Laser Shock Peening: Fundamentals and Advances 9819911168, 9789819911165

Table of contents :
Preface
Contents
1 Characteristics and Development Status of Laser Shock Peening
1.1 The Concept and Connotation of Laser Shock Peening
1.2 Characteristics of Laser Shock Peening
1.3 Early Experimental Study on Laser Shock Peening
1.3.1 Five Developmental Stages of LSP
1.3.2 Research and Development of LSP in China
1.4 Development of Industrial Applications of Laser Shock Peening
1.4.1 Applications on Aero-Engines
1.4.2 Application on Aircraft Structure
1.4.3 Application on Weld Structure
1.5 New Application Direction of Laser Shock Peening
References
2 Laser Shock Hardening Industrial Application System
2.1 Laser in China and Abroad
2.1.1 Laser Jointly Developed by University of Science and Technology of China and Jiangsu University
2.1.2 Gaia High-Energy Equivalent Pumped YAG Laser of French THALES
2.1.3 Laser Scheme with Two-Way Laser Beam Output of Labao Company
2.1.4 Laser Developed by LSPT Company
2.1.5 Laser Being Developed by Japanese Company
2.2 Design Scheme of Laser of AVIC Manufacturing Technology Institute
2.2.1 Design Scheme of the Local Oscillation Laser
2.2.2 Design Scheme of Laser Amplifier
2.2.3 Program Analysis
2.2.4 Power Supply Based on IGBT Inverter Technology
2.3 Workbench of the Strengthening System
2.4 Beam Moving Scanning System Developed by MIC
References
3 Stability Factors and Safety Protection of Laser Shock Peening
3.1 Process Stability Factors
3.1.1 Adjustable Laser Spot and Continuous Laser Path
3.1.2 Flat Confinement Layer
3.1.3 The Integrity of the Absorption Layer
3.1.4 The Quality of the Target Material
3.2 Research on the Application of Confinement Layer
3.2.1 Introduction and Application of Water Confinement Layer
3.2.2 Laser Absorbance in Water and Selection of Restrained Layer Thickness
3.2.3 Influence of Water-Confined Layer on Shock Wave
3.2.4 Parasitic Plasma
3.2.5 Optical Path Purification
3.3 Status of Damaged Tape and Absorption Layer
3.3.1 No Absorption Layer
3.3.2 Slightly Damaged Absorption Layer
3.3.3 No Damaged Absorption Layer
3.4 Effect Mechanism of Target
3.5 Strengthen Effect Improvement and Safety Protection
3.5.1 Application of Spall Prevention Technology
3.5.2 Methods to Enhance Strengthening Effect
3.5.3 Light Reflection and Explosive Fragmentation of Safety Protection
References
4 Numerical Analysis of Mechanical Effects of Laser Shock Peening
4.1 Physical Model
4.1.1 Fabbro Physical Model
4.1.2 Modified Physical Model
4.2 Numerical Analysis Steps
4.2.1 Finite Element Analysis Method
4.2.2 Numerical Model Parameter Setting
4.3 Numerical Analysis of Circular Laser Spot
4.3.1 Finite Element Model
4.3.2 Dynamic Stress–Strain Analysis of Shock Wave Loading Process
4.3.3 Study of Residual Stress Field and Surface Plastic Deformation in the Laser Shock Area
4.3.4 Verification of the Residual Stress Field in the Laser Shock Zone
4.3.5 Relationship Between Surface Profile and Residual Stresses in the Single-Spot Impact Zone
4.3.6 Residual Stress Field of Lap-Spot
4.4 Square Spot Values Analysis
4.4.1 Finite Element Model
4.4.2 Shock Wave Loading
4.4.3 Residual Stress Distribution Under Different Process Parameters
References
5 Evaluations of the Strengthening Effect of the Metals with Laser Shock Peening
5.1 Surface Profiles Induced by Square Spots
5.1.1 Spot Overlapping Patterns
5.1.2 Surface Profiles Induced by Square Spots
5.1.3 Path Planning of Spot Overlapping
5.2 Mechanical Property of High-Temperature Alloy
5.2.1 Overseas Research Status
5.2.2 Effect of Thermal Cycles on Residual Stresses of High-Temperature GH2036 Alloy
5.2.3 Fatigue Lives of High-Temperature GH30 Alloy
5.2.4 Fatigue Crack Growth Rate of High-Temperature GH30 Alloy
5.3 Mechanical Property of Stainless Steel
5.3.1 Fatigue Lives of 1Cr18Ni9Ti Austenitic Stainless Steel
5.3.2 Fatigue Lives of 1Cr11Ni2W2MoV Stainless Steel
5.3.3 Plastic Deformations of Almen Samples (SE707 Stainless Steel)
5.4 Mechanical Property of Titanium Alloys
5.4.1 Mechanical Property of TC4 Titanium Alloy
5.4.2 Mechanical Property of TC17 Titanium Alloy
5.4.3 Mechanical Property of TC21 Titanium Alloy
5.4.4 Mechanical Property of TA19 Titanium Alloy
5.5 Mechanical Property of Aluminum Alloys
5.5.1 Fatigue Lives of 1420 Aluminum–Lithium Alloy
5.5.2 Fatigue Lives of 7050 Aluminum Alloy Fastening Holes
5.5.3 Fatigue Lives of LY12(2024)T62 Riveted Aluminum Alloy Sheets
5.5.4 Fatigue Crack Growth Rate of LY12 Aluminum Alloy
References
6 Strengthening Processes and Effect Evaluations of Airplane Structures with Laser Shock Peening
6.1 Applications of Laser Shock Peening Treatment on Airplane Structures
6.2 Requirements for Strengthening Processes of Aero-Engine Blades
6.3 Spall Characteristics of Blades and Its Prevention Spall Technology
6.3.1 Spall Strength and Spall Characteristics at the Bottom Surface of Thin Sheets
6.3.2 Spall Threshold and Spall Characteristics at the Bottom Layer of Mid-Thick Plates
6.3.3 Prevention Spall Technology of Blades
6.3.4 Prevention Spall Process and Its Applications for Structures
6.4 Evaluation of Strengthening Effect of Blades
6.4.1 Evaluation of Anti-FOD Fatigue Performance at the Edge of Blades
6.4.2 Anti-bending Deformation at the Edge of Blades
6.4.3 Anti-vibration Fatigue Performance of Blades
6.5 Blisk with Laser Shock Peening
6.5.1 Laser Shock Peening with Large Inclination Angle
6.5.2 Energy Compensation Method of Laser Oblique Incidence
6.6 Plastic Forming of Wing Panels with Large-Area Laser Shock Peening
6.6.1 Bending Deformation Types of Thin Sheets
6.6.2 Upper Limit Value of Process Parameters of Convex Bending Deformation of Mid-Thick Plates
6.6.3 Convex Bending Deformation and Mechanical Property of Mid-Thick Plates
References
7 Quality Control Technology of Structures with Laser Shock Peening
7.1 Present Situation of Detection Technology for Laser Shock Peening Quality
7.2 Natural Frequency Tests of Aero-Engine Blades with Laser Shock Peening
7.2.1 Test System
7.2.2 The Changes of Natural Frequency and Residual Stresses of Blades
7.2.3 Relation Between Impact Times and Natural Frequency
7.3 Laser Shock Peening Effect Characterized by Acoustic Signal and Plasma Plume
References
8 Strengthening Processes and Effect Evaluation of Welded Structures with Laser Shock Peening
8.1 Present Situation of Weldments with Laser Shock Peening at Home and Abroad
8.2 Tensile Strength and Fatigue Lives of Argon Arc Welded GH30 Alloy
8.2.1 Micro-hardness and Residual Stress
8.2.2 Tensile Strength and Fatigue Lives
8.2.3 Fatigue Fracture Analysis
8.3 Tensile Strength and Fatigue Lives of Plasma-Welded 1Cr18Ni9Ti Alloy
8.3.1 Micro-hardness and Residual Stress
8.3.2 Tensile Strength and Fatigue Lives
8.3.3 Fatigue Fracture Analysis
8.4 Mechanical Property and Fatigue Lives of TIG Welded TC4 Titanium Alloy with Multiple Impacts
8.4.1 Micro-hardness and Microstructure
8.4.2 Tensile Properties and Fatigue Lives
8.4.3 Fatigue Fracture Analysis
8.5 Mechanical Property and Fatigue Lives of Laser-Welded TC4 Sheets Treated by Laser Shock Peening with Double Sides and Different Sequences
8.5.1 Micro-hardness and Residual Stress
8.5.2 The Comparison of Median Fatigue Life
8.5.3 Fatigue Fracture Analysis
8.6 Mechanical Properties and Corrosion Properties of TA15 Electron Beam Welds
8.6.1 Micro-hardness
8.6.2 Corrosion Properties
8.6.3 Tensile Properties
8.6.4 Tensile Fracture Analysis
References

Citation preview

Shikun Zou Junfeng Wu Ziwei Cao Zhigang Che

Laser Shock Peening Fundamentals and Advances

Laser Shock Peening

Shikun Zou · Junfeng Wu · Ziwei Cao · Zhigang Che

Laser Shock Peening Fundamentals and Advances

Shikun Zou AVIC Manufacturing Technology Institute Beijing, China

Junfeng Wu AVIC Manufacturing Technology Institute Beijing, China

Ziwei Cao AVIC Manufacturing Technology Institute Beijing, China

Zhigang Che AVIC Manufacturing Technology Institute Beijing, China

ISBN 978-981-99-1116-5 ISBN 978-981-99-1117-2 (eBook) https://doi.org/10.1007/978-981-99-1117-2 Jointly published with National Defense Industry Press The print edition is not for sale in China (Mainland). Customers from China (Mainland) please order the print book from: National Defense Industry Press. © National Defense Industry Press 2023 This work is subject to copyright. All rights are solely and exclusively licensed by the Publisher, whether the whole or part of the material is concerned, specifically the rights of reprinting, reuse of illustrations, recitation, broadcasting, reproduction on microfilms or in any other physical way, and transmission or information storage and retrieval, electronic adaptation, computer software, or by similar or dissimilar methodology now known or hereafter developed. The use of general descriptive names, registered names, trademarks, service marks, etc. in this publication does not imply, even in the absence of a specific statement, that such names are exempt from the relevant protective laws and regulations and therefore free for general use. The publishers, the authors, and the editors are safe to assume that the advice and information in this book are believed to be true and accurate at the date of publication. Neither the publishers nor the authors or the editors give a warranty, expressed or implied, with respect to the material contained herein or for any errors or omissions that may have been made. The publishers remain neutral with regard to jurisdictional claims in published maps and institutional affiliations. This Springer imprint is published by the registered company Springer Nature Singapore Pte Ltd. The registered company address is: 152 Beach Road, #21-01/04 Gateway East, Singapore 189721, Singapore

Preface

The aviation industry of China coincides with unprecedented development opportunities, both passenger aircraft and engines listed as the national major special projects, we shoulder the historical mission of building powerful aeronautical manufacturing technology. The anti-fatigue manufacturing technology belongs to the core competitive power of aviation, and laser shock peening is significant for the anti-fatigue manufacturing of aero-engine and airplane body structures; it can also be used in petroleum, chemical, energy, medical, and other industries’ key equipment or key structure long-life design. The emphasis of this book is to describe the principle and application of laser shock peening, and to discuss the theory, method, and technology of laser shock peening on metal structure surfaces. This book focuses on laser shock peening with an absorption layer and 1064 nm wavelength laser, and briefly introduces the excimer laser of micro-scale laser shock processing; direct ablative impact and underwater impact are considered in the new direction. The content of this book aims to solve the problems in the application of laser shock peening, and to provide the readers with support in basic theory, process, and application technology, the processing materials involved include aluminum alloy, titanium alloy, super-alloy, high-strength steel, and other commonly used materials in the aviation field. This book is written on a series of research results of the author on improving the fatigue properties of metal parts by laser shock peening, the equipment scheme and process details are introduced in detail, and the self-developed equipment and technology for laser shock peening of integrated bladed disks are introduced in detail; some new methods and related technologies, such as process stability control, quality control, and test analysis, are put forward. The main contents of Chaps. 2, 3, 5, 7, and 8, and parts of Chaps. 1 and 6 are original works of the author. This book is a monograph on laser shock peening. It summarizes the latest achievements in the development and application of laser shock peening technology. The whole book is divided into nine chapters: the first chapter introduces the concept and connotation, development history, development status, and prospects of laser shock peening technology, the second chapter mainly introduces the intense pulse laser and its supporting equipment, and the third and fourth chapters summarize and analyze the constraint mode and dynamic model of laser shock peening technology. v

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Preface

In Chap. 5, the laser shock peening of typical materials and structures is introduced, and in Chap. 6, the micro-scale laser shock peening and its experimental research are discussed and analyzed. Chapters 7 and 8 discuss the quality control technology of laser shock processing and the testing and analysis technology after laser shock peening. In the process of writing this book, the author has benefited many times from the exchange of scholars, colleagues, and friends at home and abroad. In the field of laser device technology, Prof. Serebrycov from the Institute of Laser and Physics of Russia has provided enthusiastic guidance. In the field of non-absorbing layer laser peening, Prof. Sano Yuji from Toshiba of Japan has provided his detailed publishing information. Thanks to many many foreign friends. Since 1996, the author’s Institute, AVIC Manufacturing Technology Institute (formerly Beijing Institute of Aeronautical Manufacturing Technology Research Institute), began to research laser shock peening; we have maintained a good cooperative relationship with Prof. Wu Hongxing and his research team, who first carried out the research on laser shock peening in China, and we subsequently established extensive exchanges with Jiangsu University and Air Force Engineering University. The authors also actively contact with aero-engine and aircraft design institute and manufacturing plant or companies, in the study to solve practical problems in production, forming a unique chapter of this book. The research report of this book comes from research programs supported by National S&T breakthroughs, National Key R&D programs, National Natural Science Foundation of China, and other ministries and commissions; we thank their 20 years of support. The book-writing process was organized by Prof. Gong Shui and supported by National Defense Industry Press. It was co-edited by Zou Shikun, Cao Ziwen, Che Zhigang, Wu Junfeng, etc. and translated into English by Prof. Ye Chang and his team from HUST. In the late period, Wu Junfeng also quoted part of the contents of his doctoral thesis in the process of submitting the manuscript. In the process of writing, we have consulted the research reports at the previous International Conference on Laser Shock Peening, as well as the monographs, academic papers, dissertations, and network information at home and abroad; we would like to thank the authors and publishers of these studies. Since 2010, the author’s Institute has carried out the application of laser shock peening on engine blisks and airplane structures in China, and accumulated some engineering experience, but laser shock peening has not been applied for a long time in domestic engineering. The aviation standard, group company standard, and enterprise standard compiled by the author or others have not been widely used in the whole industry; there is no complete theory and production system, and limited by the author’s level and knowledge, the book is inevitably improper or even wrong, any criticism from the readers is welcome. Beijing, China December 2022

Shikun Zou

Contents

1 Characteristics and Development Status of Laser Shock Peening . . . . 1.1 The Concept and Connotation of Laser Shock Peening . . . . . . . . . . . 1.2 Characteristics of Laser Shock Peening . . . . . . . . . . . . . . . . . . . . . . . . 1.3 Early Experimental Study on Laser Shock Peening . . . . . . . . . . . . . . 1.3.1 Five Developmental Stages of LSP . . . . . . . . . . . . . . . . . . . . . 1.3.2 Research and Development of LSP in China . . . . . . . . . . . . . 1.4 Development of Industrial Applications of Laser Shock Peening . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1.4.1 Applications on Aero-Engines . . . . . . . . . . . . . . . . . . . . . . . . . 1.4.2 Application on Aircraft Structure . . . . . . . . . . . . . . . . . . . . . . . 1.4.3 Application on Weld Structure . . . . . . . . . . . . . . . . . . . . . . . . . 1.5 New Application Direction of Laser Shock Peening . . . . . . . . . . . . . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2 Laser Shock Hardening Industrial Application System . . . . . . . . . . . . . 2.1 Laser in China and Abroad . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.1.1 Laser Jointly Developed by University of Science and Technology of China and Jiangsu University . . . . . . . . . 2.1.2 Gaia High-Energy Equivalent Pumped YAG Laser of French THALES . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.1.3 Laser Scheme with Two-Way Laser Beam Output of Labao Company . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.1.4 Laser Developed by LSPT Company . . . . . . . . . . . . . . . . . . . . 2.1.5 Laser Being Developed by Japanese Company . . . . . . . . . . . 2.2 Design Scheme of Laser of AVIC Manufacturing Technology Institute . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.2.1 Design Scheme of the Local Oscillation Laser . . . . . . . . . . . . 2.2.2 Design Scheme of Laser Amplifier . . . . . . . . . . . . . . . . . . . . . 2.2.3 Program Analysis . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.2.4 Power Supply Based on IGBT Inverter Technology . . . . . . .

1 1 1 4 4 9 11 11 14 15 17 21 23 23 23 26 28 31 31 32 33 33 34 38

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2.3 Workbench of the Strengthening System . . . . . . . . . . . . . . . . . . . . . . . 2.4 Beam Moving Scanning System Developed by MIC . . . . . . . . . . . . . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

45 46 50

3 Stability Factors and Safety Protection of Laser Shock Peening . . . . . 3.1 Process Stability Factors . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.1.1 Adjustable Laser Spot and Continuous Laser Path . . . . . . . . 3.1.2 Flat Confinement Layer . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.1.3 The Integrity of the Absorption Layer . . . . . . . . . . . . . . . . . . . 3.1.4 The Quality of the Target Material . . . . . . . . . . . . . . . . . . . . . . 3.2 Research on the Application of Confinement Layer . . . . . . . . . . . . . . 3.2.1 Introduction and Application of Water Confinement Layer . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.2.2 Laser Absorbance in Water and Selection of Restrained Layer Thickness . . . . . . . . . . . . . . . . . . . . . . . . . 3.2.3 Influence of Water-Confined Layer on Shock Wave . . . . . . . 3.2.4 Parasitic Plasma . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.2.5 Optical Path Purification . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.3 Status of Damaged Tape and Absorption Layer . . . . . . . . . . . . . . . . . 3.3.1 No Absorption Layer . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.3.2 Slightly Damaged Absorption Layer . . . . . . . . . . . . . . . . . . . . 3.3.3 No Damaged Absorption Layer . . . . . . . . . . . . . . . . . . . . . . . . 3.4 Effect Mechanism of Target . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.5 Strengthen Effect Improvement and Safety Protection . . . . . . . . . . . . 3.5.1 Application of Spall Prevention Technology . . . . . . . . . . . . . 3.5.2 Methods to Enhance Strengthening Effect . . . . . . . . . . . . . . . 3.5.3 Light Reflection and Explosive Fragmentation of Safety Protection . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

51 51 52 55 59 62 63 63 64 66 66 69 69 70 71 72 76 77 77 82 83 84

4 Numerical Analysis of Mechanical Effects of Laser Shock Peening . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 87 4.1 Physical Model . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 88 4.1.1 Fabbro Physical Model . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 88 4.1.2 Modified Physical Model . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 90 4.2 Numerical Analysis Steps . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 92 4.2.1 Finite Element Analysis Method . . . . . . . . . . . . . . . . . . . . . . . 92 4.2.2 Numerical Model Parameter Setting . . . . . . . . . . . . . . . . . . . . 93 4.3 Numerical Analysis of Circular Laser Spot . . . . . . . . . . . . . . . . . . . . . 97 4.3.1 Finite Element Model . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 97 4.3.2 Dynamic Stress–Strain Analysis of Shock Wave Loading Process . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 99 4.3.3 Study of Residual Stress Field and Surface Plastic Deformation in the Laser Shock Area . . . . . . . . . . . . . . . . . . . 101 4.3.4 Verification of the Residual Stress Field in the Laser Shock Zone . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 103

Contents

4.3.5 Relationship Between Surface Profile and Residual Stresses in the Single-Spot Impact Zone . . . . . . . . . . . . . . . . . 4.3.6 Residual Stress Field of Lap-Spot . . . . . . . . . . . . . . . . . . . . . . 4.4 Square Spot Values Analysis . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.4.1 Finite Element Model . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.4.2 Shock Wave Loading . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.4.3 Residual Stress Distribution Under Different Process Parameters . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5 Evaluations of the Strengthening Effect of the Metals with Laser Shock Peening . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.1 Surface Profiles Induced by Square Spots . . . . . . . . . . . . . . . . . . . . . . 5.1.1 Spot Overlapping Patterns . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.1.2 Surface Profiles Induced by Square Spots . . . . . . . . . . . . . . . . 5.1.3 Path Planning of Spot Overlapping . . . . . . . . . . . . . . . . . . . . . 5.2 Mechanical Property of High-Temperature Alloy . . . . . . . . . . . . . . . . 5.2.1 Overseas Research Status . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.2.2 Effect of Thermal Cycles on Residual Stresses of High-Temperature GH2036 Alloy . . . . . . . . . . . . . . . . . . . . 5.2.3 Fatigue Lives of High-Temperature GH30 Alloy . . . . . . . . . . 5.2.4 Fatigue Crack Growth Rate of High-Temperature GH30 Alloy . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.3 Mechanical Property of Stainless Steel . . . . . . . . . . . . . . . . . . . . . . . . 5.3.1 Fatigue Lives of 1Cr18Ni9Ti Austenitic Stainless Steel . . . . 5.3.2 Fatigue Lives of 1Cr11Ni2W2MoV Stainless Steel . . . . . . . 5.3.3 Plastic Deformations of Almen Samples (SE707 Stainless Steel) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.4 Mechanical Property of Titanium Alloys . . . . . . . . . . . . . . . . . . . . . . . 5.4.1 Mechanical Property of TC4 Titanium Alloy . . . . . . . . . . . . . 5.4.2 Mechanical Property of TC17 Titanium Alloy . . . . . . . . . . . . 5.4.3 Mechanical Property of TC21 Titanium Alloy . . . . . . . . . . . . 5.4.4 Mechanical Property of TA19 Titanium Alloy . . . . . . . . . . . . 5.5 Mechanical Property of Aluminum Alloys . . . . . . . . . . . . . . . . . . . . . 5.5.1 Fatigue Lives of 1420 Aluminum–Lithium Alloy . . . . . . . . . 5.5.2 Fatigue Lives of 7050 Aluminum Alloy Fastening Holes . . . 5.5.3 Fatigue Lives of LY12(2024)T62 Riveted Aluminum Alloy Sheets . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.5.4 Fatigue Crack Growth Rate of LY12 Aluminum Alloy . . . . . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

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6 Strengthening Processes and Effect Evaluations of Airplane Structures with Laser Shock Peening . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.1 Applications of Laser Shock Peening Treatment on Airplane Structures . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.2 Requirements for Strengthening Processes of Aero-Engine Blades . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.3 Spall Characteristics of Blades and Its Prevention Spall Technology . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.3.1 Spall Strength and Spall Characteristics at the Bottom Surface of Thin Sheets . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.3.2 Spall Threshold and Spall Characteristics at the Bottom Layer of Mid-Thick Plates . . . . . . . . . . . . . . . . 6.3.3 Prevention Spall Technology of Blades . . . . . . . . . . . . . . . . . . 6.3.4 Prevention Spall Process and Its Applications for Structures . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.4 Evaluation of Strengthening Effect of Blades . . . . . . . . . . . . . . . . . . . 6.4.1 Evaluation of Anti-FOD Fatigue Performance at the Edge of Blades . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.4.2 Anti-bending Deformation at the Edge of Blades . . . . . . . . . 6.4.3 Anti-vibration Fatigue Performance of Blades . . . . . . . . . . . . 6.5 Blisk with Laser Shock Peening . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.5.1 Laser Shock Peening with Large Inclination Angle . . . . . . . . 6.5.2 Energy Compensation Method of Laser Oblique Incidence . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.6 Plastic Forming of Wing Panels with Large-Area Laser Shock Peening . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.6.1 Bending Deformation Types of Thin Sheets . . . . . . . . . . . . . . 6.6.2 Upper Limit Value of Process Parameters of Convex Bending Deformation of Mid-Thick Plates . . . . . . . . . . . . . . . 6.6.3 Convex Bending Deformation and Mechanical Property of Mid-Thick Plates . . . . . . . . . . . . . . . . . . . . . . . . . . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7 Quality Control Technology of Structures with Laser Shock Peening . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.1 Present Situation of Detection Technology for Laser Shock Peening Quality . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.2 Natural Frequency Tests of Aero-Engine Blades with Laser Shock Peening . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.2.1 Test System . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.2.2 The Changes of Natural Frequency and Residual Stresses of Blades . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.2.3 Relation Between Impact Times and Natural Frequency . . .

231 232 238 242 243 248 257 260 263 266 275 280 286 289 290 291 292 303 316 336 341 341 346 346 349 349

Contents

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7.3 Laser Shock Peening Effect Characterized by Acoustic Signal and Plasma Plume . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 351 References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 352 8 Strengthening Processes and Effect Evaluation of Welded Structures with Laser Shock Peening . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.1 Present Situation of Weldments with Laser Shock Peening at Home and Abroad . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.2 Tensile Strength and Fatigue Lives of Argon Arc Welded GH30 Alloy . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.2.1 Micro-hardness and Residual Stress . . . . . . . . . . . . . . . . . . . . 8.2.2 Tensile Strength and Fatigue Lives . . . . . . . . . . . . . . . . . . . . . . 8.2.3 Fatigue Fracture Analysis . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.3 Tensile Strength and Fatigue Lives of Plasma-Welded 1Cr18Ni9Ti Alloy . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.3.1 Micro-hardness and Residual Stress . . . . . . . . . . . . . . . . . . . . 8.3.2 Tensile Strength and Fatigue Lives . . . . . . . . . . . . . . . . . . . . . . 8.3.3 Fatigue Fracture Analysis . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.4 Mechanical Property and Fatigue Lives of TIG Welded TC4 Titanium Alloy with Multiple Impacts . . . . . . . . . . . . . . . . . . . . . . . . . 8.4.1 Micro-hardness and Microstructure . . . . . . . . . . . . . . . . . . . . . 8.4.2 Tensile Properties and Fatigue Lives . . . . . . . . . . . . . . . . . . . . 8.4.3 Fatigue Fracture Analysis . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.5 Mechanical Property and Fatigue Lives of Laser-Welded TC4 Sheets Treated by Laser Shock Peening with Double Sides and Different Sequences . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.5.1 Micro-hardness and Residual Stress . . . . . . . . . . . . . . . . . . . . 8.5.2 The Comparison of Median Fatigue Life . . . . . . . . . . . . . . . . 8.5.3 Fatigue Fracture Analysis . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.6 Mechanical Properties and Corrosion Properties of TA15 Electron Beam Welds . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.6.1 Micro-hardness . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.6.2 Corrosion Properties . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.6.3 Tensile Properties . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.6.4 Tensile Fracture Analysis . . . . . . . . . . . . . . . . . . . . . . . . . . . . . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

353 353 355 356 359 360 361 362 363 365 366 366 372 373

377 377 379 380 384 384 386 387 387 388

Chapter 1

Characteristics and Development Status of Laser Shock Peening

1.1 The Concept and Connotation of Laser Shock Peening The absorption layer on the metal surface absorbs the laser energy and explodes and evaporates, producing plasma with high temperature (>10000 K) and high pressure (>1 GPa) when the laser with short pulse (several to tens of nanoseconds) peak power density (>GW/cm2 ) radiates the metal target. And high-intensity pressure shock wave is generated and acted on a metal surface and propagates internally when the plasma is constrained by the confinement layer. When the peak stress of the shock wave exceeds the dynamic yield limit of the metal, strain hardening occurs, which leads to the formation of compressive residual stress (CRS) in the component surface. This new technology is called Laser Shock Processing (LSP). And it is also called Laser Shock Peening (Laser Peening, LSP, LP) as its principle is similar to Shot Peening (SP). Figure 1.1 shows the schematic of LSP and SP.

1.2 Characteristics of Laser Shock Peening LSP is one of the highest peak power techniques in laser processing by using shock wave to produce plastic deformation on a metal surface. The generated plasma is equivalent to a small explosion on the surface of the material, but the thermal effect is only in the absorption layer of a few microns deep due to the extremely short action time (nanosecond order). Therefore, LSP is a cold processing technology, which can obtain smooth micron-level depressions and millimeter-level CRS layers. However, in the early stage, the principle of LSP is mainly based on phase transition hardening caused by laser thermal, which is a thermal working process. Laser shock strengthening is a cold processing technology; similar technologies include SP, extrusion, rolling, ultrasonic peening, and so on.

© National Defense Industry Press 2023 S. Zou et al., Laser Shock Peening, https://doi.org/10.1007/978-981-99-1117-2_1

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1 Characteristics and Development Status of Laser Shock Peening

Fig. 1.1 Schematic of LSP and SP

(1) Shot peening SP is a commonly used technique, which worked by driving the shot with compressed air to impact the surface of the material and make the surface plastic deformation. Compared with SP, LSP is easier to accurately control the peening area, intensity, and overlapping ratio, but it cannot eliminate the surface tool marks. In addition, the CRS layer obtained by material after LSP can reach 1 mm, which is about 2–5 times the SP. (2) Ultrasonic peening Ultrasonic peening is more efficient; it is generally suitable for welded parts with low surface requirements, because it introduces a large surface roughness of the target surface. However, by selecting the appropriate spot and overlapping ratio, the blade after LSP can maintain the surface roughness of Ra 0.4 μm. (3) Extrusion Cold extrusion is a hole treatment process commonly used in aviation, which can substantially improve the fatigue life of structural parts. However, for small holes less than ϕ 2.5 mm, the small size mandrel is very easy to break during the extrusion process. Once the mandrel is broken, it is extremely difficult to remove it from the aircraft structure. In addition, mandrel and open seam bushings are expensive; each extruded mandrel cost up to more than two hundred dollars; small diameter open seam bushings cost 1–2 dollars each, and open seam bushings are expendable process bushings; each hole will consume an open seam bushing. (4) Rolling Rolling can only treat flat surfaces, which need to be well supported.

1.2 Characteristics of Laser Shock Peening

3

In short, the characteristics of LSP are as follows: (1) High strain rate Due to the short time of the shock wave (tens or hundreds of nanoseconds), the strain rate of LSP can reach more than 106 s−1 (even 107 s−1 with femtosecond laser). Compared with other high strain rate strengthening techniques, such as SP and explosion, the strain rate of LSP is 3–4 orders of magnitude higher. The transformation mechanism, property changes, and deformation mechanism on the microstructure of material are different, and brittle materials at conventional strain rates may also produce slip lines unique to plastic deformation under LSP. (2) Deeper strain-affected layers The pressure of laser shock waves can reach several GPa or even TPa magnitude, which is difficult to achieve with conventional mechanical processing, for example, mechanical stamping is often between several tens and hundreds of MPa. Compared with SP, LSP obtains a CRS layer of up to 1 mm, which is about 2–5 times that of SP, as shown in Fig. 1.2. (3) Surface roughness Due to the protective effect of the absorption coating, the thermal effect of the laser only acts on the surface layer of the coating material, while the target material is almost not affected by the thermal effect; LSP is actually a cold treatment. The plastic deformation depth of LSP is greater than SP, but the shock wave pressure and plastic deformation are more uniform, so the surface roughness of materials after LSP is less than SP, as shown in Fig. 1.3. LSP without foreign contamination, SP may exist shot contamination (such as titanium alloy cannot use steel shot to avoid iron contamination). Fig. 1.2 Comparison of CRS between LSP and SP

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1 Characteristics and Development Status of Laser Shock Peening

Fig. 1.3 Comparison of surface profiles between LSP and SP samples

(4) Impact area and pressure can be controlled, easy to automate The processing area of LSP is limited to the spot area; the pressure of the shock wave depends on the laser power density, so the impact area and pressure can be precisely controlled; the specimen can be locally reinforced, so it is suitable for automated processes. (5) Slower residual stress release The rate of residual stress release is proportional to the degree of cold work. LSP has only a small amount of cold work, because there are only one or a few deformation cycles, and its stress release process is caused by thermal force close to deep rolling. For example, titanium alloys can reach 425 °C and In718 can reach 670 °C.

1.3 Early Experimental Study on Laser Shock Peening 1.3.1 Five Developmental Stages of LSP Allan Clauer of the American LSP Technology Company summarized the development history of LSP at the first international conference on LSP and believed that laser shock technology can be divided into five stages [1]. The first stage is to discover the phenomenon, the second stage is to discover the potential application of the phenomenon, and the third stage is to define the laser shock peening technology. The fourth stage is to introduce the technology into an application, and the fifth stage is mass production. The first stage. In 1968, Anderholm found that laser-generated shock waves have an enhanced effect under plasma confinement, and showed that the peak power density of 12 ns and 1.9 GW/cm2 can produce a peak pressure of 3.4 GPa pulse. The second stage. It was discovered that shock waves can produce plastic deformation, and Mirkin at the Russian State University discovered the plastic deformation of 10 ns laser on metal materials.

1.3 Early Experimental Study on Laser Shock Peening

5

The CGE VD-640 Q-switched neodymium glass laser in Battelle’s Columbus laboratory in Ohio, USA, was first used in LSP experiments in 1970, as shown in Fig. 1.4. The main technological parameters of the laser are that it can output 200 J laser energy, 20–30 ns pulse width, and output every 8 min. But the laser is mainly used for laser fusion research (Table 1.1 and Fig. 1.5). In the fall of 1978, Ford, Fairand, Clauer, and Gilliher of the laboratory, in conjunction with the U.S. Air Force Flight Dynamics Laboratory, conducted a study on LSP to improve the fatigue life of fasteners. The experimental results showed that the fatigue life of the 7075-T6 specimen with a thickness of 6.35 mm did not improve, the fatigue life of the 7075-T6 specimen with a thickness of 3.175 mm was improved, and the fatigue life of 2024-T3 was substantially improved, as shown in Figs. 1.6. The residual stress measurements of the 7075 aluminum alloy showed that the surface of the specimen had a very high CRS after LSP. In 1979, William F. Bates Jr. of Lockheed-Georgia CoMPany Marietta (Georgia 30063), a well-known company in the U.S. defense industry, also conducted research on LSP of 7075-T6 and 7475T73 aluminum alloys. The tensile fatigue specimens

Fig. 1.4 The CGE VD-640 Q-switched neodymium glass laser

Table 1.1 Mechanical properties of 7075 aluminum alloy after LSP Heat treatment conditions

0.2% nominal yield strength/psi

Tensile strength/psia

Elongation %

Solid solution water quenching

36,700

66,000

35

Solid solution water quenching + LSP

43,500

66,000

31

T73 over-aging

50,000

74,600

18

T73 + LSP

64,600

85,200

13

T6

78,000

85,000

14

T6 + LSP

75,000

82,000

14

a1

psi = 6.895 kPa

Peak pressure GPa

3.1

4.7 3.1

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1 Characteristics and Development Status of Laser Shock Peening

Fig. 1.5 Tensile properties and microstructure of 7050 aluminum alloy strengthened by LSP

Fig. 1.6 Fatigue life of 2024T3 treated by LSP

were 6.35 mm thick, 38.1 mm wide, and 241.3 mm long, with fatigue holes of Φ 6.35 mm, and the specimens were coated with black paint. The specimens after LSP were subjected to a normal amplitude tensile fatigue test (maximum load of 6.895 MPa, stress ratio R = 0.1), and the experimental results showed that the fatigue life of 7075-T6 specimens was increased by 1.93 times and that of 7475T73 specimens was increased by 1.91 times. The third stage. The early 1980s defined this laser technology as LSP, the earliest laser dedicated to LSP was the LSP prototype established by Battelle’s Columbus Laboratory in 1986–1987, which could split two laser beams of 50 J each, 20 ns at 0.5 Hz, as shown in Fig. 1.7.

1.3 Early Experimental Study on Laser Shock Peening

7

Fig. 1.7 LSP facility at Battelle’s Columbus Laboratory

In the 1980s, France, Germany, Israel, Japan, Russia, and Italy carried out research on the application of LSP. The application field of LSP has been rapidly expanded. In 1987, Meunier et al. in France conducted a study on the impact hardening of aluminum alloys with a pulsed laser of GW/cm2 magnitude. In the same year, the Israeli Ministry of Atomic Energy funded Salimann et al. to conduct LSP studies on epoxy, carbon–carbon, and carbon-epoxy composites. In 1990, Grevey and Maiffredy, France, et al. used laser impact to induce martensitic phase transformation in TRIP alloy steel (containing 30% Ni) to estimate the depth of the deformation layer and explain the mechanism of martensitic phase transformation. French Forget, Strude, and Jeandin M. Lu, et al. used LSP to strengthen single and polycrystalline Ni-based high-strength alloys, and the authors concluded that LSP strengthened the material and induced compressive stresses within the material, which improved the fatigue life and wear fatigue resistance of the material. In 1992, Vaccri, USA, reported that the early pulsed laser systems were bulky, and although the output energy reached 500 J, the length of the optical path system reached 45.72 m, and the pulse repetition frequency was low, about once every 8 min. Today, laser equipment used for LSP has a form factor of 1.2 m × 1.8 m (length × width), an energy of about 100 J, at 1 Hz. LSP has replaced shot peening in improving the fatigue life of metal materials. LSP can effectively strengthen carbon steel, alloy steel, stainless steel, malleable cast iron, ductile iron, aluminum alloys, and nickelbased high-temperature alloys. The Wagner Institute of Laser Technology (Wagner Laser Technologies, WLT) has proven the reliability of LSP systems and has built a number of such processing workshops. In the twenty-first century, there has been a new breakthrough in the level of LSP research and a new expansion of application areas. In the period of 2000–2002, Lawrence Livermore State Key Laboratory in the United States continuously reported that (1) for 2024-T3 aluminum alloy after high-energy LSP, the fatigue life is 50 times that of conventional shot peening, and can be used to deal with the service life of many critical components, such as the blades of jet engines, F16 fighter jet bulkhead truss upper chord, and oblique end bar joints; (2) high-energy LSP can substantially reduce the stress corrosion and crack expansion of the weld. The U.S. nuclear waste is stored

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1 Characteristics and Development Status of Laser Shock Peening

in containers welded with 22-gauge alloy and buried under Yucca Mountain, which requires the containers to be preserved for 10,000 years without leakage. However, tensile residual stresses exist in the vessel’s welds, which lead to crack expansion and accelerated corrosion. High-energy LSP converts this residual tensile stress into residual compressive stress, thus preventing crack expansion. Livermore (Livermore) State Key Laboratory in the YMP research program, the following experiments were done: two pieces of 304 stainless steel with dimensions of 3 mm × 18 mm × 75 mm put plate type specimens in the temperature of more than 120 °C MgCl2 solution for corrosion resistance experiments, the specimen without LSP in 24 h cracks, after LSP of the specimen in 7 days after the crack or corrosion has not yet occurred. Similar experiments were conducted on U-shaped 304 stainless steel specimens, the untreated U-shaped specimens fractured within 2 h, and the LSPed U-shaped specimens did not show cracks after 6 days. Similar results were obtained after experiments on 316 types of stainless steel. As a result, Livermore SKL concluded that LSP can be used not only for the treatment of nuclear waste storage vessel welds but also for improving the safety and reliability of nuclear reactors and extending the operating time of reactor components (such as internal parts, shells, bolts, and pins), resulting in longer service life and lower operating costs for boiling water reactors and pressure water reactors. The fourth stage. In May 2002, MIC USA applied LSP to a high-value jet engine blade production line to improve fatigue life. It saves millions of dollars per month in aircraft maintenance costs, millions of dollars in parts replacement costs, and also ensures reliability over the entire life cycle. Currently, the U.S. expects to save more than $1 billion in costs for the processing of military fighter blades alone. In 2002, See et al. reported: LSP Technologies, Inc. (LSPT) and General Electric Aircraft Engines (GEAE) are working on three “U.S. Air Force Manufacturing Technology Engineering” programs. By increasing the repetition frequency of the laser by 2–3 times of 0.25 Hz, the production efficiency is further improved and the cost is reduced, which is expected to reduce the cost by 50–70%. In 2004, the United States LSP Technology Company and the United States Air Force Research Laboratory carried out the LSP remanufacturing research on the damaged blade of the F119 engine. The F119 engine was equipped with the latest fighter jet F/A-22 of the United States. The blade material was TC4, and the blade prefabricated crack length was 1.27 mm. The fatigue strength is also reduced from undamaged 586.1–689.5 MPa to 206.85 MPa, which is far lower than the design requirement of 379 MPa for blade use. The fatigue strength of the damaged blade increased to 413.7 MPa after LSP, which achieved great success. The fretting fatigue life was increased by more than 25 times after laser shock treatment on the root of the blade. In 2005, Warren [2] believed that since the duration of the laser shock wave was about 40 ns, it was difficult to measure the whole process, and a large-scale numerical simulation method should be adopted for demonstration and analysis. In 2006, Benxin proposed a self-closing loop model for LSP under a water confinement layer. In 2006, Cheng et al. of the School of Mechanical and Materials Engineering at Washington State University used laser shock waves to strengthen

1.3 Early Experimental Study on Laser Shock Peening

9

single-crystal silicon brittle materials [3], proposing a multiscale dislocation dynamics theory to elaborate the strengthening mechanism. In 2006, Sano et al. [4] investigated LSP of parts in the absence of a protective layer to improve the stress corrosion resistance of 314 stainless steel. In the case of laser impact strengthening to improve the stress corrosion resistance of 314 stainless steel, the results showed a substantial increase in stress corrosion resistance in the absence of ablation. In 2007, Hatamleh [5] used laser impact 7075T7351 vibration friction welded parts, the rate of crack expansion was substantially reduced, and fatigue life was substantially increased. In 2007, Breuer concluded that LSP can produce a CRS layer 4–5 times deeper than mechanical shot peening, which can substantially improve fatigue life. The fourth stage of LSP as defined by Allan Clauer, USA, was from the late 1980s to the early 1990s, bringing LSP to the application market. The LSP project team conducted a lot of promotional activities and carried out targeted experiments for user needs, accumulating the data for engineering applications. The fifth stage is to enter the batch application stage. In 1991, the U.S. Air Force took great interest in LSP and introduced the laser impact peening project team to GE’s Aero Engines Division, and sought to adopt it on the F101 fan blade of the B1-B bomber to improve the performance of the fan blade against foreign object damage. LSP in GE’s development history is shown in Fig. 1.8. The United States is the country with the fastest development and most successful application of LSP technology, but LSP has also developed into an international technology. In 1995, Toshiba began to use LSP to strengthen the welds of nuclear reactors; Fabbro et al. in France studied the physical model of laser shock intensification. In 1996, China started the research of LSP for aeronautical manufacturing technology. In 1999, Germany cooperated with Dong Company to carry out LSP research; laser-induced shock wave intensification was studied in Australia in 1999.

1.3.2 Research and Development of LSP in China With the development of advanced manufacturing technology, the service life of key structural parts of aviation equipment is more and more demanding. The eternal goal of advanced manufacturing technology is to reduce weight and increase efficiency. Focusing on the needs of the industry, AVIC has carried out in-depth research on the anti-fatigue manufacturing technology of LSP applied to the welding structure of metal materials, fuselage hole structure, key fatigue zone conversion R, etc. and is making due contributions to the anti-fatigue manufacturing of key components in a wider range of fields. Research and development of LSP in China are summarized as follows. In 1992, the Department of Physics, University of Science and Technology of China, successfully developed China’s first LSP device, which won the second prize, Science and Technology Progress Award of the Chinese Academy of Sciences in 1999.

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1 Characteristics and Development Status of Laser Shock Peening

F118-100

B

CFM56-5BP/7 B737/A320

F110-132

GenIV

F-16 Block 60

F110-100

GenIII F-16

F110-129 F-16 F101-102

GenI/II B-1B

Fig. 1.8 Development of LSP in GE

In 1996, AVIC Manufacturing Technology Institute and Beijing University of Aeronautics and Astronautics carried out pre-research on LSP technology. In 1998, Jiangsu University began to make use of the pit effect produced by laser shock to carry out the preliminary basic research on laser shock forming, explore the changes in the mechanical properties of materials, microstructure, and formability of sheet metal during laser shock forming, and conduct a preliminary feasibility study on complex and large-size sheet metal forming. In 2004, AVIC Manufacturing Technology Institute established an experimental platform for LSP, which adopted a Q-switched neodymium glass laser with large pulse energy, and began the study on LSP of titanium alloy blades for the first time in China. In 2005, Jiangsu University successfully developed the high-power laser shock forming system. The system was successfully developed under the support of the National Natural Science Foundation of China and the key discipline construction fund of Jiangsu University, with a total investment of 2 million yuan.

1.4 Development of Industrial Applications of Laser Shock Peening

11

In 2008, AVIC Manufacturing Technology Institute established a LSP experimental base in Yanliang, Xi’an, and began to study the LSP technology of stainless steel blades. In 2009, AVIC Manufacturing Technology Institute, Liming Aero Engine Co., LTD., Shenyang Engine Design and Research Institute, etc. jointly carried out the research on LSP process of integral blade disk, and then broke through several key technologies of LSP of titanium alloy integral blade disk of the aero-engine. In 2013, on the basis of early cooperation with Chengdu Aircraft Design and Research Institute and Chengdu Aircraft Industry (Group) Co., LTD., AVIC started to install LSP for aircraft structural parts. Subsequently, in view of the high concern in the industry for the implementation of new technology, the aviation industry standard for LSP was released [6].

1.4 Development of Industrial Applications of Laser Shock Peening Laser shock peening is a new surface peening technology developed in recent years and is now widely used in the United States for surface peening of key structural parts of aero-engines. This technology has become one of the mandatory technologies for advanced engine blade strengthening. Its application has greatly improved the fatigue life of components, and has a wide range of applications in aviation, aerospace, petroleum, nuclear power, automotive, and other fields, such as the application of fastening holes, rivet holes, nuclear reactor vessel pipe welds, and stir friction welding welds, which can improve the fatigue resistance of weak parts of the material [7].

1.4.1 Applications on Aero-Engines Under the high-speed rotor rotation and strong airflow, aero-engine blade bears many kinds of load like tensile, bending, and vibration, and its working conditions are extremely harsh. In such working conditions, the blade (especially the compressor blade in the engine intake end and front fan blade) is very easy to damage (usually called foreign object damage (Foreign Object Damage, FOD)) with the airflow in the foreign object (such as sand, stone, and birds) after the impact, therefore, the engine failure usually causes accidents. FOD of the engine is the most sensitive on the first few stages of the engine’s titanium alloy blade edge, and once the formation of a gap has not been found or revised, it may lead to a sharp reduction in fatigue strength. FOD or other causes of cracking will cause serious secondary damage, and then lead to engine failure, as shown in Fig. 1.9. Foreign object damage to the blades is the primary cause of unplanned engine replacement, and FOD has a significant impact on the cost of engine maintenance. The underlying cause of such damage is

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1 Characteristics and Development Status of Laser Shock Peening

Fig. 1.9 Engine accident caused by FOD

the local notch, deformation, or crack in the blade of the front and trailing edge will formed after the blade suffered a foreign impact, and then which will result in stress concentration or directly become a source of damage, a direct threat to the safety of the blade life. The main means to resist foreign body impact is to increase the edge thickness of the blade (including fan blades, propellers, etc.), but this method will be too big a price in the aerodynamic aspects. For the longer front fan blade or the first stage of the compressor blade, there is also the use of a damping class structure, although there is an effective suppressing in blade vibration, which cannot be very effective in protecting the blade once a foreign body intrusion happened [8]. In 1995, the United States studied the sensitivity of fan blades to FOD. Three methods of laser shock strengthening, untreated, and shot peening of fan blades were compared. The results showed that the fatigue strength of the damaged F101 blades after laser shock strengthening was close to or even higher than that of the undamaged and untreated blades. The notches of 1/4 inch (1 inch = 2.54 cm) of blades strengthened by laser shock are machined or electrodischarge machined. At present, only the United States has applied laser shock peening technology to the production and maintenance fields, and has achieved great economic benefits. In 1997, GEAE Company of the United States applied laser shock peening to the B-1B/F101 engine blade production line, reducing maintenance costs by 99 million dollars. In 2002, MIC Company of the United States applied laser shock peening to the blade production line, saving millions of dollars in aircraft maintenance costs and parts replacement costs every month, and then applied it to the F-16 fighter and the most advanced F-22 fighter, as shown in Fig. 1.10. In 2004, LSPT and Air Force Research Laboratory of the United States carried out research on laser shock peening and repair of damaged titanium alloy blades of the F119 engine on F/A-22. For damaged blades with microcracks and insufficient fatigue strength, after laser shock peening, the fatigue strength was 413.7 MPa, which fully met the design requirements of 379 MPa for blade use, and achieved great success. In addition, the fretting fatigue life of the blade wedge root strengthened by laser shock is at least 25 times longer. Laser shock peening technology in the United States has been widely used in the production of F119-PW-100 engine blisks and other components. LSP company also proposed a patent for strengthening the

1.4 Development of Industrial Applications of Laser Shock Peening

13

Fig. 1.10 Laser shock peening F22 fuselage structure

riveted structure of aircraft skin, and applied mobile laser equipment to strengthen the riveted rivet and its surroundings at the aircraft assembly site, which got an obvious effect [9]. The United States can save more than 1 billion dollars by only processing the blades of military fighter aircraft. In 2004, it was applied to the blade treatment of B777 civil aircraft. In 2005, the United States gradually extended the laser shock peening to the blade treatment of large steam turbines and water turbines, as well as the reduction of key components of oil pipelines and automobiles. However, there is no exact data for these applications. It is reported that the treatment of oil pipeline welds alone can achieve benefits of more than 1 billion dollars. At present, only the United States has industrialized laser shock peening in the production lines of B-1 bombers, F-16 fighter jets, F-22 fighter jets, and B777 civil aircraft, and has formulated the technical standard AMS2546 for laser shock peening (a new technical standard is currently being developed). In the United States, laser shock peening has been applied not only to military aircraft but also to B777 civil aircraft blade treatment. In 2005, it was gradually extended to the blade treatment of large steam turbines and water turbines, as well as the reduction and life extension of key auto parts. For example, it can prolong the fatigue life of the 200 kg automobile girder by 2 times, which allows the weight to be reduced by 20 kg. This means that 8 million automobile girders are processed every year, and 285 million liters of gasoline can be saved every year. There are two advantages, one is to reduce weight, and the other one is to lower manufacturing costs: Saving fuel and using less. American Metal Modification Corporation (MIC) was approved by the Federal Aviation Administration (FAA) as the designated laser shock peening technology maintenance service station in February 2003, and the Joint Aviation Administration (JAA) as the designated laser shot peening technology maintenance service station in November of the same year. This shows that the laser shock peening technology has been mature in the United States, but there is no paper on relevant process details,

14

1 Characteristics and Development Status of Laser Shock Peening

indicating that the technology is strictly confidential. In addition to the industrial application of laser shock technology reported in the United States, we also saw the industrial application report of laser shock technology used in nuclear equipment weld processing to improve corrosion resistance in Japan. Although France, Britain, Germany, Russia, Israel, Australia, and other countries have also carried out research on laser shock since the 1980s, they have not yet seen reports of industrial applications in these countries.

1.4.2 Application on Aircraft Structure The fatigue failure of structures (often initiated from the surface) is the main form of the normal damage to aerospace vehicles. In order to improve the target life of aerospace vehicles, shot peening, extrusion, impact strengthening, and other methods are often used to improve the surface fatigue performance of key parts (such as the stress concentration part of the bearing load and the contact surface subject to wear or erosion). In the fatigue failure of aircraft skin structures, more than 90% of them are caused by fretting wear of riveted parts. In other fastening holes, the fatigue failure caused by fretting wear also affects the target life of aircraft fuselage structures. Therefore, it is very important to strengthen the structure of aircraft fastening holes and rivet holes for the target life of aircraft. Figure 1.11 shows the rivet structure strengthened by laser shock. The fastening hole is a typical stress concentration structure on the aircraft, which is prone to crack under fatigue load, especially when the size is small (ϕ < 6 mm), the strengthening effect of shot peening and cold extrusion process for hole structure or blind hole is not ideal or difficult to achieve. As a new surface strengthening Fig. 1.11 Rivet structure strengthened by laser shock

1.4 Development of Industrial Applications of Laser Shock Peening

15

Fig. 1.12 Distribution of residual compressive stress around small hole

technology, laser shock strengthening has great advantages in strengthening small holes, special-shaped holes, blind holes, etc. Focusing the laser beam into an annular light spot impacts the area around the hole, and generates residual compressive stress on the strengthened surface and subsurface [10]. As shown in Fig. 1.12, the energy and shape of the laser spot are adjusted to meet different strengthening effects, the residual compressive stress distribution obtained by the strengthening method of three annular spots from the inside out is deeper and wider, and the fatigue performance is better. In addition, the latest research results of AVIC Manufacturing Technology Institute show that for 7050 aluminum alloy, laser shock peening of its surface before drilling can also significantly improve the fatigue performance of the hole. Another major advantage of laser shock peening for small hole structural strengthening is that it can meet the requirements of on-site strengthening and has good accessibility. In the future, aircraft fuselage structures and riveted structures will be gradually replaced by welded structures or integral structures to improve the fatigue performance of the fuselage and obtain weight reduction. However, the welded structure is also a weak link in fatigue performance. Especially, titanium alloys, aluminum alloys, and other materials widely used in aircraft structures have high requirements for the welding process. Even if new processes such as laser welding and friction stir welding are used, the fatigue performance and corrosion resistance are also reduced, and laser shock peening will be a good post-weld treatment process.

1.4.3 Application on Weld Structure The early research of laser shock peening mainly focused on aluminum alloy welded joints, but there has been no substantial application. The breakthrough in engineering application was the application of the United States in aero-engine blade strengthening. Toshiba Company of Japan has developed a unique set of equipment and processes in the maintenance of reactor pressure vessels and pipe welds in nuclear

16

1 Characteristics and Development Status of Laser Shock Peening

Fig. 1.13 Schematic diagram of laser shock strengthening welds and underwater shock experimental device of Toshiba Energy Center

power plants. Use small energy and small light spot (light spot φ 0.8 mm, energy 200 mJ, pulse width 8 ns) laser is used to strengthen the welds of nuclear reactor pressure vessel and pipe joints to improve the resistance to stress corrosion cracking of welds. Due to the space limitation in the nuclear reactor, the absorption layer cannot be arranged like the conventional laser shock strengthening. The company adopts the laser shock strengthening technology without the absorption layer and adopts the way of optical fiber conducting laser. In order to eliminate the tensile stress caused by laser ablation of the surface, the company adopted laser shock strengthening with a high overlap ratio of more than 1000% and a pulse density of 36 J/mm2 . Figure 1.13 shows the schematic diagram of laser shock strengthening welds and underwater impact test devices adopted by the energy center of Toshiba, Japan, which can strengthen the inner wall of the pipe with a diameter of 9.5 mm [11]. In addition, the storage of nuclear waste and prevention of leakage are also very important. A large amount of nuclear waste must be stored in special containers and sealed by welding. The American YMP project uses laser shock peening to strengthen the 22 alloy welds of nuclear waste storage containers. The depth of the residual compressive stress layer in the strengthening area exceeds 5 mm. The goal is to ensure that nuclear waste storage containers will not leak due to stress corrosion within 10,000 years. Nuclear power is the direction of vigorous development in China in the future, and laser shock peening technology will be vigorously applied in the nuclear industry. Friction stir welding is increasingly widely used in aluminum alloy structures due to its small heat-affected zone, small deformation, and good joint strength. However, the mechanical properties and residual stress of friction stir welding joints may cause brittle fracture, fatigue fracture, and stress corrosion damage and reduce the stability of the structure. The research results of NASA Johnson Space Center show that after laser shock peening, the yield strength and tensile strength of friction stir welded joints of aluminum alloy are significantly improved (the yield strength of friction stir welded joints of 2195 aluminum alloy is increased by 60%, and the tensile strength

1.5 New Application Direction of Laser Shock Peening

17

Fig. 1.14 Residual compressive stress distribution of FSW welded joint after laser shock strengthening

is increased by 11%, as shown in Fig. 1.14), and there is grain refinement in the impact zone. AVIC Manufacturing Technology Institute applied laser shock peening to the strengthening of welded joints of laser welding and electron beam welding, significantly improving the original welding stress distribution. Laser shock peening technology is expected to become the key technology to solve the problem of large dispersion of fatigue properties of high-energy beam welded joints. Compared with shot peening, laser shock peening has great advantages in welding joints. Taking aluminum alloy as an example, the cold work hardening degree of shot peening is 30–40%, and that of laser shock strengthening is 4–9% [12]. It can be seen from the Bauschinger effect that under cyclic load, the residual compressive stress generated by laser shock peening is more stable, as shown in Fig. 1.14. In addition, laser shock peening can produce a surface quality very close to the welded joint, which is conducive to fatigue performance.

1.5 New Application Direction of Laser Shock Peening The new application directions of laser shock strengthening are as follows. 1. Laser shock forming When the thin-walled structure is strengthened by single-sided laser shock peening, the thin-walled structure will bend and deform to the side of the non-strengthened surface. Both surfaces are in a state of compressive stress. The thin-walled structure

18

1 Characteristics and Development Status of Laser Shock Peening

Fig. 1.15 Laser shock forming of aluminum alloy

can be formed by controlling the impact parameters. This technology is called laser shock forming. Figure 1.15 shows the laser shock forming of aluminum alloy. Fig. 1.16 shows the laser shock forming equipment, which has three unique advantages: (1) The composite process of shock wave strengthening and forming, which adopts optimized process parameters and paths to form high-amplitude residual compressive stress in the stress concentration area of structural parts, which can significantly improve the fatigue life. For example, the residual compressive stress of—235 to −300 MPa is formed on the surface of 2024 aluminum alloy with a thickness of 3 mm after impact forming. According to the results of our laser shock strengthening experiment, this state can increase fatigue life by more than 4 times. This is especially suitable for manufacturing sheet metal parts with antifatigue performance requirements, such as aircraft wing skin, which can be reduced by conventional strengthening processes. (2) High efficiency. The outer skin of the wing with a length of 40 m can be precisely formed by only 1–2 times of processing. (3) With accurate forming, low cost, and fast speed, it is especially suitable for the development of new products with small batches. If we only estimate the profiling of laser shock female die, we can save half of the die. This alone can save hundreds of millions of RMB for the research of a new model. The application potential of laser shock forming has the following aspects.

1.5 New Application Direction of Laser Shock Peening

19

Fig. 1.16 Laser shock forming equipment

1. Parts can be directly formed to be assembled (1) (2) (3) (4)

Smaller curvature radius is formed on a large thickness. Precision forming, no need to re-process. Better surface. Residual compressive stress on both sides of the forming surface.

2. Possibly better economic performance than shot peening (1) (2) (3) (4)

Pre-punching is allowed for precision forming. Precision molding reduces assembly time. Pre-punching can reduce fatigue and fretting wear. The residual compressive stress in a specific area can reduce the weight of the structure.

The integral wing panel has a large structure, complex profile, and stiffeners inside the panel, so the forming of the wing panel has become a major problem in aircraft manufacturing in China. The integral panel of the ARJ21 wing is formed by shot peening, but compared with shot peening, laser shock forming has larger curvature, deeper residual compressive stress, and easier control of forming parameters. Therefore, laser shock forming will be an alternative technology for shot peening forming, and it has a huge potential application prospect, for example, it can be used in an aircraft canopy (about 1 m2 , 2024 aluminum alloy, high cost of stamping die), fuel tank cap (about 0.5 m2 , small curvature forming), and wing panel (about 10 m2 ). MIC Corporation of the United States used laser shock forming for the forming of B747 wing thick wall panels. The equipment uses dual optical path transmission, with the underground optical path up to 45 m [10], and completed its first flight in 2010, as shown in Fig. 1.17. With the development of laser shock forming technology, it will be widely used.

20

1 Characteristics and Development Status of Laser Shock Peening

Fig. 1.17 Laser shock forming for B747 wing

2. Petrochemical industry The transmission pipeline of oil and natural gas is an important facility involving life, and its welding area is vulnerable to stress corrosion. Once corrosion leakage occurs, it will not only cause a huge waste of energy but also pollute the environment and may cause ecological damage. Laser shock peening can effectively improve the stress corrosion resistance fatigue life of pipelines. It is estimated that this application is expected to generate billions of dollars in economic benefits. 3. Marine vessel Submarines are prone to corrosion under the action of seawater for a long time, especially for pipe/plate weldments and hull welds. Therefore, laser shock peening will play an important role. The corrosion resistance of offshore aircraft is required to be higher, which is particularly valuable for the laser shock peening of these aircraft sheet metal parts. 4. Medical industry At present, the fretting fatigue life of implants used in human medical treatment, such as orthopedic inserts, ridge and knee substitutes, and bone fixators, can be effectively improved by laser shock strengthening. 5. Nuclear industry The hydrogen corrosion resistance of nuclear vessel welds can be effectively improved by laser shock peening. A small amount of nuclear waste can be stored in heavy water, and a large amount of nuclear waste must be stored in special containers, welded and sealed, and then buried deep in caves. This requires that no stress corrosion leakage occurs in welds for 10,000 years to ensure safety. Laser shock processing of welds can meet this demanding requirement and ensure the environment to avoid nuclear pollution. At present, the United States has begun to use this treatment technology. Nuclear power in China is the direction of great development in the future, which also indicates that there is a great application space for laser shock peening of stress corrosion resistance.

References

21

References 1. Clauer A (2008) How did we get here? A historical perspective of laser peening, Houston, Texas, USA 2. Zhang X (2010) Variations of structure and hardness of 12Cr2 Ni4 A steel induced by laser shock (in Chinese). Northwestern University 3. Cheng GJ, Shehadeh MA (2006) Multiscale dislocation dynamics analyses of laser shock peening in silicon single crystals. Int J Plast 22(12):2171–2194 4. Sano Y, Obata M, Kubo T et al (2006) Retardation of crack initiation and growth in austenitic stainless steels by laser peening without protective coating. Mater Sci Eng, A 417(1):334–340 5. Hatamleh O, Lyons J, Forman R (2007) Laser and shot peening effects on fatigue crack growth in friction stir welded 7075-T7351 aluminum alloy joints. Int J Fatigue 29(3):421–434 6. Zou SK (2013) Laser shock peening: a new generation of anti-fatigue surface strengthening technology (in Chinese). China Aviation News 7. Cao ZW, Zou SK, Gong SL (2010) The latest movement and development trend of laser shock processing (in Chinese). Aeronaut Manuf Technol 05:40–42 8. Wang J, Zou SK, Tan YS (2005) Application of laser shock processing on turbine engines (in Chinese). Appl Laser 25(01):32–34 9. Che ZG, Shi YN, Tang N et al (2013) Applications of plasma induced by laser shock on surface treatment (in Chinese). Appl Laser 33(04):465–468 10. Zou SK, Cao ZW, Zhao Y et al (2008) Laser peening of aluminum alloy 7050 with fastener holes. Chin Opt Lett 6(2):116–119 11. Sano Y, Adachi T, Akita K et al (2007) Enhancement of surface property by low-energy laser peening without protective coating. Key Eng Mater 345–346:1589–1592 12. Hatamleh O, Lyons J, Forman R (2010) Laser peening and shot peening effects on fatigue life and surface roughness of friction stir welded 7075-T7351 aluminum. Fatigue Fract Eng Mater Struct 30(2):115–130

Chapter 2

Laser Shock Hardening Industrial Application System

2.1 Laser in China and Abroad Laser shock peening requires high peak power of laser output, and there should be a power density of GW/cm2 on a large enough spot. Almost all of the first researches on laser shock peening are Q-switched Nd glass lasers, most of which are used to explore frontier physical basic problems, such as laser induced nuclear fusion (laser ignition). These lasers have low output frequency and high operating cost, it cannot be used in engineering production applications such as laser shock peening. Table 2.1 lists the main development history of lasers used for laser shock peening. Qswitched Nd glass lasers can meet the requirements of laser shock peening pulse energy very early, but only when the frequency reaches above 1 Hz can they really have application value. A production laser for laser shock strengthening, developed by Lawrence Livermore National Laboratory and LSPT Company, has an output pulse energy of 50 J, a pulse width of 20 ns, and a repetition rate of 1.25 Hz. However, the laser is large and expensive, about 550,000–700,000 dollars, and it is still a high-tech product prohibited from exporting to China. In 2004, AVIC Manufacturing Technology Institute and Jiangsu University successfully developed lasers of Q-switched Nd glass with large pulse energy, as shown in Fig. 2.1, which broke the foreign technical blockade and started the research on laser shock strengthening technology and equipment.

2.1.1 Laser Jointly Developed by University of Science and Technology of China and Jiangsu University Generally, the indicators of the laser oscillator in the laser device (such as mode, beam divergence angle, monochromaticity, pulse width, modulation performance, etc.) are good, but the output energy is low (only mJ) because the output energy is © National Defense Industry Press 2023 S. Zou et al., Laser Shock Peening, https://doi.org/10.1007/978-981-99-1117-2_2

23

24

2 Laser Shock Hardening Industrial Application System

Table 2.1 Laser for laser shock peening Year

Laser type

Pulse energy/J

Pulse width/(ns)

Frequency/Hz

Affiliated unit

1980

Nd glass

40–100

3–30

0.001

LLNL, USA

1989

Nd glass

40

7–20

0.01

LLNL, USA

1990

Nd glass

400

20

0.25

LLNL, USA

1996

Nd glass

100

20

10

LLNL, USA

Adopt solid heat capacity technology 500 J (short time) 1997

20 (short time) LLNL, USA New pumping mode and gain medium 200 (short time)

LLNL, USA

CLFA excimer

5

45

5

French Laser Research Institute

YAG excimer

3

5–9

10

LABest, China

2004

Nd glass

60

30

0.1

AVIC Manufacturing Technology Institute, China

2008

YAG

10

10–15

5

LABest, China

20

10–15

1

2010

YAG

15

15

10

AVIC Manufacturing Technology Institute, China

Fig. 2.1 Q-switched Nd glass laser developed by AVIC Manufacturing Technology Institute and Jiangsu University in 2004 a AVIC Manufacturing Technology Institute; b Jiangsu University

2.1 Laser in China and Abroad

25

Fig. 2.2 Overall optical path layout of high-power laser device

contradictory to the indicators [1]. In order to obtain good performance and high energy, enough amplifiers must be added. However, as a result, the volume of the entire laser device must be huge [2]. In order to reduce the volume of the whole laser device and increase the output energy of the laser oscillator, the number of stages of the laser amplifier can be reduced. Therefore, a multi-transverse mode laser oscillator is used [3]. Because the shock processing laser device has low requirements on various indicators, especially the basic transverse mode is not required. It is unnecessary to add small holes to select transverse modes so that the energy of the laser oscillation stage can be larger. Figure 2.2 shows the overall optical path diagram of the high repetition rate Nd glass high-power laser processing device [4–8] using a multi-transverse mode laser oscillator as the laser oscillation stage. The laser oscillator uses ∅ 8 mm × 200 mm phosphate N21 medium, the resonator is flat–flat cavity, and the output mirror is K9 flat glass. Since transverse modes are not selected through small holes, the laser rod can be larger and longer. After the resonator is adjusted, multiple transverse modes are output, and the energy can reach 2 J. Parameters of laser rods at all levels in Fig. 2.2 are as follows: (1) Laser medium in Q-switched laser oscillator (PA0 ) is ∅ 8 mm × 200 mm phosphate N21 neodymium glass rod, ∅ 12 mm × 200 mm pulsed xenon lamp double lamp pumping, the focusing cavity is a ceramic diffuse reflection cavity, and the luminous aperture of the electro-optic Q-switch is ∅ 15 mm, with matching liquid inside; (2) The laser medium in PA1 is ∅ 14 mm × 350 mm phosphate N21 neodymium glass rod, ∅ 22 mm × 350 mm pulsed xenon lamp double lamp pump, and the focusing cavity is a double elliptical metal copper silver plated cavity; (3) The laser medium in the 1st stage laser main amplifier (MAI-1 ) of circuit I is ∅ 16 mm × 350 mm phosphate N21 neodymium glass rod, ∅ 22 mm × 350 mm pulsed xenon lamp double lamp pump, and the focusing cavity is a double elliptical metal copper silver plated cavity;

26

2 Laser Shock Hardening Industrial Application System

(4) The laser medium in the primary amplifier (MAI-2 ) of channel I stage 2 is ∅ 20 mm (or ∅ 18 mm) × 350 mm phosphate N21 neodymium glass rod, ∅ 22 mm × 350 mm pulsed xenon lamp double lamp pump, and the focusing cavity is a double elliptical metal copper silver plated cavity; (5) The parameters of the II channel primary laser amplifier (MAII-1 ) are consistent with those of the II channel secondary laser amplifier (MAII-2 ). The charging voltage settings of all levels of power supply in Fig. 2.2 are as follows: The charging voltage of the Q-switched laser oscillator power supply is 1680 V, that of the preamplifier power supply is 1900 V, and that of the laser main amplifier power supply (the four power supplies are the same) is 2000 V. The laser energy (provided by the capacitor) includes: the capacitance/lamp of the laser oscillator energy is 200 μF, the capacitance/lamp of the laser preamplifier energy is 300 μF, and the capacitance/lamp of the laser main amplifier energy is 600 μF. When the circulating water cooling system starts, the temperature is 25 °C, the temperature difference is 2 °C, and the passive isolator Cr4+ YAG (transmissivity is about 70%) is added to the oscillation stage. The average output laser pulse energy of the laser is E = 42.23 J. When the laser operates at a repetition frequency of 0.5 Hz, the output laser energy decreases by 6.2%; The thermal lens effect of the medium is serious when the laser operates at 0.5 Hz, which hinders the normal operation of the laser. The experiment shows that the laser can only work normally for about 1 min when the above pump energy is used for operation.

2.1.2 Gaia High-Energy Equivalent Pumped YAG Laser of French THALES The overall dimensions of Gaia high-energy and other pumped YAG lasers of THALES company in France are shown in Fig. 2.3. Their structures are divided into oscillators, isolators, spatial filters, beam shaping devices, high-energy amplifiers, frequency doubling optical options, and overall dimensions. Gaia is one of the most efficient and compact high-energy pumping lasers in the market, with pulse energy up to 6 J@532 nm, 10 Hz. It has the following characteristics: compact design (less than 1/2 of the volume of similar lasers), high stability, simple maintenance, and square beam cross-section, which improve the coverage efficiency during material processing. Its main advantages are as follows. The output pulse energy of this product is high, which can reach 13 J or higher. The pulse is short, the pulse width is less than 15 ns, and the energy density before focusing can reach 2.6 J/cm2 . It can effectively generate plasma shock wave, thus meeting the process requirements of laser strengthening.

2.1 Laser in China and Abroad

27

Fig. 2.3 Overall dimensions of Gaia high energy isopumped YAG laser

The repetition rate is high, up to 10 Hz, which effectively ensures the speed of laser processing and meets the mass production requirements of production processing to the maximum extent. This product innovatively uses YAG ceramic as the laser crystal, which not only realizes the high energy pulse output but also ensures the stable operation of the laser with its excellent heat resistance and shock resistance. At the same time, its patented geometric design can maintain a good pattern and plane distribution when outputting high-energy pulses. In contrast, the high-energy and high repetition rate laser with YAG crystal rod is easy to produce “quenching” phenomenon, and the stability of the system is difficult to ensure. The laser system using phosphate laser neodymium glass is much more expensive than the system using YAG ceramics. According to the characteristics and requirements of laser processing, the internal structure and materials of the laser are improved to meet the requirements of the laser gun. For example, samarium doped quartz filter is used to ensure that the laser can output more than 100 million pulses. Uniform neodymium ion doping reduces the low-frequency modulation, thus optimizing the laser output spot characteristics. The completely symmetrical cavity design avoids the birefringence effect and improves the beam quality. The output laser pulse of this product adopts the unique flat top homogenization technology, and its pulse energy is evenly distributed in the whole spot, rather than the common Gaussian distribution. As shown in Fig. 2.4, the flat top square spot is conducive to the consistency of product processing in laser strengthening. The laser pulse output of this product can be selected as a square spot. In addition to the flat top characteristics of the beam, it can effectively improve the coverage efficiency of the laser beam, without the need for partial overlapping irradiation of the spot, which is conducive to improving the speed of laser enhanced impact and meeting the needs of industrial production. This product adopts modular design and is easy to maintain. It is completed by using seed source plus amplification stage. It has a simple structure and is easy for customers to maintain.

28

2 Laser Shock Hardening Industrial Application System

Fig. 2.4 Square output spot with flat top modulation

The application of this product: (1) high-energy short pulse laser shock; (2) material handling; (3) cleaning of nuclear waste; (4) scientific research; (5) pump source of multi hertz beat watt (PW) femtosecond laser system. THALES company in France can produce lamp-pumped lasers (100 [email protected] Hz and 30 [email protected] Hz) and diode-pumped lasers (170 W@10 kHz), as shown in Fig. 2.5. The company has provided two lasers (12 J and 22 J respectively) for Jiangsu University and Xi’an Aeroengine (Group) Co., Ltd. Amplitude company in France has been able to produce the laser with 56 J@532 nm 10 Hz in 2016, as shown in Fig. 2.6, the next step is to produce P60 laser (60 J@532 nm 10 Hz), and these are high-energy lasers and are prohibited products.

2.1.3 Laser Scheme with Two-Way Laser Beam Output of Labao Company The laser with two-way laser beam output designed by Labao is shown in Fig. 2.7, and the design of the laser with two-way laser output optical path is shown in Fig. 2.8. Laser output parameters: wavelength: 1064 nm; repetition frequency: 1–10 Hz; pulse width: 10–20 ns; laser mode: TEM00; coupling with mutually perpendicular polarized light; Single pulse energy: 10 J@10 Hz, 20 J@1–5 Hz; The energy stability is less than 5%. In order to meet the high energy demand of the laser beam for engineering applications, Labao will couple two laser beams to obtain high-energy laser beams. Figure 2.9 shows the design of two optical coupling paths, and Fig. 2.10 shows the

2.1 Laser in China and Abroad

Fig. 2.5 THALES laser

Fig. 2.6 Next generation P60 laser of Amplitude company in France

29

30

2 Laser Shock Hardening Industrial Application System

Fig. 2.7 Laser shape of two channels of laser output of Labao Company

Fig. 2.8 Optical path design of two-way laser outputs of Labao

Fig. 2.9 Two-way optical coupling design

impact on the pulse width after the coupling of two optical paths. After the coupling of the optical paths, the pulse width time increases and is flat topped, and the laser shock strengthening effect has been improved.

2.1 Laser in China and Abroad

31

Fig. 2.10 Effect on pulse width after coupling of two optical paths

Fig. 2.11 LSPT laser shock peening system introduced by Guangdong University of Technology

2.1.4 Laser Developed by LSPT Company The diode pumped nanosecond pulse laser recently developed by LSPT Company of the United States has a pulse energy of 10 J, an adjustable pulse width of 10– 25 ns, a compact overall design, SLM and TEM00 local oscillator design, a flat top pulse energy spatial distribution (energy fluctuation less than 9.1%), a pulse width variation of only 0.07 ns within 8 h, and an energy fluctuation less than 0.28%. Guangdong University of Technology has introduced this product, as shown in Fig. 2.11. In Weifang, Shandong, LSPT reached an agreement with Shandong Myrtle Metal Surface Technology Co., Ltd. (MTLS), which invested 150 million yuan to establish a processing unit including laser shock processing unit to introduce this product.

2.1.5 Laser Being Developed by Japanese Company Hamamatsu Photonics K. K. company of Japan is developing a 100 J laser for laser shock peening, as shown in Fig. 2.12. In view of the control technology of laser pulse shape, it is shown that small pulses near the front of the main pulse are conducive to improving the depth of the residual compressive stress layer on the strengthened surface. Its mechanism is that small pulses (Foot Pulse) generate low-density plasma

32

2 Laser Shock Hardening Industrial Application System

Fig. 2.12 Laser for 100 J laser shock peening

before the main pulse acts with the base material, In this way, most of the energy of the main pulse is used to convert into the internal energy of the plasma, and it is unnecessary to use part of the pulse width of the rising edge of the main pulse to form the initial plasma.

2.2 Design Scheme of Laser of AVIC Manufacturing Technology Institute In order to obtain the laser with high energy, good stability, and good beam quality, the existing technical means must be amplified by an amplifier. The energy level structure of the laser amplifier matches the signal light of the local oscillator, and the physical process of the laser amplifier is the same as that of the laser oscillator. Under the action of optical pumping, the working substance is in a state of particle number reversal. When the laser pulse generated from the local vibration passes through, the particles in the excited state produce strong excited radiation under the action of the external optical signal. The radiation is superimposed on the signal light and amplified, so the amplifier can output much stronger signal light radiation. In order to ensure the stability of the system and the quality of the laser output beam, the local oscillator laser uses laser diode pump and unsteady cavity mode selection to provide a high quality and stable light source for the amplifier stage. The local oscillator output laser is divided into two beams after passing through the spectroscope, which are respectively amplified by the laser amplifier. The laser amplifier adopts the six-level two-way amplification system. Beam couplers and optical rotators are added between the local oscillator laser and the amplifying stage, and between the amplifying stage and the amplifying stage to suppress the thermal lens effect, self-focusing effect and phase delay, and other harmful factors. In addition, the phase conjugated mirror compensation technique is used in the system to reduce the influence of the inhomogeneity of the optical medium and impurities on the laser beam quality.

2.2 Design Scheme of Laser of AVIC Manufacturing Technology Institute Prism

Prism

Polarizer

Nd: YAG

33 Output Mirror

Light wedge

Diaphragm

Q Switch

F-P

allreverse mirror

Fig. 2.13 Schematic diagram of the local oscillation laser

2.2.1 Design Scheme of the Local Oscillation Laser The use of laser diode pumping technology can effectively improve the output beam quality, stability, and reliability of the local vibration. The optical path of the local oscillation laser, as shown in Fig. 2.13. Prisms are used to “fold” the resonant cavity to reduce the longitudinal dimension and improve the stability of the optical path. The F-P standard is placed inside the resonant cavity and is used for spectral selection of the resonant cavity. The modulation part is proposed to use passive Q-switching. The light wedge pair can make the optical path adjustment more convenient, and at the same time can improve the reliability of the system.

2.2.2 Design Scheme of Laser Amplifier Figure 2.14 shows the laser optical path design, where the laser working substance is Nd:YAG crystal. The pump source uses a flash lamp pump; the amplification stage consists of four laser bars in series, two of which are 10 mm in diameter and the remaining two are 15 mm in diameter. The amplification stage is a two-way amplification system with a phase conjugate mirror (PCM). PCM is based on the excited Brillouin scattering effect. The amplification stage is isolated by a vacuum filter, which completes the function of matching the aperture of the laser bar, transferring the image plane and partially compensating the thermal lens effect of the laser bar; A 90° quartz rotator is placed between two rods of the same diameter to compensate for thermogenic birefringence. Optical isolation between the amplifier stage and the PCM by placing a Faraday rotator between the MO and the amplifier stage. The

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2 Laser Shock Hardening Industrial Application System

Fig. 2.14 Laser optical path design. Laser principal oscillator; 2—F isolator; 3—beam shaping system; 4—polarizer; 5—reflector; 6—laser bar; 7—90° rotator; 8—vacuum filter; 9—laser bar; 10—45° Faraday rotator; 11—optical phase conjugation system; 12—laser bar; 13—coupling system

amplification stage is a dual-range amplification system with phase conjugate mirror (PCM), the first four stages are dual-range amplification consisting of four laser rods in series, two rods of 10 mm diameter, and two rods of 15 mm diameter. The last 2 stages are direct discharge type, two rods of 20 mm diameter. The energy distribution diagram of the laser amplifier, as shown in Fig. 2.15.

2.2.3 Program Analysis Due to the high requirements of the laser in terms of wavelength, energy, pulse width, repetition frequency, beam dispersion angle, energy stability, etc. Therefore, attention is paid to the following aspects. (1) Selection of the number and diameter of laser bars According to calculations and the results of prior experiments, the energy extraction efficiency of the Nd:YAG amplifier can be as high as 75%, so an output energy of 6 J must be stored in the working substance of about 8 J. Single Nd:YAG laser rod stored energy can be up to 2–2.4 J, because the higher the stored energy, the greater the loss caused by amplified spontaneous radiation. This results in an output energy of 6 J and requires 4 laser bars. Considering the optical damage threshold, the net aperture of the working substance is at least 17 mm, corresponding to a working substance diameter of 15 mm. Direct amplification of 2 stages to extract at least 2 J of energy per stage, and the final output is not less than 10 J of energy, so the diameter of the rod is 20 mm.

2.2 Design Scheme of Laser of AVIC Manufacturing Technology Institute

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Energy distribution of dual-range amplifiers R=0.7 3mJ

1.2mJ Φ15mm G0=2.6

Φ10m G0=8.5

Φ10mm G0=8.5

Φ15mm G0=2.6

440mJ

64mJ

8mJ

6J 2.7J

4.2J

Φ20mm

8J

1.1J

Φ20mm

0.3J

10J

Energy distribution of direct discharge

Fig. 2.15 Energy distribution diagram of laser amplifier

(2) Geometric structure of the rod Slat structured laser systems are very popular nowadays with minimal aberrations and birefringence, but for this system, a bar structure is better due to the difficulty of designing a slat laser and achieving a suitable energy distribution. Calculations and experiments show that by compensating for birefringence, the requirements of the technical specifications can be met with the conventional rod structure. To reduce the temperature gradient inside the laser bar, a relatively long laser bar is required, and the main factor limiting the length of the laser bar is the cost. Calculations show that a reasonable laser bar length is about 120–140 mm for the first four stages of the amplifier. And the length of the last two stages of the laser bar is about 140–160 mm. (3) Structure of the amplification system For efficient energy extraction and to compensate for aberrations using phase conjugate mirrors, dual range amplifiers are required, including a string of amplification stages and rear reflectors. Use of optical phase conjugate mirror (PCM) as rear reflector. (4) Energy distribution of dual-range amplifier First, the energy distribution of the dual-range amplifier was quantitatively analyzed. Calculation of relevant energy storage data based on prior experiments and data simulations.

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Fig. 2.16 Structure diagram of spotlight cavity

Fig. 2.17 Laser bar cross section small signal gain coefficient. a is the cross-section perpendicular to the line connecting the rod lights; b is parallel to the rod

In order to compensate for the thermal-optical aberration well, the thermal load should be evenly distributed between the amplification stages, so it is necessary to use laser bars of the same diameter (2 × F1 5 mm and 2 × F1 0 mm) as much as possible. Compensating for thermogenic birefringence with a 90° quartz rotator. (5) Spotlight cavity Concentration cavity using quartz diffuse reflection cavity, the energy absorption rate is relatively high, and at the same time can obtain a uniform pumping energy distribution. Figure 2.16 shows the structure of the concentrator cavity, including the pulsed xenon lamp, the laser working substance, and the structure of the circulating water circuit for cooling the lamp rod. Figure 2.17 shows the small-signal gain distribution of the laser bar cross-section of the concentrator cavity. Due to the tight development and production cycle and high system stability requirements, existing mature technology solutions are used as much as possible, as shown in Fig. 2.18.

2.2 Design Scheme of Laser of AVIC Manufacturing Technology Institute

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Fig. 2.18 Laser internal structure

Fig. 2.19 Laser structure of AVIC Manufacturing Research Institute. a Nd: glass laser; b Nd: YAG laser

A pre-research project designed a lamp-pumped four-stage amplification section. Figure 2.19 shows a 12 J laser to which laser shock enhancement has been applied. The China Academy of Aeronautical Manufacturing Technology currently has Nd:Glass lasers and Nd:YAG lasers [9] for laser impact strengthening fatigue life extension treatment of components. The parameters of the two types of lasers are shown in Table 2.2.

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Table 2.2 Parameters of the laser Parameters

Nd:glass laser

Nd:YAG laser

Nd:YAG features

Conductivity/(W/mK)

1.02

14

High repetition frequency

Emission cross section/(cm2 )

3.6 ×

Saturation frequency/(J/cm2 )

6

0.62

High extraction efficiency

Damage threshold/(J/cm2 )

Phosphate/silicate 20/40

7–10

Larger aperture required

Size/(mm)

Anyone

≤ϕ 30

Many fine beams of light

10–20

Intrinsic Level Trigger Power

Cooling water signal

2.8 ×

10–19

Intrinsic grade xenon lamp drive power supply

3000V Xenon lamp1

30KV High Voltage Pulse Trigger Timing Circuit

Amplification stage trigger power supply

Amplification stage xenon lamp drive power supply 1

3000V Xenon lamp2

30KV High Voltage Pulse Discharge signal

Turning mirror signal

Amplification stage xenon lamp drive power supply 2

3000V Xenon lamp3

30KV High Voltage Pulse

Fig. 2.20 Laser impact intensification power supply composition

2.2.4 Power Supply Based on IGBT Inverter Technology 1. Laser impact reinforced power supply composition In order to trigger the pulsed xenon lamp and make its arc discharge, the laser impact intensifier power supply must have a xenon lamp trigger circuit, storage capacitor charging circuit, and storage capacitor discharge circuit. The Q tuned neodymium glass laser system consists of an intrinsic level neodymium glass laser oscillator and two amplification level neodymium glass laser main amplifiers. Therefore, the laser impact intensification power supply mainly consists of an intrinsic stage xenon lamp drive power supply, two amplified stage xenon lamp drive power supplies, and two high-voltage pulse trigger power supplies [10]. As shown in Fig. 2.20. In Fig. 2.20, the intrinsic stage and amplifier stage xenon lamp drive power supply output 2 A constant current, the output voltage 0–3000 V continuously adjustable. High-voltage pulse trigger power supply includes two channels: one for the intrinsic level xenon lamp triggering; the other channel can simultaneously trigger two amplification level xenon lamp. Trigger pulse width 20–50 μs, pulse voltage peak up to 30 kV.

2.2 Design Scheme of Laser of AVIC Manufacturing Technology Institute

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Fig. 2.21 Xenon lamp drive power supply main circuit topology

(1) Xenon lamp drive power supply main circuit topology The main circuit of xenon lamp drive power supply adopts IGBT full-bridge inverter main circuit topology, which mainly consists of single-phase bridge rectifier circuit, LC filter circuit, IGBT full-bridge inverter circuit, high-frequency transformer, and secondary full-bridge rectifier filter circuit. As shown in Fig. 2.21. Working principle: 220 V/50 Hz AC power is rectified and filtered by single-phase rectifier bridge, inductor L1, and capacitor C1 to obtain a stable DC voltage of about 310 V, then input into the full-bridge inverter circuit composed of IGBT power switching tubes VT1, VT2, VT3, and VT4 to transform into an AC square wave with a frequency of about 20 kHz, then coupled by high-frequency transformer B1 to the secondary full-bridge rectifier circuit to obtain a pulsating DC output, and finally charged by inductor L2 to energy storage capacitor C2 with a continuously adjustable charging voltage from 0 to 3000 V. The IGBT power switching tubes T1 and T4, T2 and T3 alternately turn on and off to complete the inversion process. (2) High-voltage pulse-triggered power supply topology The pulsed xenon lamp used in the Nd glass laser has an inner diameter of 2 cm, a pole spacing of 30 cm, and a breakdown voltage of about 20 kV. When the highvoltage pulse generated by the high-voltage pulse trigger power supply penetrates the xenon lamp with a peak of 30 kV, the energy storage capacitor C2 can discharge to the xenon lamp through the inductor L3 (L3 is a hollow saturated inductor). The high-voltage pulse trigger power supply mainly consists of a 1000 V DC charging power supply, capacitor CT, high-voltage pulse transformer, gas discharge tube GT1, GT2, and discharge tube trigger circuit, whose topology is shown in Fig. 2.22. After the system is powered up, the 1000 V DC power supply charges the capacitor CT through the current limiting resistor R2, and the maximum charging voltage can reach 1000 V. The DC breakdown voltage of gas discharge tubes GT1 and GT2 is 800 V, and they are connected in series to both ends of the capacitor CT through the original side of the high-voltage pulse transformer, i.e. the terminal voltage of the capacitor CT is the voltage loaded on both ends of gas discharge tubes GT1 and GT2.

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Fig. 2.22 High-voltage pulse-triggered power supply topology

When the capacitor CT is fully charged, the terminal voltage of the capacitor CT (≤1000 V) is always smaller than the series breakdown voltage of the gas discharge tubes GT1 and GT2 (about 1600 V), so the gas discharge tubes GT1 and GT2 are in a blocked state when there is no discharge trigger signal, which keeps the primary circuit of the high-voltage pulse transformer disconnected. When a discharge signal is initiated internally or externally in the system, the trigger timing circuit generates a discharge trigger pulse according to the currently set operating mode to turn on transistor VT1 in Fig. 2.22, so that the +12 V power supply discharges through the primary side of the low-voltage pulse transformer and generates a high-voltage pulse signal of approximately 1000 V on its secondary side. This high-voltage pulse signal acts on gas discharge tube GT2, causing its instantaneous breakdown conduction, so that the series breakdown voltage of GT1 and GT2 drops to about 800 V. At this point, the voltage across the energy storage capacitor CT is higher than the series breakdown voltage of the discharge tubes GT1 and GT2, and the gas discharge tube GT1 is then broken through, thus allowing the capacitor CT to achieve pulse discharge through the primary side of the high-voltage pulse transformer. The high-voltage pulse transformer is a step-up transformer with a ratio of 1:30, so when the discharge tubes GT1 and GT2 are broken down and the capacitor CT is discharged instantaneously, a high-voltage pulse trigger signal with a peak voltage of 30 kV can be generated on the secondary side of the high-voltage pulse transformer. This high-voltage pulse trigger signal is coupled to both ends of the pulsed xenon lamp through the high-frequency capacitor CP, thus breaking through the pulsed xenon lamp. In this way, the energy storage capacitor C2 of the xenon lamp drive power supply immediately arcs the xenon lamp. The pulsed xenon lamp used in the Nd glass laser has an inner diameter of 2 cm, a pole spacing of 30 cm, and a breakdown voltage of about 20 kV. When the highvoltage pulse generated by the high-voltage pulse trigger power supply penetrates the xenon lamp with a peak of 30 kV, the energy storage capacitor C2 can discharge to the xenon lamp through the inductor L3 (L3 is a hollow saturated inductor). The high-voltage pulse trigger power supply mainly consists of a 1000 V DC charging

2.2 Design Scheme of Laser of AVIC Manufacturing Technology Institute

41

power supply, capacitor CT, high-voltage pulse transformer, gas discharge tube GT1, GT2, and discharge tube trigger circuit, whose topology is shown in Fig. 2.22. After the system is powered up, the 1000 V DC power supply charges the capacitor CT through the current limiting resistor R2, and the maximum charging voltage can reach 1000 V. The DC breakdown voltage of gas discharge tubes GT1 and GT2 is 800 V, and they are connected in series to both ends of the capacitor CT through the original side of the high-voltage pulse transformer, i.e. the terminal voltage of the capacitor CT is the voltage loaded on both ends of gas discharge tubes GT1 and GT2. When the capacitor CT is fully charged, the terminal voltage of the capacitor CT (≤1000 V) is always smaller than the series breakdown voltage of the gas discharge tubes GT1 and GT2 (about 1600 V), so the gas discharge tubes GT1 and GT2 are in a blocked state when there is no discharge trigger signal, which keeps the primary circuit of the high-voltage pulse transformer disconnected. When a discharge signal is initiated internally or externally in the system, the trigger timing circuit generates a discharge trigger pulse according to the currently set operating mode to turn on transistor VT1 in Fig. 2.22, so that the +12 V power supply discharges through the primary side of the low-voltage pulse transformer and generates a high-voltage pulse signal of approximately 1000 V on its secondary side. This high-voltage pulse signal acts on gas discharge tube GT2, causing its instantaneous breakdown conduction, so that the series breakdown voltage of GT1 and GT2 drops to about 800 V. At this point, the voltage across the energy storage capacitor CT is higher than the series breakdown voltage of the discharge tubes GT1 and GT2, and the gas discharge tube GT1 is then broken through, thus allowing the capacitor CT to achieve pulse discharge through the primary side of the high-voltage pulse transformer. The high-voltage pulse transformer is a step-up transformer with a ratio of 1:30, so when the discharge tubes GT1 and GT2 are broken down and the capacitor CT is discharged instantaneously, a high-voltage pulse trigger signal with a peak voltage of 30 kV can be generated on the secondary side of the high-voltage pulse transformer. This high-voltage pulse trigger signal is coupled to both ends of the pulsed xenon lamp through the high-frequency capacitor CP, thus breaking through the pulsed xenon lamp. In this way, the energy storage capacitor C2 of the xenon lamp drive power supply immediately arcs the xenon lamp. 2. Control circuit design (1) Xenon lamp drive power double closed-loop control The xenon lamp drive power supply is a constant current and voltage limiting output characteristics, the output current is constant at 2 A, and the output voltage is continuously adjustable from 0 to 3000 V. In order to achieve the output characteristics of the xenon lamp drive power supply constant current and voltage limit, the control circuit uses a double closed-loop control method of current and voltage, that is, the inner loop control for the current negative feedback closed-loop control, the outer loop control for the voltage negative feedback closed-loop control. Xenon lamp drives power supply double closed-loop control schematic is shown in Fig. 2.23 [10].

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2 Laser Shock Hardening Industrial Application System

Fig. 2.23 Xenon lamp drive power supply double closed-loop control schematic

In the process of charging the energy storage capacitor C2 by the xenon lamp driver, when the voltage of capacitor C2 is lower than the set charging voltage, the current negative feedback closed-loop control circuit works, i.e. the xenon lamp driver will charge the capacitor with the set charging current at a constant current; when the voltage of capacitor C2 reaches the set charging voltage, the xenon lamp driver automatically switches to the constant voltage charging mode, at which time the voltage negative feedback closed-loop control circuit works to maintain the voltage at both ends of capacitor C2. (2) Trigger timing circuit When the energy storage capacitor C2 is fully charged, the pulse width of 2.5 ms discharge signal can be activated through the discharge button on the panel of the shock-enhanced power supply or the internal timer, and the signal then passes through a series of trigger timing circuits to generate a high-voltage trigger pulse to ignite the xenon lamp. Considering the strong electromagnetic interference during the discharge of the pump xenon lamp, the trigger timing circuit uses a monostable flip-flop NE555 as the core of the digital circuit to achieve this. The trigger timing circuit is composed of multiple NE555s and their peripheral circuits in cascade, mainly including the discharge signal generation circuit, the synchronization circuit of discharge signal and mirror signal, the delay signal generation circuit, and the trigger signal generation circuit. The single-stage NE555 and its peripheral circuit are shown in Fig. 2.24. In Fig. 2.24, the falling edge of the pre-stage signal is input to the trigger terminal of the NE555 via capacitor C46 and generates a pulse signal with adjustable width at its output, whose pulse width is determined by resistor R57, adjustable potentiometer VR4, and the charging time of capacitor C49, adjusting the resistance value of VR4 can change the duration of the pulse signal. (3) Working sequence of discharge trigger signal The neodymium glass laser is a rotating mirror Q-modulated laser, when the pump xenon lamp is ignited, the resonant cavity reflection loss is very large because the prism surface is not perpendicular to the cavity axis, and the Q value of the cavity is very low at this time, and no laser oscillation can be formed. During this time, the working material in the xenon lamp light pump excitation, the laser on the energy level

2.2 Design Scheme of Laser of AVIC Manufacturing Technology Institute

43

Fig. 2.24 NE555 and its peripheral circuit

reversal particle number accumulated a lot, while the prism surface also gradually turned to the position perpendicular to the cavity axis, the cavity Q value gradually increased to a certain moment to form laser oscillation, and output giant pulse laser. Therefore, in the Q mode, in order to obtain a stable high-power pulsed laser output, it is necessary to accurately control the delay time, which requires a certain delay time after the ignition of the xenon lamp to ensure that the number of reversed particles reaches a great value, while the delay time is exactly equal to the time required for the prism to turn into the cavity position (the position of the two mirrors parallel to each other), so that the formation of laser oscillation, in order to obtain the maximum laser power. The Q-mode operation timing is shown in Fig. 2.25a. Firstly, the discharge signal with a pulse width of 2.5 ms can be activated by the discharge button on the shockenhanced power supply panel or by the internal timer, and the discharge signal is fed into the trigger timing circuit together with the rotating mirror signal and synchronized, i.e. the delay signal is triggered when the discharge signal is high and at the rising edge of the rotating mirror signal; then the trigger signal (20–50 μs) is continuously adjustable at the falling edge of the delay signal, and the trigger signal drives transistor VT1 in Fig. 2.22 to conduct, so that the high-voltage pulse triggering power supply generates a high-voltage pulse to pierce the xenon lamp, thereby discharging the energy storage capacitor C2 of the xenon lamp drive power supply to the xenon lamp to achieve the excitation of the working substance. Delay time 0–1000 μs continuously adjustable, the specific delay time can be determined through field debugging. Therefore, as long as the delay time is suitable, the laser output with high peak power of huge pulses can be obtained. The Nd glass laser can also work in the free oscillation mode, and its operating timing is shown in Fig. 2.25b. In the free oscillation mode, the discharge signal is not synchronized with the rotating mirror signal, so the laser output power is smaller. The free oscillation mode is often used for the optical circuit commissioning of lasers.

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2 Laser Shock Hardening Industrial Application System

Fig. 2.25 Xenon lamp high-voltage pulse trigger signal operating timing. a Q modulation mode; b free oscillation mode

3. Test results (1) Output characteristics of laser shock intensification power supply Figure 2.27 shows the measured current waveform of energy storage capacitor C2 for the rapid discharge of the pulsed xenon lamp. The current transformer ratio used for measuring the discharge current of the pulsed xenon lamp is 1:1000, and the sampling resistance is 0.74 Ω, so the measured peak discharge current of the pulsed xenon lamp is about 5800 A [10] (Fig. 2.26).

Fig. 2.26 Measured current waveform of pulsed xenon lamp discharge waveform

2.3 Workbench of the Strengthening System

45

Fig. 2.27 Laser spot formed by impact on the surface of aluminum foil

(2) Process test Figure 2.27 shows the laser spot formed by impacting on the surface of aluminum foil using this laser power supply. The laser impact strengthening system was used to carry out the laser impact strengthening treatment of TC4 titanium alloy TIG weld head. After the LSP treatment, the number of needle-like α-phase in the area of the weld cross-section near the material surface was reduced, the number of fine isometric crystals in the area of the heat affected zone near the material surface was increased, and the tensile mechanical properties of the joint were better, with the average values of tensile strength, yield strength, and elongation after break increasing by 5.6%, 8.2%, and 66%, respectively [11].

2.3 Workbench of the Strengthening System During laser shock peening, the table has to realize the relative motion of the peening spot on the workpiece, which can be achieved in two ways, either by the workpiece motion or by the laser beam motion, or sometimes by both. Figure 2.28 shows the design of the workbench of MIC, which uses the workpiece movement method. The workpiece movement is realized by a robot, and the movement of the water jet device is realized by another robot. The light guide head in the studio is sometimes required to perform functions such as switching the light on and off and clearing the jet of water mist from the light path. LSP Technologies also introduces an automatic coating system at the work position, when the motion system is more complex. The Chinese Academy of Aeronautic Manufacturing Technology proposed the design of “a laser impact strengthening studio, 03266470.2” in the utility model patent [12], using a four-axis numerical control motion table design, three linear motion axes and a CNC rotary axis, a manual rotary axis are on top of the chuck, avoiding the influence of water on the motion guide, which is also designed with a closed studio to reduce the noise and the impact of the reflected laser on the surrounding personnel and equipment. The movement accuracy of the table reaches 800 mm × 600 mm × 400 mm, the rotation axis can be 360° arbitrary rotation, and

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2 Laser Shock Hardening Industrial Application System

Fig. 2.28 MIC robot-based table structure

Fig. 2.29 Laser shock peening system for blisk by AVIC Manufacturing Research Institute

the movement accuracy can reach 0.05 mm, but the designed load capacity is only 15 kg, only for small test pieces or a single blade laser shock peening. To achieve laser shock peening of large and complex structures, robots are generally used as the motion system, Fig. 2.29 shows the equipment for laser shock peening using a large-load robotic gripper for clamping integral leaf discs in Beijing Aerospace Manufacturing Engineering Research.

2.4 Beam Moving Scanning System Developed by MIC (1) Beam moving scanning system Previous laser shock peening generally adopts the strengthening method in which the laser beam is fixed and the workpiece moves, the implementation of flowing

2.4 Beam Moving Scanning System Developed by MIC

47

Fig. 2.30 Beam scanning system developed by MIC Corporation

constrained layers in this way is relatively convenient. However, for some large parts that are difficult to clamp (such as pipes), and some assembled parts, the strengthening method of moving parts is difficult to implement, and the method of fixed parts and moving beam must be used. The beam moving scanning system developed by MIC Corporation in the United States can realize the rapid positioning and rotation of the laser beam, etc. as shown in Fig. 2.30. MIC Corporation in the United States in April 16, 2010 laser shock intensified flexible beam delivery system patent (Patent No. US20110253690A1, Flexible Beam Delivery System For High Power Laser Systems) in the path implementation [13], the use of a gimbaled reflector (as shown in Fig. 2.31—2, 3, 4) movement to change the direction of the outgoing laser, so that make the incident scan according to the path, also cooperate with a telescope focusing system to adjust the spot size on the target surface. The shape of the laser can also be corrected by the motion of the field rotating mirrors and the column mirror set. The advantage is that only a few small parts need to control the movement of the laser on the target surface for shot peen scanning, improving the stability of the system. However, the disadvantage is that: (1) the telescope system and Stokes lens group in the light-guiding system are independent of each other, the number of lenses is large, and the composition is not simple enough; (2) the gimbal mirror motion mechanism is complicated, and it is more difficult to ensure the motion accuracy. (2) Flight optical processing system As shown in Fig. 2.32, a method and apparatus for double-sided precision forming of plates based on the laser shock wave effect patent (disclosure number: CN101249588A) [14] employs a laser shock head system A (“6” in figure) containing

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2 Laser Shock Hardening Industrial Application System

Fig. 2.31 Schematic diagram of laser shock intensified flexible beam transmission. 1—Laser input; 2—the first gimbal mirror; 3—the second gimbal mirror; 4—closed optical path; 5— optical joint manipulator; 6—flexible optical path; 7—the third gimbal mirror; 8—optical path control; 9—optical path after focusing; 10—workpiece Robot; 11—mirrors for complex workpiece; 12—strengthened parts

adjustable all-reflective mirrors 7, 8 and an adaptive pressure convex deformation mirror 17, an adaptive pressure concave deformation mirror 18, an all-reflective mirror 16, and the laser shock head system B (“28” in figure) with concave mirror 19 implements the motion of the incident laser guide system, including up and down and left and right translation, so that the incident laser follows a predetermined path for scanning the vertical plate surface for shot peening. The advantage of the flying optical processing system is its flexibility and its application in 2D and 3D laser processing; the disadvantage is the realization of the entire light-guiding system motion. First, the entire light-guiding system will take up a large space and be less flexible in order to cover the entire target surface with the motion of the light-guiding system. And it is difficult to guarantee the position accuracy of each point during the motion of the whole light-guiding system in a large range of motion, and the vibration during the motion may also have adverse effects on the components in the light-guiding system. (3) Galvo scanning system Galvanometer is an excellent vector scanning device, it is a special pendulum motor, the basic principle of the energized coil in the magnetic field to generate torque, but different from the rotary motor, its rotor through the mechanical torsion spring or electronic methods to add a reset torque, the size of the angle proportional to the rotor deviation from the balance position, when a certain current is applied to the coil and the rotor is deflected to a certain angle, the electromagnetic torque and the reversion torque are equal in magnitude and cannot rotate like an ordinary motor,

2.4 Beam Moving Scanning System Developed by MIC

49

Fig. 2.32 Flying optical processing system for laser shock strengthening double-sided forming of sheet metal. 1—Laser; 2—all-reflective mirror; 3—first optical path; 4—lens; 5—reflected laser; 6—first optical path; 7, 8—all-reflective mirror; 9—first optical path control; 10—laser control; 11—workpiece; 12, 13—all-reflective mirror; 14—second optical path; 15, 16—all-reflective mirror; 17, 18—shaper mirror; 19—concave mirror; 20—second optical path control; 21—reflex control; 22—shaper control; 23—optical path control; 24—probe; 25—fixture; 26—fixture control; 27—computer; 28—the laser shock head system B

but can only deflect, the angle of deflection is directly proportional to the current, as with a galvanometer. Figure 2.33 shows, an optical path device and method for laser peening of large workpieces patent (disclosure number: CN104923606A) [15] using X-direction Galvo scanner 1 and Z-direction Galvo scanner 2 rotating around its own axis of rotation, changing the direction of laser beam emission to hit the corresponding coordinates on the forming workpiece, so as to achieve mobile beam laser shock to strengthen the metal sheet. The advantages of this Galvo scanning system are simple arrangement of the optical path device, simple and small range of motion, ease to ensure the accuracy, the spot overlap rate and scanning speed on the target surface of the workpiece can be easily adjusted; the device occupies little space, low energy consumption of motion, and good flexibility.

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Fig. 2.33 An optical path device and method for laser shock peening of large workpieces

References 1. Lan XJ (1995) Laser technology (in Chinese). Huazhong University of Science and Technology Press, Wuhan 2. Wu HX, Guo DH, Zhou YS et al (1985) High-power Nd:glass laser device and shooting range system and their LPX experimental research (in Chinese). J Univ Sci Technol China 1:34–40 3. Yang XH, Guan HB, Wu HX et al (2010) Multiple transverse mode laser oscillator of high repetition-rate Nd: shock treatment installation (in Chinese). Laser Optoelectron Prog 47(7):105–108 4. Ren JX, Ye A, Zhang HB et al (2007) Nd:Ce:YAG laser with high peak power and large energy (in Chinese). Opt Optoelectron Technol 5(2):25–26 5. Cao SS, Wang MQ (1997) High peak power repetitive pulse solid state laser (in Chinese). Laser Technol 21(5):266–271 6. Zhou P, Xu XJ, Liu ZJ et al (2008) New technology and new configuration for high energy laser system (in Chinese). Laser Optoelectron Prog 45(1):37–42 7. Wang YQ, Mi GJ, Du T et al (2003) Nd:YAG laser with high peak power and large energy (in Chinese). Laser Infrared 33(3):188–189 8. Liu XS, Wang ZY, Yan X et al (2008) 56J high energy lamp-pumped pulsed Nd:YAG solid state laser (in Chinese). Laser Technol 32(3):237–239 9. Zou SK, Wu JF, Gong SL (2015) Laser peening systems and the effects of laser peening on aeronautical metals sheet (in Chinese). AASCIT Commun 2375–3803 10. Zhang W, Qi BJ, Xu HY et al (2013) Laser shock processing power supply topology and control (in Chinese). Trans China Weld Inst 34(2):97–100 11. Xu HY, Zou SK, Che ZG et al (2011) Influence of laser shock processing times on TC4 argon arc welding joint microstructure and properties (in Chinese). Chin J Lasers 38(3):92–96 12. Zou SK, Tan YS (2003) Process chamber for laser peening (in Chinese). CN, CN2641057Y 13. Dane BC, Lao EWH, Harris FB, et al (2011) Flexible beam delivery system for high power laser systems. USA, WO2011129921A2 14. Jiang YF, Zhang YK, Yao HB, et al (2008) A double-sided precision forming method and device for sheet metal based on laser shock wave effect (in Chinese). CN, CN101249588B 15. Hu YX, Yu XC, Yao ZQ, et al (2015) A light path device and method for laser shot peening forming of large workpiece (in Chinese). CN, CN201510197097.4

Chapter 3

Stability Factors and Safety Protection of Laser Shock Peening

3.1 Process Stability Factors With the development of laser device technology, the frequency of intense pulse laser can reach 1 Hz or even more than 10 Hz, and the processing efficiency of laser shock intensification has been greatly improved. The strengthening time of a single blade of an American F110 engine is shortened from the initial 30–12 min per blade and may be shortened to 4 min in one step. In 2001–2002, the U.S. Air Force Manufacturing Technology Association developed a specific laser shock strengthening technology for the F119 engine compressor integral disk production line, including automatic rapid coating, process parameter monitoring, and image positioning technology, which reduced the strengthening processing time of each integral disk to 8 h. LSPT, GE, MIC, and Toshiba’s LSP process control process stability has always affected its application and has been studied vigorously and made great progress. The development of LSP equipment in China is relatively slow, the engineering application experience is less, and the stability of process control technology is relatively weak. AVIC Manufacturing Technology Institute (China) has many years of research foundation of LSP technology. In 2004, it started the research of LSP with water confinement titanium alloy blades and carried out high-frequency YAG laser shock strengthening experimental research, which gradually solved the technical problems of high-frequency intense pulse LSP of aero-engine blades. Several key factors of process control process stability of engine blade after LSP were studied, and process control solutions were proposed, laying a foundation for engineering and automatic application of LSP technology [1, 2]. The LSP confinement mode contains four main elements, namely laser, confinement layer, absorption layer, and target material to be processed by impact. Therefore, the stability of the process can be divided into four aspects accordingly, as shown in Fig. 3.1 [3]. The stability of the process is meaningful only when the LSP is operated at a higher frequency under the automatic process. The other four factors also influence each other, and the following is only comparatively analyzed from © National Defense Industry Press 2023 S. Zou et al., Laser Shock Peening, https://doi.org/10.1007/978-981-99-1117-2_3

51

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3 Stability Factors and Safety Protection of Laser Shock Peening

Fig. 3.1 Four elements of constrained mode and process stability for LSP

four aspects. Since the current LSP process is dominated by water confinement, the process stability under water confinement is mainly discussed in this chapter.

3.1.1 Adjustable Laser Spot and Continuous Laser Path Laser is the energy source of LSP, so the process must ensure the unobstructed energy channel first. Although the laser spot size and energy density are very important factors, they are generally not adjusted in the LSP process and do not need to be considered in the study of process stability. However, in some cases, the control of the laser path is particularly important. One case is the control of the laser off, which is very important in process stability control, especially in abnormal feedback control. For example, when the damage of the absorption layer may lead to damage to the target, the control of the laser off is needed. Another situation is spot shape control, target strengthening process needs, such as leaf edge change spot shape can obtain the ideal form of overlapping; Another situation is the reflected light control, LSP process in the dimming path, the reflected light cannot destroy the external optical path or laser. In the patent of July 12, 2005 (Patent No. 6917012B2, Reducing electromagnetic feedback during laser shock peening) [4], GE proposed the use of Faraday isolators to prevent laser reflections from entering the laser, and this technology is now commonly used in industrial lasers. In patent No. 7137282B2 (Laser shock peening) [5], R–R proposed that in addition to the optical switch in the optical path, the mask can be used to change the shape of the spot in and out of the laser optical path, and to obtain the ideal spot form; the more commonly used spot form is the square spot, the square spot can obtain a better overlapped form, as shown in Fig. 3.2. Since round laser crystals are easy to grow and pump uniformly, most lasers currently use round laser crystals, whose output is a Gaussian-distributed round spot. For processes such as heat treatment and LSP, a square spot is needed to meet various process requirements, and a square spot can generally be obtained by using masks, special integrating mirrors, etc. But the mask method will lose some energy,

3.1 Process Stability Factors

53

Fig. 3.2 Patented by R–R Corporation, USA a Laser switch control; b laser shape control

for high-energy LSP, the use of the mask method will reduce processing efficiency. Special integral mirrors with dozens or even hundreds of processing surfaces are costly to process and easily damaged by high-energy laser. For example, π shaping mirror withstands a peak power density of only 200 MW/cm2 , and the shaping mirror cannot be used for LSP and other laser processing technology with peak power density greater than 1 GW/cm2 . In addition, the shaping mirror is expensive, Φ6, Φ12, and Φ34 shaping mirror price of 3400 euros, 6000 euros, and 15,000 euros, respectively, so there are great limitations in the use of this shaping mirror. On April 16, 2008, AVIC Manufacturing Technology Institute (China) proposed a laser beam shaping lens with no loss of laser energy, a laser peak power density higher than 3 GW/cm2 , and a low cost. (patent No.: CN101256287A, “A laser beam shaping quintal lens and quad lens”) uses a beam shaping system (quintal lens or quad lens) [6], realizes the conversion of round laser spot to square laser spot with almost zero energy loss, and makes a great breakthrough in laser external optical path shaping, as shown in Fig. 3.3. The laser shaping lens costs only a few hundred yuan, less than one percent of the π shaping lens.

Fig. 3.3 Schematic diagram of spot conversion device of AVIC Manufacturing Technology Institute

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3 Stability Factors and Safety Protection of Laser Shock Peening

The continuity of the laser path is an important technology in the stability of LSP process, especially in the process of high-frequency LSP, the continuity of the laser path is very crucial. LSP makes the sputtering speed of restrained medium water very fast, the sputtering water mist cannot touch the lens, can not block the laser path, at present at home and abroad on the cleaning of the laser path a lot of research, basically to blow. In Patent No. 6521860B2, Beam path clearing for LSP [7], the American LSP Technology Company mentioned the cleaning of laser path. Clean up the debris and water mist on the laser optical path by blowing air, diaphragm, fan, etc. In the patent No. 6713716, Reduced mist LSP, dated March 30, 2004 [8], GE proposed to use a flat nozzle (called air knife in the patent) between the final lens of the laser and the workpiece to remove water mist and other particles affecting the laser light path in the LSP area by emitting air. The gas used is air or nitrogen. The focal length of the lens, the focus angle, and the nozzle position are also calculated in the patent. During the LSP process, to ensure the continuity of the laser path, a focusing lens with a long focal length of 1 m is adopted by the AVIC Manufacturing Technology Institute to carry out the design of blowing air or pumping air on the laser path and timely clean the water mist on the optical path, as shown in Fig. 3.4. If conditions permit, we can also use the laser downward slightly inclined method, to avoid sputtering water pollution lens, but also can reduce the reflected light back laser. Although the method of a slightly downward tilt of the laser has little influence on the size of the spot, if the straight direction of the blade motion is not perpendicular to the direction of the laser incident, it is necessary to interpolate the focus position to ensure the stability of the spot position.

Fig. 3.4 The continuity of the laser path of LSP

3.1 Process Stability Factors

55

3.1.2 Flat Confinement Layer The layer covered with a laser-transparent material on the surface of the absorption layer is the confinement layer, which is used to limit vaporization and increase pulse pressure and action time. In the process of LSP, the confinement layer is the main factor determining the confinement mode, and the confinement layer is divided into solid medium and liquid medium [1, 9]. Solid media can be divided into hard medium and soft medium. Optical glass is a commonly used hard medium, as shown in Fig. 3.5a. Its advantages are less absorption of laser energy and high pressure of shock wave generated. The impact treatment of soft media on non-planar surfaces can achieve a good fit, as shown in Fig. 3.5b. However, soft media materials (such as organic materials) have a higher absorption rate of infrared laser than glass and water and are easy to breakdown, so they are not suitable for continuous LSP. Water in the liquid medium is the most economical confinement medium, and the use of water as a confinement medium must consider the matching of laser wavelength, such as the use of near-infrared wavelength laser is easy to be absorbed by water, and UV laser is easy to lead to water breakdown. The commonly used Nd: YAG glass laser with 1.06 μm, 10–50 ns pulse laser with water is feasible, and the frequency-doubling Nd: YAG laser is better. Water confinement can be divided into two ways: still water and Flowing water confinement. In the vaporization process of the absorption layer, still water is prone to pollution, as shown in Fig. 3.5c, and the shock wave will make the water surface fluctuate, affecting the next impact process. As shown in Fig. 3.5e, it takes time for water to get a flat interface in accurate treatment, so the laser shock frequency cannot be very high, resulting in an intermediate form of still water and flowing water. In Fig. 3.5d, the workpiece is placed in water, and the laser enters from the side window. The water tank can use the running water to wash away the polluted water at the treatment site, so as not to affect the next process. Water and optical glass work differently. Studies have shown that the power density of 4 GW/cm2 is required when the surface residual stress extremum −350 MPa is obtained, which is equivalent to the power density of 1.7 GW/cm2 when glass is used as the confinement layer. The optimal impact pressure value is

Fig. 3.5 LSP confinement mode is commonly used. a Optical glass hard media; b soft media; c still water; d intermediate form of still and flowing water; e flowing water. 1—Intense pulse laser; 2—focusing lens; 3—confinement layer; 4—absorption layer; 5—metal target; 6—water tank; 7—sprinkler head

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3 Stability Factors and Safety Protection of Laser Shock Peening

2.5 GPa. Beyond this value, the surface residual stress saturation and the influence of surface waves lead to a decrease in the stress value. If the confinement layer and absorption layer are not used and the plasma is generated by the metal specimen itself, the thermal effect after impact is relatively obvious, which easily leads to a tensile stress zone. After LSP, unsteady water flow is generated, and it takes time for water flow to recover and smooth, especially at the blade edge. Unsteady water flow is easy to cause uneven thickness of water confinement, and even lens effect may occur at local points, as shown in Fig. 3.6a, which may lead to high local power density or damage of the absorption layer. To ensure the stability of the confinement layer, a diversion layer can be introduced to the edge of the blade, or aluminum foil tape can be applied beyond the edge of the blade for drainage, as shown in Fig. 3.6b. In 2003 and 2005, LSP Technologies, Inc. (Patent No. 6548782, Patent No. 6841755, Patent No. 6548782, Patent No. 6841755, Overlay control for LSP). It is mentioned in overlay control for LSP. The detection laser is used to irradiate the water binding layer at a certain Angle, and the thickness of the water confinement is calculated according to the reflected light path difference between the upper and lower surfaces of the water film, to determine the thickness of the binding layer. The thickness of the water confinement layer can be controlled by adjusting the air outlet pressure and time of the air injection device until the thickness of the confinement layer meets the requirements, as shown in Fig. 3.7. To avoid the water transmission effect and the interaction between laser and water flow on the blade surface, especially on the curvature surface of the blade root, a smooth and uniform water confinement layer was obtained by the inclined jet mode on the blade surface. During continuous LSP, the flatness of the confinement layer affects whether LSP can be implemented or continuous. The IBR blade density is large, the twist Angle is large, can not be removed for separate processing, in the process of LSP, often need to manually adjust the position of the nozzle, so that the shape of the water confinement

Fig. 3.6 LSP of target surface confinement layer. a Local lens effect of unstable water flow; b stable restrained layer of aluminum foil

3.1 Process Stability Factors

57

Fig. 3.7 LSP Technologies tests the water layer thickness and constrains the layer thickness using airflow regulation from the air nozzle. 20—Work piece; 42—confinement layer; 24—coating ejector; 25—air ejector; 52—nozzle; 54—mane

in each adjustment has a great deal of randomness, affecting the consistency of the treatment process, greatly reduce the treatment efficiency. “Method for acquiring stable water film in laser shock processing of blisk” was proposed by AVIC Manufacturing Technology Institute (Patent No. CN103205546A) [10]. The blisk (2) is clamped on the 6-DOF manipulator (3). During LSP, the position of nozzle (9) remains unchanged. And by the six degrees of freedom manipulator (3) to ensure that the laser spot (5) position, nozzle (9) spray water (7) on the blade of the contact position (6) and water (7) and contact tangent plane Angle (8) unchanged, so as to ensure that the laser spot (5) position to obtain a stable water film (4). The advantage of this technique is that the laser spot avoids the turbulent area near the water contact. The Angle between the tangent plane and the horizontal plane at the position of the laser spot is 70–80°, so as to ensure that the velocity and direction of the water film covering the blade surface are also constant under certain water flow conditions, so as to obtain a stable water film. Usually, the laser head is separated from the water nozzle. The distance between the laser and the water nozzle should be more than 40 mm. After the water nozzle sprits water on the workpiece, it flows to form a flat water flow as the confinement layer, as shown in Fig. 3.8. However, when the space of the area to be treated is narrow or the surface is uneven, it is difficult to carry out LSP when the flat water flow cannot be used as the confinement layer. In addition, even for a flat surface, when water flow is used as the confinement layer, the laser transmission will be blocked by water mist formed by water splashing. Therefore, the LSP frequency cannot be very high, generally no more than 2 Hz. The matching of water flow and laser becomes an important factor restricting the range and efficiency of LSP. There are also frequency-doubling laser underwater LSP methods, but this method uses the whole workpiece immersed in water; from the side of the deionized water recharge, this method also has limitations. A water/laser coaxial device for LSP was proposed by AVIC Manufacturing Technology Institute (patent No. CN102505065A) [11]. Uses a nozzle (2) mounted on

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3 Stability Factors and Safety Protection of Laser Shock Peening

Fig. 3.8 The water confinement layer is uniformly stabilized during LSP of the blisk. 1—Laser; 2—blisk; 3—6° of freedom manipulator; 4—water film; 5—laser spot; 6—contacts; 7—water flow; 8—angle of the tangent plane; 9—nozzle. Laser shock intensifies the uniform stability of the water confinement layer in the integral disk

Fig. 3.9 Structure diagram of laser shock peening water/laser coaxial device. 1—Lens; 2—nozzle; 3—deionized water input hole; 4—open hole at lower end of conical nozzle

the lens (1) to inject deionized water into the deionized water input hole (3) on the side of the nozzle (2). The open hole (4) at the lower end of the conical nozzle is coaxial with the lens (1), so that the laser and stable water flow are coaxial output from the nozzle (2) to process the workpiece with narrow space or uneven surface, as shown in Fig. 3.9.

3.1 Process Stability Factors

59

3.1.3 The Integrity of the Absorption Layer During LSP, the function of the preset absorption layer on the surface of the metal target is to absorb laser energy to produce plasma and prevent melting and vaporization of the metal surface. Therefore, the basic requirement for the absorption layer is to choose the material with low-heat conduction coefficient and low heat of vaporization, increase its own heat absorption, and reduce the heat conduction to the target. Aluminum foil, lead, zinc, black paint, etc. are more effective surface coating materials. The integrity of the absorption layer directly affects the quality of the target, so it is an important link to ensure the quality. However, the quality of the absorption layer is uncontrollable in the process of LSP. Once the absorption layer breaks or bulges, the process of LSP needs to be interrupted. There are two main reasons for the damage of the absorption layer. First, the surface wave generated by the first laser pulse causes the local protrusion of the absorption layer at the second impact position, which leads to the damage of the absorption layer at the next impact position. Second, there are too many times of strengthening and the thickness of the absorption layer is too thin, resulting in damage. As an important defense line of the target in the LSP confinement mode, the absorption layer becomes the main monitoring target in the process stability control. Once there is an anomaly, the control system will issue a laser off instruction to the laser or laser path. LSP Technology Company of the United States mentioned the control of the absorption layer in Patent No. 6841755, overlay Control for LSP (Patent No. 6841755, overlay control for laser peening) [12] on January 11, 2005. The process is to determine the working position, paste the absorption layer, measure the thickness of the absorption layer, implement the confinement layer, and measure the thickness of the confinement layer. By testing the thickness of the coating, determine whether the coating thickness is uniform and whether the uniformity of the thickness is within the original range. If it does not meet the requirements, it is necessary to uncover the coating and rearrange the confinement layer until the coating meets the requirements. A large number of LSP experiments were conducted by AVIC Manufacturing Technology Institute. Due to unskilled operation, bubbles and folds existed in the aluminum foil absorption layer on the blade surface, as shown in Fig. 3.10a. These defects not only affected the LSP effect, but also caused the damage of aluminum foil on the blade surface or edge, as shown in Fig. 3.10b [13]. Therefore, the surface of the blade needs a smooth aluminum foil absorption layer. There are two methods to eliminate the surface defects of the blade. The tool rolls the aluminum foil absorption layer and repeatedly glues the aluminum foil to obtain a smooth and flat aluminum foil absorption layer, as shown in Fig. 3.11. AVIC Manufacturing Technology Institute has established a complete process and criteria for LSP of blades. The adhesion mode of aluminum foil directly affects the effect of LSP. Laser energy of 50 J, pulse width of 30 ns, wavelength of 1064 nm, confinement layer of 1–2 mm, aluminum foil absorption layer of 0.1–0.2 mm, and spot diameter of 4–5 mm are used for LSP on both sides of the blade. The laser power density is 8 GW/cm2 . The blade surface is bonded with two aluminum foil

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3 Stability Factors and Safety Protection of Laser Shock Peening

Fig. 3.10 Aluminum foil absorption layer on blade surface. a Defects in aluminum foil before LSP; b aluminum foil is damaged after LSP

Fig. 3.11 Smooth, flat aluminum foil

absorption layers, as shown in Fig. 3.12. However, the adhesive double layer of aluminum foil absorption layer on blade surface has two defects: first, to protect the damage of blade flanks by LSP, small energy LSP is used to treat the blade flanks, and the adhesive double layer of aluminum foil absorption layer reduces the effect of LSP blade flanks. Second, standard energy LSP was used to strengthen the blade flanks. The double-layer aluminum foil absorption layer was damaged, and the blade flanks were damaged to different degrees, as shown in Fig. 3.13. Because the blade flank is the key part of crack initiation and blade fracture, special protection and LSP are needed to improve the blade flank fatigue strength. Based on the disadvantages of the double-layer aluminum foil absorption layer mode, the blade surface is bonded with the single-layer aluminum foil absorption layer mode to improve the effect of LSP, as shown in Fig. 3.14. The single-layer aluminum foil absorption layer mode not only reduces the aluminum foil bonding time and improves efficiency, but also protects the blade flanks and avoids blade damage when the blade flanks are treated by high-energy LSP. In the mode of single-layer aluminum foil absorption layer, the two sides of the blade, the flank of the blade and the surface of the blade, are treated with the same energy LSP.

3.1 Process Stability Factors

61

Fig. 3.12 Two-layer aluminum foil blade treated by LSP

Fig. 3.13 Blade flank injury due to aluminum foil breakage during LSP

Fig. 3.14 Laser shock strengthening effect of blade double-sided adhesive single-layer aluminum foil

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3 Stability Factors and Safety Protection of Laser Shock Peening

Fig. 3.15 Pitting. a Aluminum foil surface; b blade surface

Due to the selection of different LSP parameters and adhesion modes of aluminum foil absorption layer, after LSP of the blade, pitted spots appear on the joint surface of aluminum foil absorption layer and blade, as shown in Fig. 3.15, thus affecting the surface roughness of the blade. Experimental analysis and observation showed that the generation and size of pitted spots on the surface of blade and aluminum foil enhanced by laser shock were related to laser energy, laser energy distribution, and aluminum foil absorption layer mode. The adverse effects of blade surface pitting on blade performance were eliminated, lower limits of laser energy and strengthening effect were determined, and the operation process strictly followed the standards of the aluminum foil absorption layer mode of AVIC Manufacturing Technology Institute of China.

3.1.4 The Quality of the Target Material The quality of the target material is the core of the stability control of LSP, but it is also an uncontrollable factor in the process of LSP. As the blade is a thinwalled structure, it is mainly used for deformation control in impact treatment. There are many discussions on the control of deformation and spalling in early patents. Simultaneous reinforcement on both sides can avoid deformation well, but it is easy to cause spalling. Since the target is behind the confinement layer and the absorption layer, it is difficult to control the quality of the target. The early quality assurance technology based on the capacity analysis of LSP pits (U.S. Pat. No. 5948293) [14], the critical Angle determination method of surface wave (U.S. Pat. 991) [15], ultrasonic multistage transform rotatable scanning instrument and method (U.S. Pat. No. 597489) [16], and other methods are difficult to realize online, while plasma monitoring of LSP quality control (U.S. Pat. No. 6554921 B2). Moreover, acoustic signal monitoring, natural frequency testing, and other methods are suitable for online detection of target reinforcement and have good reference value for process stability. However, process stability monitoring must be established on the basis of long-term stable process and

3.2 Research on the Application of Confinement Layer

63

good empirical function, and there are many difficulties in online feedback control. AVIC Manufacturing Technology Institute of China has established an online blade natural frequency detection system. In the process of LSP, each laser shock will cause a small change in the blade natural frequency and indirectly control the LSP effect on the target through the change in blade natural frequency. In the deformation analysis experiment of thin-walled structure of titanium alloy, it is found that when the blade only deals with the edge 8–10 mm region, the blade has such deformation as curling and torsion. Under the condition of reasonable parameter control, the deformation will not exceed 0.25 mm even if the single-sided LSP is used, and the deformation is easy to control. For parts with a thickness greater than 1.5 mm, double-sided LSP can be used successively. For parts less than 1.5 mm thick, doublesided reinforcement can be used. Under the condition that the surface compressive residual stress is satisfied, the thin-walled structure should be strengthened by lowintensity laser shock as far as possible to avoid spalling. If necessary, the absorption layer should be arranged on the back surface.

3.2 Research on the Application of Confinement Layer LSP technology can greatly improve the fatigue properties of materials and has been successfully applied in aerospace, nuclear industry, and other fields. At present, more and more researchers at home and abroad pay attention to the basic theory and basic technology of LSP. Confinement layer plays a very important role in LSP technology, which directly affects the effect of LSP. Therefore, the theoretical study of confinement layer and the selection of confinement layer have become the current research hotspot. This section introduces the research and application of water confinement layer in LSP [17].

3.2.1 Introduction and Application of Water Confinement Layer (1) Introduction of water confinement layer The confinement layer media mentioned in domestic and foreign literature include K9 optical glass, plexiglass, silica gel, synthetic resin, and water [18]. Glass confinement layer has the advantage of significantly increasing the shock pressure, but the disadvantage is only suitable for flat surface processing, and it is fragile and difficult to clean. Silica gel and synthetic resin have little binding force with target material and are difficult to be reused. The advantages of the water confinement layer are cheap, clean, good repetition effect, surface machining, and the flowing water confinement layer can take away the solid dust particles after the plasma explosion.

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3 Stability Factors and Safety Protection of Laser Shock Peening

These advantages cannot be replaced by all other confinement media [1]. The disadvantage of water-restrained layer is poor rigidity leads to less than glass restrained effect; Breakdown plasma is easy to occur at high power density. At high impact frequency, the water layer splashes, and the water droplets and water mist in the optical path scatter the laser. Through theoretical analysis and experimental verification, and with effective technology, these shortcomings of the water-confined layer can be diluted. (2) Application of water confinement layer LSP is the most mature application in the strengthening of aero-engine blades. GE Company and LLNL laboratories have successfully used water as a confinement layer in the curved structure of blades. Stable and flat water layer is closely related to the strengthening effect. When using the flow confinement layer, it should be monitored in real time to avoid water ripples at the impact, the lensing effect of which may cause the ablation of the local absorption layer. The absorption-free laser shot peening technology of Toshiba Company also uses water as the confinement layer and uses the strong penetration of 532 nm wavelength laser in water for underwater processing. Due to the direct contact between water and the workpiece, dissolved oxygen in water and electrolytic oxygen will oxidize the workpiece surface. Therefore, the water as the confinement layer needs to undergo a long period of deoxygenation. AVIC Manufacturing Technology Institute of China has been devoted to the study of water confinement layer. A large number of experimental results show that LSP using water as confinement layer can successfully strengthen titanium alloy, martensitic stainless steel, superalloy, and other materials with high yield strength.

3.2.2 Laser Absorbance in Water and Selection of Restrained Layer Thickness When the laser power density is less than the breakdown threshold of the waterconfined layer, a small part of the laser is still absorbed by water, and the absorption rate in water varies greatly with different laser wavelengths. Figure 3.16 shows the transmission spectral line of laser in pure water. The relationship between absorption length Δ and laser power density is ( x) Ix = I0 · exp − Δ

(3.1)

where I 0 incident laser power; I x Laser power density when propagating to distance x in water. Equation (3.1) is applicable to the case where the laser absorptivity is below 10– 20% [19]. The 532 and 1064 nm lasers are most commonly used in shot peening

3.2 Research on the Application of Confinement Layer

65

Fig. 3.16 Absorption length of laser with different wavelengths in pure water

technology. As can be seen from Fig. 3.16, the Δ of these two wavelengths are about 30–35 m and 0.02–0.025 m, respectively. It can be seen from Table 3.1 that when the laser energy is absorbed by 1% by water, the propagation distance of 532 and 1064 nm wavelength laser in water is about 300–350 mm and 0.2–0.25 mm, respectively. It can be seen that the absorptivity of 1064 nm infrared light in water is more than 1000 times greater than that of 532 nm green light. The green laser of 532 nm has very strong penetration in water and can be used for laser shock peening of underwater workpiece. Although the penetrability of 1064 nm infrared laser in water is not as good as that of green light, the single pulse energy generated by this wavelength laser is the largest (up to 100 J), so wavelength of 1064 nm is mostly used for laser shot peening with absorption layer [20]. The thickness of the water-constrained layer of laser shot peening with absorption layer is generally 1–2 mm. If it is too thick, it absorbs too much laser, and if it is too thin, the constrained effect is not ideal. Table 3.1 Laser propagation distance in water

Laser propagation distance in water

Proportion of laser absorbed by water/%

Proportion of laser penetrating water/%

0.0101Δ ≈ 1%Δ

1

99

0.0202Δ ≈ 2%Δ

2

98

0.0513Δ ≈ 5%Δ

5

95

0.105Δ ≈ 10%Δ

10

90

0.223Δ ≈ 20%Δ

20

80

0.693Δ

50

50

Δ

63.2

36.8

66 Table 3.2 Comparison of water and K9 glass restrained layer

3 Stability Factors and Safety Protection of Laser Shock Peening Constrained medium Acoustic impedance

(106

g/cm2

Water

1.14

0.165

τ 1 /τ 2

3.7

2.5

Z √

1.3

0.3

1.14

0.55

Z

s)

K9 glass

3.2.3 Influence of Water-Confined Layer on Shock Wave The rapidly expanding plasma is restricted by the water-constrained layer. When the laser power density is not very high, the peak pressure of the shock wave is 4–10 times that of the unconstrained layer, and the pulse width of the shock wave is 2–3 times that of the unconstrained layer. The interface acoustic impedance (Z) of the constrained layer and the absorption layer affects the shock wave pressure. The expression of the interface acoustic impedance is 1 2 = Z (Z 1 + Z 2 )

(3.2)

where Z 1 Acoustic impedance of target; Z 2 Acoustic impedance of constrained medium [21]. Table 3.2 shows the attribute comparison between K9 optical glass and water, in which the ratio of shock wave pulse width (τ 1 ) and laser pulse width (τ 2 ) is measured under the laser power density of 0.73 GW/cm2 . Aluminum foil is selected as the absorption layer for the calculation of interface acoustic impedance Z. From the results in Table 3.2, it can be seen that the acoustic impedance of K9 glass is 7 times that of water, but the shock wave pressure is proportional to the square root of interface acoustic impedance. The efficiency of K9 glass to generate shock waves is only 2 times that of water.

3.2.4 Parasitic Plasma (1) Generation mechanism and preventive measures of parasitic plasma When the power density is greater than the electric breakdown threshold of the water-constrained layer, another plasma, called “parasitic plasma” (Fig. 3.17), is generated in the water-constrained layer above the “restricted plasma” on the surface of the absorption layer. There are two main generation mechanisms of parasitic plasma: avalanche ionization mechanism (AI). The initial free electrons in the waterconstrained layer rapidly increase the plasma density through reverse bremsstrahlung

3.2 Research on the Application of Confinement Layer

67

Fig. 3.17 Parasitic plasma

absorption. This process is expressed as e− + M + hv → 2e− + M+ . The other is the multi-photon ionization mechanism (MPI). The energy absorbed by the particles of the constrained layer medium by multiple photons exceeds 12.6 eV, causing the following process: M + m · hv = e− + M− . Most scholars believe that the MPI mechanism is the main source of initial free electrons in the AI mechanism. With the change in laser wavelength, the effect of the two mechanisms also changes. The longer the laser wavelength is, the more dominant the AI mechanism is. The shorter the laser wavelength is, the more dominant the MPI mechanism is [22]. Fabbro et al. initially believed that the free electron density of 1020 /cm3 in water was the sign of parasitic plasma generation; Benxin Wu, an American scholar, believes that if the propagation distance of the laser in water is large, the free electron density when the parasitic plasma is generated is far less than 1020 /cm3 , and the electron density gradient in the direction of the optical path is large. Parasitic plasma will cause extremely unstable effective experimental parameters, so in the process of laser peening, it should be avoided to work above the electric breakdown threshold. In addition, the use of STR pulse waveform laser can improve the electric breakdown threshold, and it is also very important to select a water medium with a low dielectric coefficient, such as medical deionized water. (2) Influence of parasitic plasma on process Due to the existence of parasitic plasma, most or even all of the laser energy is shielded. With the increase in power density, the peak pressure of the shock wave reaches a saturated state, and the pulse width of the shock wave decreases. The smaller the laser wavelength, the higher the conversion efficiency α, and the greater the tendency is to produce breakdown plasma. See Table 3.3 for electric breakdown threshold I sat and saturated shock wave pressure of 532 and 1064 nm lasers. Table 3.3 Breakdown threshold and saturation pressure

Wavelength/(nm)

I sat /(GW/cm2 )

Psat /(GPa)

α

1064

10

5.5

0.25

532

6

5.0

0.40

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3 Stability Factors and Safety Protection of Laser Shock Peening

Wu et al. [23] studied the changes in laser transmissivity and shock wave pulse width when 1064 nm/25 ns laser generated parasitic plasma by comparing numerical simulation results with experimental data. The results of laser pulse width and shock wave pulse width under parasitic plasma are shown in Figs. 3.18 and 3.19. It can be seen from Figs. 3.18 and 3.19 that when the laser power density is 5.5 GW/cm2 , the laser is completely penetrated and the shock wave pulse width is the largest; When the laser power density is 6–8 GW/cm2 , the laser transmissivity and shock wave pulse width decrease fastest; When the laser power density rises above 10 GW/cm2 , the peak pressure of the shock wave reaches saturation, because the strong absorption of the parasitic plasma occurs at the rising edge of the laser pulse waveform, and the pulse width of the shock wave continues to decrease. Fig. 3.18 Laser pulse width under parasitic plasma

Fig. 3.19 Shock wave pulse width under parasitic plasma

3.3 Status of Damaged Tape and Absorption Layer

69

3.2.5 Optical Path Purification In the process of laser peening, the plasma explosion causes the water-confined layer to splash into the air. The water mist and water vapor generated by the previous laser shock absorb and scatter the energy of the next laser beam, resulting in a sharp decrease in the laser energy reaching the absorption layer. Another problem is that with the increase in the frequency of impact, the density of water droplets and mist generated on the optical path increases. In general, it takes about 4 s [7] for water droplets and steam to precipitate by gravity, so that the frequency of laser shock cannot be higher than 0.25 Hz, so that the ability of the laser cannot be fully utilized, and more importantly, the efficiency of laser shock is affected. In order to improve the impact frequency, it is necessary to eliminate the impact of solid particles and water mist on the impact frequency. The U.S. patent and some documents have proposed relevant solutions, as shown in Fig. 3.20. Figure 3.20a is to install a blowing system on both sides of the optical path to divert the impurity water vapor of the optical path from the optical path, and finally collect the water vapor. If the device in Fig. 3.20a continues to be modified, an inclined blowing system is installed under the existing blowing system to accelerate the water vapor falling back; Fig. 3.20b is a light transmitting organic film, which is set closer to the water-constrained layer, so that the space for water splashing is small and it is easy to fall back. The mobile organic film can take away the water vapor attached to it to realize the purification of the optical path.

3.3 Status of Damaged Tape and Absorption Layer TC17 titanium alloy was strengthened by laser shock with a pulse width of 30 ns and a frequency of 0.1 Hz. During laser shock strengthening, TC17 titanium alloy surface shall be pasted with aluminum foil absorption layer. The aluminum foil bonding area shall be 5 mm larger than the laser shock strengthening area, and the

Fig. 3.20 Purification of the laser shock optical path. a Air blowing system is arranged on both sides of the optical path; b organic film. 1—Laser beams; 2—absorption layer; 3—water-confined layer; 4—air outlet; 5—inspiratory port; 6—workpiece surface; 7—water vapor; 8—organic film

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Fig. 3.21 Flow chart of laser shock strengthened aluminum foil absorption layer

aluminum foil shall be pasted smoothly without bubbles, wrinkles, and scratches. A 1 mm thick uniform deionized water-constrained layer is provided to constrain the plasma expansion on the aluminum foil surface. Laser shock strengthening process parameters: square spot side length 4 mm × 4 mm, energy 50 J, and spot overlap rate 8%. The operation flow chart of aluminum foil absorption layer before and after laser shock strengthening is shown in Fig. 3.21 [24]: (1) Clean the laser shock strengthened surface of the sample with cotton soaked in acetone, and then clean the strengthened surface of the sample with alcohol. The cleaning area should be 10 mm larger than the strengthened area; (2) Paste aluminum foil on the strengthened surface of the sample; (3) The aluminum foil absorption layer surface is provided with a continuous and uniform water-constrained layer; (4) Square spot laser shock strengthening sample; (5) Remove the strengthened aluminum foil absorption layer on the sample surface; (6) Clean the strengthened surface of the sample with acetone and alcohol.

3.3.1 No Absorption Layer If the absorption layer of aluminum foil is damaged in a large area or there is no absorption layer during laser shock strengthening, the surface of the base material will be ablated by high-intensity laser pulses. Figure 3.22 shows the morphology of laser shock enhanced induced ablation without absorption layer (energy 50 J, single shock). SEM and backscatter diffraction images obtained in the same area are divided into upper edge image and lower half image. After ablation, the surface roughness of the specimen decreases. The uneven morphology of ablated surface may be caused by phase explosion and transient expansion of plasma. Figure 3.22 shows a large number of microcracks are clearly found on the ablated surface in the backscatter

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Fig. 3.22 Surface morphology and microcrack of laser shock hardened specimen without absorbing layer

diffraction image. There are few microcracks near the center of the spot and there is no directivity, but from the center of the spot to the ablation edge, microcracks become dense and directional. These microcracks may be caused by thermal stress caused by ultra-fast cooling. In addition, the density of microcracks in the ablated area increases with the increase of impact times. When a strong laser pulse ablates the material surface, a large number of free electrons will absorb the laser energy, and the free electrons will heat and transfer the heat to the surrounding lattice, thus producing a casting layer. Figure 3.23 shows the cross section of the casting layer on the material surface after direct laser radiation. Although the ablated surface is uneven, the interface between the casting layer and the base material is neat. It is easy to identify the interface between the casting layer and the base material through the microstructure. The thickness of the casting layer is 1–4 μm. Therefore, the casting surface can be polished to obtain a complete surface. Due to the ultra-fast heating and cooling of laser shock strengthening, the transition zone between the base material and the casting layer is difficult to produce. Most microcracks are only generated in the casting layer. Figure 3.23a generates bifurcations, and Fig. 3.23b discontinues at the interface between the base material and the casting layer. With the increase of laser shock strengthening times, the microcracks in the casting layer become wider, as shown in Fig. 3.23c. A very interesting and meaningful phenomenon was detected, that is, microcracks generated by multiple impacts continued to expand along the interface. Under fatigue load, microcracks induced by direct laser radiation can extend to the base material and cause fatigue fracture. With the microcracks expanding, the casting layer will peel off from the surface of the base material. In addition, there are some air holes in the casting layer, which are not conducive to the fatigue life of the parts.

3.3.2 Slightly Damaged Absorption Layer Although the aluminum foil is used as a sacrificial layer to protect the surface of the sample, ablation black spots appear on the surface of the sample due to local aluminum foil damage, especially the laser shock strengthening lap joint. The

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Fig. 3.23 Cross section of casting layer on material surface after direct laser radiation; a microcrack bifurcation; b microcrack discontinuation; c microcrack widening

aluminum foil damage may be caused by unstable laser intensity, excessive impact strengthening, or irregular aluminum foil state. The aluminum foil is damaged in the overlapping area of four square light spots, as shown in Fig. 3.24a. Four scratches are also found on the aluminum foil surface near the damaged area. The aluminum foil was damaged due to twice impact strengthening and scratch defects. Figure 3.24b, c shows SEM images of ablated black spots on the target surface. The ablated black spots are unevenly distributed and there are melting bumps of different sizes. The EDX analysis results in Fig. 3.25 show that there are 6 ablation areas of main chemical composition of about 1 μm at different locations. The main component of the melted bright bulges (spectrum 1 and 2) is aluminum. In the laser shock strengthening process of overlapping square spot, the former laser pulse damages the integrity of aluminum foil, and the subsequent laser pulse makes the broken aluminum foil fragments melt on the surface of the base material. The dark area (spectrum 3) is determined to be distributed with carbon element, which comes from the adhesive on the back of aluminum foil. The relatively flat area (spectrum 4) may be the surface of the matrix material. Except for a small amount of carbon, the chemical composition is close to the original material before laser shock strengthening. According to the microstructure of ablated black spots on the cross section of the sample, as shown in Fig. 3.24d, the original microstructure of TC17 titanium alloy extends to the surface of the sample, and it is proved that the ablated layer induced by laser radiation is very thin. TC17 titanium alloy matrix material was accidentally protected by molten aluminum layer. The residual stress at the ablated black spot (about 0.5 mm in diameter) measured by the X-ray diffraction method with a diameter of 1 mm is −118 MPa, which is lower than the standard residual compressive stress value of laser shock strengthened surface—500 MPa.

3.3.3 No Damaged Absorption Layer After the absorption layer of aluminum foil was removed, micro-indentation defects were found on the surface of laser shock strengthened samples. The white light interferometry method is used to detect the typical micro-indentation defects on the surface of the laser shock strengthened sample, as shown in Fig. 3.26. A large number

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Fig. 3.24 Damaged aluminum foil and ablated surface; a the aluminum foil in the four light spot overlapping areas is damaged; b surface ablation and scratch defects; c melting uplift; d microstructure of target surface

(a)

(b)

Fig. 3.25 SEM diagram of damaged aluminum foil and ablation surface and EDX results of ablation surface; a SEM diagram; b EDX analysis results

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3 Stability Factors and Safety Protection of Laser Shock Peening

Fig. 3.26 Concave and nick defects on the specimen surface of laser shock peening

of micro-pits and nicks appear on the surface of laser shock strengthened specimens. The average size of concave defect and the width of notch defect are 100 μm, and the notch length is several microns to millimeter. The depth of these defects is 1–2 μm. It can be obviously found that the surface roughness of the material around the defect is improved due to the formation of the defect. Therefore, these micro-indentation may reduce the surface roughness of the laser shock strengthened specimen, and fatigue cracks tend to initiate in the defect area due to the stress concentration effect [25]. When black tape or adhesive paper is used as the absorption layer, it is difficult to find such defects on the surface of laser shock strengthened specimens. By observing the surface of the aluminum foil absorption layer, similar defect patterns were found at the corresponding positions of the aluminum foil. Some defects of aluminum foil may be caused by the tearing force during aluminum foil removal. Other defects of aluminum foil may be caused by laser shock strengthening of the material surface absorption layer. The hardness of these aluminum foil defects is higher than that of the original aluminum foil. The high-speed shock wave will press these defects into the material surface, and then the material surface will form micro-indentation. In order to prevent the surface indentation defects induced by laser shock peening, a mixed melt layer was used for laser shock peening of the material surface. The bottom layer of the mixed melt layer is in direct contact with the surface of the sample. Laser shock peening of the mixed melt layer can avoid bubbles and cavities on the surface of the sample. The intense pulsed laser beam passes through the water-constrained layer and radiates on the aluminum foil surface. As shown in Fig. 3.27a, four overlapping square light spots radiate the ablated area of the aluminum foil surface. The width

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Fig. 3.27 Surface and thickness morphology of laser shock peening aluminum foil. a Ablation area of the aluminum foil surface radiated by four overlapping square light spots; b–e are the local magnified views of figure (a), respectively

of the square spot lap ablation area on the aluminum foil surface is larger than that of the laser spot lap ablation area. Figure 3.27b shows the surface morphology of aluminum foil strengthened by laser shock. Laser shock peening induces hightemperature plasma and laser to melt the aluminum foil surface rapidly. The molten fluid cannot be vaporized immediately due to high-pressure plasma. Therefore, the molten fluid boils at high heat. When the pulsed laser is turned off, the boiling fluid with a large number of bubbles and high-heat droplets will explode. Phase explosion and plasma explosion together cause cavitation and splash on the aluminum foil surface. Ren Naifei [26] and other researchers have shown that the aluminum foil is rapidly vaporized and ionized to form plasma, and the thickness of the aluminum foil gasification layer is about 10 μm. With the increase in impact times, the thickness of aluminum foil decreases gradually. Figure 3.27c–e shows the thickness of aluminum foil cross section corresponding to the marked area in Fig. 3.27a. The original thickness of aluminum foil is 120 μm. The thickness of aluminum foil after single laser shock peening is 80 μm (Fig. 3.27c), the thickness of aluminum foil after two laser shock peening is 55 μm (Fig. 3.27d), and the thickness of aluminum foil after four laser shock peening is 25 μm (Fig. 3.27e).

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3.4 Effect Mechanism of Target The schematic diagram of laser shock peening is shown in Fig. 3.28. After the target surface is polished, a layer of coating (also known as sacrificial layer, conventional organic coating, or metal foil, such as tape, zinc, or aluminum) is applied [27]. The peak power (>GW/cm2 ) and short pulse (ns) emitted by the laser pass through the surface of the transparent constrained layer (water or glass) radiation coating through the focusing lens. The coating absorbs laser energy, vaporizes, and ionizes rapidly in nanoseconds to form high-temperature and high-pressure plasma. The plasma is limited by the confined layer medium, and the explosion generates ultra-highpressure shock wave (GPa), which is transmitted into the target [28]. Restrained layer effect: (1) increase the peak pressure of shock wave; (2) extend the duration of shock wave. When the peak pressure of shock wave exceeds the elastic limit (HEL) of the target in an appropriate time, the surface layer of the target has the following characteristics: (1) dense and stable dislocations or twins; (2) strain hardening occurs on the surface; (3) formation of residual compressive stress layer. Because the elastic deformation energy induced by the shock wave is not less than the plastic or yield deformation energy of the target, the plastic deformation layer of the target is strengthened by laser shock, and the plastic deformation layer inhibits the recovery of elastic deformation energy, resulting in residual compressive stress on the surface of the target. The residual compressive stress changes the distribution of residual stress field inside the target and improves the surface fatigue performance of the material. Therefore, laser shock peening can significantly improve the corrosion and fatigue properties of materials [29, 30]. Because the target surface coating and laser shock peening have a short interaction time with the target, the laser shock peening target ignores the thermal effect and belongs to the cold processing process. Fig. 3.28 Schematic diagram of laser shock peening

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77

3.5 Strengthen Effect Improvement and Safety Protection 3.5.1 Application of Spall Prevention Technology When laser shock peening is used to strengthen thin-walled structures such as blades, due to the small attenuation of the shock wave in the propagation process, when the pressure wave with large amplitude propagates to the back of the structure, the reflection of the wave will occur, and the shock wave will be converted from the pressure wave to the tension wave, resulting in uneven distribution of the residual compressive stress on the surface of the thin-walled structure and the existence of a small compressive or tensile stress state. When the strength of the tension wave is higher than the spall strength of the material, the spalling occurs on the back of the material. In order to prevent back spalling or reduce the strength of reflected tensile wave, the back support can be used. However, for curved structures such as blades, it is difficult to have matching fixed support. In addition, for shock waves, the support body must closely fit with the supported body to have an effect. The method of absorbing wave by liquid on the back of the material can also be used. However, because the acoustic impedance of liquid is very different from that of solid material, the reflection coefficient of shock wave is still very high. Therefore, AVIC Manufacturing Technology Institute has proposed “a spall prevention structure for laser shock peening” (patent publication No.: CN101439440A), which pastes metal aluminum foil and flowing water close to the acoustic impedance of the material on the back of thin-walled structures such as blades [31]. This means the wave-absorbing layer of water flow and aluminum foil are added on the back of the material to absorb the shock wave, so that the shock wave is attenuated at the interface and converted into a tensile wave in the wave-absorbing layer material, so as to effectively reduce the stress wave reflection of the target, weaken the internal coupling strength of the reflected wave and the incident wave, and obtain a uniform residual compressive stress field on the surface. The schematic diagram of laser shock peening of the wave-absorbing layer of the blade is shown in Fig. 3.29. The medium (aluminum foil and water) of the wave-absorbing layer of the blade strengthened by laser shock is shown in Fig. 3.30 [32]. When the blade is strengthened by laser shock peening (there is no wave-absorbing layer on the back of the blade), spalling occurs on the back of the blade, as shown in Fig. 3.31. When only aluminum foil-absorbing layer is bonded on the back of the blade strengthened by laser shock peening, the schematic diagram of the absorbing layer is shown in Fig. 3.29a, and the aluminum foil on the back of the blade shows spalling, as shown in Fig. 3.32. Even though the process parameters of laser shock peening are adjusted, the aluminum foil on the back of the blade strengthened by laser shock still shows bulge, as shown in Fig. 3.33 [13]. When aluminum foil and water wave-absorbing layer are bonded on the back of blade, the schematic diagram of wave-absorbing layer is shown in Fig. 3.29b, and the morphology of strengthened blade back is shown in Fig. 3.34. The aluminum foil on the back of blade is flat and

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Fig. 3.29 Schematic diagram of wave-absorbing layer of blades strengthened by laser shock peening: a only aluminum foil is bonded on the back of the blade; b water flow and aluminum foil are on the back of the blade

Fig. 3.30 Laser shock peening medium of wave-absorbing layer of blade (aluminum foil and water)

smooth, the reflection effect of shock wave is weakened, and the transmission effect is enhanced, effectively improving the effect of laser shock strengthened blade. From the interface force balance and interface continuity conditions of shock wave [33], σ I + σ R = σT

(3.3)

U I + U R = UT

(3.4)

From the momentum conservation relation of shock wave,

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Fig. 3.31 Lamination phenomenon of blade without absorbing layer strengthened by laser shock peening: a obverse; b side

Fig. 3.32 Aluminum foil spalling on the back of blade

Fig. 3.33 Aluminum foil protrusion on the back of blade strengthened by laser shock peening

Fdt = d(mU P )

(3.5)

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Fig. 3.34 Morphology of aluminum foil on the back of blades strengthened by laser shock peening. (aluminum foil and water)

σ Adt = ρ Ad xU P σ =ρ

dx UP dt

σ = ρCU P

(3.6) (3.7) (3.8)

Combining Formulas (3.3), (3.4), (3.5), (3.6), (3.7) and (3.8), we can get σT 2ρ B C B = σI ρ AC A + ρB C B

(3.9)

σR ρB C B − ρ AC A = σI ρ AC A + ρB C B

(3.10)

Acoustic impedance formulation: Z = ρC where σI UI σR UR σT UT F m UP σ ρ C

target stress caused by incident wave; particle velocity caused by incident wave; target stress caused by reflected wave; particle velocity caused by reflected wave; target stress caused by transmitted wave; particle velocity caused by transmitted wave; loading of shock wave; target mass; particle velocity; internal stress of target; target density; acoustic wave velocity.

(3.11)

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When a laser-induced shock wave propagates from high-impedance material (blade) ρ A C A to low-impedance material (aluminum foil) ρ B C B , ρ B C B < ρ A C A , the reflected wave symbol is opposite to the incident wave symbol; when it is incident to the free surface, ρ B C B = 0, σσTI = 0, σσRI = −1; when incident to the rigid wall, ρ B C B = ∞, σσTI = 2, σσRI = 1. Parameters of laser shock strengthened blades: titanium alloy density is 4500 kg/m3 , acoustic wave velocity is 4000 m/s, aluminum foil density is 2784 kg/m3 , acoustic wave velocity is 5370 m/s, water density is 1000 kg/m3 , acoustic wave velocity is 1461 m/s, air density is 1.29 kg/m3 , and acoustic wave velocity is 340 m/s. When there is no wave-absorbing layer on the back of titanium alloy blade, the reflection condition of high-pressure shock wave on the free surface of titanium alloy is titanium alloy and air, calculated σσRI = −1, high-pressure incident shock waves are all converted into reflected tensile wave strength. The tensile wave strength is greater than the spall strength of titanium alloy, and spall occurs in titanium alloy, as shown in Fig. 3.31. When only aluminum foil-absorbing layer is bonded on the back of titanium alloy blade, the propagation and reflection of shock wave are shown in Fig. 3.25a. The reflection condition of shock wave interface is titanium alloy and aluminum foil, aluminum foil and air, shock wave at the interface between titanium alloy and aluminum foil: σσRI = −0.09, σσTI = 0.91, 91% of the incident shock wave pressure is transmitted into the aluminum foil, and the reflected wave intensity of titanium alloy is low, so as to avoid spalling of titanium alloy. At the interface between aluminum foil and air due to incident shock wave: σσRI = −1, so the strength of the reflected wave in the aluminum foil is 91% of the incident shock wave strength. When the strong reflected tensile wave in the aluminum foil reaches the interface between titanium alloy and aluminum foil, the bubble phenomenon occurs in the aluminum foil because the tensile wave strength is greater than the adhesive strength between the aluminum foil and the back of titanium alloy, as shown in Figs. 3.32 and 3.33. When the wave-absorbing layer on the back of the titanium alloy blade is a singlelayer aluminum foil and uniform deionized water is provided, the propagation and reflection of the shock wave are shown in Fig. 3.35b, and the reflection conditions of the shock wave at the interface are titanium alloy and aluminum foil, aluminum foil and water, Shock wave at the interface between titanium alloy and aluminum foil: σR = −0.09, σσTI = 0.91, 91% of the incident shock wave pressure is transmitted σI into the aluminum foil, and the reflected wave intensity of titanium alloy is low, so as to avoid spalling of titanium alloy. The shock wave is at the interface between aluminum foil and water: σσRI = −0.82, σσTI = 0.18. The transmitted wave intensity is taken away by flowing deionized water without reflection, but the reflected wave intensity in the aluminum foil is 75% of the shock wave intensity, which reduces or eliminates aluminum foil bubble, as shown in Fig. 3.34.

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Fig. 3.35 Propagation and reflection of shock wave of blades strengthened by laser shock peening. a Only aluminum foil is bonded on the back of the blade; b the blade is bonded with aluminum foil and water

3.5.2 Methods to Enhance Strengthening Effect The effect of laser shock peening on the mechanical properties of metal materials is closely related to the area of laser-induced plasma. Generally, when water or glass is used as the constrained layer, the number and volume of laser-induced plasma can be enhanced by increasing the pulse laser energy to enhance its shock wave effect. However, it is difficult to obtain high-energy pulsed laser, and the electro-optical conversion efficiency is low, only about 5%, while the efficiency of laser-induced plasma conversion into mechanical energy is 10%, and only 0.5% energy is converted into the energy with improved mechanical properties of metal materials. In addition, the domestic commonly used high-pulse energy laser shock strengthening equipment has a low working frequency or pulse energy, which is difficult to meet the needs of the rapid development of laser shock peening technology. AVIC Manufacturing Technology Institute has proposed a laser shock peening method and device for external electric field peening (patent publication No.: CN103911505A) [34], which uses electrode plate 2 to add a strong electric field 3 coaxial with the laser shock direction (or at an angle of 30–60°) around the laser shock peening-induced plasma cluster 6. As shown in Fig. 3.36, since the plasma induced by pulsed laser is in a nonsteady state, after the external electric field is added, it will accelerate electrons and positive ions with different initial energies, increase their collision probability with molecules and particles, and generate more active groups with different energies, thus increasing the number and volume of laser-induced plasma, so as to enhance the explosion shock wave effect of plasma. The externally reinforced electric field is equivalent to the secondary restraint except for the material of the restraint layer.

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Fig. 3.36 Structure of laser shock peening device with external enhanced electric field. 1—Laser; 2—electrode plate; 3—electric field; 4—sprinkler head; 5—water layer; 6—absorption layer; 7—workpiece; 8—insulation board; 9—clamp; 10—manipulator; 11—cathode; 12—high-voltage power supply; 13—anode; 14—insulation board; 15—optical platform 16—synchronization controller; 17—laser equipment

3.5.3 Light Reflection and Explosive Fragmentation of Safety Protection Laser shock peening uses a strong pulsed laser as an energy source, and there is a light reflection in the bound medium propagation. Explosive fragmentation of hard confinement layers under the action of strong shock waves, all of which require effective protection [9]. The first is the ablative effect of reflected light. When the laser is a focused rate density of GW/cm2 or more, the beam is partially reflected in the confinement layer. For optical glass, the single-sided reflectivity is generally 4%, if it is the double-sided reflection, the power density of reflected light is more than 8 × 107 W/cm2 . Assuming that the reflecting surface of the bound medium is flat, whether the reflected light is focused or diverging depends on whether the reflecting surface is in the out-of-focus or in-focus position of the incident light. If in the focus position, the reflected light is further focused, and the power density is further enhanced; if in the out-of-focus position, the reflected light is further dispersed, and the power density is weakened, as shown in Fig. 3.37. Due to the laser shock intensification, the focal length of the final focusing lens is larger, generally around lm, and the laser man angle is not large. The strengthening or weakening effect of reflected light power density is not very obvious, but the reflected light power density at a certain distance is still in the order of 107 W/cm2 , posing a threat to the surrounding devices, and even to the processing site operators. Thus, before LSP, He–Ni light should be used to align the shock position and then slightly rotate the specimen to be shocked to adjust the reflected light to a safe position.

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Fig. 3.37 Focus and divergence of reflected light (black thick line for the human light). a Focusing of reflected light; b divergence of reflected light

The reflected light poses a threat at the same time, as a binding medium of glass in the blast effect is quickly broken into pieces by the shock wave, splashing around, sputtering direction is mainly normal to the target surface direction, the speed of the spattered particles is large, and the surrounding personnel and optical equipment have a certain threat or destructive effect. Whether it is reflected light or spattered glass fragments, adding a protective shield in the shock area can avoid the threat to the surrounding personnel or equipment, while the laser light path location cannot be arranged to protect the device. It can not only make the normal direction of the target surface point to the focusing lens, but also can reduce the damage of the sputtering glass fragments to the lens. However, the non-mainstream direction of the glass fragments can still damage the lens on the transmission film, so that the absorption of the lens to the laser increased between the lens and the target with protective glass and will further intensify the reflection of the laser, reducing the energy to reach the target.

References 1. Wang J, Zou SK, Tan YS (2005) Application of laser shock processing on turbine engines (in Chinese). Appl Laser 25(01):32–34 2. Cao ZW, Zou SK, Liu FJ et al (2008) Laser shock treatment of 1CrlNi2W2MoV stainless steel (in Chinese). Chin J Lasers 35(2):316–320 3. Cao ZW, Zou SK, Che ZG et al (2011) Research on process control of laser shock peening for aeroengine blade (in Chinese). Aeroengine 37(1):60–62 4. Perozek PM, Lawrence WL (2005) Reducing electromagnetic feedback during laser shock peening. USA, US006917012B2 5. Westley JA, Jones D, Andrews I (2006) Laser shock peening. USA, US007137282B2 6. Cao ZW, Zou SK (2008) A laser beam shaping quintile lens and quarter lens (in Chinese). CN, CN101256287A 7. Clauer AH, Toller SM, Dulaney JL (2003) Beam path clearing for laser peening. USA, US006521860B2 8. Lawrence WL, Perozek PM (2004) Reduced mist laser shock peening. USA, US006713716B1

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9. Wang J, Zou SK (2002) Application research of laser shock treatment technology (in Chinese). Appl Laser 22(2):223–226 10. Gong SL, Dai FZ, Zhang YK, et al (2013) A method of obtaining stable water film in blisk by laser shock treatment (in Chinese). CN, CN103205546A 11. Zou SK, Che ZG, Cao ZW (2012) A water/optical coaxial device for laser processing shock (in Chinese). CN, CN102505065A 12. Dykes SE, Clauer AH, Dulaney JL, et al (2003) Overlay control for laser peening. USA, US006548782B2 13. Che ZG, Gong SL, Zou SK, et al (2010) Investigation on the key techniques of confined medium and coating layer for laser shock processing on aeroengine blade. Rare Met Mater Eng 527–530 14. Ilardi JM, Schwartzkopf G (1996) Cleaning wafer substrates of metal contamination while maintaining wafer smoothness. USA, EP0690483B1 15. Trantow RL, Bashyam M (1999) Determination of Rayleigh wave critical angle. Cincinnati, US005987991A 16. Trantow RL. Ultrasonic multi-transducer rotatable scanning apparatus and method of use thereof. Cincinnati, US005974889A 17. Cao ZW, Zou SK, Zhang XB et al (2007) Study of water confinement regime: application to laser peening (in Chinese). Appl Laser 27(06):461–464 18. Zhang YK (1995) Study on improving fatigue life of aviation materials by laser shock strengthening (in Chinese). Nanjing University of Aeronautics and Astronautics, Nanjing 19. Sowers BL, Birkhoff RD, Arakawa ET (1972) Optical absorption of liquid water in the vacuum ultraviolet. J Chem Phys 57(1):583–584 20. Kruusing A (2004) Underwater and water-assisted laser processing: part 1-general features, steam cleaning and shock processing. Opt Lasers Eng 41(2):329–352 21. Peyre P, Fabbro R, Berthe L et al (1996) Laser shock processing of materials, physical processes involved and examples of applications. J Laser Appl 8(3):135–141 22. Berthe L, Fabbro R, Peyre P et al (1999) Wavelength dependent of laser shock wave generation in the water-confinement regime. J Appl Phys 85(11):7552–7555 23. Wu B, Shin YC (2006) Laser pulse transmission through the water breakdown plasma in laser shock peening. Appl Phys Lett 88(4):41116 24. Cao ZW (2013) Surface defects produced by laser shock peening with aluminium foil ablative layer (in Chinese). Rare Met Mater Eng (s2):217–221 25. Thompson SR, Ruschau JJ, Nicholas T (2001) Influence of residual stresses on high cycle fatigue strength of Ti–6Al–4V subjected to foreign object damage. Int J Fatigue 23(1):405–412 26. Ren NF, Zhang YK (1997) Heat transmission in metal—material strengthening by laser shock (in Chinese). Appl Laser 03:105–108 27. Zhang WW, Yao YL Modelling and simulation improvement in laser shock processing. In: Proceedings of ICALEO, Section A, pp 59–68 28. Wang Y, Kysar JW, Yao YL (2008) Analytical solution of anisotropic plastic deformation induced by micro-scale laser shock peening. Mech Mater 40(3):100–114 29. Zhang YK, Zhou JZ, Ye YX (2004) Laser processing technology (in Chinese). Chemical Industry Press 30. Fairand BP, Wilcox BA, Gallagher WJ et al (1972) Laser shock-induced microstructural and mechanical property changes in 7075 aluminum. J Appl Phys 43(9):3893–3895 31. Zou SK, Wu JF, Gong SL, et al (2016) Anti-delamination method and device for laser impact strengthening thin-walled structure (in Chinese). CN, ZL105483359B 32. Che ZG, Gong SL, Cao ZW et al (2011) Theory analysis and experiment investigation of laser shock processing on titanium alloy blade (in Chinese). Rare Met Mater Eng S4:235–239 33. Zhang QM, Liu Y, Huang FL (2006) Dynamic behavior of materials (in Chinese). National Defense Industry Press, Beijing 34. Zou SK, Song W, Che ZG, et al (2014) Laser impact strengthening method and device for external reinforcement electric field (in Chinese). CN, CN103911505A

Chapter 4

Numerical Analysis of Mechanical Effects of Laser Shock Peening

Laser shock peening consists of 3 stages. The first stage is the light-wave conversion, which consists of the conversion of the laser beam into a plasma shock wave. In this process, fundamental scientific issues such as “kinetic process and energy conversion mechanism of laser-induced plasma/shock wave”, “detection of the whole process of shock wave ignition propagation and extinction” and “laser-induced nonlinear strong acoustic field” are involved. The second stage is wave-solid–liquid coupling, in which the laser-induced plasma shock wave interacts and couples with solids and liquids. This process involves the “theory of dynamic plastic deformation at ultra-high strain rates”, “real-time measurement of dynamic deformation processes in solids” and “microscopic deformation mechanisms and micromechanics of strong stress waves at ultra-high strain rates “The basic scientific problems. The third stage is performance growth, where the material properties are significantly improved by the laser shock wave. This process involves fundamental scientific issues such as “ultrahigh strain rate strengthening theory”, “physical mechanism of wave-solid–liquid coupling and performance growth under repeated rapid loading of shock waves” and “structural performance optimization and reliability”. Only with the mastery of these basic theories can laser technology be truly applied to the production field and control the safety and reliability. However, there are very few academic papers on the basic theoretical research in this area compared to technical and technological research, and many theoretical and physical mechanisms are not yet clear. For example, in the study of laser-induced shock waves, the laser/plasma energy conversion mechanism and the kinetic process of laser-induced shock waves have been studied abroad, and the effects of laser parameters on shock waves have been obtained. Domestic studies have focused on the generation, measurement, and propagation characteristics of laser-driven shock waves in materials, without in-depth discussions involving the physical mechanisms of energy conversion.

© National Defense Industry Press 2023 S. Zou et al., Laser Shock Peening, https://doi.org/10.1007/978-981-99-1117-2_4

87

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4 Numerical Analysis of Mechanical Effects of Laser Shock Peening

In the study of the mechanism of laser shock peening, a lot of research has been conducted abroad starting from the microstructure, and the dislocation kinetic theory has been adopted to elaborate the strengthening mechanism and explain the mechanism of phase change under the action of shock waves. In contrast, domestic research is mainly experimental, and the mechanisms of reinforcement are not well studied.

4.1 Physical Model 4.1.1 Fabbro Physical Model In 1993, R. Fabbro conducted a theoretical study of shock waves in the LSP process under the constrained model. The plasma expansion is shown in Fig. 4.1, and the following assumptions are made: (1) Uniform distribution of laser energy and uniform heating of the material surface throughout the spot range. (2) Both the confining layer and the target are isotropic homogeneous substances, and the thermophysical properties are constant. (3) Considering the plasma as an ideal gas. (4) Plasma expansion in the axial direction only. Part of the absorbed laser energy is used to increase the internal energy of the interfacial plasma, and the other part is used to do work. Let I (t) be the power density of the incident laser absorbed by the absorber layer at moment t, then the absorbed laser energy is ∫ W = 0

Fig. 4.1 Plasma expansion

t

I (t) · dt

(4.1)

4.1 Physical Model

89

In a time interval of dt, the plasma thickness will increase by dL, doing work for W1 = P(t) · dL and internal energy increase for W2 = d [Ei (t) · L]. According to the principle of conservation of energy, then ∫

t

W = W1 + W 2 =

I (t) · dt = P(t) · dL + d [Ei (t) · L]

(4.2)

0

Assuming that the plasma thermal energy ET (t) is only a fraction of its internal energy Ei (t), Let the constant α (often taken as 0.1 to 0.3), that is ET (t) = α · Ei (t). Eventually, the pressure of the shock wave under the water confining layer can be derived as a function of the laser power density and the estimation of other parameters. Peak pressure versus laser power density ( P(t) = 0.01

α 2α + 3

) 21

1

1

· Z 2 · I02

(4.3)

where α—the coefficient of conversion of internal energy into thermal energy part, generally taking a value of 0.2 to 0.3. Z-folded acoustic impedance, the value of the absorption protection layer, and the binding layer of acoustic impedance together to determine, (The acoustic impedance of the deionized water bound layer is 0.165 × 106 g/cm2 s, and the acoustic impedance of the aluminum foil tape of the absorption protection layer is 1.6 × 106 g/cm2 s). I 0 —laser power density. Surface plastic strain: ( ) P −2 HEL −1 εp = 3 λ + 2 μ HEL

(4.4)

where μ, λ—Lamé parameters. HEL—elastic limit. Residual compressive stress layer depth: ( LP =

Ce Cp τ Ce − Cp

)(

P − HEL 2HEL

)

where C e —the speed of propagation of elastic waves in the material. C p —the speed of propagation of plastic waves in the material. Surface residual compressive stress: ) √ )( ( 1+v 4 2 LP + σ0 1 − σsurf = σ0 − μεp (1 + v) √ 1−v π r 2 2/ = 1/ + 1/ Z2 Z Z1

(4.5)

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4 Numerical Analysis of Mechanical Effects of Laser Shock Peening

/ I0 = E π r 2 τ

(4.6)

where E—the output energy of the laser. σ0 —initial stress of the material. v—Poisson’s ratio. The Fabbro model has the following drawbacks. ➀ The coefficient is related to the nature of the absorbing layer and the state of the confining layer, as well as to the nature of the laser pulse such as pulse shape and pulse width. ➁ plasma at high pressures of the order of GPa, the ideal gas will no longer be a good approximation and its effect will be equivalent to a reduction and the calculated values will be smaller. ➂ The attenuation of the shock wave is not considered; ➃ The transverse expansion effect of the plasma is not taken into account.

4.1.2 Modified Physical Model Absorption layer effect on the material surface during laser shock peening of metallic materials: high-pressure plasma formation by absorption layer vaporization; surface ablation of absorption layer protection material. During laser shock peening, the absorbed layer on the material surface is not completely vaporized, as shown in Fig. 4.2. Therefore, there is a greater error in the calculation of the peak shock wave pressure using the Fabbro model [1]. Define the physical state of the gasification residual absorber layer as ➂, as shown in Fig. 4.2c. The total impedance of the shock wave is calculated as [2]. 1 3 1 1 = + + Z Z1 Z2 Z3

(4.7)

where Z1 —target shock wave impedance. Z2 —shock wave impedance of the confining layer. Z3 —residual absorbing layer shock wave impedance. Plasma thickness calculation equation: ∫t L(t) =

[u1 (t) + u2 (t) + u3 (t)]dt

(4.8)

0

The shock wave pressure equation is [3]: P = ρDi Ui = Zi ui (i = 1, 2 . . . 3) Combining Eqs. (4.7) to (4.9), it is obtained that:

(4.9)

4.1 Physical Model

91

Fig. 4.2 Schematic diagram of laser shock-enhanced plasma formation a pulsed laser beam radiation; b shock wave formation within the target; c residual absorption layer

3 dL(t) = P(t) dt Z

(4.10)

2 αEi (t) 3

(4.11)

Consider the energy equation: P(t) = I (t) = P(t)

dL(t) d [Ei (t)L(t)] + dt dt

(4.12)

Combining Eqs. (4.10) to (4.12), it is obtained that: (

Z Z + 3 2α

)(

dL(t) dt

)2 +

d 2 L(t) Z L(t) = I (t) 2α dt 2

(4.13)

Combining Eqs. (4.10) to (4.13), it is obtained that: / P = 0.01

√ 2α ZI0 (GPa) 3(2α + 3)

(4.14)

where I (t) = I0 —laser power density (square spot, uniform energy distribution). Therefore, a new shock wave pressure model is derived, which is consistent with the actual situation and has higher calculation accuracy than the Fabbro model. The I 0 in this model is the square spot true laser power density, while the I 0 in the Fabbro model is the circular spot peak laser power density.

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4 Numerical Analysis of Mechanical Effects of Laser Shock Peening

4.2 Numerical Analysis Steps The very short process from the arrival of a strong pulse at the material surface to the end of the shock wave effect generated by this strong pulse makes the experimental study of the laser shock peening process extremely difficult. Processes such as high temperature and high-pressure plasma formation, stress wave propagation and decay, ultra-high strain rate plastic deformation of shocked materials, and three-dimensional stress evolution in the reinforced surface layer are all relevant to the performance of the reinforced material, but these processes are difficult to analyze from experimental methods. Therefore, laser shock peening finite element simulation techniques have been rapidly developed in the last decade. Finite element simulation techniques allow not only to analyze the above dynamic processes, but also to optimize the residual compressive stress distribution by analyzing material properties, workpiece geometry, laser beam quality, and laser shock peening parameters. In the actual production process, laser shock peening finite element simulation technology has become an important tool for designing and optimizing the laser shock peening process. Most scholars have only studied the simulation calculations for the process of shock wave loading on the material and forming and generating plastic deformation and residual stress fields, while the process of high temperature and high-pressure plasma/shock wave generated by the interaction of strong pulsed laser with the material surface is less studied. At present, the software applied to laser shock peening finite element simulation is mainly ABAQUS, ANASYS, and other large commercial finite element software. In recent years, some scholars have also developed finite element software for the simulation of plasma/shock wave formation processes, such as LSPSIM and HELIOS.

4.2.1 Finite Element Analysis Method The time integration algorithm of ABAQUS software is divided into explicit and implicit methods. For the two time-integrated processes, explicit and implicit methods, the equilibrium is defined in terms of the external force P, the cell internal force I, and the nodal acceleration: M u¨ = P − I

(4.15)

where M—mass matrix. When determining the unit internal forces, two methods can be used to solve for the nodal accelerations, and the same unit calculations are to be applied. The biggest difference between them is the way they calculate the acceleration of the nodes. The explicit method is the state at the end of the time period and depends only on the

4.2 Numerical Analysis Steps

93

displacement, velocity, and acceleration at the beginning of this time period. The implicit method is based on a fully Newtonian iterative solution method that looks for a satisfying dynamic equilibrium at the end of the time period and solves a series of linear systems of equations to calculate the displacements. Since the shock wave loading time of laser shock peening is very short, the finite element analysis of the laser shock peening process must consider the stress wave propagation inside the material and its decay, so that the material accurately calculates the stress–strain on the material during the stress wave propagation and thus predicts the residual compressive stress field. Because the laser shock peening process is a highly nonlinear process, the display method is usually used to solve the highly nonlinear process in which the stress–strain occurs instantaneously. If the implicit method is used to solve this one process is very time-consuming and there may be cases where it does not converge and the computation cannot continue. For the oscillation process after the end of stress wave loading, until a stable residual stress field is generated, both the display method and the implicit method can be used to solve it, but the implicit method is generally used. Because the oscillatory process, which generates a stable residual stress field, is longer than the highly nonlinear process in which stress–strain occurs, it is also very time-consuming to solve using the display method. To understand the propagation of stress in the finite element model when the display method is applied, the propagation of stress waves along a rod model containing three cells is considered here, as shown in Fig. 4.3. In the first time increment segment, nodes 2 and 3 do not move because there is no force acting on them. In the second time increment period, the internal force is obtained from the stresses in cell ➀ and applied to the nodes connected to cell ➀. These cell stresses are subsequently used to calculate the dynamic balance equations for node 1 and node 2. This process continues, and by the beginning of the third time increment segment, stresses already exist in cells ➀ and ➁, as well as forces, exist in nodes 1, 2, and 3. This process continues until the end of the total analysis time.

4.2.2 Numerical Model Parameter Setting 1. Geometric model Depending on the experimental conditions and the need for results, the finite element geometry model can be built as a two-dimensional model and a three-dimensional model. When the model geometry is symmetrical and the laser spot shape is circular, the three-dimensional laser shock peening model can be reduced to a twodimensional model. The 3D model is suitable for essentially all cases of laser shock peening, such as square laser spots, lapped spots, geometrically asymmetric parts, etc. However, the 3D model costs more calculation than the 2D model. Therefore, the geometric model should be selected reasonably according to the actual situation.

94

4 Numerical Analysis of Mechanical Effects of Laser Shock Peening

Initial stage

First time incremental segment

Second time incremental segment

Third time incremental segment

Fig. 4.3 Stress propagation in the rod model

2. Material model The strain rate in the surface layer of the material during laser shock peening is up to 106 /s or more, so the material properties are very different from those under conventional conditions. In order to obtain more accurate simulation results, the dynamic yield strength, strain hardening, and strain rate of the material must be taken

4.2 Numerical Analysis Steps

95

into account in the laser shock peening finite element model [4]. Currently, there are two main approaches to laser shock peening finite element material model definition. dyn One is to define the yield strength of a material as the dynamic yield strength σY , the value of which is the yield strength at ultra-high strain rates (close to 106 /s). Second, the Johnson–Cook material model [5, 6] is used, and after neglecting the temperature influence factor, the expression of this model can be simplified as follows: [

(

ε˙ σ = (A + Bε ) 1 + C ln ε˙ 0

)]

n

(4.16)

where A, B, C, and n—material constants. ε—equivalent plastic strain of the material. ε˙ —strain rate. ε˙ 0 —strain rate at quasi-static. 3. Shockwave pressure Shock wave pressure is an important boundary condition for finite element simulation analysis of laser shock peening. There are two main ways to obtain the shock wave pressure generated by the laser shock intensification process. ➀Waveform of shock waves generated by laser shock intensification process using PVDF piezoelectric sensor or VISAR, etc. ➁The peak shock wave pressure is estimated directly using existing shock wave pressure estimation models (e.g. Fabbro model) and the pulse width of the shock wave is predicted based on the constrained layer conditions, while the shock wave history waveform needs to be obtained empirically or replaced using triangular waveforms. 4. Explicit dynamic solution time and incremental step settings The dynamic analysis module is used to simulate the elastic–plastic deformation of the target surface induced by ultra-high speed and short pulse laser shock, which requires multiple small time increment steps for efficient computation. The time increment step Δt has a great influence on the convergence of the simulation and the accuracy of the results. If the time increment step Δt is larger than the stability limit Δt stable , the simulation will be unstable, which leads to an unbounded solution. Based on a simple cell-to-cell estimation method, the stability limit equation is calculated as follows [7]: Δtstable =

Le Cd

(4.17)

E ρ

(4.18)

/ Cd = where L e —minimum cell length. C d —laser shock wave velocity inside the material.

96

4 Numerical Analysis of Mechanical Effects of Laser Shock Peening

E—modulus of elasticity of the material. ρ—material density. 5. Set damping parameters The steps of numerical analysis of laser shock peening targets are shown in Fig. 4.4. ➀ Elastic–plastic dynamic analysis of simulated laser shock-reinforced targets. ➁ Import the elastic–plastic dynamic analysis results into ABAQUS/Standard for static stress–strain analysis to obtain the static equilibrium residual stress field, where the dynamic analysis time must be longer than the laser shock wave pulse pressure duration to obtain the saturated plastic deformation of the laser shock peening reinforced target material. The simulation of residual stresses and strains in the target material reinforced by multiple laser shock peening is as follows: the residual stresses and strain fields of the laser shock peening target were simulated according to Fig. 4.4, and the results of the previous static analysis of the residual stresses and strain fields were used as the initial values for the next dynamic analysis of the laser shock peening target, and finally, the simulation results were post-processed in the ABAQUS/Viewer module.

Input file editing

ABAQUS/Explicit Dynamic Analysis ABAQUS/Explicit Static Analysis

Residual stress/strain field surface microplastic deformation Fig. 4.4 Finite element simulation flow of laser shock-reinforced target

4.3 Numerical Analysis of Circular Laser Spot

97

4.3 Numerical Analysis of Circular Laser Spot 4.3.1 Finite Element Model The use of a reasonable material model is a necessary condition for the success of numerical simulations. Laser shock peening causes plastic deformation of metallic materials with strain rates exceeding 106/s. At such high strain rates, the intrinsic relationship between the dynamic mechanical properties of the material and the strain rate is correlated. When choosing a material model, it is logical to choose a plastic material model related to strain rate, but it is difficult to introduce strain rate effects because there is less data on material properties at different strain rates, and the material is generally a variable strain rate process during the kinetic response. In view of the complexity of the strain rate and the available simulation reports, an ideal elastic–plastic material model independent of the strain rate is chosen. Here it is assumed that the laser shock deformation is a one-dimensional strain deformation under the action of a shock stress wave and the material obeys the Vons Mises yield criterion. The mechanical property parameters of TC4 titanium alloy used in the ideal elastic–plastic model are shown in Table 4.1. The dynamic yield strength σ y of TC4 titanium alloy can be based on the one-dimensional strain-elastic phase instantonal relationship: σy = HEL

(1 − 2ν) (1 − ν)

(4.19)

where HEL—elastic limit. N—Poisson’s ratio. 1. Geometric model and mesh There are two types of objects to be simulated in this section: ➀ single-spot shock; ➁ multi-spot lap single-sided shock. The shock area of the single-spot shock model is located in the center of the target material, the shock load (circular spot) and the shape of the target material have axisymmetry, so a quarter of the three-dimensional model can be selected, and the three-dimensional reduced unit (finite unit C3D8R Table 4.1 TC4 titanium alloy mechanical properties parameters

Material properties Density

ρ/(kg/m3 )

Numerical value 4500

Modulus of elasticity E/(GPa)

110

Poisson’s ratio ν

0.342

Elastic limit HEL/(GPa)

2.8

Dynamic yield strength σ y /(GPa)

1.345

98

4 Numerical Analysis of Mechanical Effects of Laser Shock Peening

Fig. 4.5 LSP simulation finite element model. a Model A; b Model B

Table 4.2 Finite element model mesh structure Finite element model

Finite unit type

Infinite unit type

Total number of units

Unit size/(mm)

A

C3D8R

CIN3D8

12,868

0.2

B

C3D8R

CIN3D8

43,817

0.2

and infinite unit CIN3D8) is selected for simulation calculations. The C3D8R is an elasto-plastic, nonlinear unit used to simulate the residual stress field distribution in the laser shock region and its stress wave affected zone. CIN3D8 is an infinite cell, this cell has only elastic material behavior, and the stress wave propagates to infinity in this cell. Class B model does not have axisymmetry, take one-half of the whole model for simulation, and still choose 3D finite reduction unit C3D8R and infinite unit CIN3D8 for simulation calculation. The finite element model dimensions are shown in Fig. 4.5. In Fig. 4.5, the radius of all laser shock zone loading areas are r = 3 mm, the thickness of the model is 4 mm, where the finite element area is 3 mm and the edge of the shock zone is 3 mm from the infinite cell. The grid structures of models A and B are shown in Table 4.2. 2. Laser-induced shock wave loading Loading of the material by the laser shock wave is a highly nonlinear process, with strain rates above 106 /s for the target material. The peak pressure of the shock wave is obtained by the Fabbro estimation model, which assumes that the plasma is an ideal gas and that the plasma expands only in the axial direction, neglecting the transverse expansion, and the peak pressure of the laser-induced shock wave is calculated in Eq. (4.14). During the finite element simulation, the spatial distribution of the shock wave is assumed to be uniform, i.e. the shock area is uniformly loaded. The variation curve

99

Shockwave relative pressure amplitude

4.3 Numerical Analysis of Circular Laser Spot

Time/ns Fig. 4.6 Shock wave relative pressure amplitude-time curve

of shock wave pressure amplitude with time is simplified as Fig. 4.6. The *AMPLITUDE, DEFINITION = TABULAR, TIME = STEP TIME, VALUE = RELATIVE options should be used in ABAQUS/Explicit to define the pressure amplitude-time curve of the shock wave. 3. Boundary conditions and others After the shock wave is loaded, the stress wave inside the material will be reflected on the free surface of the side or bottom, so the infinite cell is used at the edge of the finite element model, and the infinite cell is introduced as a “static” boundary condition to avoid the influence of the reflected stress wave as much as possible.

4.3.2 Dynamic Stress–Strain Analysis of Shock Wave Loading Process 1. Model energy allocation During the laser shock wave loading process, the total energy W t of the shock wave acting on the target is converted into kinetic energy W k , internal energy W i, and dissipative energy W v . The internal energy W i contains three types of energy, the elastic storage energy W e , the plastic dissipation energy W p , and the pseudo-strain energy W a . Figure 4.7 shows the energy change of the model during the shock wave loading and stress wave propagation generated at a laser power density of 5 GW/cm2 (Model A). At the end of the shock wave loading, the energy injection into the target is completed, and the external energy input is about 300 mJ. The elastic storage energy of the target W e is about 130 mJ at the beginning of the laser shock loading, and its value decreases sharply to less than 10 mJ after 4000 ns, while the plastic dissipation energy Wp increases significantly after 200 ns and then stably remains

4 Numerical Analysis of Mechanical Effects of Laser Shock Peening

Energy/J

Energy/J

100

Time/ns

Time/ns

Fig. 4.7 Energy change of laser shock peening model

at 75 mJ. Meanwhile, the pseudo-strain energy has been kept near 0 mJ during the whole analysis, which indicates that the grid results of the model are reasonable. 2. Analysis of model dynamic loading process

Dynamic Stress

Dynamic Stress

A single-spot shock is used as an example to analyze the propagation and attenuation of the elastic–plastic stress wave inside the material. Figure 4.8 shows the dynamic stresses in the center of the shock zone along the depth direction. Figure 4.8a and b show the dynamic stresses σ z , σ x , respectively (σ x is the stress at the material surface along the radius of the circular shock zone, σ y is the stress at the material surface perpendicular to the radius of the circular shock zone, and σ z is the stress at the material surface perpendicular to the radius of the circular shock zone). As can be seen from Fig. 4.8, the stress wave continuously decaying phase is the elasto-plastic loading phase, and it can be seen that the stress wave is continuously decaying at the beginning of loading, with the stress in the Z direction (longitudinal stress) decaying from 5.0 to 2.8 GPa and the stress in the X (Y) direction (lateral stress) decaying from 3.655 GPa to about 1.455 GPa. This result satisfies the Von Mises quasi-measure, and the relationship between longitudinal and lateral stresses

Distance from surface/mm

Distance from surface/mm

Fig. 4.8 Stress wave propagation and attenuation in the direction of the center depth of the laser shock area a dynamic stress σ z ; b dynamic stress σ x

4.3 Numerical Analysis of Circular Laser Spot

101

is as follows: σZ = σY = σX + Ys

(4.20)

The relationship between longitudinal stress and plastic strain is as follows: ( σx =

) ) ( 3 λ λ + μ εP − Ys 1 + 2 2μ

(4.21)

When the longitudinal stress amplitude decays to 2.8 GPa, after that the stress wave amplitude is maintained at a certain amplitude level, and the material only produces elastic deformation at this stage. The relationship between the longitudinal and lateral stresses satisfies: σX = σY =

v σZ 1−v

(4.22)

The relationship between longitudinal stress and plastic strain is given by: ) ( 4 σZ = K + G ε 3

(4.23)

where: G denotes shear modulus.

4.3.3 Study of Residual Stress Field and Surface Plastic Deformation in the Laser Shock Area From Eq. (4.14), the square root of the laser power density is proportional to the laser shock wave peak pressure. In order to study the residual stress field in the shock zone of TC4 titanium alloy under different laser power densities, the shock wave pressure is taken to be 2.5 ~ 8 GPa, and the residual stress distribution in the radius direction of the surface layer of the circular spot shock zone and the depth direction of the center of the shock zone are analyzed, and the simulation results are shown in Fig. 4.9. As can be seen from Fig. 4.9a, when the shock wave amplitude is 2.5 ~ 4.5 GPa, the residual stress on the surface of the shock zone increases with the increase of the shock wave amplitude; when the shock wave amplitude exceeds 4.5 GPa, the residual stress on the surface of the shock zone decreases with the increase of the shock wave amplitude, and the stress cavity phenomenon appears in the center of the shock zone (i.e. the residual stress amplitude in the center of the shock zone is lower than that in the surrounding area). The size of the stress cavity formed on the surface of the target material.

4 Numerical Analysis of Mechanical Effects of Laser Shock Peening

Residual stress

Residual stress

102

Surface X/mm

Depth Z/mm

Fig. 4.9 Residual stress distribution on the surface and depth direction of the shock area under different shock wave amplitude pressure a surface residual stress distribution; b depth direction residual stress distribution

This result is a good verification that when the shock wave amplitude pressure exceeds 2 times the elastic limit of the material (2 × 2.8 GPa), no further plastic stresses are generated in the material and the material plastic deformation reaches saturation 2HEL/(3λ + 2μ). From Fig. 4.9b, it can be seen that with the increasing amplitude of the shock wave, the deeper the residual compressive stress layer in the shock area, the maximum residual compressive stress in the depth direction keeps moving down when the shock wave amplitude pressure exceeds 2 times the elastic limit of the material, while the residual stress in the surface layer keeps decreasing. The laser shock zone surface profile test uses an optical morphometer, which uses vertical scanning interferometry to accurately measure the height nuances of the material surface through the interference fringes generated by the difference in the optical range of two beams of white light, and generates a morphology of the test zone through software on the collected data. As the laser power density increases, the peak pressure of the induced shock wave is also higher, and it can be seen from Eq. (4.4) that the higher the shock wave pressure, the higher the plastic strain generated. As can be seen from Fig. 4.10, the depth of the plastic deformation crater in the shock zone increases with increasing laser power density (3.37 ~ 4.07 GPa) (0.5 ~ 3.8 μm), while the surface roughness of the test zone also increases. The three specimens TC-14, TC4-6, and TC4-12 were shocked 1 time, shocked 2 times, and shocked 3 times, and the specific shock parameters are shown in Table 4.3. The three-dimensional surface profiles and profile curves of the three specimens TC-14, TC4-6, and TC4-12 are shown in Fig. 4.11. From the three-dimensional surface contours in Fig. 4.11, it can be seen that the plastic deformation within the three laser shock zones in Fig. 4.11a, b, and c is not uniform, and the degree of plastic deformation in the middle range of the shock zone is not very different, while there is a transition area from the shock zone to the substrate, and the plastic deformation gradually becomes smaller. The reason for the

103

Depth of crater/μm

Surface roughness/μm

4.3 Numerical Analysis of Circular Laser Spot

Shockwave peak pressure/GPa

Shockwave peak pressure/GPa

Fig. 4.10 Experimental results of crater depth and roughness variation in shock area under different process parameters a depth of crater; b surface roughness

large local plastic deformation in the shock zone cannot be explained yet and may be due to the instability of the water-bound layer or the unstable arrangement of the Al absorption layer. With the increase of the number of laser shock, the degree of plastic deformation in the shock area also increased, and the degree of plastic deformation and the number of shock is basically linear, which indicates that the shock 3 times has not yet produced cold work hardening phenomenon on the surface of TC4 titanium alloy. The calculation of the surface roughness of the whole test range (including the laser shock area and the test area outside the shock area) by the test software shows that the more the number of shocks, the greater the surface roughness of the whole test range.

4.3.4 Verification of the Residual Stress Field in the Laser Shock Zone Laser shock peening uses a water Confinement layer to increase the peak pressure of the shock wave. For a laser with a wavelength of 1064 nm, the maximum peak pressure of the shock wave generated in the water Confinement layer mode is about 5 GPa because the excessive laser power density will generate a parasitic plasma in the water Confinement layer, cutting off the subsequent laser energy input. The laser power density used in this small section is about 7.69 GW/cm2 , and its shock wave pressure amplitude is about 3.71 GPa. Figure 4.12 shows the residual stress distribution in the impact zone after single-spot laser shock peening. From Fig. 4.12a, the distribution of residual stresses along the radius of the impact zone in the x-direction to 3 mm outside the impact zone can be seen. The simulated results are in better agreement with the X-ray diffraction results, and the trend of stress distribution is the same. The maximum residual stress for the simulated results is −258 MPa and the maximum residual stress for the experimental results is −227.3 MPa, with a difference of 12% between the two results. The average value of all X-ray diffraction test point results (−140.0 MPa) differs from the simulated average value (−164 MPa) by 14%.

TC4-2

2.3 mm

−202.9 MPa

TC4-1

1.1 mm

−140.4 MPa

−180.6 MPa

1.5 mm

TC4-4 −333.9 MPa

3.8 mm

TC4-13 −227.3 Mpa

2.2 mm

TC4-14 −103.2 MPa

0.5 mm

TC4-15 −197.7 MPa

2.3 mm

TC4-16

−194.1 MPa

2.0 mm

TC4-17

−268.3 MPa

3.1 mm

TC4-18

−295.5 MPa

3.5 mm

TC4-19

Table 4.3 Test results of maximum depth of crater in impact zone and residual stress on the surface of the central area for different laser impact parameters

104 4 Numerical Analysis of Mechanical Effects of Laser Shock Peening

Depth of crater/μm

4.3 Numerical Analysis of Circular Laser Spot

105

Impact area profile (One Impact)

Depth of crater/μm

Impact area surface/mm Impact area profile (Two Impact)

Depth of crater/μm

Impact area surface/mm Impact area profile (Three Impact)

Impact area surface/mm

Fig. 4.11 Laser shock area surface morphology a single laser shock; b two laser shocks; c three laser shocks

Fig. 4.12 Residual stress distribution in the impact zone after single-spot laser shock peening a σ x direction residual stresses; b σ y direction residual stresses

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4 Numerical Analysis of Mechanical Effects of Laser Shock Peening

From Fig. 4.12, it can be seen that the X- and Y-directions in the impact zone are compressive stress states; in the transition phase between the impact zone and the unimpacted zone area,σ x has a gentle distribution, while σ y has a large stress distribution gradient; outside the impact zone,σ x within the edge of the impact zone 1 ~ 2 mm from the edge still maintains a compressive stress state, while σ y is a tensile stress state, the reason for the difference in the stress state between the two directions is that when the impact wave is loaded, the material around the impact area springs back to the center, resulting in compressive stress in the X-direction and tensile stress in the Y-direction. One of the advantages of the Laser shock peening technique is the depth of the residual compressive stress layer, which helps to reduce the crack expansion rate or even stop the crack expansion, while greatly improving the resistance of the part to foreign object damage. Figure 4.13 shows the distribution of residual stress σ x along the surface and depth directions in the laser shock zone, Fig. 4.13a the residual stress surface direction distribution curve can be seen in the residual stress area radius of 3 to 4 mm, Fig. 4.13b the residual stress along the depth direction distribution curve can be seen in the surface residual stress is the largest, the residual stress level decays continuously along the depth direction until it decays to zero. The distance from the surface to where the stress is zero is defined as the depth of the residual compressive stress layer, and with the increase in the number of impacts, the surface residual compressive stress amplitude increases continuously and the depth of the residual compressive stress layer increases continuously. The laser power densities of TC4-20 specimen were 9.28GW/cm2 and 8.86GW/cm2 for two impacts and 7.88GW/cm2 , 8.35GW/cm2, and 8.09GW/cm2 for three impacts of TC4-12 specimen, but the depth of residual compressive stress layer in the impact area of both specimens was basically the same, which was about 1.1 mm. Therefore, the laser power density and the number of impacts together determine the depth of the residual compressive stress layer in the impact area.

Fig. 4.13 Distribution of residual stress in the laser impact area along the depth direction a residual compressive stress distribution (6 mm); b comparison of residual stress distribution within 1.2 mm

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107

4.3.5 Relationship Between Surface Profile and Residual Stresses in the Single-Spot Impact Zone The test results of the maximum depth of the impact zone crater and the residual stress in the central region for different laser impact parameters are listed in Table 4.3, and the curves are plotted in Fig. 4.14, which shows that the residual stress value increases with the increase of the maximum depth of the impact zone crater. If the effect of testing error in residual stress is excluded, there is a linear relationship between the impact zone crater depth and residual stress from the available test results. The data of this experiment are limited and the utilized laser power density is only 7 ~ 9GW/cm2 , which can only roughly determine the relationship between the impact zone crater depth and the residual compressive stress, but it is a very important guideline for Laser shock peening TC4 titanium alloy. Figure 4.15 shows the simulation results of plastic deformation crater depth and maximum residual stress in the impact zone. Figures 4.14 and 4.15 shown results can be seen, due to the measurement of plastic deformation crater selected by the benchmark differences, resulting in large differences between the experimental results and simulation results, but the experimental results and simulation results of the laser peak pressure of 3 ~ 4 GPa, plastic deformation crater depth and residual stress level is basically positive correlation. In order to protect the laser and the limitation of water Confinement layer, this experiment failed to use the shock wave of more than 4.5 GPa to load the TC4 titanium alloy material, but the simulation results give the relationship between crater depth and residual stress of 3 ~ 5 GPa. In the interval of 3 ~ 4.5 GPa in Fig. 4.15, the plastic deformation crater depth is positively correlated with the residual stress level, and the residual compressive stress tends to decrease when the peak shock wave pressure exceeds 4.5 GPa.

Fig. 4.14 Experimental results of maximum depth of laser impact crater and maximum residual stress in the laser impact area

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Fig. 4.15 Simulation results of maximum depth of laser impact crater and maximum residual stress

4.3.6 Residual Stress Field of Lap-Spot In order to obtain a large surface area for impact treatment, the individual pulsed laser impact areas must be lapped together in an orderly manner to form a large residual compressive stress field on the surface of the workpiece. The latter laser shock will inevitably have an effect on the residual stresses in the previous laser shock zone, resulting in a redistribution of the residual compressive stress field. If the area of Laser shock peening is large, there is not only the influence of residual stress field in a single impact area but also the problem of stress redistribution between the next laser impact area and the previous laser impact area. So it is necessary to carry out finite element simulation analysis on the lap rate and lap form. In this paper, the loaded shock wave amplitude is 4GPa, the spot diameter is 6 mm, and the lap rates between spots are taken as 33%, 50%, and 67%, respectively. The overlap rate is defined as / η = (Φ − d ) Φ

(4.24)

where Φ—the laser spot diameter. d—two laser spot circular distances. Figure 4.16 shows the finite element simulation results of the residual compressive stress field on the surface in the X-direction generated by different spot lap rates. Line 1 shows the effect of the first spot action, line 2 shows the superimposed effect of two spots after the second spot action, and line 3 shows the superimposed effect of three spots after the third spot action. In order to approximate the result of residual compressive stress field generated after multiple spots are lapped into a row, three spots are chosen to lap at 33 and 67% lap rates, and two spots are chosen to lap at

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109

Fig. 4.16 Simulation results of residual stress distribution on the surface of large impact area with different lap rates a 33% lap rate; b 50% lap rate; c 67% lap rate. 1-First spot; 2-s spot; 3-Third spot

50% lap rate due to grid division. From the simulation results, it can be seen that the latter spot has a great influence on the residual stress in the action area of the former spot, and the spot of the 6 mm diameter laser has a significant influence on the residual stress in the 8 mm diameter area which is concentric with it, and the residual compressive stress effect after the impact of the spot lap is significantly increased, especially in the area where the two spots overlap before and after. From the results of the evolution of residual stresses in the three lap rates with the spot lap process, the greater the lap rate, the greater the effect of the latter spot on the residual stress field of the former spot, and vice versa, the smaller. In Fig. 4.16a, the σ x of the three light lap effect zones (4.10 mm in the horizontal coordinate) are uniformly distributed, and the residual compressive stress amplitude is in the range of 0.40–0.45 GPa; in Fig. 4.16b, based on the residual compressive stress results of the two spot laps, it can be inferred that the residual stress amplitude in the lap zone of the two spots with 50% lap rate is around 0.40–0.44 GPa; in Fig. 4.16c, the σ x in the three spot lap effect zones (3–7 mm in the horizontal coordinate) are also uniformly distributed with residual compressive stress amplitudes in the range of 0.45–0.50 GPa. All three lap rates can have a uniform distribution of σ x in the lap effect region, and this uniform residual stress distribution is beneficial to the effect of Laser shock peening, except that the residual compressive stress amplitude in the lap effect region increases as the lap rate increases. Figure 4.16 shows the effect of lap rate in one direction on the distribution of σ x on the surface of the impact zone. If one wants to obtain a larger area of residual compressive stress field, it is necessary to complete the spot lap in the point moving

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into line and line moving into surface for a large impact zone. For a circular spot with a diameter of 6 mm, it is reasonable to select 33% lap rate for a single row of laser impact area. The second row of impact area will also affect the residual stress distribution in the first row of impact area, so it is necessary to simulate and analyze the lap rate of the second row and the first row. Figure 4.17 shows the finite element simulation results for the two rows of lapped laser impact zones. The lap rate of 33% is selected within each row, and three lap rates of 33%, 50%, and 67% are selected between rows, corresponding to Fig. 4.17a–c three simulation results. From Fig. 4.17, it can be seen that the second row of laser impact zone also has a significant effect on the residual stress σ x in the impact zone of the first row. 33% and 50% lap rates between rows reduce the residual stress σ x in the impact zone of the first row, while 67% lap rate between rows increases the residual stress σ x in the impact zone of the first row. With the increase of row-to-row lap rate, the effect of the second row impact zone on the residual stress σ x in the first row impact zone results in better and better results. From the analysis of both process optimization and strengthening efficiency, the larger the lap ratio between rows, the lower the strengthening efficiency, but the higher the residual stress amplitude.

Fig. 4.17 Finite element simulation results of different lap rates between rows and rows a 33% lap rate between rows; b 50% lap rate between rows; c 67% lap rate between rows. 1-first spot; 2-s spot; 3-third spot; 4-fourth spot

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111

4.4 Square Spot Values Analysis Laser shock peening induces a high amplitude residual compressive stress layer on the surface layer of the target material to inhibit fatigue crack sprouting and extension. Therefore, it is very important to study the distribution of residual compressive stresses in the surface layer of Laser shock peening targets. At present, a large number of literature on numerical simulation studies on the residual stress distribution in the surface layer of round spot laser shock peening targets, but there are few reports on the numerical simulation studies on the residual stress distribution in square spot laser shock peening targets. In this section, Abaqus software was used to numerically simulate the surface residual stress distribution of square spot laser shock peening TC4 titanium alloy under different process parameters [8, 9].

4.4.1 Finite Element Model There are two types of objects to be simulated in this section: single-spot impact and multi-spot lapped single-sided impact. The single-spot impact area is located in the center of the target material, in view of the impact load (square spot 4 mm × 4 mm) and the target material shape with axisymmetry, a quarter of the three-dimensional model is selected, as shown in Fig. 4.18a, and the three-dimensional reduced unit (finite unit C3D8R and infinite unit CIN3D8) is selected for the simulation calculation. C3D8R is an elasto-plastic, nonlinear cell used to simulate the residual stress field distribution in the laser impact region and its stress wave influence zone; CIN3D8 is an infinite cell, this cell has only elastic material behavior, and the stress wave propagates to infinity in this cell as a reflection-free boundary. Multi-spot lap singlesided impact model does not have axisymmetry, take one-half of the whole model for simulation, as shown in Fig. 4.18b, and still choose three-dimensional finite reduction unit C3D8R and infinite unit CIN3D8 for simulation calculation.

4.4.2 Shock Wave Loading Assuming that the shock wave pressure is uniformly distributed on the material surface, the peak shock wave pressure under water Confinement layer condition can be estimated by Fabbro shock wave pressure model, i.e. Eq. (4.14). The width of the shock waveform generated under the water Confinement layer condition is about twice the laser pulse width, Fig. 4.19 shows the simplified shock wave pressure waveform generated by a pulsed laser with pulse widths of 30 ns and 10 ns, which has a pulse width of about 60 ns.

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(a)

(b)

Fig. 4.18 Square-spot laser shock peening finite element model. a single spot; b spot lap

Fig. 4.19 Simplified shock wave pressure waveform

The Johnson–Cook material model is used in this example, and the strain hardening and strain rate of the material during Laser shock peening are considered. The target material is TC4 titanium alloy, and the material mechanical property parameters are shown in Table 4.4.

4.4 Square Spot Values Analysis Table 4.4 Mechanical properties of the material

113

Properties

TC4 titanium alloy

Modulus of elasticity E/(GPa)

110

Density ρ/(Kgm− 3 )

4500

Poisson’s ratio/(ν)

0.342

Static yield strength A/(MPa)

1098

Strain hardening coefficient B/(MPa)

1092

Strain hardening parameter n

0.93

Reference strain rate M

1.1

C

0.014

4.4.3 Residual Stress Distribution Under Different Process Parameters (1) Single-spot residual stress field distribution Figure 4.20 shows for the spot center Laser shock peening TC4 titanium alloy depth direction elasto-plastic shock wave propagation and decay, the peak shock wave pressure is 4GPa. When the peak shock wave pressure exceeds the material elastic limit HEL, the target spot reinforced region produces axial and radial plastic deformation, however, the radial plastic deformation is limited by the material around the spot reinforced region. The axial and radial stresses during plastic deformation in the spot strengthening region follow the von. Mises criterion, with. σx = σy = σz − Ys

Fig. 4.20 Depth direction shock wave propagation and attenuation

(4.25)

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where the x–y plane is the metal surface and the z-direction is perpendicular to the x–y plane. When the peak pressure of the shock wave is lower than the elastic limit of TC4 titanium alloy HEL(HEL = 2.8GPa), the material plastic deformation ends only elastic deformation, at this time the material plastic deformation layer is called Laser shock peening material impact layer depth. Using ABAQUS finite element simulation calculated the pulse width of 30 ns, different spot form Laser shock peening produces surface residual compressive stress. The distribution of residual compressive stress S33 along the path from the center of the impact zone to the unreinforced region is shown in Fig. 4.21. The stress direction of S33 is parallel to the direction of the residual stress distribution path. When the shock wave pressure is 4 GPa, the maximum residual compressive stress values generated by the circular spot and the square spot are very close, but the degree of stress voids in the center of the square spot is much lower than that of the circular spot. When the laser shock wave pressure increased to 5 GPa, the residual tensile stress was generated in the center region of the square spot due to the stress cavity. The three-dimensional distribution of the residual stress S33 is shown in Fig. 4.22. It can be found that the area where the stress cavity occurs in the center of the square spot at a shock wave pressure of 5 GPa is larger than that at a shock wave pressure of 5 GPa. Another feature can be found in the 3D distribution of residual stresses shown in Fig. 4.22, which is that the volume of material affected by the stress voids is only distributed in the surface layer of the model. As shown in Fig. 4.22, the closer to the center of the spot, the lower the value of the residual compressive stress. Therefore, the most sensitive region of the stress cavity is the positive center of the spot. The surface finite element cell located at the positive center of the spot is taken out and used to study the dynamic stress and strain Fig. 4.21 Residual stresses from different spots and different shock wave pressures

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Fig. 4.22 Three-dimensional distribution of residual stresses generated at different spots and different shock wave pressures

variations. Figure 4.23 shows the deformation process of the plastic strain PE33 during the period from 0 to 1200 ns. Due to the effect of laser shock wave loading (generated by a 30 ns laser pulse), the plastic strain PE33 starts to be generated and rises rapidly to the maximum value at the initial stage of the curve of the plastic strain PE33 variation process. After the laser shock wave loading, the plastic strain PE33 stops increasing and remains at a certain level. The positive plastic strain PE33 indicates the outward expansion of the affected material. The material expansion is inhibited by the surrounding material, resulting in residual compressive stresses, which is the basic principle behind the generation of residual compressive stresses by Laser shock peening. Near 800 ns, the plastic strain PE33 starts to drop rapidly, which means that the previously expanded material starts to undergo reverse plastic deformation. The residual compressive stress decreases because the constraint of the surrounding material is reduced. The magnitude of the decrease in plastic strain PE33 near 800 ns depends on the applied peak shock wave pressure. When the shock wave pressure is 5 GPa, the plastic strain PE33 value is even negative. When the shock wave is loaded, a Rayleigh wave, known as Rayleigh wave, is generated at the laser spot edge, which propagates from the laser spot edge to the interior [10]. When the shock wave pressure is 4 GPa, the dynamic stress S33 distributed from the center of the spot to the unreinforced region is shown in Fig. 4.24. From Fig. 4.24, it can be seen that the position of the dynamic stress S33 moves toward the center of the spot with time, and the maximum dynamic stress S33 increases with time until its amplitude is equal to the dynamic yield strength (1345 MPa) of the TC4 titanium alloy material. Reverse plastic deformation occurs from 760 to 850 ns. It is clear that as the applied shock wave pressure increases, the time to produce reverse plastic deformation is prolonged. The shock wave pressure threshold for generating stress cavities in this model is about 3.6 to 3.7 GPa. For comparison, the plastic strain generated by the circular spot is also given in Fig. 4.23 Compared with the square spot, the reverse plastic deformation occurs first in the center of the circular spot

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Fig. 4.23 The process of plastic strain PE33

Fig. 4.24 Variation process of dynamic stress S33 (4GPa)

at the same shock wave pressure, and the decrease in plastic strain PE33 caused by the reverse plastic deformation is greater, which are due to the converging effect of Rayleigh waves at the center of the circular spot at the same time. Laser shock peening is applied with laser pulse widths ranging from a few nanoseconds to tens of nanoseconds. When the laser power density is the same but the laser pulse width is different, a larger pulse width means a higher laser pulse energy, so that the wider pulse laser produces a deeper residual compressive stress layer but

4.4 Square Spot Values Analysis

117

its surface residual compressive stress value is relatively low. The final equilibrium surface residual stresses produced by different laser pulse widths (10 ns and 30 ns) at 5 GPa shock wave pressure are shown in Fig. 4.25. Compared with the surface residual stress distribution curves generated by the pulse width 30 ns laser pulse, the pulse width 10 ns laser pulse not only does not produce stress voids in the center of the spot but also the maximum residual compressive stress rises to -700 MPa. It can be predicted that there is no reverse plastic deformation at the center of the spot for the 10 ns laser pulse. According to the material parameters given in the finite element model, it is known that when the Mises stress exceeds the dynamic yield strength of the material 1345 MPa, plastic deformation will occur. The dynamic mises stress curve along the surface at the moment when thedynamic mises stress generated by the 5 GPa shock wave reaches its maximum is shown in Fig. 4.26. In the 30 ns laser pulse case, the moment when the reverse plastic strain starts to be generated is around 710 ns. However, for the 10 ns laser pulse case, the maximum mises stress appears love you at the 769 ns moment, when the mises stress value is less than 1345 MPa, so no reverse plastic deformation is generated. The depth of the residual compressive stress layer is an important feature of Laser shock peening technique. In some stress areas in order to obtain a deeper residual compressive stress layer, relatively wide laser pulses are utilized for Laser shock peening applications, such as 30 ns laser pulses. However, laser pulses with wider pulse widths tend to produce stress voids, and the lack of residual compressive stresses will not be conducive to hindering fatigue crack sprouting and fatigue crack expansion. The finite element simulation results of the surface residual compressive stresses for a 30 ns laser pulse irradiated on the material surface followed by a 10 ns laser pulse irradiated at the same location are shown in Fig. 4.25. Compared with the 30 ns laser pulse, the surface residual compressive stresses generated by the Fig. 4.25 Comparison of residual stresses generated by 10 ns and 30 ns laser pulses

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Fig. 4.26 Dynamic equivalent stresses distributed along the surface

combined action of the 30 and 10 ns laser pulses reach a value of −800 MPa, and the stress voids in the center of the spot are effectively weakened. (2) Effect of spot shape on residual stress field The spot center residual stress hole phenomenon is generated by the boundary effect induced reverse plastic strain. Thus spot shape has a significant effect on the residual stress field on the surface of Laser shock peening targets. Figure 4.27 shows that circular spot, elliptical spot, and square spot laser shock peening induces residual stresses on the surface of TC4 titanium alloy. When the peak shock wave pressure is 4 GPa, the residual stresses in the center of circular spot, elliptical spot, and square spot laser shock peening TC4 titanium alloy spot are about −100 MPa, −300 MPa, and −300 MPa, respectively, so there is a residual stress hole phenomenon on the surface of circular spot induced target, and square spot and elliptical spot This is consistent with the experiments of Peyre et al. on cast aluminum alloys [11, 12]. The spot center residual stress hole phenomenon is attributed to the radial surface Rayleigh waves and their spot center focusing effect formed by the interaction at the impact boundary. The surface Rayleigh wave spot center focus effect leads to inverse plastic strain and stress holes in the target material [13]. Figure 4.28 shows that when circular spot laser shock peening target, all lateral surface Rayleigh waves generated at the circular spot boundary are focused at the center of the spot, thus more likely to induce residual stress hole phenomenon, but when elliptical spot and square spot laser shock peening target, lateral surface Rayleigh waves generated at the spot boundary have no Focusing point, the target surface has no residual stress hole phenomenon or residual stress hole phenomenon is not obvious, so the use of elliptical spot and square spot will weaken or eliminate the surface Rayleigh wave transient focusing effect and reduce the residual stress hole value in the center of the spot.

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119

Fig. 4.27 Different spot shape Laser shock peening TC4 titanium alloy surface residual stress distribution

Fig. 4.28 Surface Rayleigh wave propagation characteristics under different spot shapes a round spot; b square spot; c elliptical spot

(3) Residual stress hole mechanism Laser shock peening target induced surface residual stress hole formation mechanism is shown in Fig. 4.29, and the specific steps are as follows [14]. (1) The initial state of the target S0, spot center mass, laser shot peening induces a target state of S 1 , transverse displacement Ux > 0, longitudinal displacement Uz < 0, transverse strain εzp1 < 0. (2) The shock wave load stops, the spot center mass continues to extend laterally to reach the maximum value U xmax , and the target material state is S 2 . At this point, the surface Rayleigh wave caused by the spot boundary effect reaches the plasmas. (3) The surface Rayleigh waves gather at the center of the spot, the target state is S 3 , at this time the target transverse displacement Ux < 0, the longitudinal displacement Uz > 0, the transverse strain ε xp3 < εxp1 , the longitudinal strain εzp3 < εzp1 .

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4 Numerical Analysis of Mechanical Effects of Laser Shock Peening

Fig. 4.29 Mechanism of residual stress hole formation a initial state S0; b state S1; c state S2; d state S3

(4) After state S 3 , the dynamic response of the target material point stabilizes and no more plastic deformation occurs. When the target material state changes from S 2 to state S 3 , the inverse plastic deformation caused by the surface Rayleigh wave weakens the lateral constraint of the material around the mass point and reduces the residual compressive stress field in the center of the spot. Fan et al. [15] proposed the plastic strain and stress [ ] 1n , where εp is the equivalent plastic strain. Because of relationship as εp = (σ −A) B this n > 0, higher intensity surface Rayleigh waves will induce more plastic strain as the shock wave pressure increases. Therefore, the inverse plastic strain is related to the surface Rayleigh wave intensity. Under different conditions, the residual stress hole phenomenon appears in three behaviors as shown in Fig. 4.30 [14].

Fig. 4.30 Three behaviors of residual stress holes

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121

(1) The reverse plastic strain is small (Rayleigh wave is weak), and the transverse and longitudinal strains are εxp3 > 0, εzp3 < 0, and the residual stress in the center of the spot is still compressive σxres < 0 when the laser shot peening metal is in steady state. (2) The reverse plastic strain is large (Rayleigh wave is stronger), and the transverse and longitudinal strains areε xp3 = 0, εzp3 = 0 and the residual stress in the center of the spot is σ x res = 0 for the steady state of the laser shot peened metal, respectively. (3) The reverse plastic strain is very large (Rayleigh wave is very strong), and the transverse and longitudinal strains areεxp3 < 0,εzp3 > 0, and the residual stress at the center of the spot is still the compressive stress σ x res > 0 when the laser shot peening metal is in steady state. (4) Effect of power density on residual stress field Figure 4.31 shows the residual stress distribution of square spot laser shock peening TC4 titanium alloy at different laser power densities. From Fig. 4.31 it can be seen that as the laser power density increases, the surface residual stress and residual compressive stress layer depth increases simultaneously, and the residual stress hole phenomenon is more serious, and it is tentatively concluded that the residual stress hole phenomenon appears on the surface of TC4 titanium alloy when the peak shock wave pressure reaches 4 GPa. The experimental data in Fig. 4.31 were obtained by measuring the residual stress of Laser shock peening TC4 titanium alloy using XRD method with the experimental parameters of square spot with 4 mm side length and power density of 6.23 GW/cm2 . From Fig. 4.31, it can be seen that the experimental results of the residual stress field on the surface of TC4 titanium alloy are consistent with the numerical analysis results, but the numerical analysis results of the residual stress field in the depth direction are lower than the experimental results. When the laser power density is greater than 9GW/cm2 , the maximum residual compressive stress of Laser shock peening TC4 titanium alloy is located in the surface layer.

Fig. 4.31 Residual stress distribution of square-spot laser shock peeningTC4 titanium alloy at different laser power densities a surface residual stresses; b depth direction residual stresses

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Fig. 4.32 Residual stress distribution of TC4 titanium alloy under different number of impacts a Residual stress distribution in depth direction; b, c depth of residual compressive stress layer in different regions of spot

(5) The effect of the number of impacts on the residual stress field Laser shock peening of different thicknesses of the target material, the laser power density and the number of shocks must be well matched. Figure 4.32 shows the distribution of the residual stress field of square-spot laser shock peening TC4 titanium alloy at different laser power densities and number of shocks, with the increase in the number of shocks, the depth of the residual compressive stress layer increases, the reason may be that the depth of the attenuation layer induced by the purely elastic propagation of the shock wave in the pre-stressed layer is smaller. The maximum residual compressive stress on the surface increases with the increase in the number of impacts. In addition, two impacts with power density of 6 GW/cm2 induced similar surface residual compressive stress and residual compressive stress affected layer depth of TC4 titanium alloy with single impact with power density of 7 GW/cm2 . The low power density multiple impacts are suitable for enhancing the mechanical properties of fanblade with thickness less than 1 mm, which can effectively avoid the lamellar cracking phenomenon on the back side of the blade. The results of two Laser shock peening TC4 titanium alloy residual stress experiments with a power density of 6GW/cm2 are shown in Fig. 4.32a. Figure 4.32b shows that the depth of the residual compressive stress layer increases with the increase of the number of shocks and decreases with the increase of the distance from the center of the spot. The difference between the residual stress field of TC4 titanium alloy induced by two impacts and three impacts is small for a power density of 6 GW/cm2 , but the difference between the residual stress field of TC4 titanium alloy induced by two impacts and three impacts is large for a power density of 7 GW/cm2 . (6) The effect of spot lap rate on the residual stress field Multiple spot lap can Laser shock peening target a large area, and adjacent spot shock peening target can redistribute the last spot peening residual stress field. Appropriate spot lap rate Laser shock peening targets to obtain uniform residual compressive stress field. To avoid residual tensile stress fields in unreinforced regions, the minimum lap rate for circular spot laser shock peening is 20%, while the lap rate for square spot

References

123

Fig. 4.33 Spot lap Laser shock peening TC4 titanium alloy residual stress distribution a surface residual stress distribution; b depth direction residual stress distribution

laser shock peening is very small, below 5%. Figure 4.33 showsspot unlapped two square spot power density of 7GW/cm2 Laser shock peening TC4 titanium alloy residual stress distribution, the surface residual stress field near the lap region is relatively stable with residual stress values of −400 to −350 MPa. Compared with single-spot laser shock peening, lap-spot laser shock peening improves the depth of residual compressive stress layer at the edge of the spot.

References 1. Fabbro R, Fournier J, Ballard P et al (1990) Physical study of laser-produced plasma in confined geometry. J Appl Phys 68(2):775–784 2. Che ZG, Gong SL, Cao ZW et al (2011) Theory analysis and experiment investigation of laser shock processing on titanium alloy blade (in Chinese). Rare Metal Mater Eng S4:235–239 3. Zhou N (2002) Materials dynamics under pulse beam radiation (in Chinese). National Defense Industry Press 4. Peyre P, Chaieb I, Braham C (2007) FEM calculation of residual stresses induced by laser shock processing in stainless steels. Modell Simul Mater Sci Eng 15(3):205–221 5. Johnson GR, Cook WH (1983) A constitutive model and data for metals subjected to large strains, high strain rates and high temperatures. Hague, Netherlands 6. Fan YF, Duan ZP (2003) Cylinder explosive test and material model of Johnson-Cook (in Chinese). Mech Eng 25(5):40–43 7. Zhuang Z, You XC et al (2009) Finite element analysis and application based on ABAQUS (in Chinese). Tsinghua University Press 8. Cao ZW, Che ZG, Zou SK et al (2011) Numerical simulation of residual stress field induced by laser shock processing with square spot (in Chinese). J Shanghai Univ 06(6):553–556 9. Cao ZW, Che ZG, Zou SK (2013) Simulation study of stress hole on laser shock peening with square spot. Rare Metal Mater Eng 42(S2):222–225 10. Fabbro R, Peyre P, Berthe L et al (1998) Physics and applications of laser-shock processing. J Laser Appl 10(6):265–279 11. Peyre P, Merrien P, Lieurade H (1993) Optimization of the residual stresses induced by laser shock treatment and fatigue life improvement of 2 cast alum alloys. Oxford, UK 12. Peyre P, Fabbro R, Merrien P et al (1996) Laser shock processing of aluminium alloys. Application to high cycle fatigue behaviour. Mater Sci Eng A 210:102–113

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13. Masse JE, Barreau G (1995) Laser generation of stress waves in metal. Surf Coat Technol 70(95):231–234 14. Hu Y, Gong C, Yao Z et al (2009) Investigation on the non-homogeneity of residual stress field induced by laser shock peening. Surf Coat Technol 203(23):3503–3508 15. Fan Y, Wang Y, Vukelic S et al (2005) Wave-solid interactions in laser-shock-induced deformation processes. J Appl Phys 98(10):104904

Chapter 5

Evaluations of the Strengthening Effect of the Metals with Laser Shock Peening

Laser shock peening induced high-pressure shock waves into the interior of the target material, causing severe plastic deformation on the surface of the target, and introducing a high-amplitude residual compressive stress layer of about 1 mm [1]. After the laser shock wave is loaded, the surface Rayleigh wave is generated at the edge of the target light spot, and the Rayleigh wave propagates and focuses on the center of the light spot at the same time. A large tensile pulse is generated at the center of the light spot, thus reducing the residual compressive stress field near the center of the light spot and forming the stress hole phenomenon [2, 3]. Due to the absolute geometric symmetry of the circular spot, the residual stress hole induced by the laser shock-peened target has a negative effect [4]. In order to meet the special needs of industrial applications, square spot laser shock peening has been applied more and more. Compared with the circular spot, the square spot does not produce residual stress hole phenomenon, and the overlap ratio is less than 5%, which can effectively strengthen the target with large area laser shock. The highamplitude residual compressive stress of the target induced by laser shock peening plays a key role in improving the fatigue life of the target [5, 6].

5.1 Surface Profiles Induced by Square Spots Laser shock peening technology is more and more widely used in the United States, and there is already a SAE/AMS technical standard (AMS2456) for laser shock peening, which has been certified by ISO 9001 and FAA. The most important purpose of laser shock peening technology is to improve the fatigue performance of materials and reduce the notch sensitivity of materials. Therefore, factors that reduce the fatigue performance of materials should be avoided as far as possible to maximize the advantages of laser shock peening technology in improving the fatigue performance of materials [7, 8]. Because the increase of material surface roughness and waviness after LSP will offset part of the fatigue gain, it is very important to optimize the © National Defense Industry Press 2023 S. Zou et al., Laser Shock Peening, https://doi.org/10.1007/978-981-99-1117-2_5

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5 Evaluations of the Strengthening Effect of the Metals with Laser Shock …

process parameters to obtain a smooth LSP metal surface profile. Circular spot and square spot are the most commonly used laser shock peening spot shapes. Single or multiple shocks may be carried out on the material surface for different materials and applications. The purpose of multiple shocks is to obtain a deeper residual compressive stress layer and a more uniform overlap ratio [9].

5.1.1 Spot Overlapping Patterns When circular spot laser shock peening is adopted, to ensure full coverage of the area to be strengthened, there must be a certain degree of overlap ratio. There are two main arrangements for circular spot overlapping, one is square arrangement, and the other is triangular arrangement, as shown in Fig. 5.1. For example, define area coverage = sum of all spot areas/area of full coverage area. To ensure full coverage of an area, the minimum coverage of a circular facula in a square arrangement is 157%, while the minimum coverage of a triangular arrangement is 121%. In practical use, due to the orderly arrangement of squares, the spot path is easy to calculate and is a common arrangement form. In the process of laser shock peening, it is troublesome to calculate the area coverage. On the basis of ensuring the full coverage of laser shock peening, the line overlap ratio is generally adopted. The line overlap ratio = (spot diameter or side length − center distance between adjacent spots)/spot diameter or side length. The circular light spot overlap ensures that the line overlap ratio under full coverage is at least 29.3%, as shown in Fig. 5.1. Therefore, the minimum line overlap ratio is generally 30%. When the overlap ratio is 50%, the edge of the next light spot is just pressed against the center of the previous light spot, and when the overlap ratio is 67%, the edge of the third light spot is just pressed against the center of the first light spot. Figure 5.2 shows the laser shock peening arrangement of circular light spots under different line overlap ratios.

Fig. 5.1 Lap mode of circular spot: a square arrangement; b triangular arrangement

5.1 Surface Profiles Induced by Square Spots

127

Fig. 5.2 Arrangement of circular spots with different overlap rates

5.1.2 Surface Profiles Induced by Square Spots Figure 5.3 shows the relationship between shock wave pressure and plastic strain [10], where the calculation formula of material elastic limit HEL is

HEL =

1 − ϑ dyn σ 1 − 2ϑ s

(5.1)

where dyn

σs ϑ

dynamic yield strength of material under the strain rate of 106 S−1 ; Poisson’s ratio of the material.

It can be seen from Fig. 5.3 that when the shock wave pressure exceeds the elastic limit HEL of the material, plastic deformation occurs on the surface of the material, forming pits and residual compressive stress, which increase with the increase of the shock wave pressure. The optimal shock wave pressure is twice the elastic limit of the material. Not only the utilization rate of circular spot is relatively low, but also the surface profile obtained is uneven. In order to improve the utilization rate of laser shock peening spot and obtain a more flat surface, China Academy of Aeronautical Manufacturing Technology uses beam shaping technology to realize the transformation of Fig. 5.3 Relationship between shock wave pressure and plastic strain

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5 Evaluations of the Strengthening Effect of the Metals with Laser Shock …

a circular spot to a square spot through pentagram and quadrant lens, as shown in Figs. 5.4 and 5.5 [11, 12]. Through experimental verification, the lens can withstand the peak power density of 4 GW/cm2 , and obtain a square spot with nearly uniform energy distribution. The surface profile of the stainless steel specimen induced by square spot laser shock peening is a smooth pit with a depth of 3–5 μm, as shown in Fig. 5.5. Compared with the surface profile of square spot laser shock peening, although there is a smooth transition profile at the edge of a single circular spot, any overlapping circular spot will produce an unsmooth surface. A single square spot can produce a pit with a smooth bottom and steep steps. At the same time, square spot overlapping can obtain a very smooth lapping effect, and the lapping rate is only 10%. Figure 5.6 shows the simulation of the stress field generated by the impact of circular and square light spots. Figure 5.7 shows the surface profile obtained by the circular spot lapping

(a)

(b)

Fig. 5.4 Pentagram (a) and quadrant lens (b)

Fig. 5.5 Square spot effect: a light spot ablative crevices; b surface profile

5.1 Surface Profiles Induced by Square Spots

129

Fig. 5.6 The simulation of the stress field generated by the impact of circular and square light spots

Fig. 5.7 The surface profile obtained by the circular spot lapping impact treatment: a two spots; b six spots

impact treatment. It is unnecessary to replace the absorption layer for square spot continuous multi-row lapping peening. When the square light spot is used for lapping, due to certain errors in the shape and position accuracy of the square light spot and the energy distribution in the edge area, a certain lapping rate is required to ensure the lapping of the square light spot, but it is difficult to ensure that the lapping area is completely consistent in actual work. At this time, it is necessary to consider the possible errors under the ideal 100% coverage, that is, when the coverage rate is greater than 100% or less than 100%, there is overlapping or underlapping. Figure 5.8 is the schematic diagram of square spot lapping with coverage greater than 100% and less than 100%, and Fig. 5.9 is the surface morphology of laser shock-peened target with square spot lapping with coverage greater than 100% and less than 100%. When the coverage rate is less than 100%, the most prominent problem on the material surface is that the material extrusion of the "blind area" increases the waviness of the surface of the impact area, and the extrusion area is in the tensile stress state, which is called the weak link in the impact area. Therefore, it is necessary to adjust the process parameters according

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Fig. 5.8 Schematic diagram of square spot overlap: a coverage greater than 100%; b coverage less than 100%

Fig. 5.9 Surface morphology of laser shock strengthened target with square spot; four square light spots with coverage a greater than 100%; b less than 100%; c two square light spots with coverage less than 100%

to the actual strengthening pit position, so that there is no area not irradiated by laser on the material surface, which can further reduce the extrusion and improve the stress state of the extrusion area, reduce the surface roughness of the impact area, and improve the fatigue performance of the material strengthening area.

5.2 Mechanical Property of High-Temperature Alloy

131

Fig. 5.10 Ways of spot overlap: a one-way type; b round-trip type

5.1.3 Path Planning of Spot Overlapping When multi-row spot overlap is required for laser shock peening, there are different ways of spot overlap, such as one-way type and round-trip type, as shown in Fig. 5.10. To obtain a smooth transition area between the laser shock-peened area and the non-peened area, the power density of the last row of spots gradually decreases.

5.2 Mechanical Property of High-Temperature Alloy This section mainly introduces the stability of residual compressive stress induced by laser shock peening and the influence of the thermal cycle on the residual stress of superalloy GH2036.

5.2.1 Overseas Research Status The fatigue experiments at room temperature and high temperature of TC4 titanium alloy after laser shock peening were conducted by the University of California, Berkeley, and Kassel University in Germany. The parameters of laser shock peening are laser spot size 2.6 mm × 2.6 mm, laser pulse width 18 ns, laser power density 7 GW/cm2 , coverage rate 200%; the heating temperature is 450 °C (T /T m ≈ 0.4); the maximum stress of fatigue test is 460 MPa, the stress ratio is −1, and the frequency is 5 Hz. As shown in Figs. 5.11 and 5.12, whether it is an alternating fatigue load or a temperature load, the residual compressive stress in the strengthened surface of TC4 titanium alloy is released to varying degrees after laser shock peening. However, the strain hardening of the strengthened surface layer is very stable with fatigue or temperature load. The research results of the influence of laser shock peening on residual stress abroad show that the release rate of residual stress after laser shock peening is obviously lower than that of shot peening. Figure 5.13 shows the comparison results of IN100 after laser shock peening and shot peening. After holding at 650°C for

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Fig. 5.11 Fatigue properties of TC4 after laser shock peening

Fig. 5.12 Residual compressive stress release of TC4 titanium alloy strengthened surface layer

10 h, the residual stress of shot peening decreases to less than 1/2 of the level before holding, while after laser shock peening, the residual stress is basically stable, and only the top surface layer of 0.05 mm decreased by about 1/3.

5.2 Mechanical Property of High-Temperature Alloy

133

Fig. 5.13 Release of residual compressive stress on the strengthened surface of IN100 Alloy: a shot peening; b laser shock peening

5.2.2 Effect of Thermal Cycles on Residual Stresses of High-Temperature GH2036 Alloy The turbine disk material GH2036 alloy works at high-temperature conditions, so the release of residual stress at high-temperature thermal cycle after laser shock peening must be assessed. An automatic thermal cycle furnace was used to conduct high-temperature thermal cycle tests on the specimens with and without laser shock peening. The temperature uniformity of the automatic thermal cycle furnace is ±10 °C, and the heating speed is 50 °C/min. The thermal cycle furnace and thermal cycle temperature settings are shown in Fig. 5.14.

Fig. 5.14 Thermal cycle equipment and thermal cycle temperature setting

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5 Evaluations of the Strengthening Effect of the Metals with Laser Shock …

0h

1h

5h

20h

50h

100h

Fig. 5.15 GH2036 alloy samples after different thermal cycles

The high-temperature thermal cycle specimen is divided into two types: ➀ imitating Almen test piece, whose size is 76 mm × 19 mm × 2 mm, which is mainly used to determine the release of residual stress through the change of the arc height value of the specimen during the thermal cycle; ➁ the size of the specimen is 50 mm × 50 mm × 10 mm, and the laser shock zone of the sample surface is 15 mm × 15 mm; the influence of thermal cycle on the surface and depth residual stress of laser shock zone of GH2036 alloy was tested by XRD + electrolytic stripping method. Figure 5.15 shows the GH2036 sample after different thermal cycles. The residual stress test of GH2036 alloy specimens under different thermal cycle conditions was tested by the X-ray diffraction method. The test equipment conditions are shown in Fig. 5.14, and the test results are shown in Table 5.1. It can be seen from Table 5.1 that with the increase of thermal cycles, the residual stress decreases and is basically stable after 50 cycles, but still around −300 MPa, which is about 56% of the initial residual compressive stress. The main reason for residual compressive stress relaxation is the decrease of the elastic storage energy in the laser shock peening area; the increase of temperature will increase the mobility of dislocations and cause the large-scale migration of atoms and vacancies. As a result, the annihilation of dislocations is increased, and the resistance of dislocation movement is reduced. Residual stress relaxation is a thermal activation process, and it is not difficult to understand why there is a critical relaxation temperature. The critical temperature should be related to the energy state of the dislocation and is therefore affected by the magnitude of the residual stress itself. The greater the residual stress, the less its own stability, and the more susceptible it is to external temperature. Similarly, the relaxation rate should also be related to the residual stress magnitude. As shown in Fig. 5.16, the residual stress depth test results show that the residual stress depth can be maintained at about 0.3 mm after 100 thermal cycles.

5.2 Mechanical Property of High-Temperature Alloy

135

Table 5.1 Residual stress test results of GH2036 alloy surface after different thermal cycles Test specimen number 1

5

6

7

8

9

The number of thermal cycles 0

1

5

20

50

100

Test point

Test results/(MPa)

Average residual stress/(MPa)

1–1

−553 ± 88

−548

1–2

−584 ± 84

1–3

−510 ± 35

1–4

−543 ± 69

5–1

−519 ± 71

5–2

−572 ± 79

5–3

−610 ± 71

5–4

−560 ± 80

6–1

−479 ± 61

6–2

−454 ± 67

6–3

−471 ± 66

6–4

−492 ± 57

7–1

−416 ± 88

7–2

−419 ± 100

7–3

−216 ± 68

7–4

−475 ± 78

8–1

−379 ± 127

8–2

−378 ± 118

8–3

−315 ± 70

8–4

−240 ± 43

9–1

−311 ± 99

9–2

−304 ± 65

9–3

−385 ± 65

9–4

−275 ± 56

−565

−474

−381

−328

−318

5.2.3 Fatigue Lives of High-Temperature GH30 Alloy The size of the test piece taken from the static test is the same as that of the fatigue test piece, and only the tensile strength of the material measured without a hole in the middle is σb = 742MPa, the static strength of the material meets the requirements, the loading coefficient K = 0.45–0.6 is selected, and the nominal stress σn = 334−445 MPa is calculated according to the strength limit. After laser shock peening, all test pieces are processed ∅ 0.5–0.6 mm holes with an electric pulse in the center of the laser shock area. The test results are shown in Table 5.2. With the loading coefficient of 0.45, when the unpeened control specimens (1–3) have 90% confidence, the number of specimens meets the experimental requirements, and the median fatigue life is 706307. The LSP 38 specimen did not break after

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Fig. 5.16 Surface residual stress after different thermal cycles

3502,160 cycles, when the fatigue life increased by nearly 400%. In the case of a loading coefficient of 0.6, when the unpeened control specimens (5 and 6) are given 95% confidence, the number of specimens meets the experimental requirements, and the median fatigue life is 252517. The median fatigue life of LSP 34, LSP 35, LSP 37, and LSP 39 specimens is 318360, and the value of fatigue life improvement is only 26%. Table 5.2 Fatigue test results of superalloy Test specimen number

S/mm2

Pmax /KN

σn /MPa

K

N/times

lg N

Numerical treatment of specimen group [13]

Contrast 1

22.99

7.68

334

0.45

681,840

5.834

Contrast 2

23.52

7.85

334

0.45

482,720

5.684

Contrast 3

23.55

7.86

334

0.45

1,070,540

6.030

Arithmetic mean life of N 1 = 745,033 Median fatigue life, N 2 = 706,307 The number of specimens with 90% confidence meets the requirements

LSP38

23.95

8.00

334

0.45

3,502,160 +

10.66

445

0.6

107,980 (continued)

5.2 Mechanical Property of High-Temperature Alloy

137

Table 5.2 (continued) Test specimen number

S/mm2

Pmax /KN

σn /MPa

K

N/times

lg N

Numerical treatment of specimen group [13] Arithmetic mean life N*1 = 255,767 Median fatigue life N*2 = 252,517 The number of specimens with 95% confidence meets the requirements

Contrast 5

23.11

10.29

445

0.6

213,220

5.329

Contrast 6

23.19

10.32

445

0.6

241,940

5.384

Contrast 7

23.34

10.39

445

0.6

312,130

5.494

LSP34

23.94

10.66

445

0.6

351,390

5.546

LSP35

23.95

10.66

445

0.6

287,280

5.458

LSP37

23.97

10.66

445

0.6

339,390

5.531

LSP39

23.95

10.66

445

0.6

299,850

5.477

Arithmetic mean life N*1 = 319,480 Median fatigue life N*2 = 318,364 N*2/N 2 = 1.25 N*1/N 1 = 1.26

Note N value Note + indicates continuous loading without interruption; S is the sectional area of the small hole Pmax is the maximum load; σn is the nominal load; K is the loading coefficient; N is the number of cycles

5.2.4 Fatigue Crack Growth Rate of High-Temperature GH30 Alloy The fatigue life of structures can generally be divided into fatigue crack initiation life and fatigue crack propagation life. The fatigue crack initiation life is the life corresponding to the development of microscopic defects into macroscopic detectable cracks, while the fatigue crack propagation life is the life in the region where macroscopic detectable cracks propagate to critical cracks and cause damage. The strain hardening produced by laser shock peening can not only delay the development of microscopic defects into macroscopic detectable cracks but also greatly reduce the fatigue crack propagation rate and improve the fatigue life of the structure. This section studies the effect of laser shock peening on the crack growth rate of superalloy GH30 and other plates [14, 15]. 1. Test conditions The test material is 1.7 mm thick GH30. Compact tension (CT) test pieces are used. The specimen size and the location and trace of laser shock peening are shown in Figs. 5.17 and 5.18, respectively. The tensile direction is transverse (that is, the crack

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Fig. 5.17 Dimensions of CT specimens and positions of laser shock peening

Fig. 5.18 Traces of laser shock peening

propagation direction is the plate rolling direction). The technological process of specimen processing and testing is as follows: machining external dimensions; laser cutting round holes and grooves; wire cutting incision length of 8 mm; laser shock peening; precast crack; and fatigue crack propagation experiment. Laser shock peening adopts a round light spot with a diameter of 6 mm, which is impacted for three times continuously on one side along the extension line of the prefabricated crack, with a spacing of 5 mm, so that there is a 17% overlap rate between adjacent light spots. The laser pulse energy is 14 J, the pulse width is 20 ns, and the laser peak power density is 2.5 GW/cm2 . In this way, a laser shock peening area of about 16 mm long and 5 mm wide is formed on the crack propagation path. The fatigue crack growth experiment uses an 880MTS fatigue testing machine, loaded by the MTS TestStar Lis program, with loading accuracy of 0.5%, sinusoidal waveform, frequency f = 30 Hz, stress ratio R = 0.1, and the experimental environment being room temperature and air.

5.2 Mechanical Property of High-Temperature Alloy

139

The fatigue crack growth length was measured by visual method, and a reading microscope with a magnification of 30 times was used. After a predetermined number of cycles (2000–10,000 times), the cyclic load was stopped, and a certain constant load (about 0.75Pmax ) was loaded for static measurement to reduce the measurement error. During the test, the reading of the magnifying glass base and the increased cycle number ΔN were recorded and converted into the crack length a and the corresponding cycle number N. 2. Treatment method of the test results According to the measured crack length a and its corresponding number of cycles N, a series of discrete a–N points are polynomials fitted after tracing the line with Origin software. It is found that 6–8 term equations can well fit the a-N curve: Hypothesis: a = A0 + A1 N + A2 N2 + A3 N3 + . . . An Nn

(5.2)

Then: da/dN = A1 N + 2 A2 N + 3A3 N2 + . . . n An Nn−1

(5.3)

The ΔK of CT specimens can be expressed as [16] ( a )2 ⎤ (a) − 13.32 0.886 + 4.64 ΔP ⎢ W W ⎥ ΔK = ⎦ )2/3 ⎣ ( a )3 ( a )4 1 · ( BW 2 1 − Wa + 14.72 − 5.6 W W ) (a ≥ 0.2 W ( 2+

a W

)

⎡

(5.4)

where ΔP B W ΔK

Load range (N); Thickness of specimen (mm); Width of specimen (mm); Stress intensity factor (N·mm−3/2 ).

The width W of the specimen shall be large enough to ensure that the specimen is always in a small yield state (basically in the linear elasticity state) during the test process of obtaining valid da/d N data. For CT specimens, the following conditions shall be met, i.e. ) ( 4 K max (5.5) W − amax ≥ π σs The initial crack of the specimen is composed of a machining notch and fatigue crack. To ensure the validity of Eq. (5.4), the initial crack length should meet the following conditions: i.e. a(N0 ) ≥ 0.2W

(5.6)

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The width of the test piece used in this test W = 48 mm, the initial crack length a0 >10 mm, and other conditions meet the requirements that the test piece is always in a small yield state (basically in a linear elastic state) and the validity of Eq. (5.4). Taking a series of N values, such as N 1 , N 2 , N 3 … N i , the corresponding a1 , a2 , a3 … ai , (da/dN) 1, (da/dN) 2, (da/dN) 3….. (da/dN) i, ΔK 1 , ΔK 2 , ΔK 3 … ΔK i can be calculated from Eqs. (5.2) to (5.4), ΔK 1 , ΔK 2 , ΔK 3 … ΔK i , finally, the relationship curve between da/dN and ΔK is depicted. 3. Crack propagation rate of GH30 specimen The thickness of the GH30 specimen is 1.64 mm, and the loading load is P = 3300 N. The results of the performed experiments are shown in Fig. 5.19. According to Fig. 5.19, in the test range, the CT-12 specimen without LSP is in the transition range of initial crack stable growth stage I and intermediate crack stable growth stage II. It can be considered that after lgΔK = 3.0 is in stage II. At this time, the lg(da/dN) and lgΔK curves are fitted to Paris type, and the straight-line fitting results are expressed as lg(da/d N ) = −19.86463 + 5.35406 lg ΔK R = 0.03058, S D = 0.12504

Fig. 5.19 lg (da/dN)–lgΔK curve of GH30 specimen

(5.7)

5.3 Mechanical Property of Stainless Steel

141

In this test, the Paris formula of the intermediate crack stable growth stage II of GH30 plate without LSP is da/d N = 1.37 × 10−20 ΔK 5.4

(5.8)

Comparing the results of CT-10 and CT-12, LSP can significantly reduce the crack growth rate in the range of lgΔK = 2.98–3.12 (corresponding to the crack length α = 12.5–17 mm and ΔK = 955–1318 N/mm3/2 , respectively), and the maximum amplitude is about 30 times. The variation trend is similar to that without LSP. The whole range measured is in the LSP-ed zone.

5.3 Mechanical Property of Stainless Steel 5.3.1 Fatigue Lives of 1Cr18Ni9Ti Austenitic Stainless Steel 1. Experimental methods Austenitic stainless steel 1Cr18Ni9Ti, which is widely used in the aerospace field, was selected as the test material [17], and the solid solution cold-rolled stainless steel 1Cr18Ni9Ti sheet was a 1.2 mm thick hot-rolled annealed sheet; the laser impact specimens are shown in Fig. 5.20. Laser shock peening using dye-modulated Q multistage amplification of neodymium glass laser device. The laser wavelength is 1.06 μm, pulse width is 20–50 ns, single pulse energy is 10–50 J, the focused spot to the target is around 6 mm, and the power density is l–5 GW/cm2 . After laser shock peening, small holes were machined in the center of the laser shock peening area by electric pulse machining method to produce notched fatigue specimens with a central circular hole, and finally fatigue tests were carried out. The fatigue test was conducted on a low-frequency fatigue tester Instron 1253 with stress ratio R = 0.1, frequency f = 28.5 Hz, stress concentration factor K t = 2.85, and loading factor K = 0.4–0.6.

Fig. 5.20 LSP specimen size

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Table 5.3 Fatigue test results for stainless steel No.

S/mm2

Pmax /kN

σn /MPa

K

N/times

Compare 3

16.98

4.56

268

0.4

737,540

Compare 4

16.92

4.54

268

0.4

924,580+

5.00

295

0.44

422,420

5.67

335

0.5

1,422,320+

lg N

Numerical processing [13] Note “+” for changing the loading factor to continue the test if it is not broken

LSP28

16.93

6.24

369

0.55

129,940

Compare 6

17.38

6.42

369

0.55

341,190

Compare 7

17.29

6.38

369

0.55

179,020

Compare 5

17.63

6.98

396

0.59

40,800

40,611

Compare 8

17.60

6.97

396

0.59

54,630

4.737

Compare 9

17.60

6.94

394

0.59

63,160

4.800

LSP29

16.93

6.66

393

0.59

106,730

5.028

LSP30

17.04

6.79

398

0.59

221,800

5.346

LSP31

17.02

6.69

393

0.59

81,940

4.918

LSP32

16.97

6.72

396

0.59

90,250

4.955

LSP33

17.02

6.73

395

0.59

61,270

4.787

Arithmetic mean life N 1 * = 52,860 Median fatigue life N 2 * = 520,520 The number of specimens meets the requirement at 90% confidence level Arithmetic mean life N 1 * = 112,398 Median fatigue life N 2 * = 101,412 N 1 * /N 1 = 2.126 N 2 * /N 2 = 1.949

2. Fatigue test results of austenitic stainless steel 1Cr18Ni9Ti plate specimens The tensile strength of the material measured in the static tensile test σb = 670MPa, the static strength of the material meets the requirements. The loading coefficient K = 0.45–0.6 was selected, and the nominal stress σ n = 334–445 MPa was calculated according to the tensile strength. All specimens were first laser shock-peened and then electric pulses were used to punch ∅0.5mm holes in the center of the treated area, and then ∅0.45mm steel wires with abrasives were used to process the hole, and finally make the tensile-tensile fatigue test. The test results are shown in Table 5.3.

5.3 Mechanical Property of Stainless Steel

143

The fatigue life test results show that the fatigue life of these stainless steel specimens varies greatly when the loading factor ranges from 0.4 to 0.59. The fatigue life of specimen No. 28 with a loading factor of 0.5 is still significantly greater than the fatigue life of the untreated specimen with a loading factor of 0.44, while the dispersion of fatigue life at low loading factors is also quite large. When the loading factor was increased to 0.59, the fatigue life of all specimens decreased rapidly, and the dispersion also decreased. Comparing the fatigue life of laser shock peening parts with those of untreated parts, it can be found that under a high loading factor, laser shock peening can still substantially improve the fatigue life of austenitic stainless steel 1Cr18Ni9Ti by about 100%, and the dispersion is small.

5.3.2 Fatigue Lives of 1Cr11Ni2W2MoV Stainless Steel Currently, the U.S. applies laser shock peening to aircraft engine blades, which greatly improves the safety of aircraft. It is reported that the cost of laser shock peening in the U.S. has been reduced from $100 to $20 per blade owing to the improvement of the laser and the laser shock peening process. In this section, laser shock peening investigations were conducted for 1Cr11Ni2W2MoV martensitic stainless steel material [8, 18] to improve the cracking resistance of martensitic stainless steel compressor blades of an aero-engine model by laser shock peening. 1. Test materials and test equipment 1Cr11Ni2W2MoV martensitic stainless steel compressor blade is an important secondary forging, whose chemical composition is shown in Table 5.4, after a specific heat treatment process of martensitic stainless steel yield strength of 932–1008 MPa, Rockwell hardness (HRC) of 34–36, and Young’s modulus of 196 GPa. The size of the specimen after wire cutting was 20 mm × 20 mm × 3 mm, and the surface was polished. Before the laser shock peening test, the specimens were stress relieved by annealing to remove the surface processing stress. The laser shock peening test was conducted at the National Key Laboratory of High Energy Beam Processing Technology of China Academy of Aviation Manufacturing Technology. The laser shock peening system consists of four parts: laser, optical path adjustment platform, workpiece movement system, and water feeding device. The main technical specifications of the laser are as follows: wavelength 1064 nm; waveform between Gaussian waveform and short rising edge waveform; Table 5.4 Chemical composition of 1Cr11Ni2W2MoV martensitic stainless steel (% (mass fraction)) C

Mn

Si

S

P

Cr

Ni

W

Mo

V

0.10–0.16

1 × 107

1–2

202 MPa

1.313 KN

>1 × 107

1–4

220 MPa

1.430 KN

>1 × 107

Three-spot double-sided laser shock peening specimen No

Fatigue stress σ /MPa

Fatigue load F/kN

Fatigue life/cycle

3–1

300 MPa

1.950 KN

>1 × 107

a nominal composition of Al–5Mg–2Li, and it is a promising light alloy structural material for aerospace and aviation applications [69]. Under the same load-bearing conditions, 1420 Al–Li alloy can be used for riveted structures to reduce the weight of the structure by 15%. 1420 Al–Li alloy is not only suitable for welded structural parts but also for riveted structural parts. In previous research, it was found that laser shock peening of 2024 aluminum alloy rivet holes significantly improved the fatigue life of riveted structures [70]. In this section, laser shock peening tests were conducted on 1420 aluminum–lithium alloy plates, and the effects on their fatigue and surface mechanical properties were investigated [71]. 1. Test conditions The 2.9 mm thick 1420 alloy plate was made into tensile specimens, central circular hole specimens, and metallographic specimens, and the tensile specimens were used to test the static mechanical properties of the test material. The chemical composition of 1420 Al-Li alloy and the measured mechanical properties of 1420 Al-Li alloy plates used in the test are listed in Tables 5.16 and 5.17, respectively. The laser shock peening treatment was performed on some of the fatigue specimens and the high finish surface of the central circular hole and metallographic specimens, using the laser shock peening test device of the University of Science and Technology of China. The laser was a dye-modulated Q neodymium glass laser with a wavelength of 1.06 μm, and the laser output mode was quasi-basic mode with a spot energy close

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5 Evaluations of the Strengthening Effect of the Metals with Laser Shock …

Table 5.16 Chemical composition of 1420 Al–Li alloy Components

Mg

Li

Zr

Fe

Si

H

Al

Composition wt/%

5.0–6.0

1.9–2.3

0.1–0.5

0, so that the wall plate produces concave bending deformation. As the laser energy increases or the thickness of the wall plate decreases, the moment of inertia of the downward motion becomes larger and the concave bending deformation becomes larger. LPF wall plate-induced stress gradient mechanism and shock bending mechanism will have a coupling effect, when the convex bending deformation reaches a maximum, increase the laser energy or reduce the thickness of the wall plate; the role of shock bending mechanism began to strengthen and weaken the role of the stress gradient mechanism, resulting in convex bending deformation began to reduce. When the effect of the shock bending mechanism is further strengthened, the wall plate will produce a change from convex bending deformation to concave bending deformation. When the effect of stress gradient mechanism and shock bending mechanism is comparable, the wall plate will produce flat bending deformation. When the wall plate is too thin, LPF induces wall plate deep drawing phenomenon. (2) Surface Morphology Figure 6.77 shows the surface morphology of Al2024-T351 thin-walled plate formed by longitudinal LPF. From Fig. 6.77a–c, it can be seen that when there is no pad on the back of the thin-walled plate, the aluminum foil on the back of the longitudinal

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Fig. 6.77 Surface morphology of Al2024-T351 thin-walled plate by longitudinal LPF. a 3 mm thick and thin panel, 20 J laser energy, and no pad on the back of the thin panel; b 2 mm thin-walled panel, 20 J laser energy, and no pad on the back of thin-walled panel; c 1 mm thick and thin-walled panel, 20 J laser energy, and no pad on the back of thin-walled panel; d 1 mm thick and thin siding, 25 J laser energy, and a pad on the back of the thin siding

LPF thin-walled plate has a bulging phenomenon, and the height of the bulging increases as the thickness of the thin-walled plate decreases. When there is a pad on the back of the thin-walled plate, the aluminum foil on the back of the longitudinal LPF thin-walled plate is not raised because the reflected wave at the interface of the thin-walled plate and the aluminum foil is small (the shock wave has been transmitted into the pad), or the foil raised on the back of the thin-walled plate is limited by the pad [35]. As can be seen from Fig. 6.77d, there is no bulging of the aluminum foil on the surface of the longitudinal LPF thin-walled plate, which indicates that the reflected wave amplitude inside the thin-walled plate is small, because when LPF thin-walled plate is formed, the aluminum foil-absorbing layer is added on the back of the thin-walled plate to absorb the laser shock wave and avoid the reflected wave from passing into the thin-walled plate. In summary, when laser blasting forms thinwalled plate, the back of the thin-walled plate needs to add aluminum foil-absorbing layer, so as to effectively avoid the internal reflection wave amplitude of thin-walled plate and improve the thin-walled plate strengthening effect. Figure 6.78 shows the three-dimensional surface morphology of the impacted area with coverage of 66%, 128%, 240%, and 357%, respectively, and the surface of the impacted area varies widely for different coverage [73]. The coverage rate equation: C = S · N /A where

(6.22)

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Fig. 6.78 Laser shock surface three-dimensional morphology

A total Shocked Area; S single-spot area; N total number of light spots, the part of light spots outside the shock area is not counted in the coverage. The shock surface with 66% coverage has a discrete distribution of round plastic craters, and due to the uneven distribution of laser beam energy, the plastic craters have a smooth transition from the bottom to the edge with a maximum crater depth of about 20–25 μm; by increasing the coverage, the plastic craters start to lap each other, and the single complete plastic crater disappears; the shock surface with 240% coverage has a round bump, and this bump is caused by multiple. This bump is caused by the extrusion of multiple pot-bottom shaped craters during the lap process. When the coverage reaches 357%, the areas are subjected to multiple shocks to form a relatively flat and reinforced surface.

6.6.2 Upper Limit Value of Process Parameters of Convex Bending Deformation of Mid-Thick Plates 1. Research needs for upper limit value of process parameters The laser energy and laser shock times are the main factors affecting the forming curvature radius of the mid-thick panel in LPF. By establishing the relationship between laser energy, surface morphology of the mid-thick panel, and bending deformation, the upper limit of laser energy for bending deformation of the mid-thick panel is obtained, which lays a foundation for further improving the forming curvature

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radius. By establishing the relationship between laser shock times, forming radius of curvature, arc height, and plastic strain, the saturation value of bending deformation of medium thickness panel was obtained, which could be used as the basis for optimizing process parameters of medium thickness panel in subsequent laser shock forming. By studying the upper limit of laser energy and laser shock times, the forming limit of LPF panel is obtained. In this section, the surface morphology and bending deformation of Al2024-T351 aluminum alloy mid-thick panel formed by longitudinal LPF with different laser energies (15 J–30 J) are studied. The forming curvature radius, arc height, and plastic strain of Al2024-T351 mid-thick plate formed by longitudinal LPF with different laser shock times (1, 2, 4, 8, 16, and 32) are studied. The results lay a foundation for the process of LPF. (1) Test sample and laser shock forming test The selected test material is Al2024-T351 mid-thick panel (Fig. 6.79), the matrix microstructure is shown in Fig. 6.65, and the mechanical properties are shown in Table 6.8. The size of mid-thick panel is 160 mm × 40 mm (length × width), and the thickness t is 5 mm, 8 mm, 11 mm, and 12 mm, respectively. The LPF system (spot size 4 mm × 4 mm) of AVIC Manufacturing Technology Institute was used for longitudinal LPF of Al2024-T351 medium thickness panel, as shown in Figs. 6.64a and 6.80. The shock area L1 was symmetrical about the center line of the length direction of the mid-thick panel, and the forming track was from top to bottom and from left to right. The laser process parameters are as follows: the square spot size is 4 mm × 4 mm, the frequency is 2 Hz, and the moving speed of the spot in X- and Y-directions is 3.4 mm/s. The confinement layer is 1– 2 mm thick deionized water curtain, and the absorption layer is about 0.12 mm thick aluminum foil. Two kinds of longitudinal LPF experiments are as follows: (i) The surface morphology and transverse bending deformation of Al2024-T351 mid-thick panel formed by longitudinal LPF with different laser energies (15 J, 20 J, 25 J, 30 J) are studied. The thickness of the plate is 5 mm, 8 mm, and 11 mm. (ii) The arc height, curvature radius, and plastic strain of Al2024-T351 mid-thick panel formed by longitudinal LPF with different laser shock times (1, 2, 4, 8, 16 and 32) are studied.

Fig. 6.79 Al2024-T351 mid-thick panel

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Fig. 6.80 Schematic diagram of longitudinal LPF Al2024-T351 mid-thick panel

The thickness of mid-thick panel is 12 mm, the shock area is L1 = 90 mm, and the laser energy is 25 J. (2) Test analysis The three-dimensional surface morphology and profile of Al2024-T351 mid-thick panel formed by longitudinal LPF with different laser energies were analyzed by ZYGO Nex View, a three-dimensional white light interference surface topography instrument. Based on the optical platform, the spanwise bending deformation of the mid-thick panel was measured along the center line. The measuring instruments were a dial indicator and support frame. The optical platform is adjusted by an electronic level, as shown in Fig. 6.67. The adjustment parameters are shaft 300 mm and 1div = 0.02 mm/m. The plastic deformation layer depth and plastic strain of Al2024-T351 mid-thick panel formed by longitudinal LPF were measured and analyzed by ultra depth of field microscope VHX-5000 under different shock times. An arc height meter is used to measure the arc height of the Al2024-T351 mid-thick panel (the measuring length is Lc = 60 mm), as shown in Fig. 6.81. 2. Upper limit of laser energy based on surface topography and bending deformation In this section, the spanwise bending deformation of Al2024-T351 mid-thick panel formed by longitudinal LPF under different laser energies is studied, as shown in Fig. 6.81 Arc height meter used to measure the arc height of Al2024-T351 mid-thick panel

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Fig. 6.82. As can be seen from Fig. 6.82a, for the 5 mm thick Al2024-T351 midthick panel, the maximum spanwise bending deformation of Al2024-T351 mid-thick panel formed by longitudinal LPF with 30 J laser energy is 2.03 mm when L1 = 60 mm in the same shock area. The maximum spanwise bending deformation of Al2024-T351 mid-thick panel formed by longitudinal LPF with less than 15 J laser energy is 2.23 mm. As can be seen from Fig. 6.82b, for 8 mm thick Al2024-T351 mid-thick panel, the maximum spanwise bending deformation of Al2024-T351 midthick panel is 1.64 mm when the laser energy is 30 J and the shock zone L1 = 90 mm. The maximum spanwise bending deformation of Al2024-T351 mid-thick panel is 1.89 mm when the laser energy is 25 J and the shock zone L1 = 60 mm. As can be seen from Fig. 6.82c, for 11 mm thick Al2024-T351 mid-thick panel, the maximum spanwise bending deformation of Al2024-T351 mid-thick panel is 0.85 mm when the laser energy is 30 J and the shock area L1 = 120 mm. The maximum spanwise bending deformation of Al2024-T351 mid-thick panel is 1.03 mm when the laser energy is 25 J and the shock area L1 = 90 mm. To sum up, the experimental results in Fig. 6.82 show that when the laser energy is 25 J, the longitudinal LPF of Al2024-T351 mid-thick panel has the maximum transverse bending deformation. Compared with the laser energy of 25 J, the spanwise bending deformation of the Al2024-T351 mid-thick panel is decreased when the laser energy is 30 J. Even when the laser energy is 15 J, the spanwise bending deformation

Fig. 6.82 Spanwise bending deformation of Al2024-T351 mid-thick panel formed by longitudinal LPF with different laser energies. a 5 mm thick; b 8 mm thick; c 11 mm thick

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of Al2024-T351 mid-thick panel is larger than that of Al2024-T351 mid-thick panel when the laser energy is 30 J. Therefore, the maximum laser energy of the spanwise bending deformation of the Al2024-T351 medium thick wall plate in longitudinal LPF is 25 J, that is, the laser power density is 10.4 GW/cm2 . In order to explore the influence mechanism of laser energy on the spanwise bending deformation of Al2024-T351 mid-thick panel formed by longitudinal LPF, the three-dimensional surface morphology of Al2024-T351 mid-thick panel formed by longitudinal LPF with different laser energies was analyzed in this section, as shown in Fig. 6.83. As can be seen from Fig. 6.83b, d, f, when the laser energy is 15 J, 20 J, and 25 J, the pit depth in the lap area of Al2024-T351 mid-thick wall plate formed by longitudinal LPF is greater than that in the non-lap area. However, when the laser energy is 30 J, the lap pit depth of the Al2024-T351 mid-thick panel formed by longitudinal LPF is smaller than that of the non-lap area, that is, the lap area produces a bulge, as shown in Fig. 6.83g. The results show that, compared with the laser energies of 15 J, 20 J, and 25 J, the surface microplastic deformation of Al2024T351 mid-thick panel is reduced by the adjacent spots induced by the longitudinal LPF at 30 J, which can reduce the macroscopic spanwise bending deformation of Al2024-T351 mid-thick panel. In order to further quantitatively analyze the influence of laser energy on the surface profile of the lap area of the Al2024-T351 mid-thick panel in longitudinal LPF, the surface profile of Al2024-T351 mid-thick panel in longitudinal LPF with different laser energies was measured in detail in this section, as shown in Fig. 6.84. As can be seen from Fig. 6.84, when the laser energy is 15 J, 20 J, 25 J, and 30 J, the surface pit depths of the Al2024-T351 mid-thick panels are 15.4 μm, 22.8 μm, 28.6 μm, and 35 μm, respectively. When the laser energy is 15 J, 20 J, and 25 J, the surface pit depth of the lap area is greater than the surface pit depth of the non-lap area (single-spot pit depth). When the laser energy is 25 J, the surface pit depth of the lap area is 44.5 μm, which is greater than the non-lap area surface pit depth of 28.6 μm (single-spot pit depth). When the laser energy is 30 J, the pit depth on the surface of the lap area is similar to that on the surface of the non-shot peening area, and the pit depth on the surface of the adjacent spot is 22.6 μm, which is smaller than that on the front spot, which is 35 μm. It can be seen that when the laser energy is 30 J, the front spot induces a deeper pit depth, resulting in most of the laser shock wave energy of the adjacent spot being consumed in the plastic deformation of the front spot pit edge, and only part of the laser shock wave energy acts on the surface cumulative plastic deformation. Thus, the macroscopic bending deformation of Al2024-T351 mid-thick panel is greatly reduced by LPF based on surface cumulative plastic deformation. Therefore, for Al2024-T351 mid-thick panel, in order to better realize LPF, the maximum laser energy is 25 J, that is, the maximum laser power density is 10.4 GW/cm2 .

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Fig. 6.83 Surface three-dimensional morphology of Al2024-T351 mid-thick panel formed by longitudinal LPF with different laser energies. a, b laser energy 15 J; c–d laser energy 20 J; e–f laser energy 25 J; g Laser energy 30 J

Material parameters such as the surface absorption layer and confinement layer of Al2024-T351 mid-thick panel may affect the bending deformation effect of LPF. Therefore, it seems not accurate to use laser energy as the basis for judging the maximum bending deformation of Al2024-T351 mid-thick panel. LPF is aimed at inducing macroscopic plastic forming of metal materials, and it is caused by cumulative plastic deformation of metal surface induced by a single laser spot. Therefore, the surface pit depth induced by a single laser spot is used as the basis for judging the maximum bending deformation of Al2024-T351 mid-thick panel in LPF. As can be seen from Fig. 6.84, when the laser energy is 25 J, the surface pit depth induced by a single spot is 28.6 μm, and when the laser energy is 30 J, the surface pit depth induced by a single spot is 35 μm. Therefore, for Al2024-T351 mid-thick panel, the bending deformation of Al2024-T351 mid-thick panel is maximum when the surface pitting depth induced by single spot is about 28.6 μm.

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Fig. 6.84 Surface profile of Al2024-T351 by longitudinal LPF with different laser energies. a laser energy 15 J; b laser energy 20 J; c–d laser energy 25 J; e–f laser energy 30 J

3. Upper limit of shock number based on radius of curvature, arc height, and plastic strain LPF is caused by cumulative plastic deformation induced by multiple spot lapping. Because of its advantages of high-amplitude and deep compressive residual stress layer, LPF is suitable for plastic forming of small curvature radius and large-thickness panels. In order to obtain the small curvature radius of large-thickness panel, the forming curvature radius of LPF panel should be studied. The laser shock time is the main factor of the small curvature radius of the LPF panel. Therefore, it is urgent to establish the qualitative relationship between the laser shock times of laser shot forming, the height of the wall arc, and the curvature radius of the forming and obtain

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Fig. 6.85 Schematic diagram of testing the spanwise arc height of the shock region of Al2024-T351 mid-thick panel by longitudinal LPF

the upper limit of the number of impact of LPF. So far, no research results in this field have been published at home and abroad. In this section, longitudinal LPF is performed on the 12 mm thick Al2024-T351 mid-thick panel with multiple shocks. The forming trajectory is shown in Fig. 6.64a. The shock area is unified L1 = 90 mm, and the laser energy is unified 25 J. In order to better reflect the display of spanwise bending deformation of Al2024-T351 midthick panel formed by longitudinal LPF, arc height tester (Fig. 6.81) was used to test the spanwise arc height value of the impact area of Al2024-T351 mid-thick panel formed by longitudinal LPF, as shown in Fig. 6.85. Figure 6.86 shows the spanwise bending deformation profile of Al2024-T351 mid-thick panel formed by longitudinal LPF under multiple shocks. As can be seen from Fig. 6.86, with the increase of the shock times, the maximum bending deformation of Al2024-T351 mid-thick panel in longitudinal LPF gradually increases. The maximum spanwise bending deformations of Al2024-T351 mid-thick panels by 1 shock, 2 shocks, 4 shocks, 8 shocks, 16 shocks, and 32 shocks are 1.07 mm, 1.61 mm, 2.87 mm, 4.34 mm, 6.39 mm, and 8.38 mm, respectively. (1) Spanwise bending deformation Figure 6.87 shows the relationship between the shock times of Al2024-T351 midthick panel in longitudinal LPF and the arc height in the shock region. The measurement interval of arc height is 60 mm. As can be seen from Fig. 6.87, the spanwise arc heights of the shock area of the longitudinal LPF Al2024-T351 mid-thick panel with one shock, 2 shocks, 4 shocks, 8 shocks, 16 shocks, and 32 shocks are 0.2 mm, 0.34 mm, 0.55 mm, 0.85 mm, 1.24 mm, and 1.62 mm, respectively. The increase rates of spanwise arc height were 70%, 61.8%, 54.5%, 45.9%, and 30.6%, respectively. This indicates that with the doubling of the shock times, the increase rate of the spanwise arc height in the impact region of Al2024-T351 mid-thick panel by longitudinal LPF is no less than 10%, that is, the saturation value is not reached.

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Fig. 6.86 Spanwise bending deformation profile of Al2024-T351 mid-thick panel formed by longitudinal LPF under multiple shock times

Fig. 6.87 Relation curve of impact frequency of Al2024-T351 mid-thick panel and impact area arc height during longitudinal LPF

Therefore, the spanwise arc height may not be used to define the upper limit of the number of LPF. In order to better realize the visibility of LPF, this section establishes the relationship curve between the shock times of longitudinal LPF Al2024-T351 midthick panel and the radial forming curvature radius of the shock region, as shown in Fig. 6.88. As can be seen in Fig. 6.88, the curvature radius of the shock area of the longitudinal LPF Al2024-T351 mid-thick panel with one shock, 2 shocks, 4 shocks, 8 shocks, 16 shocks, and 32 shocks is 2250 mm, 1323.5 mm, 818.2 mm, 529.4 mm,

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Fig. 6.88 Relation curve between shock times of Al2024-T351 thick panel and curvature radius of impact area in longitudinal LPF

362.9 mm, and 277.8 mm respectively. The reduction rates of the curvature radius of spanwise forming are 41.2%, 38.2%, 35.3%, 31.5%, and 23.4%, respectively. This indicates that with the increase of the shock times by 2 times, the decreased rate of curvature radius of longitudinal LPF Al2024-T351 mid-thick panel in the shock region is not less than 10%, that is, it does not reach the saturation value. Therefore, the forming curvature radius of the impact region may not be used to define the upper limit of the shock times of LPF. The relationship curve between the shock times and the spanwise plastic strain of Al2024-T351 mid-thick panel formed by longitudinal LPF was analyzed, as shown in Fig. 6.89. As shown in Fig. 6.89, the spanwise plastic strain of Al2024-T351 midthick panel formed by longitudinal LPF with 1 shock, 2 shocks, 4 shocks, 8 shocks, 16 shocks, and 32 shocks is about 0.15%, 0.31%, 0.53%, 0.83%, 1.4%, and 1.95%, respectively. In order to ensure the subsequent assemblability of LPF wing panels, the spanwise plastic strain should be controlled at about 0.5% because the spanwise length of wing panels is tens of meters. Therefore, the upper limit of the number of shock times of the longitudinal LPF Al2024-T351 mid-thick panel is 4 impacts. (2) Chordal bending deformation Figure 6.90 shows the relationship between the shock times of Al2024-T351 midthick panel and the chord arc height in the shock region. The measurement interval of arc height value is Lc = 40 mm. According to Fig. 6.90, it can be seen that the chordal arc heights of the shock area of Al2024-T351 mid-thick panel formed by longitudinal LPF with 1 shock, 2 shocks, 4 shocks, 8 shocks, 16 shocks, and 32 shocks are 0.01 mm, 0.06 mm, 0.18 mm, 0.39 mm, 0.93 mm, and 1.87 mm, respectively. In this section, the length to width ratio of Al2024-T351 mid-thick panel is 4:1, which is mainly used to reduce chord bending deformation. The chord arc height induced by 8 shocks is 0.39 mm, which is relatively large, while the chord arc height induced by 4 shocks is 0.18 mm, which is a critical value. Therefore, the upper limit of the shock times of Al2024-T351 mid-thick panel is 4 times.

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Fig. 6.89 Relation curve of shock times and extensive-plastic strain of Al2024-T351 mid-thick panel during longitudinal LPF (shock area length is 78.8 mm)

Fig. 6.90 Relation curve of shock times of Al2024-T351 mid-thick panel in longitudinal LPF and chordal arc height of shock region (Lc = 40 mm)

Figure 6.91 shows the relationship between the shock times of Al2024-T351 thick panel and the curvature radius of the shock area. Figure 6.91 shows 1 shock, 2 shocks, 4 shocks, 8 shocks, 16 shocks, and 32 shocks longitudinal LPF Al2024-T351 thick panel impact area choroidal forming curvature radius of 20,000 mm, 3333.3 mm, 1111.1 mm, 512.8 mm, 215.1 mm, and 107 mm, respectively. The reduction rates of curvature radius in the shock area are 83.3%, 66.7%, 53.8%, 58.1%, and 50.3%,

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Fig. 6.91 Relation curve between shock times of Al2024-T351 thick panel and curvature radius of shock area of longitudinal LPF

respectively. This indicates that the curvature radius in the shock area of Al2024T351 mid-thick panel decreases rapidly with the increase in shock times. In order to improve the curvature radius of the longitudinal LPF of thick wall plate in a free state, the upper limit of the number of shock times of the longitudinal LPF of Al2024-T351 is 4. Figure 6.92 shows the relationship between the shock times and the chord plastic strain of Al2024-T351 thick panel formed by longitudinal LPF. As can be seen from Fig. 6.92, the chord plastic strain of Al2024-T351 mid-thick panel formed by longitudinal LPF with 1 shock, 2 shocks, 4 shocks, 8 shocks, 16 shocks, and 32 shocks is about 0.32%, 0.78%, 1.58%, 2.98%, 5.93%, and 10.02%, respectively. In order to ensure the subsequent assembly of the wing panel formed by LPF, the chord plastic strain should be controlled at about 1% because the chord length of the wing panel is several meters. According to the experimental results shown in Fig. 6.92, the upper limit of the number of shock times of longitudinal LPF can be defined by using the chordal plastic strain control requirement (chordal plastic strain control is about 1%) of the wing wall. Therefore, the upper limit of the number of shocks is 4. (3) Plate thickness direction plastic deformation Figure 6.93 shows the depth of influence layer at the edge of Al2024-T351 mid-thick panel formed by longitudinal LPF under different shock times. As can be seen from Fig. 6.93, the depth of affected layer at the edge of Al2024-T351 mid-thick panel formed by longitudinal LPF with 1 shock, 2 shocks, 4 shocks, 8 shocks, 16 shocks, and 32 shocks is about 1.7 mm, 4.94 mm, 5 mm, 6.41 mm, 7.75 mm, and 8.53 mm, respectively. This indicates that the depth of the edge affected layer of Al2024-T351 mid-thick panel increases with the increase of the number of shock times, and the depth of the edge affected layer of Al2024-T351 mid-thick panel is 6.41 mm, which is more than half of the plate thickness (12 mm). Figure 6.94 shows the relationship between the shock times and the plastic strain in the thickness direction of Al2024-T351 mid-thick panel formed by longitudinal

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Fig. 6.92 The relationship between the shock times and the chordal plastic strain of Al2024-T351 mid-thick panel formed by longitudinal LPF

Fig. 6.93 Influence layer depth at the edge of Al2024-T351 mid-thick panel by longitudinal LPF under different shock times. a one shock; b 2 shocks; c 4 shocks; d 8 shocks; e 16 shocks; f 32 shocks

LPF. As can be seen from Fig. 6.94, the plastic strain in the thickness direction of Al2024-T351 mid-thick panel formed by longitudinal LPF with one shock, 2 shocks, 4 shocks, 8 shocks, 16 shocks, and 32 shocks is about 4.7%, 7%, 10.5%, 11.3%, 13.6%, and 16%, respectively.

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Fig. 6.94 The relationship between the shock times and the plastic strain in the thickness direction of Al2024-T351 mid-thick panel formed by longitudinal LPF

6.6.3 Convex Bending Deformation and Mechanical Property of Mid-Thick Plates 1. Research needs for convex bending deformation and mechanical property To meet the plastic forming requirements of the small curvature radius (3–8 m) of the variable section panel, firstly, the forming ability of the mid-thick panel in LPF was studied to obtain the forming ability of the mid-thick panel in LPF, which laid the technological foundation for the LPF of the ribbon panels. Secondly, the bending deformation law of Al2024-T351 mid-thick panel formed by LPF with different process parameters was studied, including different laser process parameters, different forming trajectories, and different material parameters, which laid the process foundation for optimizing process parameters of LPF panel. Finally, the mechanical properties of the Al2024-T351 mid-thick panels formed by LPF were studied to lay a technological foundation for achieving the high fatigue life of the LPF panel. (1) Test specimen and LPF experiment The selected test material is Al2024-T351 mid-thick panel (the length direction is rolling direction). The matrix microstructure is shown in Fig. 6.65, and the mechanical properties are listed in Table 6.8. The sizes of the two types of samples are as follows: (i) the size of the mid-thick panel used in free-state longitudinal LPF is 160 mm length × 40 mm width, and the thickness t is 5 mm, 8 mm, and 11 mm, respectively. The forming trajectory is shown in Fig. 6.64a. (ii) The dimensions of mid-thick panels used in different prestressed longitudinal LPF are 320 mm length × 102 mm width × 24 mm thickness, and the forming trajectory is shown in Fig. 6.95. The process parameters of longitudinal LPF are as follows: the size of the square spot is 4 mm × 4 mm, the frequency is 2 Hz, and the moving spacing of the spot

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Fig. 6.95 Schematic diagram of Al2024-T351 mid-thick panel formed by longitudinal LPF with different prestressed

in the x-direction and y-direction is 3.4 mm. A beam shaping device was used to convert the flat-topped circular spot output by the laser into a flat-topped square spot radiating on the surface of the mid-thick panel [51]. The spanwise forming law of the Al2024-T351 mid-thick panel was studied by orthogonal experimental design. The mechanical properties of the Al2024-T351 mid-thick panels formed by prestressed longitudinal LPF were studied by testing and analyzing methods. The prestressed LPF device is composed of a YAG laser, a beam shaping device, a prebending device, and an industrial manipulator, as shown in Fig. 6.96. The prebending device is composed of a dynamic loading device and a supporting loading device. Both sides of the mid-thick panel are supported by a supporting loading device, and the span is L1 = 280 mm. The dynamic loading device implements elastic prestressing load on the center of the length direction of the mid-thick panel, and the prestressing bending deformation is ω. (2) Test analysis Based on an optical platform, the spanwise bending deformation of the Al2024-T351 mid-thick panel formed by LPF was measured along the center line by dial indicator and support frame. The optical platform is adjusted by an electronic level, as shown in Fig. 6.68. The adjustment parameters are shaft 300 mm and 1div = 0.02 mm/m. An arc height meter (Fig. 6.83) was used to measure the spanwise arc height in the impact area. The surface morphology of the Al2024-T351 mid-thick panel was analyzed using ZYGONex View, a three-dimensional white light interferometer. Talysurf PGI 1230, a surface topography measuring system, was used to test and analyze the surface roughness of Al2024-T351 mid-thick panels formed by different prestressed LPF. The test line length was 20 mm. An X-ray diffractometer was used to measure the residual stress on the upper and lower surfaces of the impact area of the mid-thick panel. The X-ray diameter was ϕ2mm, the Cr-Kα target, the {311} crystal plane, and the diffraction angle 2θ was 139°. The residual stress test points were 20 mm apart, and the intermediate test points of the upper and lower surfaces were located at the intersection of the spanwise chord to the center line. Wire cut the metallographic specimen in the impact area, and then inlay, grind, and polish it. Finally, professional etchants were used to corrode the samples on the surface.

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Fig. 6.96 Prestressed LPF device. a prestressed longitudinal LPF device; b schematic diagram of elastic prestressed loading

The corrosions were 95% of H2 O, 2.5% of HNO3 , 1.5% of HCI, and 1% of HF, and the corrosion was carried out at room temperature for 5 s. The microstructure of the whole section of the metallographic sample was analyzed by ZEISS optical microscope, and the grain size was calculated by Nano Measurer 1.2 software. 2. Formability of convex bending deformation Figure 6.97 shows the stiffened panel formed by LPF. Macro-bending deformation is induced by large-area LSP. The arc height of bending deformation is shown in Fig. 6.97a. The results show that LPF can realize the forming of the ribbed panel. At the same time, a large number of square pits are formed on the surface of the impact area of the LPF ribbed panel, as shown in Fig. 6.97b. To obtain the forming ability of the Al2024-T351 mid-thick panel by LPF, the Al2024-T351 mid-thick panel is formed by longitudinal LPF under different

Fig. 6.97 LPF ribbed panel. a arc height of bending deformation; b surface topography

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Fig. 6.98 Longitudinal LPF of Al2024-T351 mid-thick panel with different prestressed stress. a thickness 24 mm, curvature radius 3.3 m, prestress 269.23 MPa; b thickness 20 mm, curvature radius 2.8 m, prestress 246.8 MPa; c thickness 15 mm, curvature radius 1.6 m, prestress 336.54 MPa

prestressed conditions, as shown in Fig. 6.98. LPF process parameters: 1 impact, frequency 2 Hz, laser energy 20 J, andimpact area 240 mm × 102 mm. As can be seen from Fig. 6.98, the forming ability of the prestressed longitudinal LPF Al2024-T351 mid-thick panel has reached the plastic forming requirement of the wing panel. Figure 6.99 shows the distribution of residual stress in the depth direction of the LPF panel. Large-area LSP induces macroscopic bending deformation of the panel, as shown in Fig. 6.99a. The macroscopic bending deformation mechanism is that LSP induces high-amplitude and deep residual compressive stress layer on the surface of the panel, as shown in Fig. 6.99b, resulting in a bending moment M greater than the inherent constraint of the panel. The bending deformation and bending stress occur in the depth direction of the panels, as shown in Fig. 6.99c. Combined with the bending stress induced by bending deformation and the compressive residual stress induced by LSP, the residual stress in the depth direction of the LPF panel is formed, as shown in Fig. 6.99d. 3. Primary and secondary factors affecting radius of curvature and its influencing laws The orthogonal experiment design is based on probability theory and mathematical statistics. Through a reasonable arrangement of the experiment scheme, the experimental data can have a suitable mathematical model, reduce the influence of random error, improve the accuracy and reliability of the experimental results, and reduce the number of experiments. In this section, the impact zone forming curvature radius of longitudinal LPF Al2024-T351 mid-thick panel was studied by orthogonal experimental design to reduce the impact zone forming curvature radius of longitudinal LPF Al2024-T351 mid-thick panel. The influence factors include laser energy, plate thickness, and impact zone. To obtain the influence law of a single factor on the forming

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Fig. 6.99 Residual stress distribution in the depth direction of LPF panel. a LPF panel; b before bending and deformation; c bending deformation; d after bending and deformation

Table 6.9 Experimental factors and levels of longitudinal LPF

Factors levels

Plate thickness A/mm

Laser energy B/mm

Impact area (L1 ) C/mm2

1

11

15

120 × 40

2

8

25

90 × 40

3

5

20

60 × 40

curvature radius of the impact region of the longitudinal LPF Al2024-T351 medium thickness panel, three main levels, namely three factor 3 levels, were selected, as shown in Table 6.9. Orthogonal experimental design and test results of longitudinal LPF Al2024T351 mid-thick panel are listed in Table 6.10. The test pieces of longitudinal LPF Al2024-T351 mid-thick panel are shown in Fig. 6.100. Table 6.10 defines the parameters as follows: Ii = Sum of index value yi corresponding to the digit “1” in column i of Table 6.10; IIi = Sum of index value yi corresponding to the digit “2” in column i of Table 6.10; IIIi = Sum of index value yi corresponding to the digit “3” in column i of Table 6.10; T = sum of all experimental radius of curvature; Ri is the range of the factor; Ri = max (Ii , IIi , IIIi ) -min (Ii , IIi , IIIi ); RA, RB, and RC are obtained by calculation; and the primary and secondary order is determined.

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Table 6.10 Orthogonal test design and test results of longitudinal LPF Al2024-T351 mid-thick panel Factor

Plate thickness/mm

Laser energy/J

Impact area size/mm2

Index value yi (radius of curvature of spanwise forming R)/m

No. 1

11

15

120 × 40

2.8

No. 2

11

25

90 × 40

2.0

No. 3

11

20

60 × 40

2.3

No. 4

8

15

90 × 40

1.4

No. 5

8

25

60 × 40

0.98

No. 6

8

20

120 × 40

1.2

No. 7

5

15

60 × 40

0.82

No. 8

5

25

120 × 40

0.74

No. 9

5

20

90 × 40

0.78

Ii

7.1

5.02

4.74

II i

3.58

3.72

4.18

III i

2.34

4.28

4.1

Ri

4.76

1.3

0.64

T

13.02

Panel number

According to the orthogonal experimental design scheme in Table 6.10, the longitudinal LPF test of the Al2024-T351 mid-thick panel was carried out in a free state, and then the transverse bending deformation profile of the experimental piece was tested along the symmetry line. Figure 6.101 shows the spanwise bending deformation profile of the Al2024-T351 mid-thick panel formed by longitudinal LPF in the free state. As can be seen from Fig. 6.101, with the increase of laser energy or impact area, hmax of longitudinal LPF of Al2024-T351 mid-thick panel increases gradually. For 5 mm thick sheet metal, the maximum bending deformation hmax of Al2024-T351 No.7, No.9, and No.8 in the longitudinal LPF process is about 2.23 mm, 2.98 mm, and 3.24 mm, respectively. For 8 mm thick sheet metal, the maximum spanwise bending deformation hmax of Al2024-T351 mid-thick panels No.4, No.6, and No.5 is about 1.55 mm, 1.89 mm, and 2.15 mm, respectively. For 11 mm thick sheet metal, the maximum bending deformation hmax of No.1, No.3, and No.2 of Al2024-T351 midthick panels formed by longitudinal LPF is about 0.74 mm, 0.93 mm, and 1.03 mm, respectively. To better reflect the development of spanwise bending deformation of Al2024T351 mid-thick panel in longitudinal LPF, the spanwise forming curvature radius of the impact area is used to represent the spanwise bending deformation in the impact area. Therefore, the arc height of the longitudinal LPF Al2024-T351 mid-thick panel was measured by arc height meter, and the test interval was Lc = 60 mm. The radius of curvature of the impact area can be calculated by Eq. (6.21). By calculating the

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Fig. 6.100 Test piece of longitudinal LPF Al2024-T351 mid-thick panel

spanwise forming curvature radius of the impact area, the spanwise forming curvature radius of the impact area of Al2024-T351 mid-thick panel formed by longitudinal LPF with different process parameters is obtained. For Longitudinal LPF of Al2024T351 mid-thick panel impact area of the spanwise arc height, see Table 6.11. As can be seen from Table 6.11, for 5 mm thick sheet metal, longitudinal LPF of Al2024T351 mid-thick panels No.7, No.9, and No.8, the arc height of the impact area is about 0.55 mm, 0.58 mm, and 0.61 mm, respectively. For 8 mm thick sheet metal, the longitudinal LPF of Al2024-T351 mid-thick panels No.4, No.6, and No.5, impact area arc height values are about 0.33 mm, 0.38 mm, and 0.46 mm, respectively. For 11 mm thick sheet metal, the longitudinal LPF of Al2024-T351 mid-thick panels No.1, No.3, and No.2, impact area arc height values are about 0.16 mm, 0.2 mm, and 0.23 mm, respectively. For 5 mm thick sheet metal, longitudinal LPF of Al2024T351 mid-thick panels No.7, No.9, and No.8 in the impact area of the radial forming curvature radius is about 0.82 m, 0.78 m, and 0.74 m, respectively. For 8 mm thick sheet metal, longitudinal LPF of Al2024-T351 mid-thick panels No.4, No.6, and No.5 impact area of the radial forming curvature radius is about 1.4 m, 1.2 m, and 0.98 m, respectively. For 11 mm thick sheet metal, longitudinal LPF of Al2024-T351 mid-thick panel No.1, No.3 and No.2 impact area of the radial forming curvature radius is about 2.8 m, 2.3 m, and 2.0 m, respectively.

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Fig. 6.101 Profile of spanwise bending deformation of Al2024-T351 mid-thick panel formed by longitudinal LPF in the free state. a 5 mm thick; b 8 mm thick; c 11 mm thick

Table 6.11 Longitudinal LPF of Al2024-T351 mid-thick panel impact area of spanwise arc height and spanwise forming radius of curvature

Panel number

Spanwise arc height/mm

Radius of forming curvature R/m

No.1

0.16

2.8

No.2

0.23

2.0

No.3

0.20

2.3

No.4

0.33

1.4

No.5

0.46

0.98

No.6

0.38

1.2

No.7

0.55

0.82

No.8

0.61

0.74

No.9

0.58

0.78

The spanwise forming radius of the impact area of Al2024-T351 mid-thick panels for longitudinal LPF is filled into Table 6.10. According to 6.10, RC < RB < RA , that is, impact area is the main factor, laser energy is the second factor, and plate thickness is the second factor for reducing the radial forming curvature radius of the impact area of the longitudinal LPF Al2024-T351 mid-thick panel.

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6 Strengthening Processes and Effect Evaluations of Airplane Structures …

Figure 6.102 shows the influence law of various factors of Al2024-T351 midthick panel on the spanwise forming radius of the impact area. As can be seen from Fig. 6.102, the impact area has the most obvious influence on the radial forming curvature radius of the impact area in the longitudinal LPF process of Al2024-T351 mid-thick panels, followed by the influence of laser energy and the influence of plate thickness. As can be seen from Fig. 6.102a, with the increase of plate thickness, the impact area of longitudinal LPF Al2024-T351 mid-thick panel has a gradual increase in the radial forming curvature radius, that is, the two are directly proportional. As can be seen from Fig. 6.102b, with the increase of laser energy, the impact area of longitudinal LPF Al2024-T351 mid-thick panel has a gradual decline in the radial forming curvature radius, that is, the two are inversely proportional. As can be seen from Fig. 6.102c, with the increase of impact area, the radial forming curvature radius of the impact area of Al2024-T351 mid-thick panel in longitudinal LPF gradually increases, that is, the two are directly proportional. 4. Different prestressed bending deformation mechanisms and their influence on mechanical properties To improve the forming limit of LPF and reduce the forming curvature radius, the prestressed LPF was carried out by the elastic prestressing method. Firstly, the bending deformation mechanism of the thick panel in different elastic prestressed

Fig. 6.102 Influence law of various factors of Al2024-T351 mid-thick panel on the spanwise forming radius of impact area. a plate thickness; b laser energy; c impact zone

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Fig. 6.103 Relationship between curvature radius and arc height

longitudinal LPF is studied. Secondly, the effects of free-state and prestressed-state longitudinal LPF on the surface integrity and mechanical properties of mid-thick panels are compared. The research results lay the technological foundation for the LPF of wing panels. Figure 6.103 shows the relationship between the radius of curvature R and arc height ha . The formula for calculating linear strain ε of sheet metal surface is ε=

t × 100% 2R

(6.23)

The formula for calculating the prebending moment M [78, 79]: M=

EW t3 EI = R 12R

(6.24)

The formula of surface tensile stress σ of precurved thick plate: σ =

M ·t 2I

(6.25)

where t E I W

plate thickness; elastic modulus; moment of inertia; plate width.

(1) Different bending and deformation mechanisms of prestressing In this subsection, elastic prestressed longitudinal LPF is carried out for the Al2024T351 mid-thick panel, and the mechanism of prestressed improving the spanwise forming ability of longitudinal LPF is studied. The size of the mid-thick panel is 320 mm × 102 mm × 24 mm (length × width × thickness). Spanwise elastic prebending arc height ω = 2 mm, equal to ha = 2 mm, prestressed loading span L1 = 280 mm, equal to Lc = 280 mm. By Eq. (6.21) and Eq. (6.23)–(6.25), the

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Table 6.12 Relationship between spanwise prebending loading and spanwise curvature radius Prebending spanwise arc height ha / mm

Prebending moment M/N.mm

Surface residual stress σ /MPa

Prebending spanwise plastic strain ε/%

Radial curvature of prebending R/mm

Spanning result arc height of impact area/mm

Spanning result radius of curvature of impact area/m

0

0

0

0

∞

0.15

8.3

2

1,757,524.1

179.5

0.24%

4900

0.33

3.8

linear strain ε, the prebending moment M, and the surface residual tensile stress σ of the Al2024-T351 mid-thick panel under prestressing are obtained, and the results are listed in Table 6.12. The surface tensile residual stress of the Al2024-T351 mid-thick panel under prestressed loading is about 179.5 MPa, which is smaller than the yield strength of the Al2024-T351 mid-thick panel, indicating that prestressed loading belongs to the range of elastic loading. Figure 6.104 shows the spanwise bending deformation profile of the Al2024T351 mid-thick panel formed by prestressed longitudinal LPF. As can be seen from Fig. 6.104, convex bending deformation of the Al2024-T351 mid-thick panel is induced by longitudinal LPF forming in two initial states, which is attributed to the fact that the residual compressive stress layer of Al2024-T351 mid-thick panel induced by LPF is smaller than the plate thickness [55]. The maximum arc height in the direction of prestressed LPF Al2024-T351 medium thick wall is 2.46 mm, which is larger than 1.18 mm in the direction of free longitudinal LPF because the prebending moment M induced by prestressed LPF is larger. Arc height (Lc = 100 mm) and curvature radius of longitudinal LPF Al2024-T351 mid-thick panel were measured by arc height meter under two initial conditions, and the influence of prestress on longitudinal LPF Al2024-T351 mid-thick panel spanwise bending deformation was evaluated. The resulting arc height and curvature

Fig. 6.104 Transverse bending deformation profile of Al2024-T351 mid-thick panel formed by prestressed longitudinal LPF

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327

radius of longitudinal LPF Al2024-T351 mid-thick panels are listed in Table 6.12. As can be seen from Table 6.12, the resulting arc heights in the impact area of the prestressed longitudinal LPF and LPF Al2024-T351 mid-thick panels are 0.33 mm and 0.15 mm, respectively, and the resulting curvature radii are 3.8 m and 8.3 m, respectively. The results show that the elastic prestress loading can significantly reduce the forming radius of the impact area of the Al2024-T351 mid-thick panel and improve the forming ability of the Al2024-T351 mid-thick panel. Two-way forming is to consider the influence of transverse plastic forming on longitudinal plastic forming of LPF sheet metal. The forming theory of LPF sheets based on bending moment can be used to study the forming mechanism of prestressed LPF sheets. The formula [80] for calculating the curvature radius R of laser LPF metal is R=

E(t − δ)3 1 · 12(1 − μ) M '

(6.26)

According to the calculation formula of pure bending in two directions of sheet metal and the schematic diagram of bending deformation of prebending sheet metal as shown in Fig. 6.105, the calculation formula of curvature radius Rx in the prebending direction of prestressing LPF sheet metal is as follows: Rx =

1 E(t − δε )3 · ' 12 MY − μM X'

1−ν RX M ' t − δε 3 = ) · ' ·( M' R t −δ 1 − ν MX' MY Y

where δ ν

depth of compressive stress layer; Poisson’s ratio;

Fig. 6.105 Schematic diagram of the bending deformation of the prebent plate

(6.27) (6.28)

328

M' δε M 'y M 'x

6 Strengthening Processes and Effect Evaluations of Airplane Structures …

bending moment of the sheet at the perimeter of the unit; depth of compressive stress layer of prestressed LPF sheet metal; bending moment in prebending direction; bending moment in vertical prebending direction.

From Eqs. (6.26) to (6.28), it can be seen that, compared with free-state longitudinal LPF, the key factors for obtaining a smaller resulting curvature radius of prestressed longitudinal LPF Al2024-T351 mid-thick panels are the depth of residual compressive stress layer and resulting bending moment. Therefore, the formation mechanism of the small curvature radius of the thick panel in prestressed longitudinal LPF Al2024-T351 is as follows: (i) The surface residual tensile stress induced by elastic prestressing is beneficial to the extension of the surface material of Al2024T351 mid-thick panel in prestressed LPF, thus increasing the depth and amplitude of the residual compressive stress layer (δ < δε ). (ii) The amplitude of the surface residual tensile stress of Al2024-T351 is 1/μ times larger than that of the surface residual tensile stress of AL2024-T351. Therefore, the longitudinal compressive residual stress amplitude of the prestressed longitudinal LPF Al2024-T351 mid-thick panel is larger than that of the chord compressive residual stress amplitude. Thus, the transverse bending moment M’y is greater than the chordal bending moment M’x (M’x < M’y ). In addition, the longitudinal bending moment M’y of the prestressed longitudinal LPF Al2024-T351 mid-thick panel is greater than that M’ (M’ < M’y ) of the longitudinal LPF Al2024-T351 mid-thick panel in a free state. Therefore, according to Eq. (6.27), the curvature radius Rx of the prestressed longitudinal LPF Al2024-T351 mid-thick panel is smaller than the curvature radius R (Rx < R) of prestressed longitudinal LPF Al2024-T351 mid-thick panel. (2) Surface roughness Figure 6.106 shows the surface morphology of the Al2024-T351 mid-thick panel formed by single-spot LPF. It can be seen from Fig. 6.106b, c that square pits are formed on the surface of Al2024-T351 mid-thick panel by single-spot LPF, and the maximum depth H of pits in the x-direction and y-direction are 53 μm and 49 μm, respectively. 0.1H away from the upper limit value of the pits was defined as the measurement reference line, and the distance between the intersection point of the measurement reference line and the contour line of the pits was defined as the surface square pit size of single-spot LPF Al2024-T351. Therefore, the size of square spot pits in the x-direction and y-direction on the surface of the Al2024-T351 mid-thick panel is 4.6 mm. In order to reduce the surface roughness of the Al2024-T351 midthick panel by LPF, the distance between the x-direction and y-direction of the surface of the Al2024-T351 mid-thick panel is 3.4 mm. Figure 6.107 shows the surface peak and valley heights of the Al2024-T351 midthick panel formed by longitudinal LPF in two initial states. As can be seen from Fig. 6.107, compared with the matrix material, the peak-valley height of the Al2024T351 mid-thick panel formed by longitudinal LPF in a free state is higher, and the peak-valley height of the Al2024-T351 mid-thick panel formed by prestressed longitudinal LPF is similar to that of the Al2024-T351 mid-thick panel formed by

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Fig. 6.106 Surface morphology of Al2024-T351 mid-thick panel by single-spot LPF. a threedimensional surface topography; b contour lines in the x-direction; c y-direction contour

longitudinal LPF in the free state. The reason is that longitudinal LPF induces higher pitting depth on the Al2024-T351 mid-thick panel surface. The peak and valley heights of the substrate are about −1.25–0.75 μm, as shown in Fig. 6.107a. The peak-valley height of the Al2024-T351 mid-thick panel is about −3.5–4 μm, as shown in Fig. 6.107b. The peak-valley height of the Al2024-T351 mid-thick panel is about v4.5–5.5 μm, as shown in Fig. 6.107c. Table 6.13 lists the surface roughness values of the Al2024-T351 mid-thick panel formed by longitudinal LPF in two initial states. Compared with the matrix material, the surface roughness of the Al2024-T351 mid-thick panel is significantly increased by longitudinal LPF, which is mainly due to the serious plastic deformation of the Al2024-T351 mid-thick panel surface induced by longitudinal LPF. The surface roughness values of the matrix material are Ra 0.187 μm and Rz 1.21 μm, and the surface roughness values of the Al2024-T351 mid-thick panels are Ra 0.416 μm and Rz 2.43 μm. The surface roughness values of Al2024-T351 mid-thick panels are Ra 0.558 μm and Rz 3.11 μm. (3) Sectional microstructure To obtain the influence of longitudinal LPF under different prestressed states on the section microstructure of the Al2024-T351 mid-thick panel, the entire section microstructure of the Al2024-T351 mid-thick panel was analyzed, as shown in Figs. 6.108, 6.109 and 6.110. As can be seen from Figs. 6.108, 6.109 and 6.110,

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6 Strengthening Processes and Effect Evaluations of Airplane Structures …

Fig. 6.107 Longitudinal laser peak-valley surface heights of Al2024-T351 mid-thick panel formed in two initial states. a base material; b free-state longitudinal LPF; c prestressed longitudinal LPF

Table 6.13 Surface roughness values of Al2024-T351 mid-thick panel formed by longitudinal LPF under two initial conditions

Conditional state

Ra /μm

Rz /μm

Base material

0.187

1.21

LPF area

0.416

2.43

Prestressed LPF area

0.558

3.11

the entire section microstructure of Al2024-T351 mid-thick panel formed by longitudinal LPF under different prestressed states contains five regions, which are labeled as region I of refined grain, region II of elongated grain, region III of the original grain, region IV of elongated grain, and region V of refined grain. For LPF, the I and V regions are mainly concerned. For the base material, the formation of zone I is caused by the rolling of the top surface of the sheet. For LPF materials, the formation of zone I is caused by the grain refining effect induced by a laser shock wave, indicating that the grain refining zone I is a high-amplitude residual compressive stress. For the base material, the formation of the V zone is caused by the rolling of the bottom surface of the sheet. For LPF, the formation of the V zone is caused by the bending deformation of the Al2024-T351 mid-thick panel caused by the negative bending moment induced by large-area LPF and the high gradient residual compressive stress in the

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331

Fig. 6.108 Microstructure of the entire section of Al2024-T351 mid-thick panel

depth direction, which results in the extrusion plastic deformation of V zone and the formation of grain refining V zone. The V region of refined grain is the compressive residual stress layer. To further analyze the influence of longitudinal LPF on grain refinement of Zone I and Zone V of Al2024-T351 mid-thick panel under different prestressed states, the microstructure of Zone I and Zone V are analyzed in detail. The effect of longitudinal LPF on grain refinement in Zone I of Al2024-T351 mid-thick panel under different prestressed states was compared and analyzed. Figure 6.111 shows the upper surface microstructure of the Al2024-T351 mid-thick panel formed by two longitudinal LPF processes. In the upper surface microstructure, the severe plastic deformation layer (SPD) and mild plastic deformation layer (MPD) are generated. As can be seen from Fig. 6.111a, the thickness of the SPD layer and MPD layer of substrate material caused by rolling is about 820 μm and 888 μm, respectively. As shown in Fig. 6.111b, the depths of the SPD layer and MPD layer in the upper layer of Al2024-T351 mid-thick panel formed by two longitudinal LPF processes in the free state are about 946 μm and 959 μm, respectively. As can be seen from Fig. 6.111c, the depth of the SPD layer and MPD layer on the upper layer of Al2024-T351 mid-thick panel formed by prestressed twin-longitudinal LPF

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6 Strengthening Processes and Effect Evaluations of Airplane Structures …

Fig. 6.109 Microstructure of the entire section of Al2024-T351 mid-thick panel formed by two longitudinal LPF in the free state

Fig. 6.110 Microstructure of the entire section of Al2024-T351 mid-thick panel formed by prestressed two longitudinal LPF

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Fig. 6.111 Upper surface microstructure of Al2024-T351 mid-thick panel formed by two longitudinal LPF. a base material; b free-state LPF; c prestressed LPF

is about 1171 μm and 1286 μm, respectively. The SPD layer and MPD layer of the Al2024-T351 mid-thick panel formed by two longitudinal LPF are thicker than that of the substrate, which is due to the severe plastic strain induced by the laser shock wave. Compared with the free state, the SPD layer and MPD layer of the prestressed longitudinal LPF Al2024-T351 mid-thick panels are thicker. The reason is that the plastic strain of the upper surface layer of the LPF Al2024-T351 mid-thick panel is increased by elastic prestress loading. The plastic strain gradually decreases and the grain size gradually increases with the distance from the upper surface. Similar results have been reported [81]. In addition, the average grain size of the SPD and MPD layers of Al2024-T351 is smaller than that of the substrate. Compared with the free state, the average grain size of the SPD layer and MPD layer in the prestressed longitudinal LPF Al2024-T351 mid-thick panel is smaller. The average grain size of the SPD layer and MPD layer is about 72.34 μm and 131.23 μm, respectively. The average grain sizes of the SPD layer and MPD layer are about 71.88 μm and 130.22 μm, respectively. The average grain size of the SPD layer and MPD layer of the Al2024-T351 mid-thick panel is about 51.35 μm and 116.67 μm, respectively. The effect of longitudinal LPF under different prestressed states on grain refinement in the V zone of the Al2024-T351 mid-thick panel was compared and analyzed. The lower surface microstructure of the Al2024-T351 mid-thick panel was formed by two longitudinal LPF, as shown in Fig. 6.112. As can be seen from Fig. 6.112, the SPD layer and MPD layer are generated in the lower surface layer of the Al2024T351 mid-thick panel. As can be seen from Fig. 6.112a, the thickness of the SPD

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6 Strengthening Processes and Effect Evaluations of Airplane Structures …

Fig. 6.112 Lower surface microstructure of Al2024-T351 mid-thick panel formed by two longitudinal LPF. a base material; b free-state LPF; c prestressed LPF

layer and MPD layer on the bottom layer of substrate material caused by rolling is 362 μm and 725 μm, respectively. As can be seen from Fig. 6.112b, the thickness of the SPD layer and MPD layer in the lower surface layer of Al2024-T351 mid-thick panel formed by two longitudinal LPF in a free state is about 720 μm and 750 μm, respectively. As can be seen from Fig. 6.112c, the thickness of the SPD layer and MPD layer of the lower surface layer of Al2024-T351 mid-thickness panel formed by prestressed twin-longitudinal LPF is about 920 μm and 920 μm, respectively. Compared with the substrate material, the SPD layer and MPD layer of the lower surface layer of the Al2024-T351 mid-thick panel are thicker after two longitudinal LPF, which is due to the extrusion plastic deformation of the lower surface layer of Al2024-T351 mid-thick panel induced by LPF. Compared with the free state, the SPD layer and MPD layer of the lower surface layer of the prestressed longitudinal LPF Al2024-T351 mid-thick panels are thicker. The reason is that the elastic prestressed loading improves the extrusion plastic strain of the lower surface layer of the LPF Al2024-T351 mid-thick panel, and the plastic strain gradually decreases with the distance from the lower surface. The grain size increases gradually. Compared with the substrate, the grain sizes of the SPD layer and MPD layer in the bottom layer of the Al2024-T351 mid-thick panel formed by two longitudinal LPF are smaller. Compared with the free state, the SPD layer and MPD layer in the lower surface layer of the prestressed longitudinal LPF Al2024-T351 mid-thick panel have smaller grain sizes. The grain sizes of the SPD layer and MPD layer are about 57.33 μm

6.6 Plastic Forming of Wing Panels with Large-Area Laser Shock Peening

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Fig. 6.113 Distribution of plastic deformation layers of Al2024-T351 mid-thick panel with convex bending induced by two longitudinal LPF processes

and 110.68 μm, respectively. The grain sizes of the SPD layer and MPD layer are about 58.69 μm and 111.12 μm, respectively, in the bottom layer of the Al2024T351 mid-thick panel formed by two longitudinal LPF processes in the free state. The grain sizes of the SPD layer and MPD layer are about 51.08 μm and 102.08 μm, respectively. Based on the microstructure analysis of the entire section, the upper surface, and the lower surface, it can be concluded that the Al2024-T351 mid-thick panel has convex bending deformation after two longitudinal LPF processes, and the plastic deformation layer has been generated in both the upper and lower surface layers of Al2024-T351 mid-thick panel, as shown in Fig. 6.113. Therefore, LPF can induce grain refining layers in the upper and lower layers of Al2024-T351 mid-thick panels, which is helpful to improve the fatigue properties of Al2024-T351 mid-thick panels. (4) Residual stress Surface residual stress distribution of the Al2024-T351 mid-thick panel formed by two longitudinal LPF processes is shown in Fig. 6.114. The sample size is 320 mm × 102 mm × 25 mm (length × width × height), and the impact area curvature radius is 9000 mm. As can be seen from Fig. 6.114, both the upper and lower surfaces of the Al2024-T351 mid-thick panel produced high residual compressive stress during two longitudinal LPF processes. The high-amplitude residual compressive stress on the upper surface of the Al2024-T351 mid-thick panel during LPF can be attributed to the high-density dislocation, grain refinement, and nanocrystals in the upper surface of the mid-thick panel induced by laser shock wave [82]. In addition, the residual compressive stress values in the x- and y-directions at the center of the upper surface are larger than those in the x- and y-directions at both sides of the upper surface, due to the greater plastic strain generated at the center of the upper surface, as shown in Fig. 6.114a. The x-direction compressive residual stress on the upper surface of the Al2024-T351 mid-thick panel is about −266.4 MPa/left side, −335.3 MPa/center, and −279.5 MPa/right side, and the ydirection compressive residual stress on the upper surface is about −238.2 MPa/left side, −319.4 MPa/center, and −308.2 MPa/right, respectively. The residual compressive stress on the lower surface of the Al2024-T351 midthick panel induced by two longitudinal LPF processes is attributed to the negative bending moment M1 induced by large-area LSP, which results in bending deformation of the mid-thick panel and thus produces extrusion plastic deformation on

336

6 Strengthening Processes and Effect Evaluations of Airplane Structures …

Fig. 6.114 Surface residual stress distribution of Al2024-T351 mid-thick panel formed by two longitudinal LPF. a the upper surface; b the lower surface

the lower surface of the mid-thick panel. According to Eq. (6.23), the extruded plastic strain ε on the lower surface of the Al2024-T351 mid-thick panel induced by two longitudinal LPF is 0.139%. In addition, the elastic modulus E of the Al2024T351 panel is 66.23 GPa, so the theoretical residual compressive stress of the lower surface of the mid-thick panel is about −92 MPa. As shown in Fig. 6.114b, the residual compressive stress of the lower surface in the y-direction is about −96.8 MPa/left side, −62.7 MPa/center, and −79 MPa/right side, and the residual compressive stress of the lower surface in the x-direction is about −46.9 MPa/left side, −61.8 MPa/center, and −82.4 MPa/right side, respectively. Therefore, the theoretical values of residual compressive stress on the lower surface of the Al2024T351 mid-thick panel formed by two longitudinal LPF processes are similar to the experimental values. Based on the above experimental results, it can be concluded that longitudinal LPF produces compressive residual stress on the upper and lower surfaces of the Al2024-T351 mid-thick panel. It is well known that fatigue crack initiation occurs on the surface of materials, and residual compressive stress can effectively prevent fatigue crack initiation and propagation [83]. Therefore, the residual compressive stress on the upper and lower surfaces of the Al2024-T351 mid-thick panel can be effectively improved by longitudinal LPF.

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Chapter 7

Quality Control Technology of Structures with Laser Shock Peening

The main purpose of LSP of blades is to improve the fatigue strength, thereby extending the service life. Therefore, the most direct way to detect the quality of LSP is to conduct a high-frequency fatigue experiment (HCF) to obtain the fatigue life of the blade after LSP treatment as the basis for quality judgment. But the experiment was destructive, very expensive, and time-consuming. On the one hand, HCF testing is a random sampling technique and is a crude statistical quality measurement, and online quality testing cannot be achieved. On the other hand, the manufacturing cost of aero-engine blades is very high, some special blades cost hundreds of thousands of yuan, LSP is usually strengthened as the last process, in order to test the LSP quality of the blade and carry out destructive HCF experiments on the blade, which is impossible to promote in actual production. In actual production, it is impossible to carry out destructive fatigue test on the blades treated by LSP, and in order to make the LSP technology apply and promote, it is necessary to have non-destructive testing methods for LSP quality. At present, the relative backwardness of quality inspection technology has seriously hindered the promotion and use of LSP treatment. Therefore, the exploration and research of quality inspection technology for LSP parts is very necessary and meaningful.

7.1 Present Situation of Detection Technology for Laser Shock Peening Quality After the application of LSP technology on aircraft engine blades in 1997, LSP Technology Corporation of the United States has formed a complete set of rapid coating technology, quality control technology, online detection technology, etc. [1] after more than ten years of engineering technology development. After the process is stable and a large number of actual user data are obtained, the quality assurance technology of LSP has become the focus of development in recent years. In order © National Defense Industry Press 2023 S. Zou et al., Laser Shock Peening, https://doi.org/10.1007/978-981-99-1117-2_7

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not to compromise production continuity, quality assurance techniques for real-time non-destructive assessment are the most effective solution. Foreign LSP process has been developed for nearly 30 years, and at the end of the twentieth century, it has begun to be promoted and applied in the aviation industry, and now it has formed industrialization, and its supporting quality inspection technology has also been developed, and applied. The early foreign measurement method is to use laser probes or other measuring devices to detect the volume of surface pits left after LSP, and calculate the residual compressive stress on the metal surface according to the volume of the pit (that is, the amount of plastic deformation on the metal surface), so as to use it as the basis for the quality inspection of the workpiece. There have also been studies on the use of X-rays to directly detect the stress distribution inside the metal to detect the quality of the workpiece [2], which has been experimentally verified by the American Society for Experimental and Materials experiments. Another type of method is to use sound, plasma, workpiece vibration, and other quality characteristic signal detection devices to collect various characteristic signals generated during each laser shock, send the data to the computer, and compare with the relationship function of each feature signal and processing quality obtained in advance (this function is obtained by experiments) as a detection standard [3]. The development of domestic LSP process is rapid, but due to the late start time, the detection technology is relatively backward, in addition to direct or indirect measurement of residual stress and other traditional means, the currently published quality testing method relies on intuitive judgment, through the surface roughness of the workpiece and micro pits for intuitive observation and analysis, to identify the quality of LSP effect, so as to achieve the purpose of non-destructive testing. Detection methods such as intuitive discrimination are only based on experience to make intuitive judgments, which are inefficient and have no accuracy guarantee at all. It can be seen that domestic research in this field is basically blank. The following is a brief introduction to the existing LSP quality testing methods at home and abroad. 1. Discrimination of residual stress measured by X-ray diffraction The surface compressive residual stress of the workpiece after LSP treatment directly affects the fatigue life of the workpiece, and this method uses the mature residual stress measurement technology—X-ray diffraction method to directly measure the residual stress of the workpiece in the treated area, so as to complete the task of quality inspection. X-ray diffraction is based on the principle of X-ray diffraction, known as Bragg’s law. Bragg’s law establishes a definite relationship between the macroscopic and accurately determined diffraction angle and the crystal plane spacing in the material. The elastic strain corresponding to the stress in the material must be characterized by the relative change in the crystal plane spacing. When there is a stress σ in the material, the crystal plane spacing d must change with the relative orientation of the crystal plane and the stress, and according to Bragg’s law, the diffraction angle 2θ will also change accordingly. Therefore, it is possible to obtain the stress σ by measuring the change in diffraction angle 2θ with different crystal plane orientations.

7.1 Present Situation of Detection Technology for Laser Shock Peening Quality

343

2. Discrimination of pit volume After the workpiece is strengthened by LSP, due to the action of the shock wave, the workpiece will produce plastic deformation, forming micro pits on the surface. In general, micro pits are only a few microns to a twentieth micron deep. The depth of the micro pits is the result of plastic deformation, which reflects the magnitude of the residual compressive stress and the surface hardness. The morphology of the micro pits reflects the distribution of compressive residual stress and hardness. Therefore, the quality of LSP can be expressed by the volume of micro pits. GE of the United States uses the volume of the pits generated after impact to judge the quality of LSP. In this method, the volume of the pit can be calculated by scanning the surface of the workpiece before and after impact treatment with a laser interferometer or laser displacement sensor. The system first does some calibration experiments, that is, when the process conditions such as laser impact pulse parameters and workpiece material parameters are determined, the size of the pit volume is measured, and fatigue experiments are performed on these impact treated workpieces, so as to obtain the “pit volume-workpiece life” empirical curve under the conditions of these processing parameters, and obtain the detection standard (according to the required rated workpiece life, find the corresponding pit volume value on the curve, This value is the quality inspection standard for laser impact intensive treatment when the pit volume is used as the detection index under the processing parameter conditions). In the actual workpiece processing, the volume value obtained by the detection can be compared with the test standard value to determine whether the quality is qualified. 3. Intuitive discrimination The Institute of Acoustics of Nanjing University has published a detection method to visually observe and analyze the surface roughness and micro pits of the workpiece to determine the quality of the LSP effect, so as to achieve the purpose of non-destructive testing [4]. According to the results of analyzing a large number of LSP specimens, the surface quality of the LSP area can be divided into four grades: A, B, C, D, and the following are the surface characteristics and performance of each level: Class A: uniform spot, surface roughness of the impact zone is not greater than the surface roughness of the unimpacted zone, forming a very dense bright circle, slightly concave. The fatigue life of workpieces in this grade can be extended by up to eight times. Class B: The light spot is relatively uniform, the surface roughness is slightly increased, the thickness of the vaporization layer is uniform, and there are a very small number of vaporization spots. The fatigue life of workpieces in this grade can generally be increased by 50%. Class C: Discrete point-like, honeycomb-shaped, large plaque-like gasification, uneven gasification zone, and radial grooves centered on the hole outward. Class C no longer reflects the contribution of laser impact to the fatigue life of the material, and the fatigue life gain is almost equal to the loss caused by the decline in surface quality.

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Grade D: strong gasification, melting, coarse grains, formation of gasification pits, rough surface, with radial grooves. Class D can no longer reflect the contribution of laser impact to the fatigue life of the material, and the gain of fatigue life is no longer enough to compensate for the loss caused by the decline of surface quality, and the fatigue life gain of the material is reduced to negative. 4. Discrimination of sound pressure magnitude In the process of laser impact intensification, each laser impact produces elastic and plastic waves due to strong shock waves, and the peak size of sound pressure emitted during LSP is related to the energy of the shock wave caused by the laser pulse, and the energy size of the shock wave directly affects the processing quality of the workpiece and the residual compressive stress on the surface. Therefore, the instantaneous sound pressure emitted during LSP reinforcement can characterize the effect of laser impact reinforcement, and devices such as ultrasonic sound pressure sensors can be used to obtain sound pressure signals. The method should also be calibrated to obtain the empirical curve of “sound pressure peak-workpiece life”, or the relationship between the sound energy parameters during LSP and the residual stress, surface roughness, and other direct quality parameters should be obtained by experiments, and the detection standard should be obtained. In the actual processing, the acoustic signal detection sensing system picks up the acoustic emission signal and processes it, obtains the sound energy parameters caused by the shock wave during the current shock, and compares this value with the detection standard value to make a quality judgment. 5. Discrimination of natural frequency As discussed earlier in this paper, in the process of LSP intensification, each laser shock will deform on the surface of the workpiece, leaving pits and causing compressive residual stress on the surface, which will affect the natural frequency of the workpiece. That is to say, each laser impact will change the natural frequency of the workpiece, and this change is related to the residual compressive stress on the surface of the workpiece. As shown in Fig. 7.1, the figure is the spectrum diagram of the blade vibration response obtained in the US patent, the patent uses the spectrum analysis method to obtain the spectrum curve after the blade is impacted, and identifies its typical peak as the natural frequency of a certain order of the blade, from the figure, it can be seen that the natural frequency of the blade before and after the impact has changed. Therefore, the natural frequency change value of the workpiece during laser impact strengthening treatment can be detected, and the quality judgment of LSP treatment can be judged as a detection index. 6. Similar to methods 2 and 4, this method must also do calibration experiments to obtain the empirical curve of “natural frequency change value—fatigue life” under specific processing parameters, or do experiments to obtain the relationship curve between sound energy parameters and residual stress, surface deformation and other direct quality parameters during laser impact, so that the detection standard can be obtained (according to the required fatigue life, residual stress

7.1 Present Situation of Detection Technology for Laser Shock Peening Quality

345

Fig. 7.1 The natural frequency of certain order changes before and after the blade impact

or surface roughness, find the corresponding natural frequency change value of a certain order on each curve chart, This value is the quality inspection standard for LSP intensive treatment when the natural frequency change value of this order is used as the detection index under the conditions of the processing parameters). During actual processing, the vibration sensing system picks up the vibration signal of the blade during the LSP, analyzes the natural frequency of the blade, and subtracts it from the natural frequency value of the previous impact to obtain the natural frequency change value. By comparing this value with the quality inspection index and the detection standard, the quality inspection of the laser impact reinforced blade can be completed. 7. Discrimination of plasma spectral During LSP, an instantaneous plasma region is formed at the impact point, light emission is formed at the same time as the shock wave, and the spectral characteristics of the plasma optical radiation are related to the impact mass of each impact. In 2001, GE used an optical spectrum analyzer to analyze and measure the instantaneous spectral light intensity of plasma during each impact in the LSP process, and used this parameter as a quality indicator parameter for the LSP process. After establishing the empirical curve of this parameter and the high cycle fatigue life, the characteristic parameters of plasma radiation spectrum obtained by each impact detection and the nominal value on the experience curve can be compared to make a judgment on whether the LSP treatment quality is qualified.

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7.2 Natural Frequency Tests of Aero-Engine Blades with Laser Shock Peening The various detection methods mentioned above have imperfections, especially the current quality testing research on LSP in China is not enough, and the detection method such as intuitive discrimination method is only based on experience to make intuitive judgment, low efficiency, and no accuracy guarantee. In order to solve the contradiction between the rapid development of laser impact strengthening process and the relative backwardness of detection technology, China Academy of Aeronautical Manufacturing Technology and Beihang University conducted joint research by the Aeronautical Science Foundation in 2006, and proposed a new method for LSP quality detection of engine blades based on natural frequency testing [5].

7.2.1 Test System At present, the domestic measurement of the natural frequency of aero-engine blades mainly adopts the resonance method, and the accuracy of this method is often affected by the accuracy of components, and the actual accuracy is low. In addition, the resonance method needs to gradually change the frequency of the excitation force to find out the resonance frequency of the blade, which takes a long time, and it is difficult to achieve online measurement of the natural frequency during the laser impact strengthening of the blade. Therefore, pulse excitation is used to identify the natural frequency of the blade by performing spectral analysis on the blade response signal. And because the LSP treatment will produce a huge instantaneous impact force, which is equivalent to a pulse transient excitation of the blade, so the use of this excitation method to measure can also eliminate the design and consideration of the excitation device. The vibration signal generated by the LSP blade is collected by the eddy current displacement sensor and microphone and then reaches the USB data acquisition card through the transmitter, and the data is transmitted to the computer, and the computer processes the data, and the overall structure is shown in Fig. 7.2. The blade LSP experiment was carried out on the Q-switched neodymium glass laser of China Aviation Manufacturing Technology Research Institute, and the experimental object was 1Cr11Ni2W2MoV stainless steel compressor blade provided by Liyang Aero Engine Company. 1Cr11Ni2W2MoV is a ferritic forming element that reduces the austenitic phase zone by adding a large number of W, Mo, V, and other ferritic elements to low-carbon 12Cr steel. This steel grade has good comprehensive mechanical properties, high room temperature strength, durable strength, and good toughness and oxidation resistance. In the aviation industry, it has been widely used in the manufacture of engine blades, discs, shafts, and other important parts below 600 °C, and its main performance parameters are shown in Table 7.1. In this experiment, it is hoped that by monitoring the natural frequency and sound power of compressor blades in the LSP process, the change law of natural frequency

7.2 Natural Frequency Tests of Aero-Engine Blades with Laser Shock Peening

347

Fig. 7.2 Test system composition

Table 7.1 Blade material properties

Density ρ/(kgm−3 )

7800

Young’s modulus E/(Gpa)

198

Poisson’s ratio ν

0.3

Hugoniot elastic limit HEL/(GPa)

2.62

and sound power in the process of blade LSP is obtained, and then the residual stress of the impact point of the blade after treatment is detected, the relationship between the residual stress of the blade and the natural frequency and sound power of the LSP is analyzed, and finally, the empirical formula is derived as the quality testing standard of blade LSP. The physical and impact point distribution of the detection blade are shown in Figs. 7.3 and 7.4, respectively.

Fig. 7.3 Stainless steel compressor blades

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7 Quality Control Technology of Structures with Laser Shock Peening

Fig. 7.4 Distribution of measurement points

The experiment adopts a single spot impact mode, that is, each spot does not cover each other, and the experimental process is as follows: ➀ wipe the surface of the leaf to be impacted with alcohol, and paste aluminum foil to the part that needs to be impacted; ➁ Install the blades and sensors;➂ Adjust the laser power and spot size and position; ➃ Open the water pump to release water, start the impact treatment on the surface of the specimen, and record the energy meter reading after each impact. First, paste the aluminum foil tape slightly larger than the workpiece on the surface of the specimen, check that there are no bubbles between the aluminum foil and the specimen, put the specimen on the workbench, pre-adjust the position of the table, so that the laser spot diameter is 5 mm (the actual spot shape is an oval spot, the difference between the long axis and the short axis is not large, it is approximately circular, so take the average value of the long axis and the short axis, called the spot diameter), and record the energy meter reading after each impact. When the blade is processed by LSP, there is flowing water covering its surface, and when the impact position is different, the surface water flow state is different, which will affect the vibration of the blade, thereby affecting the measurement accuracy. Through the experiment on the blade percussion response under different flowing water conditions, it is found that the random error of the second-order natural frequency recognition of the blade under the influence of flowing water is within 0.04 Hz, and the random error of the system’s identification of the first-order natural frequency of the blade is consistent with the tuning fork calibration result, which is also 0.02 Hz.

7.2 Natural Frequency Tests of Aero-Engine Blades with Laser Shock Peening

349

Table 7.2 Natural frequency measurement results Ordinal

First-order natural frequency/(Hz)

Second-order natural frequency/ (Hz)

Residual stress in the length direction/(MPa)

Residual stress in width direction/(MPa)

1

218.76

1093.25

2

218.93

1093.33

−118

−280

3 4

219.00

1093.50

−208

−331

219.02

1093.78

−311

−418

5

218.96

1093.89

−185

−392

6

219.10

1093.97

−314

−370

7

219.11

1094.03

+22

−146

8

219.18

1094.24

7.2.2 The Changes of Natural Frequency and Residual Stresses of Blades The residual stress of the single spot impact zone is carried out in Beijing Mechanical and Electrical Research Institute, and the testing equipment is the MSF-2 M X-ray stress tester produced by Rigaku Electric in Japan, the X-ray spot of the equipment is 4 mm × 4 mm, which is the same size as the impact zone, and the measurement result is the average of the entire area. Due to the oval shape of the impact laser spot, the residual stress generated after laser impact of the blade is inconsistent in the direction of the major and minor axes. The first group detected the residual stress in the long and wide directions of the blade at impact points No. 2–7, and learned that the residual stress value in the blade length direction was large. The second group detects residual stress in the width direction of all impact points. In the first set of experiments, the displacement response signal after blade impact was analyzed to obtain the natural frequency, several typical points were selected to measure its residual stress, and finally, the experimental results were obtained in Table 7.2.

7.2.3 Relation Between Impact Times and Natural Frequency Observing the above experimental data, the relationship curves between the firstand second-order natural frequencies and the number of laser impacts of blades are listed, as shown in Figs. 7.5 and 7.6. It can be seen from Figs. 7.5 and 7.6 that the natural frequencies of the first and second orders of the blade increase with the increase of the number of blade impacts, and the second-order frequencies are more sensitive to the impact and have a better correlation. Using the linear regression analysis method, the univariate linear

350

7 Quality Control Technology of Structures with Laser Shock Peening

Fig. 7.5 Frequency relationship curve of the number of shocks and the first order

Fig. 7.6 Frequency relationship curve of the number of shocks and the second order

regression equation between the natural frequency of the blade and the number of shocks is derived as follows: First-order natural frequency experience curve: f 1 = 0.0486 n + 218.79

(7.1)

correlation coefficient r1 = 0.9142. Second-order natural frequency empirical curve: f 2 = 0.1423 n + 1093.1

(7.2)

7.3 Laser Shock Peening Effect Characterized by Acoustic Signal …

351

correlation coefficient r2 = 0.9849. The correlation coefficients of Eqs. (7.1) and (7.2) are relatively close to 1, which means that the natural frequency of the blade, especially the second-order natural frequency, has a significant linear relationship with the number of laser impacts. Equations (7.1) and (7.2) are shown in Figs. 7.5 and 7.6 as dashed lines as empirical curves of natural frequencies as a function of the number of shocks, and it can be seen that the measured values of second-order natural frequencies are in good agreement with this empirical curve. Therefore, Eq. (7.2) can be well used to predict and control the blade LSP process, that is, if the measured second-order natural frequency value of the blade deviates too much from the experience curve during a certain impact in the LSP process, it can be judged that the processing quality of the point is unqualified.

7.3 Laser Shock Peening Effect Characterized by Acoustic Signal and Plasma Plume The online detection of LSP effect is very important for the engineering application of LSP, and the main piezoelectric thin film method and X-ray diffraction method used by domestic scientific research institutions are currently the piezoelectric thin film method. Piezoelectric film method is to place PVDF piezoelectric film on the back of the workpiece, and the shock wave force signal in the workpiece into an electrical signal, through the analysis of the electrical signal to reflect the quality of this laser impact strengthening, its biggest disadvantage is that the piezoelectric film method equipment has a short service life, poor economy in practical production applications, in addition, under the same processing parameters, the measurement results are affected by the material and thickness of the workpiece. X-ray diffraction is an offline measurement method, the residual stress measured is itself incorrect, it is a destructive method, and the measured material is limited. In the process of LSP, there will be a sonic signal and plasma plume formed by plasma explosion between each pulsed laser beam and metal workpiece, and the LSP effect is judged by detecting the sonic signal and plasma feather image. China Academy of Aerospace Manufacturing Technology proposed “an online detection method and device for LSP effect” (patent publication number: CN103207178A) [6] using sound pressure sensor (8) and acoustic coupler (9) to collect the sound wave signal of each pulse laser beam (3) and convert it into an electrical signal, and the image acquisition platform (14) probe I (12) and probe head II (13) to collect plasma plume images formed on the surface of the absorbing layer (5) during laser impact intensification. As shown in Fig. 7.7, the shock wave acoustic signal induced by the intense pulsed laser beam and the plasma plume image information are measured in real-time and compared with the standard signal to jointly identify the LSP effect. The advantage of this technology is that it can determine the quality of LSP reinforced metal workpieces in real-time, quickly and accurately.

352

7 Quality Control Technology of Structures with Laser Shock Peening

Fig. 7.7 On-line detection device for laser impact enhancement effect. 3—laser beam; 5—absorbing layer; 8—sound pressure sensor; 9—acoustic coupler; 12—probe head I; 13—probe head II; 14—Image acquisition platform

References 1. Zou SK (2005) The latest development of laser shock processing (in Chinese). New Technol New Process 4:44–46 2. Wang J, Zou SK, Tan YS (2005) Application of laser shock processing on turbine engines (in Chinese). Appl Laser 25(01):32–34 3. Davis BM, Mcclain RD, Suh UW, et al (2003) Real time laser shock peening quality assurance by natural frequency analysis. USA, US006629464B2 4. Zhang YK (1997) Research on intuitive judgment and control method of laser shock peening effect (in Chinese). Chin J Lasers 24(5):467–471 5. Zou SK, Cao ZW, Yang HL (2010) Natural frequency test of turbine blades in laser shock processing (in Chinese). China Mech Eng 6:648–651 6. Gong SL, Zhang YK, Lu JZ, et al (2013) An online detection method and device for laser shock peening effect (in Chinese). CN, CN103207178A

Chapter 8

Strengthening Processes and Effect Evaluation of Welded Structures with Laser Shock Peening

8.1 Present Situation of Weldments with Laser Shock Peening at Home and Abroad Early research on LSP was mainly carried out on aluminum alloy welded joints, but there has been no substantial application. With the application research of friction stir welding, the United States has carried out LSP process research for aluminum alloy friction stir welding joint, mainly to reduce the residual tensile stress on the surface of the welded structure and solve the problem of tissue softening, so as to improve its fatigue and corrosion resistance. With the extensive development of the nuclear industry, in view of the maintenance of atomic reactor pressure vessels and pipeline welds, Toshiba Corporation of Japan has carried out laser impact strengthening process research and developed a unique set of laser impact strengthening equipment and processes [1], as shown in Fig. 8.1. Due to space limitations in nuclear reactors, the absorption layer cannot be arranged according to the conventional impact treatment method, so the company adopts laser impact enhancement technology without absorption layer and uses optical fiber conduction laser to directly impact water as a confinement medium. Before 2011, Toshiba Corporation of Japan needed to laser impact strengthening 6 ~ 10 nuclear reactors every year, and after the earthquake in Japan, the Fukushima Daiichi Nuclear Power Plant was seriously damaged, and a large amount of radioactive materials leaked to the outside. In 2013, the world’s largest nuclear power plant was permanently abandoned, and Japan’s Toshiba lost its largest domestic market, but Europe, the United States, and South Korea began to pay attention to the application prospects of LSP in the nuclear industry. In China, the units that carry out the application research of LSP technology mainly include China Aviation Manufacturing Technology Research Institute, Air Force Engineering University, Jiangsu University, etc. The China Academy of Aeronautical Manufacturing Technology and Nanjing University of Technology cooperated to carry out the research on stress corrosion cracks of LSP welds [2]. The © National Defense Industry Press 2023 S. Zou et al., Laser Shock Peening, https://doi.org/10.1007/978-981-99-1117-2_8

353

354

8 Strengthening Processes and Effect Evaluation of Welded Structures …

Fig. 8.1 Laser impact application in nuclear industry at Toshiba Energy Center, Japan

experimental material used is a 76 mm (length) × 72 mm (width) × 5 mm (thickness) 1Cr18Ni9Ti block, and TIG welding is performed in the middle of the test block along the length direction (as shown in Fig. 8.2). With black paint as the absorption layer and flowing water as the constraint layer, the LSP of the entire weld and heat-affected zone is carried out by spot lapping. The results show that the weld area manifests as compressive stress after laser impact strengthening. In order to reveal the mechanism, the China Academy of Aeronautical Manufacturing Technology analyzed the microstructure of the weld after laser impact strengthening. The results show that laser impact strengthening increases the dislocation density of the weld and the heat-affected zone, thereby improving the residual stress in the weld area. For the LSP of steel structure materials in the nuclear industry, Japan adopts the LSP method without absorbing layer, LSP relies on the weld structure itself to absorb laser energy to produce high temperature and high-pressure plasma in water, and generate laser shock wave, LSP after the weld has an obvious ablative loose structure. In China, the LSP method with absorbing layer is adopted, and the absorption layer is strengthened by LSP to avoid ablation and loosening of the weld surface. Fig. 8.2 Welded stainless steel plate

8.2 Tensile Strength and Fatigue Lives of Argon Arc Welded GH30 Alloy

355

At present, the international LSP technology has been successfully used in aviation, nuclear power, and other fields, and the domestic China Aviation Manufacturing Technology Research Institute and Suzhou Thermal Engineering Research Institute have cooperated in the research on stress corrosion crack resistance of LSP-reinforced stainless steel structure. Unlike the early nuclear power plant of Toshiba Corporation of Japan that introduced laser impact strengthening in the maintenance process, China’s nuclear power plant construction is in the early stage, but also the peak period of construction, it is very necessary to consider the application of LSP at the beginning of development, and the difficulty of LSP implementation of open welded structures will be greatly reduced. In the process of stress corrosion resistance strengthening of welded structures, China Aviation Manufacturing Technology Research Institute can adopt large spot LSP, LSP with absorbing layer, small spot non-absorbent coating impact treatment and combined with other strengthening methods according to different structural characteristics, which not only improves surface strengthening efficiency, strengthening quality and reduces operating costs, but also reserves relevant technologies for later maintenance. It is believed that through domestic development, supplemented by appropriate international cooperation, China will make breakthroughs in LSP of the welded structure of nuclear reactors to improve stress corrosion resistance, and make great contributions to the development of clean energy in China. In addition, with the development of aviation manufacturing technology, welding technology is increasingly used in China’s new aircraft and engine structures. Therefore, in the development of aviation welded structures, LSP not only has a wide range of application prospects, but also can improve the performance of joints.

8.2 Tensile Strength and Fatigue Lives of Argon Arc Welded GH30 Alloy In many welded structures, the mechanical properties of welds and heat-affected zones are generally worse than those of base metals due to recrystallization or tempering softening, which are weak links and are often strengthened by postweld heat treatment or strain hardening. The method of post-weld heat treatment is often impossible in the whole welding structure. A good alternative to improve the strength of the heat-affected zone of the weld is strain hardening, such as rolling or explosive impact on the weld, but these processes are not practical or ideal in some complex structures. Ultrasonic impact treatment and laser shock strengthening are promising post-weld treatment processes [3, 4], which can greatly improve the strength and fatigue properties of the weld structure through strain hardening. In this section, the exploratory experiment of laser shock peening was carried out on the weld microstructure of GH30 plate with a thickness of 1.66 mm. The microhardness, residual stress distribution, tensile strength, fatigue life, and fatigue fracture analysis of laser shock peening weld were studied [5, 6].

356

8 Strengthening Processes and Effect Evaluation of Welded Structures …

8.2.1 Micro-hardness and Residual Stress 1. Experimental method The butt joint of the welded specimen is welded by filling wire argon arc welding. The wire is intercepted on the base metal. After the material is welded into a rectangular plate of 200 mm × 300 mm, the appearance of the weld with a length of 300 mm is inspected. No welding cracks and obvious welding deformation are required, and then X-ray non-destructive testing is performed. The position where X-ray does not detect welding cracks and pores is selected. According to the arrangement shown in Fig. 8.3, the tensile-tensile fatigue specimen and static tensile specimen are intercepted on the weld. The dimensions of the fatigue specimen and the tensile specimen are shown in Figs. 8.4 and 8.5, respectively. In order to ensure a good contrast between laser shock peening and non-laser shock peening specimens, the specimens to be impacted and the specimens without Fig. 8.3 Interception method of group specimens

Fig. 8.4 Static tensile specimen size (mm)

Fig. 8.5 Fatigue Specimen Dimensions

8.2 Tensile Strength and Fatigue Lives of Argon Arc Welded GH30 Alloy

357

impact are intercepted in turn on the same weld. In order to reduce the influence of the weld surface profile (stress concentration at the convex or concave parts) on the fatigue experiment and reduce the dispersion of the experimental data, the convex and concave parts of the weld are polished after the specimen is cut, and the laser shock strengthening is performed on the middle weld of the specimen to be shocked. The dye Q-switched neodymium glass laser equipment is used, the pulse energy is 9 ~ 14 J, the pulse width is 20 ns, and the circular spot of Φ6 mm is impacted on one side, and superimposed twice to cover the 10 mm long weld. After laser shock peening, static tensile tests were performed to determine the effect of laser shock peening on the tensile strength of the weld. Finally, the maximum stress level of the fatigue test is determined according to the 0.35 ~ 0.65 times the tensile strength of the unimpacted specimen. The stress ratio R = 0.1 is taken to compare the fatigue test of the fatigue specimen (all the loading loads in the fatigue test are calculated by the measured values of the specimen size). In the experiment, we also tested the microhardness and surface residual stress of the weld surface. The surface residual stress was measured by X-ray diffraction. The radiation area was 4 mm × 4 mm. In order to eliminate the influence of polishing on the surface stress, the 0.1 mm surface layer was corroded by chemical corrosion before the test. 2. Microhardness Figure 8.6 shows the GH30 argon arc welding joint profile (the lower surface is shown as the back of the weld), and the weld width is 7 mm. GH30 is single-phase austenite structure before welding, and an obvious dendritic structure is formed in the weld after argon arc welding. The dendritic matrix is austenite and the axial crystal is carbide. Figure 8.7 shows the microhardness distribution of GH30 welded joints with/without laser shock peening. The hardness is measured from the center of the weld to the outside (Y direction shown in Fig. 8.8). It can be seen from Fig. 8.7 that the hardness levels of the front and back of the GH30 weld are comparable. Due to the influence of dendrites and carbides in the weld, the hardness of the measured points fluctuates greatly. The average microhardness is 170 HV, and the average microhardness of the heat-affected zone is about 230 HV. After laser shock peening, the microhardness of the welding zone is obviously improved, and the maximum value reaches 330 HV in the center of the laser shock zone (Fig. 8.7). The average microhardness of the whole impact zone is 280 HV, which is 65% higher than that before impact. According to the previous distribution law of microhardness of laser shock strengthened GH30 with layer depth under the same parameters [7], the microhardness of laser shock strengthened GH30 decreases with the increase of layer depth, and the total hardened layer depth is about 0.5 mm. Therefore, after laser shock strengthening GH30 welded joint, the microhardness and tensile strength of the welded zone are obviously improved due to the strain hardening caused by shock wave. 3. Residual stress The GH30 weld specimens with/without laser shock peening were cut into small pieces to release the macro stress of the welding structure. The X-ray residual stress

358

8 Strengthening Processes and Effect Evaluation of Welded Structures …

Fig. 8.6 GH30 welded joint section

Fig. 8.7 GH30 weld surface microhardness distribution

Fig. 8.8 GH30 weld surface metallographic diagram

test was carried out on the weld surface after removing the surface oxide layer or polishing layer by acid corrosion. The X-ray irradiation area was 4 mm × 4 mm. In the experiment, the X-ray diffraction seam reflects that there is an obvious dendrite structure in the GH30 weld, which is consistent with the metallographic results of the weld. The results of surface residual stress measured by austenite diffraction peaks are shown in Table 8.1. It can be seen from Table 8.1 that after laser shock peening, a higher residual compressive stress state can be obtained on the surface of the weld zone, especially for σy, which has a significant effect on the tensile fatigue properties perpendicular to the weld direction. In addition, according to the laser shock strengthening surface residual stress area is slightly larger than the spot ( two ϕ6mm lap) results [7], it can be seen that GH30 weld and heat affected zone in the

8.2 Tensile Strength and Fatigue Lives of Argon Arc Welded GH30 Alloy Table 8.1 Comparison of surface residual stresses of impact and non-impact welds

State

359

GH30

Unstrengthened specimen Strengthened sample

σx /(MPa)

σy /(MPa)

−93.1

155.8

−137.2

−109.8

laser shock strengthening area, so the fatigue life of GH30 welded specimens after impact did not significantly improve, the following further fatigue fracture analysis.

8.2.2 Tensile Strength and Fatigue Lives After the welding of GH30 sheet, there are obvious dendrites in the welding structure. The average microhardness of the weld seam decreases from HV210 of the base metal to HV170. The surface residual stress is tensile stress in the grinding part. After laser shock peening, σx changes from 155.8 MPa to −109.8 MPa; the average microhardness of the impact zone is HV280, and the maximum value reaches HV330, which is higher than that of the base metal before welding. The static tensile test of GH30 welded specimens was carried out. The tensile strength results of the specimens strengthened by laser shock peening and the specimens without laser shock peening were shown in Table 8.2. It can be seen from Table 8.2 that the tensile strength of the specimen after laser shock peening is significantly higher than that of the specimen without laser shock peening, indicating that laser shock peening can increase the tensile strength of the welded joint by 12%. After the static tensile test, we then carried out the tensile-tensile fatigue test of the welded specimen at room temperature. The high-frequency fatigue test machine was used to carry out the fatigue test with a stress level of 0.65 times the tensile strength of the welded specimen without laser shock peening and a frequency of 94 ~ 95.5 Hz. The experimental results are shown in Table 8.3. As can be seen from Table 8.3, due to the good welding quality of GH30, the performance of the welded specimens is relatively close, and the fatigue life dispersion of all the specimens is not large. The logarithmic coefficient of variation of the fatigue life of the specimens strengthened by laser shock is 0.024. The experimental results show that the sample size meets the requirements at 90% set degree, but there Table 8.2 Tensile strength of GH30 welded specimens Sample number

112➀

109➀

110

111

Tensile strength/(MPa)

637.9

635.9

588.2

546.2

Average tensile strength/(MPa)

636.9

➀Impacted specimens, same below

567.2

360

8 Strengthening Processes and Effect Evaluation of Welded Structures …

Table 8.3 Fatigue life of GH30 welded specimens Sample number

Loading stress σ /MPa

Fatigue life N/cycle

Logarithm

Average fatigue life

102

368.7

202,000

5.305

263,633

103

368.7

453,000

5.656

106

368.7

200,000

5.301

104➀

368.7

273,200

5.436

105➀

368.7

504,100

5.703

101➀

368.7

520,900

5.717

415,911

is a cross distribution of fatigue life between the specimens strengthened by laser shock and the specimens without shock treatment. In general, the median fatigue life of the specimens after laser shock processing is 58% higher than that of the unshocked specimens. Due to the lack of experimental data, the strengthening effect needs to be further verified.

8.2.3 Fatigue Fracture Analysis Figures 8.9 show the fatigue fracture macrograph (Fig. 8.9a) and fatigue striation (Fig. 8.9b, c) of Specimen 102 (fatigue life N = 202,000) without laser shock processing. It can be seen from the Figure that the fatigue crack originates from the weld toe in the middle of the specimen and extends smoothly to the entire weld in a fan-shaped radiation direction. Since the fatigue crack propagation is not blocked, the fatigue life is very short. Figure 8.9b is a strip of 500 μm above the fatigue source center, and the strip appears in blocks with a width of about 0.8 μm. Figure 8.9c shows the strips and indentations at 500 μm above the center of the fatigue source. It can be seen from the Figure that the dense strips and indentations in the vertical direction are about 0.4 μm in width, and the indentation width is not uniform.

Fig. 8.9 Fatigue fracture and fatigue striation of specimen 102 a Macro fatigue fracture; b fatigue striation; c fatigue striation

8.3 Tensile Strength and Fatigue Lives of Plasma-Welded 1Cr18Ni9Ti Alloy

361

Fig. 8.10 Fatigue fracture and fatigue bands of specimens strengthened by laser shock peening a Macroscopic fatigue fracture; b Fatigue striation

Figure 8.10a is the macro appearance of fatigue fracture of 104 specimens (fatigue life N = 273,000) strengthened by laser shock peening. It can be seen from the diagram that the fatigue zone of No.104 specimen is located at the lower left end of the diagram, and there is a block smooth fatigue propagation zone, but there is no obvious main fatigue source, and the radiation direction of crack propagation is not obvious. Figure 8.10b is a fatigue strip in the smooth fatigue propagation zone, with a width of about 1.2 μm and a very uniform distribution. It can be seen that the crack propagation in this area is very stable. There is no obvious main fatigue source in specimen No.104, which can be attributed to the influence of laser shock peening. Because multiple initial crack sources encounter strain hardening structure after extending a path, the crack is blocked and stops expanding. However, due to the rapid formation of penetrating cracks in the thickness direction of one end, the stress intensity factor ΔK increases sharply, which leads to the rapid expansion of fatigue cracks along the width direction to about 4.5 mm, and the effect of laser shock hardening cannot be effectively exerted. Therefore, the fatigue life of the specimen is significantly lower than that of other specimens strengthened by laser shock.

8.3 Tensile Strength and Fatigue Lives of Plasma-Welded 1Cr18Ni9Ti Alloy In this section, the exploratory experiment of laser shock peening was carried out on the weld microstructure of 1.2 mm thick 1Cr18Ni9Ti plate. The microhardness, residual stress distribution, tensile strength, fatigue life, and fatigue fracture analysis of laser shock peening weld were studied [5, 6].

362

8 Strengthening Processes and Effect Evaluation of Welded Structures …

8.3.1 Micro-hardness and Residual Stress 1. Experimental methods The butt joint 1Cr18Ni9Ti of the welded specimen is welded by plasma welding. The material is welded into a rectangular plate of 200 mm × 300 mm. After the appearance inspection of the 300 mm long weld, it is required that there is no welding crack and obvious welding deformation, and then X-ray non-destructive testing is performed. The position where X-ray does not detect welding cracks and pores is selected. According to the arrangement shown in Fig. 8.3, the tensile-tensile fatigue specimen and static tensile specimen are intercepted on the weld. The dimensions of the fatigue specimen and the tensile specimen are shown in Figs. 8.4 and 8.5, respectively. In order to ensure a good contrast between laser shock peening and non-laser shock peening specimens, the specimens to be impacted and the specimens without impact are intercepted in turn on the same weld. In order to reduce the influence of the weld surface profile (stress concentration at the convex or concave parts) on the fatigue experiment and reduce the dispersion of the experimental data, the convex and concave parts of the weld were polished after the specimen was cut, and the laser shock strengthening was performed on the middle weld of the specimen to be shocked. The dye Q-switched neodymium glass laser equipment was used, the pulse energy was 9 ~ 14 J, the pulse width was 20 ns, and the circular spot of ϕ6mm was impacted on one side, and superimposed twice to cover the 10 mm long weld. After laser shock peening, static tensile tests were performed to determine the effect of laser shock peening on the tensile strength of the weld. Finally, the maximum stress level of the fatigue test is determined according to the 0.35 ~ 0.65 times the tensile strength of the unimpacted specimen. Taking the stress ratio R = 0.1, doing comparative fatigue tests on fatigue specimens (all the loading loads in the fatigue test are based on the specimen size). In the experiment, we also tested the microhardness and surface residual stress of the weld surface. The surface residual stress was measured by X-ray diffraction. The radiation area was 4 mm × 4 mm. In order to eliminate the influence of polishing on the surface stress, the 0.1 mm surface layer was corroded by chemical corrosion before the test. 2. Microhardness Figure 8.11 is 1Cr18Ni9Ti plasma welding joint profile (the lower surface is the back of the weld) with weld width of about 5 mm. Fig. 8.11 1Cr18Ni9Ti welded joint section

8.3 Tensile Strength and Fatigue Lives of Plasma-Welded 1Cr18Ni9Ti Alloy Table 8.4 Comparison of surface residual stresses of impact and non-impact welds

State

363

1Cr18Ni9Ti σ x /(MPa)

σ y /(MPa)

Unstrengthened specimen

134.3

−109.8

Strengthened sample

–

−204.8

After welding, the microhardness of the middle part of the weld decreased from HV195 of the base metal to HV170, and the hardness level after laser shock peening increased slightly to the base metal level. This is because 1Cr18Ni9Ti produces phase transformation martensite during plasma welding, which weakens the effect of deformation martensite produced by laser shock peening on improving microhardness, so it has little effect on tensile strength. 3. Residual stress The 1Cr18Ni9Ti weld samples with/without laser shock peening were cut into small pieces to release the macro stress of the welding structure. The surface oxide layer or polishing layer on the weld surface was removed by acid corrosion and X-ray residual stress test was carried out. The X-ray irradiation area was 4 mm × 4 mm. There are obvious martensite diffraction peaks in the 1Cr18Ni9Ti weld, but it is still dominated by austenite diffraction peaks. The results of surface residual stress measured by austenite diffraction peaks are shown in Table 8.4. From Table 8.4, it can be seen that after laser shock peening, the surface of the welded zone can obtain a higher residual compressive stress state, especially for σy, which has a significant change, and σy has a great influence on the tensile fatigue performance perpendicular to the weld direction. In addition, according to the laser shock strengthening surface residual stress area is slightly larger than the spot ( two ϕ6mm lap) results [7], it can be seen that the laser shock strengthening area can fully cover the 1Cr18Ni9Ti weld and heat-affected zone, the following further fatigue fracture analysis.

8.3.2 Tensile Strength and Fatigue Lives After welding of 1Cr18Ni9Ti plate, the average microhardness of the weld joint decreased from HV195 of the base metal to HV170. X-ray diffraction showed that there were obvious martensite diffraction peaks in the welded structure, but the main structure was still austenite. The surface residual stress was tensile stress (austenite phase) at the polished part, and σ y was converted to −204.8 MPa after laser shock processing. 1Cr18Ni9Ti welding specimen static tensile test of laser shock peening specimen and without laser shock peening specimen tensile strength results are shown in Table 8.5. From Table 8.5, it can be seen that the tensile strength of 141 specimens after laser shock peening is close to that of 122 specimens without laser shock peening, but the average level of welded specimens after laser shock peening is still slightly

364

8 Strengthening Processes and Effect Evaluation of Welded Structures …

Table 8.5 Tensile strength of 1Cr18Ni9Ti welded specimens Sample number

122

123

141➀

142➀

Tensile strength/(MPa)

637.1

592.3

639.7

648.0

Average tensile strength/(MPa)

614.7

643.9

higher than that of specimens without laser shock peening. It shows that laser shock peening can improve the strength of welded joints, and the amplitude of strength increase is only 5%. This is because the phase transformation martensite is produced during the welding process of stainless steel, and the strength is close to the level of base metal, which reduces the influence of deformation Martensitic on the strength of laser shock peening. After the static tensile test, the fatigue test is carried out. The experimental results are shown in Table 8.6. It can be seen that the fatigue life value of the 201–204 specimens without laser shock peening is less dispersed, and it can meet the requirements of the comparison sample at a confidence level of 90%. At this time, the fatigue life of the laser shock peening specimen under the condition of non-fracture is still significantly higher than that of the comparison sample. The results show that laser shock peening can greatly improve the fatigue life of 1Cr18Ni9Ti, and the increased value is more than 300%. This is because laser shock peening can eliminate the residual tensile stress on the weld surface and obtain a higher residual compressive stress state, which is very beneficial to fatigue performance. Table 8.6 Fatigue life of 1Cr18Ni9Ti welded specimens Sample number

Loading stress/(MPa)

Frequency/(Hz)

Fatigue life N/cycle

119

450.7

80.0

56,500

201

321.9

93.0

418,700

202

321.9

90.0

256,700

203

321.9

84.0

111,600

204

321.9

90.1

948,500

120➀

322.0

88.1

2,319,600+

386.3

90.0

3,974,300+

450.7

90.0

1,112,800+

515.1

88.0

86,500

121➀

321.9

89.0

1,024,000+

117➀

321.9

87.6

2,986,400+

+ Refers to not fatigue fracture, continue to load

Average fatigue life 326.6

8.3 Tensile Strength and Fatigue Lives of Plasma-Welded 1Cr18Ni9Ti Alloy

365

Fig. 8.12 Fatigue fracture and fatigue striation of No. 106 specimen a Macroscopic fatigue fracture; b Fatigue striation

8.3.3 Fatigue Fracture Analysis Figure 8.12 show the fatigue fracture macrographs (Fig. 8.12a) and fatigue striations (Fig. 8.12b) of Specimen 106 ( fatigue life N = 200,000) without laser shock processing. It can be seen from Fig. 8.12 that a fatigue crack originated from the lower corner of the specimen, and the radiation direction was not obvious, but it caused the entire section edge line in the thickness direction to expand to the other side. Although there was a pore with a diameter less than ϕ0.1 mm on the fracture, it was in the later stage of crack propagation and had no effect on the path. Due to the intersection of multiple fatigue cracks, there are ridges and dimples on the fracture surface, and the width and direction of the bands around the ridges in different directions (Fig. 8.12b) are also different. Figure 8.13 is the macro appearance of the fatigue fracture of 105 specimens ( fatigue life N = 273,000) strengthened by laser shock peening. The main crack source is angular crack, but the propagation path is different from that of 104 specimen, which is blocked at the initial stage of propagation. The fatigue strips in the fatigue propagation zone have a width of 0.6 ~ 0.7 μm, and are distributed in a large area in the smooth fatigue zone, but the directions of the fatigue strips are different, and there are valley peaks and valley bottoms, indicating that laser shock peening produces a large number of dislocations in the weld structure [7], forming a network hardening structure to hinder the crack propagation, thus prolonging the fatigue life of the specimen.

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Fig. 8.13 Fatigue fracture and fatigue striation of specimens strengthened by laser shock peening

8.4 Mechanical Property and Fatigue Lives of TIG Welded TC4 Titanium Alloy with Multiple Impacts 8.4.1 Micro-hardness and Microstructure TC4 (Ti-6Al-4 V) titanium alloy is a kind of metal material which is widely used in aerospace and other industrial fields. Domestic titanium alloy parts are usually welded by conventional DC TIG welding process. However, TC4 titanium alloy welded by conventional DC TIG welding process is prone to defects such as coarse structure, which reduces the quality of the joint. In order to improve the quality of welded joints, TIG welded joints of titanium alloy are usually treated by post-weld heat treatment. However, it is difficult to carry out overall post-weld heat treatment for some large-scale titanium alloy welded components, and improper selection of post-weld heat treatment process will coarsen the joint structure and weaken the performance of the base metal [8, 9], making it difficult to greatly improve the joint quality. Laser shock peening can induce grain refinement and high residual compressive stress on the surface of metal materials, which can improve the mechanical properties of materials. Therefore, the research work of laser shock peening TC4 titanium alloy tungsten inert gas (TIG) welded joints is carried out many times. The optimized LSP treatment process is used to weld the TIG welded joints of TC4 titanium alloy, which is widely used in the aerospace industry and other industrial fields. The post-weld heat treatment process can not only avoid the problem of coarse joint structure and weakening of base metal performance caused by improper selection of post-weld heat treatment process, but also can be used to weld large welded components without space constraints. In this section, LSP technology was used to impact and strengthen the TIG welding seam of TC4 titanium alloy. The surface hardness, microstructure, tensile mechanical properties, fatigue life, and fatigue fracture of the weld were tested, and the test results

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367

Table 8.7 TC4 titanium alloy chemical composition ( %) Al

V

Fe

C

N

H

O

Ti

5.5 ~ 6.8

3.5 ~ 4.5

≤0.3

≤0.1

≤ 0.05

≤ 0.015

≤ 0.2

Balance

Table 8.8 TC4 titanium alloy TIG welding process parameters Current/(A)

165

Voltage/(V)

11

Welding speed/(mm/min)

Length of arc/(mm)

Argon flowrate/(L/min) Back

Protective cover

Nozzle

200

3

1~2

15

13

were compared with the weld without laser shock strengthening. The effect of LSP process on the mechanical properties and microstructure of TIG welding seam of TC4 titanium alloy was studied. This will promote the application of titanium alloy welded parts in aerospace and other industrial fields. 1. Experimental materials and process scheme TC4 is a medium strength α + β two-phase titanium alloy, the chemical composition of the mass fraction is shown in Table 8.7. The specification of TC4 plate welding workpiece is 200 mm × 150 mm × 2.5 mm. The conventional DC TIG welding process is used for wire-filled butt welding. The welding wire is TA1 with a diameter of 1.0 mm. The welding process parameters are shown in Table 8.8. For the TC4 titanium alloy after welding, X-ray flaw detection was performed on the weld according to the requirements of HB5484-91, and LSP treatment was performed on the welded joint whose joint quality reached grade I standard. The process parameters of laser shock strengthening weld are as follows: aluminum foil absorption layer, uniform water confinement layer, energy 40 J, pulse width 30 ns, square spot size 3.9 mm × 3.9 mm, spot overlap form is shown in Fig. 8.14, square spot overlap form is used for single laser shock strengthening, circular spot diameter is 4.5 mm, spot overlap form is shown in Fig. 8.15. Circular spot is used for multiple laser shock strengthening, perpendicular to the welding direction of the weld, and shocked five times in turn, so that the range of LSP treatment covers the weld, heat affected zone (HAZ), and part of the base metal. In order to avoid the aluminum foil absorption layer rupture and loss of protection, for the LSP treatment more than 2 times the process, the use of multiple replacement of aluminum foil for processing, and weld the front and back of the LSP treatment the same way. After LSP treatment, some samples were selected to test their surface hardness. Tensile specimens and fatigue life test specimens were prepared according to GB/T228-2002 and GB3075-82 respectively. The appearance and dimensions of the specimen are shown in Fig. 8.16. The weld is located in the middle of the specimen. The tensile mechanical properties and fatigue life of the joints were tested by INSTRON8801-50KN hydraulic servo testing machine. At the same time, welded joints without LSP treatment are tested.

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Fig. 8.14 Square spot overlap form

Fig. 8.15 Circular spot overlap form

Fig. 8.16 Specimen a Tensile specimen; b Fatigue life test specimen

2. Surface micro-hardness The Vickers hardness test of the weld needs to ensure that the surface flatness of the sample is good. The grinding metallographic process will break the surface

8.4 Mechanical Property and Fatigue Lives of TIG Welded TC4 Titanium …

369

strengthening layer of the weld treated by LSP, which obviously cannot be selected. There is almost no reinforcement on the back of the weld, and the finish and flatness are good. Therefore, the back of the weld is selected to test its Vickers hardness. The equipment for testing the Vickers hardness of the weld surface is a DHV-1000 micro-hardness tester with a dwell time of 15 s after loading. On the lower surface of the specimen, perpendicular to the welding speed direction, from the weld centerline to the base metal, a test point is selected every 0.5 mm, as shown in Fig. 8.17. The Vickers hardness of the square spot single laser shock peening weld and the weld without LSP treatment was tested with a test load of 4.9N. The experimental data obtained were plotted as shown in Fig. 8.18. From Fig. 8.18, it can be seen that in the weld zone, the surface hardness of the TIG weld zone without LSP treatment is higher than that of the heat-affected zone, and the maximum hardness difference is 710 MPa, which is related to the acicular martensite α ' phase generated during the solidification and crystallization process of the TIG weld pool of TC4 titanium alloy. After LSP treatment, the surface micro-hardness of the weld zone and the heat-affected zone has little difference, the maximum hardness difference is 200 MPa, and the micro-hardness of the welded joint tends to be consistent. Compared with the weld without LSP treatment, the microhardness of the weld surface decreases after LSP treatment, while the micro-hardness of the heat-affected zone of the weld increases. Because the laser shock wave pressure can reach GPa level, when the impact weld zone is impacted, the impact zone will produce obvious plastic deformation, and the impact zone will have an instantaneous rapid heating and rapid cooling process, which may induce the phase transformation of the impact zone material [7, 10], so that the number of acicular martensite α ' phases in the TIG welding seam Fig. 8.17 Weld micro-hardness measurement position

Fig. 8.18 Micro-hardness distribution of backside of single square spot laser shock strengthened weld from weld centerline to base metal

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8 Strengthening Processes and Effect Evaluation of Welded Structures …

of the original TC4 titanium alloy is reduced, resulting in a decrease in the surface micro-hardness of the weld zone. In the study of the effect of plastic deformation on the microstructure and properties of magnesium alloys, [11] was proposed in the literature. The rotation of the metal base surface during the extrusion process leads to the shear stress concentration, so that the lattice rotation occurs at the grain boundary, a large number of high-angle grain boundaries are generated, and the grains in the shear stress concentration area are significantly refined. When the heataffected zone is treated by LSP, laser shock will increase the number of fine grains in the plastic deformation zone, which increases the micro-hardness of the surface. Further research on LSP-induced phase transition will be carried out in the follow-up work. Circular spot multiple laser shock strengthened weld, using the test load of 500 g test weld back Vickers hardness, experimental results are shown in Fig. 8.19. From Fig. 8.19, it can be seen that compared with the TIG welded joint of TC4 titanium alloy without LSP treatment, the micro-hardness of the material surface in the weld zone decreases and the micro-hardness of the material surface in the heat-affected zone increases after 1,2,4 and 6 LSP treatments. The micro-hardness of the material surface in the weld zone and the heat-affected zone fluctuates little, indicating that the micro-hardness of the material surface in the two regions tends to be consistent. Comparing the micro-hardness of the material surface of the weld zone and the heataffected zone of the TC4 titanium alloy TIG welded joint treated with different times of LSP, it was found that the micro-hardness of the material surface of the weld zone and the heat-affected zone gradually increased with the increase of LSP treatment times. 3. Metallographic microstructure Figure 8.20a and b are the SEM microstructures of the cross section of the weld near the surface of the material without LSP treatment and the three LSP treatments of the circular spot, respectively. It can be seen that the microstructure of the weld without LSP treatment is a long needle-like α' phase and lath-like α phase throughout the β phase grains, and the needle-like α' phases are parallel to each other. High dislocation density and twins exist in Ti martensitic α' phase [12]. The appearance of needle-like α' phase structure caused a large number of phase boundaries, which strengthened Fig. 8.19 Micro-hardness distribution from the center line of weld to the base metal on the back side of weld strengthened by multi-round spot laser shock peening

8.4 Mechanical Property and Fatigue Lives of TIG Welded TC4 Titanium …

371

the hardness of the weld zone. After LSP treatment, the long needle-like α' phase penetrating the whole β phase grains is hardly observed near the material surface, and the number of needle-like α' phase decreases. Usually, α + β type titanium alloy, weld hardness increase is due to a large number of acicular martensite α' phase caused by the reduction in the number of α' phase will reduce the weld hardness. In Fig. 8.20b, the large lath and coarse acicular structure are retained in the area far away from the surface of the material, indicating that under the selected LSP treatment process conditions, the laser shock effect has little effect on the large lath and coarse acicular structure in the β phase grain far away from the surface of the material. Figure 8.20c and d show the microstructures near the fusion line of the joints without LSP treatment and after three LSP treatments, which point to the side of the base metal and close to the surface of the material. As shown in Fig. 8.20c, the joint without LSP treatment, the fusion line near the base metal area of coarse equiaxial α + β structure, and can clearly see the rolling direction of the base metal. As shown in Fig. 8.20d, the rolling direction on the base metal is almost invisible. In addition to the coarse equiaxial α + β structure, a large number of fine equiaxial grains appear. On the surface of the material, fine equiaxial grains are dispersed near the surface. In the area far from the surface of the material, fine equiaxial grains are distributed between the coarse equiaxial α + β structure. After LSP treatment, a large number of fine equiaxial grains appear on the material surface, which helps to improve the micro-hardness of the material surface.

Fig. 8.20 The microstructure of TC4 welded joints by three times of laser shock peening and non-laser shock peening with circular spot

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8.4.2 Tensile Properties and Fatigue Lives 1. Joint tensile properties According to the national standard GB/T228-2002 at room temperature 25 °C, the tensile properties of TC4 titanium alloy welded joints were tested at a rate of 1 mm/min. Single square spot and multiple circular spot laser shock strengthened welds, joints without LSP treatment, tensile fracture appears in the weld zone. The tensile fractures of the joints treated by LSP appear in the base metal area far from the weld center. The experimental data obtained are shown in Table 8.9. It can be seen from Table 8.9 that compared with the TC4 titanium alloy TIG welded joint without LSP treatment, after different times of LSP treatment, the tensile strength and yield strength of TC4 titanium alloy TIG welded joint do not increase significantly, and the elongation after fracture increases to varying degrees. Under the selected experimental conditions, compared with the joints without LSP treatment, the average tensile strength and yield strength of the joints with square spot single LSP treatment increased by 5.6% and 8.2%, respectively, and the average elongation after fracture was also greatly improved by 66%. The average values of tensile strength and yield strength of the joints treated by twice LSP with round spot are lower than 1%, but the average elongation is decreased by 27%. The average values of tensile strength, yield strength, and elongation of the joints treated by 3 LSPs are 5.5%, 8%, and 36.5% lower than those treated by 1 LSP, respectively. After LSP treatment, the tensile fracture of the joint appears on the base metal, indicating that the tensile strength of the joint reaches or even exceeds the base metal. This is because LSP treatment causes strong plastic deformation on the surface of the Table 8.9 Tensile properties of TIG welded joints of TC4 titanium alloy treated by LSP and untreated Elongation δ/(%)

Number of physical volumes of strokes

Tensile strength σb/(MPa)

Yield strength σp0.2/(MPa)

0

1017.56

901.53

6.50

1006.24

890.06

7.02

1010.09

903.25

6.61

1071.02

954.14

11.34

1066.22

986.05

11.12

1069.03

975.02

10.98

1060.45

952.26

7.64

1071.10

960.56

8.68

1063.13

970.51

8.00

999.24

879.54

7.04

1015.90

904.53

7.00

1012.35

899.06

7.21

1

2

3

8.4 Mechanical Property and Fatigue Lives of TIG Welded TC4 Titanium … Table 8.10 Fatigue life of TC4 titanium alloy TIG welded joints with LSP treatment and untreated joints

373

Processing technique

Sample number

Fatigue life/(N)

TIG

1

68,373

2

56,583

3

54,348

4

60,135

1

75,913

2

83,266

3

86,579

4

80,638

TIG + LSP

weld and heat-affected zone, resulting in changes in the microstructure of the plastic deformation zone. At the same time, LSP treatment greatly increases the dislocation density [13]. According to the Hall–Petch relationship, the strength of welded joints will be improved. 2. Fatigue life of joints The tensile fatigue properties of TC4 titanium alloy TIG welded joints before and after square spot single LSP treatment are shown in Table 8.10. The maximum stress σmax = 600 MPa, the stress ratio R = 0.1, the test frequency f = 30 Hz, and the waveform is sine wave. The experimental data on the fracture of the base metal in the experiment are not listed in Table 8.10. It can be seen from the table that the fatigue life of TC4 titanium alloy TIG welded joints can be improved by LSP process. Compared with the TIG welded joint of TC4 titanium alloy without LSP treatment, the average fatigue life of the LSP treated joint is increased by 36.3%. A large number of studies at home and abroad show that laser shock pressure up to GPa not only makes the impact zone plastic deformation, but also has a high residual compressive stress. In the process of tensile fatigue test, for the welded joint treated by LSP, due to the existence of surface residual compressive stress, it can play the role of negative average stress [14], weaken the influence of tensile stress in the process of experiment, delay the generation and propagation speed of fatigue crack, and improve the fatigue life of the joint.

8.4.3 Fatigue Fracture Analysis The fatigue fracture morphology is shown in Fig. 8.21. After careful observation and comparison of the fatigue fractures after LSP treatment and as-welded, it was found that the fatigue fractures of the joints without LSP treatment had obvious fatigue sources. The typical macroscopic fracture morphology is shown in Fig. 8.21a. The crack starts from the fatigue source, perpendicular to the stress direction, and extends to the inside of the weld. The fatigue fracture is relatively flat, indicating that the

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Fig. 8.21 Fatigue fracture panorama of TC4 welded joint

crack propagation process is less resistant. The fatigue bands in the early, middle, and late stages of fatigue crack propagation can be clearly observed from the whole fatigue fracture. The fatigue fracture of LSP treated joints is shown in Fig. 8.21b. It is difficult to observe the fatigue striations in the early and middle stages of crack propagation in the whole fatigue fracture, only the fatigue striations in the later stage of fatigue crack propagation can be found. Figure 8.22 is the fatigue striation of the two joints in the later stage of fatigue crack propagation. Compared with the LSP treated joints, the fatigue striation of the joints without LSP treatment is wider, and the width of the striation is about 1.2 μm. For LSP treated joints, the fatigue bands at the later stage of crack propagation are poorly oriented and narrow, as shown in Fig. 8.22b, indicating that the fatigue propagation rate of cracks is very small even at the later stage of crack propagation. This may be related to the fact that the dislocation density in the joint is greatly increased after LSP treatment, and the formed network structure hinders the crack propagation path. Along the weld penetration direction, the dimple morphology characteristics of the instantaneous fracture zone were observed in the area about 0.5 mm from the surface. The observed dimples are shown in Fig. 8.23. It can be seen from Fig. 8.23a that the dimples on the weld surface of the fatigue fracture zone without LSP treatment are typical shear dimples, and the coarse and fine dimples are intertwined. It shows

Fig. 8.22 Fatigue striation of fatigue fracture at later stage of crack propagation

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375

that the size of α phase and acicular α ‘ phase produced by β phase transformation is not uniform during the solidification and crystallization of TC4 titanium alloy molten pool. From Fig. 8.23b, the dimples at the tensile fracture of the TIG welded joint after one LSP treatment are a large number of equiaxial dimples because the fracture appears in the base metal area near the fusion line. The dimples shown in Fig. 8.23c and d appeared in the tensile fracture of TIG welded joints after 2 and 3 LSP treatments. However, the proportion of dimples shown in Fig. 8.23d was larger in the tensile fracture of joints after 3 LSP treatments. The tensile fracture of TIC welded joints after 5 and 6 LSP treatments are shown in Fig. 8.23e and g. Figure 8.23f and h are local enlargement of Fig. 8.23e and g A, respectively. Figure 8.23f appears convex dimples, Fig. 8.23e appears large and shallow dimples in B area, Fig. 8.23h appears uniform small and convex dimples, Fig. 8.23g appears large and shallow dimples in B area, and there are traces of intergranular fracture. During LSP treatment, the material in the impact area will produce strong plastic deformation, accompanied by instantaneous rapid heat and rapid cooling process, which makes the material in the impact area undergo phase transformation [14]. With the increase of impact times, the micro-hardness of the weld surface and the surface of the heat-affected zone increases, indicating that with the increase of impact times, the effect of plastic deformation induced by LSP treatment on material phase transformation is more obvious. The LSP treatment makes the phase transition induced by the surface deformation of the material a process of rapid heating and rapid cooling. The growth process of the equiaxial crystal produced by the phase transition is terminated rapidly, but the trace of the outward growth of the equiaxial crystal is retained, so that the dimples shown in Fig. 8.23d appear in the tensile fracture. As for the dimples shown in Figs. 8.23c and d that appear in the same joint after different times of LSP treatment, it may be due to the different plastic deformations of the α phase and the α ‘ phase in different original β phases. The degree of phase transformation induced by plastic deformation is also inconsistent, resulting in different shapes of dimples in different regions of the same tensile fracture. For the TIG welding seam of TC4 titanium alloy, the original weld microstructure is that the coarse β phase surrounds the α phase and the needle-like α' phase. In different β phases, the orientation of the α phase and the needle-like α' phase is inconsistent, which causes the angle difference between the different orientations of the α phase and the needle-like α' phase and the direction of the laser shock force. Under the same laser shock force, the α phase and the needle-like α' phase with a large angle between the direction of the laser shock force appear serious plastic deformation. And because the LSP treatment is accompanied by instantaneous quenching, instantaneous heating makes the deformation of large area to produce new equiaxial α + β recrystallized structure and grow, followed by quenching process makes the grain growth stop, so that the grain will not be seriously coarsened. In this way, a large number of convex dimples appear in the fatigue fracture transient region as shown in Figs. 8.23f and h. The decrease in the number of α' phases in the microstructure of the LSPed weld near the material surface also illustrates the correctness of this view.

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(a)

(b)

(c)

(d)

(e)

(f)

(g)

(h)

Fig. 8.23 The dimples in the instantaneous fracture zone of fatigue fracture under different impact times a Non-impact strengthening; b 1 impact strengthening; c 2 times impact strengthening; d 3 impact strengthening; e 5 times impact strengthening; f is the enlarged area A of Figure (e); g 6 times impact strengthening; h is the enlarged area A of Figure (g)

8.5 Mechanical Property and Fatigue Lives of Laser-Welded TC4 Sheets …

377

8.5 Mechanical Property and Fatigue Lives of Laser-Welded TC4 Sheets Treated by Laser Shock Peening with Double Sides and Different Sequences 8.5.1 Micro-hardness and Residual Stress The TC4 test plate was welded by laser welding, and the welding direction was perpendicular to the rolling direction of the test plate. Firstly, the test plate was welded by laser deep penetration welding, and then the front of the weld was modified by laser defocus welding. After welding, vacuum stress relief annealing was carried out. The annealing temperature was 650 °C, the vacuum degree was 9 × 10-2 Pa, and the furnace was slowly cooled after holding for 1.5 h. The welded test plate after annealing is made into fatigue specimen by wire cutting, whose dimensions are shown in Fig. 8.24. The fatigue specimens were divided into 3 groups, group A was not treated by laser shot peening; group B first strengthens the weld front, and then strengthens the weld back; group C first strengthens the weld back, and then strengthens the weld front. Nd: YAG nanosecond pulse laser with pulse energy of 25 J, pulse width of 15 ns, frequency of 1 Hz, and spot diameter of 4 mm was used in laser shot peening test. By means of spot overlap, the overlap rate between spots is 40%, and the weld zone, heat affected zone, and some base metal zone of the specimen can be completely covered. 1. Micro-hardness change on the back of the weld Figure 8.25 shows the micro-hardness distribution near the back surface of the weld [15]. It can be seen from Fig. 8.25 that there are obvious undercut defects on the back of the weld, and the stress concentration is serious. The change of microhardness in this area is of great significance to the improvement of the mechanical properties of the whole welded specimen. The micro-hardness of the fusion zone of the unstrengthened sample is significantly higher than that of the base metal, which is related to the acicular martensite α' phase formed by the TC4 titanium alloy during the rapid solidification cooling stage of laser welding. As the position of the test point moves from the weld center to the base metal, the peak welding temperature and temperature gradient experienced by the measured part gradually decrease, the content of the acicular martensite α' phase decreases continuously, and the micro-hardness decreases to 330HV. Due to the narrow fusion zone and heataffected zone of the laser welded joint, the gradient of micro-hardness decrease is Fig. 8.24 Fatigue specimen size

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Fig. 8.25 Micro-hardness distribution near the back of weld

large from the center of the fusion zone to the edge of the heat-affected zone less than 1 mm, especially in the weld undercut area of 0.6 ~ 0.8 mm from the weld center, the micro-hardness value begins to decrease greatly, which is not conducive to the inhibition of crack initiation. After two different impact sequences of double-sided laser shot peening, the micro-hardness of the undercut area was improved, and the micro-hardness of the heat-affected zone of the group C sample was higher than that of the group B sample. The increase of the micro-hardness of the weld surface is beneficial to inhibit the initiation of fatigue cracks, thereby improving the fatigue life of the welded specimen. 2. The residual stress change on the back of the weld According to EN15305-2008 standard, the residual stress distribution perpendicular to the weld direction of the back of three groups of fatigue specimens was tested on the LXRD high-power residual stress tester, and the position distribution of the test points was shown in Fig. 8.26. Figure 8.27 shows the residual stress distribution on the back of the weld [15]. It can be seen from the figure that the residual stress state on the back of the weld of the two groups of double-sided laser shock peening specimens has changed qualitatively compared with the untreated specimens, and the tensile residual stress state has changed from the tensile residual stress state to the compressive residual Fig. 8.26 Schematic diagram of the distribution of residual stress test points

8.5 Mechanical Property and Fatigue Lives of Laser-Welded TC4 Sheets …

379

Fig. 8.27 Residual stress distribution on the back of the weld

stress state. The compressive residual stress distribution on the back of the weld of group B specimen and group C specimen is obviously different, the amplitude of compressive residual stress around the weld of group C specimen is higher than that of group B specimen as a whole, and the peak value at the undercut of the weld is −564.37 MPa, the peak value of the compressive residual stress of the weld of group B specimen is −317.86 MPa, the peak compressive residual stress of the group C sample is 1.8 times that of the group B specimen, and the strengthening effect on the back of the weld is more significant. Since fatigue cracks mainly germinate in the stress concentration part of the undercut on the back of the weld, the compressive residual stress of high amplitude at the undercut can balance the maximum tensile stress of the cyclic load, and the compressive residual stress, on the one hand, weakens the influence of the maximum tensile stress, reduces the stress intensity factor ΔKmax of the crack tip, thereby reducing the crack growth rate: on the other hand, the cyclic stress ratio R value is changed, and when the minimum tensile stress is transformed into the maximum compressive stress, the R value changes from positive to negative. Empirical formula based on the effect of stress ratio R on the fatigue cracks growth threshold ΔKth [16]: ΔK th = (1 − R)γ ΔK th0

(8.1)

where ΔK th0 —the crack propagation threshold when the stress ratio R is 0; γ—the constant measured by the test, the value of which varies between 0 ~ 1; The change of R value sign causes the fatigue threshold to increase, thereby inhibiting the fatigue crack source from propagating into crack.

8.5.2 The Comparison of Median Fatigue Life Three groups of fatigue samples were subjected to tensile-tensile fatigue test on MTS810 testing machine, the test temperature was room temperature, the maximum

380

8 Strengthening Processes and Effect Evaluation of Welded Structures …

load was 520 MPa, and the stress ratio R = 0.1. In order to compare the fatigue life of the test plates under different treatment conditions, each group of fatigue samples was subjected to fatigue test under the same maximum load of 520 MPa, and the fatigue cycle times of each group of samples were counted, and the groups were compared and analyzed separately. First, the samples of group B and group A were compared, and the significance α was 5% and 10%. An F test was performed on the maternal standard deviation of the fatigue life of the two groups of specimens to determine whether there was a significant difference in the dispersion of the fatigue life of the two groups. The average fatigue life of the two groups of samples was then t-tested to determine whether there was a conditional error in the average fatigue life of the two groups. Then the confidence level γ = 95% was selected to compare the median lifetime of the two groups of samples [17], and the calculation results are shown in Table 1 [15]. Among them, the variance S2 represents the dispersion of the fatigue life of each group, and the F value is the variance ratio statistic of the fatigue life of the two groups of samples, which is used to characterize the standard deviation of fatigue life. The t-value is the average statistic of fatigue life of the two groups of samples, which is used to characterize the difference in the average fatigue life value. From Table 8.11, it can be seen that the F values of group A and group B are less than Fα, indicating that there is no significant difference in fatigue life dispersion between group A and group B. Moreover, the t-values of group A and group B were greater than tα, indicating that the difference between the average fatigue life of samples in group A and group B was significant. Taking the confidence of 95% for interval estimation, under the maximum load of 520 MPa, the median fatigue life of the samples in group B was 1.68 ~ 4.17 times that of the median fatigue life in group A. Similarly, comparing the samples of group C and group A, the median fatigue life of the samples in group C is 3.61 ~ 9.56 times that of the median fatigue life in group A. The above fatigue test results show that the median fatigue life of TC4 laser welded specimens is significantly improved after double-sided laser shock peening, and the impact sequence of group C is more significant than that of group B on the median fatigue life of the specimen.

8.5.3 Fatigue Fracture Analysis The fractures of three sets of fatigued specimens were observed under scanning electron microscopy, as shown in Fig. 8.28 [15]. The fatigue cracks of the three groups of specimens were born in the weld undercut stress concentration part, the fatigue fracture of the untreated specimen had many obvious fatigue sources, the fatigue source extended along the weld surface covered the entire weld undercut area, the location of the main crack was located at the edge of the specimen, due to the presence of edges and corners at the edge, the stress concentration was the most serious, and fatigue cracks first germinated here. However, the fatigue fracture crack source after laser shot peening only germinates at the undercut edge of the specimen,

520

520

9

9

A

9

A

C

9

B

4.2296

4.6672

4.2296

4.6995

0.0372

0.0522

0.0372

0.0939

Maximum load /(MPa) Group Quantity Log-mean lifetime Variance S 2

Table 8.11 Fatigue life test table of fatigue samples in different treatment states

1.4

2.52

4.43

4.43

3.44

3.44

α = 5% α = 10%

F = S 2 (Big) /S 2 (Small) F α

7.72 2.12

1.75

1.75

α = 5% α = 10%

tα 3.89 2.12

t

8.5 Mechanical Property and Fatigue Lives of Laser-Welded TC4 Sheets … 381

382

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Fig. 8.28 Fracture morphology of different fatigue specimens

and the crack initiation in other parts is inhibited. It can be seen from the residual stress test that a certain thickness of reinforcement layer is formed on the back of the weld after laser shot peening, the residual compressive stress is introduced, and the residual compressive stress in the tensile-tensile fatigue test plays the role of negative average stress on the external tensile load, which increases the threshold value of crack initiation, so that the expansion of fatigue crack along the surface is inhibited. Comparing group B specimens and group C specimens, it can be seen that the fatigue crack propagation distance along the surface of group C specimen is shorter, indicating that the inhibition effect of fatigue crack propagation along the weld surface is stronger after group C specimen strengthening [18]. Figure 8.29 is a local magnification of the fatigue crack initiation zone in Fig. 8.28 [15]. Compared with the path of fatigue crack initiation zone of the three groups of specimens extending along the thickness direction, the fatigue crack of the unreinforced specimen began to germinate from the undercut of the weld, perpendicular to the direction of the tensile load axis, expanded along the thickness direction of the weld, the fatigue fracture was relatively flat, and the number of fatigue steps in the crack initiation area was small, the radiation direction was relatively straight, and it was basically parallel to the thickness direction. After laser shot peening, the number of steps in the fatigue crack initiation area of the specimen gradually increased, the radiation direction of the fatigue step was deflected and twisted to a certain extent, the fatigue expansion path became more tortuous, and the degree of deflection of the group C sample was more obvious than that of the group B sample. This is due to the introduction of high amplitude residual compressive stress on the weld surface after laser shot peening, which reduces the stress intensity factor of the crack tip at the undercut of the weld, fatigue cracks are not easy to form a linear crack source on the surface of the weld, but priority is given to the point germination in the part with a large stress concentration coefficient, with the increase of the amplitude of residual compressive stress, the linear fatigue source decreases, the point fatigue

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383

Fig. 8.29 Local magnification of fatigue crack initiation area in Fig. 8.28

source increases, the radiation inclination angle of the fatigue groove increases, and the path is more torturous. Figure 8.30 shows the morphology of the transient fracture zone of fatigue samples in different treatment states [15], all three of which show the characteristics of static load transient fracture, and a large number of equiaxial ligaments are distributed in the transient region. The mechanism of rupture is cavitation aggregation, which shows the fracture characteristics of ductile materials. First, cavities are formed at the second phase by particles inside the material, and the cavities gradually grow by slipping and connect with other voids to form equiaxial tenacities. As can be seen from Fig. 8.30a, the size of the tendon socket of the unreinforced specimen is small and uniform. Compared with the unstrengthened specimen, the size of the ligament socket in the transient area after laser shot peening became significantly larger and deeper, and the size of the ligament socket of group C specimen was larger than that of group B specimen, and there were several small tendiment fossa distributed in the large tenament fossa, and the tearing ridge between the tendon sockets increased, as shown in Fig. 8.30b and c. Since the size and depth of the ligament socket are mainly affected by the size and distribution of the second phase particle and the relative plasticity of the metal material itself, under the same fracture conditions, the larger the size and deeper the ligament socket, the more tearing ridges indicate the better the plasticity of the material, which indicates that the plasticity of the material in the transient zone after double-sided laser shot peening is improved, and the plasticity of the group C sample is more obvious.

Fig. 8.30 Transient zone morphology of fatigue specimens in different treatment states

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8.6 Mechanical Properties and Corrosion Properties of TA15 Electron Beam Welds TA15 is a α-phase or α + β phase titanium alloy, which is widely used in thirdand fourth-generation fighters because of its good thermal stability and long-term operation at 500 °C [19, 20]. Electron beam welding is the preferred process for titanium alloy welding [21, 22]. Compared with the matrix material, titanium alloy electron beam welding joint characteristics, stress concentration, low stress corrosion performance, and sensitive areas are easy to lead to structural failure. Post-weld heat treatment is usually used to eliminate the residual stress of welding, however, the mechanical properties and corrosion properties of the weld are not significantly improved. Therefore, improving the mechanical properties and corrosion properties of titanium alloy welds is the focus of research. A new surface modification technology, laser shock peening is widely used in foreign aviation industry and other industrial fields. Due to the wide application prospects of LSP, this technology is applied to China’s aviation industry [23–27]. This section focuses on the micro-hardness, corrosion properties, and mechanical tensile properties of welds at 25° room temperature and 300° high temperature to optimize weld performance by laser shock peened TA15 electron beam welded joints [28].

8.6.1 Micro-hardness 1. Experimental materials and experimental methods TA15 is a near-α phase titanium alloy [29, 30], its chemical composition (wt%) is: 6.3Al, 1.7Mo, 2.0Zr, 2.0 V, 0.08O, 0.05Fe, 0.03C, 0.03N, 0.04Si, 0.012H, and other titanium. The size of TA15 specimen is 400 × 300 × 3 mm, and the electron beam welding parameters are shown in Table 8.12. The process parameters of lap laser shock peening TA15 weld fusion zone, heat affected zone and part of the base metal area are: Q conversion Nd:YAG laser, wavelength 1064 nm, pulse width 10 ns, laser energy 10 J, spot diameter 3 mm. According to the national standard GB3075-82, the size of the tensile test specimen is shown in Fig. 8.31. Gleeble1500 testing machine is used to detect the tensile properties of welds. 2. Micro-hardness DHV-1000 micro-hardness test equipment measures the surface micro-hardness of TA15 weld before and after LSP, as shown in Fig. 8.32. The micro-hardness measurement points are evenly distributed perpendicular to the weld direction at intervals of 0.5 mm. The micro-hardness measurement load is 500 g and the loading time is 15 s.

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Table 8.12 Electron beam welding parameters Welding voltage/(kV)

Focused current /(mA)

Electron beam current/(mA)

Welding speed/(m.min−1 )

140

339

8.5

0.6

Fig. 8.31 Experimental specimen size

Fig. 8.32 Micro-hardness value of laser impact strengthened weld a non-LSP; b LSP

It can be seen from Fig. 8.32a that the micro-hardness value of the surface of the untreated weld is symmetrically distributed in a saddle-shape, that is, the maximum micro-hardness value is in the heat-affected zone on both sides of the weld, and the minimum micro-hardness value is in the center of the weld and at the base metal. The micro-hardness value of the base metal is 3800 MPa, the micro-hardness value of the heat-affected zone is 4760 MPa, and the micro-hardness value of the weld center fluctuates. The micro-hardness value near the heat-affected zone is maintained at about 3500 MPa, but the micro-hardness value in the center of the heat-affected zone increases to 4550 MPa. It can be seen from Fig. 8.32b that the micro-hardness value of the surface of the LSP weld does not show a saddle-shaped distribution, the microhardness value of the weld center and the heat-affected zone is obviously different, basically maintained at about 5000 MPa, and the micro-hardness of the base metal increases to about 4500 MPa.

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8.6.2 Corrosion Properties Many aviation fuselages, wings, landing drives, and other structures that are exposed to air for long periods of time are inevitably subject to corrosive media. Due to longterm salt spray jetting or seawater corrosion, the working environment is very harsh for the carrier aircraft. In such a harsh environment, the fatigue life of titanium alloy welded joints is shortened. Therefore, the anti-corrosion treatment method of titanium alloy welded joints is very necessary to improve the reliability of aircraft. Laser shock peening induces the residual compressive stress of the metal surface [14, 31, 32], and the residual compressive stress state delays the occurrence of metal stress corrosion cracks [33]. In addition, the residual compressive stress can weaken the alternating load effects of the metal material. Before and after LSP, the corrosion performance of TA15 electron beam weld under 3% sodium chloride electrochemical corrosion is shown in Fig. 8.33. It can be seen from Fig. 8.33 that the untreated reinforced TA15 weld is relatively stable in the process of cathodic polarization, with the increase of corrosion voltage. When the voltage increases to −104.5 mV corrosion potential, the cathodic polarization process is converted to anodizing. During the anodic polarization process, the corrosion current continues to increase to 560 mV as the potential increases, and after this potential point, the current changes slowly. Due to the occurrence of passivation phenomena, the corrosion process is converted into a stable passivation area. The laser shock peened the TA15 weld, and the current slowly decreases as the potential increases during the cathodic polarization process. After the −79.5 mV potential appears, the corrosion process transitions to the anodic polarization region as the potential continues to increase. When the potential is less than 546 mV, the current floats larger. When the potential increases to 546 mV, the current hardly changes as the potential continues to increase, indicating that passivation has occurred. It can be seen from Fig. 8.33 that the TA15 weld has good corrosion performance. The corrosion performance of laser shock peened TA15 weld has not been significantly improved. Unsuitable laser process parameters such as laser energy and impact Fig. 8.33 Polarization curve in 3% sodium chloride solution

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387

times may reduce the corrosion performance of TA15. In the future, the optimization of process parameters will be further studied and optimized, and laser impact strengthening TA15 will obtain good corrosion performance.

8.6.3 Tensile Properties In room temperature and high temperature environments, five TA15 welds before and after LSP were tensile experiments. The results of the average tensile properties of the weld are shown in Table 8.13. Compared with the untreated weld, the elongation of the laser impact reinforced TA15 weld is improved at 25° and 300°, but there is no obvious difference in tensile strength. At 25°, the elongation of the laser shock peened TA15 weld was 23.8%, which was 3.5 times that of the untreated weld. At 300°, the elongation of laser impact reinforced TA15 weld is 5.6% greater than that of non-impact reinforced TA15 weld.

8.6.4 Tensile Fracture Analysis The SEM scanning position is 200 μm from the weld surface. All TA15 weld tensile fatigue fractures are ligacious brittle fractures, as shown in Figs. 8.34 and 8.35. It can be seen from Fig. 8.34 that compared with the untreated weld ligament fracture, the laser shock peened weld toughness fracture is significantly different. At 25° room temperature, the tensile fracture position of the untreated TA15 weld appears in the center of the weld. With different fracture heights, the tensile fracture surface is rough. As can be seen from Fig. 8.34a, the traces of fracture along the boundary of the β phase are obvious. The laser shock peened TA15 weld fracture location occurs at the base metal in the near-heat affected zone, and the fracture is relatively flat. It can be seen from Fig. 8.34b that the fracture traces along the boundary of the β phase cannot be clearly observed, indicating that laser shock peening improves the fatigue performance of TA15 weld. At a high temperature of 300°, the tensile fracture position of TA15 weld before and after LSP appeared at the base metal in the near heat-affected zone, and the fracture was relatively flat. Figure 8.35a and b show the tensile fracture surface SEM Table 8.13 Tensile properties of TA15 electron beam weld Experimental conditions

Tensile strength σb /(MPa)

25°C

–

1052.09

6.8

LSP

1066.13

23.8

–

885.59

30.5

LSP

877.75

36.1

300°C

Elongation δ/(%)

388

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Fig. 8.34 Morphology of TA15 electron beam weld fracture at 25 °C a non-LSP; b LSP

Fig. 8.35 Morphology of TA15 electron beam weld fracture at 300 °C a non-LSP; b LSP

diagram. Because the weld strength is quite high, tensile fractures appear at the base metal. The experimental results show that the laser shock peened TA15 weld maintains good tensile properties at both 25° and 300° temperatures.

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