : Near-Net Shape Casting, Rev. Met. Paris, 86(1989) 317-334. [9] Toshio Haga, Naotsugu Okuda, Hisaki Watari, Shinichi Nishida, Caster Equipped with Rotating Side Dam Plates, Science Forum, 1042(2021) 61-67. [10] Toshio Haga, Ryusei Tahara, Hisaki Watari, Shinichi Nishida, Casting of Wire Using a Twin Wheel Caster, Key Engineering Materials, 904(2021)137-142.

CHAPTER 3: Technologies of Welding Production

Key Engineering Materials ISSN: 1662-9795, Vol. 934, pp 111-118 doi:10.4028/p-xa55lv © 2022 Trans Tech Publications Ltd, Switzerland

Submitted: 2022-03-17 Revised: 2022-07-21 Accepted: 2022-07-21 Online: 2022-11-28

Optimization of Resistance Spot Welding Parameters in a Shop Floor Environment to Achieve Desired Spot Size in Low Carbon Steel Sheet Jagadeesh Bagali1,a*, Nanjundaradhya N.V.2,b and Ramesh S. Sharma2,c 1* 2

R & D Centre RVCE, Bengaluru, India

Department of Mechanical Engiinneering., RVCE, Bengaluru, India

*Author for correspondence [email protected], [email protected], c [email protected] Keywords: Low carbon steel sheet, Resistance Spot welding, Nugget size, novel technique to shear spot weld, Profile projector, spot welding parameters and macro and micro examination and electrode rod diameter

Abstract. Resistance Spot Welding (RSW) is extensively employed in transportation, electronics, furniture, Coach building and package industries. Although various spot weld methods are in use, this process is a versatile and easy method to adopt in simple fabrication shop. It is well known that various factors significantly influence the spot size and its quality. Designers, while selecting RSW process are particularly interested in fixing the appropriate spot size and pitch distance for a specific application. Against the above background, this study has been carried out to arrive at desired spot size by varying the parameters using a simple shop-floor spot welder which is readily available in any fabrication shop. Three levels of power input, welding current time with four weld cycle times have been adopted for 2mm thick low carbon steel sheet of IS 513 grades. A novel technique has been developed to apply adequate torque to shear the spot weld joint without any damage or distortion. Spot size has been determined by using vernier and profile projector. The nugget quality has been examined using optical microscope at different magnifications. The bond strength of the spot weld has been determined by tensile shear test. By adopting this simple technique, it is possible to achieve desired spot size which is defect free and having excellent penetration. 1. Introduction Resistance welding is ubiquitously used in fabrication for joining steel and aluminum sheets and parts for automotive sector. Spot nugget is obtained by passing high amperage current to heat the surrounding area by resistance offered by the air-gap leading to a pool of molten metal. The spot size is pre-determined by the design taking into account the geometry of work-pieces. Sufficient localized thrust is applied to achieve coalescence and restrict the nugget formation. . The parameters, that significantly influence the quality and size of nugget, include weld current, welding time, welding force, contact resistance, material properties, coatings, geometry, dimensions of the sheets and machine characteristics[1].Of these parameters, weld current, weld time, electrode force and geometry have much higher influence compared to other factors. The thickness of the sheet, pitch distance (number of welds for given length) and spot size are determined by the designer depending on the application. Over the past three decades, tremendous strides have taken place through advanced analytical tools to precisely simulate the spot welding [2]. Fixing suitable welding parameters for spot welding is difficult task. Large contact surface of the electrode due to wear makes leads to severe restriction in the parameter window. However, in a shop floor practice of spot welding, simpler techniques are required to arrive at approximate range of parameters. The spot diameter is defined as the mean of the major and minor button diameter. This is determined by peel off test of the lap joint (Figure 1).

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Figure 1: Spot weld nugget geometry It is to be recognized that spot weld is accountable for the transfer of loads, when the Resistance Spot Welded panel is subjected to a static or alternating or dynamic load. It is quite obvious that as the spot weld size increases, the static load carrying ability also increases. One general thumb rule for fixing the resistance spot-weld size is that the spot joint must possess a nuggetdiameter of 5√𝑡𝑡, where “t” is the sheet thickness. Thus, a spot weld made in two sheets, each 1 mm in thickness, would create a nugget 5 mm in diameter as per this rule[3].An ideal nugget is one which penetrates into the parent material and leaves a gap between the overlapping sheets as indicated in Figure 1. Enough care has to be taken to select appropriate electrode material and diameter to achieve the desired spot size. Generally copper or tungsten electrodes are chosen and the electrode tip should be sharpened whose diameter should be as per the following formula [5]: Electrode tip diameter = 2.5 + 2t in mm (t = thickness of the sheet). It is also observed that the nugget size varies between 75 – 90% of the electrode weld face [6]. Spot-weld diameter ranges from 3 mm to 12.5 mm. Low carbon steel is most suitable for spot welding.Higher carbon content or alloy steels tend to form hard welds that are brittle and could crack [7]. It is also to be noted that the spacing between two spot welds should be followed as per several guidelines available in the literature so as to achieve good strength, stiffness and energy absorption. Distance between two spot welds is arrived based on sheet thickness. The thumb rule governing the pitch distance is that it should be around ten times the thickness[8]. The size of the weld nuggets has a significant role in components subjected to alternating loads. Further, it is wellknown that smaller nugget size reduces fatigue life. In an effort to reduce the vehicle weight, high strength low alloy steels are finding increasing applications so as to achieve maximum fuel economy [9]. In this context, Davidson [10] has reviewed spot-weld fatigue characteristics and the performance has been analyzed. Fatigue life is re-evaluated in terms of Fracture Mechanics approach. In general, larger weld diameter showed longer fatigue life than the smaller one. Feng1 et al. (11) has investigated the quality of dual phase stainless steel and low carbon steel and found that the nugget of dual phase steel is of narrow width and can be achieved with low current and low cycle time. Mukhtar et al. (12) has studied the nugget characteristics of carbon steel for different thicknesses by using two different joint designs to explore importance of weld joint strength and nugget diameter and determined under different loading conditions. The nugget configuration has an effect on the heat dissipation and determined minimum weld area using lobe curve. Using statistical tool, Vignesh Krishnan et al. (13) studied the influence of welding parameters on tensile nugget properties and microstructure of spot welds of austenitic stainless steel and duplex stainless steel under lap shear loading. It is observed that ductile fractures occur at interfacial mode at higher welding currents. P. Muthu (15) conducted experimental investigations for optimization of welding parameters that influences the tensile shear strength of resistance spot welds stainless steel. Taguchi L27 method. He found that electrode diameter to be the most significant parameter that controls the tensile shear stress, whereas the welding current and heating time are comparatively

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less significant. The percentage contribution of electrode diameter is 66.09%, welding current is 25.08%, heating time is 2.29%, and the residual error is 6.54%. A numerical analysis based on Abaqus has been done by Safari et al (16) to determine predominant mode of failure in stainless steel and found that increase in welding parameters lead to increased propensity of pull-out type of failure. Abhishek Tyagi1 et al. (17) have studied to arrive at optimized parameters using ANOVA and highlighted that welding current is the most significant parameter and the weld quality got improved by 12%. 2. Experimentation With a view to achieve a desired spot size of 6.5mm with 2mm thick low carbon steel (IS 513) sheets, a detailed study of the influence of Electric Resistance Spot welding parameters is undertaken and the results obtained are discussed in this paper. The chemical composition and the specified mechanical properties are indicated in Table 1 and Table 2. Table 1: Chemical composition of Base material (IS 513) Element Specified Actual weight percentage Weight percentage Carbon 0.12 0.09 Manganese 0.50 0.11 Sulphur 0.035 0.010 Phosphorous 0.040 0.020 Silicon 0.009 (max) Aluminum 0.070 0.065 Nickel 0.007 0.004 Table 2: Mechanical properties of Base material (IS 513) as per Test certificate Surface Elongation σys σUTS Hardness Roughness at fracture MPa MPa HRB % (Ra) µm 240 370 31 50 1.60 (specified) (specified) (specified) (Actual) (specified) 229 371 48 1.01 (Actual) (Actual) (Actual) (Actual) Resistance spot welding trials are done on the steel sheet using a simple Resistance Spot Welder (RSW) available in any Fabrication shop. The set-up of the sheet for spot welding is shown in Figure 1.

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Figure 1: Panel in position for RSW The parameters used for the test trials are indicated in Table 3. In all, 12 trials have been carried Table 3: Experimentaltrials of RSW process parameters

Material Electrode size Levels adopted Power (kVA) Weld time (seconds)

Parameter P1 P2 P3 P4 P5 P6 P7 P8 P9 P10 P11 P12

: Low Carbon steel AISI 1023 (2mm thickness) : 7.0 mmElectrode Force : 75 Kgf

: 2.0 ; 2.5 ; 3.0 : 1.4 ; 1.5 ; 1.6

Power (kVA) 2.0 2.5 3.0 2.5 2.5 2.5 2.5 2.5 2.5 2.5 2.5 2.5

Hold time (seconds)

Weld time (s) 1.5 1.5 1.5 1.4 1.5 1.6 1.5 1.5 1.5 1.5 1.5 1.5

: 2 to 7

HoldTime (s) 3 3 3 3 3 3 2 3 4 5 6 7

out with different levels of Power, weld time and weld cycle time. The conical tip diameter of the copper electrode is maintained at 10% more than the desired spot size. The spot welds are separated by using a novel procedure as shown in Figure 2.

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Figure 3: Applying torque load to shear Spot weld joint. 3. Results and Discussions After shearing the spot weld, spot size is measured both by vernier and Profile Projector as shown in Figure 4.

Figure 5: a) Spot weld after shear b) Measurement of spot size using Profile Projector It can be seen from Figure 5 that the sheared portion is clean without any blurs or spatter and facilitates easy measurement of the spot size. This procedure is very easy to adopt in a simple shop floor where facilities are limited. The job can be held either in width direction or thickness direction. A simple hammer strike provides necessary torque on the weld and generates high shear stress that exceeds shear strength and tears apart the spot joint.The measured values are indicated in Table 4. It can be seen from Table 4 that the first 8 trials yielded a spot size lower than 6.0 mm and that necessitated an increase in hold time or weld cycle time. Of the remaining 4 trials, it is found that the trial 9 gave more consistent values of average spot size as indicated Table 5. At this point, it is important to note that thermal conductivity directly influences nugget area. The heat distribution around the contact zone is a complex combination of various factors such as electro-magnetic field developed around the electrode, weld current, electrode force and the thickness of the base metal. Hence, while working with different thickness of the base metal sheet, the weld process parameters have to be reset to get the desired nugget.

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Table 4: Measurement of spot diameter Parameter P1 P2 P3 P4 P5 P6 P7 P8 P9 P10 P11 P12

Spot diameter in mm measured by : Vernier Profile projector 6.0 6.0 6.0 5.8 5.8 5.7 5.0 6.0 6.8 6.9 6.9 7.0

5.3 5.4 5.4 5.4 5.5 5.5 5.2 5.3 6.2 6.3 6.6 6.6

Average spot diameter in mm 5.6 5.7 5.7 5.9 5.6 5.6 5.1 5.7 6.5 6.6 6.8 6.8

The weld nuggets obtained from Trial 9 are examined for its quality using optical microscope. The macro-photographs (Table 6) reveal that the bond is quite satisfactory and also showed good penetration into the parent material. Table 5: Trials for consistency of P9 parameter Spot diameter in mm measured by : Average spot Parameter diameter in Profile projector Vernier mm P9-1 6.8 6.1 6.5 P9-2

6.7

6.2

6.5

P9-3

6.8

6.3

6.5

P9-4

6.6

6.5

6.6

P9-5

6.8

6.0

6.4

P9-6

6.4

6.3

6.4

The bond shear strength is also determined by tensile shear test and the values of shear strength obtained (Table 7) is within the expected values (between 200 to 300 MPa). It is of utmost importance to assess the macro and micro-structures of the weld and HAZ in detail. The features to be focused are geometry of the nugget, uniformity of fusion line, confined region of HAZ, dendrite formation from center to outer regions and process induced weld defects like porosity, blowholes, inclusions etc. These will seriously determine the weld quality and strength of the joint. The features observed are indicated in Table 6.

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Table 6: Examination of weld Nugget with Parameter P9 Specimen

Macro photo at 20x

Macro photo at 50x

Observations

P9 - 1

The spot weld were sectioned using wire EDM to obtain clean images of the nugget. Fusion of the nugget with the base metal is complete with no porosity and cracks. Solidification pattern with weld grains oriented outwards from nugget centre ensures good penetration. It is further observed that after the welding is complete, there is an air-gap of 0.02 mm is left between the over-lapping sheets.

P9 - 2

P9 - 3

P9 - 4

P9 - 5

P9 - 6

Parameter (RSW) P9 - 1 P9 - 2 P9 - 3

Table 7: Tensile shear strength of spot welds Spot size Area Peak Elongation Shear 2 in mm mm Load in N at break strength in mm in MPa 6.2 30.19 7620 6.2 252 6.6 34.21 8550 6.3 250 6.8 36.31 9400 6.5 259

Note :τ𝑠𝑠 (𝑆𝑆ℎ𝑒𝑒𝑒𝑒𝑒𝑒𝑒𝑒𝑒𝑒𝑒𝑒𝑒𝑒𝑒𝑒𝑒𝑒𝑒𝑒ℎ) =

4. Conclusions

Avg. Shear strength in MPa 254

𝑃𝑃𝑃𝑃𝑃𝑃𝑃𝑃𝑃𝑃𝑃𝑃𝑃𝑃𝑃𝑃 𝐴𝐴𝐴𝐴𝐴𝐴𝐴𝐴

Detailed study on resistance spot welding on low carbon steel has been done to identify the parameter ranges to be adopted in a shop floor environment that yields the desired spot size with excellent structural integrity and defect free welds. The spot welding procedure (focusing mainly on electrode tip diameter, weld power and weld cycle time) is easy to maneuver using any type of simple spot weld equipment to achieve the desired spot size and weld strength. A novel procedure has been developed to shear the spot joint to enable easy determination of spot size. The procedure adopted can be extended to Resistance Spot Welding of low thickness low carbon steel panels in the range of 1 to 3mm.

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Acknowledgement The authors wish to thank RVCE management, Principal, Head of the Department of Mechanical Engineering for their support and encouragement in carrying out this work. References [1] [2]

[3] [4] [5] [6] [7] [8] [9] [10] [11] [12] [13]

[14]

[15] [16]

[17]

https://www.swantec.com/technology/resistance-welding/ 2020 Shenghan Guo, Dali Wang, Jian Chen, Zhili Feng, Weihong “Grace” Guo, “Predicting Nugget Size of Resistance Spot Welds Using Infrared Thermal Videos With Image Segmentation and Convolutional Neural Network”, Paper No: MSEC2021-61775, V002T06A011, 2021 http://www.robot-welding.com/Welding_parameters.htm AWS C1, 4M/C1, Specification for resistance welding of carbon and low-alloy steels (2009) https://www.millerwelds.com https://www.howtoresistanceweld.info/spot-welding/what-is-the-formula-to-calculate-nuggetdiameter-in-resistance-welding.html http://www.robot-welding.com/spot_welding.html https://dfmpro.com/wp-content/uploads/2016/02/DFM-Guidebook-Welding-DesignGuidelines-Issue-XVII Welding and Joining of Advanced High Strength Steels (AHSS)2015, Pages 137-165, Woodhead Publishing, edited by Mahadev Shome and MuralidharTumuluru. James A. Davidson, “A Review of the Fatigue Properties of Spot-Welded Sheet Steels by, SAE Transactions, Vol. 92, Section 1: 830008–830359 (1983), pp. 35-47 Q. B. Feng1, Y. B. Li2, B. E. Carlson3 & X. M. Lai4 “Study of resistance spot weldability of a new stainless steel” journal of Science and Technology of Welding and Joining 2019, Vol. 24, no. 2, 101–111 A. M. Al-Mukhtar1,2 and Q. Doos3 “The Spot Weldability of Carbon Steel Sheet”, Advances in Materials Science and Engineering Volume 2013, Article ID 146896, 6 pages Hindawi Publishing Corporation. Vignesh Krishnan1, Elayaperumal Ayyasamy2 and Velmurugan Paramasivam3, “Influence of resistance spot welding process parameters on dissimilar austenitic and duplex stainless steel welded joints” Proc IMechE Part E: Journal Process Mechanical Engineering 2021, Vol. 235(1) 12–23 ©IMechE2020 SAGE publications. Jin-Hee Bae1 Yeong-Do park2 and Mokyoung Lee3 “Optimization of welding parameters for resistance spot welding of AA3003 to Galvanized DP780 STEEL using Response Surface Methodology” International Journal of Automotive Technology Vol 22, No.3 pp, 585-593 (2021). P. Muthu, “Optimization of the Process Parameters of Resistance Spot Welding of AISI 316L Sheets Using Taguchi Method” published by Sciendo Mechanics and Mechanical Engineering 2019; 23:64–69. DOI 10.1007/s12239-021-0055-x P. Muthu*© 2019 M. Safari 1 & H. Mostaan2 & H. Yadegari Kh.3 & D. Asgari3 “Effects of process parameters on tensile-shear strength and failure mode of resistance spot welds of AISI 201 stainless steel”. Asgari3 International Journal of Advanced Manufacturing Technology (2017) vol 89: 1853–1863 © Springer-Verlag London 2016. Abhishek Tyagi a, Gaurav Kumar a, Mukesh Kumar a, Eram Neha b, Mohd Atif Wahid “Experimental investigation for optimization of robot spot welding parameters on low carbon steel JSC 590RN” Journal of Materials Today Proceedings 2214-7853/_2021 Elsevier publications.

Key Engineering Materials ISSN: 1662-9795, Vol. 934, pp 119-128 doi:10.4028/p-to1yyq © 2022 Trans Tech Publications Ltd, Switzerland

Submitted: 2022-03-26 Revised: 2022-05-23 Accepted: 2022-05-31 Online: 2022-11-28

Optimization of Process Variables for Prediction of Penetration Depth of HSLA Steel Welds Using Response Surface Methodology Deepak Pathak1,a*, Dilip Kumar2,b, R. P Singh3,c, and V. Balu4,d Research scholar, Sanskriti University, Mathura, India

1

Assistant Professor, Sanskriti University, Mathura, India

2

Professor and Dean Academics, ITM, Gorakhpur, India

3

Head of Department, Mechanical Engineering Department, Sanskriti University, Mathura, India

4

[email protected], [email protected] c [email protected], [email protected]

a

Keywords: arc welding, high strength low alloy steel, external magnetic field, screening, optimization

Abstract. The statistical model is created for predicting penetration depth in an alternating currentbased additional axial magnetic field controlled shielded metal arc welding of ASTM A 516 Gr.70 steel. The design for the trials is developed using the Placket-Burman design and response surface methodology. The created model determines the optimum process variables for getting excellent penetration depth. The input variables (current, magnetic field density, and magnetic frequency) are chosen for a response like penetration depth. This model can predict the main effects and the interacting effects of three process variables. The findings reveal that a higher current value with a low magnetic field density leads to deeper penetration and vice versa. Furthermore, a greater penetration depth is achieved at lower magnetic field density and higher magnetic frequency. With a desirability of 98.8%, the optimum process variables are 110 A, 0 mT, and 60 Hz. The predicted response values produced from the regression equation based upon process variables are extremely similar to the observed output, demonstrating the usefulness of second-order regression equations. For improved joint efficiency, a high level of penetration is needed. Introduction High-strength low alloy steels (HSLA) are utilized to make components for various industries, and when welded, they have the appropriate technical features. ASTM A 516 grade 70 is an HSLA steel extensively utilized in numerous automobile and petroleum industry applications because of its high demand and low price. It is a pressure vessel grade steel with exceptional impact strength and weldability, ideal for medium and low-temperature applications. Shielded metal arc welding (SMAW) technique is typically utilized in shipping, automobile, aviation, constructions, and petroleum products [1]. External magnetic field (EMF) controlled welding is an upgraded conventional welding method that comes under the advanced welding method [2]. Singh et al used EMF during SMAW process for welding mild steel plates. They found a rise in bead width and penetration depth while decrease in bead height [3]. Liu et al. considered the welding of dissimilar metals under the influence of alternating current (AC) based EMF. They found that magnetic induction intensities and alternating frequency are the most critical EMF factors for improving the joint's bead appearance and tensile strength[4]. This innovative welding process utilizes an EMF throughout the welding process to alter arc form, increase arc consistency, manage metal transmission, improve weld pool solidification, improve welds aesthetics and characteristics[5]. The design of experiments (DOE), analytical methods, and optimizing methods are now commonly employed to create a quantitative correlation and test the findings using the software package. The variables of the Fused Deposition Modelling (FDM) method are optimized by employing the Taguchi approach to get an excellent surface finish by reducing surface roughness[6]. Laser welding parameters were optimized to achieve suitable hardness and length of HAZ by employing the Taguchi methodology[7]. The Taguchi technique was paired with the Single-Response

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Performance Index (SPI) to optimize the post-processing parameters for obtaining minimum surface roughness of the coated items[8]. The turning experiments of pure iron metal utilized the central composite design (CCD) method. Soft computing optimization methods are employed to achieve the lowest surface residual stress[9].The trials are executed following Taguchi’s orthogonal array. ANOVA is then used to evaluate critical parameters. Finally, for the optimization of factors, RSM is used[10]. RSM was implemented successfully for parametric optimization and validated with a confirmation experiment. The factors are correlated, statistically reliable, and convenient for multiple optimizations[11]. A multi-objective optimization approach incorporating the RSM-based DFA is employed to optimize the FSW factors. A regression analysis with a 95 percent confidence interval is created utilizing RSM to estimate the performance of the weldment. The ANOVA approach is utilized to analyze the appropriateness of the constructed model and find the critical terms[12]. RSM determines the model's relevance for DOE by using design expert software[13]. RSM is the most often utilized method for parameterization. As compared to other time-consuming analyses, the goal of RSM for variable optimization was to minimize time & expense. RSM aims to link the factors and the responses [14]. It is employed compared to other approaches since it requires a small number of experiments [15]. It will describe the interactions between various factors and how they influence output results. Three-dimensional response surface plots might be utilized to visualize the response. The CCD is the most widely used DOE in RSM. CCDs are highly beneficial since they provide crucial insight concerning trial variable impacts and overall experimental variance in a few repetitions. RSM was implemented to optimize the welding of steels with activated tungsten inert gas welding [16]. Sood et al. forecasted and optimized the weld bead geometry employing a central composite facecentered (CCF) design method during MIG welding of stainless steel. RSM was used to visualize the findings, and the ANOVA approach was used to ensure that the model was appropriate [17]. When the number of contributing process parameters are large (6-15), it is required to adopt a screening method for selecting significant factors for modeling and optimizing purpose. Placket-Burman design (PBD) is a popular screening approach that logically identifies the significant variables having an extreme influence on the outcome after studying the main effects of contributing factors[18]. Hijazi et al. used PBD to screen five process parameters and optimize the three selected significant process parameters employing RSM [19]. There is a scarcity of research on parametric optimization for axial magnetic field (AMF) aided SMAW of ASTM A516 grade 70 steel. This study determined the optimum process variables for producing the desired reaction by optimizing AMF aided SMAW parameters of ASTM A516 grade 70. The studies were performed to gather information on the impact of process variables on penetration depth. Experiments have verified optimal solutions incorporating RSM-identified process variables. The work is unique because it uses an external AMF based on alternating current in SMAW for ASTM A516 grade 70 steel. The entire welding apparatus was built in the laboratory. Research Outline The current inquiry was organized in the subsequent subsections in order to attain the intended goals: Finding the relevant factors The six contributing factors that impact the penetration depth of weldment were recognized from the literary works [3, 20] and previous efforts[21] at the lab of the current authors, the six contributing factors which influence the penetration depth of welded joints were identified. The factors to consider are current, voltage, speed, groove angle, magnetic field density, and magnetic frequency. Establishing the factors' functional range To determine the acceptable operating limits of AMF assisted SMAW factors, a significant number of practice runs were performed using 8 mm thick plates of ASTM A 516 grade 70 steel. The operating range of each factor was determined by looking for visible flaws in the macrostructure.

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Recognizing the important factors After identifying the contributing factors, the PB screening design is used to select significant factors out of a large number of contributing factors. In this phase, three significant factors are identified. They are welding current, magnetic field density , and magnetic frequency. Creating the design matrix PB Screening design The 12-run PB designs were carried out for the AMF supported SMAW welding of ASTM A 516 grade 70 steel. The contributing factors and their practicable operative ranges are provided in Table 1. The design matrix and the penetration depth (DP) results are given in Table 2. Pareto chart (Fig. 1) shows that I, B, and F were the most significant factors affecting the response DP. Based on a Pareto chart, V, S, and α had no significant influence on the DP. Therefore, they were set to medium levels and removed from the further analysis. All the other three variables were considered. Table 1 Contributing factors and levels Level Factor Symbol Low Medium High Welding Current (A) Iw 90 100 110 Welding voltage (V) Vw 21 23 25 Welding speed (mm/min) Sw 40 60 80 Groove angle (°) α 40 60 75 Magnetic field density (mT) Bw 0 6 12 Magnetic frequency (Hz) Fw 0 30 60 Run order 1 2 3 4 5 6 7 8 9 10 11 12

Iw (A) 90 110 90 110 90 110 90 110 110 90 110 90

Table 2 PB design with Response Sw Vw (V) α (°) Bw (mT) (mm/min) 25 40 40 0 21 80 75 0 25 80 75 0 25 80 40 12 25 80 40 12 25 40 75 0 21 80 75 12 21 80 40 0 21 40 40 12 21 40 75 12 25 40 75 12 21 40 40 0

Fw (Hz) 60 60 60 60 0 0 0 0 60 60 0 0

DP (mm) 3.99 4.38 4.13 4.25 3.76 4.26 3.73 4.16 4.24 3.96 4.11 3.89

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Fig. 1 Pareto chart for DP

Fig. 2 Experimental set-up RSMCCF Optimization design Following the screening phase, three prominent factors welding current, magnetic field density, and magnetic frequency, have been identified and selected for analysis and optimization. Weld bead geometry has been analyzed and assessed using RSM. The impacts of three key parameters, namely current (I), magnetic field density (B), and magnetic frequency (F), on response, namely DP, have been examined. A faced-centered composite (CCF) layout is employed in all of the trials in the research. The design of the existing inquiry, incorporating three factors (I, B, and F), consists of 20 runs. Correspondingly, a factor's lowest and maximum bounds were encoded as +1 and 1. Each coefficient is determined with the design expert 13.0 computational software package and a CCF design. Table 3 provides the selected factors and levels. Table 4 illustrates the full design matrix with an observed response. Table 3 Significant selected factors and levels Levels Process parameters -1 0 1 Iw 90 100 110 Bw 0 6 12 Fw 0 30 60

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Std 10 18 8 9 6 17 2 20 13 5 19 3 11 12 14 4 16 7 1 15

Table 4 Design matrix and observed response Real values Run Iw (A) Bw (mT) Fw (Hz) 1 110 6 30 2 100 6 30 3 110 12 60 4 90 6 30 5 110 0 60 6 100 6 30 7 110 0 0 8 100 6 30 9 100 6 0 10 90 0 60 11 100 6 30 12 90 12 0 13 100 0 30 14 100 12 30 15 100 6 60 16 110 12 0 17 100 6 30 18 90 12 60 19 90 0 0 20 100 6 30

123

DP (mm) 4.38 3.92 5.07 3.43 5.66 3.88 5.13 3.99 4.55 4.66 3.9 3.75 4.26 3.61 4.81 4.73 3.94 4.02 4.4 3.95

Executing the trials and reporting the results A machining method was utilized to obtain the 8 mm thick ASTM A 516 grade 70 steel plates. Table 5 provides the chemical analysis of the base metal utilized in this investigation. The setup (Fig. 2) comprises the AMF creating arrangement with Ac supply to develop a magnetic field, and the motor-supported sliding work table is utilized to regulate consistent welding speed. The copper coil is circumferentially oriented across the electrode. The trial runs have been conducted in two phases (PB screening and RSM optimization). The samples were cleaned with standard metallurgical cleaning techniques and polished with 5 percent nital to exhibit the DP. A digital vernier caliper has been employed to measure the output. Material A516 Gr 70

C 0.18

Table 5 Chemical analysis of the base metal Mn S P Si Cr Cu 1.17 0.002 0.015 0.32 0.013 0.006

Ti 0.012

Al 0.043

Establishing an empirical correlation A professional statistical tool was utilized to analyze observed response (DP) and build a best-fit mathematical model. The following relationship in RSM could describe the response surface Y if all of the process variables (I, B, and F) are measurable, predictable, and consistent with slight error (Eq. 1). DP = f (I, B, F) Eq. 2 can be used to represent the desired second order polynomial for three factors:

(1)

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DP = α0+α1 (I)+ α2 (B)+ α3 (F) + α11(I2) + α22(B2) + α33(F2) + α12(IB) +α13(IF)+α23(BF)

(2)

xi is the process variable, and Y is the response, where α 0 is the average of responses and α 1, α 2... α 23 are approximation values of variables, xi is the input variable, and Y is the response. Eq. 3 represents the complete statistical relationship generated in coded factors to predict the penetration depth of welded joints following the determined significant coefficients. DP (mm) = 3.92745 + 0.471 * I + -0.293 * B + 0.166 * F + 0.0375 * IB + 0.0425 * IF + -0.0225 * BF + -0.0186364 * I2 + 0.0113636 * B2 + 0.756364 * F2

(3)

Evaluating adequacy of the created correlation The analysis of variance (ANOVA) method is applied to conclude whether the developed correlation is appropriate. The model's ANOVA test findings are shown in Table 6. According to the table, the created correlation is acceptable at a 95 percent confidence level. The correlation is significant, as indicated by the model F-value of 371.54. Prob4F values under 0.0500 imply significant association terms. According to the fitted summary, the best-fit model for DP was a quadratic correlation. I, B, F, IB, IF, and F2 are significant model terms. According to the findings, the model is substantial (P < 0.05). Lack of fit value(1.44) indicates that it has not been significantly associated with the pure error. Table 6 also shows the R2 value for the above-created model, indicating that actual and predicted values are highly correlated. The adj R2 of 0.9943 is fairly close to the pred R2 of 0.9723. Table 6 ANOVA Table Source Model Iw-Welding current

SS 6.21 2.22

df 9 1

MS 0.6901 2.22

F-value 371.54 1194.39

p-value < 0.0001 < 0.0001

1

0.8585

462.21

< 0.0001

1

0.2756

148.36

< 0.0001

0.0113 1 0.0113 6.06 0.0145 1 0.0145 7.78 0.0041 1 0.0041 2.18 0.001 1 0.001 0.5142 0.0004 1 0.0004 0.1912 1.57 1 1.57 847.03 0.0186 10 0.0019 0.011 5 0.0022 1.44 0.0076 5 0.0015 6.23 19 0.0431 R² 0.997 4.3 Ad. R² 0.9943 1 Pred. R² 0.9723 AP 72.0682

0.0336 0.0191 0.1706 0.4897 0.6712 < 0.0001

Bw-Magnetic field density 0.8585 Fw-Magnetic frequency Iw Bw Iw Fw Bw Fw Iw² Bw² Fw² Residual lof PE Cor Total SD. Mean C.V. %

0.2756

0.3484

significant

not significant

Optimizing factors The RSM-based DFA is utilized to discover the optimum process variables that yield a maximal or minimal response value. The process variables that correlate to the maximal DP are regarded optimal in this study. The process variables were optimized utilizing design-expert 13.0 software to realize maximal DP. The RSM approach provided one alternative with desirability of 0.988 for getting maximal DP in the specified range. The desirability bar graph for DP is shown in Fig. 3. In

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every welding operation, the penetration depth is crucial since it impacts the strength and efficiency of joining plates. As a result, obtaining quality bead geometry requires maximizing and optimizing penetration depth without causing flaws. Hence maximizing the DP in every welding process is essential for suitable bead geometry. The optimum response location is noticeable on the response surfaces.

Fig. 3 Desirability bar graph for DP

Fig. 4 Perturbation plot

Interpretation of Perturbation plots and response graphs Fig. 4 is a perturbation plot that depicts the variables' direct impact on the penetration depth (DP). The DP increases with an increased welding current (Iw). Increased heat input boosts arc strength, which increases the velocity of the metal transport, leading to greater penetration depth. Welding produces similar outcomes [22]. The findings show that the DP decreases as the magnetic field density (Bw) increases. This might be related to interactions underneath electromagnetic fields; the arc begins spinning and takes on a conical shape, causing the base material to heat up over a larger area, causing an expansive and shallower weld pool. As a result, the penetration depth was lowered. The results are in agreement with other researchers [23]. Results show that DP decreases when magnetic frequency (Fw) increases to its mid-level. While penetration depth increases when magnetic frequency goes beyond its mid-level level. It may be due to a conical arc having enhanced curvature reaching the middle level, leading to non-concentrated energy and enhanced weld pool spinning. DP was reduced as a result of this action. The arc begins to shrink at a frequency band beyond its middle level at the top and base. The energy density is improved due to the compression phenomena, resulting in greater penetration. Similar results are achieved by [24]. It may be demonstrated from Fig. 5(a) that a larger DP is achieved at a higher Iw and a lower level of Bw. It is revealed from Fig. 5(b) that higher DP is achieved at higher Iw and Fw levels. It can be seen from Fig. 5(c) that higher penetration depth can be achieved at a lower Bw and a higher Fw. Moreover, Current has the greatest impact on penetration depth while magnetic field density has the least impact. Magnetic frequency has a moderate effect on penetration depth.

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(a) (b) (c) Fig. 5 Response Graphs for interaction effects of (a) Iw and Bw, (b) Iw and Fw and, (c) Bw and Fw Verification of the model-created optimal solution The concluding phase is to predict and verify the performance variables of the optimized factors once the optimum values of the factors have been determined. Verification experiments were conducted on the model's optimal solutions for attaining the desired responses. Six sets of factors were arbitrarily selected from the generated solutions within the design space for verification using the identical experimental setup. Table 7 displays the projected and observed values and the related percent error (Eq. 4). The percent error is lower than 5%, demonstrating that the statistical models built for the prediction are valid. observed value−projected value

Percentage error = � Experiment no. 1 2 3 4 5 6

projected value

(4)

� X 100

Table 7 Verification test findings Process parameters Magnetic Theoretical Welding Magnetic field response frequency current (A) density (mT) (Hz) 110 0 60 5.63 100 0 60 5.18 110 10 60 5.17 110 0 7 4.92 100 0 0 4.80 99 12 60 4.49

Observed response

Percent error (%)

5.74 5.31 5.11 4.88 4.95 4.37

1.95 2.51 1.16 0.81 3.13 2.67

Conclusions The best combination of welding parameters to obtain maximum penetration depth is carried out using the RSMCCF technique. The following conclusions are obtained from this analysis. 1. During welding circumstances of 110 A welding current, 0 mT magnetic field density, and 60 Hz magnetic frequency, a higher penetration depth of 5.74 mm is attained. 2. The findings of the experiments show that current has the greatest influence on penetration depth, followed by magnetic field density and magnetic frequency.

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3. A higher current value combined with a low magnetic field density leads to a greater penetration. Current levels must be maintained higher to obtain greater penetration and vice versa. Additionally, a greater penetration depth is obtained at lower magnetic field density and higher magnetic frequency. 4. The verification trials are near empirical models at a percent error lower than 5%. The verification findings validated empirical models' capability to anticipate responses reliably. There is a direct proportional association between penetration depth and tensile strength of the joint. It may be inferences that tensile strength also improves with increasing penetration depth. Professionals would determine the optimum process variables for the greatest penetration depth using such models. The findings would be valuable and serve as a technological resource for the welding process, improving continuously. References [1] M.R. Isa, S.N. Sulaiman and O.S. Zaroog, Experimental and Simulation Method of Introducing Compressive Residual Stress in ASTM A516 Grade 70 Steel. Key Engineering Materials 803(2019) 27-31. [2] H. Wu, Y. Chang, J. Bai, Review on magnetically controlled arc welding process. The International Journal of Advanced Manufacturing Technology 91(9) (2017) 4263-4273. [3] R.P. Singh, D. Raghuvanshi, A. Pal, Effect of external magnetic field on weld width and reinforcement height of shielded metal arc welded joints. Materials Today: Proceedings 38(2021) 112-115. [4] Y. Liu, Q. Sun, J. Liu, S. Wang, J. Feng, Effect of axial external magnetic field on cold metal transfer welds of aluminum alloy and stainless steel. Materials Letters 152(2015)29-31. [5] R. Li, X. Yuan, H. Zhang, J. Yang, K. Wu, T. Li, S. Tao, Effect of axial magnetic field on TIG welding–brazing of AA6061 aluminum alloy to HSLA350 steel. Journal of Materials Research and Technology 12(2021) 882-893. [6] J. Nagendra, M. K. Srinath, S. Sujeeth, K. S. Naresh, M. G. Prasad, Optimization of process parameters and evaluation of surface roughness for 3D printed nylon-aramid composite. Materials Today: Proceedings 44 (2021) 674-682. [7] A. Behera, Optimization of process parameters in laser welding of dis-similar materials. Materials Today: Proceedings 33(2020)5765-5769. [8] M.K. Srinath and J. Nagendra, Post-processing parameter optimization to enhance the surface finish of HVOF-developed coatings. Multiscale and Multidisciplinary Modeling, Experiments and Design (2022) 1-13. [9] J. Luo, and Y. Sun, Optimization of process parameters for the minimization of surface residual stress in turning pure iron material using central composite design. Measurement 163(2020)108001. [10] A. Suresh and G. Diwakar, Optimization of process parameters in plasma arc cutting for TWIP steel plates. Materials Today: Proceedings, 38(2021) 2417-2424. [11] A. Aslan, Optimization and analysis of process parameters for flank wear, cutting forces and vibration in turning of AISI 5140: A comprehensive study. Measurement163 (2020) 107959. [12] S. M. Senthil, R. Parameshwaran, S. Ragu Nathan, M. Bhuvanesh Kumar, K. Deepandurai, A multi-objective optimization of the friction stir welding process using RSM-based-desirability function approach for joining aluminum alloy 6063-T6 pipes. Structural and Multidisciplinary Optimization 62(3)(2020)1117-1133.

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[13] Y. Koli, N. Yuvaraj, S. Aravindan, Multi-response Mathematical Modeling for Prediction of Weld Bead Geometry of AA6061-T6 Using Response Surface Methodology. Transactions of the Indian Institute of Metals 73(3)(2020) 645-666. [14] M. Ragavendran, N. Chandrasekhar, R. Ravikumar, R. Saxena, M. Vasudevan, A.K. Bhaduri, Optimization of hybrid laser – TIG welding of 316LN steel using response surface methodology (RSM). Optics and Lasers in Engineering 94 (2017) 27-36. [15] R.H. Myers, D.C. Montgomery and C.M. Anderson-Cook, Response surface methodology: process and product optimization using designed experiments, John Wiley & Sons:New York, (2016). [16] P. Vasantharaja and M. Vasudevan, Optimization of A-TIG welding process parameters for RAFM steel using response surface methodology. Proceedings of the Institution of Mechanical Engineers, Part L: Journal of Materials: Design and Applications, 232(2)( 2015)121-136. [17] Sood, S., et al. Prediction and Optimization of Weld Bead Geometry of MIG Welded Stainless Steel 202 Plates. Advances in Mechanical and Materials Technology (2022) 723-734. [18] A. Al-Sayyad, J. Bardon, P. Hirchenhahn, K. Santos, L. Houssiau, P. Plapper, Aluminum pretreatment by a laser ablation process: influence of processing parameters on the joint strength of laser welded aluminum – polyamide assemblies. Procedia CIRP 74 (2018) 495-499. [19] L. Hijazi, E. Kaiser, S. Altarazi, Pulsed green laser welding of copper materials: A statisticalbased methodology for parameters setting. Procedia Manufacturing 51(2020) 890-896. [20] M.I. Qazi,R. Akhtar, M. Abas, Q. S. Khalid, A. R. Babar, C. I. Pruncu, An Integrated Approach of GRA Coupled with Principal Component Analysis for Multi-Optimization of Shielded Metal Arc Welding (SMAW) Process. Materials 13(16)(2020) 3415. [21] D. Pathak, S. P. Pandey, R. P. Singh, V. Balu, Influence of external axial magnetic field on shielded metal arc weld properties for high strength low alloy steel. Materials Today: Proceedings 2022. [22] A. Choudhary, M. Kumar, D.R. Unune, Investigating effects of resistance wire heating on AISI 1023 weldment characteristics during ASAW. Materials and Manufacturing Processes 33(7) (2018) 759-769. [23] R. Singh, R. C. Gupta, S. C. Sarkar, The effect of process parameters on penetration in shielded metal arc welding under magnetic field using artificial neural networks. International Journal of Application or Innovation in Engineering & Management 2012. 1(4)(2012) 12-17. [24] Z.Q. Guan, H. X. Zhang, X. G. Liu, A. Babkin, Y. L. Chang, Effect of magnetic field frequency on the shape of GMAW welding arc and weld microstructure properties. Materials Research Express, 6(8)(2019) 0865e5.

Key Engineering Materials ISSN: 1662-9795, Vol. 934, pp 129-138 doi:10.4028/p-9oujj7 © 2022 Trans Tech Publications Ltd, Switzerland

Submitted: 2022-04-03 Revised: 2022-06-29 Accepted: 2022-06-30 Online: 2022-11-28

Microstructural and Mechanical Properties of Dissimilar Aluminium Alloys by Friction Stir Welding K. Sekar1,a*, P.Vasanthakumar2,b Department of Mechanical Engineering, National Institute of Technology Calicut, NIT Campus Calicut - 673601, Kerala, India *Corresponding author: [email protected], [email protected] Keyword: Friction stir welding, dissimilar aluminium alloys, microstructure, hardness

Abstract. In the present study, the dissimilar aluminium alloys AA 2014-T651 and AA 7075-T651 of 16 joints was successfully joined with (four welding speeds (50, 60, 70, 80 mm/min), four rotational speeds (900, 1000, 1100, 1200 rpm), four different weld tool profiles (T1=12/4, T2=15/5, T3=16/4, T4=20/5 in D/d ratio) by friction stir welding machine. In the optical microstructure and SEM very fine equiaxed grains are formed in the stir zone. The maximum hardness value is 150 HV obtained in the weld nugget area of the specimen 2 under the conditions of 900 rpm, 70 mm/min for T3 profile. Compare to other tool profiles T3 profile attained maximum hardness value. The minimum hardness value is attained for the specimen 13 for the tool profile T1 was 103 HV compared to all tool profiles. In post weld heat treatment conditions, hardness values were decreased compare with as weld conditions due to over heating followed by ageing. The maximum joint efficiency attained is around 90 to 95% with T2 and T3 profiles at rotational speed of 1000 rpm, 1200 rpm and welding speed of 60, 70 mm/min. The T1 profile attained least joint efficiency of 69.5 % at 900 rpm, 50 mm/min, since it has very less shoulder diameter or contact area and maintain a D/d ratio is 3. Introduction The experimental work joining of AA 2014 T651 and AA 7075 T651 has been divided into subdivision such as pre weld work, fabrication of tool, joint formation and testing. Four FSW tools have been designed by varying dimensions of shoulder and pin diameters since which can influences on quality of joint. In material characterization optical microscope (OM) and scanning electron microscope (SEM) examines the microstructure of welded specimens at different positions such as Heat Affected Zone (HAZ), Thermo Mechanically Affected Zone, (TMAZ) and Stir Zone (SZ). To get information about joint efficiency hardness test carried out in as weld and post weld conditions [1-4]. Ordered raw material of both alloys (AA 2014 -T651, AA 7075- T651) in the form of sheets, length 1400 mm, width 100mm, and thickness 3mm. In this study, have taken weld specimen size as length as 200mm, width is 50mm and thickness is 3mm by shear cutting, need to transform the lengthy plates to my necessary size by shear cutting operation. Then, to get sound joint, the weld samples were cleaned with acetone to avoid unwanted particles on the surface. Sixteen joints were made using different FSW process parameters. Experimental Details Joint formation After all pre – welding functions such as welding sample preparation, tool fabrication, hardening procedure to tool and evaluation of process parameter values. In order to get a proper joint, preweld work must be done properly. Weld formation had performed on the welding machine with computerized friction stir welding machine, as show in Fig.1.

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Fig.1 a) FSW machine, b) Joining of plates by FSW. Specimen preparation After formation of friction stir welded joints, the specimen must be cut to carry out experiments. Preparation of specimens is very crucial, in order to preserve high accuracy, the cutting was carried out using the wire cut electro-discharge machining. (wire cut EDM). The dimensions of the specimen were intended according to ASTM standard. Optical Microscope and SEM Analysis The Optical microscope test was examined to get the microstructural changes of welded specimens. Olympus optical microscope BX51 is used as an equipment to determine microstructures as shown in Fig.2. Polishing and etching have to be done before analyzing the microstructures. The dimension of the specimen is taken as 31.75 mm x 10 mm x 3mm, surface polishing carried out with different grit papers for all microstructure specimens. Kellars agent used as etching agent for the welded specimens for 190 ml of distilled water, 5 ml of nitric acid, 3 ml of hydrochloric acid and 2 ml of hydrofluoric acid and etching time for 15 sec. The specimen is now carefully taken and thoroughly wiped off with water. Microstructures were examined for these etched specimens. SEM analysis has been carried out in various positions such as advancing side, welding zone and retreating side as shown in Fig.3.

Fig.2 (a) Optical microscope BX51, (b) Specimen.

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Fig. 3 (a) SEM set up, (b) SEM specimen. Hardness test

Fig.4 (a) Vickers hardness test set up, (b) Specimen. The Vickers hardness test measurement as shown Fig.4 was performed with a few specifications, which are 25sec is dwell time, 1 kgf is load. With these values test was conducted for 16 samples at distinct places. Sixteen specimens had prepared as per ASTM E – 380 [1] standards with dimensions of 70 mm x 10 mm x 3 mm by wire cut EDM. Hardness test performed in both as weld and post weld conditions. With these hardness values, can comment on joint efficiency. Post weld heat treatment (PWHT) A method known as Post Weld Heat Treatment is conducted frequently to guarantee that material high strength of a component is maintained after welding. PWHT can be used as a technique of controlling hardness or even to improve material strength to decrease residual stresses. Specimens are treated with post weld heat treatment (T6 method), to investigate the variations in hardness value. The specimens are placed in the furnace and heated up to a temperature 485 ºC and soaking procedure is carried for 6 hours. Now, the specimen is carefully taken out from the furnace and quenched in water, followed by solutionizing process. Then specimen is placed back in the furnace and further heated to a temperature of 200 ºC and soaking is carried for 4 hrs. After the completion of above mention process the specimen is taken out from the furnace and exposed in the atmosphere for 3 days. Now the specimen is ready for the hardness test. Total post weld heat treatment carried out in computerized furnace, as shown in Fig.5.

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Fig. 5 (a) Furnace, (b) Quenching in water. Results and Discussion Optical microstructure The microstructures were examined for specimens namely 1, 4, 6, and 8 across the zones of weld, these specimens welded with tool profiles T2, T4, T1 and T3 respectively. In the Heat Affected Zone (HAZ) region, both on the advancing side and on the retreating side the grain structures were elongated, which resembles corresponding to the base metal region. Since the heat contribution by the tool in this region is low, there has been no significant shift in the grain structure and orientation. For all the experimental conditions, the same trend has been observed and can be seen in Fig.6. No significant grain coarsening in the HAZ was observed [3].

Fig. 6 Microstructure of HAZ regions of specimens - 1,4,6,8 Advancing side & Retreating side. In the Thermo Mechanically Affected Zone (TMAZ) region, the grain structure was observed to undergo a rotation with the long axis aligning in the pin direction. The grains in the advancing side of the TMAZ region bend forward in the working position whereas in the retreating side Fig.7, they bend in the reverse manner, smaller grains can be seen in TMAZ zone since much more heat generation than in HAZ. Towards the weld bottom, distinct onion ring patters were seen. The TMAZ, showed severely deformed, uncrystallized grains [4].

Fig.7 Microstructure of TMAZ regions of specimens 1,4,6,8 Advancing side & Retreating side. Severe plastic deformation occurs in the Stir Zone (SZ), which gets elevated heat output owing to friction, drastically altering the grain structure. For all conditions, fine equiaxed grains with proper material mixing have been observed at the SZ and are shown in Fig.8. In advancing side, more fine grains were observed due to severe plastic deformation because of proper heat generation in the faying surface. Microstructure examined in the SZ for different tool profiles.

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Fig.8 Microstructure of SZ regions of specimens 1,4,6 and 8. In Fig.8, the dark phase illustrates precipitation formation in the SZ, as shown in specimen 1, 4, 6 and 8. The precipitation considerably changes in the speed range 60 mm/min and 70 mm/min. Generally, when the temperature and heat exposure duration in the SZ were extended by a decrease in the welding speed, absence of precipitation formed [5-6]. SEM Mircostructure SEM analysis has been carried out to study the morphology of bonding between the dissimilar aluminium alloys. The weld specimen were taken with dimensions of 31.75 mm x 10 mm x 3 mm. Before the SEM examination, the specimens were manually polished with the different grade 100, 200, 600, 1000, 1200, 1500, 2000 grit papers and velvet polishing with the diamond paste to obtain mirror finish. SEM images examined for surface topography from the Fig.9. The SEM micrographs were obtained at three positions of the weld specimen such as advancing side, nugget zone and retreating side. The macrograph of the specimen with clear regions of three regions as shown in Fig.9. When compared with base metal morphology, it was observed that weld area with adequate precipitation mixing. Secondary phases can be observed in a few images due to less heat generation from the tool shoulder and pin to workpiece. The SEM was examined for specimens namely 1, 4, 6, and 8 across the zones of weld, these specimens welded with tool profiles T2, T4, T1 and T3 respectively. When the shoulder diameter is less, heat generation also less. T1 profile was used in sample 6, since it has very less shoulder diameter, mixing of materials was not sufficient. The microstructure in which the formation of Mg2Si are evenly precipitated in aluminium solid solution and fine precipitation of Al2Cu along with MgZn2 [7-8]. The precipitation flakes of Al2Cu and Zinc clusters can be observed in nugget zone. From all the weld specimens, it can be concluded that the morphology of the weld zone and the formation of precipitation depend on the process parameters and tool profiles of the FSW. There is no significant difference in both advancing and retreating side base metal region due to less heat generation.

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Fig.9 SEM micrographs a) AA 7075, b) Nugget Zone, c) AA 2014. From the SEM micrograph it was observed that more fine precipitates was observed in the weld nugget zone of the specimen 1 due to the tool profile geometry of shoulder diameter to pin diameter ratio as 3, whereas more coarse precipitates was observed in the weld nugget zone of the specimen 6 and 8 due to the ratio of D/d is 4. No defects have been observed in specimen 1, 4, 6 and 8. A good bonding has been observed over the advancing and retreating side. A straight and defect free interface is observed in advancing side and retreating side of the specimen 4 due to proper process parameters and less material flash in the contact region of tool shoulder to workpiece. The same were not observed in specimen 1. Hardness test for As-weld conditions Hardness test performed for all 16 samples and the values shown in the Fig.10-11. The length of the specimen was divided into a few parts with a gap of 10 mm between the weld area and base metal, those named as -30, -20, -10, 0, 10, 20 and 30 in mm. Here negative values (-30, -20, -10) belong to the retreating side of the weld specimen that is AA 7075 T651, 0 represents the welded zone or nugget zone and positive values (10,20,30) belong to the advancing side of the welded specimen of the AA 2014 T651. In microstructure such as NZ, TMAZ and HAZ distinct zones can be recognized. Hardness value varies with distinct locations and areas were labeled in the welding area. The nugget zone is 0, the thermally heat affected zone is 10, heat affected zone is 20 and remaining part is base metal of the specimen. The four distinct geometry sizes of the tool and graphs were plotted with respect to the length of the specimen can be shown in below graphs. These four tool profiles were fabricated from different ratio of shoulder diameter to pin diameter. Whenever the contact region or shoulder to workpiece is more sound joints will be created.

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When compared to all tool profiles, the T1 (12:4/3) profile had less shoulder diameter of 12mm or contact area and ratio is 3. This is due to minimum contact area of shoulder to workpiece at elevated rotational and welding speeds. Sound joint created at medium rotational speed and low welding speed using T1 tool profile with maximum time duration during welding [9]. The hardness values against distant across the weld center as shown in Fig.10(a) with the effect of tool profile T1. Except the specimen 6 the hardness values for others specimens comparatively low that is clearly from the Fig.10 (a) shown. The maximum hardness value obtained in the weld nugget area of the specimen 6 under the conditions of 1100 rpm, 60 mm/min. The minimum hardness value obtained to the specimen 13 in the weld nugget area under the conditions of 900 rpm, 50mm/min. The advancing side hardness is more that retreating side hardness due to high temperature input and strain rate than retreating side. The rubbing nature of the tool shoulder to workpiece pushes the material from the weld nugget to the advancing area.

Fig.10 Hardness values of four specimens using (a) T1 profile (b) T2 profile. The measured hardness value from the distance perpendicular to the weld center with the effect of tool profile T2, as shown in Fig. 10(b). The maximum hardness value of 151 HV attained at the weld nugget for at rotational speed of 1200 rpm and welding speed of 80 mm/min to the specimen 10 [10-14]. The minimum hardness value of 110 HV for the specimen 3 when compared to other specimens. The advancing side hardness is more that retreating side hardness due to high temperature input and strain rate than retreating side.

Fig.11 Hardness values of four specimens using (a) T3 profile (b) T4 profile.

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The hardness values against distant across the weld center as shown in Fig.11(a) with the effect of tool profile T3. Except the specimen 2 the hardness values for others specimens comparatively low that is clearly from the Fig.11(b) shown below. The maximum hardness value is 150 HV obtained in the weld nugget area of the specimen 2 under the conditions of 900 rpm, 70 mm/min. The minimum hardness value is 119 HV obtained to the specimen 16 in the weld nugget area under the conditions of 1100 rpm, 80 mm/min. The measured hardness value from the distance perpendicular to the weld center with the effect of tool profile T4, as shown in Fig.11(b). The maximum hardness value of 140 HV attained at the weld nugget for at rotational speed of 1200 rpm and welding speed of 60 mm/min to the specimen 4. The minimum hardness value of 127 - 130 HV for the specimen 5, 15 when compared to other specimens. The advancing side hardness is more that retreating side hardness due to high temperature input and strain rate than retreating side. Hardness test for Post weld heat treatment conditions

Hardness HV

160 140 120 100 80 60 40 20 0

As weld PWHT

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16

Specimen number Fig. 12 Comparison of hardness values of As-weld & Post weld heat treatment conditions. The hardness values were examined for T6 heat treated post weld conditions studied in this section. Regardless of tool profiles and process parameters, the hardness values were reduced due to absence of formation of precipitates. The maximum hardness value were attained for the specimen number 7, 8, 3 when using T3 and T4 tool profile with 135 HV and minimum hardness value attained 75 HV for the specimen 6. Comparison of hardness values of as weld conditions and postweld heat treatment conditions as shown in the Fig. 12. There is much difference between as weld hardness value and PWHT hardness value in specimen 2. All specimens attained less hardness value in PWHT condition when compared with as weld condition. Post weld heat treatment carried out at 200ºC, precipitation dissolution has formed at this temperature that leads to minimum hardness value. The different ageing temperatures of aluminium alloys 2014 and 7075 where 160 ºC and 120 ºC respectively [4-5]. The approximately 40 ºC difference of ageing leads to minimum hardness in the weld nugget area. In this study, we used T6 heat treated condition of 200 ºC for 6 hours followed by natural ageing for 72 hours, the 2014 and 7075 are basically heat treatable alloys during this T6 heat treated condition the precipitation formation of S type (Al2CuMg) and T type (Mg3Zn3Al2) phases, which decreases the strength of the joint in PWHT [9]. Conclusion In the present study, the dissimilar aluminium alloys AA 2014 T651 and AA 7075 T651 was successfully joined by friction stir welding machine. The following are main conclusions are drawn.  In the optical microstructure very fine equiaxed grains are formed in the stir zone due to severe plastic deformation and elongated grains are observed in base metal region, when the specimen is examined at different regions across the weld center.

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 In the SEM micrograph, it was observed that more fine precipitates are formed in the weld nugget zone of the specimen 1 due to the tool profile geometry of D/d ration as 3, whereas more coarse precipitates was observed in the weld nugget zone of the specimen 6 and 8 due to the ratio of D/d as 4.  The maximum hardness value is 150 HV obtained in the weld nugget area of the specimen 2 under the conditions of 900 rpm, 70 mm/min for T3 profile. Compare to other tool profiles T3 profile attained maximum hardness value.  The minimum hardness value is attained for the specimen 13 for the tool profile T1 was 103 HV compared to all tool profiles.  In post weld heat treatment conditions, hardness values were decreased for all specimens compare with as weld conditions due to overheating followed by ageing.  The maximum joint efficiency attained is around 90 to 95% with T2 and T3 profiles at rotational speed of 1000 rpm, 1200 rpm and welding speed of 60, 70 mm/min. The T1 profile attained least joint efficiency of 69.5 % at 900 rpm, 50 mm/min, since it has very less shoulder diameter or contact area and maintain a D/d ratio is 3. References [1] V. Pandian and S. Kannan, “Numerical prediction and experimental investigation of aerospacegrade dissimilar aluminium alloy by friction stir welding,” J. Manuf. Process., vol. 54, no. February, pp. 99–108, 2020. [2] K. Sekar and P. Vasanthakumar, “Microstructural evaluation of similar and dissimilar welding of aluminum metal matrix hybrid composite by friction stir welding,” Mater. Sci. Forum, vol. 979 MSF, pp. 124–128, 2020. [3] V. Saravanan, Nilotpal Banerjee, R. Amuthakkannan, S. Rajakumar, Microstructural Evolution and Mechanical Properties of Friction Stir Welded Dissimilar AA2014-T6 and AA7075-T6 Aluminium Alloy Joints, Metallogr. Microstruct. Anal. (2015) 4:178–187. [4] Sarpreet Singha, Gaurav Dhuriab; Investigation of post weld cryogenic treatment on weld strength in friction stir welded dissimilar aluminium alloys AA2014-T651 and AA7075-T651, Materials Today: Proceedings 4 (2017) 8866–8873. [5] Guven ipekoglu and Gurel Cam; Effects of Initial Temper Condition and Postweld Heat Treatment on the Properties of Dissimilar Friction-Stir-Welded Joints between AA7075 and AA6061 Aluminium Alloys, The Minerals, Metals & Materials Society and ASM International 2014. [6] S. Babu, G.D. Janaki Ram, P.V. Venkitakrishnan, G. Madhusudhan Reddy and K. Prasad Rao; Microstructure and Mechanical Properties of Friction Stir Lap Welded Aluminium Alloy AA2014, J. Mater. Sci. Technol., 2012, 28(5), 414–426. [7] Aruri Devarajua, V Kishanb; Influence of Cryogenic cooling (Liquid Nitrogen) on Microstructure and Mechanical properties of Friction stir welded 2014-T6 Aluminium alloy, Materials Today: Proceedings 5 (2018) 1585–1590. [8] C. Rajendran, K. Srinivasan, V. Balasubramanian, H. Balaji & P. Selvaraj, Identifying the combination of friction stir welding parameters to attain maximum strength of AA2014-T6 aluminium alloy joints, 2017. [9] P. Vijaya Kumar, G. Madhusudhan Reddy , K. Srinivasa Rao, Microstructure, mechanical and corrosion behavior of high strength AA7075 aluminium alloy friction stir welds e Effect of post weld heat treatment, Defence Technology 11 (2015).

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[10] Sekar. K, Vasanthakumar. P Friction Stir Welding of Al-Cu Alloy Metal Matrix Composites Reinforced with B4C and Graphite Particle Fabricated by Stir Casting and Thixoforming Method, M. S. Shunmugam and M. Kanthababu, “Advances in Additive Manufacturing and Joining,” in Editors Proceedings of AIMTDR, 2018. [11] P. Vasanthakumar, K. Sekar, and J. Jayantherababu, “Thermal prediction and experimental validation of Friction Stir Welded Aerospace Grade Aluminium Alloy,” J. Phys. Conf. Ser., vol. 1240, no. 1, 2019. [12] Shine, K., and K. Jayakumar. "Effect of tool pin profile on the mechanical and microstructural properties of dissimilar friction stir welded AA5083-H111 and AA6061-T6 aluminium alloys." JOURNAL OF THE CHINESE INSTITUTE OF ENGINEERS 45, no. 3 (2022): 227-236. [13] P. Naveen Kumar, and K. Jayakumar. "Influence of tool pin profiles in the strength enhancement of friction stir welded AA5083 and AA5754 alloys." Materials Research Express 9, no. 3 (2022): 036505. [14] Balamurugan, S., K. Jayakumar, and K. Subbaiah. "Influence of friction stir welding parameters on dissimilar joints AA6061-T6 and AA5052-H32." Arabian Journal for Science and Engineering 46, no. 12 (2021): 11985-11998.

Key Engineering Materials ISSN: 1662-9795, Vol. 934, pp 139-151 doi:10.4028/p-e4uahg © 2022 Trans Tech Publications Ltd, Switzerland

Submitted: 2022-04-07 Revised: 2022-07-22 Accepted: 2022-07-22 Online: 2022-11-28

Analysis of Thermal Fields, Weld Strength and Microstructural Studies of Friction Stir Dissimilar Weldments of AA6082 and AA7075 M.V.A. Ramakrishna1,a*, K. Srinivas2,b Department of Mechanical Engineering, Acharya Nagarjuna University, Guntur, Andhra Pradesh 522510, India

1

Department of Mechanical Engineering, R.V.R. & J.C.College of Engineering, Chowdavaram, Andhra Pradesh 522019, India

1

[email protected], [email protected]

a

Keywords: IR Thermography, Friction stir welding, Dissimilar weldments, Characterization.

Abstract. This research work aims to investigate thermal fields, weld strength and microstructural analyses of dissimilar welded structures of AA6082 and AA7075 developed using friction stir welding process. Three different tool shapes were used for joining of these plates to study the effect tool profiles on thermal cycles and their behavior on weld strength. The process parameters were considered same for three tool weldments. During welding, peak temperatures and their distribution towards along transverse direction and cooling temperature were recorded using Infra-Red (IR) thermography. The samples were inspected were X-ray radiography test to examine the defects. The tensile strength of the welded structures was evaluated by conducted tension test according to ASTM E/8 standards. Further, the microstructural behavior for three tool profile pins was studied using optical microscopy. The tool profile with cylindrical shape exhibiting higher tensile strength than the other two samples due to lower temperatures and uniformity of heat distribution during welding process. Also, the microstructural studies revealed that the fine grains with proper mixing zones were observed when cylindrical tool was employed. From this research work, IR thermography is most suitable for measuring In-situ temperature profiles even at weld vicinity during welding. Introduction Weldments should possess good strength and creep resistance to avoid the failures during their service life. Aluminium based alloys are generally used in various applications including aerospace automotive industries, and shipbuilding due to their lower weight density compared to other metals like steels, high strength metals [1, 2]. Aluminium 6000 series, especially 6082 alloys are used in shipbuilding and aircraft industries. Mg and Si are the main elements in AA6082 [3]. Also, Aluminium 7000 series, especially AA7075 (Mg-Zi alloy), are used in aircraft applications because of their high specific weight strength and corrosion resistance at high temperatures [4, 5]. Robert et al. [6] investigates the low cycle fatigue properties of AZ31- Mg in friction stir welding with constant tool rotation speed 1600 rpm and varying welding speeds between 150-900 mm/min. The low cycle fatigue test was conducted on three stages i.e. short, medium and long periods and in all these cases, the failure occurred at HAZ region. Tian et al. [7] studied the effect of welding speed of Friction Stir Welding (FSW) on mechanical properties and microstructure in stir zone using friction stir welding. With the increase of welding speed the grain size was decreases and vice-versa. By increasing welding speed, the tensile properties increased up to certain speed and then gradually decreased. The Selection of tool depends on its features like temperature, wear resistance and toughness. H13 steel tool is extensively used at high-temperature applications due to its alloying elements especially Mo, Cr and Vanadium [8]. The vanadium and Mo strengthen the H13 steel tool. H13 tool steel provides good combination of abrasion and shock resistance, and also has good ductility, machinability [9]. Krishna et al. [10] studied the effect of tool offset on aluminum AA6351 alloy friction stir weldments (similar joints). From the results, it is concluded that the tool offset has no influence on the tensile strength of the joint compared to non-offset weld joint.

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Ugender et al. [11] investigated the effect of process parameters on AZ31 magnesium alloy mechanical properties in friction stir welding. To know the effect of process parameters, Taguchi L9 orthogonal array was used. ANOVA was used to observe the significance of each parameter and it was observed that tool profile has less significance than tool transverse speed and tool rotation speed. Kulwanth et al. [12] investigated the microstructural studies and mechanical properties of AZ61 magnesium alloy joints. The hardness of the joint improved due to the finer grains in the stir zone. The hardness of the base metal is more than that of thermo-mechanically affected zones but lower than that of the stir zone. Defect free stir zone resulted because of uniform deformation at the joint. Bocchi et al. [13] analyzed the effect of process parameters on mechanical properties for the joints AA7075 and AA2024 and also studied their corrosion behaviour. The best mechanical properties were observed at intermediate speed and feed rates. There is no relation between process parameters on corrosion behaviour. Balram et al. [14] joined the dissimilar plates of AA5052 and AA6061 with ER4043 and ER5356 fillers. The welded samples were free from internal defects. Also, the welded joints possess higher strength with ER4043 filler. During the welding process, thermal cycles play an important role in evaluating welding strength, bead geometry, HAZ, residual stresses and thermal distortions [15, 16]. The temperature of any surface can be measured by contact and non-contact methods during welding. The measurement of temperature through contact and non-contact methods. The measurement of temperature through contact technique by using thermocouples mounted in the holes created on weld specimens [17, 18]. Deng et al. [19] used K-type thermocouples for measuring temperature in and outer surfaces of the SUS304 pipe line during GTAW process at a distance of 5 mm, 10 mm and 15 mm from the fusion line. However, the temperatures on the inside surface are slightly lower than the experimental measurements. Kumaresan et al. [20] used K-type thermocouples to measure thermal profiles for Aluminum weldment process. The thermal cycles were acquired during the process using a data acquisition system. The temperature measurement through thermocouple was robust and it was giving direct temperature values. But in this method, thermocouples are to be inserted at depth which is very difficult. The temperature measurement between two thermocouple positions may not provide accurate values. This can be considered a major drawback of thermocouples and this drawback can be easily covered through non-contact temperature measurement techniques. It is popularly as an in-situ thermal profile data logger that measures the surface temperature profiles by quantifying the wavelengths of emitted infrared energy [21]. Infra-Red Thermography (IRT) is a technique that detects the IR energy emitted from the surface and converts that energy to temperature, and shows a photograph of the temperature distribution. One of the first indications of faults in mechanical, electrical processes and R&D systems is an increase in temperature due to increased resistance or friction or many more reasons based on the different applications [22]. These increased temperatures can be detected and qualified using the latest thermal imaging cameras. When used as part of a predictive maintenance program or research & development program. Thermography can help to know the temperature based on infrared energy emitted from any object. The latest thermal cameras can operate in an alike way to a normal video camera. The item or component is scanned and an image of the temperature distribution is obtained [23]. These in-situ profile images can be analyzed using software (Forward-Looking Infra-Red (FLIR) tool plus) by plotting lines along transverse and longitudinal directions. Also, the point and area temperature of the objects can be analyzed [24]. Balram et al. [25] measured the peak temperatures, thermal heating and cooling cycles during welding of similar and dissimilar weldments of Monel 400 and AISI 316 using IRT at different time intervals and at different zones (fusion and HAZ) in order to validate the thermal fields predicted from ANSYS workbench. Higher temperature values were recorded on the steel side because of its lower thermal conductivity. Yelamasetti et al [26] recorded thermal cycles during welding Aluminum welded joints developed with TIG welding process. The emissivity ofthe weld surfaces of the given input to the IR camera in order to capture in-situ visuals. The results show that the peak temperatures were observed in the final pass of welding. Also, it was recorded that non-uniform distributions were recorded in pass-1.

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In this present work, dissimilar plates of AA6082 and AA7075 are joined by three different tool profiles, cylindrical, tapered and square. High strength H13 steel tool type was selected as a tool metal. The optimum process parameters, feed rate and tool rotational speed, were selected from research paper of own authors [27]. In-situ temperature profiles are recorded using IR thermography at different time intervals. Further, the welding strength and microstructural studies were evaluated by using tension test and optical microscope respectively. The effect of tool pin on temperature distribution, formation of grains and weld strength was studied. Materials and Experimentation The dissimilar metals, AA6082 and AA7075, of 150 × 80 × 6 mm were joined using FSW process with three different tool shapes. The tool shapes such as cylindrical, tapered and square were used for welding of dissimilar joints. The chemical composition of base metals is listed in Table 1. The process parameters used for FSW process are mentioned in Table 2. For three geometrical shapes, the parameters were kept constant for developing dissimilar joints. H13 tool material was used for welding dissimilar joints. The developed welded structures are shown in Fig. 1. Table 1. Standard composition (chemical) of AA 6082 and AA 7075 alloys Alloy Element Cu AA 6082 0.1 Wt. (%) AA 7075 1.2

Cr 0.25 0.18

Si 0.9 0.4

Fe 0.5

Ti 0.1 0.2

Zn 0.2 5.1

Mn 0.4 0.3

Mg others Al 0.8 --balance 2.1 0.15

Table 2. Process parameters used for FSW technique Sl. No.

Material

01 02 03

AA7075- AA6082 (6 mm thick)

Tool Rotational Tool feed Speed (rpm) (mm/min) 900 900 900

40 40 40

Type of tool

Pin height (mm)

Cylindrical Tapered Square

5.8

Fig. 1. Friction stir welded samples developed by (a) cylindrical, (b) Tapered and (c) square tools Infrared thermography: The images captured from IR camera were studied using FLIR Tool plus software. The data was captured at 50 fps giving a colour image of temperature distribution/profiles. This is stored as an image in an external hard disk for analyzing the profiles. These files serve as real-time monitoring of temperature profiles. The initial temperature of base metals is considered as 27oC. The emissivity value of base metals is entered into the camera in order to record the temperature distribution. From the literature, the emissivity of the AA 7075 and AA 6082 has been taken as ε = 0.38, ε = 0.30 respectively. The visual images of weld zone of three tools is shown in the Fig. 2.

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Fig. 2. IR Thermography images of FSW process with (a) cylindrical tool, (b) tapered tool and (c) square tool Characterization of weldments: The friction stir weldments were examined by conducting X-ray radiography test for any internal defects. In the current paper, CEREM 235-based X-Radiography machine was used for examining the welded samples for any internal defects. The characterization FSW weldments were done according to ASME SEC-VIII standards. The welded plates were exposed for a period of 120s at shorter wavelengths of X-rays. A 150 keV source voltage and 3.5 mA current were maintained for identifying the defects during the test. After Non-Destructive Testing (NDT) was completed, the weldments were sliced transversely into different welding coupons to study the microstructural and mechanical properties. Tensile tests were conducted for evaluating the weld strength and an optical microscopy technique was used to study the metallurgical aspects of the dissimilar weldments. Results and Discussions Thermal analysis: The peak temperatures and their distribution along transverse direction in each welding pass were analyzed by using FLIR tool plus software. During the welding process and after welding process, the images were analyzed and presented in Figs. 3-5. In cylindrical tool weldment, the temperature at different time intervals were recorded and analyzed data presented in Table 3. In this weldments, the maximum temperature at 90 s was recorded as 528 oC at weld zone. The maximum temperature of 351.4 oC was recorded at HAZ of AA7075 whereas the temperature of 315.8oC was recorded at HAZ of AA6082. The transient effect can be seen at weld vicinity and this could be achieved due to proper thermal cycles. In tapered tool weldment, the temperature at different time intervals were recorded and analyzed data presented in Table 4. In this weldments, the maximum temperature at 90 s was recorded as 551 oC at weld zone. The maximum temperature of 359.7 oC was recorded at HAZ of AA7075 whereas the temperature of 311.5 oC was recorded at HAZ of AA6082. The non-uniform thermal fields were observed at time interval of 60 s in this weldment due to the improper mixing or formation coarse grains at HAZ of both metals. In square tool weldment, the temperature at different time intervals were recorded and analyzed data presented in Table 5. The peak temperature of 576.6 oC was recorded at the weld zone. The maximum temperature of 342.9 oC was recorded at HAZ of AA7075 whereas the temperature of 316.4 oC was recorded at HAZ of AA6082. The improper thermal gradients and development of peak temperatures were observed to low when cylindrical shape tool was used. It is also evident that, uniform distribution of both the metals were observed from Fig. 1. After welding, the samples were cooled in normal air-condition and the IR image of cylindrical tool weldment is shown in Fig. 6. It is observed that the welded samples were got cooled after 25 minutes. Table 3. IR temperature values at different time intervals when cylindrical shape tool is used Time intervals 30 s 60 s 90 s

Temperature at HAZ of AA6082 303.2 °C 335.2 °C 315.8 °C

Temperature at Weld zone 498.5 °C 509.7 °C 528.1 °C

Temperature at HAZ of AA7075 336.8°C 352.1 °C 351.4 °C

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Table 4. IR temperature values at different time intervals when tapered shape tool is used Time intervals 30 s 60 s 90 s

Temperature at HAZ of AA6082 329.5 °C 299.2 °C 311.5 °C

Temperature at Weld zone 530.4 °C 532.9 °C 551.0 °C

Temperature at HAZ of AA7075 302.2 °C 321.8 °C 359.7 °C

Table 5. IR temperature values at different time intervals when square shape tool is used Time intervals 30 s 60 s 90 s

Temperature at HAZ of AA6082 304.9 °C 299.2 °C 316.4 °C

Temperature at Weld zone 502 °C 522.9 °C 576.6 °C

Temperature at HAZ of AA7075 314.8 °C 321.8 °C 342.9 °C

Fig. 3. 3-D temperature profiles during FSW of dissimilar weld joined by cylindrical tool at different time intervals (a) 30 s, (b) 60 s and (c) 90 s

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Fig. 4. 3-D temperature profiles during FSW of dissimilar weld joined by tapered tool at different time intervals (a) 30 s, (b) 60 s and (c) 90 s

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Fig. 5. 3-D temperature profiles during FSW of dissimilar weld joined by square tool at different time intervals (a) 30 s, (b) 60 s and (c) 90 s

Fig. 6. 3-D temperature profiles FSW weldments after cooling process

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Tensile strength: The tensile properties of welded structures evaluated by conducting tension test and the obtained results are listed in Table 6. The fractured samples under tension test are shown in Fig. 7. The higher tensile strength of 265 MPa was observed for dissimilar weldment developed with cylindrical shape with optimum parameters and proper mechanical mixing. Also, the higher yield strength of 166 MPa was observed when cylindrical tool was employed. The weld tensile strength of 179 MPa and 121 MPa was observed for weld specimen of developed with tapered and square tools respectively. The yield strength to weld strength ratio is observed as 0.62, 0.76 and 0.79 for cylindrical, tapered and square tool weldments respectively. The stirring of base metals and proper mechanical mixing was observed when cylindrical tool was mentioned. It is also evident that, the peak temperature recorded at weld zone is minimum in comparison of other tool profiles. The thermal gradients towards the base metals were observed to be uniform in cylindrical shape tool which could be attributed to formation of fine grains in the vicinity of weld zone. The tensile stress-strain graph three welded samples is shown in Fig. 8. It is observed from this figure, the higher % of elongation can be achieved when tapered shape tool was employed when compared with other two tool profiles. Table 6. Tensile properties of dissimilar joint AA6082 and AA7075 S. No

Base material & type of tool

Yield strength (MPa)

Ultimate tensile strength in (MPa)

% of Elongation

1. 2. 3.

AA7075- AA6082 (Cylindrical) AA7075- AA6082 (Tapered) AA7075- AA6082 (Square)

166 137 96

265 179 121

6.26 9.60 5.06

Fig. 7. Tensile fractured samples (a) cylindrical tool, (b) tapered cylindrical tool and (c) square tool

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Tensile stress (MPa)

900 rpm 40 mm/min.

AA6082-AA7075 Cylindrical Tool Tapered Tool Square Tool

250

147

200 150 100 50 0

0

2

4

6

8

10

12

14

Tensile strain (%)

Fig. 8. Tensile stress graph for three tools friction stir welded samples Microstructural studies: From the XRT results it is observed that all the samples were free from weld internal defects. The microstructures of three friction stir welded plates are taken using optical microscopy and the structures at various zone are shown in Figs. 9-11. From Fig. 9 shows the macro and microstructures of cylindrical tool weldment at various zones. It is observed that, the microstructure consists of fine and homogeneously dispersed AlZn2 precipitates with in the base material of AA7075, HAZ, and TMAZ and in nugget zone. Whereas, the microstructure of AA6082 shows AlMg2Si precipitates with in the base material and in HAZ and showing TMAZ, HAZ at equal magnifications. From Fig. 10 shows the macro and microstructures of tapered tool weldment at various zones. It is witnessed that, the onion rings were observed at nugget zone. The microstructures contains of equi-sized grains and homogeneously dispersed AlZn2 precipitates with in the HAZ of AA7075, HAZ, and TMAZ and in nugget zone. Whereas, the microstructure of AA6082 shows AlMg2Si precipitates with in the base material and in HAZ and showing TMAZ, HAZ at higher magnifications. From Fig. 11 shows the microstructures of square pin tool weldment at various zones. It is revealed that, the microstructure consists of in- homogeneous dispersed AlZn2 precipitates and slight unmixed zones with in the base material of AA7075 and in nugget zone. While, the microstructure of AA6082 shows AlMg2Si precipitates in the HAZ and at TMAZ. The unmixed zones were observed in both sides of base metals which could be reasoned due to development of high temperature values recorded in this weldment which could be attributed to failure at TMAZ under tension test.

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Fig. 9. Microstructures of cylindrical tool friction stir weldment of AA6082 and AA7075

Fig. 10. Microstructures of tapered tool friction stir weldment of AA6082 and AA7075

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Fig. 11. Microstructures of square tool friction stir weldment of AA6082 and AA7075 Conclusions The following conclusions are drawn from this research work which includes measurement of temperature, weld strength and microstructural studies; • The dissimilar weldments of AA6082 and AA7075 developed with three tools such as cylindrical, tapered and square shape are developed successfully without any macro/micro defects. • The proper thermal gradients in the base metals, lower temperature values were recorded in cylindrical tool weldment. The peak temperature of cylindrical tool was decreased by 9% and 4.3% when compared to the square and tapered tools respectively. • The weld strength of cylindrical tool weldment is increased by 54.4 % and 32.4% in comparison of square tool and tapered tools respectively. Also, the higher yield strength of 166 MPa is observed than the other two tool weldments. Due to uniform distribution of thermal gradients towards the both base metals, higher weld strength properties were observed. • The microstructural studies revealed that fine and homogeneously dispersed AlZn2 precipitates at HAZ of AA7075 and at nugget zone. Whereas, the microstructure of AA6082 shows AlMg2Si precipitates at HAZ of AA6082 and near to the TMAZ

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References [1] Gadallah N, Sabry I, Ghafaar MA (2020) A Summarized Review on Friction Stir Welding for Aluminum Alloys. Acad Res Community Publ 4:1 https://doi.org/10.21625/archive.v4i1.695 [2] Santhosh kumar S, Senthil Kumar KL, Kalil Rahiman M, Mathankumar P (2020) A review on friction stir welding of aluminium alloys and the effects on tool geometry. IOP Conf. Ser. Mater. Sci. Eng. 764. https://doi.org/10.1088/1757-899X/764/1/012009 [3] T. Bajpei, H. Chelladurai, M.Z. Ansari, Experimental investigation and numerical analyses of residual stresses and distortions in GMA welding of thin dissimilar AA5052-AA6061 plates, J. Manuf. Process. 25 (2017) 340–350. https://doi.org/10.1016/j.jmapro.2016.12.017 [4] Balram Yelamasetti, Venkat Ramana G, Sandeep Manikyam, Kuldeep K. Saxena, Multiresponse Taguchi grey relational analysis of mechanical properties and weld bead dimensions of dissimilar joint of AA6082 and AA7075, Adv. Mater. Process. Technol. (2021). https://doi.org/10.1080/2374068X.2021.1946340 [5] Ganesh, M.R.S., Reghunath, N., J.Levin, M. et al. Strontium in Al–Si–Mg Alloy: A Review. Met. Mater. Int. 28, 1–40 (2022). https://doi.org/10.1007/s12540-021-01054-y [6] Robert Kosturek, Janusz Mierzyński, Marcin Wachowski, Janusz Torzewski, Lucjan Śnieżek. The influence of tool traverse speed on the low cycle fatigue properties of AZ31 friction stir welded joints. Procedia Structural Integrity Elsevier B.V. 2021; 36:153-158; https://doi.org/10.1016/j.prostr.2022.01.017. [7] Tian DING, Hong-ge, YAN1, Ji-hua CHEN, Wei-jun XIA, Bin SU, Effect of welding speed on microstructure and mechanical properties of Al−Mg−Mn−Zr−Ti alloy sheet during friction stir welding. Transactions of Nonferrous Metals Society of China. 2021; 31:12; 3626-3642 https://doi.org/10.1016/S1003-6326(21)65753-9. [8] R.P. Verma, K.N. Pandey, Y. Sharma, Effect of ER4043 and ER5356 filler wire on mechanical properties and microstructure of dissimilar aluminium alloys, 5083-O and 6061-T6 joint, welded by the metal inert gas welding, Proc. Inst. Mech. Eng. Part B J. Eng. Manuf. 229 (2015) 1021–1028. https://doi.org/10.1177/0954405414535771. [9] G. Venkat Ramana, B. Yelamasetti, T. Vishnu Vardhan, Effect of FSW process parameters and tool profile on mechanical properties of AA 5082 and AA 6061 welds, Mater. Today Proc. (2021). https://doi.org/10.1016/j.matpr.2020.12.801 [10] G. G. Krishna, T. Mahender, S. Reddy, and R. S. U. Rao, “The effect of offset tools on aluminum AA6351 alloy friction stir welds,” in Materials Today: Proceedings, 2021, vol. 46, pp. 320–324, doi: 10.1016/j.matpr.2020.08.180. [11] Ugender, Influence of tool pin profile and rotational speed on the formation of friction stir welding zone in AZ31 magnesium alloy, Journal of Magnesium and alloys. 2018; 206-213. https://doi.org/10.1016/j.jma.2018.05.001. [12] Kulwant Singh, Gurbhinder Singh, Harmeet Singh, Investigation of microstructure and mechanical properties of friction stir welded AZ61 magnesium alloy joint, Journal of Magnesium and Alloys 6 (2018) 292–298, https://doi.org/10.1016/j.jma.2018.05.004. [13] Sara Bocchi, Marina Cabrini, Gianluc, D’Urso, the influence of process parameters on mechanical properties and corrosion behavior of friction stir welded aluminum joints, Journal of Manufacturing Processes, 35 (2018) 1–151526-6125 The Society of Manufacturing Engineers. Published by Elsevier Ltd. https://doi.org/10.1016/j.jmapro.2018.07.012 [14] Balram Yelamasetti, Venkat ramana G, Vishnu vardhan T, Weldability and mechanical properties of AA5052 and AA7075 dissimilar joints developed by GTAW process, Mater. Today Proc. (2021). https://doi.org/10.1016/j.matpr.2021.04.446.

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[15] Balram Y, Sridhar Babu B, Vishnu Vardhan T, et al. Residual stress analysis of dissimilar tungsten inert gas weldments of AISI 304 and Monel 400 by numerical simulation and experimentation. In: Materials Today: proceedings. Elsevier Ltd, 2019. 478–483. [16] Balram Yelamasetti, Venkat Ramana G, Sandeep Manikyam & Vishnu Vardhan T (2021): Thermal field and residual stress analyses of similar and dissimilar weldments joined by constant and pulsed current TIG welding techniques, Advances in Materials and Processing Technologies, https://doi.org/10.1080/2374068X.2021.1959114. [17] Vemanaboina H, Edison G, Akella S, et al. Thermal analysis simulation for laser butt welding of inconel625 Using FEA. Int. J. Eng. Technol. 2018;7(4.10):85 [18] Vemanaboina H, Edison G, Akella S. Validation of residual stress distributions in multi-pass dissimilar joints for GTAW process. J. Eng. Sci. Technol. 2019; 14:2964–2978. [19] Deng, D. and Murakawa, H. (2008), ‘Finite element analysis of temperature field, microstructure and residual stress in multi-pass butt-welded 2.25Cr-1Mo steel pipes’, Computational Materials Science, 43(4), 681–695. [20] Kumaresan, D., Asraff, A. K. and Muthukumar, R. (2011), ‘Numerical Investigation on Heat Transfer and Residual Stress in a Butt Welded Plate’, Journal of Pressure Vessel Technology, 133(4), 041-206. [21] Yelamsetti B, Rajyalakshmi G, Thermal stress analysis of similar and dissimilar welded joints, U.P.B. Sci. Bull., Ser. D. 80 (2018). [22] Y. Balram, T. Vishu Vardhan, B. Sridhar Babu et al., Thermal stress analysis of AISI 316 stainless steels weldments in TIG and pulse TIG welding processes, Materials Today: Proceedings, https://doi.org/10.1016/j.matpr.2019.06.695 [23] K.C. Ganesh, M. Vasudevan, K.R. Balasubramanian, N. Chandrasekhar, P. Vasantharaja, Thermo- mechanical analysis of TIG welding of AISI 316LN stainless steel, Mater. Manuf. Process. 29 (2014) 903– 909, https://doi.org/10.1080/10426914.2013.872266. [24] M. Vasudevan, M. N. Chandrasekhar, M. V Maduraimuthu, A. K. Bhaduri, and B. Raj, Realtime monitoring of weld pool during GTAW using infra-red thermography and analysis of infra-red thermal images, Weld. World, vol. 55, no. 7–8, pp. 83–89, 2011. [25] Balram Y, Rajyalakshmi G. Thermal fields and residual stresses analysis in TIG weldments of SS 316 and Monel 400 by numerical simulation and experimentation. Mater Res Express. 2019; 6:0865e2. [26] Balram Yelamasetti, Deepak Kumar, Kuldeep K Saxena & Rajyalakshmi G (2021): Experimental investigation on temperature profiles and residual stresses in GTAW dissimilar weldments of AA5052 and AA7075, Advances in Materials and Processing Technologies. https://doi.org/10.1080/2374068X.2021.1927641. [27] Ramakrishna M V a & Srinivas K (2021): Grey relational analysis of frictionstir welding parameters for the development of dissimilar joints between AA6082 and AA7075, Advances in Materials and Processing Technologies, https://doi.org/10.1080/2374068X.2021.1959112

Key Engineering Materials ISSN: 1662-9795, Vol. 934, pp 153-160 doi:10.4028/p-51a271 © 2022 Trans Tech Publications Ltd, Switzerland

Submitted: 2022-05-16 Revised: 2022-06-10 Accepted: 2022-06-13 Online: 2022-11-28

Mechanical and Metallurgical Characteristics of Rotary Friction Welded Low Carbon Steel Plate/Rod Joints T. Dhamotharakannan1,a*, P. Sivaraj2,b, M. Seeman3,c, V. Balasubramanian4,d Research Scholar, 2,3Associate Professor, 4Professor

1

Centre for Material Joining and Research (CEMAJOR), Dept of Manufacturing Engineering,

1234

Annamalai University, Tamil Nadu, India. [email protected], [email protected], [email protected], d [email protected]

a

Keywords: rotary Friction Welding (RFW), low carbon steels, unsymmetrical joint, microstructure, tensile properties.

Abstract. This work aims to study the mechanical properties and microstructural characteristics of rotary friction welding of the unsymmetrical (plate/rod) joints of AISI 1018 low carbon steel plate and AISI 1020 low carbon steel rod. The mechanical properties (tensile properties and hardness) were studied. The fractured surface of the tensile specimen was examined by Scanning Electron Microscope (SEM). The tesnile properties (strength and elongation) are higher than the AISI 1018 plate but slightly lower than the AISI 1020 rod due to coarse ferrite grains in the HAZ region of the AISI 1018 plate. The hardness varied from the fully deformed zone (FDZ) to the base metal. The average value of the ultimate tensile strength of the friction welded joint is about 452 MPa. The average value of hardness at fully deformed zone is about 252 Hv, which 32% higher than the base metal. Introduction Shafts, lightly stressed gears, chains, pins and other structural parts were made of AISI 1020 low carbon steel [1], gears, non-critical components of tool, dowels, pins, tool holders, pinions, ratchets, machine parts, chain pins are the important applications of AISI 1018 low carbon steel [2]. Rotary friction welding (RFW) is a solid-state joining process which works by rotating one workpiece relative to another while under a compressive axial force. The friction between the surfaces produces heat, causing the interface material to plasticise. The compressive force displaces the plasticised material from the interface, expelling the original surface oxide layer and other contaminants and promoting metallurgical and/or surface interlocking joining mechanisms [3]. Many researchers [4, 5] have done work on RFW of similar and dissimilar materials. RFW was used [3] to study the influence of rotational speed on the metallurgical and tensile characteristics of dissimilar metals such as AISI 304 and AISI 4140. In recent study, authors [4] used rotary friction welding to study the mechanical characteristics of mild steel with stainless steel joints. The author found that the variation in hardness from the fully deformed zone (FDZ) to the base metal is due to the recrystallization of grains. Another document [5] published on effects of process parameters on microstructural characteristics and observed that axial pressure has the greatest impact. The authors stated that the applied higher forging pressure in the weld region produces fine grains, equiaxed grains, and recrystallized grain structures, which lead to superior tensile strength and microhardness. Based on published data [6], the tensile strength of the joints is influenced by the duration of friction. The authors noticed a decrease in tensile strength when friction time increased from a minimum to a maximum and then rapidly increased. The upset force has no beneficial influence on joint tensile characteristics. Similarly, the upset force has little influence on joint tensile characteristics. The upset force has opposite effect on tensile strength. The study [7] optimised the process parameters of RFW of medium carbon steel (AISI 1040) and stainless steel (AISI 304) dissimilar joints using a response surface approach. The document [1] explains the mechanical and microstructural characteristics of friction welding of AISI 4140 low alloy steel and

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ASTM A106 B carbon steel pipe joints. The authors reported that the higher hardness was produced in the interface of weld due to the bainite and fine pearlite at grain boundaries From the literature review, it is understood that many researchers have studied the joining of the rod with the rod using the rotary friction welding process. There is no study on the joining of unsymmetrical components (Plate/rod) using rotary friction welding. Therefore, an attempt was made to join the AISI 1018 carbon steel plate with the AISI 1020 carbon steel rod. The main objective of this study is to investigate the mechanical and microstructural characteristics of RFW of AISI 1018 low carbon steel plate and AISI 1020 low carbon steel rod joints. Experimental Procedure Materials and fabrication process Table 1 shows the chemical composition of the low carbon steel materials (plate and rod). Table 1. Chemical composition (wt%) of low carbon steels Material

Cr

Mn

C

Mo

Si

S

P

Ni

Fe

AISI 1020 (Rod)

0.172

0.580

0.285

0.034

0.196

0.037

0.045

0.108

Bal

AISI1018 (Plate)

0.129

0.617

0.172

0.024

0.194

0.021

0.048

0.124

Bal

Initially, many trials were conducted to find good RFW parameters for joining the AISI 1018 low carbon steel plate and AISI 1020 low carbon steel rod. To produce the defect-free joint, process parameters such as 1200 rpm rotational speed, 26 MPa forging pressure, 26 MPa friction pressure, and 7 secs of forging time and friction time were optimized. AISI 1018 low carbon steel plate with dimensions of 15×32×32mm and AISI 1020 low carbon steel rod with a 12×120 mm were used. Figure 1 a,b) displays the photographs of the specimens before welding. The specimens were welded using a numerically controlled RFW machine. From various trials, the optimised process parameters were used to produce the sound joint. Figure 1c) shows the photographs of the joints after welding.

Fig. 1.a, b) Photographs of AISI 1018 Plate and AISI 1020 rod specimens and c) friction welded unsymmetrical steel joints

Fig.2. a) Dimensions of miniature tensile specimen and b) photograph of friction welded unsymmetrical joint after tensile specimen’s extraction

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Mechanical properties evaluation Vickers microhardness tester was used to measure the friction welding of unsymmetrical joints of low carbon steel. The microhardness sample was extracted from the cross section of unsymmetrical joints of low carbon steel. The specimen was mirror polished to get the fine surface. The hardness measurement was done from weld interface to base metal (plate and rod side) with a 500 gms load and 10 secs dwell time was maintained for each indentation. The fabricated unsymmetrical joints were tested by extracting miniature tensile specimens from the welded specimens as per standard ASTM E8–M. The dimensions of the tensile specimen are shown in Figure 2 a). The specimens are tested on a servo-controlled Tinius Olsen (H50KL) Universal Testing Machine. The details of specimen extraction from a welded specimen are shown in Figure 2 b). Three tensile specimens were tested, and their average values are reported for analysis. Macrostructure and Microstructural characterization The mirror polished cross section of RFW of unsymmetrical low carbon steel joints was subjected to metallurgical examination. To evaluate the different weld zones and their structures, the macrostructure and microstructure of welded samples were analysed using stereo zoom microscope and optical microscope at various magnifications. A Scanning Electron Microscope (SEM) was used to examine the fracture surface of a tensile specimen. Results and Discussion Tensile properties Figure 3 a) shows the photographs of tensile samples (before and after testing). The Tensile properties (tensile strength and yield strength) of the unsymmetrical joints are shown in Figure 3 b). Additionally, table 2 presents the average values of the base materials and weld specimen. The unsymmetrical joint of low carbon steel showed an ultimate tensile strength (UTS) of 452 MPa and yield strength (YS) of 362 Mpa. The observed tensile properties are closer to the base material properties, which are confirmed by the 94% joint efficiency of the unsymmetrical joint. The samples failed from heat affected zone (HAZ) to the base metal (plate side) during tensile testing of unsymmetrical low-carbon steel joints. The lower strength of the RFW unsymmetrical joint is mainly due to the lower hardness of the HAZ of the plate, which caused the failure in HAZ of the specimen. The published data [8] explores the mechanical and microstructural properties of friction welded aluminium alloy joints. The authors observed higher UTS, YS, and elongation in friction welded aluminium joints due to fine grain α-Al matrix in the weld metal. RFW was used to join the carbon steel [9]. The authors observed better tensile properties in the welded joint than in the base metal due to recrystallization of fine grains and a hard bainitic structure. Microhardness survey The microhardness was evaluated at the cross section of the RFW unsymmetrical joint. Figure 4 a) shows the hardness distribution of different zones of an unsymmetrical low carbon steel friction welded joint. The RFW unsymmetrical joint showed higher hardness at the weld interface. The hardness of the weld interface is about 252 Hv, which is higher than the other regions due to more deformation and finer grain bainitic structure. The hardness decreased gradually from the weld interface to other weld zones (partially deformed zone (PDZ), HAZ, and base metal). The average microhardness values of the PDZ range between 251 HV and 219 HV for the rod and from 251 Hv to 203 Hv for the plate. This indicates that the amount of deformation of the joint that occurs in the rod is greater than in the plate [10]. In the HAZ of the rod, the hardness value is about 218 Hv to 180 Hv. On the other hand, the HAZ of the plate shows 202 Hv to 160 Hv. Here also, there is higher hardness on the rod side. It is mainly due to the higher deformation on the rod side when compared to the plate, and the size of the grains on the rod side is smaller than the plate in RFW unsymmetrical low carbon steel joints. The hardness of the weld interface is 32% greater than that

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of the base metal. The higher hardness at the weld interface is mainly due to the finer bainite. The hardness decreased from the weld zone to the HAZ and base metal due to microstructural variations. The grain size was increased from the FDZ to the HAZ and base material, which led to a reduction in hardness on both sides of the unsymmetrical low carbon steel RFW joints. Compared to the rod side, the plate side resulted in coarser grain, which led to failure in the HAZ to base metal region during tensile testing. The tensile properties (discussed in the previous section) of RFW unsymmetrical joints are consistent with hardness results. The strength and elongation of the RFW unsymmetrical joint showed higher values than the AISI 1018 plate but slightly lower values than the AISI 1020 rod due to coarse ferrite grains in the HAZ region of the AISI 1018 plate. Similar behaviour was observed in recently published work [11]. The authors observed higher hardness in the weld interface than in other zones due to the differences in microstructural features between the weld interface and base metal. a)

b)

Fig.3. a) Photographs of tensile specimens and b) tensile properties Table 2. Tensile properties Specimen

Tensile strength (MPa)

Yield strength (MPa)

Elongation (%)

Joint efficiency (%)

Welded specimen

452

362

27.08

94

AISI 1020 (rod)

478

383

29.50

----

AISI 1018 (plate)

440

341

26.00

----

Macrostructure and Microstructural Characterization The macrostructure was examined from the cross section of the RFW of the unsymmetrical joint of low carbon steel. Figure 4 b) shows the macrostructure of an unsymmetrical joint of low carbon steel with different zones. A higher flash was observed on the rod side of the joint. This is mainly due to the smaller surface area of the contact area when compared to the plate side, which led to higher flash. The macrostructure is free from cracks and other visual defects. The microstructural analysis was done from the cross section (Figure 4 b)) of RFW of the unsymmetrical joint of low carbon steel. The weld metal is formed by four zones: FDZ, which is the weld centre region; PDZ, which is nearer to the weld centre region; HAZ, which is closer to the PDZ; and the last one is the base metal (BM). The primary microstructures of base materials such as AISI 1020 rod and AISI 1018 plate are depicted in Figures 5 a) and b). The BM primarily composed of equiaxed ferrite grains with an average grain size of 22 µm and a pearlite structure. Figure 5 c) shows the weld interface, or FDZ, and the PDZ of the plate and rod.

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b)

Fig.4.a) Microhardness distribution of friction welded joint and b) Macrostructure a)

b)

d)

e)

g)

f)

h)

Fig.5. Optical Micrographs of RFW unsymmetrical joint, a) Base metal – AISI 1018 (plate), b) Base metal – AISI 1020 (rod), c) Weld region, d) FDZ or weld zone, e) PDZ AISI 1018 (plate), f) PDZ of AISI 1020 (rod), g) HAZ of AISI 1018 (plate) and h) HAZ of AISI 1020 (rod). Figure 5 d) shows the microstructure of the FDZ with different microstructural features when compared to other zones. This zone mostly consists of fine bainitic structures with elongated ferrite. This recrystallization of grains is due to chemical elements and temperature that occur during cooling [12]. The grains become 15 µm finer as a result of dynamic recrystallization at the weld

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interface. There is also a temperature distribution over the weld interface region to the base metal. When frictional heat at the interface is dissipated via the parent material, it transfers temperature, resulting in distinct microstructures in various zones of the material [13]. The microstructure of the PDZs on both sides is shown in Figure 5 e) and f), which shows partial grain refinement due to less heat input and pressure over the grains. The microstructure of the HAZ region of both the plate and rod is shown in Figures 9 g) and h). The HAZ of both the rod and plate is composed of ferrite and pearlite. The grain size of the rod side is finer compared to the plate side. The interstitial free steels were joined by RFW [14].The authors observed microstructural variations from the weld metal to the base metal. The FDZ is characterised by the bainite structure, and the grain size and structure were changed from FDZ to PDZ and HAZ. The HAZ was composed of a coarser microstructure than the PDZ. The authors reported that the hardness decreased from FDZ to HAZ and base metal due above microstructural changes. Fractography The fracture surface of both the plate and rod was examined by a scanning electron microscope. Figure 6 shows the fracture surfaces of the base materials and the friction welded unsymmetrical low carbon steel joint were characterized by ductile areas with dimples and micro voids, indicating that the failure mode of the unsymmetrical low carbon steel RFW joint is the ductile mode. The fracture surface of the unsymmetrical joint showed fine and deeper dimples with a more ductile region. This suggests that the joint efficiency is almost nearer to the base material. Similar behaviour was observed in reference [15]. The authors observed higher strength and elongation in the friction welded aluminium alloy similar joint than the parent metal due to the finer and deeper dimples.

Fig. 6. SEM Fractographs of RFW unsymmetrical joint, a) AISI-1018 plate,b) AISI-1020 rod,c) RFW Joint Conclusions The study attempts to understand the mechanical and metallurgical characteristics of RFW of unsymmetrical joints of low carbon steel. The RFW technique is found to be very much suitable for producing low-carbon steel unsymmetrical components. There are no major defects in the unsymmetrical low carbon steel joint (confirmed from macrostructure) fabricated by the RFW process. The tesnile properties (strength and elongation) are higher than the AISI 1018 plate but slightly lower than the AISI 1020 rod due to coarse ferrite grains in the HAZ region of the AISI 1018 plate. The hardness slightly decreased from the weld interface to the base metal. The weld interface of the low-carbon steel unsymmetrical joint fabricated by RFW has a higher hardness than the base metal. This is mainly due to the hard bainite and grain refinement in the weld interface.

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Acknowledgements The first author express his gratitude to Centre for Material Joining and Research (CEMAJOR), Department of Manufacturing Engineering, Annamalai University Annamalai Nagar, India, for their technical assistance. References [1]

A. Seshu Kumar, Sk. Abdul Khadeer, V. Rajinikanth, S. Pahari, B. Ravi Kumar, Evaluation of bond interface characteristics of rotary friction welded carbon steel to low alloy steel pipe joints, Mater. Sci. Eng. A. 824 (2021) 1-14.

[2]

B. Prasanna Nagasai, S. Malarvizhi and V. Balasubramanian, Effect of welding processes on mechanical and metallurgical characteristics of carbon steel cylindrical components made by wire arc additive manufacturing (WAAM) technique, CIRP. J. Manu. Sci. Tech. 36 (2022) 100–116.

[3]

G. Subhash Chander, G. Madhusudhan Reddy, A. Venugopal Rao. Influence of rotational speed on microstructure and mechanical properties of dissimilar metal AISI 304-AISI 4140 continuous drive friction welds, J. iron. Steel Res. Int. 19(2012) 64-73.

[4]

D. Anantha padmanaban, V. Seshagiri Rao, Nikhil Abraham, K. Prasad Rao, A study of mechanical properties of friction welded mild steel to stainless steel joints, Mater. Des. 30 (2009) 2642–2646.

[5]

H.T. Nu, N.H. Loc, P. Luu, Influence of the rotary friction welding parameters on the microhardness and joint strength of Ti6Al4V alloys, J. Eng. Manu. 235 (2020) 795-805.

[6]

R. Winiczenko, Effect of friction welding parameters on the tensile strength and microstructural properties of dissimilar AISI 1020-ASTM A536 joints, Inter. J. Adv. Manu. Tech.84 (2015) 1-16.

[7]

R Paventhan, P.R. Lakshmi narayanan, V. Balasubramanian, Optimization of Friction Welding Process Parameters for Joining Carbon Steel and Stainless Steel, J. Iron. Steel. Res. Int. 19 (2012) 66-71.

[8]

P. Hariprasath, P. Sivaraj, V. Balasubramanian, A Critical Assessment on Rotary Friction Welded High Strength Armor Grade Aluminum Alloy Joints, Phy. Meta. Metall. 122 (2021) 1401–1408.

[9]

J. Alex Anandaraj, S. Rajakumar, V. Balasubramanian, S Kavitha, Influence of process parameters on hot tensile behavior of rotary friction welded In 718/AISI 410 dissimilar joints, CIRP. J. Manu. Sci. Tech. 35 (2021) 830-838.

[10] H.R. Lashgari, S. Li, C. Kong, M. Asnavandi, Sh. Zangeneh, Rotary friction welding of additively manufactured 17-4PH stainless steel, J. Manu. Pro. 64 (2021) 1517–1528. [11] R. Paventhan, P.R. Lakshmi narayanan, V. Balasubramanian, Optimization of Friction Welding Process Parameters for Joining Carbon Steel and Stainless Steel, J. Iron. Steel. Res. Inter. 19 (2012) 66-71 [12] Y.S. Kong, M. Cheppu, Y.W. Park, Effect of heating time on thermo mechanical behaviour of friction-welded A105 bar and A312 pipe joints, Trans. Ind. Inst. Met. 73 (2020) 1433-1438.

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[13] GK. Padhy, CS. Wu, S. Gao, Friction stir based welding and processing technologies-process, parameters, microstructures and applications: a review, J. Mater. Sci. Tech. 34 (2018) 1-38. [14] E. Bayraktar ,D. Kaplan , L. Devillers, J.P. Chevalie. Grain growth mechanism during the welding of interstitial free (IF) steels, Journal of Materials Processing Technology 189 (2007) 114–125. [15] P. Sivaraj, P. Hariprasath, C. Rajarajan, and V. Balasubramanian, Analysis of grain refining and subsequent coarsening along on adjacent zone of friction stir welded armor grade aluminum alloy joints, Mater. Res. Express. 6(2019) 1-9.

CHAPTER 4: Building Materials

Key Engineering Materials ISSN: 1662-9795, Vol. 934, pp 163-170 doi:10.4028/p-385381 © 2022 Trans Tech Publications Ltd, Switzerland

Submitted: 2022-08-02 Accepted: 2022-09-06 Online: 2022-11-28

Pervious Concrete Utilized as a Base Slab of Pavement Edge Drains Jul Endawati1,a*, Aditia Febriansya2,b Politeknik Negeri Bandung, Indonesia

1,2 a*

[email protected], [email protected]

Keywords: Pervious, concrete, base slab, edge drains, clay soil.

Abstract. The drainage base plate is generally made of normal concrete. The application of pervious concrete as base-slab of pavement edge drain will then be possible to infiltrate storm water runoff on the pavement surface that is not infiltrated through the pavement surface. The edge drain does not only accommodate storm water flow, but also transmits water to the ground so that it can restore the ground water level or can be reused through a reservoir in the ground underneath. The effectiveness of the application of pervious concrete as a drainage base plate on silt-clay soils was conducted experimentally. The thickness of the porous concrete slab in the modeling box is 7 cm, laid above 20 cm height of silty clay soil layer. Reference soil conditions as well as contours and cross-sectional characteristics of the edge drain are taken based on the condition of one of the existing drainages in West Bandung. Based on the results of numerical simulations using Geostudio SEEP/W, it can be seen that there is an increase in pore water pressure in the area of the pervious concrete slab and the surrounding soil. The pore water pressure at the bottom of the plate is in the range of 18.64 kPa. Overall, the pore water pressure in the first modeling tends to be higher than in the second modeling. On the third day of testing, the difference in pore water pressure between the two modeling conditions is 27.08%. Introduction Pervious concrete can be used for a number of applications, as pavement in light traffic areas, parking lots. sidewalks, tennis courts, sub-bases for conventional concrete pavements, slope stabilization, well linings, hydraulic structures, pool decks, curbside drains, and noise barriers [1]. On the other hand, the watershed plan should include measures to control the output volume avoiding the displacement of downstream impacts, in accordance with the modern concept of urban rainwater drainage systems. Drainage control can be regulated through catchment areas and infiltration ditches, permeable pavements, and drainage systems. Pervious concrete has a network of interconnected cavities that allow exfiltration of water to the sub-base below [2]. In the absence or minimum of fine aggregate in the pervious concrete mixture, the cement paste binds the coarse aggregate by leaving an interconnected air void of 15-25%, so that it becomes highly permeable Within the development of its application, various research activities have been carried out in balancing the infiltration ability and acceptable compressive strength [3-5]. The infiltration rate testing method on pervious concrete, mostly is carried out based on ASTM 1701 [6], ASTM 1781 [7], ISO 17785-1 [8] and the falling head method [2,9,10]. The recommended permeability coefficient range is in the range of 0.11 cm/s [11] to 1.75 cm/s [9,12,13]. ACI [14] set this parameter from 0.14 to 1.22 cm/s. In several publications related to the infiltration capability of pervious concrete, the experimental results are only related to the characteristics of individual infiltration capabilities, for example: the effect of compositional parameters on the void content and permeability of permeable concrete [15]. The use of pervious concrete as a pavement allows rainwater to seep directly into the ground, reducing runoff into the city's drainage system. Its application also contributes to the maintenance of underground aquifers as well as the reuse of stored rainwater. In general, the drainage base is made of conventional concrete, suchlike, that the drainage runoff only relies on the channel discharge capacity and flow velocity in the horizontal direction. The application of pervious concrete at the bottom of the channel is expected to increase the runoff flow, taking into account that the water flow in the drainage is not only in the horizontal direction but also in the vertical direction. However, the use of pervious

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concrete must have such a system, which is in accordance with the conditions of the soil beneath it [16]. Seepage through the soils may affect the stability of geotechnical structures as well. Seepage in soils is established by Darcy’s Law according to which flow through a soil is laminar while flow velocity depends on the hydraulic gradient [17]. The characteristics and conditions of the groundwater table will also affect the performance of pervious concrete in its application as drainage base slab. Seepage through soil can affect geotechnical structures as well. The objective of this research is to measure the effectiveness of pervious concrete layer in the drainage base in transmitting storm water runoff but still retaining the soil condition underneath. Materials and Methods Mix Proportion. The mixture consists of 70% coarse aggregate and 30% binder. It’s a no-fines aggregate concrete mixture which used a 100% local coarse aggregate. Proportion of the binder consists of 63% Portland cement (PCC), 25% fly ash (FA) and 12% silica fume (SF) used based on the previous research [18]. The properties of fresh pervious concrete were improved by using a chemical admixture PV-100. Modeling. The study was carried out in a modeling box consists of a pervious concrete slab measuring of 800 mm x 200 mm x 70 mm and a silty-clay layer which equipped by perforated base of 5 mm diameter and 5 cm apart. A 20 cm height soil layer having a 2% of slope was prepared. Soil was obtained from the reference location and was prepared by adding water until the water content matches the lost water content. It was prepared in the modeling box after thoroughly mixed and compacted. The density of the compacted soil was carried out every 10 cm at 3 points. Fig. 1 shows the pervious concrete and modeling box used in this research.

Fig. 1. Pervious concrete and modeling box used in the experiment. Hydrological analysis carried out using the last ten years rainfall data of Cemara rain station and Husein Sastranegara airbase rain station. The consistency test on the rainfall data was carried out using the double mass curve method, while the distribution characteristics were tested by annual method maximum series. Frequency analysis principal Analysis of the parameters obtained by using: (1) Gumbel; (2) Normal; (3) Log Normal and (4) Log Pearson III distribution to determine type of the frequency distribution [19]. The correctness of the rainfall distribution data is tested by the compatibility test using the Chi-Square method and the Kolmogorov-Smirnov method [20]. Resume of the frequency distribution that is most likely to be used is given in Table 1.

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Table 1. Determination of the type of distribution. Type of distribution Gumbel Normal Log Normal Log Person III

Provision Cs Ck Cs Ck Cs = Cv³+3Cv Ck = Cv⁸+6Cv⁶+15Cv⁴+16Cv²+3

≈ 1,1396 ≈ 5,4002 ≈0 ≈3 0.147 3.038

In addition to the above provisions (Cs=0)

Result -0.063 2.242 -0.063 2.242 -0.204 2.266 -0.204 2.266

Description inadequate inadequate inadequate adequate

The rainfall intensity used refers to the Jakarta Intensity Duration Frequency (IDF) curve by Van Breen. Result of the translation or shift of the Jakarta IDF based on 2 years, 5 years, and 10 years showed that the value of the rain intensity to be used is of the 10-year return period. The compatibility test calculated using Talbot, Sherman, and Ishiguro equation as shown in Table 2 shows that the Talbot equation has the smallest average deviation of 0,06. The rain intensity equation is then determined using the Talbot equation. Table 2. The compatibility test based on Talbot, Sherman, and Ishiguro equation. No 1 2 3 4 5 6 7

Duration I-location 5 150.39 10 134.39 20 119.39 40 92.39 60 77.39 120 47.39 240 26.39 Total Average

Talbot 152.18 138.77 117.98 90.78 73.77 47.23 27.46

Deviaation Isherman Deviation Iishiguro Deviation 1.78 188.40 38.00 189.61 39.22 4.38 139.28 4.89 143.50 9.11 -1.41 102.97 -16.42 106.78 -12.62 -1.61 76.13 -16.27 78.40 -13.99 -3.62 63.80 -13.59 65.12 -12.27 -0.16 47.17 -0.23 47.11 -0.28 1.07 34.87 8.48 33.87 7.47 0.42 4.86 16.64 0.06 0.69 2.38

Results and Discussions Concrete Materials Characteristics. The specific gravity of the binder materials was tested by using Le Chatelier flask, which resulted in 3.042 gr/cm3, 2.15 gr/cm3 and 2.13 gr/cm3 respectively for Portland PCC Cement, fly ash and silica fume. The coarse aggregate was prepared in a certain narrow gradation within the range of 4.75mm and 12.5 mm, having a bulk specific gravity (dry) of 2.59 gr/cm3, water absorption of 3.08% and dry density of 1545 kg/m3. In Situ Soil Parameters. The existing soil is clay silt with the original soil permeability coefficient of 1.83 x10-7 cm/sec, the porosity is 54.692% and the pore number is 1.207. The depth of the groundwater table in the reference location is 3.93 m. Designed Discharge. The drainage selected as a reference in the observation has an area of land cover and building layout served of 39999 m2. The designed discharge calculation is summarized in in Table 3. Table 3. Summarized of the design discharge calculation. Catchment area (A) Surface runoff coefficient (C) Concentration time (tc) Rain intensity (I) Runoff discharge

39999 m² 0,44 4,62 min mm/hr 0,761 m³/s

Determination of Flow. Flood thickness used in modelling was designed for an edge drain of 20 cm width, with a 2,1 cm water level of 0,003313 m/s flow and of 1,4 x 10-5 m3/s discharge in scale. The infiltration tests. The infiltration tests were carried out in 3 consecutive days, 2 hours per day. Water infiltrated through layers of modeling materials was measured in 10-minute interval. The test

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conducted in two modeling-conditions, consisted of :(1) pervious concrete slab-soil, (2) pervious concrete slab-non woven geotextile-soil. Layers average infiltration capacity can be seen in Fig.1.

Fig. 2. Layers average infiltration capacity. The average infiltration capacity as presented in Fig. 2 shows a similar pattern between the two conditions, decreasing in the second cycle test and started to increase in the third cycle test. The overall average infiltration capacity shows a smaller result in the second modeling condition with a 10,38% difference of average infiltration capacity between the first and the second condition. Data gathered from the infiltration test was used in numerical analysis and simulation using Geostudio SEEP/W software. Cumulative Water Volume. Results of numerical analysis and simulation show that the amount of infiltration volume of the first modeling condition is 0.3386 m3 . Fig. 3 shows the infiltration volume of the modeling curve. The increase in the volume of water absorbed is of 0.1137 m3, 0.1132 m3, and 0.1117 m3 for each test cycle of the first modeling condition. Meanwhile, the second modeling condition showed that the amount of infiltration volume after three consecutive cycle test is 0.2938 m3 as presented in the infiltration volume curve in Fig. 4.

Fig. 3. Cumulative water volume of the 1st modeling condition.

Fig. 4. Cumulative water volume of the 2nd modeling condition.

It can be seen from the graph that the water continued to infiltrate and filled the pore spaces in the soil. Infiltrated water along with the provision of water flow in the drainage caused the volume of the water absorbed to increase. Infiltration Discharge. According to the results of numerical analysis and modeling by Geoseep Studio, a slight fluctuation of infiltration disharge changes was occured, but end up in almost the same discharge at the last of each cycle test for both modeling condition. The first ten minutes of water discharge of the first modeling condition indicated the same amount in every cicle of the simulation, as shown in Fig. 5. On the other hand, water discharge of the second condition modeling showed a slight fluctuation as can be seen in Fig. 6, although the condition tend to be stable at the end of each cycle test. The first ten-minutes discharge of the third cycle is 12% lower compares to the first condition of modeling, Overall, the total water discharge of the second condition of modeling is lower at the third cycle test compares to the first modeling condition.

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Fig. 5. Water discharge of the 1st modeling condition.

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Fig. 6. Water discharge of the 2nd modeling condition.

Total Water Head. The total water head indicates the presence of water pressure caused by water that infiltrated and can be defined as the stress represented in units of length at a certain elevation. Elevation represents the effect of gravity. Water in the pore media will flow when there is an energy imbalance. Water flows from a place of high energy level to a place of lower energy. Groundwater flow in the actual state does not change, the flow is influenced by the principles of hydraulics that have passed through the aquifer which in general as a flow medium, so that Darcy's law can be applied.

(a)

(b)

Fig. 7. Total water head (a) of the first modeling condition; (b) of the second modeling condition, after three cycle infiltration test. Fig. 7 shows the condition of the water total head surrounding the drain and the soil underneath the pervious concrete slab which changes along with the increment of the water infiltrated. These changes occurred due to the water pressure increment at the bottom of the drain that compressed a lower pressure part of the soil underneath. The red color in the numerical simulation results shows the part of the soil that experiences the greatest stress due to the influence of absorbed water, while the dark blue color shows the part of the soil that experiences the least stress. The blue dotted line shows the boundary of the area that has become saturated. Pore Water Pressure. During dry periods, the pore water pressures become more negative. The pore water pressure 'acts in the water and in the solid in every direction, produce both normal and tangential forces at particle contacts. The effectiveness of stress will be influenced by the presence of bulk water inside soil pores. The pore-water pressure condition changes due to the infiltrated water. The water pressure in the soil layer below the water table increases along with the increasing of the infiltration time. The condition of the pore-water pressure at the base of the pervious concrete plate and the soil surrounding changes due to the increase in pore water pressure.

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(b)

(a)

Fig. 8. Pore water pressure at Node 149 : (a) node position (b) first condition; (c) second condition.

(c) Based on the results of numerical simulations, it can be seen that there is an increase in pore water pressure in the area of the pervious concrete slab and the surrounding soil. The addition of water pressure occurs along with the flow of water in the drainage channel. Water will seep through the pervious concrete slab and the underlying soil layer, moving downwards in the direction of gravity. Once there is no flow in the drainage channel, the water will press the soil layer in all directions to fill the gaps in the surrounding pores. The pore water pressure at the bottom of the plate is in the range of 18.64 kPa. The higher pore water pressure for three cycle test is obtained at the first condition of simulation. As can be seen in Fig. 8. On the first day of testing, both the first modeling and the second condition of the pore water pressure were negative. When the water has started to infiltrate, in the first 10 minutes the water pressure changes to positive, with a very significant amount. However, with increasing test time, water begins to fill the surrounding pore spaces and begins to flow in the direction of gravity. then the pore water pressure at the bottom of the plate returns to negative, although its value is close to zero. With the increase in soil moisture, then on the second and third day of testing, the pore water pressure until the end of the test time was completely positive. The pore water pressure in the first modeling tends to be higher than the second modeling. On the third day of testing, the difference in pore water pressure between the two was 27.08%. Summary The average infiltration capacity based on the infiltration test shows a smaller result in the second modeling condition with a 10,38% difference of average infiltration capacity between the first and the second condition. The results of the infiltration test simulated with a numerical approach using the

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Geostudio SEEP/W software show that the total water discharge of the second condition of modeling is lower at the third cycle test compares to the first modeling condition. Based on the results of numerical simulations, it can be seen that there is an increase in pore water pressure in the area of the pervious concrete slab and the surrounding soil. The pore water pressure at the bottom of the plate is in the range of 18.64 kPa. Overall, the pore water pressure in the first modeling tends to be higher than the second modeling. On the third day of testing, the difference in pore water pressure between the two-modeling condition is 27.08%. Acknowledgments We sincerely thank P3M Politeknik Negeri Bandung, for funding this research, and Civil Engineering Department Politeknik Negeri Bandung which provides facilities for this research. References [1]

Tennis, P.D.; Leming, M.L.; Akers, D.J. Pervious Concrete Pavements; Portland Cement Association: Skokie, IL, USA, 2004; p. 28.

[2]

Kevern, J.T.; Wang, K.; Schaefer, V.R. Test Methods for Characterizing Air Void Systems in Portland Cement Pervious Concrete. J. ASTM Int. 2009, 6, JAI102451.

[3]

Huang, B.; Wu, H.; Shu, X.; Burdette, E.G. Laboratory evaluation of permeability and strength of polymer-modified pervious concrete. Constr. Build. Mater. 2010, 24, 818–823. [CrossRef]

[4]

Deo, O.; Neithalath, N. Compressive response of pervious concretes proportioned for desired porosities. Constr. Build. Mater. 2011, 25, 4181–4189. [CrossRef]

[5]

Deo, O.; Neithalath, N. Compressive behavior of pervious concretes and a quantification of the influence of random pore structure features. Mater. Sci. Eng. 2010, 528, 402–412. [CrossRef]

[6]

ASTM International. ASTM C1701/C1701M-17a, Standard Test Method for Infiltration Rate of In Place Pervious Concrete; ASTM International: West Conshohocken, PA, USA, 2017.

[7]

ASTM International. ASTM C1781/C1781M-18e1, Standard Test Method for Surface Infiltration Rate of Permeable Unit Pavement Systems; ASTM International: West Conshohocken, PA, USA, 2018.

[8]

International Organization for Standardization. ISO17785-1: 2016: Testing Methods for Pervious Concrete-Part 1: Infiltration Rate; International Organization for Standardization: Geneva, Switzerland, 2016.

[9]

Schaefer, V.R.; Wang, K.; Suleiman, M.T.; Kevern, J.T. Mix design development for pervious concrete in cold weather climates. In Final Report, National Concrete Pavement Technology Center; Iowa State University: Ames, IA, USA, 2006

[10] Ong, S.K.; Wang, K.; Ling, Y.; Shi, G. Pervious Concrete Physical Characteristics and Effectiveness in Stormwater Pollution Reduction. In 2016 Final Report, Institute for Transportation; Iowa State University: Ames, IA, USA, 2006. [11] Yahia, A.; Kabagire, K.D. New approach to proportion pervious concrete. Constr. Build. Mater. 2014, 62, 38–46. [CrossRef] [12] Chindaprasirt,P.;Hatanaka,S.;Chareerat,T.;Mishima,N.;Yuasa,Y.Cement paste characteristics and porous concrete properties. Constr. Build. Mater. 2008, 22, 894–901. [CrossRef] . [13] Cackler, E.T.; Ferragut, T.; Harrington, D.S. Evaluation of U.S. and European Concrete Pavement Noise Reduction Methods; National Concrete Pavement Technology Center, Iowa State University: Ames, IA, USA, 2006; p. 96.

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[14] American Concrete Institute. ACI522.1 - 13Specification for Pervious Concrete Pavement; American Concrete Institute: Farmington Hills, MI, USA, 2013; ISBN 9780870318221. [15] Kishore,R.;Guntakal,S.N.Experimental Investigationson Properties of Pervious Concretefor Its Application in Rural Pavements. Int. J. Eng. Res. Technol. 2017, 8, 111–120 [16] Sonebi, M.; Bassuoni, M.; Yahia, A. Pervious Concrete: Mix Design, Properties and Applications. RILEM Tech. Lett. 2016, 1, 109–115. [CrossRef] [17] Obla, K.H. Pervious Concrete—An overview. Indian Concr. J. 2010, 84, 9 [18] Endawati, J, Diasti, L., Enung, Characteristics of Pervious Concrete with Environmental Friendly Based Binder, Applied Mechanics and Materials, ISSN: 1662-7482, Vol. 865, pp 263-269, 2017 [19] Vedat Batu, Aquifer Hydraulics: A Comprehensive Guide to Hydrogeologic Data Analysis, John Wiley & Son, 1998 [20] Soewarno, Hidrologi, Penerbit NOVA, 1995

Key Engineering Materials ISSN: 1662-9795, Vol. 934, pp 171-178 doi:10.4028/p-1r6f9d © 2022 Trans Tech Publications Ltd, Switzerland

Submitted: 2022-07-27 Accepted: 2022-09-06 Online: 2022-11-28

Experimental Study on Multi-Segmented Circular or Non-Circular Pipes to Determine the Pipe Stiffness Devanand Chelota, Priyank Upadhyaya*,b Department of Mechanical Engineering, Birla Institute of Technology & Science Pilani, Dubai Campus, P.O. Box 345055, Dubai, United Arab Emirates

1

[email protected], *, [email protected]

a

Keywords: GFRP pipes, unitary, multi-segmental, Ring Stiffness test, tongue and groove joint Adhesive, Strength, Stiffness

Abstract. A study was carried out to test unitary as well as multi-segmental circular and noncircular glass-reinforced plastic pipes under compression. Three pipe shapes were tested (box, semi-elliptical and round) under identical conditions using the ring stiffness method of testing. Hand layup technique was used for manufacturing the pipes. The pipes were designed with inner layers, central core and outer layers. The inner layer acts as the corrosion barrier layer, the core consists of blended sand impregnated with resin and the outer layer acts as the external layer which protects the pipe from impact. Based on the experimental results, it was concluded that multi-segmental pipes are stiffer and more flexible than unitary pipes because of the loading carrying capacity of the tongue and groove joint (TGJ). Introduction Underground infrastructure, such as sewer collection systems, usually shows signs of emergency for rehabilitation only after the damage is done. This results in a lack of standard maintenance actions, unless an actual failure occurs. The maintenance, repair, and rehabilitation of sewer system infrastructure is no small task. Hence, there is a strong incentive and need to find alternative rehabilitation technologies for existing pipes, which are cost-effective and require as little disruption as possible [1]. A wide range of civil infrastructure such as decks, storage tanks, bridges, and pipelines are constructed using glass-reinforced plastic (GRP) due to its low weight-to-strength ratio and corrosion resistance properties [2]. In order to repair damaged sewage systems, GRP pipe manufacturers receive requests from many clients to fabricate circular and noncircular pipes. As part of the manufacturers ability to satisfy the customers' requirements, extensive testing, especially for newly requested shapes is performed. Because this process is laborious and costly, manufacturers require a reliable and costeffective alternative. To maintain the effectiveness of the system and maintain the purity of the groundwater, damaged sewer pipelines should be repaired. Typical rehabilitation methods require intensive trenching, long operating periods, and are often inconvenient for traffic and the community. Using a thin-walled GRP tube to regenerate sewers is one of the methods used when relining sewers. The procedure to rehabilitating a sewer involves inserting or pulling a new pipe into an existing pipeline [3]. The space between the panel and the channel is filled with injection mass, which works well for various materials used for construction of the channel [4]. There are currently a large number of sewers with noncircular shapes, which requires noncircular pipes with special design criteria, fabrication tools, and experimental work in the laboratory. It is often desirable to manufacture panels in two or more longitudinal sections due to transportation problems or access problems (e.g. only allowing entry through manholes), or in large structures with complicated interiors. Assembling a multi-segmental panel may take place above or below ground using a structural adhesive that is applied in the tongue and groove joints (TGJs), arranged at intersection points, i.e. at zero bending points. Whenever bigger diameter structures are repaired, multi-segmented pipes result in a substantial reduction in transportation costs. The manufacturing of the pipes in multiple segments allows for nested loading, thereby increasing the number of sections

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that can be transported in one container. In situations where access to an existing pipe structure is limited, multi-segment rehabilitation is often the only option. Depending on the site conditions, pipe segments may be joined before entering the existing structure adjacent to the pipe structure to be rehabilitated, or within the pipe structure itself. For rehabilitation of underground sewers, one of the most important tests is to calculate the stiffness, which is the measure of deflection for a given load [5]. All buried pipe will experience vertical deflection as a result of installation and imposed loads, with several parameters determining the amount of pipe deflection [6]. These parameters include solid load, live load, native soil characteristics, trench width, hunching, and pipe stiffness, with the pipe stiffness being the most adjustable. Extensive testing of GRP unitary circular pipes have been conducted in the past with different pipe geometry [7], however, there are very few experimental data on GRP multi-segmental pipes with non-circular shapes under ring stiffness test [8]. This study indicates the results from an experimental study conducted on several multi-segmental and unitary pipe shapes, a study that can be applied to comparing equivalent pipe stiffness between unitary and multi-segmental pipes of different shapes. Experimental Study Manufacturing of pipes/segments. The fabrication of the GRF pipes was done using hand layup process [9] [10]. The advantage of the hand layup technique is that various shapes and sizes can be fabricated without any apparatus or large equipment. The pipes were designed with inner layers, central core and outer layers. The inner layer acts as the corrosion barrier layer consisting of a surface tissue and chopped strand mat impregnated with an isophthalic resin. The core consists of blended sand impregnated with dicyclopentadiene (DCPD) resin. The outer layer is made up of unidirectional fabric and chopped strand mat which is impregnated with an isophthalic resin. For the current study, isophthalic resin Crystic-491 E PA and DCPD resin Crystic 2-451 were procured from Scott Bader. The manufacturing process involves the pipes being made in multiple sections, ranging from 2 sections to 4 sections, depending upon the application. A tongue and groove joint (TGJ) was adopted to join these segments, which is glued together with an adhesive [11] [12] [13]. In a previous work, we have concluded that 65% ARF mix gives better performance for TGJ joints [14]. Therefore, in the current work all the TGJ were glued together using ARF 65%. With the introduction of joint sections, these areas are prone to cracking, crazing and delamination that can damage and lower the mechanical performance of these pipes [15] [16] [17]. To counter this scenario, the tongue section, which replaced the weaker core material, is fabricated from solely using unidirectional fibers, thus making it more resilient than the pipe wall layup [18] [19]. The assembling of a two-piece box culvert and two-piece semi elliptical pipes followed the same procedure in which the crown (top section) was lowered down and joined into the invert (bottom section) of the pipe. Prior to inserting the crown in the invert, the 65% ARF mix was applied to the tongue and groove section on the segments. Immediately following the application of the adhesive, both segments are glued together and cured for a period of at least three hours. However, a TPC and an FPC pipe were joined together by placing the pipe specimens vertical and by joining the tongue and groove section of the pipe in longitudinal direction as shown in Fig. 1. The different shapes that are used for the experimental study are Unitary Circular (UC1, UC2), Unitary Box (UB), Unitary Semi Elliptical (USE), Three-Piece Circular (TPC), Four-Piece Circular (FPC), Two-Piece Box culvert (TPB), Two-piece Semi Elliptical (TSE). The dimensions of the all the pipes considered here are listed in Table 1. Here, UC1 and UC2 are both circular but of different dimensions. The entire manufacturing process is continuous and since it is performed manually, there is complete control to ensure the bonding of the different layers. After manufacturing long pipes, sample with smaller length as per the instructions provided in ASTM D2412 were cut and conditioned [20] [21].

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Assembling of multi-piece pipes

Figure 1: (a) Segments of TPC before joining (b) TGJ using ARF mix (c) Final TPC pipe after the joint is cured (d) Inside view of the cured final TPC pipe (e) Cut specimen of the TPC pipe for the ring stiffness test Testing Procedure The ring stiffness tests on various shapes consisted of testing 305-mm-long (+/2mm) GRF pipe specimens for internal diameter (𝑑𝑑𝑖𝑖 ) or height (ℎ) less than 1524mm. 20% of the 𝑑𝑑𝑖𝑖 or ℎ was calculated for obtaining the length of the specimens with 𝑑𝑑𝑖𝑖 or ℎ greater than 1524mm as per the requirements and specifications listed in ASTM D2412 [21]. The procedure subjects the pipes to compressive point loads using steel plate of which strain, load, and displacement measurements were recorded. The wall thickness was measured by dividing the pipe into eight equally spaced halves and the average wall thickness was calculated. For the circular shape pipes, the test was repeated for three orientations. The location of the minimum thickness was marked as a reference (𝜃𝜃 = 0°). In the first test run, pipe was loaded along the axis that passes through the reference line. In the subsequent test runs, the pipe was rotated by 35° and 70° and before applying the compression load. Table 1: Dimensions of the Ring Stiffness test specimens Shapes UC1 & TPC UC2 & FPC UB & TPB

USE & TSE

Dimension Internal Diameter: 2200[mm] Length: 440[mm] Wall Thickness 32[mm] Internal Diameter: 4876.80[mm] Length: 975.36[mm] Wall Thickness 90[mm] Height: 316[mm] Width: 536[mm] Length: 305[mm] Wall Thickness: 25.4[mm] Height: 1346.2[mm] Width: 1541.78[mm] Length: 305[mm] Wall Thickness: 28[mm]

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The 𝑑𝑑𝑖𝑖 or ℎ of the pipe were used to compute the percentage of deflection for the specimen. The compression strain for these tests was kept at 5% as per the specifications stated in ASTM D2412 [21]. The schematic presentation of the test procedure is shown in Fig. 2 and the actual experimental setup is shown in Fig. 3. The joint sections in Fig. 3 are highlighted inside a green color oval. The upper plate was brought into contact with the specimen and a small initial load was applied that was necessary to keep the specimen in place as prescribed in the ASTM D2412 [21].

Figure 2: Schematic representation of the TPC test specimen positioned at (a) 0-degree; (b) 35-degree; (c) 70-degree Compression of the specimen was set at a constant rate of 12.5mm/min as per ASTM D3262 and ASTM D2412 [20]. Load and deflection measurements were taken relative to the movement of the bearing plates. Any occurrence of liner cracking, crazing, wall delamination, surface cracks or structural damage observed during the test was monitored. Results and Discussion. In all scenarios, the load is seen to increase as the deflection of the pipe specimen increases. The experimental data were used in Eq. 1 defined in ASTM D2412 to determine the pipe stiffness, PS.

𝑃𝑃𝑃𝑃 = 𝑆𝑆𝑆𝑆 = 𝐸𝐸𝐸𝐸 = 0.149

𝐹𝐹

(1)

𝛥𝛥𝛥𝛥 𝐹𝐹𝑟𝑟 3 𝛥𝛥𝛥𝛥

= 0.149𝑟𝑟 3 (𝑃𝑃𝑆𝑆)

(2)

where “E” is the pipe ring modulus of elasticity, “I” the moment of inertia for unit length of pipe wall for ring bending, “EI” the flexural stiffness, and “r” is the vertical dimension or (𝑑𝑑𝑖𝑖 ) or (ℎ) divided in half. The recorded load deflection data is used to determine the difference in the EPS by comparing a unitary pipe and a multi-segmental pipe. The experimental results of EPS obtained for all specimens is defined in Table 2.

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Figure 3: Experimental setup of Ring Stiffness test depicting the position of the TGJ in (a) UC1; (b) TPC at 0-degree; (c) TPC at 35-degree; Table 2: Experimental results for the Ring Stiffness test Shape

Max Load (kN)

Stiffness (MPa)

UC1

11.022

0.227

TPC

15.530

0.321

UC2

62.405

0.262

FPC

69.597

0.305

UB

20.619

4.295

TBC

22.390

4.709

USE

7.045

0.345

TSE

11.606

0.568

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Ring Stiffness Test: Circular

18

Ring Stiffness Test: Circular

80

UC2 - 0deg

UC1 - 0deg

16

70

UC1 - 35deg

UC2 - 35deg UC2 - 70deg

UC1 - 70deg

14

60

TPC - 0deg

FPC - 0deg FPC - 35deg

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10 8 6

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Figure 4: Load displacement data from the experiments Fig. 4 shows the load vs displacement data which in all scenarios indicate that the load resilience of the multi-segmental pipes increases due to its stronger TGJ which is manufactured with only unidirectional fibers. It is worth mentioning that TPC and FPC pipe specimens showed a higher stiffness than the UC1 and UC2 respectively. Conclusion This study aims to determine the pipe stiffness of unitary and multi-segmental GFRP pipes of different shapes and sizes. Testing live size pipes using the ring stiffness test procedure established by ASTM D2412 [21] was used to calculate the stiffness of the pipes and comparisons were made between unitary and multi-segmented pipes. In light of the improved TGJ system for the multisegmental pipes, the test results demonstrate that multi-segmental pipes have higher pipe stiffness due to the increased load carrying capacity than unitary pipes. Acknowledgement Authors acknowledge the support from Channeline International Fiberglass Manufacturing LLC from Dubai for providing the materials and test facilities for the research work.

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References [1] J. A. V. A. M. A. a. A. H. K. Santiago M. Reyna, CONSTRUCTION TECHNOLOGIES FOR SEWER REHABILITATION, J. Constr. Eng. Manage. 1994.120:467-487. [2] T. K. Mukhopadhyay, "Rehabilitation of old water supply pipes," in AFFORDABLE WATER SUPPLY AND SANITATION. [3] C. Macey, A. Braun, K. Zurek and a. S. Cournoyer, Large-Diameter, Non-Circular Trunk Sewer Rehabilitation Using GRP Composites, Pipelines, 2016. [4] B. S. Bednarz, "Numerical Verification Of The Thickness Of Grp Panels For Modernization Of Large-Diameter Sewage Collectors With Non-Circular Cross-Sections," Means of Distributed Fiber Optic Sensors (DFOS)." Sensors 21.19 (2021): 6607. [5] R. R. a. M. R. Habibagahi, "On The Stiffness Prediction of GFRP Pipes Subjected to Transverse Loading," KSCE Journal of Civil Engineering, 2018. [6] S.P.G.S.M. J.S.P. G. &. A.P.M.S. Nimish Kurien Thomas, "Stress Analysis of Underground GRP Pipe Subjected to Internal and External Loading Conditions," International Journal of Advanced Mechanical Engineering. [7] T. Masada, "Improved Solution for Pipe Stiffness as Measured by Parallel-Plate Load Test Method," 2006. [8] I. K. A. M. T. W. &. R. F. A.C. Seibi, "Shape Factor for Glass-Reinforced Plastic Pipes With Noncircular Shapes Under Diametral Loading - An Experimental Study," 2011. [9] D. B. C. W. A. C. &. K. P. M. Elkington, "Hand layup: understanding the manual process," Advanced Manufacturing: Polymer & Composites Science, 2015. [10] WRc Engineering, "Information and Guidance Note No. 4-34-02," WRc WAA 4-34-02, no. 1, 1986. [11] K. W. Robert D. Zimmerly, "MULTI-PIECE PIPE CLAMP". United States Patent 4568115, 4 February 1986. [12] J. L. G. W. J. M. Joseph D. Melogranaa, "Adhesive tongue-and-groove joints between thin carbon fiber laminates and steel," Composites: Part A 34 (2003) 119–124, 2002. [13] J. Z. O. C. George J. Dvorak*, "Adhesive tongue-and-groove joints for thickn composite laminates," Composites Science And Technology, 2001. [14] A. T. S. K. &. P. U. Devanand Chelot, "Enhanced Strength and Toughness of Repurposed Glass Fiber Reinforced Adhesive Joints for Sewage Applications," International Journal of Adhesion and Adhesives, 2022. [15] B. C. e. al, "Mechanical behavior of novel GFRP foam sandwich adhesive joints," Composites Part B, 2017. [16] O. C. e. al, "Fatigue strength estimation of adhesively bonded tongue and groove joint of thick woven composite sandwich structures using genetic algorithm approach," International Journal of Adhesion & Adhesives, 2011. [17] I.I.,. S. S.-S. a. E. J. B. Lorena M. Fernandez-Canadas, "Effect of adhesive thickness and overlap on the behavior of composite single lap joints," MECHANICS OF ADVANCED MATERIALS AND STRUCTURES, 2019.

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[18] O. C. e. al, "The effect of design on adhesive joints of thick composite sandwich structures," Journal of of Achievements in Materials and Manufacturing Engineering, 2008. [19] M. R. A. e. al, "Improvement of an adhesive joint constructed from carbon fiber fiber reinforced," Composites Part B, 2016. [20] ASTM International, "ASTM D3262 - Standard Specification for “Fiberglass” (Glass-FiberReinforced Thermosetting-Resin) Sewer Pipe," D 3262 – 06, 2008. [21] ASTM Standard, "Standard Test Method for Determination of External Loading Characteristics of Plastic Pipe by Parallel-Plate Loading," ASTM D2412.

Key Engineering Materials ISSN: 1662-9795, Vol. 934, pp 179-188 doi:10.4028/p-l09342 © 2022 Trans Tech Publications Ltd, Switzerland

Submitted: 2022-03-28 Revised: 2022-06-06 Accepted: 2022-06-08 Online: 2022-11-28

Coal Ash and Rice Husk Ash Binder for Manufacturing Hollow Blocks Zircon S. Culilang School of Engineering, Northern Iloilo Polytechnic State College, Estancia, Iloilo, 5017 Philippines [email protected] Keywords: Hollow Blocks, Coal Ash, RHA, Binder, Alkali Activator

Abstract. This paper presents the development of a binder from coal ash and rice husk ash for manufacturing hollow blocks. Experiments were conducted to determine the compressive strength of the coal ash and rice husk ash as binders using cylindrical specimens at different curing periods. The mixture proportion with the highest compressive strength in the experiment was adopted in making the 4” hollow blocks. The compressive strength of the hollow blocks with the coal ash and rice husk ash binder was also determined at different curing periods. Cylindrical specimens with four different alkali activators were made and subjected to compression testing on the 3rd, 5th, and 7th day of curing. The mean compressive strength of the cylindrical specimen with all sodium hydroxide activator (RS0CS0) after 7 days curing is 2.54 MPa as compared to 3.05 MPa with the 2.5 sodium silicate to sodium hydroxide ratio (RS1CS0). As the amount of sodium silicate from rice husk was doubled (RS2CS0), the average compressive strength of the 7 days old cylindrical samples increased from 3.05 MPa to 3.49 MPa. Meanwhile, the 7th-day compressive strength of the cylindrical specimen with commercial silica (RS0CS1) was 5.5 Mpa. Hollow blocks were also manufactured based on the material proportion that utilizes RHA with the highest compressive strength or the design mixture RS2CS0. These blocks have a mean compressive strength of 1.22 MPa on the 3rd day, 1.88 MPa on the 7th day, and 2.7 MPa on the 14th day of the curing period. Introduction Cement is the most common binder in concrete, and its demand increases as infrastructure development continue to grow rapidly [1]. In the Philippines, the concrete hollow block is one of the most commonly used building materials and is produced using cement as its binder. The production of cement, however, is an energy-intensive process that constitutes a significant portion of anthropogenic carbon dioxide emissions and other greenhouse gases [2]. Numerous studies have been conducted to explore industrial wastes as supplementary cementitious materials in concrete without sacrificing their long-term performance and reliability [3]. One of the promising industrial wastes that could replace Portland cement in the near future is coal ash from thermal power plants. It focuses on the activation of coal ash using an alkali solution to be used as binder materials for concrete. French scientist Joseph Davidovits first coined these materials as inorganic polymers, geopolymers, and alkali-activated concrete [4]. This study explores the development of coal fly ash and rice husks ash as a binder for manufacturing hollow blocks so that this could bring a more valuable usage of the said industrial and agricultural waste materials. Fly ash-based binder offers several economic benefits over Portland cement. The cost of one ton of fly ash is only a tiny fraction, if not free in some parts of the world, compared to the cost of one ton of Portland cement [5]. The eco-friendly hollow block uses coal-ash-based geopolymer (alkali-activated) as its binder. Depending on the raw material selection and processing conditions, alkali-activated coal ash in concrete can exhibit a wide variety of properties and characteristics, including high compressive strength, low shrinkage, fast or slow setting, acid resistance, fire resistance, and low thermal conductivity, although not inherent to all geopolymeric formulations [6]. The reaction of coal ash with an alkali solution containing sodium hydroxide and sodium silicate results in a material bond with a three-dimensional polymeric chain and ring structure consisting of Si-O-Al-O [7].

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The main components in making the geopolymer are coal ash, sodium hydroxide, and sodium silicate [8]. Of the three (3) raw materials needed for the binder, sodium silicate or water glass is not commonly available in the Philippine local market. Many researchers reported that sodium silicate plays a vital role in the alkali activation process [ 9, 10, 11, 12, 13]. The alkali solution used in the study was made by mixing the caustic soda with tap water and the sodium silicate from rice husk ash. The method used in the production of sodium silicate from RHA was adapted from the procedure outlined in studies of De Sousa et al., [14] and Todkar et al. [15]. Sodium silicate was made by boiling rice husk ash in sodium hydroxide solution to dissolve its silica content. Cylindrical specimens of various design mixes and curing periods were used in testing to evaluate the effect of sodium silicate from RHA on the compressive strength of concrete using the coal ash as the binder. The material proportion of the best design mix of the binder in terms of compressive strength was then adopted to produce the hollow blocks. Raw materials in the production of the hollow blocks were river washed sand and a binder made from coal ash and rice husk ash. The performance of the hollow blocks was then evaluated in terms of their compressive strength in the material laboratory. This study was conducted to develop a binder from coal ash and rice husk ash for manufacturing hollow blocks. Specifically, it seeks to determine a good material mix and calculate the compressive strength of the hollow blocks with the coal ash and rice husk ash binder at different curing periods. Materials and Methods This study was divided into three stages: first, different designs mix was tested in the laboratory to determine its compressive strength at different curing periods; second, the hollow blocks were produced based on the material composition with the highest compressive strength; and lastly, the compressive strength of the hollow blocks with the coal ash and rice husk ash binder was determined. Characterization of Raw Materials Coal Fly Ash. The coal fly ash used in this study was taken from Panay Energy Development Corporation (PEDC) coal-fired power plant in Iloilo City, Philippines. Data from the power plant reveals that the coal fly ash was class C (CaO>20%). The chemical compositions of the coal fly ash are all presented in Table 1. Table 1. Chemical Composition of Coal Fly Ash. Chemical Composition Silicon Dioxide (SiO2) Aluminum Oxide (Al2O3) Iron Oxide (Fe2O3) Titanium Dioxide (TiO2) Calcium Oxide (CaO) Magnesium Oxide (MgO) Sodium Oxide (Na2O) Potassium Oxide (K2O) Sulfates

Coal Fly Ash 33.05 % 18.59 % 10.06 % 0.45 % 20.76 % 3.64 % 0.17 % 1.36 % 10.42 %

As shown in Table 1, the SiO2/Al2O3 ratio is 1.8, a little bit short of the range of values (2.0-3.5) for very good geopolymerization [17]. It was also noted that the coal ash contains high calcium oxide or lime. The coal fly ash was not processed but was used in the study as received. Rice Husk Ash (RHA). The whole raw rice husk used in this study was sourced from a local rice mill in Estancia, Iloilo, Philippines. The rice husk was sun-dried and burned in an open pit furnace to gather the ash. Typical rice ash husk ash contains 86-97% silica after complete combustion. Some researchers from the Philippines and Thailand reported that silicon dioxide (SiO2) in RHA ranges from 70-92% [7, 18, 19]. The important compound in rice husk ash is silicon dioxide, and this will be reacted with sodium hydroxide to yield sodium silicate.

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Aggregates. The fine aggregates used in the experiment were a natural siliceous, fine aggregate river washed sand. The sand was washed before use to ensure that dust and organic materials were not present. The aggregate was soaked in water for a few minutes and was placed on plastic sheets to dry naturally to reach the saturated and surface-dried condition. Sodium Hydroxide. The alkaline liquid that was used in this research was a 12M sodium hydroxide solution. It was made by dissolving 480 grams of pure caustic soda (NaOH) powder, a little at a time, into a large volume of water and then diluting the solution to make 1 liter. The solution was kept and let settled in a borosilicate glass container after mixing for 24 hours to fully dissolve the caustic soda powder. The mixture density was recorded at 1.14 kg/l. The caustic soda used in the study bears a 99% purity label on its packaging and was purchased from a local soap ingredients supplier in a 25 kg sack from a local supplier. Sodium Silicate. Rice husk ash was used to produce sodium silicate. Silica in RHA was dissolved in NaOH (3M) with 200 gms of rice husk ash per liter of the solution. The process of making the solution is discussed in the preparation of the alkali activator presented in the next section. For comparison, commercial silica gel was also prepared to make a sodium silicate solution with SiO2/NaOH=1 and solid/liquid=1. The silica gel was bought from a retailer in OLX (a local trading web portal). Water. Tap water from a local water utility was used to make sodium hydroxide solution and to enhance the workability of fresh composite concrete as needed. Preparation of the Alkali Activator A sodium hydroxide solution with a concentration of 12 M was prepared by dissolving 480 grams of caustic soda powder in 1 liter of water. The alkali solution was prepared a day (24 hours) before the mixing process to fully dissolve the caustic soda in water. Sodium silicate was produced from the rice husk ash. In a stainless-steel vessel, rice husk ash was added to sodium hydroxide (3M) in 1:5 ratios (i.e., 35 gm of ash in 175ml of NaOH). It was then kept for heating (boiled) on an electric stove for 60 minutes. This involves the digestion of the rice husk ash with caustic at specific conditions. In this process, the silica in the ash gets extracted with caustic soda to form a sodium silicate solution. The density of the mixture was recorded at 1.17 kg/l with 200 gms of rice husk ash per liter of the solution. For RS0CS1, which was used as controlled samples, the sodium silicate was prepared by dissolving 200 gms of caustic soda in 500 ml of water in a beaker; then it was heated to boil before 200 grams of silica gel was added. The beaker was heated until all the solid particles dissolved in the solution and with occasional stirring. Cylindrical Specimen Preparation The details of material composition used to prepare each 15 specimens of the four (4) types are presented in table 2. Table 2. The proportion of Mixtures Used for Cylindrical Specimen [kg]. Design Mix No. RS0CS0 RS1CS0 RS2CS0 RS0CS1

Water to Binder (FA+Alkali Sol.) Ratio 0.4 0.4 0.4 0.4

Aggregates Gravel 12.4 12.4 12.4 12.4

Fine Sand 6.7 6.7 6.7 6.7

Coal FlyAsh

Sodium Hydroxide

5.5 5.5 5.5 5.5

1.48 0.42 0.42 0.42

Sodium Silicate (from RHA) 0 1.06 2.12 0

Sodium Silicate (from Silica Gel) 0 0 0 1.06

Mix RS0CS0 means rice husk ash sodium silicate equals zero and commercial sodium silicate is also zero, while RS1CS0 means approximately 1 kg sodium silicate from rice husk and with no commercial sodium silicate. For all design mixes, coal fly ash were kept at 5.5kg, and water to binder ratio was kept at 0.4.

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During the preparation, dry raw materials were kept in sacks or buckets while the alkali solutions were stored in a high thermal plastic container. The weights of the raw materials were measured in the weighing scale and were used to calculate the proportion of the different design mix to make fifteen (15) specimens (7mm xm150mm cylinder). The method adapted for making an alkali-activated concrete had been used by Davidovits et.al, Mustafa et. al., and Soltaninaveh [9,10,12] generally consisted of mixing the dry materials for a few minutes followed by adding alkaline activators and continuing the mixing for another couple of minutes. The steps in the preparation of cylindrical specimens with the coal fly ash and rice husk ash binder are shown in figure 1. It includes preparation of sodium hydroxide and sodium silicate from rice husk, measuring and mixing of dry and liquid raw materials, molding, storing, demolding, curing, and specimen testing. Preparation of Sodium Hydroxide

Preparation of Sodium Silicate

Storing of Liquid Solutions

Measuring (Weighing) of Raw Materials

Mixing of Dry Raw Materials

Mixing of Liquid Raw Materials

Moulding of the Mixture

Thorough Mixing of Dry & Liquid Raw Materials

Storing (24 hrs)

De-moulding of the Specimen

Measuring and Testing

Curing

Figure 1. Process Flow in Preparing the Cylindrical Specimen. After the raw materials were ready, dry components were mixed all together uniformly on the mixing floor while the wet components were also poured in a stainless steel bucket one after the other and were stirred for a few minutes. The liquid solution was then added to the dry mixture and continued the mixing until homogeneity was observed. The fresh mixture was then placed in a cylindrical mold 75 mm in diameter and 150 mm in height and was compacted with a steel rood. The specimens were kept in the laboratory room for 24 hrs before the molds were removed. The cylindrical samples were left inside the laboratory room at ambient condition and room temperature until it was used for compression testing on the 3rd, 5th, & 7th days curing period. Test Procedure of the Cylindrical Specimen The hardened cylindrical samples were analyzed according to their compressive strength. The Compressive strengths of the prepared specimen at different curing periods were tested based on the ASTM C39 - Standard Test Method for Compressive Strength of Cylindrical Concrete Specimens [16]. Five (5) sample replicates for each design mix were tested on the 3rd, 5th & 7th day curing.

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Instrumentation and Performance Parameters The compression test was conducted on the specimen prepared using a microcomputer hydraulic universal testing machine (model WEW-1000B manufactured by Jinan Liangong Testing Technology Co., LTD) at the NIPSC Engineering Materials Laboratory. This test method consists of applying a compressive axial load to the molded cylinders at a rate that is within a prescribed range of ASTM C39 until failure occurs. The compressive strength of the specimen is calculated by dividing the maximum load attained during the test by the cross-sectional area of the specimen [16].

Figure 2a & 2b. Cylindrical specimens were measured and tested in the universal testing machine. CHB Prototype Production The hollow blocks with the coal ash and rice husk ash binder were made at NIPSC Engineering Materials Laboratory. In making the hollow blocks, the same procedures were followed in the preparation of the cylindrical specimen that was used except for the water that was added a little at a time to make the mixture a little drier that suited the hollow block mold. By weight, the Liquid-CFAAggregate proportion that was used in making the hollow blocks is approximately 1:4:12. Nine (9) specimens were made for laboratory testing to determine the compressive strength of the hollow blocks with coal ash and rice husk ash binder on the 3rd, 7th & 14th day of the curing period.

Figure 3a & 3b. The raw materials used and the molding of the hollow blocks. Evaluation Procedure and Instrumentation in Testing the Hollow Block The compressive strength and density of the hollow block with coal ash and rice husk ash binder were tested based on ASTM C140 [21] at NIPSC engineering materials laboratory on the 3rd, 7th, and 14th-day curing period. For each hollow block specimen, the face-shell and web thickness were measured using a caliper, and then the net cross-sectional area was calculated and recorded. The hollow block specimen was then brought to the universal testing machine to measure the maximum load it can carry before material failure occurs. The maximum load per net cross-sectional area was then calculated and tabulated as the maximum compressive strength.

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Figure 4a & 4b. The hollow block’s face shell and web thickness were measured using a caliper and tested in the universal testing machine. Results and Discussion The Compressive Strength and Density of the Cylindrical Specimen As previously stated, four (4) different types of mixtures were made during the experimental program concerning the amount of sodium silicate solutions from RHA. The primary objective of the test was to determine the compressive strength of the binder using coal fly ash and rice husk ash which is suitable for the hollow block that meets the Philippines standard requirement. Five (5) specimens from each type of mixture were measured and tested on the 3rd day, 5th day, and 7th day of curing at room temperature in the laboratory. The diameter, height, weight of each cylinder, and the maximum load it can carry were recorded. The data was tabulated in a spreadsheet and the compressive strength of the cylindrical specimens was calculated. The results were summarized and presented in table 3. Table 3. The Mean Compressive Strength [MPa] and Density [kg/m3] of Various Mix Design of the Cylindrical Specimen on 3rd, 5th, & 7th Day Curing Period. Design Mix No.

Density (3 Days)

1 2 3 4

2,190 2,178 2,219 2,250

Compressive Strength (3 Days Curing) 0.94 1.06 1.20 3.07

Density (5 Days) 2,169 2,167 2,171 2,187

Compressive Strength (5 Days Curing) 1.18 1.46 1.62 3.79

Density (7 Days) 2,123 2,106 2,111 2,137

Compressive Strength (7 Days Curing) 2.54 3.05 3.49 5.50

As expected, all the compressive strength of the different design mixtures increases with the curing period. The 7 days mean compressive strength of the cylindrical specimen with all sodium hydroxide activators is 2.54 MPa (RS0CS0) as compared to 3.05 MPa (RS1CS0) with the 2.5 sodium silicate to sodium hydroxide ratio. As the amount of sodium silicate from rice husk was doubled, the average 7-day compressive strength increased from 3.05 MPa (RS1CS0) to 3.49 MPa (RS2CS0). When commercial silica was used (RS0CS1), the 7th-day compressive strength of the cylindrical specimen was 5.5 MPa as against 2.54 MPa with the sodium hydroxide only activator. Furthermore, the data revealed that the compressive strengths of the cylindrical geopolymer were far lower than the strength reported by other researchers (>10 MPa) on the same curing time even with RS0CS1 (5.5 MPa), the one with commercial silica. This is because fly ash varies in the sources of raw materials and combustion technology. In terms of criteria presented by Davidovits et al., (2014), the coal ash used has SiO2/Al2O3 ratio equals 1.8, a little bit short of the range (2.0-3.5), and contains high calcium oxide, which is not suitable for high strength application.

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On the other hand, the density of the four (4) different material compositions, as shown in table 3, variation is minimal. The density of AAC specimen ranges from 2,190 kg/m3 to 2,250 kg/m3 after 3 days and ranges from 2,123 kg/m3 to 2,137 kg/m3 after 7 days. The one with commercial silica being the densest and the one with only sodium hydroxide as the lightest; all specimens, however, are classified as a normal weight type of concrete of 2000 kg/m3 or more [20]. Differences in Compressive Strength The compressive strengths of the cylindrical specimen are proportional to the curing period as expected. Figure No. 1 shows the trends as the compressive stress at different curing times of the different design mixtures were compared against each other.

Figure 5. The Mean Compressive Strength of Various Mix Proportions of the Cylindrical Blocks at Different Curing Periods. Figure 5 shows that the compressive strength of the cylindrical specimen increases with the addition of sodium silicate from rice husk ash. To determine if this improvement is significant, the differences in compressive strength between mix RS0CS0 and mix RS1CS0 were tested using Mann– Whitney U. Similarly, the differences in compressive strength between mix RS1CS0 and mix RS2CS0 were also tested to ascertain the effect of increasing the sodium silicate from rice husk ash. The results are summarized in Table 4. Table 4. Differences in 7 days Compressive Strength between Design Mixtures Using the Mann–Whitney U Test. Mixture RS0CS0 vs RS1CS0 RS1CS0 vs RS2CS0

MannWhitney U 0 1.5

Significance Level 0.009 0.021

Interpretation Significant Significant

The Mann-Whitney U test reveals that the increase in 7 days compressive strength of the cylindrical specimen due to the addition of sodium silicate from rice husk ash (RS1CS0) is significant. The computed probability is 0.009 (0.9%), which is less than the assumed 5% level of significance. The increase in compressive strength of RS1CS0 implies that the addition of sodium silicate from rice husk ash enhances its mechanical characteristic substantially. It also affirms the data in the study of Davidovits J. [9]; Mustafa et al. [10]; Reddy et al. [11]; Soltaninaveh [12]; Rangan et al. [13] that sodium silicate enhances the strength of geopolymer concrete. Likewise, by increasing the content of mix RS1CS0 with sodium silicate from rice husk ash by 100%, the compressive strength will be enhanced significantly with a 2.1% chance of being wrong. This also suggests that if the content of sodium silicate from rice husk ash of mix RS1CS0 is doubled, its compressive strength will further increase significantly.

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The Compressive Strength of Hollow Blocks with Coal Fly Ash and Rice Husk Ash Binder The main objective of this study was to develop a binder from coal ash and rice husk ash for manufacturing hollow blocks. Based on the 7-day compressive strength of cylindrical specimen with coal fly ash and rice husk ash binder, mix RS2CS0 has the highest (3.49 MPa). Although mix RS0CS1 was the strongest (5.5 MPa) overall on the 7th day, considering the prohibitive cost of silica, it will make the said material composition not a viable option. The 4"x8"x16" hollow blocks specimen with the coal fly ash, and rice husk binder was therefore made based on design mix RS2CS0. Twelve (12) replicates of the material composition were prepared, and the compressive strengths of hollow blocks were determined on the 3rd, 7th, and 14th day of its curing period. The results of the testing at the different curing periods are presented in table 5. Table 5. The Compressive Strength of the 4” Hollow Blocks with CFA and RHA Binder at different curing periods. Sample

3rd Day

7th Day

14th Day

1 2 3 Mean

1.06 0.74 1.68 1.16

2.10 1.40 2.14 1.88

2.26 3.00 2.91 2.72

As shown in table 5, the 3 days compressive strengths of the hollow blocks range from 1.06-1.68 MPa and a mean of 1.16 MPa. As compared to the 3-day cylindrical specimen (RS2CS0), its mean compressive strength is slightly lower by 0.04 MPa. The hollow blocks reach their 3 days strength since the difference with their cylindrical counterpart is seems unremarkable. The 7th-day compressive strengths of the hollow blocks range from 1.40-2.14 MPa and a mean of 1.88 MPa. As compared to the 7 days compressive strength (3.49 MPa) of the cylindrical specimen with mix RS2CS0, the difference is 1.61 MPa. The wide discrepancy could probably be attributed to the slightly lower water to binder ratio since the water was intentionally added a little at a time during the mixing for workability reasons with the mold. In addition, manual mixing also plays an important factor since a large number of raw materials were mixed all together. Furthermore, the 14th-day compressive strengths of the hollow blocks with CFA and RHA Binder range from 2.26-3.00 MPa and a mean of 2.72 MPa. A notable observation with the hollow blocks during the preparation is the rapid hardening. This had been reported by Davidovits et al. [17]. Rapid hardening or flash setting is caused by the high calcium content of the coal fly ash. Because of the flash setting of the mixture, water was added once in a while to augment its workability with the hollow block mold. The enhancement of the mixture with water might have caused the wide ranges in compressive strength of the hollow blocks. The 14 days compressive strength of hollow blocks exceeds the Philippine Trade Standard Specifications of 350 psi (2.41 MPa) but falls short of the ASTM C129 [22] standard which is 3.45 MPa (500 psi). This implies that for the Philippine market, the hollow blocks with coal fly ash and rice husk ash binder are viable in terms of strength. Summary This study explores a binder from coal ash and rice husk ash for manufacturing hollow blocks. Experiments were conducted to determine the compressive strength of the coal ash and rice husk ash as binders using cylindrical specimens at different curing periods. The design mix with the highest compressive strength was adopted in making the 4” hollow blocks. The compressive strength of the hollow blocks with the coal ash and rice husk ash binder was also determined at different curing periods. The findings revealed that the replacement of sodium silicate from RHA somehow affects the compressive strength of alkali-activated concrete. At 7 days curing, the mean compressive strength

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of the cylindrical specimen with all sodium hydroxide activators (RS0CS0) is 2.54 MPa as compared to 3.05 MPa with the 2.5 sodium silicate to sodium hydroxide ratio (RS1CS0). As the amount of sodium silicate from rice husk was doubled (RS2CS0), the average 7-day compressive strength increased from 3.05 MPa to 3.49 MPa. When commercial silica was used (RS0CS1), the 7th-day compressive strength was 5.5 MPa as against 2.54 MPa with the sodium hydroxide only activator. After the experiment was conducted, twelve (12) hollow blocks based on the mixture that utilizes RHA with the highest compressive strength were manufactured. The mean compressive strength of the hollow blocks with coal fly ash and rice husk ash binder is 1.22 MPa on the 3rd day, 1.88 MPa on the 7th day, and 2.72 MPa on 14th day curing period. In terms of technical viability, the compressive strength of the hollow blocks with coal fly ash and rice husk ash binder was compared with the standards. The strength of the specimen after 14 days of curing falls short of the ASTM C129 standard, which is 3.45 MPa (500 psi) but exceeds the Philippine Trade Standard Specifications of 350 psi (2.41 MPa) References [1] X. Shi, Z. Yang, Y. Liu, & D. Cross, Strength and corrosion properties of Portland cement mortar and concrete with mineral admixtures, in: Construction and Building Materials (2011). 25. 10.1016/j.conbuildmat.2011.03.011. [2] Lei, Y., Zhang, Q., Nielsen, C., & He, K., An inventory of primary air pollutants and CO2 emissions from cement production in China, 1990-2020, in: Atmospheric Environment (2011), 147154. doi:doi:10.1016/j.atmosenv.2010.09.034 [3] N.S. Xie, Up-cycling of Waste Materials: A “Green” Binder Prepared with Pure Fly Ash: submitted to ASCE Journal of Materials in Civil Engineering (2015). [4] J Davidovits, Geopolymer Chemistry and Properties, Compiegne, France (1988): First European Conference on Soft Mineralurgy, pp. 2-48. [5] V. Rangan, D. Hardjito, S. Wallah, & D. Sumajouw, Studies on fly-ash based geopolymer concrete, in: Geopolymer, Green Chemistry, and Sustainable Development Solution, Saint-Quintin, Geopolymer Institute, France (2005), pp. 134-139. [6] P. Duxson, A. Fernandez-Jimenez, J. Provis, G. Lukey, & A. Palomo, Geopolymer technology: the current state of the art: submitted to Journal Material Science (2007), 42, 2917-2933. doi:10.1007/s10853-006-0637-z [7] H. Nguyen, S. Gallardo, F. Bacani, H. Hinode, Q. Do, M. Do, & M. Promentilla, Evaluating the thermal properties of geopolymer produced from red mud, rice husk ash, and diatomaceous earth: submitted to ASEAN Engineering Journal (2014). AUN/SEED-Net. [8] A.M. Aleem, & P.D. Arumairaj, Geopolymer Concrete - A Review: submitted to International Journal of Engineering Sciences & Emerging Technologies (2012), 1(2), 118-122. [9] J. Davidovits, Geopolymer Chemistry and Applications, 2nd ed., Institut Geopolymer, France (2008) [10] A. B. Mustafa, H. Kamarudin, M. BinHussain, I. Khairul Nizar, Y. Zarina, & A. Rafiza, The Effect of Curing Temperature on Physical and Chemical Properties of Geopolymers: in SciVerse ScienceDirect (2011), 22, 286-291. [11] A. Reddy Narender, D. Anitha, & U. Venkata Tilak, Performance of Alkali Activated Slag and Alkali Activated Slag + Fly Ash with various Alkali Activators: submitted to International Journal of Engineering and Technical Research (2014) [12] K. Soltaninaveh, The Properties of Geopolymer Concrete Incorporating Red Sand as Fine Aggregate, Curtin University of Technology, Australia (2008).

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[13] V. Rangan, D. Hardjito, S. Wallah, & D. Sumajouw, Studies on fly-ash based geopolymer concrete: in Geopolymer, Green Chemistry, and Sustainable Development Solution. Saint-Quintin, Geopolymer Institute, France (2005), pp. 134-139. [14] A. M. De Sousa, L. Visconte, C. Mansur, & C. Furtado, C, Silica Sol Obtained from Rice Husk Ash: in Chemistry & Chemical Technology (2009), 3(4). [15] B. S. Todkar, O. A. Deorukhkar, & S. M. Deshmukh, Extraction of Silica from Rice Husk, submitted to International Journal of Engineering Research and Development (2016), 12(3), 69-74. [16] ASTM C39, Standard Test Method for Compressive Strength of Cylindrical Concrete Specimens, ASTM International. West Conshohocken, PA, US (2016). [17] J. Davidovits, M. Izquierdo, X. Querol, D. Antennuci, H. Nugteren, V. Butselaar-Orthlieb, & Y. Luna, The European Research Project GEOASH: Geopolymer Cement Based On European Coal Fly Ashes: in Technical Paper #22. Saint-Quentin, Geopolymer Institute Library, France (2014) [18] M. E. Kalaw, A. Culaba, A, H. Hinode, W. Kurniawan, S. Gallardo, & M. A. Promentilla, Optimizing and Characterizing Geopolymers from Ternary Blend of Philippine Coal Fly Ash, Coal Bottom Ash and Rice Husk Ash: in Materials (2016), 9(7), 580. doi:10.3390/ma9070580 [19] P. Kamhangrittirong, P. Suwanvitaya, P. Suwanvitaya, & P. Chindaprasirt, Synthesis and Properties of High Calcium Fly Ash Based Geopolymer for Concrete Applications: in 36th Conference on Our World in Concrete & Structures: 14 - 16 August 2011, Singapore (2011) [20] J. Clarke, Structural Lightweight Aggregate Concrete. CRC Press, 1993. [21] ASTM C140, Standard Test Methods for Sampling and Testing Concrete Masonry Units and Related Units, ASTM International. West Conshohocken, PA, US (2015). [22] ASTM C129, Standard Specification for Nonloadbearing Concrete Masonry Units, ASTM International, West Conshohocken, PA, US (2014).

Key Engineering Materials ISSN: 1662-9795, Vol. 934, pp 189-195 doi:10.4028/p-n4v65b © 2022 Trans Tech Publications Ltd, Switzerland

Submitted: 2022-05-31 Revised: 2022-07-12 Accepted: 2022-07-12 Online: 2022-11-28

Flexural Behaviors of Modified Recycled Aggregate Concrete Reinforced with Pultruded GFRP SALLEHAN Ismail1,a* and MAHYUDDIN Ramli2,b Department of Built Environment Studies and Technology, Faculty of Architecture, Planning & Surveying Universiti Teknologi MARA Perak Branch, Seri Iskandar Campus, 32610 Seri Iskandar, Perak, Malaysia 1

School of Housing, Building and Planning, Universiti Sains Malaysia, 11800 Pulau Pinang, Malaysia

2

[email protected], [email protected]

a

Keywords: Fibre reinforced concrete, Flexural behaviour, Recycled aggregate concrete,

Abstract. Corrosion of steel reinforcement in underwater concrete structures continues to be a problem in the modern era. The possibility of employing pultruded glass fibre reinforced polymer (GRFP) I beams to replace traditional steel reinforcement in an encased beam is still being explored. The purpose of this research is to further investigate the effects of incorporating treated recycled concrete aggregate (RCA) and discrete synthetic fibres, specifically Polyolefin (PO) and Fibrillated Polypropylene (PP) fibres, in the production of concrete beam reinforced with GFRP I beam for underwater construction. Five beams were constructed and tested in this investigation using a fourpoint loading arrangement. The load–deflection curve, stress–strain curve, and initial and ultimate fracture loads are all included in the test parameter set. The results indicate that treating RCA and fibres greatly increases the flexural capacity of modified recycle aggregate concrete (RAC) beams compared to unmodified RAC beams. Introduction The use of recycled concrete aggregates (RCA) in place of natural aggregates in concrete production is a viable solution to the problem of aggregate deficiency and can help reduce concrete waste. To verify that recycled aggregate concrete (RAC) meets the structural design criteria of both real and civil construction applications, additional full-scale studies and assessments are required. This is because the RCA characteristic factor is inferior to the natural aggregate material. A weak and porous adherent mortar in RCA particles is recognised as the main reason of RAC's lower mechanical qualities compared to NAC, even though both types were made using the identical mixing percentage [1]. As a result, the unfavourable features of RCA may cast doubt on RAC's structural application capabilities and performance. The RAC structure's flexural performance has been extensively studied [2-6]. The majority of research concluded that the flexural behaviour of reinforced RAC beams with RCA content is comparable to that of NAC beams. Precisely at the level of 100% replacement of virgin aggregates with RCA, regardless of the water–cement ratio, various research [2, 3, 5] have emphasised caution. As a result, deflection and fracture width increase, ultimate load falls, and reinforcement-concrete bond strength diminishes. Another study [7] found that using fine RCA in structural concrete increased flexural performance but not significantly. Recent recommendations include improving structural performance and reducing the disadvantages of RC structural designs using RAC. Maruyama et al. [8] and Fonteboa et al. [9] proposed strengthening the design of a reinforcing steel system in concrete to increase the flexural capacity of a RAC beam. Another study found that adding discrete fibres to reinforced RAC improves structural ductility and overcomes the negative effects of RCA [4, 6]. The effects can be amplified by mixing discrete fibre with mineral admixtures like silica fume. With the proposed mix proportioning approach, namely equal mortar volume, Fathifazl et al. [10]. The flexural strength of RAC beams is equivalent or even superior to that of natural aggregate beams, according to the above study.

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The current study aims to extend prior research on the impacts of RAC with treated coarse RCA when exposed to flexural loads. The treatment has been found to be effective in decreasing the negative effects of RCA on concrete performance [11]. This treatment can only strengthen the binding between aggregate and new cement paste, not between aggregate and old cement paste. As a result, the latter components can cause the concrete to fail through the aggregate particle. Fibre insertion may employ a coupling effective strategy to stop microcrack propagation [12]. Adding fibre to concrete can reduce brittleness and increase ductility [13]. The present study also discusses the advantages of employing treated coarse RCA mixed with two types of short synthetic fibres, PO and PP, for structural beams. We also submitted a design for a marine-grade RC structural beam. The beam was developed and manufactured with GFRP I beams instead of steel bars to prevent corrosion. In order to compare the flexural behaviour of structural beams reinforced with GFRP and made of modified RAC with similar structures and normal aggregate concrete, such research is required. Experimental Program Materials, Sample Preparation and Testing: A standard Portland cement (ASTM C150 Type 1) is used in this study. This experiment's coarse RCA is made from a waste-tested concrete cube and separated into two types: treated and untreated. The therapy includes altering the RCA surface structure in two steps. The coarse RCA is first soaked in 0.5 M hydrochloric acid for 24 hours to break down the mortar substances on the RCA surface. The aggregates were then rinsed and drained. An overnight soak in wollastonite (calcium metasilicate) solution increases absorption. This procedure coated the coarse RCA with wollastonite particles to cover the pores and fissures on its surface. Then the coarse RCA was dried for 24 hours at 105°C. Our earlier research [11] describes the materials, experiments, and treatment processes. Other materials include crushed granite and river sand. Table 1 shows the parameters of coarse natural aggregate, treated coarse RCA, and untreated RCA. Contrast the water absorption, specific gravity and crushing and impact parameters of coarse recycled concrete with coarse natural material (Table 1). This is due to the presence of old adhering mortar on the coarse RCA particle. A reduction in water absorption and porosity of the coarse RCA. Table 2 summarises fine aggregate qualities. Two synthetic fibres of varying kinds and qualities were used: PO and PP. Table 3 lists both fibres' specifications. This study uses a super plasticising admixture based of sulphonated naphthalene polymers manufactured to ASTM C494 to achieve a goal slump of 50 mm. The replacement ratio of coarse NA in all RAC mixes is kept constant at 60% by using untreated or treated RCA [11]. On day 28, the concrete mix proportions should achieve a compressive strength of 50 MPa. The tests for cubic compressive strength (fcu) are carried out in line with BS EN 12390-3. It is maintained at 1.2 percent volume fraction of PO and PP fibres in the respective RAC mix (by volume of cement). Our prior research studies [14] indicated the optimum compressive strength with the identical mixture configuration. Table 4 shows the overall concrete specimens' mix proportions and compressive strengths. A total of five beam specimens were prepared for testing. Figure 1 shows the detailed specification of these beam specimens. All specimens were 1500 mm long with a rectangular cross-section of 175 mm × 250 mm. Each specimen was made of different concrete mixes (Table 4) and reinforced by GFRP (I beam section). GFRP is the product of TRUGRID and manufactured through the pultrusion method. GFRP mainly consists of polymer and resin materials, namely, isophthalic polyester and vinyl ester, and reinforced by glass fibres. Tables 5 and 6 present the detailed GFRP product specifications from the supplier. The GFRP surface is smooth. Kwan and Ramli [15] recommended installing shear studs on the flange of GFRP to address the bond slip problem between GFRP and the concrete matrix (see Figure 1). The preceding study also found that the installation of shear studs enhances the ductility and ultimate load of the composite beam. In this study, shear studs are made of steel bolts, with a diameter of 5.0 mm and length of 30.0 mm, and hexagonal nuts (12.0 mm) fixed and screwed to the flange. The tested beams were prepared and casted by using timber formwork. Thereafter, fresh concrete was poured into the formwork in three layers. Each layer was vibrated by using a poker vibrator to ensure proper compaction of the mix. The specimens were stored under laboratory conditions for 28 days before they were subjected to the testing stage. Both

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sides of the beam surface were coated with white paint to ensure that cracks were clearly visible during and after testing. The beam was simply supported and tested under four-point loading conditions to investigate its flexural behaviour. Such beam was tested by using a TORSEE testing machine (with a maximum loading capacity of 50 tons). The loading was monotonically applied in setting with small increments. The load applied to the specimen was sensed by the load cell attached to the crosshead of the machine and connected to a data logger. A complete set of linear variable displacement transducers was installed to measure the displacement and curvature at the mid-span of the beam. Two electrical strain gauges were installed to measure the compressive strain of the beam. One gauge was attached at the top and another at the side surface of the beam centre. Strain (Ɛ) on beams is obtained by the following equation: Ɛ = M/ZE where, M is bending moment, Z is section modulus and E is young‘s modulus. Table 1. Properties of the coarse aggregate Properties of aggregate % Mortar Content Particle density (Mg/m3) Water absorption (%) Agg. crushing value (%) Agg. impact value (%)

Natural aggregate 2.64 0.65 24.32 13.98

Untreated RCA 34 2.57 5.01 29.15 21.78

Treated RCA 2.61 3.98 28.34 19.26

Table 2. Properties of sand Properties

Value

Particle density (Mg/m3)

2.67

Water absorption (%)

1.08

Moisture content

0.31

Fineness 75 µm (%)

0.93

Table 3. Specification of PO and PP fibres Item Product types Average Length (mm) Tensile Strength (MPa) Specific Gravity (g/cm3) Modulus of elasticity (GPa)

PO Barchip 54 27 640 0.92 10

PP Fibrillated 15 310-420 0.90 3.5

Melting Point (°C)

159 - 179

160 - 170

Ignition Point (°C)

> 450

590

Table 4. Types of concrete mix Mix

Types of mix

Cement kg/m3

Water kg/m3

N0

Normal

512

210

Gravel 956

R0

Untreated RCA

512

210

T0

Treated RCA

512

TB

Treated RCA + single fibre

TP

Coarse aggregate kg/m3

Vol. fraction of fibre (%) PO PP fibre fibre

fcu/MPa

Sand kg/m3

SP (%)

RCA -

722

0.2

-

-

58.89

382

574

722

0.2

-

-

49.18

210

382

574

722

0.2

-

-

51.64

512

210

382

574

722

0.25

1.2

-

52.73

512

210

382

574

722

0.25

-

1.2

55.05

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Table 5. Dimension of the GFRP I beam section

d

w

tw

tf

Rf

Area

150mm

75mm

6.25mm

6.25mm

7.53mm

1827mm2

Table 6. Physical and engineering properties of the GFRP I beam Properties Barcol hardness Density Water absorption (24h) Tensile strength Tensile modulus Flexural strength Flexural strength Dielectric strength Elongation

Specification 45 1930 kg/m3 0.6% Max. 437.6 MPa 17.2 GPa 206.8 MPa 11 GPa 13.78 KV/cm 6.29% 1500 1400

A 50 150 50

A Side view

Section A-A

Fig. 1. Specification of the beam specimens Results and Discussions Load Deflection Characteristic. Figure 2 plots all tested beams' load versus central deflection curves. In this test, the beam deflection is affected by the load magnitude. According to Kwan and Ramli [15], the load–deflection curves of GFRP-RC beams can be classified into three stages: (1) before beam cracking, (2) after initial cracking but before FRP yielding, and (3) after FRP yielding. The deflection curves of the study's specimens show these stages (see Figure 2). A steep slope is shown in Figure 2, proving that treatment and inclusion of fibre has no effect on the flexural property before cracking because it is strongly governed by the tensile strength of the concrete matrix [16]. At this stage, the slope load–deflection curves of each beam are nearly identical. When the static load increases, a small decrease occurs, indicating the onset of specimen cracks. The TB specimens had the highest stiffness during the first cracking load, followed by T0, TP, R0, and N0. Up to a certain load level, all specimens show a similar trend as the static load increases. The curve then becomes mildly steep towards the end load point. Figure 2 also shows that the control and modified RAC specimens behave the same. The R0 specimens show the deflection curve's low gradient slope. The

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second stage curve deviation corresponds to the development of shear forces. The specimens show new shear cracks propagating. With increasing shear force, new and existing cracks diagonally propagate. Cracks reduce beam stiffness and thus the slope of the load deflection curve [15]. The specimens have a high deflection rate before failure (N0 = 14 mm), whereas R0 and TP have the lowest rate (12 mm). Due to its lower modulus of elasticity, RAC is known to deform more than conventional concrete. Similar effects have been seen before [5]. So higher deflections were expected due to R0's higher deformability. In the third stage, the applied load dropped dramatically as the concrete beams reached their ultimate load capacity, indicating GFRP fracture. The load dropped to nearly the same level as the R0 beam, then TP and N0. However, the T0 and TB residues showed a slight delay in their final failure stage. The stiffness of the T0 specimens in sustaining the load may be attributed to the treated RCA strengthened the aggregate and concrete matrices bond. The presence of PO fibres (at 1.2%) in the random TB mix may have bridged cracks and slowed crack propagation. Regardless of the concrete mix, the tensile strength and elasticity of GFRP reinforcement strongly influenced the overall behaviour of beams in this stage. Static Load (kN)

200 N0

150

R0

100

T0 TB

50 0

TP 0

5

10 15 Deflection (mm)

20

Fig. 2. Load–deflection curve of all beam specimens Intial Crack and Ultimate Load: The initial crack and ultimate loads for all tested beam specimens are listed in Table 7. All RAC-concrete beams have a larger initial cracking load than conventional aggregate concrete beams (N0). This effect could be due to the rough and angular surface of RCA, which promotes interfacial bonding between the aggregate and cement paste. As a result, the RAC beam's toughness under the first fracture stress is increased. The addition of treated RCA and fibre contributes marginally to the improvement of RAC performance. This conclusion is evident in the modified RAC beam's higher initial cracking load when compared to the unmodified RAC beam (R0). The TB beam has the highest initial cracking load, followed by T0 and TP. The normal variance in specimen ultimate loads was shown to be connected to changes in mechanical strength attributes. The results indicate that the control beam's ultimate load is 174.68 kN. The inclusion of the unmodified RCA has a clear effect on the reduction in load carrying capability of the R0 beam specimens. These specimens achieve a load that is 19% less than the control beam specimens. This finding is due to the low quality of the untreated coarse RCA, which has a detrimental effect on the R0 beam's loadcarrying capacity. The result indicates that the ultimate load capacity of all modified RAC beams, including T0, TB, and TP, is comparable to that of the control specimens. Table 7. First crack and ultimate load of the tested beam Specimens N0 R0 T0 TB TP

First Crack Load (kN) 22.8 26.9 27.9 30 27

Comparison load over N0 1.00 1.18 1.22 1.32 1.18

Ultimate Load (kN) 174.68 141.42 176.15 172.92 177.33

Comparison load over N0 1.00 0.81 1.01 0.99 1.02

Stress-Strain Relationship. The flexural stress vs compressive strain curves for all tested beam specimens are shown in Figure 3. All beams with different types of concrete mix exhibit a similar stress–strain relationship. Typically, the beam curve is linear with the chance of failure. This finding is consistent with findings from other researchers. Because GFRP is not steel, when used to reinforce RC beams, the beams display linear stress–strain behaviour until the ultimate load failure occurs [16].

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Flexural Stress (MPa)

Additionally, Figure 3 demonstrates that the strain at peak stress varied between 1471 and 2903 microstrains for all specimens examined in this investigation. The results indicate that the R0 beam constructed from untreated RCA is brittle in comparison to conventional aggregate concrete (N0). The stresses at peak stress are 1471 v and 1942 microstrains for the TP and T0 beams, respectively, which are less than those for the N0 beam. This behaviour may be explained by the fact that these sorts of beams have larger failure loads than the N0 beam. The results reveal that the beams are quite stiff at failure. Due to the addition of PO fibre, the concrete compressive strain at failure of the TB beam is increased to 2567 microstrains. Flexural loads are absorbed in this example by bridging fibres following fracture formation. Following that, the bending moment is redistributed to all of the beam element's components. However, due to the failure mechanism of beams, the total minimal compressive strain is recorded in all beams when the design is reinforced with GFRP. The majority of beams broke in the flexural–shear mode due to the low modulus elasticity, brittle nature of the GFRP material, and inadequate shear resistance specified in the beam design. Thus, the beam's flexural–shear failure may occur as a result of the FRP rupturing, followed by the crushing of concrete, before the beam's flexural capacity can be fully developed. 22 20 18 16 14 12 10 8 6 4 2 0

N0 R0 T0 TB TP 0

250 500 750 10001250150017502000225025002750 Compressive Strain (micro strain)

Fig. 3. Average mortar stress-strain curves at different strain rates Summary The following conclusions can be drawn on the basis of the structural performance of the modified RAC: The coupling effect inclusion of the treated RCA and fibre slightly contributes to the enhancement of the first cracking load of the modified RAC beam compared with the unmodified RAC beam (R0). The ultimate load of all modified RAC beams, such as T1, TB and TP, is comparable to the control specimens. This result confirms that modifications in the production of RAC by the addition of treated RCA and introduction of fibre in the mix significantly enhanced the ultimate load behaviour of the RAC specimens. The effect of the modification in the capability of RAC mix to retain the load was negligible after the ultimate failure of the T1, TB and TP beams. This phenomenon may be attributed to the beam that was strongly governed by the fracture behaviour of the GFRP reinforcement at this moment. The inclusion of PO fibre results in high concrete compressive strain at the failure of the TB beam, the strain of which at peak stress is recorded as 2567 microstrains. Acknowledgement We would like to thank the Universiti Sains Malaysia Penang and University Teknologi MARA (UiTM) for providing us with research facilities and financial support throughout our research work.

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References [1] W.H. Kwan, M. Ramli, K.J. Kam, M.Z. Sulieman, Influence of the amount of recycled coarse aggregate in concrete design and durability properties, Construction and Building Materials 26(1) (2012) 565-573. [2] J. Pacheco, J. de Brito, J. Ferreira, D. Soares, Flexural load tests of full-scale recycled aggregates concrete structures, Construction and Building Materials 101 (2015) 65-71. [3] M. Arezoumandi, A. Smith, J.S. Volz, K.H. Khayat, An experimental study on flexural strength of reinforced concrete beams with 100% recycled concrete aggregate, Engineering Structures 88 (2015) 154-162. [4] D. Gao, L. Zhang, Flexural performance and evaluation method of steel fiber reinforced recycled coarse aggregate concrete, Construction and Building Materials 159 (2018) 126-136. [5] S. Seara-Paz, B. González-Fonteboa, F. Martínez-Abella, J. Eiras-López, Flexural performance of reinforced concrete beams made with recycled concrete coarse aggregate, Engineering Structures 156 (2018) 32-45. [6] W. Alnahhal, O. Aljidda, Flexural behavior of basalt fiber reinforced concrete beams with recycled concrete coarse aggregates, Construction and Building Materials 169 (2018) 165-178. [7] H. Dong, Y. Song, W. Cao, W. Sun, J. Zhang, Flexural bond behavior of reinforced recycled aggregate concrete, Construction and Building Materials 213 (2019) 514-527. [8] I. Maruyama, M. Sogo, T. Sogabe, R. Sato, K. Kawai, Flexural Properties of Reinforced Recycled Concrete Beams, in: E. Vázquez, C.F. Hendriks, G.M.T. Janssen (Eds.) International RILEM Conference on the Use of Recycled Materials in Buildings and Structures, RILEM Publications, Barcelona, Spain, 2004, pp. 526-535. [9] B. González-Fonteboa, F. Martínez-Abella, Shear strength of recycled concrete beams, Construction and Building Materials 21(4) (2007) 887-893. [10] G. Fathifazl, A.G. Razaqpur, O.B. Isgor, A. Abbas, B. Fournier, S. Foo, Flexural performance of steel-reinforced recycled concrete beams, ACI Structural Journal 106(6) (2009) 858-867. [11] S. Ismail, M. Ramli, Mechanical strength and drying shrinkage properties of concrete containing treated coarse recycled concrete aggregates, Construction and Building Materials 68(0) (2014) 726739. [12] B. Li, Y. Chi, L. Xu, Y. Shi, C. Li, Experimental investigation on the flexural behavior of steelpolypropylene hybrid fiber reinforced concrete, Construction and Building Materials 191 (2018) 8094. [13] M. Gümüş, A. Arslan, Effect of fiber type and content on the flexural behavior of high strength concrete beams with low reinforcement ratios, Structures 20 (2019) 1-10. [14] S. Ismail, M. Ramli, Effects of adding fibre on strength and permeability of recycled aggregate concrete containing treated coarse RCA, International Journal of Civil, Architectural, Structural and Construction Engineering 8(8) (2014) 890-896. [15] W.H. Kwan, M. Ramli, Indicative performance of fiber reinforced polymer (FRP) encased beam in flexure, Construction and Building Materials 48 (2013) 780-788. [16] G. Ma, Y. Huang, F. Aslani, T. Kim, Tensile and bonding behaviours of hybridized BFRP–steel bars as concrete reinforcement, Construction and Building Materials 201 (2019) 62-71.

Keyword Index A Additive Manufacturing Alkali Activator Arc Welding

49, 59 179 119

B Base Slab Binder Build Plate Bonding Buried Oxide Layer (BOx)

163 179 49 15

C CAD/CAM Characterization Clay Soil Coal Ash Coatings Concrete

59 139 163 179 59 163

G GFRP Pipes Globular Structure Groove Joint Adhesive Groove Wheel

171 81 171 103

H Hardness High Speed Twin Roll Caster High Strength Low Alloy Steel Hollow Blocks Hot Forging Process

129 81 119 179 95

I InP Based TFET IR Thermography

15 139

L D Dielectric Dissimilar Aluminium Alloys Dissimilar Weldments Dividing Of Dendrite

3 129 139 81

Laser Manufacturing Laser Metal Deposition (LMD) Low Carbon Steel Sheet LPG Sensor

75 59 111 23

M E Edge Drains Effect of Solvents Electrode Rod Diameter End Mill Process External Magnetic Field

163 37 111 87 119

F Fabry Perot Scattering Studies Feed Driving System Fibre Reinforced Concrete Finite Element Modeling Flexural Behaviour Friction Stir Forming Friction Stir Welding Functional Materials

3 87 189 95 189 67 67, 129, 139 67

Machine Tools Macro Examination Mechanical Joining Medium Carbon Steels Micro Examination Microstructure Model-Based Systems Engineering Molecular Semiconductor Multi-Segmental

87 111 67 153 111 129, 153 75 37 171

N Nematic Liquid Crystal Novel Technique to Shear Spot Weld Nozzle Nugget Size

3 111 103 111

198

Advances in Materials and Technologies

O Optical Fiber Sensor Optical Properties Optimization Organic Semiconductors

67 37 75, 119 37

P Pervious Photoluminescence Plastic-Metal Joint Polarizing Optical Microscopy Polymer Power Consumption Praseodymium Pre-Treatment Profile Projector Puddle Holding

163 37 75 3 3 87 23 49 111 103

Q Qform

95

R Recycled Aggregate Concrete Resistance Spot Welding RHA Ring Stiffness Test Rotary Friction Welding (RFW)

189 111 179 171 153

S Screening Semisolid Simulation SNCM8 SnO2 Spot Welding Parameters Spray Pyrolysis Stiffness Strength Sustainable Manufacturing Technologies

119 81 75 95 23 111 23 171 171 87

T Tensile Properties TFET Thin Film TiAl6V4

153 15 23 49

Titanium Tongue Toolpaths Twin Wheel Caster

67 171 59 103

U Uniform Doping (UD) Unitary Unsymmetrical Joint Upper Ball Joint UV-Vis Spectroscopy

15 171 153 95 37

W Welding Parameters Wire

111 103

Author Index A Alvarez, P. Aoyama, E.

59 87

J Jacobs, G. Joshi, M.

75 15

B Bagali, J. Balasubramanian, V. Balu, V. Bay, C. Beohar, A. Berges, J.M. Berroth, J.

111 153 119 49 15 75 75

C Chelot, D. Culilang, Z.S.

171 179

K Khosla, S. Krishnan, P.K. Kumar, A. Kumar, D.

3 23 37 119

M Mahr, A. Mani, S. Mathew, R. Mizuguchi, Y. Montealegre, M.A.

49 3 15 87 59

D Deepa, S. Dhake, P. Dhamotharakannan, T. Döpper, F.

23 15 153 49

163

F Febriansya, A.

163

G Ghosh, J.

15

81, 103 87 49

Ohashi, T. Okuda, N. Ortiz, I.

67 103 59

P Pathak, D. Pradhan, M.

119 3

Rai, P. Ramakrishna, M.V.A. Ramli, M.

3 139 189

S

I Ismail, S.

111 103 67

R

H Haga, T. Hirogaki, T. Hofmann, A.

Nanjundaradhya, N.V. Nishida, S. Nishihara, T.

O

E Endawati, J.

N

189

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