Tailor Welded Blanks for Advanced Manufacturing

Tailor welded blanks for advanced manufacturing © Woodhead Publishing Limited, 2011 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 ...

Author: Brad Kinsey | Xin Wu

387 downloads 1232 Views 12MB Size Report

This content was uploaded by our users and we assume good faith they have the permission to share this book. If you own the copyright to this book and it is wrongfully on our website, we offer a simple DMCA procedure to remove your content from our site. Start by pressing the button below!
Report copyright / DMCA form

DOWNLOAD PDF

Tailor welded blanks for advanced manufacturing

© Woodhead Publishing Limited, 2011

1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

Related titles: Materials, design and manufacturing for lightweight vehicles (ISBN 978-1-84569-463-0) Research programs into the manufacture of lightweight automobiles have led to the consideration of a variety of materials, such as high strength steels, aluminium alloys, magnesium alloys, plastics and composites. This research is driven by a need to reduce fuel consumption to preserve dwindling hydrocarbon resources without compromising other attributes such as safety, performance, recyclability and cost. This important book enables engineers to not only learn about the materials being considered for lightweight automobiles, but also to compare their characteristics and properties. It also covers issues such as crashworthiness and recycling. Welding and joining of magnesium alloys (ISBN 978-1-84569-692-4) This book covers all aspects of the welding and joining of magnesium alloys. The use of magnesium and its alloys is increasing due to their light weight and other properties. Part I covers welding metallurgy, preparation methods and welding materials as well as corrosion protection. Part II covers the various welding and other technologies that can be used for joining magnesium alloys. Minimization of welding distortion and buckling (ISBN 978-1-84569-662-7) Weld distortion and buckling is caused by expansion and contraction of the weld and base metal during the heating and cooling cycle of the welding process. This book reviews ways of understanding and modelling welding residual stress and distortion. It also discusses a range of techniques for minimizing bowing, buckling and angular distortion. Details of these and other Woodhead Publishing materials books can be obtained by: • •

visiting our web site at www.woodheadpublishing.com contacting Customer Services (e-mail: [email protected]; fax: +44 (0) 1223 832819; tel.: +44 (0) 1223 499140 ext. 130; address: Woodhead Publishing Limited, 80 High Street, Sawston, Cambridge CB22 3HJ, UK)

If you would like to receive information on forthcoming titles, please send your address details to: Francis Dodds (address, tel. and fax as above; e-mail: francis. [email protected]). Please confirm which subject areas you are interested in.


Tailor welded blanks for advanced manufacturing Edited by Brad L. Kinsey and Xin Wu


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

Published by Woodhead Publishing Limited, 80 High Street, Sawston, Cambridge CB22 3HJ, UK www.woodheadpublishing.com Woodhead Publishing, 1518 Walnut Street, Suite 1100, Philadelphia, PA 19102-3406, USA Woodhead Publishing India Private Limited, G-2, Vardaan House, 7/28 Ansari Road, Daryaganj, New Delhi – 110002, India www.woodheadpublishingindia.com First published 2011, Woodhead Publishing Limited © Woodhead Publishing Limited, 2011 The authors have asserted their moral rights. This book contains information obtained from authentic and highly regarded sources. Reprinted material is quoted with permission, and sources are indicated. Reasonable efforts have been made to publish reliable data and information, but the authors and the publisher cannot assume responsibility for the validity of all materials. Neither the authors nor the publisher, nor anyone else associated with this publication, shall be liable for any loss, damage or liability directly or indirectly caused or alleged to be caused by this book. Neither this book nor any part may be reproduced or transmitted in any form or by any means, electronic or mechanical, including photocopying, microfilming and recording, or by any information storage or retrieval system, without permission in writing from Woodhead Publishing Limited. The consent of Woodhead Publishing Limited does not extend to copying for general distribution, for promotion, for creating new works, or for resale. Specific permission must be obtained in writing from Woodhead Publishing Limited for such copying. Trademark notice: Product or corporate names may be trademarks or registered trademarks, and are used only for identification and explanation, without intent to infringe. British Library Cataloguing in Publication Data A catalogue record for this book is available from the British Library. ISBN 978-1-84569-704-4 (print) ISBN 978-0-85709-385-1 (online) The publisher’s policy is to use permanent paper from mills that operate a sustainable forestry policy, and which has been manufactured from pulp which is processed using acid-free and elemental chlorine-free practices. Furthermore, the publisher ensures that the text paper and cover board used have met acceptable environmental accreditation standards. Typeset by RefineCatch Limited, Bungay, Suffolk Printed by TJI Digital, Padstow, Cornwall, UK


Contents

Contributor contact details Preface

ix xi

Part I Processing and modeling

1

1

3

Weld integrity of tailor welded blanks M. M. LI, TWB Company, USA

1.1 1.2 1.3 1.4 1.5 1.6 2

Introduction Typical weld imperfections Testing methods Quality control in production Conclusions References

3 4 11 19 22 22

Deformation of tailor welded blanks during forming

24

K. NARASIMHAN, IIT Bombay, India and R. G. NARAYANAN, IIT Guwahati, India

2.1 2.2 2.3 2.4 2.5 2.6 2.7 2.8 2.9 2.10

Introduction Estimation of the constitutive behavior of the weld region Methods to evaluate the weld width (or cross-sectional area) in tailor welded blanks (TWBs) Forming limits of TWBs: influence of weld orientation Weld line movement Design considerations for TWB forming Simulation of TWB forming behavior Conclusions Acknowledgment References

24 25 26 32 36 39 40 43 45 45

v © Woodhead Publishing Limited, 2011

1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

vi

Contents

3

Mechanics-based modeling of tailor welded blank forming

48

B. L. KINSEY, University of New Hampshire, USA

3.1 3.2 3.3 3.4 3.5 3.6 3.7

Introduction Thickness and strength ratio analysis Determination of weld line movement and forming height Determination of material draw-in ratios Determination of non-uniform binder force Conclusions References

48 49 51 54 63 66 66

4

Numerical simulation modeling of tailor welded blank forming

68

A. A. ZADPOOR, Materials Innovation Institute (M2i) and Delft University of Technology, the Netherlands and J. SINKE and R. BENEDICTUS, Delft University of Technology, the Netherlands

4.1 4.2 4.3 4.4 4.5 4.6 4.7 4.8 4.9 4.10

Introduction Finite element method (FEM) modeling of the welded zones Material models Theoretical failure prediction of tailor welded blanks (TWBs) Some topics in design and optimization of TWBs Conclusions Future trends Sources of further information and advice Acknowledgments References

68 71 75 82 85 88 89 89 90 90

Part II Applications

95

5

97

Lightweight metal alloy tailor welded blanks R. PADMANABHAN, VIT University, India and University of Coimbra, Portugal and M. C. OLIVEIRA and L. F. MENEZES, University of Coimbra, Portugal

5.1 5.2 5.3 5.4 5.5 5.6

Introduction Lightweight metal alloy tailor welded blanks (LWMA TWBs) LWMA TWB formability LWMA TWB benefits/recycling Sources of further information and advice References

97 101 109 114 115 115

6

Advanced high-strength steel tailor welded blanks (AHSS-TWBs)

118

X. WU, Wayne State University, USA

6.1

Introduction to advanced high-strength steel (AHSS)


118

Contents

6.2 6.3 6.4 6.5 6.6 6.7 6.8 7

Types of advanced high-strength steels and their characteristics Fabrication of advanced high-strength steels for tailor welded blanks (AHSS-TWBs) Properties and formability of AHSS-TWBs Understanding the evolution of microstructure and its impact on properties of AHSS-TWBs Other manufacturing processes related to AHSS-TWBs Conclusions References Tailor welded blanks for the automotive industry

vii

120 123 127 149 157 159 160 164

B. L. KINSEY, University of New Hampshire, USA

7.1 7.2 7.3 7.4 7.5 7.6 7.7 7.8 7.9 7.10 7.11 8

Introduction Door inner example Historical perspective Advantages of tailor welded blanks (TWBs) Disadvantages of TWBs Research efforts for TWBs TWB forming methods Welding processes for TWBs Materials used to produce TWBs Conclusions References

164 166 167 168 170 174 174 177 177 178 178

Tailor made blanks for the aerospace industry

181

J. SINKE, Delft University of Technology, the Netherlands, A. A. ZADPOOR, Materials Innovation Institute (M2i) and Delft University of Technology, the Netherlands and R. BENEDICTUS, Delft University of Technology, the Netherlands

8.1 8.2 8.3 8.4 8.5 8.6

Introduction The tailor made blank (TMB) concept and the aircraft industry Future trends Conclusions Acknowledgements References

181

Index

203


184 198 199 200 200

1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

Contributor contact details

(* = main contact)

Editors

Chapter 2

Professor Brad L. Kinsey Mechanical Engineering Department University of New Hampshire 33 Academic Way Durham NH 03824 USA

Professor K. Narasimhan* Department of Metallurgical Engineering and Materials Science IIT Bombay Powai Mumbai 400 076 India

E-mail: [email protected]

E-mail: [email protected]; nara@ met.iitb.ac.in

Professor Xin Wu Department of Mechanical Engineering Wayne State University Detroit MI 48202 USA E-mail: [email protected]

R. G. Narayanan Department of Mechanical Engineering IIT Guwahati Guwahati 781 039 Assam India E-mail: [email protected]

Chapter 1

Chapter 3

Dr Michael M. Li TWB Company, LLC 1600 Nadeau Rd Monroe MI 48162 USA E-mail: [email protected]

Professor Brad L. Kinsey Mechanical Engineering Department University of New Hampshire 33 Academic Way Durham NH 03824 USA E-mail: [email protected]

ix © Woodhead Publishing Limited, 2011

1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

x 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

Contributor contact details

Chapter 4

Chapter 7

A. A. Zadpoor, Jos Sinke* and R. Benedictus Faculty of Aerospace Engineering Delft University of Technology (TUDelft) Kluyverweg 1 Delft 2629HS The Netherlands

Professor Brad L. Kinsey Mechanical Engineering Department University of New Hampshire 33 Academic Way Durham NH 03824 USA

E-mail: [email protected]; [email protected]; r. [email protected]

Chapter 5 Professor R. Padmanabhan*, M. C. Oliveira and L. F. Menezes CEMUC Department of Mechanical Engineering University of Coimbra Coimbra 3030 788 Portugal

E-mail: [email protected]

Chapter 8 Jos Sinke*, A. A. Zadpoor and R. Benedictus Faculty of Aerospace Engineering Delft University of Technology (TUDelft) Kluyverweg 1 Delft 2629HS The Netherlands E-mail: [email protected]; [email protected]; [email protected]

E-mail: [email protected]; [email protected]

Chapter 6 Professor Xin Wu Department of Mechanical Engineering Wayne State University Detroit MI 48202 USA E-mail: [email protected]


Preface

In several industries (e.g. the automotive and aerospace industries), the need for lightweight and cost-effective products with exceptional performance is essential for success. Tailor welded blanks (TWBs) offer an excellent means to meet these competing and seemingly contradictory demands. Thirty years after this technology was introduced, and with the increasing concerns in energy conservation and environmental protection, now is a perfect time to publish this book to summarize the current knowledge and state-of-the art in forming TWBs. The goal is to promote sustained TWB applications for new materials, new processes and new concepts in design for manufacturing. To create the structural and skin components for vehicles, sheet metal is used due to the superior strength-to-weight ratio compared with bulk material products. Traditionally, to create a sheet metal assembly with various components, individual sheet metal parts are formed and then subsequently welded (e.g. spot welded) together. Alternatively, in TWBs, multiple sheet metals are seam welded (or bonded) together prior to the deformation process, thus requiring only one forming operation. The multiple sheets could be of various material alloys, thicknesses and/or surface treatments (e.g. galvanized versus non-galvanized) in order to ‘tailor’ the location of specific material properties into the final part. This technology offers numerous advantages over the traditional multiple forming and subsequent spot welding method including reduced cost, component mass and vibrations/noise in the assembly as well as improved material utilization, structural integrity, corrosion resistance and dimensional accuracy. However, formability concerns (e.g. reduced strain at failure and weld line movement) are created due to the dissimilar mechanical properties of the weld seam material and heataffected zone and the various sheet metal strengths in the TWB combination. The benefits, though, are so desirable that means to fabricate TWBs are of significant interest to multiple industries despite the formability concerns. In this book, various aspects of TWB forming, analysis, production and application are presented. To facilitate the presentation of information, the book is split into two parts: Part I, ‘Processing and modeling’ and Part II, ‘Applications’. In Chapter 1, weld integrity concerns are described, test methods (both destructive xi © Woodhead Publishing Limited, 2011

1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

xii 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

Preface

and non-destructive) are presented and prevention and monitoring strategies for weld defects are highlighted. The resulting forming conditions for TWBs are discussed in Chapter 2 with specific details on the effect of weld orientation, properties, etc. In addition, design considerations to address forming concerns are provided. In Chapter 3, mechanics-based modeling of TWB forming is discussed with respect to the allowable thickness/strength ratio of TWB components; the determination of the weld line movement and forming height during the process; and the desired non-uniform binder force ratio to fabricate a TWB component successfully. Then, in Chapter 4 (which concludes Part I), numerical simulation strategies are presented. In particular, methods to address modeling of the weld line in the finite element analyses and the design of the die set and the TWB are discussed. The remaining chapters are in Part II of the book. These include Chapters 5 and 6 on lightweight metal alloy TWBs and advanced high-strength steel TWBs respectively, as well as information in both chapters on dissimilar metal TWBs. In each of these chapters, information related to specific material TWBs, formability, defects, welding methods, etc., are presented. Finally, Chapters 7 and 8 discuss TWBs for the automotive and aerospace industries respectively. While similar challenges exist for the diverse applications, unique means to address these are highlighted. Our hope is that this book will provide knowledge and assistance for the implementation of TWBs for various applications. We would like to thank our colleagues who have contributed their expertise and knowledge for the creation of book chapters. In addition, Ming Shi from US Steel and Rick Vanker from The TWB Company contributed tremendous insights into the content of the book. Finally, we would like to acknowledge the publishing team for their tireless efforts to assure the success of this text. Brad L. Kinsey Durham, NH, USA Xin Wu Detroit, MI, USA


1 Weld integrity of tailor welded blanks M. M. LI, TWB Company, USA

Abstract: The quality of a laser weld is sensitive to its geometrical integrity because of the small fusion zone, especially for those tailor welded blanks (TWBs) that are subject to post-weld forming. Most weld imperfections can be visualized from their appearance; however, not all of them are visible from the surface, such as porosity. Understanding the causes of weld imperfections helps the prevention of their creation. Nonetheless, knowing the characteristics of weld imperfections and their impact on the end application of the product allows the decision to be made regarding which weld imperfections are to be controlled or not. Proper weld monitoring systems and quality processes can be developed to control the quality of TWBs once the above factors, critical imperfections and their causes, are identified. Key words: weld defect, laser welding, laser processing, concavity, pinhole, lack of penetration, lack of fusion, porosity, inclusion, mismatch, sporadic weld, laser triangulation, machine vision camera, camera vision, micro-structure, ultrasonic, eddy current, flux leakage, X-ray.

1.1

Introduction

Tungsten inert gas (TIG), metal inert gas (MIG), electron beam and laser welding processes have been used for creating tailor welded blanks TWBs. However, due to the small heat-affected zone (HAZ) and fusion zone, the laser and electron beam welding processes produce less impact on the material properties than others. Laser welding has been the most frequently used process for producing TWBs due to the lower cost and greater flexibility compared to those of electron beam welding. For this reason, laser welding will be the focus of the material in this chapter. Most weld imperfections are related to the weld geometry or chemical contents of the base materials. Improper fit-up between the two joining metal sheets is a major cause of imperfections related to weld geometry, such as concavity, mismatch and sporadic welds (i.e. a weld that does not have consistent weld geometry and is mostly a mixture of good weld and bad weld in short sections). Impurities or gaseous elements, such as oxygen, nitrogen and hydrogen, embedded in the material or introduced to the weld during the welding process can also create imperfections in the weld, such as porosity, pinholes or craters. The acceptance of a weld imperfection is normally determined by the application of the welded product. A pinhole or crater can be a defect for one product that undergoes a post-weld deep forming process and an accepted weld imperfection if no forming is required. Examples used in this chapter are based on a typical TWB joining two blanks of 1 mm and 2 mm in thickness. 3 © Woodhead Publishing Limited, 2011

1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

4 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


Tests and measurement of weld integrity are divided into two categories, destructive and non-destructive. The most common destructive tests on TWBs are cross-sectioning, micro-hardness test, tensile test, bend test and formability test. These tests are used to understand the mechanical and metallurgical properties as well as the geometrical integrity of the weld. Destructive tests are normally time consuming. In some cases, fatigue tests are performed on welded parts that are subject to applications under cyclic loading. Non-destructive testing is desired for production part inspection, especially in mass production. Most non-destructive testing systems are designed to determine the geometrical integrity of the weld in the fusion zone. These systems are normally vision, ultrasound or electromagnetic based. When a non-destructive testing system is able to perform the inspection faster than the welding speed, such a system can then be used for real-time weld quality monitoring.

1.2

Typical weld imperfections

The most common imperfections seen in TWBs are discussed in this section. When the size of these imperfections becomes significant they are considered as defects for a part that requires a subsequent forming process performed on or around the weld seam. Splits inside or adjacent to the fusion zone can be found after the forming process if the weld imperfection is larger than the critical dimension. It is crucial to identify the imperfections of interest so that efforts can be focused on finding proper solutions to reducing and monitoring these imperfections. The critical dimension of imperfections is normally determined by the material grades and thickness. For example, for a low carbon steel weld with a gauge combination of 1 mm and 2 mm in its base materials, an imperfection with 0.5 mm in its largest linear dimension can be a defect if a post-weld forming process is required. The geometry of the imperfection is also an important factor in determining the critical dimension. Figure 1.1 shows a typical good weld cross-section with the thick gauge material on the left side and thin gauge material on the right side.

1.1 The cross-section of a typical good weld.


Weld integrity of tailor welded blanks

5

1.2.1 Pinhole and crater A pinhole is formed during the welding process when part of the molten material is ejected away from the weld pool due to a burst of excess pressure in the keyhole or molten pool. When the laser beam hits a cluster of low temperature elements embedded in the base material or weld seam, due to the sudden increase of the pressure contributed by quick vaporization of the low temperature material, part of the molten material can be ejected away from the weld pool and forms a temporary cavity in the molten pool. A pinhole or crater is formed as a consequence of insufficient molten material filling into the cavity during the solidification process. The typical size of a pinhole is around 1.0 mm, while most pinholes have their largest dimension ranges from 0.5 mm to 2.0 mm. A crater is sometimes referred to as a partial pinhole or a blind pinhole. Additive materials such as aluminum, magnesium and zinc, which are popularly used in steel processing with vaporization temperatures far lower than that of steel, have a potential to vaporize rapidly and create a burst of pressure in the keyhole of a laser weld and result in a pinhole or crater, which is a partial pinhole. Figure 1.2 shows the top view of a pinhole. While breaking the pinhole along the welding direction, a cross-sectional view of the pinhole is illustrated in Fig. 1.3.

1.2 A pinhole located in a laser weld, tip view.

1.3 Cross-section of a pinhole (Fig. 1.2) along the welding direction.


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

6 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


Frequent cleaning of the welding equipment, especially the contact areas and the vicinity of the welding spot where fume and welding slag accumulate, may reduce the formation of pinholes. In addition, controlling incoming material to avoid unexpected inclusions and maintain weld edge cleanness also helps the reduction of pinhole formation. Most pinholes have round edges and a smooth inside surface. In some cases a small pinhole opens up as a larger hole after the forming process, though in many cases the existence of pinholes does not cause splitting in a forming process. A weld split is expected when the size of a pinhole is significant or deep and the forming strain is high.

1.2.2 No-weld ‘No-weld’ happens when the laser beam energy applied to the material is not high enough to create fusion for a short segment or the entire length of the weld seam. No-weld is normally considered as a weld defect due to the significant reduction of the weld strength. When a short segment of no-weld is formed in a weld seam, the ends of the no-weld serve as sharp cracks during subsequent forming processes and result in a split in the weld. Losing laser power during welding can directly result in no-weld. A laser with stable output power is able to prevent this from happening. Another possible cause of no-weld is that the plasma plume, created by the welding process, is thick and thus blocks the majority of the laser energy. When no-weld happens, no new plasma plume forms; therefore, the laser beam becomes un-blocked. As a result, this type of plasma plume-induced no-weld normally shows in a short segment of several millimeters. Applying proper assisting gas to keep the plasma plume away from the laser beam path is the key to the prevention of plasma plume induced no-welds. In some cases, a thick contamination layer on the surface of the weld seam, such as weld debris, machine oil falling off from the equipment or foreign material carried by the incoming material, can momentarily reduce the total laser energy hitting the weld seam and result in no-weld. Frequent equipment cleaning and incoming material quality control play important roles in preventing this type of no-weld from happening.

1.2.3 Lack of fusion Lack of fusion (LOF) refers to a misaligned fusion zone caused by favoring the laser beam more on one side of the base metal sheets. As a result, the weld seam, regardless of having full penetration or not, is not completely joined together. In most cases, the top of the weld shows complete fusion across the weld seam but the root of the weld does not; see Fig. 1.4(a). For an extreme case, the top and root sides of the seam are fully welded with a small portion of seam in the middle not fused. This type of defect is not visible from the surfaces and can only be visualized by cross-sectioning the weld at the LOF (see zone, Fig. 1.4(b)) or by a non-destructive physical technique.



7

1.4 Lack of fusion: (a) visible from the root of the weld, (b) not visible from the surface of the weld.

1.2.4 Lack of penetration Similar to no-weld, lack of penetration (LOP) is caused by insufficient laser energy applied to the material. The difference between no-weld and LOP is that no-weld has no fusion created during the process and LOP is the result of partial material fusion. A no-weld can be inspected from both sides of the weld and a LOP can only be recognized from the opposite side of the weld where no fusion is found. LOP reduces the formability of the weld significantly and is normally considered as a defect for a weld where forming is one of its subsequent processes. The formation of LOP is related to insufficient heat applied to the material and, thus, full penetration is not achieved. An unstable laser power source is one of the possible causes. In addition, excess amounts of media, such as plasma plume, oil or water contamination, that can block or resist laser light from reaching the material can also reduce the overall energy applied to the material. Normally, porosity is also present when water causes the LOP. Applying proper cross-jet gas, such as inert gas, carbon dioxide or nitrogen, to blow away the plasma plume can prevent the plasma plume from blocking the laser beam. The impact of the plasma plume on the penetration of the laser beam is associated with the wavelength of the laser and the composition of the plasma plume. Figure 1.5 demonstrates lack of penetration on the lower half of the weld with a small gap between both sides of the base material. In some cases, the gap may not exist and a fine seam line will be present where lack of penetration occurs, instead of an open gap as shown in this figure.

1.2.5 Concavity This is one of the most common weld imperfections, especially for those welding processes that do not use filler material. Weld concavity is a result of lack of molten material filling in the fusion zone. Most weld edges are not perfectly


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

8 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


1.5 A weld cross-section with top concavity and lack of penetration.

square or straight. When butting two metal sheets with sheared edges together, finding a gap in the seam is common because most sheared edges are not square. A concavity is expected if the molten thick sheet is not sufficient to fill in the gap. Better edge quality is needed to improve this situation by enhancing the straightness and squareness of the sheet edges. While welding materials with low surface tension, such as aluminum, part of the molten material may drip away from the root of the weld pool during the process, if the addition of gravitational pull to the pressure provided by the shielding gas on the top of the weld is higher than the effect of the surface tension on the molten material. Some welding equipment suppliers may suggest the use of a root-shielding gas to push the molten material back into the seam. A typical top concavity profile can be found in Fig. 1.5. Concavity is not a critical weld imperfection if its existence does not significantly affect the formability of the weld. It is measured as the vertical distance from the surface of the thin sheet to the bottom of the concave surface in the weld. Normally, a concavity with a depth of 10% of the thickness of the thin sheet is acceptable. For parts that do not require subsequent forming or only undergo minor subsequent forming, a higher ratio may be acceptable as long as the weld is stronger than the base material. Li (1999) showed that a 21% total concavity, top and bottom, is the critical value that determines if the formability of the weld is higher than that of the thin material.

1.2.6 Undercutting Undercutting is normally produced by welding high surface tension materials caused by the material properties and welding parameters. Undercut forms at both sides of the weld line adjacent to the HAZs in both parent materials as a shallow and small notch along the weld (see Fig. 1.6). The undercut grooves are weld imperfections and may reduce the formability and strength of the weld due to the stress concentration at the groove and the reduction of the effective thickness.



9

1.6 An illustration of undercuts along the weld edges.

1.2.7 Porosity Porosity is a material discontinuity either totally subsurface or partially exposed to the surface. Most porosity is related to the change of the ability of the base material to absorb gaseous elements such as hydrogen, oxygen and nitrogen (Lancaster, 1980: 110). While molten, the material absorbs more gaseous elements. These elements are released and move toward the top surface to escape from the molten pool during the solidification process. Porosity is caused by gas trapped in the molten pool when there is insufficient time for the gas to escape from the metal. A pore is normally spherical or elliptical and, if fractured, has a smooth internal surface. Figure 1.7 shows multiple pores found in a weld. These small pores are actually shown on the surface of several connected larger pores in the laser weld. In Fig. 1.8, a small pore is associated with a planar inclusion in the thick material. High aluminum and oxygen content was found in this particular inclusion. It is very likely that this inclusion was part of the oxygen-killing aluminum added during the casting process and was trapped inside the material. After rolling, an inclusion will normally be flattened. Porosity was observed while

1.7 A longitudinal cross-section of a laser weld with small pores on the walls of large pores.


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

10 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


1.8 An inclusion-induced pore.

using CO2 as the shielding gas with a welding speed of 0.48 m/min or lower (Linnert, 1994: 575). When lap welding galvanized steel, porosity is normally seen in the weld because of the vaporization of zinc. Such porosity can be reduced by pulsing the laser beam. For example, using a laser with an average power of 1020 watts, pulses of 1660 watts for 2.06 milliseconds at a frequency of 300 Hz allowed travel speed to be reduced to 20 in/min and resulted in less surface porosity than continuous-power laser welding (Oates et al., 1998: 151).

1.2.8 Hard weld The welding process is composed of rapid heating and cooling cycles with melting and solidification happening in a significantly short time. During the heating process, the material in the weld pool is molten and the material in the HAZs is tempered or annealed, depending on the distance from the weld and the amount of cold work applied to the base material prior to welding. The heating process is followed by a cooling process during which solidification happens in the weld pool and continuous cooling happens in the solidified weld pool and the HAZ. The hardness of the material is normally elevated in the fusion zone as the result of a fast cooling process. However, depending on the original material cold work status and chemical composition, some softening effects such as tempering effect and recrystallization may or may not happen in the HAZ. Combining the softening effects with the hardening effect, residual stresses and quenching effect, differences in the HAZ hardness may result. It may become a problem in forming when the weld or HAZ hardness is elevated too much, for example twice the hardness in the weld as that in the base material. Using the hardenability data to predict the hardness in the HAZ and the fusion zone of a material (Avner, 1974: 302) is important to ensure that the welded assembly meets the hardness requirement.



1.3

11

Testing methods

Like many other industries, material testing methods for TWBs are categorized as destructive and non-destructive testing methods. Destructive testing methods are normally used for off-line tests to determine the quality of products by sampling a small fraction of products from the overall production. This type of testing method is normally time consuming and requires destruction of parts, and thus is less employed in a mass-production environment when an on-line testing method is available. However, despite their disadvantages, destructive testing methods can, mostly, provide more detailed information about product quality. As a matter of fact, most non-destructive testing methods are built based on the knowledge and experience accumulated from numerous destructive tests.

1.3.1 Off-line testing, destructive Cup test Cup test is the generic name of the Erichsen test, the Tinius Olsen test or other similar tests. It uses a clamping device to hold the test sample with the weld seam at the center, and a spherical punch is used to push the center of the sample until it breaks. The test is normally stopped at the moment the fracture occurs. The dome height and load are normally recorded at the moment the fracture happens. The fracture location with respect to the weld is also observed to determine if the weld allows higher formability. For a TWB, it is normally expected to fracture outside of the weld. If the fracture is in the HAZ or inside the weld, a weak weld is normally the cause and, often, an excess amount of concavity or undercut can be found. Cup testing a laser weld with severe concavity will normally result in a split in the weld (left picture of Fig. 1.9). This cup test was performed on the same sample shown in Fig. 1.5. For a weld in low- to mid-carbon steel, it indicates a good weld when the split is away from the fusion zone and HAZ (right picture of Fig. 1.9). For materials that have extremely hard or softened weld, the fracture may initiate in the weld during the test. In such cases, the standard test procedure,

1.9 Cup test results with a split in the middle of the weld (left) and a split outside of the weld (right).


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

12 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


measuring the dome height ratio of the weld to the thin base material, is normally suggested to determine the quality of the weld. Tearing Similar to the cup test, a tearing test is used to determine if the weld is more durable than the base material, in most cases a thin gauge material, when undergoing deformation. This process is conducted by clamping on both sides of the weld at the edge and tearing the material in the direction perpendicular to the blank surface until the material breaks. The purpose of this test is to observe the fracture location. When the fracture location is outside the weld or the HAZ, the weld passes the test. Otherwise, a weak weld is found and further analysis of the weld imperfection is needed to reveal the root cause. This method is not suggested for TWBs of aluminum alloys and some advanced high strength steel, such as DP 980 and martensitic steel, where the weld is typically weaker than the base material. Cross-sectioning Cross-sectioning is a straightforward way to determine the geometrical profile and the micro-structure of the weld. With proper preparation, such as grinding, polishing and etching, the micro-structure can be developed for further metallurgical analysis. A magnifying device such as a microscope is required for a close-up view of the cross-section of the weld. In some cases, the information of the weld geometry is needed in a quick fashion. Cross-sectioning the laser weld with rough sanding can provide good information about the weld geometry (see Fig. 1.10) in a short time. This process is short and can be adapted as part of the quality certification in mass production.

1.10 The cross-section of a concave weld, prepared without sanding and polishing for fast inspection.



13

Micro-hardness test The same sample used for micro-structural study of the weld cross-section can also be used for micro-hardness tests. A micro-hardness tester produces indentations in a microscopic scale under a certain load. The dimensions of the indentations indicate the hardness at its location. Normally, a linear array of micro-indentations is created from one base material to the other across the weld and HAZs. The hardness variation across this test line is therefore observed. Vickers hardness and Knoop hardness testing heads are the most popular. The Vickers testing head creates a square indentation, and the Knoop testing head creates a rhombus indentation. In Fig. 1.11, an array of micro-hardness tests show that the hardness increases in both the HAZ and the fusion zone.

1.3.2 On-line testing, non-destructive In a mass production environment, non-destructive testing methods are normally used to inspect the quality of products. There are several testing methods used for monitoring the quality of a TWB. These testing methods are used to determine the dimensions, surface conditions, cross-sectional geometry and material continuity.

1.3.3 Surface inspection Most tools used for surface inspection are to replace human inspections. That is, those weld imperfections that human eyes can recognize have the potential to be inspected by a vision system.

1.11 A sample micro-hardness test across a laser weld with different base material hardness.


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

14 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


Machine vision camera Using a camera to inspect the surface and dimensions of products is becoming a popular method in many industries, especially for those that employ a high degree of process automation. With the cost reduction in computers and cameras, machine vision has become a popular choice for product or process validation. A basic application of a machine vision camera in TWB production is to use it for measuring the dimensions of products and validating the product quality based on the information acquired. Material surface irregularities such as rust or improper coating can also be observed by this technology. Some developers use machine vision cameras to inspect surface imperfections such as craters or roughness which are normally difficult for vision tools. Most vision systems are composed of cameras and a computer running a program to acquire and analyze images from the object of interest. The results of the analysis are normally communicated to a control unit via digital and/or analog outputs or the network. There are smart cameras available for this type of application. A smart camera is, as its name implies, a camera with the ability to process images and interact with other devices through wired or wireless communications for those devices to react based on the inspection results. Lighting, field of view determination, lens and camera selections are critical to the success of acquiring good quality images. Having a good quality image is the basis of a successful machine vision practice. A machine vision camera can be used to inspect imperfections such as product dimensions and surface finish. Weld width for lack of penetration (LOP) Due to its miniature geometry, LOP is difficult to discover. It is only visible from the root side of the weld as a fine straight line with missing molten material. When the length or depth of the LOP is large, tools to inspect material discontinuity such as ultrasound and eddy current devices have a better chance to detect it. However, when the LOP is small, such as 0.5 mm in length or 0.1 mm in depth, it becomes very difficult to observe. Unfortunately, if the welded part is undergoing a subsequent forming process, a small LOP can be a good origin of a split. It is possible to detect a LOP by using a high speed camera dedicated to measuring the root width of the weld by the surface difference between fully penetrated and partially penetrated parts. Laser triangulation The basic idea of laser triangulation is to use a laser line to scan the object in three dimensions and a camera to observe the laser line projected on the object as illustrated in Fig. 1.12. By moving the object of interest against the laser line, a three-dimensional profile of the object can be constructed. The cross-sectional profile of a laser weld can be used to determine its quality, for example concavity



15

1.12 An illustration of laser triangulation with a laser line projected across the weld developing the top cross-sectional profile of the weld at the laser.

and undercut can be observed by this technology. Mismatch of both base materials in the thickness direction can also be easily identified by this method.

1.3.4 Material discontinuity inspection Material discontinuity mentioned in this section is especially for discontinuity in the welding direction, such as pinholes, porosity, lack of penetration and craters. Ultrasound A traditional ultrasound device is used to detect the material discontinuity by transmitting a sound wave into the material from its surface and detecting the timing of returning waves. While a sound wave moves in the metal, it bounces back on an interface that has transmittance change, normally associated with a wave transmitting media change, such as a void or a pinhole in the weld. In Fig. 1.13, a void reflects the sound wave back to the ultrasound transducer located on the top of the sheet material. As a result, the ultrasound receiver detects a signal before the reflected signal from the bottom of the sheet. In order to provide better wave front manipulation, some ultrasound devices employ an array of transmitters to create an array of wave sources to steer the direction and geometry of the wave

1.13 An ultrasonic transducer emitting sound waves into the material with a void reflecting the sound wave.


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

16 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


front by providing a phase-shift time delay in the wave emission, such as phased array ultrasound devices. Generally speaking, an ultrasound device is not sensitive to imperfections that are close to the surface of the weld, top or bottom, due to the fact that the top and bottom surfaces are also the interfaces of material discontinuity and, thus, the ultrasound receiver tends to mix up the signal from the material surfaces with that from a small imperfection near them. This situation is more significant on a rough surface. As a result, an ultrasound device is normally used for parts with large volumes, or thicknesses of 5 mm or higher. Using a coupling media, such as oil or water, is also a disadvantage of the ultrasound device. Nonetheless, the traditional ultrasound device is, in many cases, the least expensive device. It is often considered as the first-line, non-destructive detecting device because of the affordable cost. Electromagnetic acoustic transducer (EMAT) An electromagnetic acoustic transducer (EMAT) is also an ultrasound device but with a different way of generating sound waves and sensing returning waves. Instead of generating a sound wave which is perpendicular to the incidence surface, an EMAT device generates a shear sound wave that is parallel to the surface of the material. When an obstacle, or interface, exists in the path of the sound wave, the sound wave bounces back and can be detected by a receiving device (see Fig. 1.14). By arranging the sensing coils at a few specific locations, an EMAT device is able to detect material discontinuity close to the surfaces of the material. The EMAT

1.14 Illustration of the principle of an EMAT system.



17

device offers better monitoring capabilities for imperfections in a TWB than a traditional ultrasound device because of the fact that most TWBs have base material thicknesses ranging between 0.6 mm and 3.0 mm. The sensing coil of an EMAT device does not have to contact the material. A fixed distance between the sensing coil and the material surface is required. As a result, a thin layer of metal acting as a spacer and a wear surface, which contacts the inspected material, is employed to maintain a constant coil-to-surface distance. This wear surface is considered as a consumable. Eddy current An eddy current testing device uses a current-carrying coil, i.e. a primary coil, placed in proximity to the test sample, which is electrically conductive. By alternating the current in the coil, a changing magnetic field is created. This changing magnetic field interacts with the test sample and generates eddy currents in the test sample. Variations in the phase and magnitude of these eddy currents indicate possible material discontinuity on or under the surface of the test sample. These eddy currents can be measured by using a secondary sensing coil or detecting the current changes in the primary coil. An eddy current device is sensitive to imperfections on or near the surface. Its sensitivity to an imperfection drops when the imperfection is located away from the surface. Magnetic flux leakage detector (MFL) When applying a magnetic field through the material along its thickness, most of the magnetic flux flows within the material. When a material discontinuity is found in the test sample, a flux leakage can be found by the discontinuity. The basic principle of a magnetic flux leakage (MFL) detector is to use a Hall-effect sensor to detect the leakage while applying a uniform field of magnetic flux across the weld (see Fig. 1.15). Similar to the eddy current and ultrasound devices, an MFL detector does not require direct contact between the sensor and the test

1.15 A magnetic flux leakage sensor detecting a defect in the material.


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

18 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


sample. However, due to the need to maintain a fixed distance between the sensor and the test sample surface, a non-magnetic wear surface is desired to maintain the distance and protect the sensor. The MFL detector is able to detect flux leakages caused by a discontinuity within a few millimeters from the surface. This depth of penetration is sufficient for most TWBs. The major drawback of this technology is that it is only effective on ferromagnetic materials. Besides, the magnetic field used to apply a uniform magnetic field through the test sample creates an attracting force and may become a problem while moving the test sample against the sensor, or vice versa. By scanning the Hall-effect sensor on a ferromagnetic sample, the sensor is able to detect a defect as small as 0.34 mm in diameter (O’Connor et al., 2002). X-ray Similar to other materials, a negative X-ray image of a metal sample shows a light area for the portion of the part that has less mass and a dark area for heavier mass. When a port, pinhole or crater exists in the material, a sudden change in the image grayscale intensity is observed. This sudden change can be processed using image processing tools to identify its geometrical information, especially its size, and used for quality decision purposes. Using an X-ray device to inspect the interior quality of metal parts has been used in industries for products that are sensitive to internal imperfections, voids, cracks or internal holes. Figure 1.16 shows a defect in a laser weld, located at the center of the image horizontally. Being a positive image, any material discontinuity or lack of material shows up as a darker area. Thus, the lower part of the picture is the thick material and the upper part the thin material. Most of these applications are performed at an X-ray job shop as an off-line process. In the food industry, it is becoming popular to use the X-ray device to

1.16 An X-ray image showing a defect located in the laser weld.



19

inspect food quality while the objects are being transported on the conveyors. The X-ray intensity required for inspecting metal is a lot higher than that for fruits and meat. Thus, the difficulty of inspecting metal parts as a real time process using an X-ray device is higher than that for the food industry. Micro-focus X-ray can be a good candidate for this application due to the fact that the X-ray is highly focused in a line for good penetration even though the total intensity is less. By placing a line scanning camera under the part, an X-ray image can be constructed. Light emission Light emission detection is one of the earliest technologies used to inspect the quality of laser welds. Photodiode sensors coupling with designated filters allow the researcher to select the wavelength of light in a specific range for the inspection (Miyamoto, 1995).

1.4

Quality control in production

Obviously, the best way to ensure good product quality is to produce as many good products as possible. Therefore, the most effort involved in the continuous improvement of product quality should be applied to producing fewer defects. However, it is almost impossible to produce zero defects in TWBs. Quality monitoring systems are, as a result, normally employed to detect product imperfections, especially for an automated production process.

1.4.1 Prevention of weld imperfections Equipment cleaning Equipment cleaning is a critical contribution to the reduction of pinhole formation by preventing weld debris from entering the seam. Depending on the processed material and its coating, the frequency of cleaning varies significantly and can be arranged from once a few hundred to a few hundred thousand meters of weld. Optics also need to be examined on a regular basis and cleaned if necessary. Laser quality When referring to laser quality, the temporal and spatial energy distributions are mostly considered. The temporal laser quality is the stability of a laser to hold its energy at a special level for a certain period of time. This also includes the time required by the laser to rise to its set power. The spatial laser quality is normally represented by plotting the energy distribution across the laser beam before the focusing optic. A Gaussian distribution is expected to be the perfect laser energy distribution. However, a laser with a perfect Gaussian distribution is not necessarily


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

20 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


suitable for the butt welding process due to the high energy concentration at the center of the focus spot. This results in a smaller effective focal spot and, thus, less gap tolerance. Such lasers normally have good performance in cutting, drilling and lap welding. On the contrary, a flat-top energy distribution is preferred for butt welding for the larger effective focal spot, which has 86% of its total power (Steen, 1998: 81). Weld edge preparation The quality of a butt weld is sensitive to its weld seam fit-up, which is significantly affected by the weld edge quality. Edge squareness and straightness of the edge are the main factors that affect the fit-up of the seam. Assuming that the total amount of molten material per unit length of a weld seam remains the same after welding, i.e. no material loss during the welding process, a gap existing in the weld seam may result in a concavity if the volume of the void is more than the volume of the material melted down from the portion of the thick material located above the thin material. For a larger gap, such as 0.2 mm or higher, the laser focus may fall into or through the gap. As a result, an insufficient molten material is expected and, normally, a concavity or partially joined seam may be found. Filler material Adding a filler material makes the process more tolerant to poor fit-up. The welding process is, thus, easier to control while the edge quality of the material is less than perfect. However, a wire feeder and proper filler materials are needed for products of various types of materials. This also implies that additional cost and equipment management are involved. From a mass production point of view, using a filler material is not necessarily a better solution to making better TWBs. The manufacturer needs to choose between maintaining better edge quality of incoming blanks and maintaining proper filter material feeding to the weld, such as material type, feed rate and so on.

1.4.2 Monitoring of weld imperfections Some weld imperfections are equipment or material dependent. It is important to understand the types of imperfections associated with the equipment, material and process setup used in the manufacturing process before selecting quality monitoring systems. Once the weld imperfections of interest are selected, quality monitoring systems can then be determined. In order to successfully select proper systems, the following information is needed:

• •

production performance data on the weld imperfections of interest quality monitoring system effectiveness against imperfections.



21

In many situations, a TWB manufacturer will end up choosing more than one device for their product quality monitoring due to the fact that each device has its specialties and none of the mentioned devices can detect all types of imperfections effectively. A technology combining multiple sensor systems, including cameras, with each system specialized in detecting one or some weld imperfections, has been well adopted by the industries and is sometimes referred to as ‘Sensor Fusion’ (Sun et al., 1999). A sample production performance matrix is shown in Table 1.1. Once the production performance is identified, the effectiveness of the quality monitoring systems is needed to complete the selection. Table 1.2 shows sample quality system performance. Select light emission, laser triangulation, eddy current and machine vision for defect inspection. Table 3.1 shows the result of undefected weld imperfections for each system in parts per million (ppm) by multiplying the values in Tables 1.1 and 1.2 for the selected systems. The best performing system for inspecting each weld imperfection shows as the lowest number in the corresponding column of Table1.3. The resultant value Table 1.1 Sample production performance matrix (showing defect occurrence as a percentage) Mismatch Lack of Lack of Concavity Pinhole, Internal Dimensions fusion penetration crater pore 0.05

0.01

0.01

0.10

0.50

0.01

0.20

Table 1.2 Sample quality system error table (with percentages that a system will miss various defects) Mismatch Lack of Lack of Concavity Pinhole/ Internal Dimensions fusion penetration crater pore Light barrier

100.00

100.00

100.00

100.00

0.20

100.00

100.00

Audible sound

100.00

100.00

100.00

100.00

0.30

100.00

100.00

Light emission

0.50

0.40

0.40

0.30

0.30

0.50

100.00

Laser triangulation

0.10

100.00

100.00

0.10

0.30

100.00

0.50

Ultrasound

0.50

0.50

0.50

0.50

0.20

0.30

100.00

Eddy current

0.50

0.50

0.50

0.50

0.10

0.30

100.00

Flux leakage

0.40

0.30

0.30

0.40

0.10

0.20

100.00

EMAT

0.40

0.30

0.30

0.40

0.10

0.20

100.00

X-ray

100.00

0.50

0.50

0.40

0.10

0.10

100.00

Machine vision

100.00

100.00

100.00

100.00

0.20

100.00

0.10


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

22 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


Table 1.3 Result table, showing defects not detected in ppm (parts per million) Mismatch Lack of Lack of Concavity Pinhole/ Internal Dimensions fusion penetration crater pore Light emission

2.5

0.4

0.4

3.0

15.0

0.5

2000.0

Laser triangulation

0.5

100.0

100.0

1.0

15.0

100.0

10.0

Eddy current Machine vision Result

2.5

0.5

0.5

5.0

5.0

0.3

2000.0

500.0

100.0

100.0

1000.0

10.0

100.0

2.0

0.5

0.4

0.4

1.0

5.0

0.3

2.0

of such a system is placed in the bottom row as the result of the combined performance of the selected systems on such weld imperfection. Adding the individual results together gives the final performance as 9.6 ppm. This final performance is achieved when all systems perform as expected and each part has only one type of weld imperfection. As mentioned previously, not all weld imperfections are considered as defects to a TWB. Whether the imperfection is a defect or not is determined by the subsequent processes the weld is going to experience and the end application of the product.

1.5

Conclusions

The integrity of a laser weld is significantly affected by the fit-up of the seam, base material quality, laser quality, shielding gas, weld seam contamination and laser alignment. Various weld imperfections are found to be associated with the abovementioned factors. Understanding the types of imperfections that are unacceptable to a certain process is the first step of TWB quality management. Secondly, properly selecting quality monitoring systems enables the ability to contain defective products and prevent them from reaching the customers. Learning the root causes of these imperfections helps a manufacturer develop procedures and equipment to reduce the production of defective products. This is the ultimate goal for almost all industries.

1.6

References

Avner, S. H. (1974), Introduction to Physical Metallurgy, McGraw-Hill, Inc. Lancaster, J. F. (1980), Metallurgy of Welding, George Allen & Unwin. Li, M. (1999), ‘Real Time Weld Quality Monitoring of Laser Welded Blanks – Weld Profile Monitoring’, Proc. Int. Congress on Applications of Lasers and Electro-optics (ICALEO) 1999, 87, 2E:1–8. Linnert, G. E. (1994), Welding Metallurgy – Volume 1: Fundamentals, 4th edn, American Welding Society.



23

Miyamoto, I and Mori, K. (1995), ‘Development of in-process monitoring system for laser welding’, Proc. Int. Congress on Applications of Lasers and Electro-optics (ICALEO) 1995, 80: 759–67 Oates, W. R. and Saitta, A. M. (1998), Welding Handbook – Volume 4: Materials and Applications, Part 2, 8th edn, American Welding Society. O’Connor, S., Clapham, L. and Wild, P. (2002), Measurement Science and Technology, 13: 157–202. Steen, W. M. (1998), Laser Material Processing, 2nd edn, Springer. Sun, A., Kannatey-Asibu, E. Jr., & Gartner, M (1999), ‘Sensor systems for real-time monitoring of laser weld quality’, Journal of Laser Applications, 11: 153–68.


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

2 Deformation of tailor welded blanks during forming K. NARASIMHAN, IIT Bombay, India and R. G. NARAYANAN, IIT Guwahati, India

Abstract: The reliability of tailor welded blanks (TWBs) in industrial applications depends on many factors affecting TWB formation. The impacts of base metal thickness ratios, strength ratios and welding conditions are explored. The forming behavior of TWBs can be simulated either by incorporating the properties of the weld zone or by assuming the weld region as a weld line without considering the properties of constituent base metals. Weld line movement is a common phenomenon during the drawing of TWBs and can lead to significant reduction in formability. Weld zone properties can be evaluated by tensile testing of subsize samples using the rule of mixtures technique feeding into finite element simulation models for accurate forming behavior predictions. Key words: tailor welded blanks, rule of mixtures, finite element method, forming limit curve, weld region, weld zone, drawbead, blank holder force.

2.1

Introduction

The deformation and forming behaviors of tailor welded blanks (TWBs) are affected in a synergetic manner by many factors. Important factors include base metal properties, thickness/strength ratio, weld process and weld orientation. Prediction of forming behavior of TWBs can significantly reduce the lead time for the manufacturing of sheet metal components using TWBs. The accuracy of such predictions depends greatly on proper estimation of the properties of the weld, in addition to the base metal properties. The following topics will be addressed in this chapter. The estimation of the constitutive behavior of the weld and the effect of the weld parameters on the tensile behavior of TWBs will be detailed. A section will cover factors affecting the forming limit strains of TWBs, with emphasis on the effect of the various TWB parameters. Weld line movement is a common phenomenon observed during the forming of steel TWBs. The factors affecting weld line movement and its control will be discussed. Finally, design issues pertaining to the forming of TWBs will also be discussed and the chapter will conclude with a section on the simulation issues of TWB forming.

24 © Woodhead Publishing Limited, 2011


2.2

25

Estimation of the constitutive behavior of the weld region

TWBs are typically manufactured by laser welding processes, which minimize the weld-affected zone (i.e. weld pool and heat-affected zone HAZ), usually in the order of a few millimeters. In TWBs of steel sheets, the strength and hardness of the weld is significantly higher than the base metal properties. Correspondingly, the work hardening exponent of the weld zone is much lower than that of the base steel sheets. In order to better understand and simulate the forming behavior of TWBs, it is essential to have an accurate description of the constitutive behavior of the weld zone in the TWBs. In this section, the procedure based on rule of mixture is discussed for estimating the weld zone properties. The constitutive behavior of the weld region is important from both the formability and simulation points of view in TWBs. Evaluating the constitutive behavior of the weld region is not a new issue, as it was performed even before the advent of tailor welded blanks (TWBs). But most applications use arc welding processes where the weld region is in the order of 15 mm, so that the weld properties can be obtained quite easily. In the case of TWBs, since laser welding is widely used for welding, obtaining the properties of a smaller weld zone (1–2 mm) accurately is difficult and cumbersome. Longitudinal welded blank Few published works1–6 deal with the methodology to evaluate the weld zone properties. In Abdullah et al.,2 the rule of mixtures (ROM) technique with TWBs having a longitudinal weld (see Fig. 2.1) is used to evaluate the stress–strain behavior of the TWB and weld region. The stress–strain relationship of base metals is first obtained by tensile testing and fitting an appropriate hardening law like the Hollomon equation giving,

σ 1 = K1 ε 1n1 and σ 2 = K2ε 2n2

[2.1]

where ‘K’ and ‘n’ refer to the strength coefficient and strain hardening exponent of the two base metals, 1 and 2 respectively. According to ROM, the total load, P,

2.1 Schematic of load sharing by base metals and weld region in longitudinal TWBs.


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

26 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


applied is distributed on the three areas – namely the base metal 1 (P1), base metal 2 (P2) and the weld zone (Pw). In the tensile testing of the welded specimen, the total load ‘P’ on the sample is represented as, P = P1 + Pw + P2 = σ 1 A1 + σw Aw + σ2 A2

[2.2]

where ‘A’ refers to the cross sectional area, ‘w’ refers to the weld region and subscripts 1 and 2 refer to base metal 1 and base metal 2 respectively. Substituting σ1 and σ 2 from equation [2.1] in [2.2] gives, [2.3] where σ–w represents the average stress in the weld region, as the hardening behavior of weld may vary along its length. Longitudinal strain is assumed to be constant in all of the regions during deformation and hence,

ε 1 = ε 2 = εw

[2.4]

Substituting equation [2.4] in [2.3] results in the relationship [2.5] Equation [2.5] defines the stress–strain relation in the weld region. Here, with the exception of σ–w and εw, all other values can be obtained using subsize tensile tests on the base metal. Area of the weld region ‘Aw’ can be calculated by knowing the area of cross-sections of base metals 1 and 2 or from weld microstructure. Measuring the area of the weld region is critical to the successful implementation of this method. In Ghoo et al.,3 the same method was followed to establish an understanding of the effect of different tensile specimen sizes and offset weld position on the stress–strain behavior of weld region. It was found that tensile specimen size shows little effect on the stress–strain behavior of TWBs. Also, the offset position of weld in the longitudinal TWB showed negligible effect on the ‘K’ value obtained because the same total load absorbed by the TWB is independent of the offset values. In Auger et al.,7 the same methodology of ROM is used to determine the weld stress–strain behavior.

2.3

Methods to evaluate the weld width (or crosssectional area) in tailor welded blanks (TWBs)

Evaluating the constitutive behavior of TWBs by ROM depends on the accuracy of measuring the cross-sectional area of the weld region. The following are the few methods available to perform this:



27

1. The cross-sectional area of the weld region can be obtained by multiplying the thickness of the weld region by the weld width. Here, weld width can be obtained by subtracting the widths of base metals (base metal 1 and 2) from the total sample width. This method is subjected to approximations in measuring dimensions. 2. Another reliable method is by measuring micro-hardness, which is found in many published papers.1–10 The weld region exhibits different hardness when compared to that of the base metal. One can perform hardness measurements perpendicular to the weld region. The sudden hardness change can yield weld width and hence the weld cross-sectional area can be obtained. This is also an approximate method as the weld thickness is assumed to be the same as the average base metal thickness (say 1 mm), without considering the actual shape of the weld region. 3. An accurate way to obtain weld cross-sectional area is through a micrograph. An optical micrograph with an image analyzer can yield accurate measurement of the exact shape of the weld region and hence the cross-sectional area of the weld zone. This method is followed in Bhagwan et al.11 to study the effect of weld properties and geometry in numerically predicting the forming behavior of Al TWBs.

2.3.1 Application of rule of mixtures (ROM) As described in the methodology, chosen weld properties are evaluated from the tensile testing of welded blanks and base material. Subsize 3 type samples made from welded blanks under desired or optimized welding conditions (welding power = 3.5 kW, welding speed = 5.5 m/min, 300 mm focal length mirror and 360 microns focal diameter in donut mode) were tested for this purpose. In this study, the TWB is produced by CO2 laser welding of the same steel sheet (interstitial free steel (IFS)) and therefore both the base metals 1 and 2 will have identical tensile behaviors. Figure 2.2 shows the engineering stress–strain behavior of base material, welded blank (with longitudinal weld) and weld region only. Here, weld constitutive behavior is obtained from the ROM method. Figure 2.3 shows the true stress–strain characteristics of the same data. Table 2.1 compares some of the important mechanical properties of base metal, welded blank and weld region. It is clear from Fig. 2.2 and Table 2.1 that the base metal has higher ductility and lower strength when compared to the welded blank. The welded blank contains a harder weld zone, as a result of which it requires more load to deform at any point of progression compared to the base metal. The presence of a heterogeneous weld zone in TWB reduces the ductility when compared to that of base material. It is also observed that a peak hardness of about 450 VHN develops in the weld region, as compared to about 140 VHN in the base metal area. Table 2.1 also shows that the welded blank strength coefficient, ‘K’, is


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

28 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


2.2 Engineering stress–strain behavior of base metal, welded blank and weld region (subsize 3 type sample).

2.3 True stress–strain behavior of base metal, welded blank and weld region (subsize 3 type sample).



29

Table 2.1 Comparison of chosen mechanical properties of base metal, welded blank and weld zone Sample

Properties

Base metal

σys (MPa) UTS (MPa) eu (%)

n

135

300

38

0.31

Welded blank – longitudinal weld 250

365

25

0.2

Weld region only (from ROM)

800

700

6 (app.) 0.06

K (MPa) 592 611 1015

approximately twice that of the base metal. The weld region, as expected, is stronger and less ductile compared to the welded blank and the base metal. The same pattern is also reflected in the hardness distribution, with the weld zone showing higher peak hardness than the base metal. The ‘K’ and ‘n’ values shown in Table 2.1 are evaluated from the ln (true stress)–ln (true strain) plot of data shown in Fig. 2.3, as followed in the case of standard tensile testing. The weld zone, being the least ductile and the strongest, results in very low ‘n’ value of 0.06 and very high ‘K’ value of 1015 MPa. Table 2.1 also gives ‘K’ and ‘n’ values of the base metal and the welded blank. As described in the methodology, the evaluated weld properties are validated for their accuracy by predicting the tensile stress–strain behavior of welded blanks with two different weld orientations, viz. 30° and 45° to the loading direction, by finite element method (FEM) simulations. The predicted tensile behavior is compared to experiments. Figure 2.4 shows the tensile behavior of welded blanks with 30° and 45° weld orientations. It is clear that the predicted and experimental stress–strain curves are comparable to each other in both the weld orientations. This demonstrates that the tensile weld properties obtained from subsize tensile sample following the iso-strain principle can be incorporated into an FE model of welded blanks to predict their forming behavior.

2.3.2 Tensile behavior: influence of weld orientation American Society for Testing and Materials (ASTM) standard samples were used for tensile testing as explained in the previous methodology. The longitudinal and transverse welded blanks were cut (by electrical discharge machining) from parent TWBs made from laser welding with the desired or optimum welding conditions (power = 3.5 kW and speed = 5.5 m/min). The tensile behavior comparison between TWBs with varied weld orientations and an un-welded blank is shown in Fig. 2.5. Table 2.2 gives the same information. It is clear from Table 2.2 and Fig. 2.5 that welded blanks with transverse welds


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

30 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


2.4 Experimental and predicted engineering stress–strain behavior of TWBs with 30° and 45° weld orientations.

2.5 Experimental and predicted tensile behavior of TWBs having longitudinal and transverse welds, and an un-welded blank (failure patterns are shown in insets; F = failure location).



31

Table 2.2 Comparison of tensile properties of TWBs (from experiments) having longitudinal and transverse welds, with un-welded blank Properties Sample

σys (MPa) UTS (MPa) eu (%)

Base metal

133

Transverse welded blank

135

299

38

Longitudinal welded blank 250

365

25

298

40

and un-welded IFS base metals exhibit the same tensile behavior. Yield strength (133, 135 MPa), ultimate tensile strength (UTS) (298, 299 MPa) and uniform elongation (40, 38%) are practically the same. The reason is because the failure location is the same in both the cases, i.e. failure occurs in base material only. In the case of transverse welded blanks, since the weld is stronger and samples were fabricated with the optimized welding conditions, weld failure is avoided. Hence, failure occurs far away from the weld region in the base metal, such that the weld does not participate in the plastic deformation at all. On the other hand, longitudinal welded blanks display entirely different tensile behavior. Ductility of 25% is seen in this case, showing reduced formability when compared to that of transverse welded and un-welded blanks. Moreover, the strength properties, yield strength (250 MPa) and UTS (365 MPa), have increased tremendously. The increase in strength properties and ductility reduction compared to the un-welded and transverse welded blank is due to the presence of the stronger (hence requiring more load to deform) and less ductile (hence an overall reduction in the TWB elongation) weld zone in the TWB. In summary, longitudinal welded blanks possess reduced ductility when compared to that of un-welded and transverse welded blanks, hence lower formability is expected in this case. The simulation shows similar trends as the experimental data shown in Fig. 2.5. Similar conclusions are seen in Cheng et al.12 where TWBs from an Al 5754-O alloy were studied for their formability.

2.3.3 Tensile behavior: other influencing factors In the previous section, the effect of weld orientation on tensile behavior was described. In addition, the thickness ratio, strength ratio, hardening ratio and weld strength also affect the tensile behavior. In general, it is observed that for transverse welded TWBs the region which has the lowest load bearing capacity determines the overall deformation behavior. As an example, for steel


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

32 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


TWBs, the base metal that has lower strength and/or thickness (effectively lower load bearing capacity) will control the overall deformation behavior. The stronger base metal and the weld simply act as a rigid body during deformation. The difference observed in the tensile behavior of such TWBs can be rationalized as simply the gauge effect of the test sample. If the TWB weld has lower strength and hardness compared to the corresponding base metals, then the entire deformation tends to concentrate and localize in the weld region with minimal deformation in the base metal. This scenario leads to a significant drop in the overall ductility and strength of the TWB compared to the corresponding base metal.

2.4

Forming limits of TWBs: influence of weld orientation

Similar to tensile behavior, formability of TWBs was studied by limiting dome height (LDH) tests. Figure 2.6 compares the influence of weld orientation on the forming limit curve (FLC) of welded blanks with that of un-welded blanks. It is clear that both transverse and longitudinal welded blanks show lower FLC than the un-welded blank FLC, showing reduced formability. But for transverse welded blanks, limit strains are more or less the same (< 3% variation) in all strain paths except in the near plane strain condition. This 3% variation is within the experimental error bar (more than 5% is considerable) and hence can be

2.6 Comparison of experimental FLC of TWB having longitudinal and transverse welds, with an un-welded blank (IF steel sheet).



33

seen as a negligible effect. This is consistent with the observations in tensile testing. When transverse and longitudinal welded blanks are compared, unlike tensile behavior where longitudinal welded blanks show reduced ductility, FLCs of TWBs are practically the same irrespective of weld orientation. Cheng et al.4 showed that longitudinal welded blanks (Al 5754-O) show a considerably reduced FLC when compared to that of un-welded blanks. Longitudinal welded blanks exhibit less ductility and formability when compared to transverse TWBs. But from the FLC point of view, both transverse and longitudinal welded blanks have practically the same limit strains (< 5 % variation).

2.4.1 Experimental forming limits of steel TWBs The forming limits of two steel TWBs were studied in detail. SPCC 440, IF and DP590 steel sheets were used in this work. The basic mechanical properties of these three steel sheets are presented in Table 2.3. The experimental FLCs of these three base steel sheets are given in Fig. 2.7. IF–DP and IF–SPCC TWB combinations were undertaken for further studies. Figures 2.8 and 2.9 compare the experimental FLC levels of the base metals with those of the TWBs for DP 590–IF and SPC 440–IF combinations respectively. It can be observed that TWBs have FLC levels in-between those of the base metals for the drawing region of the FLC. For the stretching region, the TWB shows lower formability than the corresponding base metals. Most of the literature suggests that the FLC of a steel TWB is typically in-between that of the corresponding base steel sheets. Our results show a similar trend in the drawing region but in the stretching zone, the FLCs of the TWBs are marginally

Table 2.3 Mechanical properties of IF, DP and SPCC steel sheets IF

DP 590

SPCC 440

Initial yield strength, σys (MPa)

137

415

340

Ultimate tensile strength (MPa)

301

619

444

40

22

20

Uniform elongation, euniform Strain hardening index, n Strength coefficient, K (MPa)

0.3 589

0.20 1052.3

0.17 696

Anisotropy r0

1.68

0.91

1.6

r45

1.08

0.73

1.22

r90

1.84

1.21

1.02

R-bar

1.26

0.84

1.47

Thickness (mm)

1.0

1.6

1.6


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

34 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


2.7 The experimental FLCs level of three base steel sheets.

2.8 Comparing the experimental FLCs of TWBs of DP–IF steels with the base metal.

lower than the base steel sheet FLCs. It is felt that the weld quality could be responsible for this observation. In fact, the simulation shows that the predicted TWB FLCs lie in between those of the corresponding FLCs of the base steel sheets.



35

2.9 Comparing the experimental FLCs of TWBs of SPC-IF steels with the base metal.

2.4.2 Other factors affecting forming limits of TWBs The factors that affect the tensile and general deformation behavior also affect the forming limits of TWBs. The thickness ratio of the sheets welded, the ratio of base sheet metal mechanical properties (e.g. the yield strength, work hardening exponent, etc.) and weld characteristics (e.g. strength, ductility, hardening, size and orientation) have significant effect on the formability and FLC of the TWBs. In most conditions, if the weld is sound, narrow and stronger than the base sheets welded (as in the case of laser welded steel sheets), it does not impact significantly on the forming behavior of the TWBs. The exception to this could be a situation where the weld is located in a critical region of a component where peak strains may develop during forming. In such cases, it may lead to a drop in the forming limit strains due to the weld tending to assist the development of the strain localization process. In case the weld tends to be softer (much lower strength than the base sheets), then the forming limit strains will drop significantly as compared to the base sheets due to the preferred strain localization in the weak, soft weld region. Assuming a sound, harder weld, then only the other factors listed earlier will have an influence on the formability of the TWB. The effect is such that in most situations the limit strains of the TWB usually lie in-between the corresponding base sheet metal FLCs. As an example, the limit strain of TWB of sheets of different thickness will lie between that of the two sheets welded.


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

36


2.5

Weld line movement

Weld line movement is a common phenomenon during the deformation of TWBs. This movement can impact the formability of TWBs. The main origins of such weld line movement lie in the differential draw characteristics of the base sheets that are welded. The imbalance in the draw resistance of the base sheet metal leads to different levels of deformation and draw of the corresponding base metals leading to the movement of the weld line. Another way to understand the phenomenon is that the weld line movement is due to the heterogeneity in plastic deformation achieved in the two base metals during forming. The weld line movement possesses a great challenge as it can lead to reduced formability and tool design issues. Therefore, it is essential to understand and control the weld line movement during forming of TWBs. Figure 2.10 illustrates the typical weld line movement during the deep drawing of a square cup. The weld line movement is typically represented in terms of its displacement with respect to its initial location in the un-deformed sheets, as shown in Fig. 2.11. This figure shows the extent of weld line movement for a square cup based on FEM predictions. The major factors that affect the weld line movement are the thickness ratio, anisotropy properties and mechanical properties of the base metals welded. In addition, the initial weld location also impacts the movement. It is observed that as the thickness ratio increases the weld line movement increases, since the thinner blanks deform more. Hence the weld zone moves towards the thicker blank in the cup wall region. In a simple symmetric cup drawing process, as the initial weld line is farther from the symmetry line of the cup, the extent of the weld line movement increases (Fig. 2.11).

2.10 Illustration of the weld line movement during square cup deep drawing.



37

2.11 Predicted weld line movement for a square cup for different initial weld locations in the un-deformed TWB.

2.5.1 Control of weld line movement As pointed out earlier, weld line movement is a product of the heterogeneity in plastic deformation attained by the thinner or weaker material and thicker or stronger material during forming. This weld line movement is mostly seen during deep drawing of cups, and to some extent in stretching processes, if it is placed at some offset. In stretching, one can visualize the ‘weld line movement’ as a ‘bowing of the weld’. Many factors show compounding effects on the weld line movement, as discussed in previous sections, and the deep drawing behavior deteriorates because of this. Weld line movement should be controlled and minimized during the forming of TWBs. Excess weld line movement not only leads to a drop in formability, but also creates problems with the accommodation in the stepped tool. Weld line movement can be controlled in several ways. One of the best practices would be to use variable blank holder pressure at different locations of the TWBs to have a local variation in the resistance to draw and essentially to compensate for the inherent variation in the steel sheet’s draw resistance. In today’s scenario, the advanced forming presses provide servo hydraulic control with multiple cushions to precisely vary the clamping force locally in a pre-programmed manner to achieve any desired variation in clamping force, both temporally and spatially. Ahmetoglu et al.13 have used this approach to control the weld line movement during the deep drawing of a cylindrical cup from a TWB


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

38 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


made of AKDQ steel of 0.8 mm and 1.8 mm thickness. Blank holder force of 10 tons is applied on the thinner side and 1 ton is applied on the thicker side which resulted in a deep drawn cup with the weld line at the center, the same location in which it was placed before forming. Another approach in controlling the weld line movement is to resist the movement of the thinner (or weaker) blank relative to that of the thicker (or stronger) blank by providing drawbeads. This results in equivalent plastic deformation and draw-in levels on both thinner and thicker blanks, and hence weld line movement is reduced. For example, Heo et al.14 show that drawbeads of different dimensions (drawbead radius and height) are installed in the blank holder and restraining forces are applied on the thinner blank side of TWB and as a consequence weld line shift is minimized (Fig. 2.12). Also, by increasing the radius and height of the drawbead the weld line movement is minimized (Fig. 2.12). It is seen from the figures that weld line shift is reduced from 4 mm (without a drawbead) to 2 mm (R = 3 mm) and 0.5 mm (R = 5 mm), by providing a drawbead on the thinner blank. By increasing the drawbead height from 5 mm to 7 mm, the weld line shift is minimized from 1 mm to 0.2 mm approximately. This approach is useful, as the drawbead location and geometry for a given TWB combination and a component to be manufactured can be optimized. In case the component shape changes, or the material combination changes, it may necessitate the reworking of the drawbead to optimize for the changed conditions. This approach also does not require sophisticated metal forming equipment capable of applying variable blank holder pressure. An improved method to control weld line movement is by providing specially designed local adaptive controllers.15–19 This method involves restricting weld line movement by subjecting nearby areas of weld region to local hydraulic pressure or clamping force. Hydraulic pressure (or clamping force) can be applied at one or two locations or along the entire weld line by designing a multi-local hydraulic pressure system. As deformation proceeds, the local adaptive pressure system adjusts pressure accordingly, resulting in control of weld line movement.

2.12 Effect of presence of drawbead and drawbead dimensions (R, H) on the weld line shift.14



2.6

39

Design considerations for TWB forming

TWBs typically consist of sheets with differing strengths and/or thicknesses. TWBs having sheets of different thickness pose a challenge while designing the tools to form them. Tools made for conventional single sheets cannot be directly used for forming TWBs, as the thickness difference in the sheet cannot be accommodated. If the TWB is made of dissimilar thickness blanks, uniform blank holder pressure cannot be applied on the blank, as uniform contact between the thinner blank and the tools is absent. When dissimilar thickness blanks are used, the thickness difference should be compensated, otherwise wrinkling of the thinner metal, tearing near the weld region and/or weld line movement will occur, which will deteriorate the formability of the welded blanks. This problem will not arise if same thickness blanks are used, as uniform contact is established. Compensating the thickness difference can be achieved in several ways – (1) a shim can be provided, and a shim is nothing but a piece of sheet inserted below or above the thinner blank as shown in Fig. 2.13. (2) Another method, which is widely used in industries, is providing a stepped die or blank holder (or lower, upper binder). Here, the die or blank holder is segmented into different regions, which has a step near the location where the weld zone is in contact with the die or blank holder. The following are the problems that should be addressed while designing the stepped tools for the forming of TWBs.

• •

Sheet and tool tolerance should be considered, otherwise non-uniform holding pressure may lead to compressive stresses near the weld region, resulting in failure of weld zone during forming. It should be ensured that the lower or upper tool is ‘slackened’ enough so that stresses can be avoided near the weld region, allowing the weld to move freely (Fig. 14).20

The location of the weld in a TWB is an important design decision to be made while designing TWB-based components. One aspect that is obviously considered is to ensure that the differential property requirement (load bearing capacity etc.) of the component is addressed. The selection of the weld location goes beyond the formability requirement and should address the final performance of the component in an assembled structure, as in a car. Computer-aided engineering

2.13. Schematic of providing a shim for thickness compensation in TWBs.


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

40 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


2.14 Schematic of stepped binder which is slackened near the weld region.20

analysis of the assembled structure (consisting of the TWB component) should be carried out to ensure that the weld location is acceptable.

2.7

Simulation of TWB forming behavior

Finite element (FE) simulation is a widely used numerical technique to analyze manufacturing processes like forming, machining, etc.21–23 FE simulation of the forming of TWBs is a critical issue because of the presence of the weld region, which has relatively different mechanical properties to the base metals. The forming behavior of TWBs can be simulated by two methods – either by incorporating the properties of the weld zone (σys, K, n, r, etc.) separately, like base metals, or by assuming the weld region as a weld line without considering properties such that the weld line is just a boundary between two adjacent base metals (Fig. 2.15). The constitutive behavior of the weld zone is obtained by conducting tensile tests on subsize specimens with a longitudinal weld and following the ROM technique, as discussed in a previous section. These assumptions (weld line or weld zone) incorporated when modeling TWB deformation play a vital role in predicting the forming behavior.

2.15 Schematic of weld region representation in FE simulations of TWB (90° weld orientation).



41

2.7.1 Representation of weld region in the FE simulation of TWBs In most of the published literature, TWBs are modeled without incorporating weld properties, i.e. the weld line assumption is followed. The reasons for this are twofold: (1) evaluating weld properties is difficult and cumbersome, as laser welded blanks (which are widely used) possess a weld width of the order of just 2 mm, and (2) the weld line assumption can predict the forming behavior of TWBs with acceptable accuracy, as the weld region occupies very little area of the entire TWB part produced by laser welding, say a door inner, fender, bi-axial stretching (200 × 200 mm) strain path in LDH testing, etc. If the weld zone assumption has to be followed, then the weld properties should be evaluated accurately like those of the base materials. With the available methods, only a few weld properties like yield strength, strain hardening exponent, strength coefficient (tensile testing of subsize samples with longitudinal weld and following ROM) and weld width (micro-hardness measurement, microstructure) can be evaluated. Other weld properties are assumed while modeling TWBs. Though the weld line assumption can predict forming behavior to acceptable accuracy, the weld region is a separate zone wherein the mechanical properties are quite different than that of the base metal. So, in principle, a TWB modeled with the weld zone assumption should predict the forming behavior of TWBs more accurately. Also, relative mechanical properties of the weld region with that of base material, say a softer or harder weld zone, will affect the forming behavior prediction considerably. This issue cannot be addressed if the weld line assumption is followed in FE simulations. If the weld region is incorporated with the same mechanical properties as the base metal during FE simulations, one can expect that the predictions will be the same as those of TWBs modeled with a weld line assumption or un-welded blank. But, this is practically impossible. Even if the weld properties are considered for simulation, identifying the weld shape and incorporating it in the simulations is a task by itself. This will become slightly complex if different thickness sheets are modeled, which also have a profound influence on the forming behavior prediction. The weld line assumption is followed in many studies1, 13–19, 24–27 of the forming behavior of TWBs like strain distribution, weld line movement control, deep drawability, etc. Few studies have aimed at including weld properties into simulations11, 28–34 also. For instance, in several studies29–31 the weld properties, viz. yield strength, strength coefficient and strain hardening exponent, are included in modeling the weld region with some variation within or along with HAZ properties. The HAZ (4 mm) properties are given a triangular variation in the twodimensional FE model which contains linear elements and gradual stepped variation inside the HAZ three-dimensional model (containing rectangular elements).29 A door inner panel and hemispherical dome stretching are simulated. The thickness strain distribution and weld line movement predicted correlates


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

42 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


well with that from experiments. In studies by Rojek et al.,30, 31 the HAZ and weld properties are included into the model, by obtaining the yield strength of the weld zone from the following relation:

σyweld = (σysheet HV weld)/(HV sheet)

[2.6]

where, σyweld and σysheet are the yield strengths of weld and base metal respectively. HV weld and HV sheet are the hardness of the weld zone and base material respectively. Some of the work performed on modeling issues compares the importance of weld line and weld zone assumptions in predicting forming behavior.11, 32–34 This issue is very important and critical from the accuracy point of view in predicting TWB forming behavior. For example, Raymond et al.32 compared the TWB models with and without weld properties using ASTM tensile test simulation, in-plane test and LDH test simulations. Weld line movement and plastic strain are monitored for different thickness ratios. It was found that there are a number of relatively subtle effects in which the weld is modeled and most of them are related to the constraining effect of the weld region during forming. This indirectly means that weld line assumption (excluding weld properties in simulations) is sufficient to model TWBs. But this conclusion depends on the forming behavior that is monitored during the simulation. In Raymond et al.,32 weld line displacement is monitored, which may not be affected by the methods by which the weld is modeled. The weld line moves because of the change in plastic deformation levels attained by the thinner and thicker metals during forming. Weld line or weld zone assumptions may not affect the weld line movement prediction much. Instead, forming behavior such as load progression curves, limit strains (or FLC) will be affected. Also, in Raymond et al.,32 TWBs with different thickness ratios and transverse welds are simulated. In this case, only the thinner metal will deform without much happening in the thicker blank and hence it is quite obvious to expect that weld line modeling is sufficient to predict the forming behavior accurately. In Bhagwan et al.,11 weld line shift and equivalent plastic strains are monitored by performing LDH test simulations using LS-DYNA FE software by including weld properties and geometry. It was found that weld geometry is more important than weld properties in predicting the weld line shift. But this was concluded by taking only one set of weld properties (i.e. n, σys) without any variations. Also, here the σys ratio (σys-weld /σys-base) is approximately 1.06, which reflects that the weld exhibits approximately the same yield strength as that of the base metal. The relative property ratio plays a vital role in deciding the participation of the weld region in the deformation. Suppose, if the weld is softer than the base metal (e.g. σys ratio = 0.5), the weld properties will dominate the overall forming behavior of the TWBs. It is obvious, as said earlier, that for different thickness combinations the weld line assumption may be sufficient. So, one has to look into the worst case scenario of considering the same material



43

and thickness to study the importance of weld zone representation in FE simulations. It is pointed out by Kampus and Balic33 that weld region modeling with weld properties predicts the actual shape of a deep drawn cup in the weld region and base metal more accurately than a TWB modeled without weld properties. Similar results are seen in Zhao et al.34 In Zhao et al.,35 six different TWB models, viz. weld parallel to bending moment – shell element with weld, shell element without weld, solid element with weld and the same three models for weld normal to bending moment are compared. Free bend tests, stretch bend tests and LDH tests are simulated and load–stroke and spring back behaviors are monitored. It is concluded from the analyses that a weld region modeled with the HAZ properties (by shell elements) improves the prediction of forming behavior. To summarize, the importance of weld line or zone modeling depends on the weld conditions like orientation, properties, location, deformation mode or part shape, thickness combination, etc. In general, weld line modeling is sufficient for most TWB conditions, but not always. Some of the TWB conditions demand weld zone assumption, so that the forming behavior of TWBs can be predicted accurately. In fact one can also study the modeling aspect of welding36 and subsequent stamping processes required for making TWBs as performed by Zimniak et al.37 Narayanan et al.38, 39 addressed the importance of weld region modeling by identifying the domain of weld conditions explicitly wherein the weld line or the weld zone assumption needs to be followed. The weld conditions considered are weld properties (n, σys), weld orientation (transverse, longitudinal) and weld width. Tensile and LDH tests (plane strain, biaxial stretching strain paths) are simulated. Same thickness and material (IF steel) is modeled in the simulations. The load–stroke behavior, maximum load reached and progression at failure are monitored and comparison is done between weld line and zone assumptions. It is found from the analysis that the weld line assumption is not always sufficient to model TWBs, and that this depends on the weld conditions and mode of deformation. Narayanan et al.38,39 propose that appropriate representation of weld region in simulations depends on whether the weld region is participating in the deformation or not. For example, if the weld region is soft and is normal to the major straining direction, failure occurs only in the weld region without much happening in the base metal. In this case, the weld zone assumption is essential. But if the weld is harder than the base metal, then base metals accommodate most of the deformation and hence the weld line assumption is sufficient here.

2.8

Conclusions

The effect of various parameters on the forming behavior of TWBs was discussed in this chapter. Also, some of the important issues in modeling TWBs, controlling weld line movement and evaluating the mechanical properties of weld zone were discussed. The above discussion can be summarized as follows:


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

44 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

•

•

•

•

Tailor welded blanks for advanced manufacturing The TWB parameters like thickness ratio, strength ratio and various weld conditions affect its forming behavior in a synergetic manner. The weld zone may be either hard or soft, and loading may be parallel or normal to the weld region which decides the formability of TWBs. Assuming same material and thickness, with a ‘soft’ weld, loading parallel to the weld may not be a problem, but loading normal will result in weld failure, because of the strength differential. In the case of a ‘hard’ weld, loading parallel to the weld can reduce the forming limits of TWBs, as the weld region possesses less ductility. Transverse loading of a TWB with a hard weld does not pose a significant problem as the weld will not deform plastically. In the case of different thickness or strength combinations, the weaker region, either the blank or the weld, decides the forming limit of TWBs if the weld is normal. In the case of parallel loading, formability of a TWB is decided both by the weld and base materials. In general, thickness ratio or strength ratio dominates over other parameters in deciding the forming behavior, with weld properties, say σys and n, showing significant effect depending on whether the weld zone is participating in the deformation or not. In cases where the weld is stronger than the base sheets welded, then the FLCs of such TWBs are typically in-between those of the corresponding base metals. Weld line movement is a common phenomenon during the drawing of TWBs, and it takes place due to differential draw resistance of the sheets welded. Weld line movement can lead to significant reduction in the formability of TWBs. Weld line movement can be controlled by adapting techniques like applying differential blank holding force (BHF), drawbead restraining forces, designing special local adaptive controllers etc. The main idea here is to control the plastic deformation and draw-in level attained by thinner or weaker material, so that it will be nearly equivalent to that of thicker or stronger material, which eventually minimizes the weld line movement. TWBs consisting of sheets of differing thicknesses need special design considerations for forming. The thickness difference needs to be accommodated either by using a shim, stepped die or blank holder. The stepped tool design should provide sufficient slack for allowing free movement of the weld. In a component, the weld line should be appropriately located, considering not only formability issues, but also the overall performance of the component in an assembled structure. Modeling a TWB is a critical issue depending on how the weld region is represented during FE simulations. The weld region can be treated simply as a weld line, without incorporating weld properties, or as weld zone by considering weld properties during modeling. The weld line assumption is not accurate for all the TWB conditions. For some TWB conditions and deformation modes, the weld zone assumption should be followed; for example, TWBs with softer weld zones show accurate forming behavior predictions with weld zone modeling. The important properties of weld zone,



45

σys, n, K, can be evaluated by tensile testing of subsize samples and following the ROM methodology. These weld properties can be input to FE simulation models for more accurate forming behavior predictions.

2.9

Acknowledgment

The authors are thankful to the CAR–TIFAC funding (Grant No. 05TI002, IIT Bombay, India) that supported a significant part of the work reported in this chapter.

2.10

References

1. F. I. Saunders, R. H. Wagoner, ‘Forming of tailor welded blanks’, Metallurgical and Material Transactions A, Vol. 27A, September 1996, p. 2605. 2. K. Abdullah, P. M. Wild, J. J. Jeswiet, A. Ghasempoor, ‘Tensile testing for weld deformation properties in similar gage tailor welded blanks using the rule of mixtures’, Journal of Materials Processing Technology, Vol. 112, 2001, p. 91. 3. B. Y. Ghoo, Y. T. Keum, Y. S. Kim, ‘Evaluation of the mechanical properties of welded metal in tailored steel sheet welded by CO2 laser ’, Journal of Materials Processing Technology, Vol. 113, 2001, p. 692. 4. C. H. Cheng, L. C. Chan, C. Y. Tang, ‘Determination of true stress–strain curve for the weldment of aluminum laser-welded blanks’, Journal of Laser Applications, Vol. 17, No. 3, August 2005, p. 159. 5. S. Liu and Y. J. Chao, ‘Determination of global mechanical response of friction stir welded plates using local constitutive properties’, Modelling and Simulation in Materials Science and Engineering, Vol. 13, 2005, p. 1. 6. M. Tolazzi, M. Merklein, ‘Determination of tensile properties for the welding line of tailored welded blanks’, Proceedings of the 8th ESAFORM Conference on material forming, ESAFORM 2005, Cluj-Napoca, Romania, April 2005, p. 281. 7. Marc Auger, Kassim Abdullah, Jack Jeswiet, Peter Wild, Lynann Clapham, ‘Determination of weld line characteristics in tailored blanks’, SAE paper, 2000–01–2661. 8. R. W. Davies, H. E. Oliver, M. T. Smith, G. J. Grant, ‘Characterizing Al tailor welded blanks for automotive applications’, Journal of Metals, Vol. 51, Issue 11, November 1999, p. 46. 9. K. B. Min, S. S. Kang, ‘A study on resistance welding in steel sheets using a tailor welded blank (2nd Report): Evaluation of flash weldability and formability’, Journal of Materials Processing Technology, Vol. 103, 2000, p. 218. 10. K. B. Min, K. S. Kim, S. S. Kang, ‘A study on resistance welding in steel sheets using a tailor welded blank (1st report): Evaluation of upset weldability and formability’, Journal of Materials Processing Technology, Vol. 101, Issues 1–3, 2000, p. 186. 11. Amit V. Bhagwan, Ghassan T. Kridli, Peter A, Friedman, ‘Influence of weld characteristics on numerically predicted deformation behavior of aluminum tailor welded blanks’, SAE paper, 2002-01-0386 (DOI: 10.4271/2002-01-0386). 12. C. H. Cheng, L. C. Chan, C. L. Chow, T. C. Lee, ‘Experimental investigation on the weldability and forming behavior of aluminum alloy tailor welded blanks’, Journal of Laser Applications, Vol. 17, No. 2, May 2005, p. 81.


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

46 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


13. Mustafa A. Ahmetoglu, Dirk Brouwers, Leonid Shulkin, Laurent Taupin, Gary L. Kinzel, Taylan Altan, ‘Deep drawing of round cups from tailor welded blanks’, Journal of Material Processing Technology, Vol. 53, 1995, p. 684. 14. Youngmoo Heo, Youho Choi, Heon Young Kim, Daegyo Seo, ‘Characteristics of weld line movements for the deep drawing with drawbeads of tailor welded blanks’, Journal of Materials Processing Technology, Vol. 111, 2001, p. 164. 15. Brad Kinsey, Zhihong Liu, Jian Cao, ‘A novel forming technology for tailor-welded blanks’, Journal of Materials Processing Technology, Vol. 99, 2000, p. 145. 16. Brad Kinsey, Zhihong Liu, Jian Cao, ‘New apparatus and method for forming tailor welded blanks’, Journal of Materials and Manufacturing, SAE paper, 1999–01–0681, Vol. 108, Section 5, p. 653. 17. Brad Kinsey, Vikram Viswanathan, Jian Cao, ‘Forming of aluminum tailor welded blanks’, SAE paper, 2001–01–0822, p.1. 18. Brad Kinsey, Nan Song, Jian Cao, ‘Analysis of clamping mechanism for tailor welded blank forming’, SAE paper, 99IBECC–26 19. Brad L. Kinsey, Jian Cao, ‘Enhancement of sheet metal formability via local adaptive controllers,’ Transactions of NAMRI/SME, Vol. XXIX, 2001, p. 81. 20. Klaus Siegert, Edgard Knabe, ‘Fundamental research and draw die concepts for deep drawing of tailored blanks,’ SAE paper, 950921, p. 866. 21. Z. Zimniak, A. Piela, ‘Finite element analysis of a tailored blanks stamping process’, Journal of Materials Processing Technology, Vol. 106, 2000, p. 254. 22. Yu-Ning Liu, E. Kannatey-Asibu, Jr., ‘Finite element analysis of heat flow in dual beam laser welded tailored blanks’, Journal of Manufacturing Science and Engineering, Vol. 120, May 1998, p. 272. 23. Hongjong Kim, Youngmoo Heo, Naksoo Kim, Heon Young Kim, Daegyo Seo, ‘Forming and drawing characteristics of tailor welded sheets in a circular drawbead’, Journal of Materials Processing Technology, Volume 105, Issue 3, 2000, p. 294. 24. A. Buste, X. Lalbin, M. J. Worswick, J. A. Clarke, M. Finn et al., ‘Prediction of strain distribution in Aluminum tailor welded blanks’, Proceedings of the 4th International Conference and Workshop on Numerical simulation of 3D sheet forming process, NUMISHEET’99, Besançon, France, 1999, p. 455. 25. T. Meinders, A. van den Berg, J. Huetink, ‘Deep drawing simulations of tailored blanks and experimental verification’, Journal of Materials Processing Technology, Vol. 103, 2000, p. 65. 26. Ghassan T. Kridli, Peter A, Friedman, Andrew M. Sherman, ‘Formability of aluminum tailor welded blanks’, SAE paper, 2000–01–0772, p. 1. 27. M. A. Ahmetoglu, D. Brouwers, G. Kinzel, T. Altan, ‘Deep drawing of round cups from laser-tailor welded blanks’, Proceedings of Laser Assisted Net Shape Engineering 1994, LANE ‘94, Vol. 1, Meisenbach Bamberg, p. 167. 28. D. Dry, D. Hughes, R. Owen, ‘Methods of assessing influence of weld properties on formability of laser welded tailored blanks’, Ironmaking and steelmaking, Vol. 28, No. 2, 2001, p. 89. 29. B. Y. Ghoo, S. J. Back, Y. T. Keum, S. Y. Kang, ‘Finite element analysis of tailored sheet forming processes considering laser welding zone’, Metals and Materials, Vol. 4, No. 4, 1998, p. 862. 30. Jerzy Rojek, Eugenio Onate, Antoni Piela, Laurentiu Neamtu, ‘Numerical modeling and simulation of tailor welded blanks’, Proceedings of the 5th International Conference and Workshop on Numerical simulation of 3D sheet forming process, NUMISHEET’02 Jeju Island, Korea, 2002, p. 177.



47

31. Antoni Piela, Jerzy Rojek, ‘Experimental study and modeling of tailor welded blanks’, Proceedings of the 5th International Conference and Workshop on Numerical simulation of 3D sheet forming process, NUMISHEET ’02, Jeju Island, Korea, 2002, p. 225. 32. Scott D. Raymond, Peter M. Wild, Christopher J. Bayley, ‘On modeling of the weld line in finite element analyses of tailor-welded blank forming operations’, Journal of Materials Processing Technology, Vol. 147, Issues 1, 2004, p. 28. 33. Z. Kampus, J. Balic, ‘Deep drawing of tailored blanks without a blankholder ’, Journal of Materials Processing Technology, Vol. 133, Issues 1–3, 2003, p. 128. 34. K. M. Zhao, B. K. Chun, J. K. Lee, ‘Finite element analysis of tailor welded blanks’, Finite Elements in Analysis and Design, Vol. 37, 2001, p. 117. 35. K. M. Zhao, B. K. Chun, J. K. Lee, ‘Numerical modeling technique for tailor welded blanks’, SAE paper, 2000–01–0410, p. 151. 36. Yu-Ning Liu, E. Kannatey-Asibu, Jr., ‘Finite element analysis of heat flow in dual beam laser welded tailored blanks’, Journal of Manufacturing Science and Engineering, Vol. 120, May 1998, p. 272. 37. Z. Zimniak, A. Piela, ‘Finite element analysis of a tailored blanks stamping process’, Journal of Materials Processing Technology, Vol. 106, 2000, p. 254. 38. R. Ganesh Narayan, V. Vijay Bhaskar, K. Narasimhan, ‘Effect of weld conditions on the deformation behavior of tailor welded blanks (TWB)’, Proceedings of the 8th International Conference on Numerical Methods in Industrial Forming Processes, NUMIFORM ’04, Columbus, Ohio, 2004, p. 856. 39. R. Ganesh Narayan, K. Narasimhan, ‘Weld region representation during the simulation of TWB forming behavior,’ International Journal of Forming Processes, Vol. 9 (4), 2006, p. 491.


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

3 Mechanics-based modeling of tailor welded blank forming B. L. KINSEY, University of New Hampshire, USA

Abstract: Tailor welded blanks (TWBs) offer several notable benefits including decreased component weight, reduced manufacturing costs, increased environmental friendliness and improved dimensional accuracy. In order to take full advantage of these benefits, designers must be able to predict the reduced formability of TWBs and the unique attributes related to TWB forming early in the design process. In this chapter, analytical models to predict the limiting thickness ratio of the TWB combination and the forming behavior of TWBs will be presented. In these studies, comparisons with numerical simulations and experimental results will be used to validate the accuracy of these models. Such analytical models provide designers with valuable tools to determine material and process parameters for the application, such as the location of steps on the die surface to accommodate the weld displacement, the non-uniform binder force ratio to prevent weld line movement and the potential forming height for a TWB forming process. Key words: analytical model, tailor welded blanks, weld line movement, material characterization, non-uniform binder force.

3.1

Introduction

Analytical process models are very useful at the early stage of process development to provide engineers with initial design feedback regarding forming techniques and processes without time consuming and expensive experimentation or finite element analyses (FEA). With respect to TWBs, a number of analytical models have been developed. These include determining the limiting thickness and strength ratio for a given TWB application (Shi et al. 1993), calculating the forming limit curve for TWBs (Cayssials et al. 2000, Davies et al. 2001) using a Marciniak–Kuczynski (M–K) analysis, providing analytical analysis of the weld properties for TWBs (Doege et al. 1996 and Davies et al. 2000) and using twodimensional (2-D) cross-sectional analyses to determine a non-uniform binder force ratio to improve TWB formability and the weld line movement and forming height in the process (He et al. 2001, Kinsey and Cao 2003, respectively). As with analytical models in general, mathematical expressions of material behavior and deformation mechanisms are required. In this chapter, details with respect to three of these analytical models (Shi et al. 1993, Kinsey and Cao 2003 and He et al. 2001) will be presented. Throughout this chapter the term ‘forming height’ refers to the drawing depth and the term ‘weld line movement’ refers to the weld line 48 © Woodhead Publishing Limited, 2011


49

displacement in the direction normal to the weld line within the sheet plane. The positive direction is toward the thicker sheet side.

3.2

Thickness and strength ratio analysis

One of the first analytical models developed with respect to TWB forming was by Shi et al. (1993), which determined the limiting thickness and strength ratio feasible for a TWB application. Figure 3.1 shows a cross-section of a TWB with varying thicknesses (tA and ta for the thicker and thinner sides of the TWB, respectively). Throughout this chapter an orthogonal coordinate system is used with the 1- and 2-directions denoted as perpendicular and parallel directions to the weld line segment respectively within the sheet plane (for both flat and contoured sheets) and the 3-direction is in the sheet thickness direction. While Fig. 3.1 represents the TWB with varying thicknesses, recall that TWBs can be produced with sheets of varying strength as well. For this configuration, force (F) equilibrium must be satisfied in the direction perpendicular to the weld seam: F1A = F1a

[3.1]

This equation can be rewritten assuming a uniform width in the 2-direction as:

σ1AtA = σ1ata

[3.2]

where σ is the true stress. The thickness can also be expressed as: t = toeε3 = toe−(ε1+ε2)

[3.3]

where to is the initial thickness and ε is the strain in the various directions (see Fig. 3.1) based on the definition of true strain and assuming incompressibility of the material. Substituting this expression into Eq. [3.2] gives:

σ1AtoAe−(ε1A+ε2A) = σ1atoae−(ε1a+ε2a)

[3.4]

Assuming that there will not be a discontinuity in the strain value along the weld line,

ε2A = ε2a

[3.5]

This has been shown experimentally to be the case in past studies (e.g., Shi et al. 1993 and Bravar et al. 2007). Figure 3.2, from Shi et al. (1993),

3.1 Schematic of cross-section of weld line for a TWB with varying thicknesses in sides A and a. For the coordinate system used, the 2-direction is in the width direction (i.e. parallel to the weld seam).


50


3.2 Strain values longitudinal and transverse to the weld line for a TWB application (Shi et al., 1993; reprinted with permission from SAE paper 930278 © 1993 SAE International).

demonstrates a constant strain longitudinal to the weld line. Thus, Eq. [3.4] can be simplified to:

σ1AtoAe−(ε1A) = σ1atoae−(ε1a)

[3.6]

The engineering stress (S) can be substituted into Eq. [3.6] to give: S1AtoA = S1atoa

[3.7]

This equation provides a relationship between the initial thickness ratio and the engineering stress ratio of the material at any stage of deformation in the TWB application. For a TWB application, achieving some amount of plastic strain in both sheet materials is desirable for part rigidity. Using Eq. [3.7], a ratio of thickness or strength values to assure this condition is achieved can be obtained. Assume material A has a greater initial thickness or higher strength for the TWB application. Then, S1a would be the limiting stress for the application. This value can be set to the tensile strength of the material (STa). Note that since uni-axial tension is not necessarily the loading condition, this is an assumption in the model. At the loading state that material a is at failure, assume material A is yielding (SYA). Thus, the limiting thickness ratio (LTR) can be determined from the strengths of the two parent materials:



51

[3.8] This LTR value would represent the minimum thickness ratio where the thicker material in a TWB application would experience some plastic deformation. Similarly, a limiting strength ratio (LSR) can be determined from the desired thickness ratio: [3.9] Shi et al. (1993) presented an example of an AKDQ steel door inner application to demonstrate the effectiveness of their model. The TWB was created with 1.8 mm and 0.8 mm materials with the thicker material used to eliminate the hinge and window reinforcements. The thickness ratio (toA/toa) was thus 2.25 for this TWB application. The yield strength for the thicker material (SYA) was 186.9 MPa while the tensile strength for the thinner material (STa) was 326.8 MPa. According to their model, the minimum LTR to achieve some plastic deformation in the thicker material would be 326.8/186.9 which is equal to 1.75 (see Eq. [3.8]). Since the calculated LTR is less than the one used in the application, the thicker material is predicted not to plastically deform during the operation. This was confirmed by measuring the strain transverse to the weld line (see Fig. 3.2). Since the transverse strain values on the 1.8 mm side of the TWB are nearly zero, this example supports the analytical model prediction.

3.3

Determination of weld line movement and forming height

In addition to the LTR and LSR, other parameters are of interest in TWB forming applications. For example, the amount that the weld line will move during the operation and the forming height when a specified strain at the weld line would be achieved are important parameters to determine during the design phase. Kinsey and Cao (2003) developed an analytical model to predict these parameters. The details for their analytical model are as follows. A 2-D sectional analysis was used to simplify the calculations. This is a standard technique used in analytical modeling of metal forming processes (e.g., Saran et al. 1991, Brooks et al. 1991 and Yao et al. 2000). Figure 3.3 shows a 2-D profile for a TWB application with thicker material on the left side of the blank and thinner material on the right side. The orientation of the weld line is perpendicular to this 2-D cross-section and the location of the analysis would be at a position where the value of the weld line movement and strain in the neighborhood of the weld line are of interest. Key locations on the 2-D cross-section are identified with capital letters on the thicker material side, from A at the weld line to E at the die radius on the binder area, and lower case letters on the thinner material side. © Woodhead Publishing Limited, 2011

52


3.3 2-D cross-sectional representation of a TWB formed part with key locations identified (adapted from Kinsey and Cao, 2003).

(Sections are also numbered I to V, which will be used when describing the He et al. (2001) model.) The 1-direction follows the contour of the 2-D cross-section and is a local coordinate axis. While a 2-D cross-section is used, several analyses could be performed to provide information for a 3-D geometry. Similar to the model by Shi et al. (1993), the model is based on force equilibrium but also includes bending theory and frictional effects. Other assumptions in the model include plane strain, no shear stress, incompressibility and negligible thickness stress (i.e., σ3 = 0). An input into the analytical model is the true strain in the 1-direction at location a in the thinner material at the weld line. This value of the strain could be set based on the known plane strain failure limit or the desired strain amount to achieve the desired part stiffness. From this ε1a value, the forces, T1, and strains, ε1, in the 1-direction are calculated at all locations shown in Fig. 3.3 from the weld line to the binder area. Assuming a unity width and using the relationship of Eq. [3.3]: T1a = σ1a tote−ε1a

[3.10]

where the original thickness of the thinner blank and σ1a is the stress in the 1-direction of the material. This relationship also holds at all locations in the 2-D cross-section with the corresponding substitution for the location subscripts. The constitutive behavior of the material is modeled using a power hardening law, t ot is

[3.11] where σ– is the equivalent stress, K is the strength coefficient, n is the strain hardening exponent, and ε– is the equivalent strain. The K and n values are assumed to be the same for the thicker and thinner materials in the TWB combination. Force equilibrium must exist across the weld line, i.e., T1a = T1A. Thus, the strain at A, ε1A, can be calculated from:



53 [3.12]

where tOT is the original thickness of the thicker blank. (Note a discrepancy as the power hardening law for uni-axial tension was used in this plane strain case; however, this assumption produces reasonable results as will be shown.) Force equilibrium also exists at all of the locations in the 2-D cross-section of Fig. 3.3. For a straight section, a simple force equilibrium can be used, e.g. T1a = T1b. Additional force terms for bending, ∆TB, and friction, ∆TF , are required for sections that include a radius. The force relationships that apply to all of the sections on the thinner side of the TWB are:

[3.13]

Whether the material initially located at location b on the 2-D cross-section flows onto the punch face (negative) or into the formed wall (positive) determines the sign of the ∆TB and ∆TF terms in the equation for T1c. For location e, the material is assumed to flow from the binder area into the formed wall, which is typically the case for sheet metal forming operations. Identical relationships exist for the thicker side as well with the appropriate substitutions for the 2-D cross-sectional location subscripts. Several past analytical models have been developed to assess the forces and strains associated with plane strain bending (e.g., Zhang and Hu 1997, Wood et al. 1985 and Wang and Wenner 1978). For the model by Kinsey and Cao (2003), the methodology proposed by Swift (1948) was employed to calculate ∆TB, i.e.:

[3.14] where σY is the yield strength of the material, t is the material thickness and R is the bend radius. Note that the increase in strain due to bending is not accounted for in this model or the effects of bending and unbending during material flow. The ∆TF term is related to the friction coefficient, µ, and the angle of wrap around the radius, θ, through:

∆TF = Teµθ

[3.15]

With these relationships, the forces and strain values in the 1-direction and the stress values in the 2-direction at all sectional locations can be calculated. See the theoretical model details in the appendix of Kinsey and Cao (2003).


54


3.4

Determination of material draw-in ratios

From the force values at E and e, a ratio of the material draw-in from under the binder area into the formed wall on the thick (xT) and thin (xt) sides of the TWB 2-D cross-section are assumed to follow the expression: [3.16] The rationale for this assumption is that the forces in the 1-direction at e and E are the forces that draw the material into the formed wall and therefore their ratio is directly proportional to the ratio of the material draw-in. This relationship though would depend on the initial location of the weld line. Equation [3.16] applies to cases where the initial weld line position is in the center of the 2-D cross-section. If the initial weld line is offset from the center of the blank, an alternative relationship is: [3.17] where XWo is the initial weld line offset position (positive if moved towards the thinner material and negative if moved towards the thicker material) and Lo is the total initial length of the blank. While Eq. [3.16] and [3.17] provide a means to calculate the draw-in ratios, expressions to determine the actual draw-in values based on the geometric parameters and the strain values calculated in the 1-direction are required. These draw-in values are related to the forming height for the application (xH), which is a parameter interest for this model. For the thinner side of the 2-D cross-section, the material draw-in can be calculated from: xt = lf − lo − (lf × ebe) ± xb

[3.18]

where lf is the final length of the formed wall, lo is the length of the material above the formed wall prior to the process including the punch and die radii, ebe is the engineering strain in the 1-direction from section b to section e of the formed wall and xb is the movement of the material initially at location b onto the punch face (positive) or into the formed wall (negative). Note that lf and lo are geometric parameters based on the application and lf is a function of the forming height. See Fig. 3.4 for a graphical representation of these variables on the 2-D cross-section. The correct value of xb for the application is not known a priori so a value of xb is assumed to allow calculations to proceed. The weld line movement (xW) can also be calculated from the strain values, the geometric parameters and the assumed xb value from: xW = (lab × eab) ± xb


[3.19]


55

3.4 Schematic of thin side of the 2-D cross-section showing geometric parameters (Kinsey and Cao, 2003).

A similar expression to Eq. [3.18] can also be written for material draw-in on the thicker side of the TWB, which includes the weld line movement and the stretching of the material on the punch face: xT = lf − lo − (lf × eBE) − (lAB × eAB) − xw

[3.20]

Equations [3.18] through [3.20] can be used to create a ratio of draw-in values, xt and xT , which can be equated to the draw-in ratio value calculated in Eq. [3.16] or [3.17]. Since an xb value was assumed, the only parameter which is unknown is the lf value, which is a function of the forming height, xH. Thus, a plot of xH versus xb values can be created (e.g., see Fig. 3.5). Note the local minimum in the graph.

3.5 Relationship between forming height and the material movement at location b for the 2-D cross-section (Kinsey and Cao, 2003).


56


3.6 Flowchart for the analytical model (adapted from Kinsey and Cao, 2003).

Based on practical forming experiences, one unique forming condition is achieved for each set of process parameters. This is often assumed to be the minimum energy solution. Thus, the xb value where the minimum forming height is achieved is assumed to be the actual movement of the material onto or off of the punch face for the given process. Figure 3.6 shows a flowchart summarizing the analytical model.

3.4.1 Comparison with 2-D numerical simulation results To assess the effectiveness of the analytical model, results were compared with both numerical simulations and experiments. For the FEA, two geometries were investigated using ABAQUS/Standard (Kinsey and Cao 2003). The FEA models were plane strain and used four-node, reduced integration elements. Figure 3.7 shows one of the geometries investigated including various locations for the initial weld line position, i.e., XWo. The Coulomb friction coefficient was assumed to be 0.15 and the material, Al 5182-H00, was modeled using a power hardening law with K = 570 MPa and n = 0.3 with 2 mm and 1 mm material thicknesses for the simulations. The forming height for the model was set to 30 mm and the binder force was varied to achieve an ε1a value of 8% as this has been shown to be the forming limit for this particular TWB combination (Viswanathan et al. 2001). See Kinsey and Cao (2003) for further details related to the FEA performed. Table 3.1 lists several cases for the geometry in Fig. 3.7, including ones with different initial weld line positions and thickness combinations. Also listed in the table are results comparing the weld line movement and forming height. As is evident, the analytical model provided reasonable results compared to the 2-D FEA simulations. In all cases, the analytical model under-predicted the weld line movement and forming height values which would lead to conservative designs. Discrepancies observed are due in part to the assumptions used in the analytical model, e.g., the simplified bending model.



57

3.7 2-D cross-sectional geometry for numerical simulations and analytical model including initial weld line location xWo (adapted from Kinsey and Cao, 2003).

Table 3.1 Comparison of results from 2-D numerical simulations and analytical model TWB thickness combination (mm)

Initial xW weld line 2-D FEA position (mm) xWo (mm)

xW analytical model (mm)

xW xH difference 2-D FEA (mm) (mm)

xH analytical model (mm)

xH difference (mm)

2.0/1.0

−24

5.53

4.18

1.35

25.00

5.00

−16

4.92

3.61

1.31

25.03

4.97

−8

4.47

3.05

1.42

25.12

4.88

0

3.74

2.53

1.21

25.36

4.64

8

2.99

1.92

1.07

25.63

4.37

16

2.43

1.35

1.08

26.26

3.74

0

2.38

2.23

0.15

24.76

5.24

1.5/1.0

30

30

Source: adapted from Kinsey and Cao, 2003.

Other parameters can also be compared with respect to the analytical model and numerical simulation results, e.g., the forces and strains in the 1-direction, T1 and ε1. Figures 3.8 and 3.9 show normalized results with respect to values at locations A and a for these parameters respectively for the 2.0/1.0 thickness combination case and the initial weld line location in the center of the blank. For these plots, reasonable agreement exists at all locations except for the strain values at locations D and d. This is due to the simplified bending model used for the analytical results as stretching caused by bending and unbending was not included. Note that the


58


3.8 Results for normalized force in the I-direction for both 2-D FEA and analytical model at key locations in the 2-D cross-section (Kinsey and Cao, 2003).

3.9 Results for normalized strain in the I-direction for both 2-D FEA and analytical model at key locations in the 2-D cross-section (Kinsey and Cao, 2003).



59

discrepancy on the thicker material side at location D is only 0.4%, but appears significant due to the normalized parameter used.

3.4.2 Comparison with experimental and 3-D numerical simulation results The analytical model by Kinsey and Cao (2003) was also compared with experimental and 3-D numerical simulation results. For these analyses, instead of calculating the weld line movement, a slight modification was incorporated into the analytical model so that the final weld line position (Xf) was entered instead of the initial weld line position. This would be more representative to the design process, as the engineer would specify the desired final weld line location in the part and the initial weld line position would be determined. Also, another modification to the analytical model was that the strains in the 2-direction were entered at all locations on the cross-section. This eliminates the plane strain assumption, which is more realistic for forming operations. The ε2 values were determined from the given experimental or 3-D numerical simulation case, but could be predicted based on past experience by designers. The geometric and material parameters for the symmetric, steel experimental case and the non-symmetric, aluminum 3-D numerical simulation case considered are given in Table 3.2 where lG is the gap between the punch and die, lP is the length of the punch face not including radii and RD and RP are the radii on the die and punch respectively. For the pan shaped, symmetrical experimental case, the material used was DP600GI steel with thicknesses of 1.37 mm and 0.66 mm, and dimensions for each blank were 260 mm by 600 mm. The punch dimensions were 304.8 mm by 381 mm with the weld line centered on the 304.8 mm dimension. Figure 3.10 shows a formed TWB for this steel case, and Fig. 3.10 shows a graph of ε2 values from the experiments and numerical simulations, which were conducted in DYNAFORM-PC. Note the reasonable FEA prediction of these ε2 values in Fig. 3.11. See Krishnan and Cao (2004) for further details related to the experiments and the numerical simulations. The results comparing the experiments to the analytical model for both the plane strain case and the case where ε2 values were entered into the model are Table 3.2 Tooling dimensions and material properties for cases investigated Case Experimental symmetric, steel FEA nonsymmetric, Al

lG (mm) 6.35 30

xH (mm) lP (mm) RD (mm)

RP (mm) K (MPa) N

50

279.4

12.7

12.7

648

0.23

80

460

20

20

570

0.3

Source: adapted from Bravar et al., 2007.


60


3.10 Specimen from symmetric steel case (Krishnan and Cao, 2004).

3.11 Strain in the 2-direction for the various cases (Bravar et al., 2007).

presented in Table 3.3. While reasonable predictions were obtained with the original model, which included the plane strain assumption, the elimination of the plane strain assumption created significantly better results. Figure 3.12 shows the geometry for the non-symmetric, aluminum case (which is a crude representation of a TWB door inner). Also shown on this figure is the location of the 2-D cross-section for the analytical model, which is where the maximum weld line movement was observed. Geometric and material parameters are also listed in Table 3.2. The numerical simulations were conducted using DYNAFORM-PC with Belytschko-Tsay shell elements, and the punch and die



61

Table 3.3 Comparison of weld line movement and forming height for experimental, analytical model (both with and without plane strain assumption) and FEA results TWB case and condition

Final weld line location XWf (mm)

Analytical model using plane strain assumption and XWf

Analytical model using FEA ε2 and XWf

FEA results

Analytical model using plane strain assumption and XWf

Analytical model using exp. ε2 and XWf

Exp. results

Weld line parameter

Experimental, symmetric steel TWB case FEA, non-symmetric Al TWB case

10.7

10.7

10.7

85.7

85.7

85.7

Weld line 10.7 movement xW (mm)

10.2

7.3

15.7

16.0

18.0

0.5

3.4

70.0

69.7

67.7

−0.5

−3.4

0.3

2.3

Initial weld 0 line location XWo (mm) Difference compared to XW exp. (mm)

–

–

Source: Adapted from Bravar et al., 2007.

were modeled as rigid bodies. Complete details with respect to the numerical simulations can be found in Kinsey and Cao (2003). Figure 3.11 shows the ε2 values which were obtained from the numerical simulations and entered into the analytical model along this cross-section. Unfortunately, values of ε2 were not obtained on the test specimens, which is what prevented the analytical model from being compared to experimental results. However, good agreement between the numerical simulations and experimental results has been demonstrated, e.g., see Fig. 3.13 (Kinsey and Cao 2001). The results for 3-D numerical simulations and analytical model (with and without the plane strain assumption) are presented in Table 3.3. Again, while the model provided reasonable predictions with the plane strain assumption, the elimination of the plane strain assumption provided better predictions compared to the 3-D FEA case.


62


3.12 Geometry for the 3-D non-symmetric, aluminum case (adapted from Bravar et al., 2007). All dimensions in mm.

3.13 Results from (a) numerical simulations (percent thickness reduction) and (b) experimental investigation of a non-symmetric aluminum TWB case (Bravar et al., 2007).



3.5

63

Determination of non-uniform binder force

He et al. (2001) used a 2-D analysis with a plane strain assumption to determine a non-uniform binder force ratio to control weld line movement for a TWB application. For their analytical model, the 2-D geometry was discretized into sections, from I to V (see Fig. 3.3). For their model, the draw-in of the material and the weld line movement were specified. This allowed the weld line movement desired for the process (e.g., no weld line movement) to be specified. Force equilibrium between sections and across the weld line was assumed. Force values were used to calculate the strain in the material at various locations in the 2-D geometry. The stress in the 1-direction at the edge of the blank is zero and was assumed to be linearly distributed within Section I. The normal punch contact pressure was ignored in Section IV. The effects of bending and friction were included using Eq. [3.14] and [3.15] respectively. The weld line movement was adjusted by using different blank holding forces (BHF) in the thicker and thinner sides of the sectioned strip, noted as BHF and BHF ′, respectively. The forces along the section (in the 1-direction) consist of the frictions from binder and die contact surfaces in an outward direction and punch friction forces that are inward to the neutral section. The two BHFs provide restraining forces (through the binder/die surface friction) that contribute to the force balance in the 1-direction of the strip. The desired BHFs in the thicker and thinner sides of the binder can be determined. For a sectioned strip under plastic stretching by a punch in the center portion and restrained by two binder regions at the two ends, the tension forces on the strip include the binder restraining force that is always in the outward direction, the die shoulder friction force that is also in the outward direction and the punch friction forces at various locations (punch side wall, shoulder radial corner and bottom surface) that are in the direction towards the neutral section, either located in one side of the side wall, denoted as Pattern 1, or on punch bottom surface, denoted as Pattern 2. Flow Pattern 1 may not occur in regular stamping with a homogeneous blank but it is possible with a TWB and with different blank holding forces applied on different binder regions, so that one side of metal may flow outward over the punch radius. From the calculated force values, the BHF can be calculated for the two flow patterns. For flow Pattern 1: [3.21] where TV is the tension force at the weld line (Section V), TB and TF are the tensions generated from bending and friction, respectively, in Section V of the geometry (see Fig. 3.3), µ is the coefficient of Coulumb friction at the interface between the blank and binder and µ1 is the friction at the interface between the blank and the die. Flow Pattern 1 can apply to either the thicker or thinner side of


64


the TWB combination. However, this pattern cannot co-exist in both sides. For flow Pattern 2: [3.22] See Fig. 3.3 for the locations of Sections II and IV on the 2-D cross-section. Accordingly, the difference in the two BHFs can be determined. For example, for the case where the neutral surface is on the punch bottom (i.e., a flow Pattern 2 condition), the non-uniform binder force can be calculated from: [3.23] where the prime marks designate the values on the thicker material side. From this analytical model, the required non-uniform binder force can be calculated as a trajectory of the binder forces on the thicker and thinner sides of the TWB versus punch displacement. As an example, a TWB case was considered with zero weld line movement, a draw-in on the thinner material side of 17.5 mm, a TWB thickness ratio of 1.5 mm to 0.7 mm, a width (W) of 50.8 mm, frictional coefficients of 0.1, n = 0.23, K = 648 MPa, length of Section I = 75 mm, length of Section II = 0.25 mm and RD and RP = 12.7 mm. Figure 3.14 shows the normalized BHF trajectories for this example. Note that a higher binder force ratio must be applied as the process progresses in order for the weld line movement to be eliminated for the application. The ratio for this case is approximately 1.7 at higher punch displacements. This case was implemented in a FEA software package, LS-DYNA3D to assess the predicted non-uniform binder force value from the analytical model. The weld line was modeled as a set of rigid nodes with six degrees of freedom. See He et al. (2001) for further details related to the numerical simulations. For simplicity, a constant non-uniform BHF ratio was employed in the FEA model as opposed to the BHF trajectories in Fig. 3.14. Figure 3.15 shows that a constant BHF difference

3.14 Normalized non-uniform blank holder force versus punch displacement trajectories from analytical model (He et al., 2001).



65

3.15 Effect of constant non-uniform binder force on weld line movement in numerical simulations (He et al., 2001).

between 1 and 2 tons prevents weld line movement for various BHF values on the thicker material side (BHF ′). This agrees well with the analytical model results. A similar geometry (Lo = 203.2 mm and W = 25.4 mm) was also investigated experimentally through strip drawing. Due to machine limitation, a BHF trajectory could not be employed, so constant BHF ratios were applied. Figure 3.16 shows that a constant non-uniform BHF ratio between 1 and 2 creates zero weld line movement in the process. In addition to the analytical model, numerical simulation and experimental strip drawing cases presented here, box drawing and a door inner geometry were also investigated in He et al. (2001). Reasonable agreement was again obtained. But similar to the results by Bravar et al. (2007), it is anticipated that elimination of the plane strain assumption would provide better predictions by the analytical model.

3.16 Effect of constant non-uniform binder force on weld line movement for strip drawing experiments and numerical simulations (He et al., 2001).


66


3.6

Conclusions

In this chapter, three analytical models to characterize the deformation and limitations during TWB forming were presented. Parameters such as the limiting thickness ratio, weld displacement, forming height and non-uniform binder force ratio to prevent weld line movement were determined. Such analytical models provide a valuable tool for engineers early in the design process. But these models include assumptions which adversely affect the predictions, e.g., a plane strain condition along the weld seam and constant frictional conditions over a surface, which may not physically exist during the process. Still, such models provide insight prior to costly numerical simulations or physical production. Other analytical models may also be of interest with respect to TWB applications (e.g., Davies et al. 2001, Cayssials et al. 2000 and Doege et al. 1996).

3.7

References

Bravar, M., Krishnan, N., and Kinsey, B.L. (2007), ‘Comparison of Analytical Model to Experimental Results and Numerical Simulation for Tailor Welded Blank Forming’, Journal of Manufacturing Science and Engineering, 129(1), 211–15. Brooks, A., Jhita, R., and Ni, C.M. (1991), ‘Applications of a Two-Dimensional Metal Forming Analysis Tool to Production Stamping Problems’, SAE Transactions, Paper No. 910775, 100(5), 751–8. Cayssials, F. (2000), ‘An Industrial Application of Specific Forming Limit Curves for Tailor Welded Blanks’, Proceedings of the 2000 International Deep Drawing Research Group, Ann Arbor, MI, June 17–22. Davies, R.W., Grant, G.J., Oliver, H.E., Khaleel, M.A., and Smith, M.T. (2001), ‘FormingLimit Diagrams of Aluminum Tailor-Welded Blank Weld Material’, Metallurgical and Materials Transactions A: Physical Metallurgy and Materials Science, 32(2), 275–83. Davies, R.W., Smith, M.T., Oliver, H.E., Khaleel, M.A., and Pitman, S.G. (2000), ‘Weld Metal Ductility in Aluminum Tailor Welded Blanks’, Metallurgical and Materials Transactions A: Physical Metallurgy and Materials Science, 31(11), 2755–63. Doege, E., Dohrmann, H., and Kosters, R. (1996), ‘Simulation and Optimisation of the Forming Process of Tailored Blanks’, Proceedings of Numisheet ’96, 199–204. He, S., Wu, X., and Hu, S.J. (2001), ‘Formability Enhancement for Tailor-Welded Blanks Using Blank Holding Force Control’, Journal of Manufacturing Science and Engineering, 125, 461–7. Kinsey, B.L. and Cao, J. (2001), ‘Numerical Simulations and Experimental Implementation of Tailor Welded Blank Forming’, Second Global Symposium on Innovations in Materials Processing & Manufacturing: Sheet Materials, 2001 TMS Annual Meeting. Kinsey, B.L. and Cao, J. (2003), ‘An Analytical Model for Tailor Welded Blank Forming’, Journal of Manufacturing Science and Engineering, 125(2), 344–51. Krishnan, N. and Cao, J. (2004), ‘Experimental Investigation and Comparison of Variable Blank-holder Force Trajectories Using an ARMA Model’, JUSFA 2004, Denver, Colorado, July 19–21. Saran, M.J., Keum, Y.T., and Wagoner, R.H. (1991), ‘Section Analysis with Irregular Tools and Arbitrary Draw-in Conditions for Numerical Simulation of Sheet Forming’, International Journal of Mechanical Science, 33(11), 893–909.



67

Shi, M.F., Pickett, K.M., and Bhatt, K.K. (1993), ‘Formability Issues in the Application of Tailor Welded Blank Sheets’, SAE Transactions, Paper No. 930278, 102(5), 27–35. Swift, H. W. (1948), ‘Plastic Bending under Tension’, Engineering, 166, 333–59 Viswanathan, V., Kinsey, B.L., and Cao, J. (2001), ‘Forming of Aluminum Tailor Welded Blanks’, SAE Paper No. 2000-01-0822, Proceedings of the Society of Automotive Engineers. Wang, N.M., and Wenner, M.L. (1978), ‘Elastic-viscoplastic Analysis of Simple Stretch Forming Problems’, in Koistinen, D.P., and Wang, N.M. (eds), Mechanics of Sheet Metal Forming, Plenum Press, New York, p. 367. Wood, R.D., Mattiasson, K., Honnor, M.E., and Zienkiewicz, O.C. (1985), ‘Viscous Flow and Solid Mechanics Approaches to the Analysis of Thin Sheet Forming’, in Wang, N.M., and Tang, S.C. (eds), Computer Modeling of Sheet Metal Forming Processes, AIME. Yao, H., Kinsey, B.L., and Cao, J. (2000), ‘Rapid Design of Corner Restraining Force in Deep Drawn Rectangular Parts’, International Journal of Machine Tools and Manufacture, 40(1), 113–31. Zhang, Z.T., and Hu, S.J. (1997), ‘Mathematical Modeling in Plane Strain Bending’, SAE Transactions, Paper No. 970439, 106(5), 63–76.


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

4 Numerical simulation modeling of tailor welded blank forming A. A. ZADPOOR, Materials Innovation Institute (M2i) and Delft University of Technology, the Netherlands and J. SINKE and R. BENEDICTUS, Delft University of Technology, the Netherlands

Abstract: Numerical modeling of TWBs is more complicated than the modeling of conventional sheet metal forming processes primarily due to the change of mechanical properties caused by the welding process and because the welded blanks are not uniform in thickness and/or material. Moreover, there are certain new issues such as the modeling of the weld line movement and determination of the blank holding forces (in the case of segmented-die forming) that are specific to TWBs. This chapter reviews the numerical modeling of TWBs and addresses several important issues in this regard. The topics covered include finite element method modeling of the weld zone, the material models used for numerical modeling of TWBs, theoretical failure techniques and their application to TWBs, and the design and optimization of TWBs. After making some concluding remarks, the anticipated trends of future research are discussed and the reader is referred to the related literature for further reading. Key words: numerical modeling, tailor welded blanks, finite element method, material models, theoretical failure prediction, optimal design.

4.1

Introduction

Tailor welded blanks (TWBs) are sheet metal assemblies that are composed of blanks with different materials, thicknesses, coatings, etc. The blanks are welded prior to the forming process that brings the assembly to its final structural shape. The additional endeavor needed for tailoring is justified, because air and ground vehicles need to make as efficient use of the structural material as possible to simultaneously minimize several cost functions. The most important cost function is probably the weight of the structure, which needs to be minimized without compromising (and often even improving) the structural integrity of the vehicle. However, other cost functions such as production cost, material cost and production time also play a role. The possibility of having different materials and thicknesses in one single assembly enables designers to optimally distribute the material within the structure and practically design the material for the component instead of selecting it. The success of any such optimal design is dependent on the possibility of predicting the forming behavior of TWBs. Reliable and feasible numerical modeling is the most important design tool that engineers need to have at their disposal. 68 © Woodhead Publishing Limited, 2011


69

Even though the numerical modeling of conventional sheet metal forming processes is still being researched on many fronts, reasonably accurate modeling of conventional processes is possible nowadays owing to several decades of research. However, the numerical modeling of TWBs is associated with some additional difficulties that distinguish it from the modeling of conventional processes. In this introductory section, we mention those difficulties. Some of the difficulties pointed out in this section will be reviewed in the following sections. Note that if a joining method other than welding is used to join sheets with dissimilar thicknesses and/or materials, the resulting assembly is called a tailor made blank (TMB) instead of a tailor welded blank. For more information see Chapter 9. The first difficulty originates from high welding temperatures the blanks experience during the welding processes that are applied for their joining. The high welding temperatures of fusions welding processes (Sharma and Molian, 2009; Zadpoor et al., 2007) and even the relatively lower temperatures of solid state joining processes such as friction stir welding (FSW) (Mishra and Ma, 2005) significantly change the mechanical properties of the blanks. Therefore, the mechanical properties of the weld seam and its surrounding area are different from those of the base metals, and the material of the blanks is no longer uniform. The non-uniform distribution of the mechanical properties within the blanks causes several difficulties in the numerical modeling of TWBs. First, experiments are needed to determine the local mechanical behavior of the different parts of the blanks. The yield strength, plasticity parameters (including the parameters of the strain hardening rule and anisotropy parameters) and tensile strength need to be determined through a sophisticated series of mechanical tests. Of particular importance is the need for specialized strain measurement techniques, because the normal strain measurement procedures that are based on uniform strain distribution do not work anymore. Even strain gauges are not able to provide fine enough spatial resolutions that are needed for characterization of the strongly heterogeneous material of the weld. The most successful and widely used strain measurement technique for measurement of the local mechanical properties of TWBs is the digital image correlation (DIC) technique. In the DIC technique, a random speckle ink pattern is applied on the surface of the blanks. The morphology of the ink pattern is continuously captured by one digital camera (or two cameras in case of threedimensional analysis) during the mechanical testing. The restriction is that the cameras should be able to ‘see’ the ink pattern, and that the ink pattern itself should not be damaged, blurred or rubbed during the process. Subsequently, local deformations are determined by a computer program that uses the captured digital photos to follow the change of the ink pattern during the test. Since the test machine is correlated with the cameras, the force corresponding to each digital photo can be read from the log file of the test machine and local stress–strain curves can be obtained. The required local mechanical properties can be extracted then from the local stress–strain curves. In addition to the mechanical properties,


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

70 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


the geometry of the weld seam may also have changed due to the welding process. Similar optical techniques can be used for determination of the geometry of the weld seam, particularly in the areas close to the thickness step, where the geometry of the weld can play an important role in the severity of the stress concentration. Once the local mechanical properties and the geometry of the weld seam are determined, one needs to decide about the level of the implementation of the weld details in the numerical model. On the one hand, implementation of too many details results in a complex numerical model that is not only difficult to build but is also expensive to run. On the other hand, implementation of too few details compromises the accuracy of the numerical model. The modeling of the weld line and the level of the weld details implementation is therefore one of the most important topics in the numerical modeling of TWBs. The next section of this chapter is devoted to that topic. The thermal evolution (thermomechanical evolution in the case of FSW) of the material during the welding process has more consequences. The material models that may have been appropriate for the modeling of the base metals may not be appropriate for the modeling of the welded blanks. As an example, one may consider the welding of high strength aluminum or steel alloys. In contrast to the conventional forming of (advanced) high strength aluminum and steel alloys, where ductile fracture seems to be the dominant fracture mechanism (Zadpoor et al., 2009b), instability may play an important role in the failure of welded TWBs. This is because of two reasons: first, sometimes the weld seam and heataffected zones (HAZ) of TWBs have substantially different microstructure and are significantly softer than the base metals. Therefore, the material close to the weld line may behave like more ductile materials where instability plays an important role in the failure of the blank. Second, there are too many variations in the mechanical properties of the different locations close to the weld line (Zadpoor et al., 2008b). The areas with lower strength play the role of imperfections that are assumed in the Marciniak–Kuczynski theory of instability (Marciniak and Kuczynski, 1967). Therefore, larger imperfections are present in TWBs as compared to the base metals. The two above-mentioned mechanisms increase the likelihood of instability acting as a major contributing factor to failure. Therefore, one may need to take into account not only ductile fracture but also instability in the numerical analysis of the failure of those TWBs. Moreover, welding may result in a drastically more porous material. More advanced models such as porous metal plasticity models may be needed for modeling of such a porous material. As a result of these changes, the type of material model used for numerical modeling of TWBs may be quite different from the material models which would have been used for the modeling of the base materials. The third section of this chapter focuses on the material models than can be used for the numerical modeling of both monolithic sheets and tailor welded blanks. As already discussed, the failure mechanism of TWBs can be partially or totally different from that of the base metals. Furthermore, the forming limit diagrams



71

(FLDs) of the weld seam, HAZs and the surrounding areas tend to be different from those of the base metals. Different modeling techniques may therefore be needed for a theoretical prediction of the straining limits of TWBs. The fourth section of this chapter will discuss the theoretical failure prediction of TWBs. Finally, the fact that TWBs need to be designed such that certain cost functions are optimized calls for new design and optimization approaches. The fifth section of this chapter will briefly review some topics in the design and optimization of TWBs.

4.2

Finite element method (FEM) modeling of the welded zones

One of the important challenges of the finite element method (FEM) modeling of TWBs is the modeling of the weld area. Two approaches are common in this regard. The first approach is to model the weld area accurately. In this approach, the geometry and properties of the weld and the HAZ are taken into account. Several researchers have adopted this approach in their studies (Raymond et al., 2004; Zhao et al., 2001; Chang et al., 2002; Chien et al., 2003; Iwata et al., 1995; Kampus and Balic, 2003; Reis et al., 2004; Hetu and Siegert, 2005; Jain, 2000; Roque et al., 2005b; Roque et al., 2005a). The second approach excludes weld properties and/or geometries from the FEM model. This approach is also extensively used in the literature (Jiang et al., 2004; Zhao et al., 2001; Ahmetoglu et al., 1995; Buste et al., 2000; Kampus and Balic, 2003; Meinders et al., 2000; Buste et al., 1999). Many other papers, in which the procedure of numerical modeling is not sufficiently described, have apparently used the second approach. While it is desirable to make the numerical model as accurate as possible, the reason for disregarding the weld line and the HAZ is either added computational cost or lack of experimental data for the mechanical behavior of the weld metal. The lack of material parameters should no longer be a problem, as a growing number of studies are dealing with measurement of the weld properties and new measurement techniques such as DIC are emerging. Therefore, the main debate would be the necessity of implementing the weld properties in the model. When the geometry and/or material properties of the weld area are excluded from the model, some kind of modeling technique is required to make the connection between the two different parts of the TWB. One may choose to use rigid links (spot welds) for the connection (Jiang et al., 2004; Buste et al., 2000). As far as steel TWBs are considered, it seems reasonable to use this method because, due to the high strength of the weld metal in steel TWBs, the failure is more likely to happen in the base metals. This modeling method is particularly efficient while dealing with laser welded blanks, because the weld area in laser welding is very narrow (about 1–2 mm). However, the model may not be efficient for other welding methods such as gas tungsten arc welding (GTAW), mash seam welding and FSW, which create a relatively wide weld. The method is also not


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

72 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


justified when aluminum TWBs are being modeled. It is because the weld area in the aluminum TWBs is sometimes weaker than the base metal and the failure is more likely to happen in the weld area. Therefore, a good description of the metal behavior near the weld line is necessary for aluminum TWBs. There is an important point to be noticed here. When modeling weld zone with shell elements that simply share nodes along the weld line, the modeled TWB is different than the physical TWB. That is because the modeled TWB of different thicknesses will share a common centerline (i.e. the thickness difference will be split between the top and bottom surfaces of the TWB). That is different from the physical TWBs that share a common surface (i.e., the thickness difference will be either on the top or bottom). This is due to the positioning of the blanks during welding. Several techniques are used to model the weld area in tailor welded blanks. One method is to use beam elements for representing the weld (Iwata et al., 1995; Nakagawa et al., 1993). However, beam elements limit the geometry that can be represented and the mesh refinement in the weld zone (Raymond et al., 2004). The second method uses shell elements to model both the weld zone and base metal (Zhao et al., 2001; Roque et al., 2005b). Nevertheless, shell elements are not able to describe the geometry of tailor welded blanks adequately. The third method is to model the weld zone by solid elements. Concerning this method, two strategies can be used. The first strategy uses solid elements for both the base metal and the weld zone (Zhao et al., 2001; Roque et al., 2005b). There is a high computational cost associated with this method, because several through-thickness solid elements are required for a good representation of the bending behavior in the forming simulation of the base metal (Raymond et al., 2004). The second strategy is to use solid elements for the weld zone and shell elements for the base metals (Raymond et al., 2004). The solid elements are constrained to move with the parent materials, meaning that movement of the dependent nodes is interpolated from the movement of a set of independent nodes on the base metal mesh. This strategy does not need a fine solid mesh for the base metal and can significantly reduce the associated computational load. In most studies, the geometry of the weld line is only roughly approximated; accurate representation of the weld geometry is rarely found (Chien et al., 2003). In that study, the geometry of the weld included two shallow notches. It was shown that the notches caused stress concentration and, ultimately, failure of the TWB. As previously stated, there is a debate on the value of the implementation of the weld area in the FEM model. Saunders reported that additional costs associated with the implementation of the weld area are not justified (Saunders and Wagoner, 1996; Saunders, 1994). Zhao et al. compared three models: one not implementing the weld area and two including the weld area (Zhao et al., 2001). They found that the solid model with the HAZ and the shell model with the HAZ increase the reaction forces by 25% and 7% respectively. However, implementation of the HAZ had little effect on the springback. The CPU time was increased from 0.615 hour for the simplest model to 16.26 hours for the three-dimensional solid element



73

model. Kampus and Balic used two different models: one without the weld area and one accounting for the hardening caused by the metal inert gas (MIG) welding (Kampus and Balic, 2003). They found that while the differences between the forming forces are minimal, the second model could give a better approximation of the actual shape of the deformed part. Raymond et al. concluded that though there are a number of differences between models with the weld area and models without the weld area (Raymond et al., 2004), the differences are subtle. Roque et al. examined four different models: combinations of shell or solid elements on one hand, and with or without HAZ on the other hand (Roque et al., 2005b). They showed that only the solid element with the HAZ could provide validated thickness distributions. Buste et al. used rigid links for making connections between adjacent nodes of thick and thin sheets, and found that the model is lacking accuracy in predicting the strain distribution near the weld line (Buste et al., 2000). Zadpoor et al. studied the FEM modeling of FSW TWBs (Zadpoor et al., 2009a) and investigated the effects of the implementation of the mechanical properties of the different weld zones (i.e. HAZ and weld nugget) on the strain distribution and springback behavior of the TWBs. It was revealed that the implementation of the mechanical properties of the weld zones is important for accurate prediction of the failure, strain field and springback behavior of the studied FSW TWB (Fig. 4.1). However, it was discovered that a smart choice of the implemented mechanical properties is necessary in order to keep the computational costs low. For example, for accurate prediction of the strain distribution, the mechanical properties of both HAZ and weld nugget had to be implemented in the FEM. However, implementation of the mechanical properties of the weld nugget was sufficient for accurate prediction of the springback behavior. Therefore, if one is only interested in the accurate prediction of the springback behavior of the FSW TWB, one may consider omitting the mechanical properties of the HAZ. An example of such a case is in the implementation of numerical algorithms that predict the springback of sheet metal parts and accordingly modify the die design to compensate for it. Nevertheless, the abovementioned assumption may not hold for all base materials and one needs to examine the validity of the assumption for the specific application at hand. In summary, whether the weld area should be implemented in the model is highly dependent on specifications of the problem. As a rule, the errors caused by excluding the weld area from the model are minimal when the weld line is located in the low strain regions and the FEM model is used for steel laser welded blanks. The errors are higher while dealing with materials such as high strength aluminum alloys and when the material near the weld line is expected to undergo significant deformation. Aside from the modeling of the weld zones, there are a few other considerations that are specific to the FEM modeling of TWBs. The first issue is the lack of symmetry in the FEM models of TWBs. One can benefit from symmetry in the numerical modeling of conventional sheet metal forming in order to decrease the size of the FEM model and to decrease the computational expense of the modeling.


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

74 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


4.1 Simulation results for a limiting dome height test of a FSW TWB are presented. The major (a) and minor (b) in-plane strains along the blank’s y-axis are presented. Model 1 corresponds to a mode in which both HAZ and weld nugget are excluded from the FEM model. Model 2 includes the mechanical properties of only the weld nugget. The mechanical properties of both zones are implemented in model 3. One can see that implementation of mechanical properties diacritically affects both strain distribution and failure prediction of this FSW TWB (reprinted from Zadpoor et al., 2009a; with permission from Elsevier).

For two reasons, that is not possible in many cases of TWBs. Firstly, the thickness and/or material changes from one side of the TWB to the other side and therefore the degree of symmetry of the part is significantly decreased. Secondly, even in the case of same-thickness and same-material combinations, there may be a considerable asymmetry in the distribution of the mechanical properties around the weld centerline (Fig. 4.2). One example is FSW where the material on the advancing side experiences different conditions as compared to the material on



75

4.2 Hardness profile around the weld centerline of a FSW TWB made of 7075-T6 aluminum alloy. The hardness values are asymmetric around the centerline. WN: weld nugget, HAZ: heat-affected zone, TMAZ: thermomechanically-affected zone (reprinted from Zadpoor et al., 2008b; with permission from Elsevier).

the retreating side (Zadpoor et al., 2008b). It has been shown that the asymmetry of the mechanical properties in FSW TWBs can result in considerable change of the forming behavior of the TWB (Zadpoor et al., 2009a). Another issue in the FEM modeling of TWBs is the meshing of the weld zone. Different meshing strategies may be needed in the areas close to the weld centerline. That is due to the change of the thickness and/or material at the transition line. A sufficiently fine mesh may be needed in order to capture the stress concentration that occurs due to the change of the thickness at the transition line. Moreover, the in-plane deformation may be accompanied by out-of-plane deformation (bending) (Zadpoor et al., 2008a; Zadpoor et al., 2009c), because the change of thickness at the transition results in a displacement of the neutral axis. Through-thickness stress gradient also contributes to the bending effect (Zadpoor et al., 2008a). The presence of any significant amount of bending can complicate the meshing, particularly in the case of solid elements, because several through-thickness elements are normally needed in order to capture the bending behavior (Zhao et al., 2001). Commercial FEM packages are widely used for modeling of the forming of TWBs. The punch is normally taken as a rigid body and Coulomb’s friction model is adopted. In most forming processes like stamping or bending, it is sufficient to adopt the constant friction coefficient for contact between the base metal and the tool. Tolazzi and Merklein showed that considering friction coefficients of the weld seam could considerably improve the accuracy of the FEM model (Tolazzi and Merklein, 2005).

4.3

Material models

Generally speaking, the same type of material models that are used for the numerical modeling of monolithic sheets can also be used for TWBs. In this


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

76 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


section, the most important material models that are commonly used for monolithic sheets and TWBs will be discussed. Some comments will be made about the specific applications of the material models in the previous studies of TWBs. The two most important models that are needed for the numerical modeling of forming processes are the yield function and the strain hardening law. The category of each of these two models can be subdivided according to various criteria. One such important criterion is the degree of the coupling of the material models with damage models. There are three possibilities in this regard. The first class is the class of material models that do not have any coupling with damage laws. Secondly, there may be some weak coupling between the damage criteria and the yield function. In that case, the damage model is simulated independently from the material models and the forming limits of the material are determined. The strains and/or stresses calculated through the FEM models are then compared with the forming limits predicted by the damage model and the onset of failure is detected based on that comparison. The problem of the second class is that the damage initiation and evolution are detached from the specific problem at hand. Therefore, such important aspects, such as the effects of the strain path (Hiwatashi et al., 1998) on the forming limits, are often neglected. The third class of the models, which is the most advanced class, is the one for which there is strong coupling between the damage and material models. For example, the yield surface may be dependent on the amount of damage sustained by the material during the deformation. The damage models that can be coupled with the material (plasticity) models will be discussed later in this chapter. The material models can be categorized also according to the nature of the model as being physical, semi-physical, or phenomenological. Physical models are based on the microstructural physics of the problem and generally cannot be expressed in a closed form. Some sort of averaging is often needed to determine the behavior of a reference volume element (RVE) of the material and connect the micro-scale to larger scales. Examples of the physical models are crystal plasticity and dislocation dynamics models. Semi-physical models are the models that are often formulated as closed-form functions describing the yield function or strain hardening law. The model is physical in the sense that the functional form of the model is determined through physical modeling. Once the functional form is derived, the relationship of the model with the physics is not in a direct manner anymore. The Bergström strain hardening model (Bergström, 1970) is an example of the so-called semi-physical models. The phenomenological models are the models that do not have much connection with the microstructural physics of the material but are instead based on the concepts and assumptions of the continuum mechanics. Phenomenological models are mostly given as closed-form functions. Widely used yield functions such as von Mises and Hill 1948, and the Hollomon, Swift and Nadai hardening laws are all examples of phenomenological models. The phenomenological models are the most widely used models in real practice due to their simplicity, smaller number of parameters, easier process of parameter



77

identification and drastically lower computational cost. Nevertheless, one needs to notice that the boundaries between these three categories are somewhat arbitrary. A model that is regarded as physical in one context and/or application may be regarded as phenomenological in another context. Generally speaking, there is hardly any purely physical model, and some degree of non-physical assumption needs to be made in order to make the models manageable from computational and/or parameter identification viewpoints. Table 4.1 presents a summary of a representative selection of the many numerical studies of TWBs. The type of material models (yield function and strain hardening law) as well as the base material and the welding technique are mentioned. The most frequently used material models (see Table 4.1) will be presented in the two following subsections of this section. For the reasons mentioned above, phenomenological models are the most widely used models for numerical modeling of TWBs. We will therefore limit our discussion to the phenomenological models. Table 4.1 A summary of some of the numerical studies of TWBs

Bayley and Pilkey, 2005

Yield

Hardening Material

Welding

GTN1

Voce

Aluminum (5754)

NVEB2 Laser

Seo et al., 2000a

Hollomon

Steel (SPC1)

Heo et al., 2001b

Hollomon

Steel (SPC1)

Hollomon3 Steel

Jiang et al., 2004

Hill 1948

Seo et al., 2000

von Mises Swift

Kim et al., 2000b Seo et al., 2000

Barlat

Raymond et al., 2004 Zhao et al., 2001

Swift, Hollomon

Steel (cold-rolled, hot-rolled, SAPH38P)

Hollomon

Aluminum (6111-T4)

Hollomon

Steel (AISI 1005)

von Mises Hollomon

Chang et al., 2002

Steel (SPC1)

Hollomon

Laser Laser Laser

Laser

Steel (SPCEN, SPRC)

Laser

Steel (SCP1)

Laser

Aluminum (5754)

NVEB, laser

Buste et al., 2000

Barlat

Shen et al., 2005

Hill 1948

Swift

Steel (JIS G3141)

Laser

Chien et al., 2003

Hill 1948

Voce

Aluminum (AA5754)

Laser

Kinsey and Cao, 2003

Hill 1948

Hollomon

Aluminum (5182-H00) Laser

Kinsey et al., 2004

Barlat 1989 Hollomon

Aluminum (5182-H00) Laser

Davies et al., 2001

Hosford

Davies et al., 2001

Hollomon

Aluminum (5182,5754) GTAW4

Hollomon

Aluminum (5182)

Hollomon

Steel (mild)

Laser

Steel (X5CrNi18-10, RRStW23, 20MnCr5, Ck15)

MIG6, laser

Iwata et al., 1995

Gotoh5

Kampus and Balic, 2003

von Mises Hollomon

GTAW

Continued


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

78 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


Table 4.1 Continued Yield (Meinders et al., 2000)

Hill 1948

(Reis et al., 2004)

Hill 1948

Hardening Material

Hollomon

Welding

Steel (HS and IF)

Laser

Steel (low carbon FeP06)

Laser Laser

(Tolazzi and Merklein, 2005)

Hollomon

Steel (FeP04, DP450)

(Dry et al., 2002)

Swift

Steel (0.8HSIF, 2.0DP) Laser

(Panda and Kumar, 2008)

Barlat

(Shen, et al., 2008)

Hollomon

Steel (IF)

Laser

Hollomon

Steel (CRDQ, DP800, annealed stainless 301)

Laser

Laser

(Shen et al., 2008)

Hill 1948

Hollomon

Steel (ST1605)

(Gaied et al., 2009)

Hill 1948

Hollomon

Steel (mild, HSLA)

Laser

(Lee et al., 2009)

Yld2000, Cazacu

Voce, Hollomon

Aluminum (6111, 5083), Steel (DP590), Magnesium (AZ31)

FSW7

(Padmanabhan et al., 2007) Hill 1948

Swift

Steel (DC06, DP600)

Laser

(Padmanabhan et al., 2007)

Hollomon

Aluminum (5182-O, 5454-O, 5754-O, 5052-O)

(Bhagwan et al., 2003)

Barlat

Hollomon

Steel (IF)

Laser

(Rodrigues et al., 2004)

Hill 1948

Hollomon

Aluminum (6016-T4, 5182-H111)

FSW

(Zadpoor et al., 2009a)

von Mises Hollomon

Aluminum (2024-T3)

FSW

1Gurson–Tvergaard–Needleman

(GTN) (Gurson, 1977; Tvergaard, 1982; Tvergaard, 1987). 2Non-vacuum electron beam (NVEB) welding. 3Wherever strain hardening and strength coefficients are presented without any reference to the specific strain hardening model, it is assumed that Hollomon’s hardening law is adopted. 4Gas tungsten arc welding (GTAW). 5For more information see Gotoh, 1985b; Gotoh, 1985a; Gotoh, 1985c. 6Metal inert gas (MIG). 7Friction stir welding (FSW).

4.3.1 Strain hardening laws In general, there are three types of hardening models, namely isotropic hardening, kinematic hardening and combined hardening. In case of isotropic hardening, the yield surface expands (contracts in the case of strain softening) as the material is strained. In case of kinematic hardening, the yield surface displaces due to the deformation. In the sheet metal forming community, kinematic hardening is often referred to as the Bauschinger effect. The combined hardening model is a combination of isotropic and kinematic hardening. Most numerical studies of



79

TWBs have neglected the kinematic hardening of the material during deformation. Whether or not the contribution of the kinematic hardening should be taken into account depends on the specific problem at hand and particularly on the deformation path of the TWB. One important case where kinematic hardening needs to be taken into account is when there is a significant change of the straining direction. In this chapter, we will only introduce isotropic hardening models due to their wider application in the numerical modeling of TWBs. Different isotropic hardening models can be used for FEM modeling of metal forming processes. However, Hollomon’s power law is the simplest and by far the most widely used model. Hollomon’s law can be expressed as: [4.1] where σ– and ε–p are, respectively, the equivalent stress and strain. The parameters K and n are called the strength and strain hardening coefficients, respectively. The Hollomon’s law gives a relatively good approximation of the actual stress–strain curve. A modification of the model is Ludwik’s law, which can be expressed as: [4.2] Ludwik’s law includes the yield stress, σy. Both equations have been extensively used in modeling of the metal forming processes including modeling of TWBs. Both the base and the weld metal can be approximated by these laws. Similar strain hardening laws are also used by some researchers. If the material needs to be considered as pre-strained, Ludwik’s law can be modified to give: [4.3] This is the generalized form of the Swift’s strain hardening law. The Voce law is an exponential law and is expressed as [4.4] Strain rate dependency of steel and aluminum are often omitted in FEM analysis of TWBs. However, one can simply take the strain rate effect into account by . multiplying the strain hardening law by a term like ε m.

4.3.2 Yield criteria Many different phenomenological yield criteria have been developed since 1950 in an attempt to accurately describe the yield behavior of engineering materials in general and metallic materials in particular. Many of those yield criteria were developed having the particular application of sheet metal forming in mind. Sheet metals exhibit significant anisotropy due to the roll forming process that is normally used in their production. Many yield criteria are developed such that they can capture anisotropy. Depending on whether or not a yield criterion can take the anisotropy into account, yield criteria are categorized as being either isotropic or anisotropic. Among


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

80 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


the yield criteria discussed in this sub-section, the only isotropic ones are von Mises and isotropic Gurson–Tvergaard–Needleman (GTN). The rest are anisotropic. von Mises yield criterion According to this criterion, a material yields when its second deviatoric stress invariant, J2, reaches a critical value: [4.5] where the second deviatoric stress invariant is given as: [4.6] and δij is the Kronecker delta. The summation convention of index notation applies throughout the chapter, unless otherwise stated. In the case of plane stress, the yield function can be written as: [4.7]

Hill’s 1948 yield criterion In a 1948 paper, Hill proposed that the yielding condition for an anisotropic material could be expressed as (Hill, 1948): [4.8] where F, G, H, L, M, and N are experimentally determined coefficients. In the case of plane stress, one can write the Hill’s 1948 yield criterion as: [4.9] where σy,1 is the uniaxial yield stress and R0 and R90 are the anisotropy parameters (Lankford coefficients) in the directions parallel with and perpendicular to the rolling direction respectively. The anisotropy parameters are related to the experimentally determined parameters of the yield locus as: [4.10] In the case of planar isotropy (i.e. normal anisotropy), the plane stress version of Hill’s 1948 yield criterion can be written as: [4.11]



81

One may notice that the last equation will reduce to the plane stress of von Mises yield function for an isotropic material, i.e. for R = 1. Hosford yield criterion Later, in 1979, Hill generalized his yield function as (Hill, 1979): [4.12] Several classes of yield criteria can be derived from this general formulation by assuming certain combinations of the experimental parameters. For a detailed description of the possibilities see Chu (1995). Hosford proposed his yield criterion independently from Hill (Hosford, 1972; Logan and Hosford, 1980). His yield criterion is, however, a special case of the generalized Hill criterion and can be obtained from that by assuming L = N = 0. For a planar isotropic material, we have F = G and L = M. The Hosford yield criterion is then obtained as: [4.13] If the material is not planar isotropic, the Hosford yield criterion can be given as: [4.14] Barlat yield criterion Barlat’s 1991 yield criterion (Barlat et al., 1991) is a generalization of the Hosford yield criterion for the case where the directions of the orthotropic axes are not coincident with the directions of the principal stresses: [4.15] where θ = arccos(I3/I23/2). The second and third invariants of the stress determinant (I2 and I3) are defined using the Bishop-Hill notation (Barlat et al., 1991) as follows: [4.16]

[4.17]


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

82 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


where A = σ22 − σ33, B = σ33 − σ11, C = σ11 − σ22, F = σ23, G = σ31, and H = σ12. The coefficients a, b, c, f, g, and h are weighting coefficients that represent the anisotropy of the material. All these weighting coefficients are equal to unity for an isotropic material and the yield criterion reduces to the Barlat’s isotropic yield function. Barlat and co-workers have developed several other widely used yield criteria, e.g. see (Barlat and Lian, 1989; Barlat, 1987; Barlat et al., 2003). Gurson–Tvergaard–Needleman (GTN) yield function Isotropic and anisotropic GTN yield functions are, respectively, generalizations of the von Mises and the Hill 1948 yield functions. These two yield functions belong to the class of yield criteria that are coupled with the damage law. The damage law in the case of GTN models is based on the physical ductile fracture mechanism, i.e. void nucleation, growth and coalescence. The whole population of the voids in the material is represented by the void volume fraction, ϕ (for a detailed description of the GTN models see Gurson (1977), Tvergaard (1982), Tvergaard (1987) and Zadpoor et al., 2009b)). In the isotropic case, the GTN yield function is given as:

[4.18] where σeq, σH and σy are the effective von Mises stress, hydrostatic pressure and yield stress of the fully dense matrix material, respectively. The other parameters, namely q1, q2 and q3, are experimentally determined material parameters. Liao et al. modified the GTN model to include the anisotropy according to the Hill 1948 yield criterion (Liao et al., 1997): [4.19] where R is the anisotropy parameter. The only other difference is that the equivalent stress is calculated according to the Hill 1948 yield criterion and not according to the von Mises criterion.

4.4

Theoretical failure prediction of tailor welded blanks (TWBs)

One of the challenges in the numerical modeling of TWBs is the theoretical formability prediction of TWBs. Theoretical formability prediction, if successful, can greatly facilitate the design process of all sheet metal forming processes, and TWBs are no exception. That is because an accurate theoretical prediction of the



83

forming limits removes the need for the elaborate, time consuming and expensive experiments that are otherwise needed for the determination of the forming limits. The theoretical failure prediction of monolithic sheets is already a difficult task, let alone the failure prediction of TWBs where different material and different thicknesses coexist in one single assembly and several failure mechanisms can potentially compete with each other. However, a significant number of studies are devoted to this topic, primarily because experimental determination of the forming limits of TWBs can be much more expensive than for monolithic sheets. As already discussed, many different zones with significantly different mechanical properties and forming limits exist in one single TWB. Therefore, not one FLD but several FLDs are needed to describe the failure of TWBs. In addition to the problem regarding the larger number of needed FLDs, there are at least three other difficulties that make the experimental determination of these FLDs formidably expensive. Firstly, the strain field needs to be determined with a very high spatial resolution so that the strains taking place in the multiple narrow zones around the weld line can be distinguished from each other. Secondly, various tests or various specimens with different geometry are needed to make sure that the FLDs of different zones can be determined. That is because a particular design of specimen is only useful for determination of the FLD of its least-formable zone. Thirdly, the forming limits determined for the different zones of TWBs tend to show much more scatter than for a monolithic sheet. The scatter is due to the uncertainties introduced by the welding procedure and the uncertainties introduced by the lowest accuracy in experimental determination of the forming limits of individual zones of TWB. Due to the above-mentioned difficulties, theoretical determination of the forming limits of TWBs is a very attractive topic. In general, theories that are used for the failure prediction of TWBs are the same as the ones that are used for monolithic sheets. We have recently reviewed the detailed formulation of these techniques elsewhere (Zadpoor and Sinke, 2010). In this chapter, we only review the main concepts and classifications of these techniques and give certain examples of their applications for theoretical formability prediction of TWBs. The theories used for prediction of the formability of TWBs can be divided into three main categories: variants of the Marciniak–Kuczynski theory, bifurcation techniques and ductile fracture theories. In the Marciniak–Kuczynski (M–K) theory of sheet metal instability (Marciniak and Kuczynski, 1967), it is assumed that there are pre-existing imperfections in the material that can be modeled by imperfection zones with smaller thicknesses (compared to the uniform sheet thickness). Due to the smaller thickness, the strain level in the imperfection zone is higher than the strain level of the uniform zone as the deformation takes place. At some point, the strain localizes in the imperfection zone and the ratio of the strain in the imperfection zone to that of the uniform zone increases rapidly to very large numbers. Once the ratio of the strain in the imperfection zone to that of the uniform zone reaches a presumed threshold,


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

84 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


the sheet is assumed to have failed. Several researchers have used the M–K theory for determination of the forming limits of gas tungsten arc welded (Davies et al., 2001) and FSW (Lee et al., 2009; Zadpoor et al., 2009a) TWBs. The second category of the failure prediction techniques is the category of the bifurcation theories. In these theories, the J2 deformation theory is applied for determination of the forming limits instead of the conventional flow theory. Stören and Rice used the vertex formation assumption to derive the forming limits of a monolithic sheet (Stören and Rice, 1975). According to that assumption, localized necking takes place once a vertex is developed on the subsequent yield surface and the bifurcation emerges from the state of uniform deformation. The bifurcation theories were later modified and extended by Zhu (Zhu et al., 2001) and Chow (Chow et al., 2003). The bifurcation theories have been used for failure prediction of laser welded steel TWBs (Cheng et al., 2007; Jie et al., 2007). The third category of the failure prediction techniques is the category of ductile fracture theories. These theories try to describe the failure based on the physical mechanism of ductile fracture, i.e. nucleation, growth and coalescence of the voids. The ductile fracture theories can be categorized as being either physical or phenomenological. Physical models of ductile fracture directly work with the voids and try to model their evolution during the deformation. Most of these models are based on the GTN material models which were briefly discussed in the section on yield criteria. The phenomenological models of ductile fracture do not directly work with the evolution of the voids and try to describe the ductile fracture mechanism by using continuum assumptions and concepts. In the phenomenological models of ductile fracture, it is often assumed that the ductile fracture takes place once a particular function of stresses and strains reaches a critical value. A presentation and discussion of many such phenomenological ductile fracture criteria can be found in Vallellano et al. (2007), Vallellano et al. (2006), Vallellano et al. (2008), Wierzbicki et al. (2005), Xue and Wierzbicki (2008), Xue and Wierzbicki (2009) and Zadpoor et al. (2009b). Bayley and Pilkey used GTN type models for the failure prediction of aluminum TWBs made by non-vacuum electron beam welding (Bayley and Pilkey, 2005; Bayley and Pilkey, 2006). Kim et al. used a phenomenological ductile fracture criterion (Oyane et al., 1980) to predict the bursting failure of a tailor welded tube made by high frequency electric resistance welding (HF-ERW) of rolled steel (Kim et al., 2004). The choice of failure prediction theory for a particular application depends on the dominant failure mechanism. Among the three above-mentioned categories of failure prediction techniques, the first two (M–K and bifurcation theories) are based on the assumption that the sheet metal fails due to instability. The ductile fracture theories, third category, assume that the failure is due to void nucleation, growth and coalescence. The theories based on the sheet instability may be used for the materials that predominantly fail due to instability such as draw quality mild steels and 5000 and 6000 series automotive aluminum alloys. The ductile fracture theories are more appropriate for the materials that fail predominantly



85

through ductile fracture such as (advanced) high strength steel and aluminum alloys. The two classes of models can be combined together to describe the failure of the materials that exhibit both types of failure mechanisms combined, such as some aluminum alloys (Zadpoor et al., 2009b). For example, Chien et al. combined M–K theory with GTN type models of ductile fracture for failure prediction of laser welded aluminum (Chien et al., 2003).

4.5

Some topics in design and optimization of TWBs

Successful design and manufacturing of TWBs involves optimal design of the TWB itself as well as proper modification of the design of the die set. Numerical techniques are often needed in order to achieve the optimal and even functional design of the TWB and die set. The current section is divided into two subsections, each of which covers the application of numerical modeling in one of these two areas.

4.5.1 Design of the die set The conventional design of the die set needs to be modified in order to successfully manufacture TWBs. The modifications are needed for two primary reasons. Firstly, the thickness step that exists in TWBs calls for modification of the punch such that the difference in the thickness can be properly accommodated. Therefore, the punch and binder may need a thickness step. Even though the thickness step complicates the design and, thus, the geometric modeling of the punch, it does not have serious consequences in terms of FEM modeling. The only major difficulty is that the mesh of the punch needs to be modified. The modification of the punch to accommodate the thickness step of the TWB makes the punch more complicated and expensive. Production techniques with a soft/flexible forming tool such as rubber forming are particularly attractive, because they do not need any such modification. Numerical modeling of processes with soft/flexible tools is generally more complicated and computationally expensive, as the forming tool cannot be considered rigid anymore. Secondly, the weld line movement, which takes place during the forming process of TWBs, has some consequences in terms of the design of the die set. The weld line movement can result in different modes of failure depending on the specifications of the problem (Kinsey and Cao, 2003). It is therefore needed to control the weld line movement. One of the effective ways for controlling the weld line movement is adjustment of the blank holding force. The adjustment of the blank holding force can be done by segmenting the die or by adding additional tooling to the die set (Hetu and Siegert, 2005; Kinsey et al., 1999a; Kinsey et al., 1999b; Kinsey et al., 2001; Siegert and Knabe, 1995). Some of these modifications apply different blank holding forces to different areas of a TWB and therefore need a model of the process. The model should be able to determine the distribution of the blank holding force such that the weld


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

86 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


line movement is minimized (or is otherwise brought under control). The modeling should therefore include an optimization process where the optimization algorithm tries to minimize the weld line movement by adjusting the blank holding force. The optimization algorithm needs to be connected to a FEM model of the forming process. The FEM model is also somewhat more complicated, because the die needs to be segmented with different blank holding forces applied to the different segments of the die. For each iteration of the optimization algorithm, the optimization algorithm determines the distribution of the blank holding force around the periphery of the blank. The FEM program then solves the model, determining the weld line movement. The optimization algorithm reads the FEM solution and, based on the amount of weld line movement, modifies the distribution of the blank holding force for the next iteration. Another solution for controlling the material flow is to apply drawbeads as local constraints (Heo et al., 2001a; Heo et al., 2001b). It is shown that the height and the size of the drawbeads play an important role in controlling the material flow (Heo et al., 2001a; Heo et al., 2001b). Again, one needs to connect the FEM model to an optimization algorithm so that the optimal height and size of the drawbeads can be determined.

4.5.2 Design of the TWB The main rationale for application of TWBs, despite their complexity and less formability, is their potential for lowering the weight of the structural parts of air and ground vehicles. As mentioned earlier, in the TWBs technology, the dilemma changes from selecting the material to designing it. One should therefore select the material, thickness and shape of the blanks that constitute the TWB such that the cost functions of the design are minimized and the constraints of the design are satisfied. The cost and constrain functions of the design can be categorized as being either related to the functionality or the manufacturing of the product. On the functionality side, the major cost function is the weight of the structure. The structure must meet certain other conditions as well. For example, it must exhibit the minimum level of stiffness that is determined according to the loading scenario. Moreover, the first natural frequency of the structure must be above a certain threshold. In most cases, the weight of the structure is taken as the cost function and the minimum stiffness and minimum natural frequency are applied as optimization constraints (Lee and Kang, 2007; Song and Park, 2006; Pan et al., 2010; Lee et al., 2003). There is also a vector of the optimization variables that includes two types of variables: the thickness and the geometrical parameters (i.e. size, shape, etc.) of the constituting blanks. If the constituting blanks are not only from different alloys but also from different materials (welding of different materials is possible using FSW, etc.), there will be a third category of optimization variables, i.e. the mechanical (e.g. elasticity modulus) and physical (e.g. density) properties of the constituting blanks.



87

On the manufacturing side, a concern is manufacturability, meaning that the production of the optimized TWB must be possible. There are two main concerns in this regard. Firstly, the strains must remain below the safety margin of the FLDs of the different zones of the TWB. Secondly, the product should meet the design tolerances after springback. Given the discrepancy in the springback behavior of the different constituting blanks of the TWB, meeting the design tolerances may be a difficult task to accomplish. The other parameters such as the production cost and production time also play a role, either as a cost function in a multi-objective optimization process or as a constraint to keep the production cost and production time below a certain (reference) limit. In addition to the above-mentioned criteria, one needs to take the weld line movement into account. Research has shown that the design parameters such as the initial location of the weld line (Heo et al., 2001a; Seo et al., 2000) and the blank shape (Seo et al., 2000) affect the weld line movement. Therefore, one may need to consider the weld line movement in the design of TWBs. The easiest way of formulating the optimization problem in the design of a TWB is to use one single cost function, i.e. weight of the structure, and apply all other important factors as optimization constraints. The optimization problem can be then formulated as: Find the vector of the design variables b(ti, lj, Ei, ρi, Ki, ni, Ri, . . .) to minimize the weight of the TWB; W(b), where ti, lj, ρi, and Ri are, respectively, the thickness, geometrical parameter, density and anisotropy coefficient of the i th blank. The optimization constraints can be divided into two categories: functional constraint and manufacturing constraints. The most important functionality constraints are (Song and Park, 2006): 1. Kzi = fi (K, stiffness matrix; zi, displacement vector; fi, loading vector). This is the linear relationship between the applied load scenario and the displacements of the TWB. 2. K(b)y − λM(b)y = 0 (λ, first natural frequency; y, first natural mode). This is the governing equation for determination of the natural frequencies of the TWB. 3. |TWBδi| − | Refδi | ≤ 0 (TWBδi, the displacement of the TWB at the point of the application of the i th load; Refδi, the reference values of the same displacement). For the TWB to be considered stiff enough, its displacements must be below some specified reference values. 4. λ ≥ λRef (λRef, the reference value of the first natural frequency). The first natural frequency of the TWB must be above a reference value to prevent unforeseen resonance. The most important manufacturing constraints are: 1. The constitutive equations of the plasticity model must be satisfied. 2. ε1,i < (1 − α)ε1,FLD, ε2,i < (1 − α)ε2,FLD (ε1,i, ε2,i , major and minor principal strains of the ith element (respectively); ε1,FLD, ε2,FLD, major and minor principal strains


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

88 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


of the FLD (respectively); α, safety margin). Since there are several different failure strains, this criteria would have to be satisfied from all combinations of major and minor strains. This check would have to be iterative. The forming strains must remain below the FLD by at least the specified safety margin. 3. Di,j ≤ RefDi,j (Di,j, RefDi,j, the vector of the springback displacement of the ith element in the j direction and the reference value of the vector, respectively). The springback displacements of the TWB must be below the acceptable level of springback that is dictated by the design tolerances. 4. ∆i,j ≤ Ref∆i,j (∆i,j, Ref∆i,j, the vector of the weld line movement of the ith element in the j direction and the reference value of the vector, respectively). The weld line movement of the TWB must be below the specified reference limit. Once the optimization problem is formulated, one needs to use an optimization algorithm to optimize the objective function. Some of the above-mentioned constraints may be relaxed for the sake of simplicity. In that case, separate rounds of checking may be required to make sure that the optimization results are feasible. Engineering judgment may replace the optimization process for some design variables. Various optimization techniques have been used for optimal design of TWBs (Lee and Kang, 2007; Song and Park, 2006; Pan et al., 2010; Lee et al., 2003; Zhu et al., 2008; Kim et al., 2000). In most cases, a subset of the optimization problem formulated above has been investigated with the vast majority of the researches concentrating on the functionality side of the optimization problem.

4.6

Conclusions

Numerical modeling of TWBs was reviewed in this chapter. The thermal (or thermomechanical in the case of FSW) evolution of the base material during the welding process results in a heterogeneous structure with varying mechanical properties, particularly in the areas in and around the weld line. One of the most important problems in the FEM modeling of TWBs is therefore the modeling of the weld zone. One needs to know whether or not the mechanical properties of the different weld zones need to be implemented in the FEM model. The review of the previous researches showed that the answer is strongly dependent on the specifications of the problem. For example, there may not be a need for implementation of the full details of the weld zone for a steel laser welded TWB. However, the mechanical properties of an FSW aluminum need to be implemented in the FEM model. Material models, theoretical failure prediction techniques and their application to the case of TWBs were also discussed. It was pointed out that essentially the same type of material and failure prediction models that are used for monolithic sheets can be used for TWBs. The only major difference is that there are a number of zones in a TWB which may need independent modeling. The optimal design of TWBs was also discussed. It was noticed that numerical modeling is needed for the optimal design of both the die set and the TWB. A



89

general case of the optimization was formulated and the related cost and constraint functions were presented.

4.7

Future trends

One of the general areas where further research is needed is the numerical modeling of the relatively new types of TWBs such as FSW TWBs and multiplematerial TWBs. With growing interest in these new areas, it is expected that more research funding will be directed towards studying the numerical modeling of the new types of TWBs. Accurate failure prediction of TWBs is also an important topic which needs further research, because accurate prediction of the forming limits of the different zones of TWBs is still not possible. Since experimental determination of the forming limits of TWBs is very difficult, it is expected that more attention will be paid to the numerical prediction of the FLDs. Application of more advanced material models, particularly the material models that are based on physical models, is also anticipated. The physically based models are particularly appropriate for the new types of TWBs such as the FSW aluminums, where there is a strong interaction between the forming behavior of the TWB and the thermomechanical evolution of the material during the welding process. Ultimately, the real advantage of TWBs is in their capability of reducing the structural weight and production cost of the vehicle through optimization of the material distribution and the production process respectively. So far, there has not been enough research on the optimal design of the TWBs. Our knowledge of the forming behavior of the TWBs seems to have reached the level that warrants a firm study of the optimal design of the TWBs. More research is therefore expected in this direction.

4.8

Sources of further information and advice

The interested reader is advised to browse through the issues of relevant journals such as the Journal of Materials Processing Technology, the International Journal of Plasticity, the Journal of Automobile Engineering (Proceedings of the Institution of Mechanical Engineers – Part D) and the Journal of Engineering Materials and Technology (Transactions of the ASME). For a review of the theoretical failure prediction of TWBs, see our recent book chapter on the subject (Zadpoor and Sinke, 2010). For a detailed explanation of the most widely used material models see the book edited by Banabic (Banabic, 2000). Tang and Pan (Tang and Pan, 2007) have recently published a useful book on the specific subject of the FEM modeling of the sheet metal forming process. The book by Dixit (Dixit and Dixit, 2008) is a recently published treatise on the general subject of the numerical modeling of the metal forming processes. The two above-mentioned books can be consulted for further information regarding the material models and numerical techniques discussed in this chapter.


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

90


4.9

Acknowledgments

The research was carried out under project number MC1.05224 in the framework of the Strategic Research Programme of the Materials Innovation Institute M2i (http://www.m2i.nl), the former Netherlands Institute for Metals Research. A part of the text of section 2 is taken from one of our papers (Zadpoor et al., 2009a) with permission from the publisher Elsevier.

4.10

References

Ahmetoglu, M. A., Brouwers, D., Shulkin, L., Taupin, L., Kinzel, G. L. et al. (1995) ‘Deep drawing of round cups from tailor welded blanks’, Journal of Materials Processing Technology, 53, 684–694. Banabic, D. (Ed.) (2000) Formability of Metallic Materials: Plastic Anisotropy, Formability Testing, Forming Limits, Berlin, Germany, Springer Verlag. Barlat, F. (1987) ‘Crystallographic texture, anisotropic yield surfaces and forming limits of sheet metals’, Materials Science and Engineering, 91, 55–72. Barlat, F., Brem, J. C., Yoon, J. W., Chung, K., Dick, R. E. et al. (2003) ‘Plane stress yield function for aluminum alloy sheets–Part 1: Theory’, International Journal of Plasticity, 19, 1297–1319. Barlat, F., Lege, D. J. and Brem, J. C. (1991) ‘A 6-Component yield function for anisotropic materials’, International Journal of Plasticity, 7, 693–712. Barlat, F. and Lian, K. (1989) ‘Plastic behavior and stretchability of sheet metals. Part I: A yield function for orthotropic sheets under plane stress conditions’, International Journal of Plasticity, 5, 51–66. Bayley, C. J. and Pilkey, A. K. (2005) ‘Influence of welding defects on the localization behaviour of an aluminum alloy tailor welded blank’. Materials Science and Engineering A, 403, 1–10. Bayley, C. J. and Pilkey, A. K. (2006) ‘A bifurcation criterion for predicting weld-line failures in AA5754 alloy tailor welded blanks’, Materials Science and Engineering: A, 435–436, 62–70. Bergström, Y. (1970) ‘A dislocation model for stress–strain behaviour of polycrystalline -Fe with special emphasis on variation of densities of mobile and immobile dislocations’, Materials Science and Engineering, 5, 193–200. Bhagwan, A. V., Kridli, G. H. and Friedman, P. A. (2003). ‘Formability improvement in aluminum tailor welded blanks via material combinations’, in Proceedings of NAMRC XXXI, SME paper MF03–155. Hamilton, Ontario, Canada. Buste, A., Lalbin, X., Worswick, M. J., Clarke, J. A., Altshuller, B., et al. (2000) ‘Prediction of strain distribution in aluminum tailor welded blanks for different welding techniques’, Canadian Metallurgical Quarterly, 39, 493–501. Buste, A., Lalbin, X., Worswick, M. J., Clarke, J. A., Altshuller, B., et al. (1999) ‘Prediction of strain distribution in aluminum tailor welded blanks for different welding techniques’, International Symposium on Light Metals as held at the 38th Annual Conference of Metallurgists of CIM; Quebec City, Quebec; Canada; 22–26 Aug. 1999. Canadian Institute of Mining, Metallurgy and Petroleum. Chang, S. H., Shin, J. M., Heo, Y. M. and Seo, D. G. (2002) ‘Springback characteristics of the tailor welded strips in U-bending’, Journal of Materials Processing Technology, 130, 14–19.



91

Cheng, C. H., Jie, M., Chan, L. C. and Chow, C. L. (2007) ‘True stress–strain analysis on weldment of heterogeneous tailor welded blanks – a novel approach for forming simulation’, International Journal of Mechanical Sciences, 49, 217–229. Chien, W. Y., Pan, J. and Friedman, P. A. (2003) ‘Failure prediction of aluminum laser welded blanks’, International Journal of Damage Mechanics, 12, 193–223. Chow, C. L., Jie, M. and Hu, S. J. (2003) ‘Forming limit analysis of sheet metals based on a generalized deformation theory’, Journal of Engineering Materials and Technology – Transactions of the ASME, 125, 260–265. Chu, E. (1995) ‘Generalization of Hill 1979 anisotropic yield criteria’, Journal of Materials Processing Technology, 50, 207–215. Davies, R. W., Grant, G. J., Khaleel, M. A., Smith, M. T. and Oliver, H. E. (2001) ‘Forminglimit diagrams of aluminum tailor welded blank weld material’, Metallurgical and Materials Transactions A, 32A, 275–283. Dixit, P. M. and Dixit, U. S. (2008) Modeling of Metal Forming and Machining Processes, London, UK, Springer Verlag. Dry, D., Waddell, W. and Owen, D. R. J. (2002), ‘Determination of laser weld properties for finite element analysis of laser welded tailored blanks’, Science and Technology of Welding & Joining, 7, 11–18. Gaied, S., Roelandt, J-M., Pinard, F., Schmit, F. and Balabane M. (2009), ‘Experimental and numerical assessment of tailor welded blanks formability’, Journal of Materials Processing Technology, 209, 387–395. Gurson, A. L. (1977) ‘Continuum theory of ductile rupture by void nucleation and growth, 1. Yield criteria and flow rules for porous ductile media’, Journal of Engineering Materials and Technology – Transactions of the ASME, 99, 2–15. Heo, Y., Choi, Y., Kim, H. Y. and Seo, D. (2001a) ‘Characteristics of weld line movements for the deep drawing with drawbeads of tailor welded blanks’, Journal of Materials Processing Technology, 111, 164–169. Heo, Y. M., Wang, S. H., Kim, H. Y. and Seo, D. G. (2001b) ‘The effect of the drawbead dimensions on the weld-line movements in the deep drawing of tailor welded blanks’, Journal of Materials Processing Technology, 113, 686–691. Hetu, L. and Siegert, K. (2005) ‘Hydromechanical deep drawing of tailor welded blanks’, Steel Research International, 76, 857–865. Hill, R. (1948) ‘A theory of the yielding and plastic flow of anisotropic metals’, Proceedings of the Royal Society of London Series A – Mathematical and Physical Sciences, 193, 281–297. Hill, R. (1979) ‘Theoretical plasticity of textured aggregates’, Mathematical Proceedings of the Cambridge Philosophical Society, 85, 179–191. Hiwatashi, S., Van Bael, A., Van Houtte, P. and Teodosiu, C. (1998) ‘Prediction of forming limit strains under strain-path changes: Application of an anisotropic model based on texture and dislocation structure’, International Journal of Plasticity, 14, 647–669. Hosford, W. F. (1972) ‘Generalized isotropic yield criterion’, Journal of Applied Mechanics, 39, 607–669. Iwata, N., Matsui, M., Nakagawa, N. and Ikura, S. (1995) Improvements in finite-element simulation for stamping and application to the forming of laser welded blanks’, Journal of Materials Processing Technology, 50, 335–347. Jain, M. (2000) ‘A simple test to assess the formability of tailor welded blanks’, International Journal of Forming Processes, 3, 185–212.


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

92 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


Jiang, H. M., Li, S. H., Wu, H. and Chen, X. P. (2004) ‘Numerical simulation and experimental verification in the use of tailor welded blanks in the multi-stage stamping process’, Journal of Materials Processing Technology, 151, 316–320. Jie, M., Cheng, C. H., Chow, C. L. and Chan, L. C. (2007) ‘Limit dome height and failure location of stainless steel tailor welded blanks’, Proceedings of the Institution of Mechanical Engineers Part C – Journal of Mechanical Engineering Science, 221, 1497–1506. Kampus, Z. and Balic, J. (2003) ‘Deep drawing of tailored blanks without a blankholder ’, Journal of Materials Processing Technology, 133, 128–133. Kim, J., Kim, Y.-W., Kang, B.-S. and Hwang, S.-M. (2004) ‘Finite element analysis for bursting failure prediction in bulge forming of a seamed tube’, Finite Elements in Analysis and Design, 40, 953–966. Kim, J. Y., Kim, N. and Huh, M. S. (2000) ‘Optimum blank design of an automobile subframe’, Journal of Materials Processing Technology, 101, 31–43. Kinsey, B., Liu, Z. and Cao, J. (1999a) ‘New apparatus and method for forming tailor welded blanks’, SAE Transactions: Journal of Materials & Manufacturing, 108, 653–660. Kinsey, B., Song, N. and Cao, J. (1999b) ‘Analysis of clamping mechanism for tailor welded blank forming’, SAE Transactions: Journal of Materials & Manufacturing, 108, 1062–1068. Kinsey, B., Viswanathan, V. and Cao, J. (2001) ‘Forming of aluminum tailor welded blanks’, Society of Automotive Engineers, SAE Transactions: Journal of Materials & Manufacturing, 110, 673–679. Kinsey, B. L. and Cao, J. (2003) ‘An analytical model for tailor welded blank forming’, Journal of Manufacturing Science and Engineering-Transactions of the ASME, 125, 344–351. Lee, K. H. and Kang, D. H. (2007) ‘Structural optimization of an automotive door using the kriging interpolation method’, Proceedings of the Institution of Mechanical Engineers Part D – Journal of Automobile Engineering, 221, 1525–1534. Lee, K. H., Shin, J. K., Song, S. I., Yoo, Y. M. and Park, G. J. (2003) ‘Automotive door design using structural optimization and design of experiments’, Proceedings of the Institution of Mechanical Engineers Part D – Journal of Automobile Engineering, 217, 855–865. Lee, W., Chung, K.-H., Kim, D., Kim, J., Kim, C., et al. (2009) ‘Experimental and numerical study on formability of friction stir welded TWB sheets based on hemispherical dome stretch tests’, International Journal of Plasticity, 25, 1626–1654. Liao, K. C., Pan, J. and Tang, S. C. (1997) ‘Approximate yield criteria for anisotropic porous ductile sheet metals’, Mechanics of Materials, 26, 213–226. Logan, R. W. and Hosford, W. F. (1980) ‘Upper-Bound anisotropic yield locus calculations assuming (111)-Pencil glide’, International Journal of Mechanical Sciences, 22, 419–430. Marciniak, Z. and Kuczynski, K. (1967) ‘Limit strains in the process of stretch-forming sheet metal’, International Journal of Mechanical Sciences, 9, 609–620. Meinders, T., Van Den Berg, A. and Huetink, J. (2000) ‘Deep drawing simulations of tailored blanks and experimental verification’, Journal of Materials Processing Technology, 103, 65–73. Mishra, R. S. and Ma, Z. Y. (2005) ‘Friction stir welding and processing’, Materials Science and Engineering R – Reports, 50, 1–78. Nakagawa, N., Ikura, S., Natsumi, F. and Iwata, N. (1993) ‘Finite element simulation of stamping a laser welded blank’, Sheet Metal and Stamping Symposium. Detroit, USA, Society of Automotive Engineers.



93

Oyane, M., Sato, T., Okimoto, K. and Shima, S. (1980) ‘Criteria for ductile fracture and their applications’, Journal of Mechanical Working Technology, 4, 65–81. Padmanabhan, R., Oliveira, M. C., Alves, J. L. and Menezes, L.F. (2007), ‘Influence of process parameters on the deep drawing of stainless steel’, Finite Elements in Analysis and Design, 43, 1062–1067 Pan, F., Zhu, P. and Zhang, Y. (2010) ‘Metamodel-based lightweight design of B-pillar with TWB structure via support vector regression’, Computers & Structures, 88, 36–44. Panda, S. K. and Kumar, D. R. (2008), ‘Improvement in formability of tailor welded blanks by application of counter pressure in biaxial stretch forming’, Journal of Materials Processing Technology, 204, 70–79. Raymond, S. D., Wild, P. M. and Bayley, C. J. (2004) ‘On modeling of the weld line in finite element analyses of tailor welded blank forming operations’, Journal of Materials Processing Technology, 147, 28–37. Reis, A., Teixeira, P., Duarte, J. F., Santo, A., Da Rocha, A. B. et al. (2004) ‘Tailored welded blanks – an experimental and numerical study in sheet metal forming on the effect of welding’, Computers & Structures, 82, 1435–1442. Rodrigues, D. M., Menezes, L. F., Loureiro, A. F. and Fernandes, J.V. (2004), ‘Numerical study of the plastic behaviour in tension of welds in high strength steels’, International Journal of Plasticity, 20, 1–18 Roque, A. P., Jorge, R. M. N., Parente, M. P. L., Valente, R. A. F. and Fernandes, A. A. (2005a) ‘Influence of the heat affected zone on hydroforming with tailor welded tubular blanks’ in Onate, E. and Owen, D. R. J. (eds.) VIII International Conference on Computational Plasticity, Barcelona. Roque, A. P., Jorge, R. M. N., Parente, M. P. L., Valente, R. A. F. and Fernandes, A. A. (2005b) ‘Numerical study of hydroforming with tailor welded tubular blanks’, NUMISHEET 2005: 6th International Conference and Workshop on Numerical Simulation of 3D Sheet Metal Forming Processes. Detroit, USA, American Institute of Physics. Saunders, F. I. (1994) ‘Forming of tailor welded blanks’, Department of Mechanical Engineering, Ohio State University. Saunders, F. I. and Wagoner, R. H. (1996) ‘Forming of tailor welded blanks’, Metallurgical and Materials Transactions A, 27A, 2605–2616. Seo, D., Choi, Y., Heo, Y. and Kim, H. Y. (2000) ‘Investigations of weld-line movements for the deep drawing process of tailor welded blanks’, Journal of Materials Processing Technology, 108, 1–7. Sharma, R. S. and Molian, P. (2009) ‘Yb:YAG laser welding of TRIP780 steel with dual phase and mild steels for use in tailor welded blanks’, Materials & Design, 30, 4146–4155. Siegert, K. and Knabe, E. (1995) ‘Fundamental research and draw die concepts for deep drawing of tailored blanks’, SAE Transactions: Journal of Materials & Manufacturing, 104, 866–876. Song, S. I. and Park, G. J. (2006) ‘Multidisciplinary optimization of an automotive door with a tailored blank’, Proceedings of the Institution of Mechanical Engineers Part D – Journal of Automobile Engineering, 220, 151–163. Stören, S. and Rice, J. R. (1975) ‘Localized necking in thin sheets’, Journal of the Mechanics and Physics of Solids, 23, 421–441. Tang, S. C. and Pan, J. (2007) Mechanics Modeling of Sheet Metal Forming, Washington, USA, SAE International. Tolazzi, M. and Merklein, M. (2005) ‘Precise material properties as a prerequisite for FE-analysis of the hydroforming of tailored welded blanks’, Steel Research International, 76, 915–919.


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

94 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


Tvergaard, V. (1982) ‘Material failure by void coalescence in localized shear bands’, International Journal of Solids and Structures, 18, 659–672. Tvergaard, V. (1987) ‘Effect of yield surface curvature and void nucleation on plastic flow localization’, Journal of the Mechanics and Physics of Solids, 35, 43–60. Vallellano, C., Guzman, C. and Garcia-Lomas, F. J. (2007) ‘Prediction of ductile failure in the stretch-forming of AA2024 sheets’, AIP Conference Proceedings, 908, 135–140. Vallellano, C., Guzman, C. and Garcia-Lomas, J. (2006) ‘Failure prediction in stretched sheets of aluminium 2024-T3’, Materials Science Forum, 526, 91–96. Vallellano, C., Morales, D. and Garcia-Lomas, F. J. (2008) ‘A study to predict failure in biaxially stretched sheets of aluminum alloy 2024-T3’, Materials and Manufacturing Processes, 23, 303–310. Wierzbicki, T., Bao, Y. B., Lee, Y. W. and Bai, Y. L. (2005) ‘Calibration and evaluation of seven fracture models’, International Journal of Mechanical Sciences, 47, 719–743. Xue, L. and Wierzbicki, T. (2008) ‘Ductile fracture initiation and propagation modeling using damage plasticity theory’, Engineering Fracture Mechanics, 75, 3276–3293. Xue, L. and Wierzbicki, T. (2009) ‘Numerical simulation of fracture mode transition in ductile plates’, International Journal of Solids and Structures, 46, 1423–1435. Zadpoor, A. A. and Sinke, J. (2010) ‘Weld metal properties and their influence on the formability of tailor welded blanks’ in Sun, X. (ed.) Failure mechanisms of advanced welding processes, Cambridge, UK, Woodhead Publishing. Zadpoor, A. A., Sinke, J. and Benedictus, R. (2007) ‘Mechanics of tailor welded blanks: an overview’, Key Engineering Materials, 344, 373–382. Zadpoor, A. A., Sinke, J. and Benedictus, R. (2008a) ‘Experimental and numerical study of machined aluminum tailor made blanks’, Journal of Materials Processing Technology, 200, 288–299. Zadpoor, A. A., Sinke, J. and Benedictus, R. (2008b) ‘The mechanical properties and microstructure of friction stir welded tailor made blanks’, Materials Science and Engineering A, 494, 281–290. Zadpoor, A. A., Sinke, J. and Benedictus, R. (2009a) ‘Finite element modeling and failure prediction of friction stir welded blanks’, Materials & Design, 30, 1423–1434. Zadpoor, A. A., Sinke, J. and Benedictus, R. (2009b) ‘Formability prediction of high strength aluminum sheets’, International Journal of Plasticity, 25, 2269–2297. Zadpoor, A. A., Sinke, J. and Benedictus, R. (2009c) ‘The mechanical behavior of adhesively bonded tailor made blanks’, International Journal of Adhesion and Adhesives, 29, 558–571. Zhao, K. M., Chun, B. K. and Lee, J. K. (2001) ‘Finite element analysis of tailor welded blanks’, Finite Elements in Analysis and Design, 37, 117–130. Zhu, P., Shi, Y. L., Zhang, K. Z. and Lin, Z. Q. (2008) ‘Optimum design of an automotive inner door panel with a tailor welded blank structure’, Proceedings of the Institution of Mechanical Engineers Part D – Journal of Automobile Engineering, 222, 1337–1348. Zhu, X. H., Weinmann, K. and Chandra, A. (2001) ‘A unified bifurcation analysis of sheet metal forming limits’, Journal of Engineering Materials and Technology – Transactions of the ASME, 123, 329–333.


5 Lightweight metal alloy tailor welded blanks R. PADMANABHAN, VIT University, India and University of Coimbra, Portugal and M.C. OLIVEIRA and L. F. MENEZES, University of Coimbra, Portugal

Abstract: Demand for lightweight materials in the automotive, aerospace, construction and other related industries has increased multifold in recent years due to environmental concerns, government regulations and consumer demand. Lightweight metal alloys are the preferred choice of these industries due to their low density and high specific strength, as well as other attractive features such as corrosion resistance, dimensional stability, etc. Lightweight metal alloy tailor welded blanks are widely considered nowadays for their functionality and aesthetics. With recent advances in the relevant manufacturing technologies, production of these tailor welded blanks has become feasible at low cost. The salient features and functions of different lightweight metal alloys and their role in tailor welded blank manufacturing are discussed in this chapter. Key words: lightweight metal alloys, aluminum, magnesium, titanium, tailor welded blank.

5.1

Introduction

The need for energy-efficient, high performance systems for propulsion and transportation is driving the development and use of lightweight metal alloys (LWMA). Lightweight materials such as aluminum, magnesium, titanium, beryllium, composites, etc., have received much wider attention over the past few years due to their high strength-to-weight ratio/specific strength. In addition to their excellent stiffness and corrosion resistance, other attractive features of lightweight materials include damping capacity, thermal properties, fatigue properties, dimensional stability, etc. The use of beryllium is limited due to related health hazards, although it has a very high stiffness to weight ratio and a high heat absorption capacity.

5.1.1 Aluminum Aluminum is available in abundance on the Earth’s crust, largely in the form of cryolite or bauxite. Aluminum alloys are widely used in day-to-day life as thin foils in food packages and beverage cans, as structural members in public transport systems, aircraft parts, etc. This wide range of applications is possible due to properties such as superior corrosion resistance, natural and chemical inertness, recyclability and the ease with which a variety of parts can be produced. Aluminum also has high thermal and electrical conductivity, emissivity, strength-to-weight ratio, fracture toughness, energy absorption capacity, cryogenic toughness, fatigue strength, 97 © Woodhead Publishing Limited, 2011

1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

98 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


etc. It does not oxidize progressively because a hard, microscopic oxide coating forms on the surface and protects the metal from corrosive environments. With additional electro-coating, either anodic or cathodic, greater protection can be achieved in applications where the metal comes in direct contact with a corrosive environment. The strength of aluminum can be increased by alloying it with manganese, silicon, copper, magnesium, zinc, etc. Therefore aluminum grades are identified based on the alloying element and heat treatment, using a four digit representation, from 1XXX to 8XXX. The 1XXX series of aluminum has low yield strength while the 5XXX and 6XXX series have yield strengths equivalent to mild steel (Fig. 5.1a) and 7XXX series yield strengths are equivalent to high strength steels (AluMatter, 2010). Figure 5.1b shows the strengths of aluminum alloys of

5.1 (a) Comparison of Al 5083-H34 and mild steel strengths; (b) aluminum alloy strengths according to different grades (AluMatter, 2010).


Lightweight metal alloy tailor welded blanks

99

different grades. Cold working also enhances the strength of aluminum to approximately double. The main advantage of using aluminum is the ease with which it can be recycled to produce high quality parts (Abu-Farha and Khraisheh, 2008). Applications The bodies of beverage cans are made of aluminum alloy (Al) 3004, while the ends are made of Al 5182, making it the largest volume alloy combination in the industry. Al 5454 has been widely used for rail car body construction for heavy load applications, and Al 5083/5383 has also been used in high speed single-hull ships, like the ‘Proserio’. Al–Mg alloys such as Al 5454, Al 5086, Al 5083/5383 are used for the welded structures of offshore oil rigs and platforms applications to protect against high humidity and salt water exposure. In the automotive industries, Al 5754 has been used for body-in-white and Al 6111-T4 is used for external body panels. Al 7XXX series are used in guard rail and truck bumpers due to their toughness. Current applications of Al- tailor welded blanks (TWB) include body parts for high performance cars, such as Lamborghini Gallardo, and generally in the manufacture of bonnets, front door inners, rear rail inners, body side outers, etc. With recent advances in aluminum welding technologies, applications of Al-TWB are being extended to aircraft structures and body panels (Schubert et al., 2001).

5.1.2 Magnesium The reasons for using magnesium in parts manufacture are twofold: first, magnesium is the lowest density (1.74 gm/cm3) metal available for engineering use; secondly, it is easy to obtain and there is an abundant supply of magnesium ore to the industry. The significant properties of magnesium for structural applications are its low density, high damping capacity, highest machinability index, high corrosion resistance in an alkaline environment and the fact that it is inert to chromic and hydrofluoric acids. However, magnesium is less resistant to acidic or salt-laden environments. Presently, magnesium is widely used in applications that utilize its chemical and metallurgical properties. Magnesium-lithium alloy is the lightest structural alloy available; hence, it can potentially be used in automotive, aircraft, aerospace, marine and other applications that necessitate weight saving in order to reduce fuel consumption. The magnesium alloy has low ductility at room temperature due to the small number of slip systems available, whereas its ductility is enhanced at elevated temperatures between 100 and 300°C (Fig. 5.2). Therefore, warm and hot stamping processes are commonly used to deform magnesium alloy blanks. In the warm and hot stamping, formability is improved by the local heating of the flange portion by means of a die and blank holder with a built-in heater (Zhang et al., 2006; Yoshihara et al., 2003a). However, the cold deep drawing of magnesium alloy blanks is made possible by annealing the sheet for one hour at 500°C, provided oxidation is controlled (Mori and Tsuji, 2007). Cold forming of


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

100 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


5.2 Stress–strain curve of magnesium at different temperatures (Chen and Huang, 2003).

magnesium can be used to produce shallow parts such as the shells of lightweight laptop computers, music players, etc. Strong crystallographic textures evolve during rolling, such that the majority of grains are oriented with their basal planes in the sheet plane. These strong textures are responsible for an intrinsic plastic anisotropy, low elongation and limited formability. Grain refinement, which promotes grain boundary sliding and controlled texture, can improve the room temperature formability of magnesium. The addition of lithium to magnesium exhibits a remarkable improvement in room temperature ductility owing to the enhanced activity of non-basal slip, particularly, slip mode. Furthermore, the addition of lithium to magnesium is believed to reduce the plastic anisotropy, typical of commercial magnesium alloys (Al-Samman, 2009). Weak texture, which promotes sheet formability, can be achieved in magnesium alloy sheets by the addition of rare earth elements such as Ce, Nd and Y (Hantzsche et al., 2010). Applications Magnesium, in the current state of technology, is used to make laptop casing, mobile phone casing, camera parts, sporting goods, etc.

5.1.3 Titanium Titanium alloys offer exceptional strength-to-weight ratio, mechanical properties and corrosion resistance, are biocompatible and have non-magnetic properties. A distinct advantage of titanium over other LWMAs is that it retains its strength at



101

moderate service temperatures. Its range of operation is from cryogenic temperatures to around 540°C. An important characteristic of titanium-base materials is the reversible transformation (or allotropy) of their crystalline structure from an alpha (α) (hexagonal close-packed) structure to a beta (β) (body-centered cubic) structure when the temperature exceeds a certain level. Diverse microstructures and properties can be imparted in the material by varying the alloy contents, hence the allotropy, and thermo-mechanical processing. Until now, titanium alloys have been excluded from applications involving large volume production due to high production costs. However, considering the reliable and more durable systems and components, which in many situations have substantially exceeded performance and extended service life expectations, titanium alloys have been preferred over other materials. Applications Titanium is used in critical applications for the aerospace, marine, food processing, nuclear power, sports equipment and medical (implants and prostheses) industries.

5.1.4 Metal matrix composites Applications such as aircraft parts need high strength per unit of weight. The requirement for such parts can be satisfied by metal matrix composite (MMC) materials consisting of a hard reinforcement, such as fiber, particulates, whiskers or platelets, in a soft matrix such as aluminum, magnesium or titanium. MMCs offer a superior strength-to-weight ratio with wear resistance and low coefficient of thermal expansion providing the necessary characteristics to produce lightweight and dimensionally stable structures. Continuous fiber and particulate reinforced MMCs are attractive in many applications that require performance rather than low manufacturing and material costs. Applications Structural applications such as airframes for hypersonic aircraft, automobile, marine and gas turbine components; sports equipment and medical devices are some potential applications for MMC TWBs.

5.2

Lightweight metal alloy tailor welded blanks (LWMA TWBs)

Environmental concerns, fast depletion of fossil fuels and increased customer demand for improved fuel efficiency have forced the automotive industry to implement strict government regulations. Thus, the industry is driven to look for high performance and lightweight materials in order to improve fuel efficiency, while also improving the aesthetics and reducing the cost of vehicles. One way of


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

102 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


handling this issue is to increase the application of TWBs in the production of automotive body parts. Through the proper location of appropriate materials by tailored welding, the required function of a part can be accomplished at reduced weight and nominal cost. TWBs are made of similar or dissimilar grade materials, and uniform or non-uniform thickness sheet materials that have potential applications in the automotive, aerospace, construction and other industries. In both similar or dissimilar TWBs, the welding and the difference in grades and/or thickness of the blanks cause differential material flow during forming processes. Also, the presence of a weld line, with different material properties compared to the base material, results in the reduced formability of the TWB. The main difference between steel TWBs and LWMA TWBs is in the weld metal properties. Increased mechanical properties are observed at the weld metal in steel TWBs (Boronski, 2006) while reduced mechanical properties are observed in LWMA TWBs (Stasik and Wagoner, 1996). The weld quality in aluminum TWB can be improved by adding filler materials such as silicon and magnesium (Braun, 2006). In the following sections, the defects and formability characteristics of LWMA TWBs with similar and dissimilar material properties and thicknesses will be discussed.

5.2.1 Similar material TWBs LWMA TWBs with blanks of similar materials and dissimilar thicknesses are widely used in applications that require different strengths at different sections and proper weight distribution. This change in thickness creates a discontinuity in the TWB that alters the stress state and consequently the strain that occurs during forming (Cheng et al., 2007). The stress concentration becomes responsible for failures in the transition zone. Furthermore, the limit strain decreases and is governed by the minimum thickness blank. For example, the weld line movement reduces with decreasing thickness mismatch (Heo et al., 2001). Fewer defects arise with decreasing mismatch in the TWBs and these defects occur mostly due to irregular straining during the forming process. The material flow during the forming process is also affected by the blank anisotropy. The anisotropy, which is prevalent in the pre-processed sheet segment, influences the subsequent deformation and failure patterns of tailor welded blanks. For instance, Al 6111-T4 and Al 5754-0 TWBs exhibit adequate ductility when loaded along the weld line, given sufficient specimen width. Al 5754-0 performs much better than Al 6111T4 under transverse loading, because of the softening of the heat-affected zone in heat-treatable Al 6111-T4 (Stasik and Wagoner, 1996).

5.2.2 Dissimilar material TWBs Dissimilar material TWBs refers to TWBs consisting of different families of materials, for example, various LWMAs tailor welded with steels – one of the



103

most widely used materials in structural applications. Vast applications of these TWBs are currently used by TWB parts manufacturers in transportation and propulsion industries and largely consist of joining aluminum, magnesium and titanium to steel; aluminum superstructures to steel hulls for ships; aluminum or magnesium to steel TWBs for automotive components; titanium to steel sheets for aerospace parts; and many other emerging applications are dependent on dissimilar material joints for reducing the total component weight and/or improving performance. Figure 5.3 shows a typical application of a TWB in a car door inner panel. In this application, rigidity at the hinge section and light weight at other sections are necessary for light weight and longer service life. Even in this lighter section, a dissimilar gauge sheet can be used, with a thick sheet at the bottom and a thin sheet at the window portion. As mentioned earlier, the formability of dissimilar-gauged TWBs is mainly affected by the thinner part so that the forming limit level decreases as the thickness ratio increases. In such a case, the thicker blank can be replaced by a higher strength material of different composition. Dissimilar material joining presents challenges significantly different to those explored during similar material joining, due to the difference in chemical, mechanical and thermal properties of component blanks. Joining material combinations such as aluminum and steel pose a number of problems viz., formation of brittle intermetallic compounds, poor wetting behavior, difference in physical and chemical properties of the base metals, etc. Since the melting temperatures of LWMAs and steel are quite different, conventional fusion welding processes do not yield a sound joint in this case of dissimilar metals joining. In

5.3 Typical application of a dissimilar material TWB.


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

104 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


addition, fusion welding produces a high concentration of intermetallic compounds that are detrimental to the joint. Friction stir welding (FSW) produces a solid state bond between the base metals through the frictional heat generated between the tool and the base metals. Thus, the formation of intermetallic compounds can be limited to produce a sound weld.

5.2.3 Different manufacturing techniques The formability of LWMA TWBs is highly influenced by the welding method used to produce the TWB, especially under high straining regimes such as bending and stretching modes (Panda and Kumar, 2008; 2009; 2010). The welding process parameters also play a major role in determining the mechanical properties of the weld as well as in the TWB’s formability. Thus, the choice of a welding method to produce TWBs depends on its application, aesthetic requirements, cost, formability concerns, etc. There are a variety of welding processes in use to produce TWBs including laser, FSW, resistance mash seam welding, electron beam, roll bonding, etc., each producing unique characteristics in the TWB. Laser welding In laser welding, a laser beam is focused and irradiated on the specimen to generate heat. Laser wavelength, power and welding speed can be varied to generate the required heat in the specimen. A steep temperature gradient is obtained in the weld metal resulting in a refined grain micro-structure with improved properties. Laser welding features keyhole penetration, resulting in the rapid solidification of the weld metal. A slight concavity results at the joint which can be reduced using filler material or beam weaving, which in turn will result in a wider heat affected zone (HAZ) (Auto/Steel, 1995; Ribic et al., 2009). With suitable power output, a two beam Nd: YAG laser process can be used to weld dissimilar thickness blanks, treating thick and thin sheets separately. The same approach can also be used to weld dissimilar material combinations. However, intermetallic compounds form at the weld region as a result of high heat input and fusion welding, thus affecting the formability of the TWB. In addition, common defects such as hot cracking, porosity, loss of alloying elements and grain boundary melting in the HAZ arise from the laser beam welding of aluminum alloys (Pastor et al., 2000; Pastor et al., 2001). In aluminum alloy welding, the laser beam reflection (the polished surface of aluminum will behave like a mirror) is a major concern, as it reduces the efficiency of the process. The laser beam reflectivity of Al 5052 aluminum alloy reduces drastically as the material changes from solid state to molten state (Wang et al., 2007). In recent years, laser absorption has been improved by increasing the power density of the focused spot, which is achieved through higher average power, improved beam focusing system and decreased beam reflectivity on the



105

workpiece surface. The main advantage of laser welding processes is the narrow weld line produced (e.g. 5 mm). Friction stir welding (FSW) The recently developed FSW process is a solid-state joining process that was invented at The Welding Institute (TWI) in 1991 (Thomas et al., 1991). FSW produces a solid state bond between the base metals through the frictional heat generated between the tool and the base metals. The heat generated at the stir zone is far less than that produced in a fusion welding process. Thus, a sound weld between soft light metal alloys and strong steel can be produced by limiting the formation of intermetallic compounds (Manesh and Taheri, 2003; Bach et al., 2005). FSW produces a fine re-crystallized grain structure in the stir zone and the welds exhibit better tensile, bend and fatigue properties than fusion welds (Watanabe, 2003; Chen and Kovacevic, 2004; Elrefaey et al., 2005; Uzun et al., 2005). It has several advantages over other techniques since it has low energy input, short welding time, low distortion and relatively low welding temperatures, which are essential criteria for LWMA dissimilar TWB manufacturing. Additionally, this process reduces manufacturing costs incurred due to the elimination of common defects, such as porosity, filler materials, shielding gases and expensive weld preparation. Hence, taking advantage of these positive factors in FSW, this process has already been applied to the construction of Al structures, e.g. the external fuel tank of rockets, high speed boats, etc. However, the FSW produces a wider weld line (e.g. 10 mm) than laser welding. Resistance mash seam welding Resistance mash seam welded joints are produced at lower temperatures compared to laser joints. Hence, they have better formability due to less martensite formation. However, the HAZ is approximately twice the width of the weld and some thickness increase at the joint is inevitable. Hot planishing of the joint reduces this thickness increase and results in enhanced formability than cold planishing (Saunders and Wagoner, 1996). Other processes, such as high frequency induction welding developed jointly by Elva Induction and Volvo, electron beam (non-vacuum) welding introduced by IKE Research Center at the University of Stuttgart, are also being developed exclusively for TWB manufacturing (Auto/Steel, 1995). The main advantages of electron beam welding over laser welding are higher welding speed, lower heat inputs and greater depth-to-width aspect ratios. Dissimilar sheet materials can also be welded using laser roll welding (Rathod and Kutsuna 2002; Marya et al., 2007). Using bimetallic transition inserts and/or coating the dissimilar materials prior to welding can help to alleviate the problem of intermetallic material formation. Thus, bimetallic inserts, with Al–steel and other combinations,


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

106 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


produced by non-fusion techniques such as rolling, flash welding or hot pressure welding, are used to produce dissimilar joints. Coating sheets with appropriate metal deposition and welding can also produce dissimilar joints. However, this method cannot be used to produce a strong bond and hence can be applied for aesthetic or sealing purposes only.

5.2.4 Defects Intermetallics Fusion welding of dissimilar materials can produce high concentrations of intermetallic compounds that are detrimental to the joint. For example, joining aluminum with steel can produce Al-rich and Fe-rich intermetallic compounds (FeAl2, Fe2Al3, FeAl3, FeAl, FE3Al, etc.) (Yasuyama et al., 1996, Ozaki et al., 2010). These brittle intermetallic layers lead to a fast rupture of the joint under stress. Generally, the thickness of the intermetallic layer determines the ductility and tensile strength of the joint. Intermetallic layers below 10 µm in thickness have no significant influence on the strength of the joint and are acceptable for engineering applications (Bach et al., 2005). Post welding thermal history is also a dominant factor in the growth of intermetallic compounds. Therefore, the growth of intermetallic compounds can be limited through the use of exact energy and heat supply, surface activating flux, etc. Internal and external defects TWBs typically exhibit decreased formability relative to their parent materials, arising from the combined influence of the weld surface roughness, the weldmetal hardening behavior and the internal porosity. The rough weld metal surface and the internal voids pose severe restriction to the weld metal ductility, thus limiting the formability of TWBs. Inadequate surface preparation in aluminum alloys and improper welding process parameters selection will result in internal porosity. The internal porosity that results from a laser welded joint has been classified into three categories (Katayama et al., 1998, Zhao et al., 1999), namely:

• • •

keyhole collapse macro-porosity due to the entrapment of gas bubbles micro-shrinkage porosity.

Keyholes are present during the welding process of aluminum TWB and keyhole stability plays an important role in defect formation. The weld-metal zone is characterized by a rapidly cooled microstructure with columnar grains growing inward from the HAZ and equi-axed grains at the center of the weld. Throughout the weld-metal zone, a very fine uniform population of micro-shrinkage porosity and larger macro-porosity can be observed. The micro-porosity is brought about



107

by solidification shrinkage of the as-cast dendritic structure of the welds, while macro-porosity arises from differences in hydrogen solubility between the liquid and solid states. Internal porosity produces a greater reduction in TWB failure strain when compared with a similar sized external surface defect (Bayley and Pilkey, 2005). As a result, when the major-straining direction is along the weldline, as shown in Fig. 5.4, the sheet thickness ratio is of far less importance. Magnesium alloys tend to produce more stable keyholes compared with aluminum due to a much higher equilibrium vapor pressure, lower boiling temperature and lower surface tension. In FSW, there has been growing interest in faint line patterns such as the ‘kissing bond’, as shown in Fig. 5.5, which are sometimes present in the stir zone of the friction stir weld due to the inclusion of oxide particles (Sato et al., 2005).

5.4 Straining directions in a TWB.

5.5 Kissing bond in FSW aluminum (Cao and Jahazi, 2008).


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

108 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


5.2.5 Failure prevention As with other TWB combinations, controlling material flow in LWMA and dissimilar material TWBs, with differential strength component blanks, is a major challenge to the manufacturers. It is highly desirable to minimize the weld line movement to reduce failure and improve the formability of TWBs. Various techniques have been developed to minimize the weld line movement utilizing drawbeads (Heo et al., 2001), controlling blank holding force (BHF) and/or constraining the weld line directly by using clamping pins (Kinsey et al., 1999; Kinsey et al., 2000; He et al., 2003; Tang et al., 2007). The drawbead (Fig. 5.6), positioned preferentially, adds restraining forces to the blank in addition to the blank holding force, so that the weld line movement is reduced. The drawbead geometry, such as die radius (Rd), bead radius (Rb), bead depth (Db), bead position (Bp), channel radius (Rc), channel width (Wc), etc, can be adjusted to obtain adequate material flow and to restrict weld line movement. In the case of differential strength TWBs, a low BHF is imposed on the high strength blank, while the low strength blank is imposed with a higher BHF. Thus, uniform material flow is achieved in the formation of the TWB by reducing the extent of stretching in the thinner/weaker blank. The weld line clamping method (Kinsey et al., 1999; Kinsey et al., 2000), as shown in Fig. 5.7, can also be used to restrict weld line movement. A rubber bushed piston end, attached to a hydraulic cylinder, is used to apply clamping forces to specific locations along the weld line of the TWB during the forming process. In this method, the determination of the clamping location is straightforward and non-straight weld lines can also be accommodated.

5.6 Drawbead design (Numisheet Benchmark, 1995).



109

5.7 Weld clamping method (Kinsey et al., 2000).

5.3

LWMA TWB formability

The formability of the metallic materials strongly depends on their deformation mode, loading history and on the plastic behavior and anisotropy of the rolled sheets. Thus, its prediction becomes more complex in TWBs with LWMAs and dissimilar materials due to the presence of differential strength and weld metals. The formability of TWBs is highly affected by the weld characteristics, the strength and thickness mismatch between the TWB components and the orientation of the blank sheets rolling direction relative to the weld line. Thus, additional care in process design for TWB forming is required. Generally, the formability limits are established by determining the limiting strains at the onset of localized necking, from plastic instability analysis and plotting forming limit curves (FLCs). Although the forming limit curve (FLC) of the TWB is closer to weaker material FLC, the FLC of the parent metal cannot be used to assess the onset of necking in TWBs due to the geometric discontinuities of the weld line, the presence of weld defects and the decreased work hardening rate of the weld metal. When straining occurs along the weld line, failure happens at the weld metal due to its low ductility, while straining perpendicular to the weld line (see Fig. 5.4) causes a weak base material to fail. In mash seam and laser welding, weld metal characteristics are insignificant (Saunders and Wagoner, 1995). But in FSW welded TWBs with wider weld lines, the formability is significantly influenced by the weld properties. The formability decreases further as the weld line location moves towards critical forming areas, such as cup corners. This effect is stronger as the strength ratio between the base materials increases. Minimization of weld displacement can be achieved by placing the weld in a region with low strains perpendicular to the weld line or by choosing appropriate base materials with equal load bearing capacities (Bhagwan and Kridli, 2004). Thus, the strength mismatch and the weld direction with respect to the major straining direction are the two major concerns in LWMA and dissimilar material TWB forming. Fusion welding processes do not yield a sound joint in some cases of dissimilar metals joining, such as in the case of the aluminum–steel TWB. The welding © Woodhead Publishing Limited, 2011

1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

110 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


produces a high concentration of intermetallic compounds that are detrimental to the joint, as well as to the formability limits. As previously mentioned, intermetallic layers below 10 µm in thickness have no significant influence on the strength of the joint and are acceptable for engineering applications. FSW makes it possible to weld aluminum and steel and produce a solid-state bond between the base metals through the frictional heat generated between the tool and the base metals. Thus, the formation of intermetallic compounds can be limited to produce a sound weld between soft aluminum and strong steel. Low energy input, short welding time, low distortion and relatively low welding temperature are essential criteria for dissimilar material welding, such as Al–steel welding, and these are characteristics of FSW. The use of Al–steel TWBs can be exploited in producing components such as a car door inner panel, as shown in Fig. 5.8, where rigidity at the hinge portion and lightweight at other portions is a necessity. Forming two different materials simultaneously, as in the case of Al–steel TWBs, requires careful determination of the process parameters. The Young’s modulus of aluminum is almost one third of that of steels; hence the elastic recovery of aluminum is comparatively larger than that of steels. The difference in

5.8 (a) Dissimilar materials TWB; (b) formed car door panel.



111

material anisotropy has a direct effect on the thickness distribution among aluminum and steel for the same forming conditions. The friction coefficient of steel is marginally higher than that of aluminum. The tendency of the lower strength material to split parallel to the weld line is higher (Auto/Steel, 1995). Figure 5.9a shows the ideal square cup geometry of an Al–steel friction stir welded TWB, with a typical wide weld line. In the actual forming process, the differential material strengths and hence the material flows, cause the flange shape to vary between different combinations of aluminum with different grades of steel blanks, as shown in Fig. 5.9b. A weak blank tends to flow more compared to a strong blank, leading to failure in the former. The failure in a weak blank can be reduced by allowing increased material flow in the strong blank, through the use of reduced blank holder force (BHF) at the strong blank section. Thus, the application of different blank holder forces on aluminum and steel blank segments can help to control the flange draw-in. Yet, a difference in the draw-in, between the different combinations of Al–steel TWBs, can be observed in Fig. 5.9b. The steel blank segment is subjected to more draw-in compared to the aluminum blank segment, except for the Al–mild steel (Al–DC) combination. In the case of

5.9 (a) TWB square cup with weld line; (b) flange deformation for different dissimilar material TWBs. DP, dual-phase.


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

112 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


Al–DC TWB, the flow characteristics of both materials are almost the same and hence very little difference in the drawn-in can be observed. On the other hand, a large difference can be observed in the case of Al–dual-phase steel (Al–DP) TWB where there is an appreciable draw-in at the steel side. An overall shift in the blank periphery is observed towards the aluminum side at the flange, as shown in Fig. 5.9b. Excess deformation is observed at the flange midsections very close to the weld line as a consequence of the materials strength mismatch and differential material flow characteristics. This can be restricted by optimizing the TWB shape prior to deep drawing operation and/or by using draw beads. Figure 5.10a shows the weld line movement across the cup. During the deep drawing of dissimilar material TWBs, the weld line shifts across the geometry,

5.10 (a) Weld line movement; (b) thickness variation.



113

depending upon the strength of the base materials. Generally, the weld line moves towards the stronger material side due to increased material flow in the weaker side. As shown in the figure, a larger shift in the weld line can be observed in Al–DP TWB than Al–Mild steel TWB. Figure 5.10b shows the thickness variation along a mid-cut section in the cup. A double neck near the punch radius, a typical phenomenon that occurs in deep drawing, can be observed on the steel side cup corner. The application of different BHFs reduces, but does not completely eliminate, the tendency of lower strength material to split parallel to the weld line. Al–DP TWBs show thinning tendency in aluminum parallel to the weld line. The aluminum sheet segment is subjected to large strains and, consequently, thinning along the weld line, indicating a critical zone as shown in the figure. Similar TWB welds of Al 5182 alloy have increased mechanical properties and good ductility relative to the Al 5182 base material. On the other hand, similar welds of Al 6016 alloy blanks display reduced mechanical properties relative to the Al 6016 base material and also have lower ductility. In an FSW Al 5182–6016 alloys TWB, the weld has increased yield stress and equivalent tensile strength to the weakest base material, i.e., the Al 6016 alloy. The ductility of the weld metal is lower than that of both base materials. Though the weld line is wider in an FSW TWB, the microstructure and mechanical properties of the HAZ are similar to the base materials, both in the advancing and the retreating sides of the similar and dissimilar welds. Similar TWBs of both grades of aluminum produce little weld line movement, while the dissimilar welds of Al 5182–6016 alloys TWBs produce irregular material flow at the cup edge close to the weld line as shown in Fig. 5.11 (Rodrigues et al., 2010).

5.11 Similar and dissimilar aluminum TWB cups: S55 (similar: AA5182–AA5182), S66 (similar: AA6016–AA6016) and D56 (dissimilar: AA5182–AA6016) (Rodrigues et al., 2010).


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

114


5.4

LWMA TWB benefits/recycling

The need to reduce weight is increasingly shifting the vehicle composition from steel to light metals and plastics. A typical car weighing 1500 kg currently contains about 1000 kg of steel, 125 kg of Al, 115 kg of polymeric materials and 5 to 6 kg of magnesium. It is estimated that a mass reduction of 22.5 kg would improve fuel efficiency by around 1%; thus, automotive manufacturers worldwide have goals to increase the Mg content of automobiles to between 45 and 160 kg (Pollock, 2010). While aluminum engine blocks have been slowly replacing cast iron blocks and saving around 66% in weight, magnesium equivalents take this reduction further to around 75% (Tharumarajah and Koltun, 2007). A V6 3.0 liter engine made of magnesium weighs 30 kg, compared to 39 kg and 84.6 kg, respectively, for an aluminum and cast iron block. Thus, aluminum has successfully penetrated the automotive market in the form of castings. However, an aluminum sheet of the proper alloy is still too expensive to penetrate significantly except in the production of components where lower weight has extra value (e.g., large hoods or deck lids). Further research is needed to either lower the cost of the LWMA currently used for body sheet or to develop methods to use less expensive alloys. In fact, casting typically uses lower cost aluminum alloys than sheet stamping. Also, joining technologies need to be improved to lower their cost while improving their quality, particularly for thin sheets (Hadley et al., 2000). On the other hand, the shift from steel to lightweight materials poses challenges in terms of its disposal after service life. For example, in a 2000 automobile, non-ferrous metals comprise almost 10% of the vehicle weight, but account for more than half of the scrap material value. The nonmetallic components still have a small negative value, reflecting the cost of disposal charged by the landfills for use of the residue as the daily landfill cover. In the recycling of both aluminum and magnesium, recycling processing costs are a small fraction of the prime reduction costs. The energy consumption per ton of a recycled aluminum ingot is approximately 2 kWh/kg, about 5% of the original production cost. The cost of re-melting scrap aluminum or magnesium is lower than that of the electric arc furnace production of steel from scrap. This is mainly due to the lower melting temperatures of the light metals compared to steel. This comparison is especially attractive for light metals on a per-volume basis. When considering weight reductions, material substitution is never done on a kg/kg basis; it is more often closer to a cc/cc basis. In this way, recycling significantly favors the substitution of light metals for steel in lightweight applications (Gesing, 2004). The environmental benefit of reducing weight with LWMAs to save on weight and consequent tailpipe emissions needs to be appraised. The production of LWMAs can be more energy intensive than steel/cast iron and, consequently, its environmental impact can be higher. However, this can be compensated for by the use of less energy due to lower weight and the cumulative saving in the spent energy.



5.5

115

Sources of further information and advice

ASM Handbook (1990), Properties and Selection: Nonferrous Alloys and Special-Purpose Materials, 10th Edition, ASM International. Dieter G (1967), Mechanical Metallurgy, New York, McGraw-Hill. Kutz M (2002), Handbook of Materials Selection, New York, John Wiley & Sons, Inc.

5.6

References

Abu-Farha FK and Khraisheh MK (2008), ‘An integrated approach to the superplastic forming of lightweight alloys: towards sustainable manufacturing’, International Journal of Sustainable Manufacturing, 1(1–2), 18–40. Al-Samman T (2009), ‘Comparative study of the deformation behavior of hexagonal magnesium–lithium alloys and a conventional magnesium AZ31 alloy’, Acta Materialia, 57, 2229–2242. AluMatter (2010), Wrought aluminium alloys, Available from: http://aluminium. matter. org.uk [Accessed 19th August 2010]. Auto/Steel Partnership (1995), Tailor welded blank design and manufacturing manual. Available from http://www.a-sp.org/publications.htm. Bach FW, Beniyash A, Lau K, Versemann R (2005), ‘Joining of steel–aluminium hybrid structures with electron beam on atmosphere’, Advanced Materials Research, 6–8, 143–50. Bayley CJ, Pilkey AK (2005) ‘Influence of welding defects on the localization behaviour of an aluminum alloy tailor-welded blank’, Materials Science and Engineering A, 403, 1–10. Bhagwan AV, Kridli GT (2004), ‘Formability improvement in aluminum tailor-welded blanks via material combinations’, Journal of Manufacturing Process, 6(2), 134–140. Boronski D (2006), ‘Cyclic material properties distribution in laser-welded joints’, International Journal of Fatigue, 28, 346–354. Braun R (2006), ‘Nd:YAG laser butt welding of AA6013 using silicon and magnesium containing filler powders’, Materials Science and Engineering A, 426, 250–262. Cao X, Jahazi M (2008), ‘Effect of welding speed on lap joint quality of friction stir welded AZ31 magnesium alloy’, Trends in Welding Research, Proceedings of the 8th International Conference, ASM International, 72–80. Chen CM, Kovacevic R (2004), ‘Joining of Al 6061 alloy to AISI 1018 steel by combined effects of fusion and solid state welding’, Int J Mach Tool Manu, 44, 1205–1214. Chen FK and Huang TB (2003), ‘Formability of stamping magnesium-alloy AZ31 sheets’, Journal of Materials Processing Technology, 142, 643–647. Cheng CH, Chan C, Chow CL (2007), ‘Weldment properties evaluation and formability study of tailor-welded blanks of different thickness combinations and welding orientations’, Journal of Materials Science, 42, 5982–5990. Elrefaey A, Takahashi M, Ikeuchi K (2005), ‘Friction-stir-welded lap joint of aluminum to zinc-coated steel’, Quarterly Journal of Japanese Welding Society, 23(2), 186–193. Gesing A (2004), ‘Assuring the continued recycling of light metals in end-of-life vehicles: A global perspective’, Journal of the Minerals, Metals and Materials Society, 56(8), 18–27. Hadley SW, Das S, Miller JW (2000), ‘Aluminum R&D for automotive uses and the Department of Energy’s role’, ORNL/TM–1999/157, Oak Ridge National Laboratory, USA.


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

116 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


Hantzsche K, Bohlen J, Wendt J, Kainer KU, Yia SB et al. D (2010), ‘Effect of rare earth additions on microstructure and texture development of magnesium alloy sheets’, Scripta Materialia, 63, 725–730. He S, Wu X, Hu SJ (2003), ‘Formability enhancement for tailor welded blanks using blank holding force control’, Journal of Manufacturing Science and Engineering, Trans. ASME, 125(3), 461–467. Heo YM, Wang SH, Kim HY, Seo DG (2001), ‘The effect of the drawbead dimensions on the weld-line movements in the deep drawing of tailor welded blanks’, Journal of Materials Processing Technology, 113(1–3), 686–691. Katayama S, Matsunawa A, Kojima K (1998), ‘CO, laser weldability of aluminium alloys (2nd Report): Defect formation conditions and causes’, Welding International, 12(10), 774–789. Kinsey B, Liu Z, Cao J (2000), ‘A novel forming technology for tailor welded blanks’, Journal of Materials Processing Technology, 99, 145–153. Kinsey B, Song N, Cao J (1999), ‘Analysis of clamping mechanism for tailor welded blank forming’, 1999 SAE International Body Engineering Conference, SAE Paper No. 1999–01–3192, Detroit, MI, USA, Sept. 28–30. Manesh HD, Taheri AK (2003), ‘Bond strength and formability of an aluminum-clad steel sheet’, Journal of Alloys and Compounds, 361(1–2), 138–143. Marya M, Rathod MJ, Marya S, Kutsuna M, Priem D (2007), ‘Steel-to-Aluminum joining by control of interface microstructures: Laser-roll bonding & magnetic pulse welding’, Materials Science Forum, 539–543, 4013–4018. Mori K, Tsuji H (2007), ‘Cold deep drawing of commercial magnesium alloy sheets’, Annals of the CIRP, 56(1), 285–288. Ozaki H, Kutsuna M, Nakagawa S, Miyamoto K (2010), ‘Laser roll welding of dissimilar metal joint of zinc coated steel and aluminum alloy’, Journal of Laser Applications, 22(1), 1–6. Panda SK, Kumar DR (2008), ‘Improvement in formability of tailor welded blanks by application of counter pressure in biaxial stretch forming’, Journal of Materials Processing Technology, 204, 70–79. Panda SK, Kumar DR (2009), ‘Study of formability of tailor-welded blanks in plane-strain stretch forming’, International Journal of Advanced Manufacturing Technology, 44, 675–685. Panda SK, Kumar DR (2010), ‘Experimental and numerical studies on the forming behavior of tailor welded steel sheets in biaxial stretch forming’, Materials and Design, 31, 1365–1383. Pastor M, Zhao H, DebRoy T (2000), ‘Pore formation during C.W. NdrYAG laser welding of aluminum alloys for automotive applications’, Revista de Metalurgia, 36(2), 108–117. Pastor M, Sagaro R, Cabrera R (2001), ‘Modelling of the welded profile and the keyhole of a laser welded Al 5182 alloy’, Revista de Metalurgia, 37(6), 643–652. Pollock TM (2010), ‘Weight loss with magnesium alloys’, Science, 328(5981), 986–987. Rathod J, Kutsuna M (2004), ‘Joining of aluminum alloy 5052 and low-carbon steel by laser roll welding’, Welding Journal, 83(1), 16–26. Ribic B, Palmer TA, DebRoy T (2009), ‘Problems and issues in laser-arc hybrid welding’, International Materials Reviews, 54(4), 223–244. Rodrigues DM, Leitão C, Menezes LF (2010), ‘A multi-step analysis for determining admissible blank-holder forces in deep-drawing operations’, Materials and Design, 31(3), 1475–1481.



117

Sato YS, Takauchi H, Park SHC, Kokawa H (2005), ‘Characteristics of the kissing-bond in friction stir welded Al alloy 1050’, Materials Science and Engineering A, 405, 333–338. Saunders FI, Wagoner RH (1995), ‘The use of tailor-welded blanks in automotive applications’, in Shen SF and Dawson PR (eds.) NUMIFORM 95, Simulation of Materials Processing: Theory, Methods And Applications, 5th International Conference on Numerical Methods in Industrial Forming Processes, 157–164. Saunders FI, Wagoner RH (1996), ‘Forming of tailor-welded blanks’, Metallurgical and Materials Transactions A, 27(9) 2605–2616. Schubert E, Klassen M, Zerner I, Walz C, Sepold G (2001), ‘Lightweight structures produced by laser beam joining for future applications in automobile and aerospace industry’, Journal of Materials Processing Technology, 115, 2–8. Stasik MC, Wagoner RH (1996), ‘Forming of tailor-welded aluminum blanks. In: Aluminum of magnesium for automotive applications’. A Publication of TMS, Warrendale, 69–83. Tang BT, Zhao Z, Wang Y (2007), ‘One-step FEM-based evaluation of weld line movement and development of blank in sheet metal stamping with tailor-welded blanks’, The International Journal of Advanced Manufacturing Technology, 35(3–4), 268–279. Tharumarajah A, Koltun P (2007), ‘Is there an environmental advantage of using magnesium components for light-weighting cars’, Journal of Cleaner Production, 15, 1007–1013. Thomas MW, Nicholas ED, Needham JC, Murch MG, Templesmith P et al. GB Patent Applications No. 9125978.8, December 1991; US Patent No. 5460317, October 1995. Uzun H, Donne CD, Argagnotto A, Ghidini T, Gambaro C (2005), ‘Friction stir welding of dissimilar Al 6013-T4 to X5CrNi18–10 stainless steel’, Materials and Design, 26, 41–46. Wang J, Takenaka Y, Hongu T, Fujii K, Katayama T (2007), ‘Laser-MIG arc hybrid welding of aluminium alloy – Comparison of melting characteristics between YAG laser and diode laser ’, Welding International, 21(1), 32–38. Watanabe T, Takayama H, Kimapong K, Hotta N (2003), ‘Joining of steel to aluminium alloy by interface activated adhesion welding’, Materials Science Forum, 426–432, 4129–34. Yasuyama M, Ogawa K, Taka T (1996), ‘Spot welding of aluminum and steel sheet with insert of aluminum clad steel sheet – Part 1’, Quarterly Journal of Japan Welding Society, 14(2), 314–320 (in Japanese) Yoshihara S, Nishimura H, Yamamoto H, Manabe K (2003a), ‘Formability enhancement in magnesium alloy stamping using a local heating and cooling technique: circular cup deep drawing process’, Journal of Materials Processing Technology, 142, 609–613. Yoshihara S, Yamamoto H, Manabe K, Nishimura H, (2003b), ‘Formability enhancement in magnesium alloy deep drawing by local heating and cooling technique’, Journal of Materials Processing Technology, 143–144, 612–615. Zhang KF, Yin DL, Wu DZ (2006), ‘Formability of AZ31 magnesium alloy sheets at warm working conditions’, International Journal of Machine Tools & Manufacture, 46, 1276–1280. Zhao H, White DR, DebRoy T (1999), ‘Current issues and problems in laser welding of automotive aluminium alloys’, International Materials Reviews, 44(6), 238–266.


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

6 Advanced high-strength steel tailor welded blanks (AHSS-TWBs) X. WU, Wayne State University, USA

Abstract: The automotive industries are continually searching for ways to reduce manufacturing costs while preserving (and improving) vehicle safety and fuel efficiency. Tailor welded blanks (TWBs), sheet metals fabricated from combinations of lightweight materials such as advanced high-strength steels (AHSS), offer an attractive solution and are the subject of intensive research and development. Types, characteristics and fabrication methods of AHSS-TWBs are discussed in this chapter. Formability of TWBs depends upon thermal behaviors during welding and the microstructural properties of component materials. Common test methods to assess TWBs are microhardness measurements across the weld seam, tensile tests, limit dome height tests and microstructural analyses. Key words: tailor welded blank, advanced high-strength steels, thermomechanical processing, formability, deformation, weld seam, fusion zone, heat-affected zone, ferritic, martensite, bainite, austenite, microstructure, microhardness.

6.1

Introduction to advanced high-strength steel (AHSS)

Advanced high-strength steels for tailor welded blanks (AHSS-TWBs) refer to the stamping blanks or hydroforming tubular blanks that are made from parent blanks of dissimilar materials containing at least one AHSS, or from the same AHSS but with different thicknesses. AHSS-TWBs have received great interest from manufacturing automotive bodies and other structural components where light weight is of primary consideration. The motivation to tailor a blank with dissimilar materials or thicknesses is to better distribute mass/weight and strength in a component, so that required structural functionalities (e.g. strength and safety, crashworthiness) can be better achieved at a reduced cost (by reducing the number of stamping parts and saving trimmed materials). For automotive structure components it is more reasonable and effective to use a wide range of AHSS combinations to further improve TWBs’ functionality for energy saving. For this reason, in recent years many research activities have been undertaken, almost in parallel with AHSS material development itself. Some examples of applied researches on AHSS-TWBs can be found in Malakondaiah et al.,41 Ribolla et al.,42 Múnera et al.,35 Jeswiet et al.43 and Sharma and Molian.44 It is expected that the applications of AHSS-TWBs for automotive body structures will become a common practice and in many cases the first choice. 118 © Woodhead Publishing Limited, 2011

Advanced high-strength steel tailor welded blanks

119

There are several ways to categorize AHSS to reflect their characteristics in different aspects that are also important for TWBs. According to WorldAutoSteel – the automotive group of the World Steel Association1 – AHSSs may be categorized in three different ways:

•

• •

Based on their metallurgical features associated with the thermo-mechanical processing and resulting microstructure. For example, dual phase (DP) steel, steel with phase-transformation induced plasticity (TRIP steel), etc. This feature is important for fabricating AHSS-TWBs since the welded region, after experiencing a different thermo-mechanical process from welding, needs to be related and compared with the base metals for understanding the forming behavior of TWBs. Based on the strength of AHSS, which is important for designing the configurations of AHSS-TWBs in autobody components. Based on plastic behaviors of AHSS such as total elongation, strength hardening exponent (n-value), hole expansion ratio, etc., which are also important for forming process design.

More details on AHSS and their applications can be found in the AHSS Guidelines.1 All of these characteristics need to be addressed in AHSSTWBs. For better understanding AHSS-TWBs, the relationship between process, microstructure and properties of current interest is briefly discussed here. The welding process in making TWBs has a great impact on the formability of produced blanks, and it is more so for AHSSs. This is because this family of materials often contains super-fine and hard second-phase particles (often martensite phase particles) within a ductile ferrite matrix, produced by a specific thermomechanical process for each type of AHSS. Melting and solidification in a fusion welding process will, in most cases, cause remarkable changes not only in the grain size distribution (as will occur in conventional TWBs) but in the phase configuration as well. In this chapter the research on welding methods, the process conditions and their effects on the properties of welded AHSS will be described and discussed. Formability is one of the major concerns even for monolithic AHSS and for regular TWBs, and it is more so for AHSS-TWBs. For various grades of steels the strength and ductility often show an inverse relationship – the formability decreases with increasing strength, though AHSSs often show better formability than regular steels of the same strength grade. A TWB consists of three constituents – two parent blanks (of dissimilar materials or thicknesses) and one weld seam. There are more potential formability problems for AHSS-TWBs than for TWBs of regular materials (non-AHSS) and for monolithic AHSS blanks. Based on the forming mechanics involved, failure tends to occur at the weak region of the deformation zone, which can be within any one of the parent blanks or within their welding seam. Depending on the strength competition among the constituents the following cases exist: 1

Failure occurs in the thinner/weaker portion of one base metal blank: Due to non-homogeneous deformation, localized necking and fracture may occur


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

120 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


within the thin/weaker part of TWB. This is a general problem for TWBs, and it has been covered by other chapters of this book. This chapter will not discuss this macroscopic forming mechanics issue, but rather will only address some intrinsic formability issues of AHSS. 2 Failure occurs within the weld seam and its neighboring region: Due to the change of thermal history, new material constituents will form within the weld seam and its vicinity, the heat-affected zone (HAZ) that did not melt in welding but changed phase configurations and grain structures that are very different from the parent materials. This is a special issue related to AHSS-TWBs that has not been addressed by other chapters and thus needs to be further studied in this chapter. We will discuss this issue from two aspects: (a) The TWB fabrication processes, cutting and welding, and their effects on the microstructure and initial properties of welded seams; (b) The effect of forming condition on the formability of AHSS-TWBs. For the failure to occur in the welded area, the formability problem is related to both intrinsic properties of the welds and the principal loading directions relative to the weld line direction. There are two extreme cases in which the welded region becomes a limiting factor of the formability. (a) If the welded region is weaker than the two base metals, due to the annealing effect during a slow cooling rate (that often happens in HAZ), and if the major principal stress/strain direction is perpendicular to the weld line, the weld area will fracture first. (b) If the weld line is more brittle than the base metals and the major principal strain is parallel to the weld line, fracture may initiate within the weld seam and further propagate into the base AHSS material. In addition, the existence of dissimilar thicknesses and properties around the welded seam will generate stress concentration and strain inhomogeneity, which can cause a premature failure during forming. In production, more complex combinations of material properties and an arbitrary weld line direction relative to an applied stress field exist, and a numerical analysis is needed to determine the limiting factor and forming limit of AHSS-TWBs based on a reliable failure criterion, which is a challenging task. In this chapter, the types of AHSSs and the blank welding/tailoring process for fabricating AHSS-TWBs are to be briefly described first, and then the previous studies on characterization of AHSS-TWB properties and formability are reviewed with the focus on understanding AHSS-TWB materials, processes and failure mechanisms. The current status of AHSS-TWBs related to both process and product/component designs are also provided.

6.2

Types of advanced high-strength steels and their characteristics

Most AHSSs contain dual or multiple phases with one or more hard phases (e.g. martensite, bainite or retained austenite) in the form of super-fine phase



121

particles that are surrounded by the ductile ferritic phase. These kinds of finecomposite-like microstructures are produced in situ through special thermomechanical schedules during rolling and post-rolling heat treatment, and are also related to their special chemical and metallurgical designs. The hard phase particles play the role of composite reinforcements that significantly enhance the strength of materials, while the ductile ferritic phase as matrix provides required ductility for plastic deformation in stamping, so that a combined strength and ductility/formability can be achieved. The fundamentals of AHSS development can be seen in the book by Bhadeshia and Honeycombe.2 These superior properties combined with strength and ductility make them especially useful for AHSS-TWBs and provide further opportunities for improved structural integrity and for overall weight reduction. On the other hand, the thermal process involved in the welding can significantly change the as-received microstructure and properties within the fusion zone and HAZ. To address AHSSTWBs some common AHSS materials are briefly introduced below, extracted from the descriptions given by WorldAutoSteel.1

6.2.1 Dual phase (DP) steels This type of steel consists of a ferritic matrix and a hard martensitic second phase in the form of islands. Increasing volume fraction of the hard second phases generally increases the strength of the DP steel. DP steels may be produced by hot-rolling followed by controlled cooling from the austenite phase (in hot-rolled products), or by heat-treating cold-rolled sheets that contains a step to rapidly cool the material from a two-phase regime (ferrite plus austenite phases), to convert the austenite to martensite. The process can be achieved by a continuous product line with continuously heating/annealing and hot-dip (for quenching and surface coating).

6.2.2 Transformation-induced plasticity (TRIP) steels The microstructure of a TRIP steel contains retained austenite particles that are embedded in a primary matrix of ferrite. In addition to a minimum of 5% by volume of retained austenite, hard phases such as martensite and bainite are often present in varying amounts. TRIP steels typically require the use of an isothermal hold at an intermediate temperature, which produces some bainite. The higher silicon and carbon contents of TRIP steels also result in significant volume fractions of retained austenite in the final microstructure. During the sheet forming process the applied stress or strain induces transformation from retained austenite to martensite phases, associated with a volume expansion, which contribute to a certain amount of additional plasticity in stamping when a tensile mean stress presents.


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

122 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


6.2.3 Complex phase (CP) steels CP steels have high ultimate tensile strengths, and their microstructure contains small amounts of martensite, retained austenite and pearlite within the ferrite/ bainite matrix. An extreme grain refinement is created by retarded recrystallization or precipitation of microalloying elements like Ti or Cb. In comparison with DP steels, CP steels show significantly higher yield strengths at equal tensile strengths of 800 MPa and greater. CP steels are characterized by high energy absorption and high residual deformation capacity.

6.2.4 Martensitic (MS) steels The austenite that exists during hot-rolling or annealing is transformed almost entirely to martensite during quenching. The MS steels are characterized by a martensitic matrix containing small amounts of ferrite and/or bainite. This structure can also be developed with post-forming heat treatment. MS steels provide the highest strengths (above 1.6 GPa) among the multiple phase steels. In order to improve ductility, MS steels are often subjected to post-quench tempering.

6.2.5 Ferritic–bainitic (FB) steels FB steels have improved edge stretch capability and thus sometimes called stretch flangeable (SF) or high hole expansion (HHE) steels. FB steels have a microstructure of fine ferrite and bainite. Strengthening is obtained by both grain refinement and the second phase hardening with bainite. FB steels are available as hot-rolled products. Compared to HSLA steels with the same level of strength, FB steels also have a higher strain hardening exponent (n) and increased total elongation. Because of their good weldability, FB steels are considered for tailored blank applications. These steels are characterized by both good safety performances and good fatigue properties.

6.2.6 Twinning-induced plasticity (TWIP) steels TWIP steels contain the entire austenitic phase at room temperatures, due to the high manganese content (17–24%) that postpones the phase transformation during cooling. During plastic deformation the principal deformation mechanism is twinning inside the grains. The twinning process refines the microstructure and results in a high value of the instantaneous hardening rate (n value), similar to the role of grain refinement. TWIP steels combine extremely high strength with extremely high formability. The n value increases to a value of 0.4 at an approximate engineering strain of 30% and then remains constant until a total elongation around 50%. The tensile strength is higher than 1000 MPa.



123

6.2.7 Hot-formed (HF) steels As one solution to improve formability, reduce springback and increase die life, HF steels are developed. The stamping is performed at temperatures above the austenitic region (900–950°C), at which good formability (elongation above 50%) can be achieved, and then quench hardened within a water-cooling die, and a yield strength above 600 MPa and an ultimate strength above 1300 MPa at room temperature can be obtained. Typical stamping cycle time is 20 to 30 seconds for each press cycle. However, several parts can be stamped at the same time so that two or more parts can be obtained per cycle. Hot-forming boron steels are most commonly used for safety and structural parts. A surface coating, such as aluminum–silicon, is often applied that forms a dense Fe-Al-Si surface layer at elevated temperatures and effectively reduce oxidation.

6.3

Fabrication of advanced high-strength steels for tailor welded blanks (AHSS-TWBs)

The majority of TWBs are produced by laser cutting and laser welding from dissimilar materials or different thicknesses for structural optimization. Therefore, laser processing will be emphasized here. Laser beam provides a high energy density that is concentrated within a small processing zone, and the solidification can be completed at a high cooling rate behind the laser welding zone at a high welding speed. Thus, for conventional steels the produced welding region (including both fusion zone and HAZ) usually has a fine-grained microstructure, and the strength is often higher than the base metal. Therefore, excellent welding quality and blank formability can be produced by laser welding. Whether this can be applied to AHSS as base material is one of the topics under investigation in AHSS-TWBs, and will be reviewed in this section. A large amount of research has been conducted on the effect of welding parameters on the AHSS microstructures, and the results can be applied to both fabrication of AHSS-TWBs and to auto-body assembly involving welding of AHSS components. In this section the cutting and welding methods are described/ summarized, and their results are given in section 6.4.

6.3.1 Laser cutting Most laser cutting systems use CO2 laser which offers a high power at relatively lower cost. One study on laser cutting of AHSS was studied by Lamikiz et al.3 Their results show that the blank thickness has an important effect on the cutting process and edge quality: the thicknesses above 1 mm and below 1 mm have very different cutting behaviors that require different settings of cutting parameters. This means that a thicker sheet requires a higher energy input, and that the heat dissipation rate is slower as compared with a thinner thickness sheet. Thus, there


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

124 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


exists a critical laser line-power and laser speed combination in order to obtain a high quality edge in terms of its geometry, making the thickness a sensitive parameter in laser cutting. The effect certainly also depends on the type of steels and lasers, as well as the welding parameters used. The surface coating on the sheet steels is also found to have an important influence on the cutting process, which can be explained due to its effect on the chemical composition and melting point. Interestingly, from this study the type of materials did not show large influence on the variation of the cutting parameters, probably because the differences in chemical composition and microstructure are less significant in terms of the melting process during laser cutting. To optimize the process, the best laser beam focus position was found to be underneath the sheet surface. Pulsed mode laser cutting of sheets for AHSS-TWBs was reported by Bagger and Olsen.4 The laser cutting process can be optimized to give a good edge quality required for subsequent laser welding TWBs. Systematic laboratory experiments were conducted, and the effects of the processing parameters were analyzed that included the cutting speed, assisted gas pressure, average laser power and pulse energy. The cutting quality was evaluated based on the edge squareness, roughness and dross attachment of laser cut blanks. High quality edges with a squareness offset as small as 0.015 mm were obtained, which is well suited for subsequent laser welding for TWBs.

6.3.2 Laser welding Laser welding has become a major technique used for fabricating TWBs, including AHSS-TWBs, due to its highly concentrated energy within a restricted processing zone and its high welding quality–productivity combination. For metal cutting and welding, high power and energy density are needed. Based on present technology, CO2 laser has the highest power, lowest capital cost and operation cost per power over other types of lasers, and is used most commonly. YAG laser is next, but it has a lower frequency that reduces energy loss from light reflection of workpiece, which is important for welding aluminum and many other lightweight metals. These two types of laser apply the focused light power that not only melts the metal, but generates evaporated metal gas at a focused point, or commonly called ‘keyhole’. Due to the high energy concentration at a high line velocity, less heat is released to the base metal, resulting in improved mechanical properties of the welds. Another type of laser, diode laser, has received rapid development in recent years due to the increased power through stacking semiconductor units, but due to less energy concentration the keyhole will not form and so the heat conduction is relatively high. It is also seen in metal welding applications. The main interests in welding AHSS-TWBs include the welding quality at a high welding speed, the mechanical properties of as-welded materials, and the welding process optimization for obtaining highquality welds.



125

CO2 laser Han et al.5 performed a study on butt-welding of 800 MPa class TRIP steel with the focus on the effect of welding speed on microstructure and porosity of welds. It was reported that the high welding speed did not cause a perceptible effect on microhardness values, indicating that the increased cooling rate from the high welding speed did not cause martensite transformation. In addition, a lower porosity weld was achieved at a high welding speed. Thus, a high welding speed is proffered for TRIP steels. Chatterjee et al.6 used a 3.5 kW slab CO2 DC power source for welding thin interstitial-free (IF) and thick DP steels with a thickness ratio of 2. The final beam diameter was 360 µm focused with a 300 mm mirror, with the laser beam position offset by 100 µm towards the thicker DP side to balance the energy and the laser beam focused at a point 25% of the thickness above the sheet bottom. Pure argon was used as a shielding gas at the flow rate of 30 l/min. YAG laser Yan and Gallagher7 used a Nd:YAG laser to weld several regular and advanced high-strength steels (including HSLA300, DP600, M900 and M1310 steels) at various speeds, and found that the tensile strength of the welds was dependent on steel grade, weld penetration and weld width. This is understandable, since DP steel is more sensitive to the cooling rate that varies the amount of martensite transformed as well as the grain size, while for regular high-strength steels the main effect of welding speed is to change grain size. The weld penetration and weld width are related to the line power or input energy per weld length, which has a similar effect on the produced microstructures and properties. In addition, based on their report, the fatigue strength of the welds was more closely related to the sheet thickness than to the steel grade or microstructure. In another study of laser welding of TRIP780 and DP980 by Sharma and Molian,8 a 2 kW Trumpf TRUDISK 6002® Yb:YAG laser beam was utilized to join 1 mm thick TRIP780 with 1.5 mm thick DP980 and with 1 mm thick mild steel, respectively. Results indicate that the laser welds exhibit excellent strength and hardness with minimal defects. Diode laser Diode laser has relatively lower power density as compared with CO2 and YAG lasers, but it has a compact size that offers a high flexibility in welding application. Farabi et al.9 used a diode laser with a rectangular shaped beam and a focal length of 90 mm for welding DP 600. Ultra-high purity argon was used as a shielding gas on both top and bottom surfaces of the blanks. The welding speed was at 1 m/min in full penetration bead on a plate mode.


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

126 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


Xia et al.10 used a Nuvonyx ISL4000L diode laser for welding high strength low alloy (HSLA) and DP980 (dual phase, 980 MPa) sheet steels, that was operated at 4 kW and the welding speed varying from 1.0 m/min to 1.6 m/min. With the relatively large rectangular focal spot of 0.5 × 12 mm, and at the maximum travel speed of 1.6 m/min, a fully penetrated weld could be produced. High purity argon at a flow rate of 30 l/min was applied for shielding the top surface only.

6.3.3 Other methods for welding AHSS Friction stir welding (FSW) This is a solid-state material joining process that uses a spining tool head to penetrate the workpiece mating pairs, during which frictional heat and deformationinduced heat are generated that soften and mix adjoining pairs. This technique was initially developed for welding aluminum alloys and can be considered for replacing riveting of aircraft component joining, but in recent years it has been extended to a wide range of metals and applications. Miles et al.11 reported a study on FSW of AHSS-TWBs. More details on FSW can be found in a review by Nandan et al.12 Gas metal arc welding (GMAW) This welding method can operate at a low cost, but the amount of heat dissipated to the base metal is high; not only is it not energy effective but the method compromises the material properties. Kapustka13 studied GMAW for joining several AHSSs, including coated DP780, DP980, TRIP600, and TRIP780 of the thicknesses ranging from 1.34 to 1.54 mm. Although the welding was for similar materials and thicknesses, the welding behaviors could be correlated to the TWB application. Welding aluminum–steel TWBs Joining aluminum and steel is of great interest for potential aerospace and automotive applications but is very challenging due to the following difficulties:14, 15

• • •

Al and Fe tend to form brittle intermetallic compounds. Fe4Al13 and Fe2Al5 are the predominant intermetallics compounds found in the Al–steel weld area, and have led to a fast rupture of the joint under stress. Poor wetting behavior of aluminum to steel, causing difficulties in mixing molten Al and Fe. Very different physical and chemical properties of the two base metals, for example the significantly different melting temperatures of aluminum and steel, which reduces the welding temperature window, if it does exist.



127

Therefore, conventional fusion welding processes do not produce a sound joint, and special welding techniques are needed. The successful joining of Al to steels has been reported using the following methods:

• • • • • • •

friction stir welding14, 16 electron beam welding17 and combined fusion and solid-state joining15, 18–20 impact joining21 interface-activated adhesion welding22 ultrasonic butt welding23 cladding (by rolling or explosive)24 electrical discharge welding,25 etc.

In these studies it was found that, in general, a thickness of the intermetallic layer below 10 µm thickness has no significant influence on the strength of the joint and is acceptable for engineering applications. The growth of intermetallic compounds can be limited by controlling the exact energy and heat supplication, and by using a surface activating flux.

6.4

Properties and formability of AHSS-TWBs

The welding qualities of AHSS-TWBs are usually tested by:

• • • •

Microhardness measurements across the welding lines so that the local properties of the welded zones (fusion zone, heat-affected zone and the base metals) can be obtained and compared. Tensile tests with a dog-bone specimen geometry and with the welding line either in the parallel or perpendicular direction to the tensile axis. Other mechanical tests, such as fatigue and impact tests, to characterize the performance of welded or further formed components. Various formability tests, including a cup dome test that uses a hemispherical punch head of 1″ in diameter on fully clamped TWBs.

Due to the dissimilar materials or thicknesses, the stress and strain condition may not be under a balanced biaxial tension. Conventional limit dome height (LDH) test with a 4″ diameter hemispherical head can also be used for TWBs, in which the initial weld line is placed at the symmetrical axis. By using fullyclamped blanks or partially clamped blanks of smaller width, and placing the weld line in two perpendicular orientations of the major principal direction, a forming limit diagram (FLD) for TWBs can be obtained. Unlike conventional steels and other alloys, various AHSS materials contain dual or multiple phases that experience complex changes during melting and solidification. Thus, the properties and formabilities of AHSS-TWBs are controlled by the specific type of AHSS, and by the response of the material to the applied welding process conditions. In general, laser welded steel blanks often show higher hardness in the fusion zone and lower hardness in two HAZs relative


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

128 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


to the base metals, being the results of the different cooling rates in the two zones. Depending on the thermo-mechanical history and initial/final microstructures and properties, the AHSS-TWBs need to be studied on a case-by-case basis. For this reason, this section needs to be discussed based on the individual material system and welding process used.

6.4.1 Laser-welded DP600 Microhardness distribution and microstructures across the laser welded line Farabi et al.9 studied the properties of laser-welded DP600 steel having a thickness of 1.2 mm with a galvannealed (GA) coating (46 g/m2 at the top and 47 g/m2 at the bottom). This study, although not directly focused on tailor welded blanks, characterized the mechanical properties of laser-welded DP600 that are applicable to AHSS-TWBs. The obtained microhardness distribution across the laser weld line is shown in Fig. 6.1. The fusion zone has significantly high hardness due to the high cooling rate from laser welding that promotes martensite transformation, and also due to a good quenching ability of DP steel to form the martensitic phase, as evident from the microstructure observation in Fig. 6.2. In contrast, in the HAZ with the lowest microhardness, bainites and tempered martensites were formed from initial martensite in base metal. This explains the reason for softening.

6.1 Microstructure and mechanical properties of laser-welded DP600 steel joints (courtesy of Farabi et al.9).



129

6.2 SEM micrographs showing the microstructural change of a laserwelded DP600 steel joint: (a) fusion zone, (b) outer HAZ (soft zone), and (c) base metal, where M, F, B, and TM stand for martensite, ferrite, bainite, and tempered martensite, respectively (courtesy of Farabi et al.9).

Tensile properties of welded coupons The stress–strain curves of DP600 base metal and welded joint coupons in the transverse direction to the weld seam are shown in Fig. 6.3, at various strain rates, along with a typical failed specimen. The curves of base metals were smooth and continuous at all the strain rates (Fig. 6.3a), while the welded DP600 joints showed yield point phenomena at all the strain rates as shown in Fig. 6.3b. All the welded samples failed at the outer portion of the HAZ, see Fig. 6.3c. It was verified that the onset of yielding occurred in the softened HAZ first, and then the majority of the plastic deformation was accumulated in that zone until final failure. The welded joints had higher yield strength than the base metal but the ultimate tensile strength of the welded specimens was almost the same or slightly lower than that of the base metal. The total elongation to failure of welded specimens was significantly lower than that of the base metals, due to non-homogeneous properties and stress concentration outside the hardened weld zone.


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

130 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


6.3 Engineering stress–strain curves of DP600 steel tested at different strain rates, for (a) base metal and (b) laser-welded joints, where the weld line was perpendicular to the tensile axis of the dog-bone specimen; (c) the typical failure location in the tested specimens (courtesy of Farabi et al.9).

Fatigue of AHSS-TWBs From a tension–tension fatigue experiment of dog-bone specimens with the weld line perpendicular to the tensile axis, the S–N curves of DP600 are shown in Fig. 6.4, together with photographs of two typical fractured specimens. The fracture occurred near the HAZ but within the base metal for the stress amplitudes higher than 250 MPa, due to stress concentration near the hardened weld line. For the stress amplitudes lower than 250 MPa, the fracture occurred in the base metal away from the weld line, and the stress concentration did not cause damage initiation and final failure. Fatigue crack initiation occurred from the specimen surface, formed striation in stable crack propagation and finally, in the fast propagation area, left dimples and deformation bands on the fractured surfaces.



131

6.4 (a) Fatigue S–N curves, and the failure locations for (b) failure at high amplitude above 250 MPa and (c) below 250 MPa (courtesy of Farabi et al.9).

6.4.2 Laser-welded IF–DP steels for TWBs One of the comprehensive studies on laser-welded TWBs made of interstitial-free and dual-phase steels, IF–DP, was provided by Chatterjee et al.6 and is reviewed here. Laboratory scale laser-TWBs were produced for welding IF, 300 MPa, and DP, 590 MPa, at the thickness ratio of 2, under the situation that the DP side’s material strength was almost two times stronger and its thickness was two times larger. Microhardness distribution The microhardness distribution across the weld line is shown in Fig. 6.5. Within the fusion zone a significant increase in hardness up to 350 HV was obtained from laser welding, and the hardness was rapidly decreasing away from the central peak location, and two significantly different hardness distribution profiles were shown for the two sides. A reduced hardness in the IF HAZ (about 175 HV) relative to the IF base metal, and an increased hardness in the DP HAZ relative to the DP base metal were observed. Microstructures The above hardness distribution characteristics are closely related to the microstructure evolution. The microstructures of the base metals, weld and HAZs are shown in Fig. 6.6. The increased peak hardness in the fusion zone can be related to the rapid cooling rate from laser welding, which might be faster than the


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

132 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


6.5 Microhardness distribution across the weld line (courtesy of Chatterjee et al.6).

6.6 Microstructures of (a) IF, (b) DP590, (c) welded zone (courtesy of Chatterjee et al.6).



133

cooling condition of base DP steel under its production condition, so a high martensite volume fraction was produced that brought about the hardening effect in the fusion zone. This phase transformation hardening did not apply to IF steel due to its low carbon content. The reduced hardness of the IF HAZ relative to its base metal can be due to grain size increase in the HAZ, a commonly seen situation. Tensile and formability testing Tensile tests along the directions parallel and perpendicular to the weld line direction were performed, see Fig. 6.7(a,b) for the testing configurations, and Fig. 6.8 for the results. As expected, the tensile strength of the two TWB specimens tested in two orientations fell between the two base metals, and the tension in the parallel direction of the welds was stronger than that in the transverse direction, which is inconsistent with the upper bond and lower bond analyses. The total elongation of the two TBW tests showed more complex features with a combined effect from the two parent base materials. In the case of tension in the transverse direction, the weld region was only a small portion of total gauge length, within which the majority of the weld region was stronger than the base metal of each side, so the total elongation from the three constituents (the weld and the two base metals) in series was somewhere between the two base metals, and in this case it was closer to that in DP steels. While in the case of tension in the longitudinal

6.7 Subsize tensile test specimens: (a) weld line transverse and (b) weld line longitudinal; (c) balanced biaxial tension test with 4” punch (also called limit dome height test) (courtesy of Chatterjee et al.6).


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

134 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


6.8 Engineering stress–strain plots of parent IF–DP steels and welded blanks (courtesy of Chatterjee et al.6).

welding direction, regardless of the tailoring configurations with different materials or gauge thicknesses, or both, all of the three constituents must be stretched to the same amount, and the welded region almost always has lower strain to failure, so the observed fracture started from the welded region at a lower elongation and propagated to its neighboring region, leading to failure of the entire specimen at an elongation that is lower than that of the base metal (DP in this case). In formability testing, LDH tests were performed with a different sheet width that is smaller than the punch diameter so that the deformation path was in a drawing mode (under tension–compression in the second quadrant of principal strain space), see Fig. 6.9 for two failed specimens in the two orientations, respectively. The load vs. displacement curves of welded strips of 50 mm width from LDH tests are shown in Fig. 6.10, and the two monolithic base metals also shown in the figure that provides a good comparison of the punch forces and the dome heights to fracture among two parent base metals and the two welded TWB strips in two orientations. The magnitudes of the punch forces follow the order from large to small as: DP base metal, DP–IF welded longitudinal, DP–IF welded transverse and IF base metal. The total dome height to failure follows the reversed order; again, the welded strips of two orientations were between the two parent base metals, and the transverse direction had a higher dome height to failure than that of the longitudinal direction, being consistent with the results from the above-mentioned tensile tests (but now with slightly wider sheet width under punch stretching).



135

6.9 Crack formation in (a) transversely welded blanks and (b) longitudinally welded blanks (courtesy of Chatterjee et al.6).

6.10 Load vs. dome height plot for different types of blanks of 50 mm width (courtesy of Chatterjee et al.6).

The FLDs for welded blanks and base metals are shown in Fig. 6.11, obtained from the digital image correlation (DIC) technique. The FLD (limit strain in the plane strain condition) for transversely welded blanks was at about 39% major strain, that is between that of the two base metals, whereas for a longitudinally


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

136 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


6.11 Forming limit diagrams of parent steel sheets for various welded blanks tested. A DIC automatic strain analysis system was used (GOM optical measuring technique coupled with Argus software version 6.1.04; GOM mbH, Braunschweig, Germany) (courtesy of Chatterjee et al.6).

welded blank it was at 36.5% major strain, very close to that of DP. With the use of fully circular blanks, the strain path was equal or close to balanced biaxial tension (not exact if the macroscopic anisotropy of tailored blanks is considered), and the two weld-line orientations became identical. In this case the limit strains of two tailored blanks were significantly lower than either one of the parent materials. With the use of narrow strips the strain paths were close to the uniaxial tension condition, and the results showed that the longitudinal weld was the worst case, the DP base metal and the transverse weld were similar and the IF base metal was the best. This sequence is consistent with the results in uniaxial tension (Fig. 6.9), but is inconsistent with the results of the limit dome height at fracture, for strip specimens (Fig. 6.11), where the two welded samples were worse than two base metals. This inconsistency may be explained as follows: the uniaxial tension test and punch stretching results describe the average properties and material responses within the gage length or the clamped portion of the strips, for both monolithic base metals and the welded samples that are the macroscopic mixture of three constituents; while the FLD results (Fig. 6.11) reflect the intrinsic material fracture



137

strains under different strain paths and stress/strain states. With the use of DIC, even more local strain information is provided than from the traditional circle– grid method, so the differences between average properties and local properties become more clearly seen. However, Fig. 6.11 did not provide detailed information on the locations where the fracture strains were measured, which could be interesting for TWBs. On the other hand, the difference between uniaxial tension and punch stretching results can be due to two different testing configurations, with the latter involving superimposed bending and punch friction effects. Fatigue and impact tests Fatigue tests with a dog-bone specimen were performed and the resulting curves of stress versus the number of cycles to failure (or S–N curves) are shown in Fig. 6.12. Only high cycle fatigue testing was carried out (at above 103 cycles) on the transverse weld specimens. The specimens had two pre-notches at the two ends of the weld segment in the transverse direction, and the notch root radius was 0.2 mm and the notch depth was 2 mm (see the inset in Fig. 6.12). The notches were used to restrict the fatigue failure in the weld and not in the thinner IF sheets. To avoid buckling only tension–tension mode was used at the stress ratio r = 0.05. The same geometry was also used for fatigue testing of parent DP and IF steels for comparison. The results show that the weld had a shorter fatigue life than the parent DP or IF steel, except for one data point of IF base steel at 225 MPa and at

6.12 S–N curves of IF and DP base metals and TWBs. The fatigue test specimen used in high-cycle fatigue (HCF) testing is also shown. The test was performed in a magnetic resonance testing machine (courtesy of Chatterjee et al.6).


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

138 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


about 104 cycles, which was shorter than for the welded material. The reduced fatigue life of the welded region can be related to the hardening effect from the laser welding and rapid cooling (Fig. 6.5 and 6.6), which produced a microstructure containing a higher volume fraction of martensite in the fusion zone, and consequently gave lower toughness and reduced fatigue life. Also seen in Fig. 6.12, the endurance limits (the stress at 106 cycles) of the welded region were lower than that of the parent DP and IF. It is notable that this set of fatigue tests used the pre-notched specimens to confine the damage zone within each constituent of very different strengths; the S–N curves in Fig. 6.12 do not reflect the applied stress relative to the individual strength of each constituent. If we normalize the applied stress with the constituent strength, the S–N curve of the welded seam would be much lower than the two base metals, i.e. have a much shorter fatigue life and a lower endurance limit. In addition, the fatigue endurance limit is defined by the stress at 106 cycles of the S–N curve, below which it is traditionally considered to have no further fatigue damage. This endurance limit is set at 106 cycles not only based on the observed material properties but also based on the testing techniques within the affordable time window. From Fig. 6.12 the trend of the obtained S–N curves does not seem to suggest that the damage process would be stopped at 106 cycles. This ‘no further damage’ limit is questionable, as it is seen for many other high-cycle fatigue materials tested in modern high frequency fatigue testing systems. Impact tests measure the energy absorption in the event of structure crashing (e.g. a car body collision). The above pre-notched specimen geometry was also used for the impact tests. At room temperature the impact energies for the notched samples of IF (0.8 mm), DP (1.6 mm) and a transverse welded TWB were at 7.8 J, 20 J and 5.4 J, respectively, indicating that the total energy needed for the crack to propagate within the weld and cut through the welded cross-sectional area was the lowest of that of the two base metals. The results were not sensitive to temperature within the range from 27°C to −25°C, indicating no brittle–ductile transition occurred in this temperature region. For a reference, a review of historical development of impact testing, the problem of traditional Sharpy test for thin sheets and a new industrial impact testing system and testing method for welded specimens that simulates the real crashing process can be seen in Bayraktar et al.26

6.4.3 YAG laser-welded TRIP780–mild steels and TRIP780–DP980 steels The two dissimilar AHSS-TWBs were studied by Sharma and Molian.8 A 2 kW Yb:YAG laser beam was used under the material/thickness combinations listed in Table 6.1. The weld profile and microstructures were characterized with optimal microscopy and the mechanical properties were evaluated by microhardness, tensile and fatigue tests.



139

Table 6.1 Materials and weld theme Material 1

Pre-strain

Thickness Material 2

Thickness Label

TRIP780 CR and GI

Unstretched

1 mm

Mild steel CR

1 mm

TRIP780-0/MS

TRIP780 CR and GI

10% stretched 1 mm

Mild steel CR

1 mm

TRIP780-10/MS

TRIP780 CR and GI

Unstretched

1 mm

DP980 CR

1.5 mm

TRIP780-0/DP980

TRIP780 CR and GI

10% stretched 1 mm

DP980 CR

1.5 mm

TRIP780-10/DP980

CR = cold rolled, GI = galvanized iron. Source: Sharma and Molian (2009).8

Microhardness distribution Microhardness distributions across the weld line for all of the above combinations of tailoring cases were tested, and one set of results is given here to compare the welding of TRIP780 to mild steel and to DP980, and TRIP780 to DP980 with and without pre-strain (TRIP-10 and TRIP-0). The results show that (see Fig. 6.13):

•

•

The fusion zone hardness: In all three cases of TRIP-0 or TRIP-10 with mild steel and with DP980, the fusion zones exhibited much increased hardness as compared with the base materials, and in the case (Fig. 6.13b) of TRIP-0 to DP980 seems to have the highest hardness but with large variation. It is not so clear why 10% pre-strain case (Fig 6.13c) showed a lower hardness than zero pre-strain (Fig 6.13b), possibly due to the release of the stored energy in the TRIP780 side that eased the melting process of TRIP and made the molten pool to have a chemical composition closer to TRIP, which reduced the hardening in the cooling process. The HAZ harnesses: For mild steel, the hardness not only did not decrease, as is commonly seen in welding regular steels, but became harder than the base material, probably due to alloying effect from TRIP and the high cooling rate effect. For TRIP780 and mild steel, the HAZ had very limited softening. For DP980, however, some softening occurred in HAZ that made it even softer than the pre-strained TRIP780. This may potentially cause the failure of TWB to occur in the DP980 side (local HAZ area) but not in the pre-strained TRIP780 side.

Tensile tests For specimens of TRIP780 with 10% pre-strain welded with DP980 and with the weld line in the transverse direction to the tensile axis, it is interesting to see that,


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

6.13 Typical microhardness values observed for AHSS-TWBs from: (a) TRIP780 (10% pre-strain)–mild steels; (b) TRIP780 (0% pre-strain)–DP980; (c) TRIP780 (10% pre-strain)–DP980 (courtesy of Sharma and Molian8).

140 Tailor welded blanks for advanced manufacturing



141

6.14 Tensile test samples of TRIP-10% pre-strain welded to DP980 with weld in transverse direction, showing that all fractures occurred in TRIP780 in the distance away from the welds (courtesy of Sharma and Molian8).

in Fig. 6.14, all the samples failed on the TRIP780 side and away from the weld line. Although the strain hardening effect in the welded zone may be eliminated by fusion welding, the resulting properties of the weld zone and its neighborhood would not be the same as its strain-free base metal and pre-strained base metal. Since the tensile necking occurs in the weakest portion of the specimen, this result is possible only when the welded zone has equivalent or higher strength than the pre-strained base TRIP780. In the case of equivalent strength, the pre-strained TRIP780 base metal has a lower remaining strain and will fail before the welded zone. In the case of higher strength, the necking will be in the weaker base metal side. This is consistent with the tensile result. The total elongations of the four welding configurations are given in Fig. 6.15, which shows that:

•

TRIP780 welded to mild steel has much higher elongation than when welded to DP980. It is understood that when the two welded constituents are stretched in series, the total elongation is contributed from each constituent, and the mild steel certainly makes more of a contribution to the total elongation than DP980. © Woodhead Publishing Limited, 2011

1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

142 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


6.15 Elongations obtained in TWBs consisting of TRIP780 without pre-strain (TRIP-0) or with 10% pre-stretching (TRIP-10), welded to mild steel or DP980 (courtesy of Sharma and Molian8).

•

Pre-strain of TRIP780 decreases the total elongation. Although the strain hardening effect in the welded zone may be eliminated by fusion welding that involves melting and solidification, the resultant properties in the weld zone would not be the same, and its neighborhood would not be the same as its strain-free base metal. These results are consistent with the above microhardness results, since no softening in the HAZ was observed.

6.4.4 YAG laser-welded TRIP/DP TWBs The mechanical properties and stretch-formability of YAG laser-welded TRIP/DP TWBs have been investigated by Nagasaka et al.27 Based on the report, the two coldrolled DP steels with the chemical compositions of (0.1–0.3)C-1.5Si-1.5Mn (mass%) and 0.14C-0.22Si-1.78Mn (mass%) were prepared, quenched (called MDP0) and further tempered (called MDP4), then tailor welded with TRIP steel using YAG laser. The formability was evaluated by a tensile test and by a punch stretching test using the maximum stretch height (H-max) as the formability index. There was no difference in H-max between the two DP heat-treating conditions (MDP0 and MDP4), though the tensile strength of MDP4 is significantly lower than MDP0. This was considered due to the contribution of the TRIP for the tailored blanks.

6.4.5 Diode laser-welded DP980 In the study by Xia et al.,10 laser welding of HSLA steel sheets and DP980 steel sheets (both are the same material welding) were compared, and the results of produced microstructures, microhardness testing across the welds, tensile tests, © Woodhead Publishing Limited, 2011


143

limit dome height tests, etc., are of interest for understanding diode-laser welding DP980 that may be used in TWB fabrication. The diode laser (a 4-kW diode laser was used in this study) usually provides a power density that is between traditional arc welding and CO2 or YAG lasers. The microhardness profiles across the weld line are shown in Fig. 6.16. It shows that for both DP980 and HSLA steels, the

6.16 Hardness profiles of (a) DP980 and HSLA at welding speed of 1.0 m/min and (b) DP980 and HSLA welds with different welding speeds indicated (courtesy of Xia et al.10).


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

144 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


welding resulted in an increased hardness within the fusion zone with limited welding speed effect, and at the same time resulted in a decreased hardness in HAZ, with the higher speed softening less and the softening zone appearing at the locations closer to the welding central line. The tensile behaviors of welded DP980 and HSLA are shown in Fig. 6.17. The tensile strength in the longitudinal direction is the weighted sum of the base metal

6.17 (a) Typical stress–strain curves for DP980 and HSLA with the welding speed of 1.3 m/min, with the fractured coupons shown in inset; (b) the ultimate strength vs. total engineering strain to failure at the welding speeds of 1.0, 1.3 and 1.6 m/min (courtesy of Xia et al.10, with minor modification).



145

and the weld region containing both hardening and softening zones, and the overall effects from their experiments showed as increased ultimate tensile strength (UTS) and reduced total elongation relative to the base metals (not shown). The tension in the transverse direction, on the other hand, is controlled by the weak part of the coupon, so the overall effects were the reduced UTS and total elongation, and the failure locations corresponded to the softened HAZ identified from hardness tests. For both HSLA and DP780 the welded microstructures (see Fig. 6.18) show the formation of plate-like martensite phase formation in the fusion zones that are

6.18 The microstructures of: (a) base HSLA, (b) fusion zone of HSLA, (c) base DP780, (d) fusion zone of DP780, (e) HAZ of welded DP780 (courtesy of Xia et al.10).


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

146 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


remarkably different from their base metals. However, in the HAZ (see Fig. 6.18e for DP780), martensite cannot be clearly identified, with a morphology that is neither in initial round particle shape nor in the welded sharp plate-like shapes. It is interesting to compare the formabilities of DP780 and HSLA steels, for both base metals and for laser-welded coupons, which are provided from LDH tests, as shown in Fig. 6.19. For DP780 base metal, the fracture was initiated slightly away from the dome apex and propagated always along the rolling direction (RD), while for the welded coupon the fracture always initiated from the

6.19 Comparison of failed specimens from limit dome height (LDH) test, for DP780 (a and b) and HSLA steels (c and d), and for base metals (a and c) and laser-welded specimens (b and d), respectively (courtesy of Xia et al.10).



147

outer area of the HAZ near the dome apex and propagated parallel to the weld line direction, regardless of the welding direction with respect to RD or TD, and the welding significantly reduced the LDH. This means that the welding process significantly affected (reduced) the formability. In contrast, for HSLA, in terms of crack path and limit dome height obtained there were no major differences between the base metal and the welded coupon, both demonstrated very good formability. Despite the similarities, the welding did affect the fracture location relative to the weld line. For the welded coupon the fracture initiated near the welding area (apparently in the HAZ) and the main crack propagated perpendicular to the weld line direction. Note that this is consistent with the previous limit strain sequence in uniaxial tension (see Fig. 6.17 for the tensile curves for HSLA), in which the limit strain was lower in the transverse direction than that in the longitudinal direction of the weld line. This is because a balanced biaxial stress state appears in the LDH test only at the dome apex, so a crack initiated at the early stage of the LDH test would result in a crack in the transfer direction, as that seen in welded DP780 (see Fig. 6.19 top right). But with the good limit strain of HSLA, the fracture occurred at a much higher LDH, at which the two principal stresses are not equal: the radial stress is larger than the hoop stress, resulting in the major crack to open in the radial direction, or the cracking in the perpendicular direction to the weld line. In the dome side region a secondary crack was initiated from the welded area at the side center, and propagated along the radial direction, again consistent with the major principal stress direction, now in the hoop direction.

6.4.6 Other studies on welding AHSS and applications of AHSS-TWBs Arc welding various AHSS of similar types Joining AHSS by gas metal arc welding (GMAW) was studied by Kapuska,13 with the objective of characterizing the effects of welding process conditions on the microstructure and mechanical properties of welded DP and TRIP sheet metal, with the focus on the effect on HAZ properties, since it is the weak part of the welds controlling the performance. The process conditions studied include material pre-strain, cooling rate (the result of welding heat input and fixture heatsinking), filler-metal selection, dilution of alloying elements and post-baking. The following findings were reported:

• • • •

For the TRIP600, no softening was observed in the welds. For the TRIP780, lower peak HAZ hardness than for DP780 was observed. The DP980 had the greatest peak HAZ hardness but also the highest degree of HAZ softening (relative to the base DP980), due to its higher hardenability and martensite content. Post-baking has no effect on HAZ hardness.


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

148 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

•

• •


The fusion zone microstructure and corresponding hardness was found to be largely influenced by dilution of the aluminum or silicon content of the base material. The high aluminum content of TRIP780 promoted the formation of harder constituents in the weld metal. All of the TRIP780 and DP780 butt-joints failed in the softened region of the HAZ during the static and dynamic tensile testing. Filler metal strength did not affect the static and dynamic tensile properties.

From the above findings it can be seen that the properties of the HAZ from arc welding of AHSS are determined by competition of the properties of the HAZ and the base metal, which are developed from different thermo-mechanical histories and with modified chemical compositions in the HAZ after melting and solidification. If the original material (base metal) has a relatively low volume fraction of hard martensite phase, such as that in DP600 and TRIP 780, the strength of the HAZ is less likely to be weakened, so the welding has less influence to the welded structure. On the other hand, for a material with a high strength and high martensite volume fraction, such as DP780 and DP980, the HAZ softening would occur due to reduced martensite volume fraction from heating and slow cooling rate. The local concentration of alloying elements will redistribute in the welding thermal cycle, which also causes properties to vary with respect to the base metal. With arc welding that uses fillers, additional alloying effects will occur, and the actual effects depend upon the fillers used and the formation of microstructure constituents under the welding conditions used. In comparison with laser welding, the arc welding, with relatively slower cooling rate, introduces more uncertainty and variation of weld properties from both melting/cooling rate and chemical composition aspects. Aluminum–AHSS TWBs Welding aluminum to AHSS was studied based on a finite element analysis (FEA) simulation by Padmanabhan et al.,28 to verify the advantages and formability of Al-AHSS TWBs using an internal code DD3IMP. To determine the formability characteristics of aluminum-steel TWBs, the aluminum AA6016-T4 welded with four steels was considered to form four different Al-steel TWBs. The steels include DP600 and three other steels, i.e. mild-steel (DC06) and two regular high-strength steels (AISI–1018, HSLA–340). A segmented blank holder was considered to allow the application of different forces on the aluminum and steel sheet segments in the TWBs. The formability and weld line movement were improved by using different blank holder forces across the weld line. In an earlier paper by the same group,29 the effect of anisotropy relative to the orientation of weld line on the formability of TWBs was studied for DC60/DP600 TWBs, also based on FEA simulation, in a half-squared box drawing process. In the simulation they used Hill’s 1948 formulation for anisotropic yield criterion, and the data of the base metal properties from literature, and discussed the effect



149

of blank texture orientation and welding direction combinations on the formability of TWBs. The results indicate that a proper combination of textured TWB can significantly improve the formability of the TWBs in the deep-drawing of the square cup. Friction stir welding (FSW) FSW was initially developed for welding aluminum alloys and was considered as an alternative joining process to the rivet joining for aircraft components. In recent years the process has been extended to a wide range of materials, including steels and polymers; see a review article by Nandan et al.12 One of the FSW studies on high and ultra-high-strength steels for automotive body assembly is reported by Miles et al.11 Although FSW is capable of generating superfine (submicron) grain sizes and considered as one of the approaches for producing nano-grained materials, at this stage the application of FSW for TWBs has not been reported. This is probably due to the very high requirement for TWBs from both strength and formability. The weldment properties directly produced from FSW without further treatments are generally weaker and more brittle than the base materials, due to the defects introduced from this severe deformation process.

6.5

Understanding the evolution of microstructure and its impact on properties of AHSS-TWBs

Due to the existence of dual or multiple phases in AHSSs that are produced from specially designed chemistries and thermo-mechanical processes, fusion welding of two dissimilar materials with different chemical composition will certainly cause changes of chemical composition of the fusion zone materials, as well as the thermo-mechanical history and microstructures for both fusion zone and HAZ. This alters the distribution of mechanical properties around the weld line and remote area of TWBs. Thus, it becomes a central issue to understand the evolutions of chemical compositions, microstructures and properties that occurred in the fabrication and forming of TWBs, and their relationship to the forming mechanics of TWBs.

6.5.1 Phase diagram A phase diagram defines, for a given state of chemical composite and temperature (under a normal pressure), the existing phase(s) (crystal structure) and constituent(s) (microstructure) of the system; the amount ratio or weight percent of these phases and constituents; and the microstructures formed during a heating or cooling condition. For an AHSS with various alloying elements, the multi-dimension equilibrium phase diagram is very complex. Considering the rapid heating and


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

150 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


cooling that appears in a TWB fabrication process, a non-equilibrium condition occurs and metastable intermediate phases may form. Despite the complexity of the problems, a basic Fe-C phase diagram and the effect of the major alloying element on the change of the Fe-C phase diagram can still be a basic foundation to understand the thermo-mechanical process, and modern simulation techniques in computational phase diagrams provide further opportunities to understand the trends of the alloy effects and the thermo-mechanical history effects. For TRIP steels, an aluminum alloying element is used as a ferrite phase stabilizer. A computed phase diagram of TRIP780 with 1.8w% Al (higher than nominal) by Kapustka et al.30 is shown in Fig. 6.20. Based on this phase diagram, a certain amount of ferrite phase (F) can exist and remain stable from room temperature all the way to melting temperature in heating and cooling. The austenite phase (A) is formed in a wide range of intermediate temperatures (700– 1400°C) and is fully or partially retained when cooling down to room temperature, in the form of retained austenite. This results in no discontinuous volume change during heating and cooling for this portion of constituent. The retained austenite is a metastable phase at room temperature, and further application of stress or strain will induce the martensite phase transformation.

6.20 Equilibrium phase diagram of TRIP780 with 1.8w% Al, calculated with Thermo-Calc™ software (courtesy of Kapustka et al.30).



151

Direct martensite transformation occurs for martensite steels (MS), DP steels and possibly regular carbon steels, depending on their chemical compositions and cooling rates (determined by the welding methods, welding parameters and blank geometries used). Lower cooling rates promote equilibrium phase transformation from austenite to ferrite. Both martensite and ferrite transformations involve a crystallographic structure change from face-centered cubic (fcc) to body-centered cubic (bcc), with an atomic packing density change from 0.74 to 0.68, resulting in a discontinuous volume expansion. However, in the case of martensite transformation a distorted bcc lattice forms (sometimes it is also considered as a tetragonal phase) with a slightly larger unit cell than that in a regular bcc ferrite phase. Superimposed over the discontinuous volume change from a phase transformation is the regular continuous thermal expansion/ contraction from atomic thermal vibration. Other than direct phase identification with X-ray diffraction patterns widely used in material processing studies, indirect methods such as monitoring volume changes during heating and cooling simulating the welding thermal cycle can be used. A comparative study on heating at a constant heating rate from room temperature to 1000°C and natural cooling of DP780 and TRIP780 was performed30 using a Gleeble test, closely monitoring actual measured temperature and dilatation (volume change), and the results are shown in Fig. 6.21. It shows that for DP780 there is a decrease in the actual cooling rate at 600°C, associated with the heat release in the austenite-martensite phase transformation, but for TRIP780 a relatively smooth temperature change was observed. The corresponding dilatations for the two steels were also obtained. For TRIP780 steel there is a regular thermal expansion/contraction during heating/cooling (except for a minor discontinuous expansion in heating at about 700°C), while for DP780 there is a superimposed volume contraction starting at 700°C during heating that is associated with both ferrite and martensite to austenite transformation, and a volume expansion starting at 600°C during cooling that is associated with austenite to ferrite and probably martensite transformation. Note that for both materials the volumes did not return to their original values, but with a certain reduction, which was more significant for DP780. This is an indication that the heating and cooling cycle introduced microstructure changes, possibly with reduced martensite volume fraction after the heating/cooling cycle (due to a lower density of martensite than that of ferrite).

6.5.2 Microstructures As described previously, various product phases and microstructures may form from welding that depend on material chemical composition, the welding process (especially the cooling rate) and geometry. Since the chemical compositions and thermal histories are generally not uniform in AHSS-TWB fabrication processes, neither is the microstructure distribution across the weld line from fusion zone


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

152 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


6.21 Comparison of DP780 and TRIP780 steels during heating to 1000°C at 25°C/s and natural cooling. (a) Measured temperature–time curves; (b) measured dilatation corresponding to the above temperature cycle (courtesy of Kapustka et al.30).

center to the proximity outside the HAZ and between the two sides of welding pairs (for dissimilar materials and perhaps also for dissimilar thicknesses of the same material). Such a non-uniform microstructure distribution feature has been shown earlier in Fig. 6.2 (DP600/DP600), Fig. 6.6 (DP590/IF), Fig. 6.18 (DP780/ HSLA). In the fusion zone the products of welding often consist of various phases at different volume fractions and grain size distributions, for example:

•

Mostly martensite phase, if original base metals tend to form martensite and welding is under a high cooling rate (often with laser), as evidenced by Rizzi et al.31 on laser welding of TRIP, DP and MS steels.



•

•

153

Ferrite phase dominating, if cooling rate is low or if there exist alloys that stabilize ferrite formation. Examples of this case include laser-welded Al-alloyed TRIP steel (Fe-1.5C-1.73Al-2Mn) by Xia et al.,32 who reported that ferrite was one of the predominant phases in the fusion zone due to the ferrite-stabilizing property of Al. Retained austenite phase dominating, for TRIP steels with a cooling rate that is not too fast.

In the HAZ the microstructures may remain the same as in the base metal but with significant grain size/shape changes due to the heating effect, or they may even change to different phases such as different martensite volume fractions. A high gradient of microstructure variation may also show in the proximity of the fusion zone. In lower strength steels such as DP600, austenite formation will be higher at closer proximity to the fusion zone and the percentage of martensite in the HAZ decreases with increased distance away from the fusion zone. This has been verified by Yan and Gallagher7 on DP590 materials. In the HAZ of higher strength materials such as M1310 steels, tempered martensite or ferrite could form, resulting in softened regions. When two materials of widely different alloy compositions are laser welded, microstructures that emerge depend upon the composition of the fusion zone. This has been reported by Yan and Gallagher7 who found the formation of Widmanstätten ferrite during welding DP590 with EDDS (an IF steel). Scanning electron microscope (SEM) micrographs are shown in Fig. 6.22(a–e). The DP HAZ indicates rather fine ferrite grains (average grain size is 4 µm) that decreased from original grain size, while the IF HAZ clearly shows large ferrite grains (average grain size is 18 µm) that are larger than the original size. In addition, the grains in the IF HAZ are very much elongated along the direction perpendicular to the weld line. The very large grain size with elongated shape in the IF HAZ was analyzed and explained by the limited nucleation rate and dominated grain growth process. The thermal history has a significant impact on a special excessive grain growth phenomenon for IF steels observed in welding, during which grain growth dominates nucleation for the γ → α phase transformation during cooling from the austenite temperature.

6.5.3 History of thermal profile in welding The temperature distribution during welding is important for analyzing the phase transformation and microstructure evolution. The temperature distribution over location and time during welding follows the basic heat transfer equation, but for welding thin sheets along a one-dimensional line at a given speed, the temperature distribution along the line in the transverse direction at the distance x away from the line center can be described by a simplified analytical equation, based on the derivation by Grong,33


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

154 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

Tailor welded blanks for advanced manufacturing [6.1]

where θ(x,t) and θ0 are the initial and final temperatures respectively, x is the distance from the fusion line, a is the thermal diffusivity, ρ is the density, c is the specific heat, v is the welding speed, Q is the weld power, d is the sheet thickness, and η is an efficiency factor of the process. This equation has been used by Bayraktar et al.34 for GTAW welding of IF steels; see Fig. 6.23 for some numerical solutions and a validation. The formula

6.22 Microstructures of the DP–IF welded zones: (a) total welded zone, (b) fusion zone DP side, (c) fusion zone IF side, (d) DP HAZ, and (e) IF HAZ (courtesy of Chatterjee et al.6).



155

6.23 Temperature profile in gas metal arc welding of IF steel: (a) calculated temperature θc(x, t) where x is the distance from fusion line; (b) measured θm(x, t); (c) comparison of calculated and measured values at x = 1.8 mm from fusion line. TIG welding, E = 0.9 kJ/cm, v = 40 cm/min (courtesy of Bayraktar et al.34).


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

156 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


6.24 Temperature profiles in the two HAZs of laser-welded IF (0.8 mm)– DP (1.6 mm) steels (courtesy of Chatterjee et al.6).

has also been used by Chatterjee et al.6 for IF–DP laser-welded TWBs; see Fig. 6.24 for the temperature profile over time at the HAZ in the IF and DP sides, respectively. The thermal modeling provides a numerical description of thermal history over locations across a weld line, which is very important information for predicting microstructure evolution and mechanical behaviors of TWBs. In the paper by Chatterjee et al.,6 the thermal analysis was able to provide a better understanding of microstructure evolution. The temperature profile produced in IF–DP laser welding resulted in excessive grain growth in the IF HAZ, with the elongated grain developed and the long axis aligned in the temperature gradient direction (i.e. perpendicular to the weld line). A high phase boundary mobility of IF steel was considered to contribute the grain growth in γ to α phase transformation. On the other hand, in the DP side, with a higher alloy content that reduces the mobility of grain/phase boundaries, greater thickness increases the heat transfer so the peak temperature and overall temperature over time were lower than that in the IF side (see Fig. 6.24). As a result, a fine-grain microstructure was produced in the DP HAZ. In summary, despite complexity of multi-phased microstructures and various dissimilar material combinations in AHSS-TWBs, the understanding of phase diagrams of each material system, the thermo-mechanical history during welding in fabricating TWBs, and the resulting microstructures and corresponding mechanical properties are critical to establishing a solid foundation in the understanding of the formability of TWBs. A systematic approach is helpful for design and manufacturing of AHSS-TWBs.



6.6

157

Other manufacturing processes related to AHSS-TWBs

6.6.1 Hot stamping of boron steels for auto parts requiring enhanced crash resistance Martensite steels are strong but difficult to form and so two steps of forming and heat treatment are commonly needed. Hot stamping combines the two steps in one operation and offers a solution to various problems for forming super-highstrength steels along the direction of today’s R&D trend of material processing. The hot stamping offers various advantages including greatly enhancing formability for high strength and lightweight materials (which often show a poor formability under traditional forming conditions), reduces or eliminates springback and reduces forming load for overcoming press load limitation. It also allows the formation of complex stamping part geometries or integrated parts, so that the total number of stamped parts and tools can be reduced (as a means of reducing tooling cost). In this process blanks are pre-heated to around 900–1000°C in an austenite temperature range, and quenched inside a water-cooled stamping die set to produce martensite-based microstructures. To prevent oxidation at elevated temperatures an aluminum-boron coating is commonly applied on the blanks. Based on two reports by Múnera et al.35 and Pic et al.,36 in order to meet safety requirements for a crash from front, side, or rear directions, and under a rollover condition without compromising vehicle weight, monolithic and TWB boron steels are under development in the steel industry. For example, at ArcelorMittal a two-step development program is under way. The first step is to use monolithic boron steels for analyzing vehicle crashworthiness performance of three stamping parts: side member, B-pillar reinforcement and front-end bumper. The second step is to consider a TWB solution. A comparison between monolithic solutions and TWB solutions was presented to determine the extra advantages (e.g. mass saving, energy absorption) that hot stamping TWB can offer. A new steel grade for the most demanding crash zones was under development that made it possible to significantly expand the application area of press-hardened TWBs, but no detailed information is available. Most likely the hot stamping TWBs need to use boron steel blanks of different thicknesses instead of using different materials other than boron steel, unless both parent materials can withstand elevated temperatures without severe oxidation. For fabricating boron steel TWBs, the feasibility of using a hybrid welding process with a laser in the leading position and the arc welding trailing was studied by Koganti et al.,37 for welding aluminized coated boron steel. In this process the laser welding can provide high speed, deep penetration and low heat input within the concentrated weld zone, but it requires good edge preparation without large gap variation. MIG welding uses filler material to bridge the gap and allows a large edge fit-up tolerance, but has much lower welding speed and higher heat


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

158 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


input, which tends to develop unfavorable microstructures. Taking advantages from both processes, this hybrid welding process uses a focused laser spot to create a vapor capillary ‘keyhole’ that increases the arc stability and welding speed. Aluminized coated boron steels (also referred to as USIBOR) are used for oxidation protection in hot stamping and corrosion protection in service. Although this study used the same material/thickness in welding, the results from the hybrid welding process and the boron steel welding are applicable and are briefly described here. In this study the butt welding (‘bead on plate’) of boron steels at 1.0 mm, 1.6 mm and 2.0 mm were studied and for lap welding, the boron steel sheets at 1.6 mm and 2.0 mm thicknesses were studied. Good weldability was achieved for all studied cases. However, based on tensile coupon tests and microhardness measurements, significant strength reductions of HAZs relative to the base metal were observed for all welding cases studied. The results indicate that the product strength is controlled by the last step welding process – the arc welding with filler that gave a lower cooling rate, even the process can increase welding speed significantly. The welded components, with a weak HAZ from this hybrid welding process, may raise severe strength concerns if used for final products; however, if used for producing boron steel TWBs for hot stamping, they may not be a problem. This is because the preheating and hot-deformation at the austenite temperature range may be able to eliminate the inhomogeneity of weld zone microstructure and chemical composition. At this time there is no literature available on forming tailored boron steel blanks. The formability of TWB boron steels in hot stamping deserves further investigation.

6.6.2 Tailor made blanks Although tailored blanks are commonly fabricated by line welding, they can also be made from other processes, such as tailor rolled blanks. Kleiner et al.38 reported the R&D activities involving some product and process developments at the University of Dortmund, including forming semi-final products with tailored blanks of ultra-high-strength steels. Lightweight constructions can only be achieved when all aspects of the technologies are integrated, including the use of the full potential of the material, the design of load-optimized component geometries, and the use of rational manufacturing procedures as part of an integrated process. A load-adapted material distribution is the key to successful lightweight components. Unlike that in bulk forming, in which a complex material distribution can be easily achieved with forging or extrusion processes, in sheet metal forming this is not applicable. In this paper a new approach to using tailor-made semi-finished products was presented to achieve cost-efficient production of weight and load-optimized workpieces. The process chain for using tailor rolled blanks to make lightweight components is shown in Fig. 6.25a. Variable wall thicknesses were manufactured



159

6.25 (a) Manufactured crash structure from tailor rolled blanks; (b) process chain for manufacturing of load-optimized profiles (courtesy of Kleiner et al.38).

by rolling (tailor rolled blanks), leading to a continuous thickness transition. The potential of forming tailor rolled blanks into a lightweight crash structure (e.g. car cross-member) was investigated (Fig. 6.25b). It was produced by a series of processes: flexible rolling to variable thicknesses, production of profile-shaped structural elements by means of air bending to manufacture semi-finished profile components, subsequent further processing by laser welding to close the profiles and finally three-roll-bending to the desired target radii. More details can be found in another paper from this group.39, 40

6.7

•

•

Conclusions

For energy and cost savings and for structure safety, TWBs have been widely used in sheet metal stamping. Advanced high-strength steels (AHSS) offer a more efficient way to redistribute mass/strength of structures with minimum weight, and AHSS-TWBs are under rapid development. There are three important aspects to the formability of AHSS-TWBs: (1) the thermal history during welding, determined by temperature distribution over time; (2) the microstructure evolution in the fusion zone and HAZ that is highly dependent upon AHSS chemistry and microstructure; and (3) the


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

160 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

•

•

• • • •


applied stress and strain field against the material’s intrinsic ductility and formability. Laser welding is a major approach for fabricating AHSS-TWBs, with various laser types applicable due to the faster thermal cycles and reduced weld and HAZ widths. Other manufacturing routes such as roll-made tailored blanks are also viable. The properties and formability of AHSS-TWB are the result of competing processes in a non-homogeneous tailored structure consisting of two parent base metals of different material types or dimensions, a welded seam and HAZs. Failure will occur at the weak link of this system among these constituents, which is a function of the materials used, the tailoring methods and processing parameters and the forming conditions. The most common methods to assess the properties of AHSS-TWBs (and others) are microhardness measurements across the weld seam, tensile tests, limit dome height tests and microstructural analyses. The enhancement of formability in AHSS-TWBs can be achieved by optimizing welding process design, AHSS material selection and tailoring design. The global stress and strain distribution and process design associated with TWBs is a mechanical problem to be covered by other chapters in this book. This chapter only addresses special issues related to AHSS. More research efforts are needed for AHSS-TWBs in order to achieve optimal manufacture of energy-efficient structures.

6.8

References

1. WorldAutoSteel, Advanced High-strength steel (Ahss) Application Guidelines, Version 4.1, World Steel Association, June 2009. 2. Bhadeshia, H. and R. Honeycombe, Steels: Microstructure and Properties. Oxford (UK): Butterworth-Heinemann, 2006. 3. Lamikiz, A., L. N. L. d. Lacalle, J. A. Sánchez, D. d. Pozo, J. M. Etayo et al. ‘Co2 Laser Cutting of Advanced High-strength steels (Ahss),’ Applied Surface Science, vol. 242, pp. 362–368, 2005. 4. Bagger, C. and F. O. Olsen, ‘Pulsed Mode Laser Cutting of Sheets for Tailored Blanks,’ Journal of Materials Processing Technology, vol. 115, pp. 131–135, 2001. 5. Han, T., S. Park, K. Kim, C. Kang, I. Woo et al., ‘Co2 Laser Welding Characteristics of 800 Mpa Class Trip Steel,’ ISIJ International, vol. 45, pp. 60–65, 2005. 6. Chatterjee, S., R. Saha, M. Shome, and R. K. Ray, ‘Evaluation of Formability and Mechanical Behavior of Laser-Welded Tailored Blanks Made of Interstitial-Free and Dual-Phase Steels,’ Metallurgical and Materials Transactions a-Physical Metallurgy and Materials Science, vol. 40A, pp. 1142–1152, 2009. 7. Yan, B. and M. Gallagher, ‘Strength and Fatigue of Laser Butt Welds for If, Hsla and Dual Phase Sheet Steels,’ presented at the International symposium on advanced highstrength steels for the ground transportation industry, Mater Sci Technol (MS&T), 2006. 8. Sharma, R. S. and P. Molian, ‘Yb:Yag Laser Welding of Trip780 Steel with Dual Phase and Mild Steels for Use in Tailor Welded Blanks,’ Materials & Design, vol. 30, pp. 4146–4155, 2009.



161

9. Farabi, N., D. L. Chen, J. Li, Y. Zhou, and S. J. Dong, ‘Microstructure and Mechanical Properties of Laser Welded Dp600 Steel Joints,’ Materials Science and Engineering a-Structural Materials Properties Microstructure and Processing, vol. 527, pp. 1215– 1222, 2010. 10. Xia, M., N. Sreenivasan, S. Lawson, and Y. Zhou, ‘A Comparative Study of Formability of Diode Laser Welds in Dp980 and Hsla Steels,’ Journal of Engineering Materials and Technology-Transactions of the ASME, vol. 129, pp. 446–452, 2007. 11. Miles, M. P., T. W. Nelson, R. Steel, E. Olsen, and M. Gallagher, ‘Effect of Friction Stir Welding Conditions on Properties and Microstructures of High Strength Automotive Steel,’ Science and Technology of Welding and Joining, vol. 14, pp. 228– 232, 2009. 12. Nandan, R., T. DebRoy, and H. Bhadeshia, ‘Recent Advances in Friction-Stir Welding – Process, Weldment Structure and Properties,’ Progress in Materials Science, vol. 53, pp. 980–1023, 2008. 13. Kapustka, R. N., ‘Effect of Material and Gmaw Process Conditions on Ahss Welds,’ Thesis (M.S.), Ohio State University, Columbus, 2006. 14. Fukumoto, S., H. Tsubakino, K. Okita, M. Aritoshi, and T. Tomita, ‘Friction Welding Process of 5052 Aluminium Alloy to 304 Stainless Steel,’ Materials Science and Technology, vol. 15, pp. 1080–1086, 1999. 15. Chen, C. M. and R. Kovacevic, ‘Loining of Al 6061 Alloy to Aisi 1018 Steel by Combined Effects of Fusion and Solid State Welding,’ International Journal of Machine Tools & Manufacture, vol. 44, pp. 1205–1214, 2004. 16. Uzun, H., C. D. Donne, A. Argagnotto, T. Ghidini, and C. Gambaro, ‘Friction Stir Welding of Dissimilar Al 6013-T4 to X5crni18–10 Stainless Steel,’ Materials & Design, vol. 26, pp. 41–46, 2005. 17. Bach, F., A. Beniyash, K. Lau, and R. Versemann, ‘Joining of Steel–Aluminium Hybrid Structures with Electron Beam on Atmosphere,’ Adv Mater Res, vol. 6–8., pp. 143–50, 2005. 18. Uematsu, Y., K. Tokaji, Y. Tozaki, and Y. Nakashimac, ‘Fatigue Behaviour of Dissimilar Friction Stir Spot Weld between A6061 and Spcc Welded by a Scrolled Groove Shoulder Tool,’ Procedia Engineering, vol. 2, pp. 193–201, 2010. 19. Taban, E., J. E. Gould, and J. C. Lippold, ‘Dissimilar Friction Welding of 6061-T6 Aluminum and Aisi 1018 Steel: Properties and Microstructural Characterization,’ Materials & Design, vol. 31, pp. 2305–2311, 2010. 20. Jiang, W. H. and R. Kovacevic, ‘Feasibility Study of Friction Stir Welding of 6061-T6 Aluminium Alloy with Aisi 1018 Steel,’ Proceedings of the Institution of Mechanical Engineers Part B – Journal of Engineering Manufacture, vol. 218, pp. 1323–1331, 2004. 21. Date, H., S. Kobayakawa, and M. Naka, ‘Microstructure and Bonding Strength of Impact-Welded Aluminium Stainless Steel Joints,’ Journal of Materials Processing Technology, vol. 85, pp. 166–170, 1999. 22. Watanabe, T., H. Takayama, K. Kimapong, and N. Hotta, ‘Joining of Steel to Aluminum Alloy by Interface-Activated Adhesion Welding,’ Thermec’2003, Pts 1–5, vol. 426–4, pp. 4129–4134, 2003. 23. Tsujino Jiromaru, Hidai Kazuaki, Hasegawa Atsushi, Kanai Ryoichi, Matsuura Hisanori et al., ‘Ultrasonic Butt Welding of Aluminium, Aluminium Alloy and Stainless Steel Plate Specimens,’ Ultrasonics, pp. 371–4, 2002. 24. Danesh, M. H. and T. A. Karimi, ‘Bond Strength and Formability of an Aluminium– Clad Steel Sheet,’ J Alloy Compd, pp. 138–143, 2003.


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

162 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


25. Matsugi, K., Y. Wang, T. Hatayama, O. Yanagisawa, and K. Syakagohri, ‘Application of Electric Discharge Process in Joining Aluminum and Stainless Steel Sheets,’ Journal of Materials Processing Technology, vol. 135, pp. 75–82, 2003. 26. Bayraktar, E., D. Kaplan, F. Schmidt, H. Paqueton, and M. Grumbach, ‘State of Art of Impact Tensile Test (Itt): Its Historical Development as a Simulated Crash Test of Industrial Materials and Presentation of New “Ductile/Brittle” Transition Diagrams,’ Journal of Materials Processing Technology, vol. 204, pp. 313–326, 2008. 27. Nagasaka, A., K. I. Sugimoto, M. Kobayashi, K. Makii, and S. Ikeda, ‘Press Formability Yag Laser Welded Trip/Dp Tailored Blanks,’ Journal De Physique Iv, vol. 115, pp. 251–258, 2004. 28. Padmanabhan, R., M. C. Oliveira, and L. F. Menezes, ‘Deep Drawing of AluminiumSteel Tailor-Welded Blanks,’ Materials & Design, vol. 29, pp. 154–160, 2008. 29. Padmanabhan, R., A. J. Baptista, M. C. Oliveira, and L. F. Menezes, ‘Effect of Anisotropy on the Deep-Drawing of Mild Steel and Dual-Phase Steel Tailor-Welded Blanks,’ Journal of Materials Processing Technology, vol. 184, pp. 288–293, 2007. 30. Kapustka, N., C. Conrardy, S. Babu, and C. Albright, ‘Effect of Gmaw Process and Material Conditions on Dp 780 and Trip 780 Welds,’ Welding Journal, vol. 87, pp. 135s–148s, 2008. 31. Rizzi, P., S. Bellingeri, F. Massimino, D. Baldissin, and L. Battezzati, ‘Microstructures in Laser Welded High-strength steels,’ presented at the The 13th international conference on rapidly quenched and metastable materials, 2009. 32. Xia, M., Z. Tian, L. Zhao, and Y. Zhou, ‘Fusion Zone Microstructure Evolution of Alalloyed Trip Steel in Diode Laser Welding,’ Mater Trans, vol. 49, pp. 746–753, 2008. 33. Grong, Ø., Metallurgical Modelling of Welding. London: The Institute of Materials, England, 1994. 34. Bayraktar, E., D. Kaplan, L. Devillers, and J. P. Chevalier, ‘Grain Growth Mechanism During the Welding of Interstitial Free (If) Steels,’ Journal of Materials Processing Technology, vol. 189, pp. 114–125, 2007. 35. Múnera, D. D., L. Lacassin, and F. Pinard, ‘Very and Ultra High-strength steels Based Tailored Blanks: A Step Further Towards Vehicle Crash Performances Improvement,’ Revue De Metallurgie-Cahiers D Informations Techniques, vol. 104, pp. 613–624, 2007. 36. Pic, A., D. D. Múnera, L. Cretteur, F. Schmit, and F. Pinard, ‘Innovative Hot-Stamped Laser Welded Blank Solutions,’ Stahl Und Eisen, vol. 128, pp. 59–66, 2008. 37. Koganti, R., S. Angotti, A. Joaguin, and E. Stiles, ‘Laser Hybrid Welding Joining of Aluminized Coated Boron Steel for Automotive Body Construction,’ in SAE vol. 08M–189, ed, 2008. 38. Kleiner, M., S. Chatti, and A. Klaus, ‘Metal Forming Techniques for Lightweight Construction,’ Journal of Materials Processing Technology, vol. 177, pp. 2–7, 2006. 39. Kleiner, M., S. Chatti, B. Heller, R. Kopp, C. Wiedner et al., ‘Umformung Und Weiterverarbeitung Von Flexibel Gewalzten Stahlblechen (Tailor Rolled Blanks) Fur Leichtbaustrukturen,’ November 2002. 40. Chatti, S., B. Heller, M. Kleiner, and N. Ridane, ‘Forming and Further Processing of Tailor Rolled Blanks for Lightweight Structures,’ in Advanced technology of plasticity: Proceedings of the Seventh Internatioranal Conference on Technology of Plasticity (ICTP), Yokohama, Japan, 2002, pp. 1387–1392. 41. Malakondaiah, G., M. Srinivas, and P. Rama-Rao, ‘Ultrahigh-strength low-alloy steel with enhanced fracture toughness,’ Prog. Mater. Sci., vol. 42, pp. 209–242, 1997.



163

42. Ribolla, A., G. L. Damoulis, and G. F. Batalha, ‘The use of Nd:YAG laser weld for large scale volume assembly of automotive body in white,’ Journal of Materials Processing Technology, vol. 164–165, pp. 1120–1127, 2005. 43. Jeswiet, J., M. Geiger, U. Engel, M. Kleiner, M. Schikorra et al., ‘Metal forming progress since 2000,’ CIRP Annals, Journal of Manufacturing Science and Technology, vol. 1(S), pp. 2–17, 2008. 44. Sharma, R. S. and P. Molian, ‘Yb:YAG laser welding of TRIP780 steel with dual phase and mild steels for use in tailor welded blanks,’ Materials & Design, vol. 30, pp. 4146–4155, 2009.


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

7 Tailor welded blanks for the automotive industry B. L. KINSEY, University of New Hampshire, USA

Abstract: One of the main users of sheet metal components capitalizing on the advantages of tailor welded blanks (TWBs) is the automotive industry. These advantages include reductions in weight, cost and noise with simultaneous increases in crashworthiness, dimensional accuracy and corrosion resistance. However, there are formability concerns that are present due to the welding process. In this chapter, these advantages and disadvantages are discussed along with TWB forming methods, welding processes and materials used in the automotive industry. Specifics on many of these topics are covered in more detail within other chapters of this book. Key words: tailor welded blanks, automotive industry.

7.1

Introduction

The competitive and dynamic nature of the automotive industry requires that companies continuously improve their products and operations. Due to customer demands and governmental regulations (e.g., Corporate Average Fuel Economy [CAFE] standards in the United States), there is a concerted focus on reducing the vehicle weight for environmental and performance reasons. As for potential environmental benefits, each 68 kg (150 lb) of weight reduction creates a 0.43 km/ liter (1 mpg) fuel efficiency improvement in the vehicle (Brooke and Evans, 2009). In addition, manufacturing cost is always a critical concern when producing a mass production product. Sheet metal stamping is one of the main manufacturing processes to create structural and ‘body in white’ components for automobiles due to the high production rates, efficient material utilization and low manufacturing costs. To produce such weight improvements using more advanced lighter weight materials, which are more expensive, is not always the solution as the material cost can be approximately 80% of the final part cost in sheet metal forming applications (Natsumi et al., 1991). Finally, performance parameters such as crashworthiness and structural rigidity are critical requirements for both customer satisfaction and government regulations. In summary, automakers must devise innovative solutions to reduce manufacturing costs, improve product performance and reduce vehicle weight. One such design and manufacturing technology to achieve these seemingly conflicting goals is the tailor welded blank (TWB). Figure 7.1 shows various TWB applications in an automobile. In the 1990s through early 2000s, TWBs were a considerable topic of interest; thus, many classic references are from this 164 © Woodhead Publishing Limited, 2011

Tailor welded blanks for the automotive industry

165

7.1 Exploded view of current and/or potential TWB body components (courtesy of TWB Company, LLC).

period and are highlighted in this chapter. Such work led to a significant increase in the use of TWBs with production rates doubling to 110 million parts from 2000 to 2005 (Das, 2000; Rooks, 2001). Current production rates worldwide are estimated to be 200–250 million laser welded blanks per year (Vanker, 2010). Note that while automotive applications are highlighted here, other industries such as aerospace, manufactured appliances, etc., also incorporate TWBs into their designs to obtain similar benefits. Typically, to create e.g. a structural member for an automobile chassis, several sheet metal stampings are formed separately and then welded together to produce the component. This allows the sheet metals to be varied so that the material properties and characteristics are located in the areas of the structural member where required. For example, for a door inner component, stronger material is required near the hinge to support the weight of the door when the door is opened. However, the material near the door handle can be lighter weight for mass reductions and so less force is required to close the door. As opposed to forming multiple components and subsequently welding these stampings together to create the structural member, flat sheet metal blanks (i.e., workpieces) can be welded together prior to the forming process. Thus only one forming operation line is required to fabricate the component. This is referred to as a TWB, as the designer is able to tailor the location of the material properties within the final structural member where desired. These material differences can be with respect to the grade/strength of the material, the thickness of the sheet metal, or the coating, e.g., galvanized versus ungalvanized. Figure 7.2 shows schematically the concepts of both traditional fabrication of an assembly by


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

166 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


7.2 Flowcharts for fabrication of a door inner by (a) a conventional process and (b) a TWB process.

forming multiple components and then welding these pieces (see Fig. 7.2a) together and TWB fabrication where multiple blanks are welded together and then subsequently formed (see Fig. 7.2b).

7.2

Door inner example

Some of the automotive applications for TWBs include body side panels (see Fig. 7.3), motor compartment rails, center pillar inner panels and wheelhouse/ shock tower panels. As mentioned previously, a ubiquitous example of a TWB is a door inner (see Fig. 7.4). As an example of the potential cost savings available for a TWB component, General Motors (1991) estimated that a TWB door inner saved the company $4.9 million in U.S. dollars through the elimination of fourteen dies, weld fixtures and check stands (American Machinist, 1992). This estimate was even conservative, as other costs such as reductions in die storage, materials handling and die maintenance were not considered. Another technical cost analysis showed a saving of $6.3 million for the manufacturing of a separate hypothetical TWB door inner (Trogolo and Dieffenbach, 1998). To provide an example of the potential structural improvements for the implementation of a TWB component, an automotive manufacturer reported a 9% improvement in the elastic sag and 30% improvement in the plastic set when a 90 kg load was applied to the latch of an assembled TWB door (Irving, 1995).



167

7.3 A body side TWB panel (courtesy of TWB Company, LLC).

7.4 A door inner TWB panel (courtesy of TWB Company, LLC).

7.3

Historical perspective

Prompted by the 1973 oil embargo, the United States Congress enacted the Energy Policy and Conservation Act in 1975, which established a Corporate Average Fuel Economy (CAFE) standard for the fleets of automakers. Today, the US National Highway Traffic Safety Administration (NHTSA) also considers greenhouse gas emissions, which lead to global climate change, and the nation’s need to conserve energy when it sets the level of the CAFE requirement. The CAFE standard was instrumental in more than doubling the fuel economy level for US cars from 1973 to 1988, from an average of 14.2 mpg to 28.6 mpg respectively (DeCiccio, 1995). But only modest increases in fuel economy have been achieved since then, e.g., 32.6 mpg in 2009 (US Department of Transportation, 2009). These fuel economy gains were obtained through technological advances to produce a more fuelefficient automobile and by weight reductions through part optimization and the


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

168 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


increased use of plastics and other lightweight metal components. To continue this trend of improved fuel economy, technologies such as TWBs have been implemented, beginning in the 1980s. In addition to weight savings, TWBs have demonstrated other essential advantages in the competitive global automotive market, such as improved part consistency, reduced manufacturing cost, etc., which will be covered in the following section.

7.4

Advantages of tailor welded blanks (TWBs)

TWBs offer numerous advantages that will assist automakers in meeting tighter fuel economy standards while also improving other key sheet metal part characteristics. Thus, automakers attempt to incorporate TWBs into their vehicles whereever possible. For example, the Chrysler Pacifica uses eight TWBs for large body-side, door and lift gate inners to achieve various benefits (Trem, 2004).

7.4.1 Reduced vehicle weight TWBs lower vehicle weight through reinforcement part elimination, part consolidation and material overlap reduction for the welding process. As mentioned previously, reduced vehicle weight is essential for automakers to meet ever-constricting CAFE regulations. Front and rear rails reduced the mass in the Chevy Malibu by 2.5 kg (Trem, 2004). For other examples from the past, a motor compartment rail in the 1992 Seville saved General Motors 1.3 kg in total vehicle weight (American Machinist, 1992), and a hypothetical door inner analysis estimated a potential weight saving of 1.5 kg (Trogolo and Dieffenbach, 1998). These weight savings may seem insignificant; however, if several TWB components are incorporated into a vehicle, these reductions will make a considerable contribution to improve the fuel economy of the automobile.

7.4.2 Reduced manufacturing costs TWBs offer a significant cost saving due to reduced manufacturing costs and improved material utilization. Referring again to Fig. 7.2, there are several cost savings apparent from a comparison between the fabrication flowcharts for a conventional process and a TWB process. First, the number of required production dies is significantly reduced for the TWB process. Lexus reduced the number of required production dies from twenty to four for a side body panel when it switched to a TWB (American Machinist, 1992). The cost of these dies, while in the hundreds of thousands of dollars, is not a major source of piece price since the tooling is amortized over the life of the product. However, the scrap that is produced in each forming process significantly contributes to the piece price. Thirty to fifty percent of sheet metal purchased ends up as scrap (Venkat et al.,



169

7.5 Amount of material overlap for (a) spot, (b) mash seam and (c) laser welding processes (adapted from Wang et al. 1995).

1997), and as previously mentioned the material accounts for 80% of the cost of stamped parts (Natsumi et al., 1991). Also, it may be possible to use the material trimmed off during blanking from one forming operation as one of the blanks to produce a TWB for another application (Ogmen, 1995). Figure 7.5 shows the amount of material overlap necessary for a spot weld used in a conventional fabrication process and mash seam and laser welds used to fabricate TWBs. This further demonstrates the advantage of improved material utilization through TWBs. In addition, elimination of fixturing for spot welding and quality control provides for additional cost savings. It should be noted however that cost savings from TWBs may be initially offset by the need to purchase welding equipment to produce TWBs or bring a TWB supplier on-line, increased start-up quality control activity and the need to make changes in material handling equipment to accommodate TWBs. However, the long term cost savings overshadow these concerns. Also, note that the references with respect to cost savings are primarily from the 1990s, when TWBs were a new technology that was being considered for implementation.

7.4.3 Improved crashworthiness and part stiffness The structural integrity of the automobile is also improved through the use of TWBs. This is due to the fact that mash seam and laser welds are generally stronger than the spot welds used in the conventional component assembly process. In experiments with axially loaded TWB rails produced with mash seam and laser welds, the energy absorption of the TWB rail was greater than for a comparable spot welded rail, without a penalty of increased peak load (Wang et al., 1995). In addition, no failures occurred in the mash seam or laser welds, while two out of sixteen samples failed for the spot welds. Similar results were obtained under bending loads as well. Therefore, improved crashworthiness (and corresponding part stiffness) is obtained through the use of TWBs.


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

170 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


7.4.4 Decreased noise Due to the improved structural integrity of the TWB components, reductions in noise are also achieved through improved vibration dampening. For a BMW 1-Series firewall, the noise was reduced by 0.9 dB while also obtaining a mass saving of 5.1% with no detrimental effects to crashworthiness (Florentin et al., 2009; Smock 2009). Using TWBs in such an application would prevent the necessity of expensive damping sheet treatments, which are commonly used to reduce the amplitude of vibrations and noise. Thus, the cost savings from TWB applications are often beyond what is measured simply by material savings.

7.4.5 Improved dimensional accuracy and consistency Spot welding inherently has large dimensional inaccuracies and variations due to clamping-induced part distortion and due to location/fixturing error from tool wear-out during the welding process. Welding prior to the forming process on flat blanks is a simpler and more consistent process. Even if there were no other advantages to using TWBs, improvements in fit and finish of the final product from 0.5 mm for spot welding to 0.08 mm for laser welding, one of the preferred welding methods for TWBs, would provide enough justification to make the transition (Irving 1995). These fit and finish improvements would have a considerable effect on the quality of the assembled vehicle. (Note that the shear edge tolerances for laser welded joints are much more stringent than for spot welded joints; however, these values are more readily achieved in a precision blanking operation to assure the desired fit and finish.)

7.4.6 Improved corrosion resistance The longevity of an automobile is a mark of quality and is used as a major selling point in today’s competitive automobile market. By eliminating the lap joints of spot welds through the consolidation of reinforcements, the corrosion resistance of the body increased substantially due to the sealing of seam welds over spot welds (Kubel, 1997; Scriven et al., 1996).

7.5

Disadvantages of TWBs

As is often the case, there are trade-offs with several formability and processing obstacles to overcome when implementing TWBs.

7.5.1 Decreased formability Changes in the mechanical properties of the material due to welding affect the formability of TWBs. In particular, the potential elongation prior to tearing failure



171

is significantly reduced. Typically for steel, the ductility of the material in the weld decreases while the strength of the weld increases, and there is no distinct heat-affected zone (HAZ). Several researchers (Saunders and Wagoner, 1996; Scriven et al. 1996) have measured reductions in the elongation both longitudinal and transverse to the weld line of approximately half what is found for the base material. More recent work has focused on such reductions in formability for TWBs fabricated with advanced high-strength steels (AHSS) (e.g., Kapustka et al., 2008; Sharma et al., 2009; Farabi et al., 2010) due to the interest in this class of materials for further weight reductions. For aluminum, the weld will generally have a slightly lower ductility than the base material; however, significant reductions in the HAZ elongation are common. Venkat et al. (1997) and Stasik and Wagoner (1996) investigated aluminum alloys 5754-O and 6111-T4, which have the combination of being both weldable and formable. Reductions of 20% and 75% for strains longitudinal and transverse to the weld line respectively were reported for Al 6111-T4 for both CO2 and YAG laser welding. These decreases in the HAZ for aluminum are due to vaporization of strengthening elements from the HAZ, overaging due to weld heat and solidification cracking due to liquidization and re-solidification (Venkat et al., 1997). Furthermore, Wu et al. (2004) conducted various material tests (i.e. optical microscope, atomic force microscope, energy dispersive spectroscopy, microindentation and nanomechanical tester) for aluminum TWBs. This reduced elongation effect was attributed to softer material adjacent to the grain boundary (as opposed to the center of the grain), which was discovered during nanoindentation tests with in situ AFM imaging. Al 5182-H00 (AlMg5Mn) is often used for automotive TWB applications such as door inners and floor pans (Auto/Steel Partnership, 1995; Vollertsen et al., 1996) due to its superior deep drawability and weldability. Kridli et al. (2000) conducted tensile tests with Al 5182-H00 on standard size TWB specimens with transverse weld lines. Also, miniature specimens that consisted of only the gas tungsten arc welding (GTAW) material to represent the potential formability (i.e., strain) of the weld material loaded longitudinally were tested. Similarly, Viswanathan et al. (2001) fabricated TWBs from Al 5182-H00. The welding was performed with a 3 kW YAG laser and a welding speed of 100 mm/ sec. Forming limit curves (FLCs) were created for the Al 5182-H00 base material and TWBs based on limit dome height tests. The potential plane strain elongation for the base material and TWBs was found to be 22% and 6%, respectively (Viswanathan et al., 2001). More recently, Moulten and Weckman (2010) developed a double-sided arc welding process for Al 5182-H00 and reported no decrease in strength for the weld area but a decrease in the ductility. It should be noted that results of formability testing on aluminum, as well as steel, are highly dependent on the alloy (or grade) tested and welding parameters such as weld speed, heat applied, welding process used (e.g. CO2 laser welding, YAG laser welding, mash seam welding, electron beam and TIG welding) and whether a filler metal is used.


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

172 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


Also, forming parameters affect the results for TWB applications. For example, Davies et al. (2002) conducted tensile tests of 5000-series Al weld material at superplastic forming temperatures and strain rates. The results showed that while increased ductility in the weld material was possible, a disparity still existed between the base material and the weld material for this Aluminum alloy. In addition to creating TWBs with similar materials, friction stir welding has been used to create blanks with dissimilar materials. For example, Shigematsu et al. (2009) created a TWB with Al 5052P-O aluminum and AZ31B-0 magnesium alloys and investigated the strength and elongation for various welding parameters. For the best case, a 30% reduction in the tensile strength and a 2% reduction in the potential elongation were obtained. Similarly, investigations of welding steel with aluminum blanks have been conducted (Padmanabhan and Oliveira, 2008; Chen and Kovacevic, 2004).

7.5.2 Weld line movement As was previously indicated, a popular application of TWBs is to weld materials with different strengths, whether obtained from variations in the material thickness or grade, together to eliminate reinforcements and improve material utilization. A forming problem is created by this practice if strain transverse to the weld line is present. The material deformation of the process will concentrate in the thinner, weaker material causing the weld line to ‘move’ towards the thicker, stronger material. Weld line movement is affected by the material draw-in from the binder area (e.g., by varying the draw bead in the process [Heo et al., 2001]) and the initial weld line position and part geometry (Choi et al., 2000). Weld line movement creates several concerns. First, the deformation concentration in the weaker material may cause tearing failure due to the excessive strain created and potential elongation reductions in the HAZ, particularly for aluminum TWBs. Another concern of weld line movement is die wear as the harder weld material is displaced across the tooling surfaces (Auto/Steel Partnership, 1995). In addition, the weld line movement combined with a step in the binder area to accommodate the varying thickness of the TWB may cause a tearing concern if the step prevents the weld line movement (see Fig. 7.6b). Finally, weld line movement limits the ability of the automotive designer to position the material properties in the stamping where desired.

7.5.3 Wrinkling in die addendum As mentioned previously, steps in the binder force tooling are necessary to accommodate TWBs with thickness differences. Since clearance is required in the stepped binder area for weld line movement, an unconstrained material condition exists which could lead to wrinkling in the die addendum (Auto/Steel Partnership, 1995) (see Fig. 7.6a for a schematic of this concern). Also, wrinkling may be



173

7.6 Schematic of weld line movement causing (a) wrinkling and (b) tearing due to the stepped binder area.

produced from the weld line movement itself as positive strain will not be created in the thicker, stronger material (Nakagawa et al., 1993; van der Hoven and Ostyn, 1994). This can be addressed through the use of a segmented binder (Siegert and Knabe, 1995) to apply varying perimeter forces on the TWB. Kinsey et al. (2004) developed a methodology to determine the segmented binder force ratio for a TWB application to reduce wrinkling and also a means to quantify the wrinkling behavior.

7.5.4 Other TWB concerns Other operations in the forming process may have difficulties with the harder weld line area and different thicknesses of TWBs as well, e.g. trimming, flanging and hemming (Auto/Steel Partnership, 1995). However, deep drawing is the most critical of the forming operations for formability concerns such as tearing and wrinkling. Another concern with TWBs is springback. Reduced plastic strain in the thicker, stronger material will lead to increased springback. Also, uneven springback between the thicker, stronger material and the thinner, weaker material will create distortion in the stamping. However, the uneven springback can be greatly reduced if springback compensation is designed into the tooling. Finally, process concerns such as destacking and blank transfer of TWBs will need to be addressed (Auto/Steel Partnership, 1995).


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

174


7.6

Research efforts for TWBs

Due to the industrial interest in TWBs, substantial research has been conducted in several areas. For example, improvements to the welding process and various welding methods have been studied to minimize material property changes and produce consistent weld properties. The potential elongation of the weld and the change of ductility are affected by the material used. Minimizing these effects to produce a consistent deformation pattern in the final part is desirable. Research in this area has included the significance of welding parameters on the formability of TWBs (Eisenmenger et al.; 1995; Bhatt et al., 1995), analytical and numerical simulations of weld line properties (Doege et al., 1996) and the effect of post weld processing, e.g. hot and cold planishing, on the formability of TWBs (Lee et al., 1996). Sato et al. (2004) compared friction stir welding to GTAW for three alloys (5182-O, 5754-O and 6022-T4). These materials and this welding method are currently used in automotive applications. They found that the 5xxx series alloys had similar tensile ductility and formability regardless of the welding process used. However, for the 6022-T4 sheets joined with friction stir welding, the potential elongation increased due to less softening in the heat-affected zone compared to the GTAW. Other research has concentrated on the sheet metal itself, investigating the formability of TWBs created from popular grades and alloys of material as well as searching for materials that are less susceptible to the potential negative effects of welding. In addition, research on numerical and analytical modeling of TWB forming has been conducted as discussed in Chapters 3 and 4 of this book. Finally, studies have investigated alternative processes to improve the formability of TWBs. The following section highlights some of these alternatives.

7.7

TWB forming methods

7.7.1 Strategic locating of weld line on part As was mentioned previously, reductions in the elongation longitudinal to the weld line are created due to material property changes in the weld. A common design practice to address this concern is to orient the length of the weld line in a location where this longitudinal strain is not large enough to cause failure (Auto/Steel Partnership, 1995; Ahmetoglu et al., 1995). This however limits the ability of the automotive designer to take full advantage of material utilization improvements offered by TWBs.

7.7.2 Non-uniform binder force To address the concern of transverse strains to the weld line causing a deformation concentration in the weaker material and weld line movement for TWBs, more material draw-in is allowed through the die addendum area of the stronger



175

material. Thus the strain level in the weaker material is reduced since weld line movement does not occur. This increased material draw-in can be obtained through modification of the draw bead geometry, if utilized in the process, or through a non-uniform binder force such as generated by a segmented binder (Pepelnjak et al., 1998; Ninforge and Dawance, 1998; Siegert and Knabe, 1995) or multiple, independently controlled binder force mechanisms. Ahmetoglu et al. (1995) investigated forming round cups with TWBs. In their work, six nitrogen cylinders produced the binder force for a circular blank holder plate, three under the thicker, stronger material half and three under the half with the thinner, weaker material (see Fig. 7.7). The nitrogen cylinders varied the blank holder force applied to the two different material gauges. The thicker material was subjected to a lower blank holder force, thus allowing more material to flow into the die cavity. This process modification was successful at reducing the weld line movement and delaying tearing failure along the weld line compared to the case where a uniform binder force was applied to the TWB. This practice is also not ideal, however, since only limited strain will be produced in the stronger material. Also, determining the values of the non-uniform binder force that will lead to no weld line movement currently must be obtained through a time consuming trial and error process. Finally, this concept will only work for TWBs where the stronger material is on one side of the TWB blank and the thinner material is on the opposite side. Consider the door inner geometry of Fig. 7.2. Here, the non-uniform binder force method could be used because the stronger material is on the hinge side of the door and the weaker material is on the opposite side. However, the straight linear weld line of the door inner shown in Fig. 7.1 not only places the stronger material at the hinge of the door where desired but also near the top of the door inner. This straight linear weld line is necessary to control the material draw-in over the entire hinge side of the door, but does not optimize on material utilization since the stronger material is not required near the

7.7 Application of non-uniform binder force to improve TWB forming (adapted from Ahmetoglu et al. 1995).


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

176 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


top of the door. Thus, the non-uniform binder force method is not ideal with respect to material utilization. Concerning a different component such as a floor pan, the stronger material is located in the center of the part; therefore, the nonuniform binder force could not assist in the forming of this particular application whatsoever.

7.7.3 Means to clamp on weld line In another effort to improve TWB formability, Munzen (1997) used the lower die cushion from a production triple action stamping press to clamp on the weld line during the forming process (see Fig. 7.8). The die cushion supported the weld line when the blank was initially placed across the die cavity. The part of the punch opposite the die cushion would then contact the blank prior to or just as deformation of the blank occurred. Once clamped, the part was formed entirely before the weld line was released. This clamping prevented the weld line from moving during forming. However, this process required that the weld line be at the highest point on the stamping in order for the clamping to occur successfully. Also, it was not possible to release the weld line prior to the end of the forming process, which is desirable for part integrity since it creates a limited amount of strain in the material. Kinsey and Cao (2001) used a separate system other than the die cushion to clamp the weld line during deformation. A 20% increase in the potential depth of a formed part and a 32% reduction in the strain perpendicular to the weld line were achieved with this technique without the drawbacks listed above for the die cushion method.

7.8 Sequence of forming process using lower die cushion to clamp on the weld line (adapted from Munzen, 1997).



7.8

177

Welding processes for TWBs

The most popular processes for welding TWBs are friction stir, mash seam and laser welding. Friction stir and mash seam welding are solid state forging processes, and laser beam welding is a fusion process. Historically, European automakers seem to prefer solid state welding while the Japanese and United States automakers use laser welding more frequently (Irving, 1995). As a general rule, mash seam welds are acceptable for non-exposed, non-sealing components, and laser welds are required for all other applications (Kubel, 1997). There are advantages and disadvantages to each of these welding methods. The HAZ of the laser weld is smaller than the mash seam weld; therefore, the decreased formability discussed previously may be less severe in laser welds. However, research has shown equivalent formability of the two weld types if the mash seam weld is hot planished, i.e. post weld heat treating (Bhatt et al., 1995; Saunders and Wagoner, 1996). Conversely, cold planishing, i.e. cold working to diminish the material build up found in mash seam welds, severely decreases the formability of TWBs (Lee et al., 1996). A disadvantage of laser welds is the necessity for a straight, precision-sheared edge for fit up of the blanks prior to the welding process (Kubel, 1997). While the edge straightness of the steel coil from the mill is currently almost accurate enough to meet requirements, i.e., 0.05 mm deviation, conventional shearing methods may not be adequate (Natsumi et al., 1991). Also, the initial start up cost of laser welding is more expensive than mash seam welding. However, as the production volume increases, the cost between mash seam and laser welding becomes competitive (Baron, 1997). In general, neither welding system is consistently better for all product designs. The choice of welding method should be made on a part-by-part basis (Baron, 1997). For aluminum TWBs, laser welding is difficult because of high surface reflectivity and thermal conductivity (Kubel, 1997). Therefore, the power density of the laser required is higher than that for steel welding (Venkat et al., 1997; Vollertsen et al., 1996). Aluminum welding is also more sensitive than steel welding to process conditions and so parameters must be kept constant to avoid pores in the weld, which lead to failure (Nagel et al., 1997; Vollertsen et al., 1996). YAG laser welding is more popular than CO2 welding for aluminum because it is less sensitive to process variations.

7.9

Materials used to produce TWBs

Various materials are used for TWBs, including various steel, AHSS, aluminum and other lightweight alloys. The specific effects with respect to fabricating and forming TWBs are material dependent. For example, in general, the weld area in steel TWBs is stronger than the base material due to microstructural changes. Alternatively, for aluminum TWBs, the weld material is typically weaker than the


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

178 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


base material. This will affect the general deformation behavior of TWBs produced from these materials. In addition, the most effective welding method is material dependent. For example, due to the reflectivity of aluminum, laser welding is less effective than it is for steel TWBs. Chapter 5 on light weight metal alloys and Chapter 6 on AHSS discuss TWBs with respect to various materials to provide an understanding of such material effects.

7.10

Conclusions

In this chapter, specific applications of TWBs for the automotive industry were presented focusing on classic references in the area, but also highlighting more recent work and topics. Due to the notable benefits of TWBs, e.g., reduced component weight, decreased cost, less noise, improved crashworthiness and increased dimensional accuracy, TWBs have attracted significant interest from the automotive industry. To address related formability concerns due to variations in the material properties of the weld zone, process modifications were also discussed in this chapter. Finally, typical welding processes and materials used were presented. While this chapter mentioned several key topics with respect to TWB forming, other chapters of this book provide more detailed information and should be consulted as needed.

7.11

References

Ahmetoglu, M. A., D. Brouwers, L. Shulkin, L. Taupin, G. L. Kinzel et al., 1995, ‘Deep Drawing of Round Cups from Tailor-welded Blanks’, Journal of Material Processing Technology 53: 684–94. American Machinist, 1992, ‘More Parts from Welded Blanks’, May: 19–21. Auto/Steel Partnership, 1995, Tailor Welded Blank Design and Manufacturing Manual, A/SP Technical Report. Baron, J.S., 1997, ‘A Cost Comparison of Weld Technologies for Tailor Welded Blanks’, Welding Journal 76: 39–45. Bhatt, K.K., M. Eisenmenger and M.F. Shi, 1995, ‘Formability of Mash Seam Welded Blanks: Effects of Welding Set-Up Conditions’, SAE Technical Paper Series Paper No. 950923: 183–9. Brooke, L. and H. Evans, 2009, ‘Lighten Up!’, Automotive Engineering International, March: 16–26. Chen, C.M. and R. Kovacevic, 2004, ‘Joining of Al 6061 alloy to AISI 1018 steel by combined effects of fusion and solid state welding’, International Journal of Machine Tools and Manufacture, 44 (11): 1205–14. Choi, Y.M., Y.M Heo, H.Y. Kimand and D.G. Seo, 2000, ‘Investigations of Weld-line Movements for the Deep Drawing process of Tailor Welded Blanks’, Journal of Materials Processing Technology, 108 (1): 1–7. Das, S., 2000, ‘Aluminum Tailor Welded Blanks’, Advanced Materials and Processing, 157 (3): 41–2. Davies, R.W., J.S. Vetrano, M.T. Smith and S.G. Pitman, 2002, ‘Mechanical Properties of Aluminum Tailor Welded Blanks at Superplastic Temperatures’, Journal of Materials Processing Technology, 128 (1–3): 38–47.



179

DeCicco, J. M., 1995, ‘Projected Fuel Savings and Emissions Reductions from LightVehicle Fuel Economy Standards’, Transportation Research – Part A: Policy and Practice, 29A (May): 205–28. Doege, E., H. Dohrmann and R. Kosters, 1996, ‘Simulation and Optimisation of the Forming Process of Tailored Blanks’, Proceedings of NUMISHEET ’96: 199–204. Eisenmenger, M., K.K. Bhatt and M.F. Shi, 1995, ‘Influence of Laser Welding Parameters on Formability and Robustness of Blank Manufacturing: An Application to a Body Side Frame’, SAE Technical Paper Series Paper No. 950922: 171–82. Farabi, N., D.L. Chen and Y. Zhou, 2010, ‘Fatigue properties of laser welded dual-phase steel joints’, Procedia Engineering, 2: 835–43. Florentin, J., F. Durieux, Y. Kuriyama, and T. Yamamoto, 2009, ‘A Steel Solution for a Firewall Using a Hybrid Test/CAE Approach’, SAE Technical Paper, 2009-01-1547. General Motors Research Laboratories, 1991, ‘Automotive Safety and CAFE: Proposed Changes in Standards Have “Weighty” Implications for Vehicle Occupants’, Search 26 (3). Heo, Y.M., S.H. Wang, H.Y. Kim and D.G. Seo, 2001, ‘The Effect of the Drawbead Dimensions on the Weld-line Movements in the Deep Drawing of Tailor-welded Blanks’, Journal of Materials Processing Technology, 113 (1–3): 686–91. Irving, B., 1995. ‘Welding Tailor Blanks is Hot Issue for Automakers’, Welding Journal (August): 49–52. Kapustka, N., C. Conrardy, S. Babu, and C. Albright, 2008, ‘Effect of Gmaw Process and Material Conditions on Dp 780 and Trip 780 Welds,’ Welding Journal, 87: 135s–148s. Kinsey, B.L. and J. Cao, 2001, ‘Enhancement of Sheet Metal Formability via Local Adaptive Controllers’, Transactions of the North American Manufacturing Research Institute of SME, Vol. XXIX: 81–88. Kinsey, B.L., N. Krishnan and J. Cao, 2004, ‘A Methodology to Reduce and Quantify Wrinkling in Tailor Welded Blank Forming’, International Journal of Materials and Product Technology, 21 (1/2/3): 154–68. Kridli, G.T., P.A. Friedman and A.M. Sherman, 2000, ‘Formability of Aluminum TailorWelded Blanks’, SAE Paper No. 2000-01-0772. Kubel, E., 1997, ‘Manufacturers Want More Tailored Blanks’, Manufacturing Engineering (Nov.): 38–45. Lee, A.P., E. Feltham and J. Van Deventer, 1996, ‘Tailor Welded Blank Technology for Automotive Applications’, Sheet Metal Stamping for Automotive Applications: SAE Technical Paper Series Paper No. 960817: 91–102. Munzen, W., 1997, ‘Stretch Controlled Forming Mechanism and Method for Forming Multiple Gauge Welded Blanks’, US Patent No. 5,600,991. Nagel, M., R. Fischer, J. Lowen and O. Straube, 1997, ‘Production and Application of Aluminum Tailored Blanks’, Proceedings of IBEC ’97 29: 87–91. Nakagawa, N., S. Ikura and F. Natsumi, 1993, ‘Finite Element Simulation of Stamping a Laser-Welded Blank’, SAE Technical Paper Series Paper No. 930522: 189–97. Natsumi, F., K. Ikemoto, H. Sugiura, T. Yanagisawa, and K. Azuma, 1991, ‘Laser Welding Technology for Joining Different Sheet Metals for One-Piece Stamping’, JSAE Review 12 (3): 58–63. Ninforge, D. and J. Dawance, 1998, ‘Improvement of Tailored Blank Stamping by using a Control and Localisation of the Blank-holder ’, Proceedings of the 1998 International Deep Draw Research Group, Genval, Belgium: 511–9. Ogmen, M., 1995, ‘CO2 Laser Welding and Stamping of Aluminum Alloys for Tailor Blanking’, Proceedings of Sixth International Conference on Aluminum Weldments, American Welding Society, April: 325–31.


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

180 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


Padmanabhan, R., M.C. Oliveira and L.F. Menezes, 2008, ‘Deep drawing of aluminiumsteel tailor-welded blanks’, Materials and Design, 29 (1): 154–60. Pepelnjak, T., Z. Kampus, and K. Kuzman, 1998, ‘Non-conventional Methods of Deep Drawing of Tailored Blanks’, Proceedings of the 1998 International Deep Draw Research Group, Genval, Belgium: 77–88. Rooks, B., 2001, ‘Tailor Welded Blanks Bring Multiple Benefits to Car Design’, Assembly Automation, 12 (4): 323–8. Sato, Y.S., Y. Sugiura, Y. Shoji, S.C. Park, H. Kokawa et al. 2004, ‘Post-Weld Formability of Friction Stir Welded Al Alloy 5052’, Materials Science and Engineering A, 369 (1–2): 138–43. Saunders, F.I. and R.H. Wagoner, 1996, ‘Forming of Tailor-Welded Blanks’, Metallurgical and Materials Transactions, 27A (Sept.): 2605–16. Scriven, P.J., J.A. Brandon and N.T. Williams, 1996, ‘Influence of Weld Orientation on Forming Limit Diagram of Similar/Dissimilar Thickness Laser Welded Joints’, Ironmaking and Steelmaking, 23 (2): 177–82. Sharma, R.S. and P. Molian, 2009, ‘Yb:YAG laser welding of TRIP780 steel with dual phase and mild steels for use in tailor welded blanks’, Materials and Design, 30: 4146–55. Shigematsu, I., Y-J. Kwon and N. Saito, 2009, ‘Dissimilar friction stir welding for tailorwelded blanks of aluminum and magnesium alloys’, Materials Transactions, 50 (1): 197–203. Siegert, K. and E. Knabe, 1995, ‘Fundamental Research and Draw Die Concepts for Deep Drawing of Tailored Blanks’, SAE Transactions: Journal of Materials and Manufacturing, 104: 866–76. Smock, D., 2009, ‘Tailor-Welded Blanks Cut Noise, Weight in Autos’, Design News, June. Stasik, M.C. and R.H. Wagoner, 1996, ‘Forming of Tailor-Welded Aluminum Blanks’, Aluminum and Magnesium for Automotive Applications, The Minerals, Metals and Materials Society: 69–83. Trem, R., 2004, ‘The Future of Automaking: Tailor-Welded Blanks’, Welding Magazine, http://weldingdesign.com/consumables/news/wdf_10713/ Trogolo, J.M. and J.R. Dieffenbach, 1998, ‘Evaluation of Tailor Welded Blanks Through Technical Cost Modeling’, SAE Technical Paper Series, Paper No. 980446. U.S. Department of Transportation – National Highway Traffic Safety Administration, 2009, Summary of Fuel Economy Performance. van der Hoeven, J. and K.M. Ostyn, 1994, ‘Stamping of Tailor Made Blanks’, Proceedings IBEC ’94, 7 (Sept.): 52–9. Vanker, R., 2010, personal correspondence, TWB Company, Nov. 17. Venkat, B.S., C.E. Albright, S. Ramasamy and J.P. Hurley, 1997, ‘CO2 Laser Beam Welding of Aluminum 5754-O and 6111-T4 Alloys’, Welding Journal, 76 (7): 275–82s. Viswanathan, V., B. Kinsey and J. Cao, 2001, ‘Forming of Aluminum Tailor Welded Blanks’, 2001 SAE International Congress and Exposition, Paper No. 2001-01-0822., Detroit, USA. Vollertsen, F., M. Schultz and M. Geiger, 1996, ‘Formability of Tailored Blanks from Steel and Aluminum Alloys’, Advanced Sheet Metal Forming: Proceedings of 19th IDDRG Biennial Congress: 337–46. Wang, B.Y., M.F. Shi, H. Sadrnia and F. Lin, 1995, ‘Structural Performance of Tailor Welded Sheet Steels’, SAE Technical Paper Series, Paper No. 950376. Wu, N.Q., C. Xia, M. Li, N. Perrusquia and S.X. Mao, 2004, ‘Interfacial Structure and Micro- and Nano-mechanical Behavior of Laser-welded 6061 Aluminum Alloy’, Journal of Engineering Materials and Technology, 126 (1): 8–13.


8 Tailor made blanks for the aerospace industry J. SINKE, Delft University of Technology, the Netherlands, A. A. ZADPOOR, Materials Innovation Institute (M2i) and Delft University of Technology, the Netherlands and R. BENEDICTUS, Delft University of Technology, the Netherlands

Abstract: This chapter presents potential applications of tailor made blanks (TMBs) created from adhesive bonding, (chemical) milling and welding for the aircraft industry. The combination of the TMB concept with typical forming processes like bending and rubber forming offer a high potential for weight and cost savings in metallic aircraft structures. Although thus far the applications are very limited, typical aircraft parts such as wing ribs may offer benefits in cost and weight savings of the order of 50% and 15% respectively, as proven by a demonstrator project. Further development of the concept may even increase the weight reduction, in particular when specific guidelines are used like separating strength and stiffness in the part design. Key words: tailor made blanks, adhesive bonding, welding, (chemical) machining, aircraft industry, bending, rubber forming.

8.1

Introduction

This chapter is focused on the potential use of tailor made blanks (TMBs) in the aircraft industry. Thus far, the use of TMBs is very limited, partly due to lack of research (Zadpoor, 2010) and partly because the main focus of the aircraft industry is currently on composites.

8.1.1 The aircraft industry and its manufacturing principles Market and stakeholders The civil aircraft manufacturing industry consists of a limited number of companies. Airbus and Boeing sell large commercial planes, while Bombardier, Embraer and a few smaller companies or consortia, such as ATR, build regional aircraft. The handful of Russian manufacturers mainly produces aircraft for the domestic market, though a few more companies will emerge in Russia and China in the near future. An estimate of the market size is set out in Table 8.1, which was presented in Airbus Industrie’s Global Market Forecast for 2009–2028 (Airbus Industrie, 2009). From this table, it can be deduced that about 20 000 aircraft were used by the airlines in 2008, and that the figure will reach 36 000 by the year 2028. 181 © Woodhead Publishing Limited, 2011

1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

182 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


Table 8.1 Market analysis of Airbus for the period 2009–2028 Fleet 2008

New aircraft deliveries 2009– 2018

50-seat 5444 869 70/85-seat 1305 1881 100-seat 1553 1407 125/210-seat 9254 7422 Small twin 2261 2150 aisle Intermediate 924 667 twin aisle VLA 24 496 Total 20 765 14 892

New aircraft deliveries 2019– 2028

New Remaining Fleet aircraft in service 2028 deliveries 2009– Recycled With same 2028 operator

1599 1729 836 7312 1947

2468 3610 2243 14 734 4097

1613 274 70 2677 273

58 169 118 636 84

4139 4053 2431 18 047 4454

1038

1705

114

42

1861

822 15 283

1318 30 175

0 5021

0 1107

1318 36 303

Note: Other companies like Boeing provide similar market forecasts for 10–20 year periods, which are generally identical, but may differ in the details. Source: Airbus Industrie’s Global Market Forecast 2009–2028 (http://www.airbus.com).

This means that about 30 000 new aircraft need to be built in a 20-year period in order to replace old aircraft and meet future market growth. Besides the companies already mentioned, many others supply a wide variety of products, parts and services to the Original Equipment Manufacturers (OEM), and therefore also depend on the aircraft market. Examples are the manufacturers of aero engines like Pratt and Whitney (P&W), Rolls Royce (RR) and General Electric (GE), and manufacturers of undercarriages, systems, interiors, etc. In addition to the shareholders of these companies, other stakeholders in the aircraft industry include hundreds of thousands of workers and suppliers. Moreover, millions of customers use aircraft for business trips or leisure, and society as a whole benefits from the technological achievements of the industry. Trends and drivers Current trends in the aircraft manufacturing industry include shifts towards new materials, developments in propulsion systems and improvements in assembly. These trends are driven by the everlasting objectives of weight reduction and cost reduction. The first trend relates to the introduction of new materials. The use of composite materials in aircraft increased steadily until the end of the twentieth century, though most applications were not in primary structures. However, this changed at the beginning of the twenty-first century. In 2005, the large Airbus A380 made



183

its first flight. A significant number of structural parts of this aircraft are made of composite materials (Rackers, 2006): full composites using both thermoset and thermoplastic polymers, and metal/composite materials named GLARE (Vlot et al., 2001). In 2009, another large aircraft (the Boeing 787) with a composite wing and composite fuselage, made its maiden flight. In 2013 Airbus will launch a similar aircraft: the Airbus-350 XWB. In the meantime, developments in lightweight alloys are still being made. The inclusion of scandium or lithium improved the properties of aluminium alloys, and therefore also improved their competitiveness. In addition, new production processes for aluminium alloys, like high speed machining, laser beam welding and friction stir welding, reduced the production costs of metal structures. The development of more efficient propulsion systems is another trend. In recent decades aircraft engines have become increasingly fuel efficient as engine noise and pollution have been reduced significantly. A further decrease in fuel consumption, noise and emissions is planned for the coming decades and, at the same time, a search for alternative fuels is at hand. A third trend is towards new developments in manufacturing. One such development is related to the transition from metal to multi-material structures, which involves a number of new processes. New organizational principles are also being introduced, such as those related to lean manufacturing (Murman et al., 2002). Manufacturing principles Aircraft structures are assembled from many parts (order of 1000 to 10 000 – fasteners not included), which are made from various materials like composites, metal alloys and hybrid materials. A wide range of different production processes are used in order to manufacture these parts. Since TMBs are metallic, further discussions are focused on the manufacturing principles for metals. Metal structures are assembled from sheet metal parts that have been cut and formed into the desired shape. The forming processes are often universal; much of the applied tooling is not product related. These tools are applied to a variety of different products and are very suitable for limited product series. The universal processes perfectly match the needs of the aircraft industry, where the diversity in parts is huge and the production quantities are low (in the order of 1000). Aircraft structures are also characterized by variation in materials and thicknesses. Each area is dimensioned by several load cases, which result in different materials and material conditions, and a specific distribution of thicknesses over the entire structure. Changes from one material to another, or from one thickness to another, require separate parts that have to be either joined or made as integral parts. The first option is often used for metal structures, which results in joints that add weight, time and costs during assembly. The option of integral parts, often applied to composites, has advantages with respect to weight, assembly time and costs, but is more expensive and often more costly to maintain.


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

184 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


8.1.2 Tailor made blanks and their relevance for the aircraft industry The concept of TMBs fits very well with the design and manufacturing principles of the aircraft industry. Because a low operation empty weight (OEW) is required, the structure should be optimized locally for the right material and thickness. On the other hand, there is a need to reduce the number of joints and thereby the weight and assembly costs. When two or more materials and/or thicknesses can be combined in one blank, the TMB concept fulfils both requirements. A second favorable aspect is the particular (universal) processes used. In the automotive industry, the press forming processes use matching dies. In case of TMBs with thickness steps, the thickness differences should be incorporated into the dies as well. This is not necessary for processes like rubber forming, where the soft rubber tool adapts easily to the thickness variations. Taken together, these aspects suggest there is a new opportunity for sheet forming parts in the aircraft industry. In section 8.2.4, a demonstrator is discussed and the advantages will be quantified and explained in more detail.

8.1.3 Chapter outline In this chapter, the following subjects will be dealt with: In section 8.2, the potential of the TMB concept for the aircraft industry is discussed in more detail. After a brief description of current production processes, the potential of the TMB principle for these manufacturing processes is discussed. The advantages and disadvantages of the TMB concept are briefly mentioned in this section. Sections 8.2.1 to 8.2.3 focus on the different variants of the TMB concept that can be used in the aircraft industry: machined (section 8.2.1), adhesively bonded (section 8.2.2) and welded (section 8.2.3) TMBs respectively. The basic ideas and the potential for the aircraft industry are discussed for each type. In section 8.2.4, a case study is presented. In this study, a machined wing rib is replaced by a TMB wing rib, which is made by rubber forming. For this replacement, the rib had to be redesigned and the technology change had a number of consequences for the manufacturing sequence as well as for the features of the rib itself. The last sections mention a few future trends that might be seen for TMBs in the aircraft industry section 8.3), give the conclusions (section 8.4) and conclude with references (section 8.6).

8.2

The tailor made blank (TMB) concept and the aircraft industry

The potential of the TMB for the aircraft industry can be summarized by the following keywords: reducing weight and cost by using different materials, variations in thickness and universal processes. In this section, the potential is



185

further detailed by a brief discussion of the typical aircraft parts, the way they are manufactured, and a description of some universal processes. The main substructures of an aircraft are the fuselage, the wing and the horizontal and vertical tail planes. These structures are designed using the same general concept: a load carrying skin supported by a back-up structure that consists of a grid of stiffening elements. The main elements of a metallic fuselage are the skin, frames and stringers (see Fig. 8.1). The skin is a mildly curved sheet, either single or double curved, usually made by stretch forming. After stretch forming, the skin is often etched chemically in order to remove material and to reduce the weight of the skin. The frames have a specified radius and curved flanges and are made either by rubber forming or by a roll forming process. When additional details are required, like stiffening elements or lightening holes, rubber forming is preferred. Roll forming is also used for the stringers if the quantity is large enough. When the quantity is small, or when the diversity in required stringers is large, they are made by press brake bending. The wing and tail planes have similar structures (see Fig. 8.2): the main parts of the wing (and tail sections) are the skin, the wing ribs, the stringers and a few spars. The skin of the wing is roll or stretch formed, the wing ribs are often rubber formed or machined from solid material, and the stringers are made by roll forming or bending processes. Often, the spar is a sub-assembly, assembled from plates and profiles. The large and coherent structures like the fuselage and wings are assembled from the structural elements previously mentioned, and the applied joints are often mechanical joints (either riveted or bolted). Adhesive bonding may be used as an alternative joining method, but unlike in the automotive industry, the application of welding as a joining method is still rare. The main reason is that welding disrupts the carefully constructed microstructure of high-strength

8.1 The grid of the back-up structure in the fuselage of a metallic aircraft.


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

186 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


8.2 Typical structural layout of a wing.

aluminium alloys, and the mechanical properties are reduced significantly, the tension and fatigue properties in particular. In Fig. 8.3, six variants of the TMB concept that could be used in aircraft structures are presented. These variants are based on different joining methods, sheet thicknesses and types of material/alloy. The variants consist of a TMB made by machining (variant 6), two TMBs made by adhesive bonding (variants 2 and 5) and three TMBs made by welding (variants 1, 3 and 4). The potential use of the welded TMB is explained by the growing interest in welding techniques, particularly in friction stir welding (FSW) (Mishra et al., 2005) and laser beam welding (LBW) (Zhao et al., 1999). The LBW process is already used in commercial aircraft: in the Airbus A318 and A380 the lower fuselage skin panels are made by LBW welding the stringers on top of a skin. The FSW process seems to have an even higher potential, because it is a solid state process, and because the material properties after welding do not deteriorate as much as in fusion welding processes. Figure 8.4 demonstrates

8.3 Different variants of TMB. The different gray colours indicate different alloys.



187

8.4 The principle of the friction stir welding (FSW) process.

the principle of FSW: the shoulder of a rotating tool creates friction with the metal sheets and the probe or pin stirs the hot, soft material to create a FS weld. One disadvantage of FSW is the high forces needed to introduce sufficient friction during the welding process, thus requiring heavy and expensive machinery. Finally, a few comments are made about universal processes as used in the aircraft industry. Typical examples of these universal processes are:

• • •

Rubber forming processes (see Fig. 8.5), which are press forming processes operating with a rigid, product related tool and a universal soft tool. Bending, where the bending tools can be used for different products. Stretch forming, where the stretch forming machine can manufacture multiple different shells, each shell having only one product related tool.

In rubber forming (ASM, 2006), the upper die consists of a metal casing/ container with a rubber pad inside. The rigid product related tool is placed on the lower bed of the press. During rubber forming the rubber pad forms the blank over or in the rigid die. After release of the press force the rubber springs back to its original shape. The metal blank, however, is plastically deformed and shows very little (elastic) spring back. This type of rubber forming, as shown in the figure, is sometimes called rubber pad forming. Another type of rubber forming process uses a fluid cell covered with a rubber membrane instead of a rubber pad. The rubber forming process is ideal for TMBs because the soft rubber pad adapts itself to any variation in material and/or thickness. For other press forming processes there is a need to adjust the (rigid) tools; for example, the thickness steps in the TMB. Rubber forming is typically used for rather flat and shallow parts, or ‘flanged parts’, which have a flat web plate and curved flanges at the periphery (see Fig. 8.6). In the web plate and the flanges, specific details might be present. The wing rib described in section 8.2.4 is a typical ‘flanged part’. For this product


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

188 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


8.5 Schematic of rubber forming using a rubber pad.

the most dominant deformation is bending, although locally (in flanges, for example), some in-plane straining also takes place. The press brake bending process (Lange, 1985) is another universal process often used for the manufacture of sheet metal parts. In this case, the upper and lower dies have a radius and a v-shaped cavity respectively, and are both universal. Depending on the sheet thickness and type of material, the right combination of dies is selected for the bending operation, which is often controlled by computer. During bending the deformations are concentrated in the bend zone, a narrow line in the product, with no deformation taking place outside that zone.

8.6 A wing box with several wing ribs, typical examples of flanged parts (including curved flanges and collared holes and stiffening beads).



189

Bending of a TMB made from different materials (same thickness) is limited by the material with the worst formability properties. That material will define the bend radius. In case of a TMB with different thicknesses, the combination with the worst formability defines the minimum bend ratio (r/t). The spring back in these bending operations is different at both sides of the transition line, so compensation is needed when a constant product angle is required. Simple tool adaptations may provide the necessary equipment for the right product geometry. Stretch forming (ASM, 2006) (see Fig. 8.7) can also be regarded as a universal process. However, in this process the in-plane deformations are much larger than in bending and rubber forming. Therefore, the process will be much more sensitive to the positions and orientations of the transition lines (weld lines or thickness steps in the undeformed blank). The strain distribution in a stretch formed sheet is not uniform. The largest strains are concentrated on the top of the stretch forming shape. Therefore, the transition line must run parallel to the direction of stretching (parallel to the center line in the top view). But, because the major strain is mainly in one direction (in the direction of stretching), the potential application is less than for the other two processes. Another problem in stretch forming of TMBs with thickness steps is that usually the sheet is stretched over a concave die, so the outside of the skin should be smooth. The inside of the skin, with the thickness steps, will be in contact with the stretch forming die. In that case, the tool has to be adapted locally and this adds complications.

8.2.1 Machined tailor made blanks The machined TMB suits the needs of the aerospace industry very well. For many parts, weight is reduced by removing material, like in the machining of lightening

8.7 Sketch of the stretch forming process of sheet (side, front and top view).


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

190 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


holes in sheets, milling pockets in thick sections and creating pockets in sheets by chemical milling. The current state of the art process for rubber forming of parts like wing ribs involves milling lightening holes before the forming operation. The blank is made in a net shape and after forming no additional trimming operations are necessary. Local thinning of the part by (chemical) milling is not an option yet for rubber forming parts. The proposed concept of machined TMBs adds this possibility to the existing methods of weight reduction. The machined TMB offers only a variation in thickness; the material and its condition are the same for the entire TMB. Pocketing or local thinning is performed only on skins for the fuselage and wings of aircraft. The most common procedure is that the sheets undergo stretch forming followed by chemical milling. Stretch forming is preferred because a significant tensile force is combined with the bending moments. Due to these tensile loads the entire cross section of the sheet is in tension and the residual stresses after unloading are minimized. As a result, the subsequent etching or chemical milling process, aimed at the local removal of some thickness, has no effect on the shape of the skin. Machined TMBs are made by removing a thin layer of material over a large area. Both chemical milling and conventional milling can be used to remove the material. In chemical milling, the layer is removed by submerging the material in a chemical bath. Those areas that need to be etched are exposed to the chemical agent; the remainder of the surface is covered by a polymer that prohibits chemical etching. The length of time the sheet is exposed to the chemical determines the thickness of the removed layer. The other method, conventional milling, removes a thin layer of the surface e.g. by slab milling. During milling special care is required over internal stresses in the sheet and flapping. When, upon removal of a surface layer, internal stresses influence the shape of the product, pre-stretching of the sheet is applied prior to milling in order to eliminate the internal stresses. The formability of the machined TMB is not as good as the formability of the parent sheets (Zadpoor et al., 2008a). Due to the milling, thickness steps are created in the blank, which cause stress concentrations during the forming process. The impact of these stress concentrations is maximized when the major stress and strain are perpendicular to the thickness step or transition line (see orientation B in Fig. 8.8). The effect is minimal when the major stress and strain are parallel to the transition line (orientation A). In the worst case (orientation B), the stress concentration will reduce the ultimate strain in the thin sections, although this reduction depends on the thickness ratio r (= t1/t2). At the same time, the strain in the thicker section remains well below the failure limit, and the formability potential of that part of the blank is not used. Although the transition line has a significant impact on the failure strain (B orientation), the failure strength of the part or specimen remains nearly constant and is hardly affected by the thickness ratio r.



191

8.8 Sketch of the transition line in a machined TMB and the orientation of the major strain (A or B).

Surface roughness can also have an impact on the formability of machined TMBs. During the milling process traces are often created on the surface, or some roughness is present after chemical milling, which may induce a local strain increase (strain localization) and thereby a reduction in formability (Marciniak and Kuczynski, 1967). The two effects, stress concentration at the transition line and strain localization at surface imperfections, are sometimes in competition. The thinner the machined side of the TMB, the more severe the surface imperfections become, and thereby the strain concentrations, which may result in failures away from the transition line (Zadpoor et al., 2008a). Another aspect to be mentioned here is that during loading of the TMB material perpendicular to its transition line, a shift in the neutral axis takes place, which may introduce some bending at the transition line if this is not suppressed by the selected forming process (as in rubber forming). Machined TMBs can also be heat treated. The part that has been selected for the demonstrator was heat treated by a solution treatment followed by quenching. This treatment improved the limiting failure strain of the aluminium alloy by a factor of 1.5 to 2. Bending tests on machined TMBs showed that the minimum bend ratios correspond to the minimum ratios of the parent materials (see Fig. 8.9). Only when the sheet has been machined to create large thickness ratios (r > 1.5), and stretching of the sheet was required before machining, is some increase in the minimum bend ratio reported: an increase of about 10–12% (Zadpoor et al., 2009a). Finally, the different thicknesses in the blank also result in different levels of spring back. The thinner section, when bent over the same radius, will result in more spring back. When a constant product angle is required, the radius and bend angle should be defined for both materials and thicknesses, and will result in different bend radius/bend angle combinations.


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

192 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


8.9 Minimum bending radii calculated using FEM and experimental values for machined tailor made blanks (based on 2024-T3, t = 2.5 mm) with different thickness ratios.

8.2.2 Adhesively bonded tailor made blanks Another type is the adhesively bonded TMB. Adhesive bonding is a well-known process in the aircraft industry. Bonding has a number of advantages when compared to other joining processes: the ability to join dissimilar materials, the limitation of stress concentrations, and the high structural efficiency. On the other hand, the elaborate process of surface treatment and bonding, the limited temperature range for bonded TMBs and the inability to disconnect bonded joints, are disadvantages. Thickness steps in metallic parts can be created by either removing or adding material. Adhesive bonding is a process that creates thickness variations in parts and structures by adding sheets to a base sheet. In these TMBs, the layers may consist of dissimilar metal alloys. Typical examples in aircraft structures are so-called waffle plates; pocketed sheets that provide additional thickness underneath stringers, frames or ribs. Local reinforcements, named doublers, also strengthen the sheet in areas where the loads increase. An example that is currently applied is given in Fig. 8.10: the window belt area of the fuselage is locally reinforced with doublers. Adhesively bonded TMBs are aimed at extending this idea to smaller substructures or even individual components. The forming limits of adhesively bonded TMBs depend also on the orientation of the transition lines (run-out of layers) with respect to the main loading and deformation directions (Zadpoor et al., 2009b). When the transition line is perpendicular to the major strain, the end of the bond line is loaded with shear and peel forces (see Fig. 8.11). Whether the joint fails depends on the type of



193

8.10 Sketch of the window belt area with the skin (1), a doubler (2), a waffle plate (3) and a number of window cut-outs (4).

adhesive and the design of the overlap, but regardless of this it limits the formability of the TMB. In particular, the peel stresses may induce early failure of the bond line. Although the assembly does not fail in a catastrophic manner, it should be regarded as failed and be replaced. The main cause of failure, the peel stress, is dependent on a number of parameters like the type of adhesive, its thickness and the joint geometry. By tuning these parameters, the effect of peeling can be reduced. When the major strain is applied parallel to the transition line, the forming limit is determined by the properties of the constituents, particularly the failure strain. Often the adhesive has the smallest failure strain and is therefore responsible for failure. Nevertheless, one could select an adhesive with sufficient failure strain and thereby increase the formability of the bonded TMB in case of loading parallel to the transition line. Other phenomena, like secondary bending and spring back, are similar to those mentioned for machined TMBs. Unlike the other two types of TMB, the machined

8.11 Peel and shear stress distribution at the end of a bonded layer.


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

194 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


and FSW TMB, the bonded TMB cannot be heat treated. The temperatures for these treatments are well beyond the maximums for the adhesives.

8.2.3 Friction stir welded tailor made blanks The third and last type mentioned here is the FSW TMB. FSW is a solid state welding process that offers the potential to weld different alloys and a limited reduction in mechanical properties like yield and ultimate stress after welding. In addition, a heat treatment can be used to improve the formability before the forming process and to restore part of the original properties of the parent materials after forming (Totten and MacKenzie, 2003). The FSW process is still under development and needs to be certified before it can be applied to aircraft structures. Until recently, welding was not used in aircraft structures, because the aluminium alloys had poor welding properties. Since the introduction of some new alloys, like the Al-6013 alloy, welding is applied in some parts of aircraft structures that are not dominated by fatigue loadings: e.g. the lower skin panels of the Airbus 318 and 380 fuselages are made of stringers welded onto the skin. In this case however, LBW is used; a fusion welding process that results in larger reductions of the mechanical properties. For the FSW process, smaller reductions are predicted. Unless the thickness is varied too, the forming limits in FSW blanks are dominated by the weld lines. The weld line itself consists of several areas: the weld nugget (WN), the thermo-mechanical affected zones (TMAZ), the heataffected zone (HAZ) and the parent or base material (Zadpoor et al., 2008b). These lines, which act as transition lines, reduce the formability in particular for the orientation perpendicular to the weld line. In such tests the strain to failure at the weld decreases significantly (in the order of 50%) when compared to the parent materials. This reduction in failure strain can be counteracted by a heat treatment, which eliminates some of the unfavorable microstructures. Preliminary tests have proven that the failure strain can be restored following heat treatment of the aluminium alloys Al-7075 and Al-2024 (Zadpoor and Sinke, 2010). The heat treatment used in this case was solution treatment at 495°C followed by quenching. The test specimen had been tested within half an hour after quenching. A test specimen with a weld line parallel to the loading direction also initiates cracks at rather low values of strain, although these failures are not catastrophic for the entire specimen. Nevertheless, in those cases a heat treatment can still improve the formability significantly. An FSW TMB requires similar design rules as the machined and adhesively bonded TMB (i.e., the orientation of the transition lines with respect to the major strain direction is important, the spring back should be related to the parent materials and that there is secondary bending when the parent materials have an unequal thickness).



195

In addition, the FSW TMB offers the combination of different alloys side by side in one blank and the possibility to improve the formability limits by heat treatments.

8.2.4 A rubber formed TMB wing rib: a case study To demonstrate and quantify the potential of the TMB for the aircraft industry, a case study is presented in which several aspects of the TMB concept are highlighted. This study was performed in cooperation with Stork Fokker Aerospace in the Netherlands. Forming trials have been performed at their production plant at Papendrecht. The main objective of this demonstration was to replace a current product and to investigate the advantages and disadvantages of the TMB concept, in particular with respect to the production process. In the redesign, the aim was to create a TMB part that was easy to manufacture using existing processes. The part selected for this study was a rib in the horizontal stabilizer (horizontal tail wing) of a business jet – the original and redesigned wing rib are presented in Fig. 8.12. The wing rib was selected from four other structural parts because, in a

8.12 Schematic figure of the original wing rib (top) and a drawing of the redesigned TMB wing rib (bottom).


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

196 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


preliminary evaluation of the parts, the rib showed the highest potential for reductions in manufacturing costs and time. The original wing rib was machined from a 25 mm thick Al 7075-T351 plate. During the milling operation about 90–95% of the material is removed. The TMB wing rib, made from a 2.8 mm 2024-T42 sheet, is a redesign based on the features and dimensions of the original product. The change in material is related to the change in the manufacturing process: machining is a good choice for the manufacture of a part of a 7000-series alloy, but forming is a better choice for a 2000-series alloy. The wing rib itself is not highly loaded or fatigued; its primary properties should be good stiffness with regard to compressive and shear forces. As stated, the geometrical features of the original were redesigned to make the rib by a sheet forming process. For instance, the blade stiffeners of the machined example were replaced by formed stiffening beads and the vertical edges at the periphery of the holes were replaced by formed collared holes with angles of 45°. The 2.8 mm thick sheet was milled to a TMB by reducing the thickness in some areas to 2.6 and 1.6 mm. During the rubber forming process, the TMB is formed over a positive or male die, with the thickness steps at the side of the soft tool. The face of the sheet without steps makes contact with the forming die, which simplifies the manufacture of the tool (i.e., no steps in the tool surface to accommodate for the thickness differences). On a few occasions, a special feature was required, like a ‘reversed joggle’, as explained below. The transition lines are not interfering with the stiffening beads and the collared holes. The flanges at the periphery of the web plate have different thicknesses. Since the outside of the wing rib should have smooth contour lines, the thickness steps at the flanges (the transition line is perpendicular to the bend line) should be compensated for by joggles. In this case, ‘reverse joggles’ are needed (see Fig. 13). As stated before, the focus of the demonstrator was on manufacturability. The original wing rib was milled from a thick plate (25 mm). The machining was the only manufacturing process involved except for a number of finishing processes such as quality control, surface treatments and painting. The manufacturing process of the TMB wing rib had a number of processing steps. First, the sheet was milled to the required dimension of the blank with the contours made by edge milling and the thickness steps created by slab milling. Subsequently, the sheet was heat treated to improve the formability of the as-received material (in T3-condition). The standard heat treatment used was a solution heat treatment (495°C for 30 minutes) followed by quenching. The next step was the forming process. The sheet was rubber formed over a positive die and all features were made in one process cycle, including the overall shape, and all local details like the stiffening beads reversed joggles and collared holes. After forming, the sheet was heat treated again (ageing) to obtain the right mechanical properties, and finally the sheet was finished using inspection, surface treatments and painting.



197

8.13 (a) Location and (b) a conventional joggle (A–A) and (c) a ‘reversed joggle’ (B–B) as used for TMBs.

Since the demonstrator was used to investigate the potential of the TMB concept, a time and cost analysis was performed based on the experience of the Stork Fokker Company. The results of this study are presented in Table 8.2. This compares the original manufacturing processes for the wing rib (i.e., the machining from plate) to the alternative processes (i.e., the forming from a TMB sheet). In the table the total values for the original product are used as reference. It is clear that a decrease in time and costs of 24% and 56% respectively was achieved.


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

198 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X


Table 8.2 Comparison between the time and costs of the original and TMB processes Original process Item or activity Material Processing plate Handling sheet Milling sheet Heat treatment (ST/Q) Rubber forming Heat treatment (ageing) Q-control, anodizing Painting Total

Time (%)

85

TMB process Costs (%)

Time (%)

60 34

Costs (%) 13

7 8

4 2

30 12 15 10 2 2 5

100

100

76

13 4 7 3 1 1 2 44

Although the number of manufacturing steps for the TMB rib is higher than for the machined rib, the costs are more or less equal. The biggest difference in cost between the two concepts is due to the material: the metal plate (25 mm thickness) is about four to five times more expensive than the sheet. The milling of the wing rib is a rather time consuming process, which accounts for the time difference. The difference in quality control and painting is caused by the fact that the machined rib is more susceptible to cracks and therefore inspection and surface treatment should be more thorough. The weight of the two wing ribs was also compared. The TMB rib made of sheet metal was 15% lighter than the original one, despite the fact that the sheet metal rib was not optimized for weight. During the study, this wing rib proved the feasibility with respect to manufacturing costs and manufacturing time. The next move is to optimize the wing rib for weight: further reducing the weight while maintaining the same level of performance.

8.3

Future trends

The concept of TMBs has great potential for the manufacture of sheet metal parts in the aircraft industry, although thus far (in 2010), the number of applications is limited. As described in this chapter, the match between TMBs and a universal process like rubber forming is ideal since the TMB part can be made without a significant increase in (tooling) costs. A second advantage is that the application of the right metal alloy and thickness can be optimized much further than before, because a TMB consists of more than one alloy and/or one thickness. It might be feasible to combine alloys and metals in such a way that functions are separated in a part, e.g. a part in which strength and corrosion resistance are the dominant design requirements for different sections in the product. In such a case, a TMB



199

8.14 Schematic picture of a rib-like sheet metal part with stiff (gray) and strong (white) areas.

could provide a suitable solution: a strong sheet joined to a corrosion-resistant sheet. As explained in this chapter, such a blank can be manufactured. Furthermore, the TMB concept can be exploited when during the analysis of a part the strength and stiffness of that part can be (partially) separated. To illustrate this, a wing rib has been used in the case study. In Fig. 8.14, a schematic of the wing rib is presented with a number of distinctive features: flanges, collared holes, stiffening beads and web plate. In this figure the gray areas (i.e., the collared holes, flanges and the beads) are providing stiffness to the rib with respect to shear, bending and compression. The stiffness in these elements is primarily provided by their shape. The thickness and the elastic modulus of the material are of secondary importance. When the stiffness needs to be increased, the best way is to adapt the shape of (one of) these elements. The flat areas in between should not buckle by compression or shear, nor fail. The buckling resistance of these areas is increased tremendously by reducing their size. This example demonstrates the fact that for further exploitation of the TMB concept, new design principles should be developed. However, design and development of optimized TMBs is more complicated than for monolithic parts, because the number of variables involved increases. In a TMB all three items, the design, the choice of materials and the applicable manufacturing processes, should be joined in such a way that the resulting structure is cost effective and lightweight. The best result is obtained when these three entities are taken into account in an iterative optimization process.

8.4

Conclusions

In this last section, this chapter is summarized in a few statements:

•

The concept of TMBs is ideal for the aircraft industry, because this industry has suitable manufacturing processes, uses different mixes of materials and the reduction of weight is a very important basic rule.


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

200 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

• • • •


TMBs offer the opportunity to reduce the number of parts and joints, which further reduces the weight. Several manufacturing processes of the aircraft industry are universal, applying only one product related tool. These processes are also favorable for the application of the TMB concept. The best results will be obtained when the design of the new TMB is optimized by taking into account the materials, the design and the manufacturing process in an iterative manner. Another important and potential idea is the tailoring of the sections in a product to the specific tasks of those sections (i.e., aiming for stiff sections, strong sections and fatigue resistant sections, impact resistant sections, etc).

8.5

Acknowledgements

The text of this chapter is mainly based on research that was carried out under project number MC1.05224 of the Strategic Research Program of the Materials Innovation Institute M2i (http://www.m2i.nl), the former Netherlands Institute for Metals Research. The industrial partner in this research was Stork Fokker Aerospace, the Netherlands.

8.6

References

Airbus Industrie, (2009), ‘Global Market Forecast 2009–2028’, http://www.airbus.com. ASM Handbook, (2006), ‘Metal working, Sheet forming’, volume 14B, ASM International. Lange, K., (1985), Handbook of Metal Forming, McGraw-Hill, New York. Marciniak, Z. and Kuczynski, K., (1967), ‘Limit strains in the process of stretch-forming sheets’, Int. Journal of Mechanical Science, 9, 609–620. Mishra, R. S. and Ma, Z. Y., (2005), ‘Friction stir welding and processing’. Materials Science and Engineering R-Reports, 50, 1–78. Murman, E. et.al., (2002), Lean enterprise value: Insights from MIT’s Lean Aerospace Initiative, Palgrave, New York. Rackers, B., (2006), ‘Composite materials in the Airbus A380’, Congress on Plastics in Automotive Engineering, Mannheim, Germany. Totten, G.E. and Mackenzie, S., (2003), Handbook of Aluminium, Physical metallurgy and processes, Marcel Dekker, Switzerland. Vlot, A. and Gunnink, J.W., (2001), Fiber Metal Laminates, an Introduction, Kluwer Academic Publishers, Dordrecht, the Netherlands. Zadpoor, A.A., (2010), ‘Tailor-Made Blanks in the aircraft industry’, PhD thesis, TU Delft, the Netherlands. Zadpoor, A.A. and Sinke, J., (2010), ‘Effects of post-weld heat treatment on the mechanical properties of similar- and dissimilar-alloy friction stir welded blanks’, to be published. Zadpoor, A. A., Sinke, J. and Benedictus, R., (2008a), ‘Experimental and numerical study of machined aluminium tailor-made blanks’, Journal of Materials Processing Technology, 200, 288–299.



201

Zadpoor, A. A., Sinke, J. and Benedictus, R., (2008b), ‘The mechanical properties and microstructure of friction stir welded tailor-made blanks’, Materials Science and Engineering A, 494, 281–290. Zadpoor, A. A., Sinke, J. and Benedictus, R., (2009a), ‘Bendability of machined aluminum tailor-made blanks’, International Journal of Material Forming, 2, S821–S824. Zadpoor, A. A., Sinke, J. and Benedictus, R., (2009b), ‘The mechanical behaviour of adhesively bonded tailor-made blanks’, International Journal of Adhesion and Adhesives, 29, 558–571. Zhao, H., White, D. R. and Debroy, T., (1999), ‘Current issues and problems in laser welding of automotive aluminium alloys’, International Materials Reviews, 44, 238–266.


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

Index

ABAQUS/Standard, 56 adhesive bonding, 185, 192 advanced high-strength steel tailor welded blanks (AHSS-TWBs), 118–60 AHSS types and their characteristics, 120–3 complex phase steels, 122 dual phase steels, 121 ferritic–bainitic steels, 122 hot-formed steels, 123 martensitic steels, 122 transformation-induced plasticity steels, 121 twinning-induced plasticity steels, 122 diode laser-welded DP980, 142– 7 failed specimens from limit dome height test for DP780 and HSLA steels, 146 hardness profiles of DP980 and HSLA, 143 HSLA and DP780 welded microstructures, 145 tensile behaviours of welded DP980 and HSLA, 144 fabrication, 123–7 laser cutting, 123–4 laser-welded DP600, 128–31 engineering stress–strain curves, 130 fatigue, 130–1 fatigue S–N curves, 131 microhardness distribution and microstructures across the laser welded line, 128–9 microstructural change, 129 microstructure and mechanical properties, 128 tensile properties of welded coupons, 129–30 laser-welded IF–DP steels for TWBs, 131–8 crack formation, 135 engineering stress–strain plots of parent IF–DP steels and welded blanks, 134 fatigue and impact tests, 137–8 forming limit diagrams of parent steel sheets for various welded blanks tested, 136

load vs dome height plot for different types of blanks of 50 mm width, 135 microhardness distribution, 131 microhardness distribution across the weld line, 132 microstructures, 131, 133 microstructures of IF, DP590 and welded zone, 132 S–N curves of IF and DP base metals and TWB, 137 subsize tensile test specimens, 133 tensile and formability testing, 133–7 laser welding fabrication methods, 124–6 CO2 laser, 125 diode laser, 125–6 YAG laser, 125 other methods for welding AHSS, 126–7 friction stir welding, 126 gas metal arc welding, 126 welding aluminum–steel TWBs, 126–7 other related manufacturing processes, 157–9 boron steels hot stamping for auto parts requiring enhanced crash resistance, 157–8 manufactured crash structure from tailor rolled blanks, 159 tailor made blanks, 158–9 other studies on welding AHSS and applications of AHSS-TWBs, 147–9 aluminum–AHSS TWBs, 148–9 arc welding various AHSS of similar types, 147–8 friction stir welding, 149 properties and formability, 127–49 understanding microstructure evolution and its impact on properties, 149–56 comparison of DP780 and TRIP780 steels during heating and natural cooling, 152 equilibrium phase diagram of TRIP780 with 1.8w% Al, 150 history of thermal profile in welding, 153–6 microstructures, 151–3

203 © Woodhead Publishing Limited, 2011

1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

204 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

Index

microstructures of DP–IF welded zones, 154 phase diagram, 149–51 temperature profile in gas metal arc welding of IF steel, 155 temperature profiles in the two HAZs of laser-welded IF–DP steels, 156 YAG laser-welded TRIP/DP TWBs, 142 YAG laser-welded TRIP780-mild steels and TRIP780-DP980 steels, 138–42 elongations obtained in TWBs consisting of TRIP780 welded to mild steel of DP980, 142 materials and weld theme, 139 microhardness distribution, 139 microhardness values observed for TRIP780-mild steels and TRIP780-DP980, 140 tensile test samples of TRIP–10% pre-stain welded to DP980, 141 tensile tests, 139, 141–2 advanced high-strength steels (AHSS), 171 Airbus–350 XWB, 183 aircraft industry adhesively bonded tailor made blanks, 192–4 peel and shear stress distribution at the end of bonded layer, 193 sketch of the window belt area with the skin, 193 friction stir welded tailor made blanks, 194–5 future trends, 198–9 schematic of rib-like sheet metal part with stiff and strong areas, 199 machined tailor made blanks, 189–92 minimum bending radii, 192 transition line in a machined TMB and orientation of the major strain, 191 manufacturing principles, 183 market analysis for Airbus for the period 2009–2028, 182 market and stakeholders, 181–2 rubber formed TMB wing rib, 195 comparison between time and costs of original and TMB processes, 198 location and conventional joggle and a ‘reversed joggle as used for TMBs, 197 original wing rib and drawing of redesigned TMB wing rib, 195 tailor made blanks, 181–200 tailor made blanks concept, 184–97 different variants of TMB, 186 grid of the back-up structure in fuselage of metallic aircraft, 185 principle of friction stir welding process, 187 rubber forming using a rubber pad, 188 stretch forming process of sheet, 189 typical examples of flanged parts, 188 typical structural layout of a wing, 186 TMBs and their relevance, 184 trends and drivers, 182–3

Al 5083/5383, 99 Al 5182-H00, 171 Al 7075-T351, 196 allotropy, 101 aluminized coated boron steels, 158 aluminum, 97–9 Al 5083-H34 vs mild steel strengths and Al alloys strengths according to grades, 98 applications, 99 aluminum–silicon, 123 aluminum–steel TWBs, 126–7 American Society for Testing Materials, 29 analytical model, 52 ArcelorMittal, 157 automotive industry advantages of tailor welded blanks, 168–70 decreased noise, 170 improved corrosion resistance, 170 improved crashworthiness and part stiffness, 169 improved dimensional accuracy and consistency, 170 material overlap for spot, mash seam and laser welding processes, 169 reduced manufacturing cost, 168–9 reduced vehicle weight, 168 disadvantages of tailor welded blanks, 170–3 decreased formability, 170–2 other TWB concerns, 173 schematic of weld line movement, 173 weld line movement, 172 wrinkling in die addendum, 172–3 materials used to produce TWBs, 177–8 research efforts for TWBs, 174 tailor welded blanks, 164–78 body side TWB panel, 167 door inner example, 166–7 door inner TWB panel, 167 exploded view of current and or potential TWB body components, 165 flowcharts for fabrication of door inner, 166 historical perspective, 167–8 TWBs forming methods, 174–6 application of non-uniform binder force to improve TWB forming, 175 means to clamp on weld line, 176 non-uniform binder force, 174–6 sequence of forming process using lower die cushion to clamp on the weld line, 176 strategic locating of weld line on part, 174 welding processes for TWBs, 177 bainite, 121 Barlat yield criterion, 81–2 Bauschinger effect, 78 Belytschko-Tsay shell elements, 60 bending, 187 bending theory, 52


Index Bergström strain hardening model, 76 Bishop-Hill notation, 81 blank holder force, 63–5, 111 chemical milling, 190 clamping force see hydraulic pressure CO2 laser, 125 complex phase (CP) steels, 122 concavity, 7–8 constitutive behaviour, 25–6 longitudinal welded blank, 25–6 load sharing schematic, 25 conventional milling, 190 Corporate Average Fuel Economy, 167 Coulomb friction coefficient, 56 Coulumb friction, 63 cross-sectional area method, 27 cross-sectioning, 12 cup test, 11–12 deformation tailor welded blanks during forming, 24–45 constitutive behaviour estimation, 25–6 design considerations, 39–40 forming limits, 32–5 simulation, 40–3 weld line movement, 36–8 weld width evaluation method (or cross-sectional area), 26–32 digital image correlation, 69 diode laser, 125–6 DP 590, 33 dual phase (DP) steels, 131 ductile fracture theories, 84 DYNAFORM-PC, 59, 60 eddy current, 17 electromagnetic acoustic transducer, 16–17 Erichsen test, 11–12 ferritic–bainitic steels, 122 finite element method, 71–5 hardness profile, 75 simulation results, 74 simulations, 29, 40–3 forming limit curve, 32–3, 109 forming limit diagram, 70–1, 127, 135–6 free-bend test, 43 friction stir welding, 69, 101, 105, 110, 126, 186–7 other studies on welding AHSS and applications of AHSS-TWBs, 149 gas metal arc welding (GMAW), 126 gas tungsten arc welding (GTAW), 174 gaussian distribution, 19–20 GLARE, 183 Gleeble test, 151 Gurson–Tvergaard–Needleman (GTN) yield function, 82

205

heat-affected zone, 70, 171 high frequency electric resistance welding, 84 high hole expansion steels see ferritic–bainitic steels Hill’s 1948 yield criterion, 76, 80–1 Hollomon equation, 25 Hollomon’s power law, 79 Hosford yield criterion, 81 hot-formed steels, 123 hydraulic pressure, 38 in-plane test, 42 interstitial free steel, 27 J2 deformation theory, 84 ‘keyhole,’ 124 ‘kissing bond,’ 107 lack of penetration, 7, 14 laser beam welding, 186 laser cutting, 123–4 laser roll welding, 105 laser triangulation, 14–15 illustration, 15 laser welding, 104–5 lightweight metal alloy tailor welded blanks (LWMA TWBs), 97–114, 101–9 aluminum, 97–9 Al 5083-H34 vs mild steel strengths and Al alloys strengths according to grades, 98 applications, 99 benefits/recycling, 114 different defects, 106–7 intermetallics, 106 internal and external defects, 106– kissing bond in FSW aluminum, 107 straining directions in a TWB, 107 different manufacturing techniques, 104–6 friction stir welding, 105 laser welding, 104–5 resistance mash seam welding, 105–6 dissimilar material TWBs, 102–4 failure prevention, 108–9 drawbead design, 108 weld clamping method, 109 formability, 109–13 dissimilar materials TWB and formed car door panel, 110 similar and dissimilar aluminum TWB cups, 113 TWB square cup with weld line and flange deformation for different dissimilar material TWBs, 111 weld line movement and thickness variation, 112 magnesium, 99–100 applications, 100 stress–strain curve at different temperatures, 100


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

206 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

Index

metal matrix composites, 101 similar material TWBs, 102 titanium, 100–1 applications, 101 typical application of dissimilar material TWB, 103 limit dome height (LDH) test, 32, 127 limiting strength ratio, 51 limiting thickness ratio, 50–1 LS-DYNA FE software, 42 LS-DYNA3D, 64 Ludwik’s law, 79 magnesium, 99–100 applications, 100 stress–strain curve at different temperatures, 100 magnetic flux leakage detector, 17–18 manufacturability, 87 Marciniak–Kuczynski theory, 70, 83–4 martensite, 121 martensitic steels, 122 material models, 75–82 numerical studies of TWBs, 77–8 strain hardening laws, 78–9 yield criteria, 79–82 MDP4, 142 mechanics-based modelling material draw-in ratios, 54–62 3-D non-symmetric, aluminum case, 62 2-D numerical simulation results, 56–9 3-D numerical simulation results, 59–2 analytical model flowchart, 56 forming height and material movement, 55 numerical simulations and analytical model, 57 numerical simulations and experimental investigation results, 62 results for normalized force, 58 results for normalized strain, 58 results of numerical simulations and analytical model, 57 strain in 2-direction for various cases, 60 synthetic steel case specimen, 60 thin side of the 2-D cross-section showing geometric parameters, 55 tooling dimensions and material properties, 59 weld line movement and forming height vs FEA results, 61 non-uniform binder force, 63–5 constant non-uniform binder force for strip drawing experiments, 65 constant non-uniform binder force in numerical simulations, 65 normalized non-uniform blank holder force versus punch displacement, 64 tailor welded blanks, 48–66 material draw-in ratios, 54–62 thickness and strength ratio analysis, 49–51

strain values longitudinal and transverse to the weld for a TWB, 50 weld line cross section for a TWB, 49 weld line movement and forming height, 51–3 2-D cross-sectional representation of a TWB formed part, 52 metal matrix composites, 101 micro-hardness method, 27 micro-hardness test, 13 sample across a laser weld, 13 numerical simulation modeling design and optimization topics, 85–8 design of the die set, 85–6 TWB design, 86–8 finite element method, 71–5 hardness profile, 75 simulation results, 74 material models, 75–82 numerical studies of TWBs, 77–8 strain hardening laws, 78–9 yield criteria, 79–82 tailor welded blank, 68–89 future trends, 89 theoretical failure prediction, 82–5 Nuvonyx ISL4000L diode laser, 126 optical micrograph, 27 orthogonal coordinate system, 49 phenomenological models, 76–7 physical models, 76 porosity, 9–10 inclusion induced pore, 9, 10 power hardening law, 52–3 press brake bending, 185, 188 production quality control, 19–22 weld imperfections monitoring, 20–2 weld imperfections prevention, 19–20 equipment cleaning, 19 filler material, 20 laser quality, 19–20 wedge edge preparation, 20 ‘Proserio’, 99 reference volume element, 76 resistance mash seam welding, 105–6 reversible transformation, 101 roll forming, 185 rubber forming process, 187 rubber pad forming, 187 rule of mixtures, 25, 27, 29 engineering stress–strain behavior, 28 experimental and engineering stress–strain behavior, 30 mechanical properties comparison, 29 true stress–strain behavior, 28 semi-physical models, 76 see also Bergström strain hardening model


Index sensor fusion, 21 sheet metal stamping, 164 SPCC 440, 33 spot welding, 170 strain hardening laws, 78–9 stress–strain behavior, 25–6 stretch-bend tests, 43 stretch flangeable steels see ferritic–bainitic steels stretch forming, 185, 187, 188, 190 Swift’s strain hardening law, 79 tailor made blanks, 69, 158–9 aircraft industry, 181–200 future trends, 198–9 industry and its manufacturing principles, 181–3 TMBs and their relevance, 184 TMBs concept, 184–97 tailor welded blanks advanced high-strength steel, 118–60 fabrication of AHSS-TWBs, 123–7 other manufacturing processes related to AHSS-TWBs, 157–9 properties and formability of AHSS-TWBs, 127–49 types and their characteristics, 120–3 understanding microstructure evolution and its impact on AHSS-TWBs properties, 149–56 automotive industry, 164–78 advantages of tailor welded blanks, 168–70 disadvantages of tailor welded blanks, 170–3 door inner example, 166–7 historical perspective, 167–8 materials used to produce TWBs, 177–8 research efforts for TWBs, 174 TWBs forming methods, 174–6 welding processes for TWBs, 177 design, 86–8 during forming, deformation, 24–45 constitutive behavior estimation, 25–6 design considerations, 39–40 forming limits, 32–5 shim for thickness compensation, 39 simulation, 40–3 stepped binder scheme, 40 weld line movement, 36–8 weld width evaluation method, 26–32 forming limits, 32–5 experimental FLC of TWB comparison, 32 experimental FLCs comparison of TWB, DP-IF steels, 34 experimental FLCs comparison of TWB, SPCC-IF steels, 35 experimental FLCs levels of 3 base steel sheets, 34 IF, DP and SPCC steel sheets mechanical properties, 33

207

other factors, 35 steel TWBs, 33–5 weld orientation influence, 32–5 lightweight metal alloys, 97–114 aluminum, 97–9 LWMA TWB benefits/recycling, 114 LWMA TWB formability, 109–13 LWMA TWBs, 101–9 magnesium, 99–100 metal matrix composites, 101 titanium, 100–1 mechanics-based modeling, 48–66 material draw-in ratios, 54–62 non-uniform binder force, 63–5 thickness and strength ratio analysis, 49–51 weld line movement and forming height, 51–3 numerical simulation modeling, 68–89 design and optimization topics, 85–8 finite element method, 71–5 future trends, 89 material models, 75–82 theoretical failure prediction, 82–5 simulation, 40–3 finite element simulation, 41–3 weld region representation, 40, 41–3 weld integrity, 3–22 production quality control, 19–22 testing methods, 11–19 weld imperfections, 4–10 tearing, 12 tensile behavior, 29–32 other factors, 31–2 weld orientation influence, 29–31 experimental and predicted tensile behavior, TWBs, 31 TWB tensile properties comparison, 31 tensile test simulation, 42 testing methods, 11–19 material discontinuity inspection, 15–19 eddy current, 17 electromagnetic acoustic transducer (EMAT), 16–17 light emission, 19 magnetic flux leakage detector, 17–18 ultrasonic transducer emitting sound waves, 15 ultrasound, 15–16 x-ray, 18–19 off-line testing, destructive, 11–13 cross-section, concave weld, 12 cross-sectioning, 12 cup test, 11–12 micro-hardness test, 13 tearing, 12 on-line testing, non-destructive, 13 surface inspection, 13–15 laser triangulation, 14–15 machine vision camera, 14 weld-width, lack of penetration (LOP), 14


1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

208 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 7 8 9 20 1 2 3 4 5 6 7 8 9 30 1 2 3 4 5 6 7 8 9 40 1 2 43X

Index

Tinius Olsen test, 11–12 titanium, 100–1 applications, 101 transformation-induced plasticity steels, 121 TRUDISK 6002 Yb:YAG laser 125 twinning-induced plasticity steels, 122 ultimate tensile strength, 30 ultrasound, 15–16 undercutting, 8–9 illustration, 9 USIBOR, 158 Voce law, 79 von Mises yield criterion, 76, 80 waffle plates, 192 weld imperfections, 4–10 concavity, 7–8 cross-section of good weld, 4 hard weld, 10 lack of fusion, 6–7 lack of penetration, 7 cross-section with top concavity, 8 monitoring, 20–2 result table, 22 sample production performance matrix, 21 sample quality system error table, 21 no-weld, 6

pinhole and crater, 5–6 cross-section, 5 pinhole in laser weld, 5 porosity, 9–10 inclusion induced pore, 9, 10 longitudinal cross-section, 9 prevention, 19–20 undercutting, 8–9 illustration, 9 weld integrity tailor welded blanks, 3–22 production quality control, 19–22 testing methods, 11–19 weld imperfections, 4–10 weld line movement, 36–8, 54 control, 37–8 presence of drawbead and drawbead dimension, 38 illustration, 36 square cup predicted weld line movement, 37 weld width evaluation methods, 26–32 rule of mixture (ROM) application, 27–9 tensile behavior, 29–32 WorldAutoSteel, 119 X-ray, 18–19 image of a sample defect, 18 YAG laser, 125 yield strength, 30


Tailor Welded Blanks for Advanced Manufacturing

Welded

Smart Devices and Machines for Advanced Manufacturing

Design of Advanced Manufacturing Systems: Models for Capacity Planning in Advanced Manufacturing Systems

Memory's tailor

Memory’s Tailor

Changeable and Reconfigurable Manufacturing Systems (Springer Series in Advanced Manufacturing)

The Tailor of Panama

Tinker Tailor Soldier Spy

The Tailor of Panama

Tinker, Tailor, Soldier, Spy

Emerging nanotechnologies for manufacturing

Effective Resource Management in Manufacturing Systems: Optimization Algorithms for Production Planning (Springer Series in Advanced Manufacturing)

Advanced Design and Manufacturing Based on STEP (Springer Series in Advanced Manufacturing)

ANEMONA: A Multi-agent Methodology for Holonic Manufacturing Systems (Springer Series in Advanced Manufacturing)

Advanced manufacturing: an ICT and systems perspective

The Tailor of Panama

Tinker, Tailor, Soldier, Spy

Frontiers in Computing Technologies for Manufacturing Applications (Springer Series in Advanced Manufacturing)

Effective Resource Management in Manufacturing Systems: Optimization Algorithms for Production Planning (Springer Series in Advanced Manufacturing)

Condition Monitoring and Control for Intelligent Manufacturing (Springer Series in Advanced Manufacturing)

The Tailor of Panama

The Tailor of Panama

Tinker Tailor Soldier Spy

Tinker, tailor, soldier, spy

Tinker, Tailor, Soldier, Spy

Tinker, Tailor, Soldier, Spy

welded joint design

Welded Design: Theory and Practice

Cumulative damage of welded joints

Particle Control for Semiconductor Manufacturing

Tailor Welded Blanks for Advanced Manufacturing

Welded

Smart Devices and Machines for Advanced Manufacturing

Design of Advanced Manufacturing Systems: Models for Capacity Planning in Advanced Manufacturing Systems

Memory's tailor

Memory’s Tailor

Changeable and Reconfigurable Manufacturing Systems (Springer Series in Advanced Manufacturing)

The Tailor of Panama

Tinker Tailor Soldier Spy

The Tailor of Panama

Tinker, Tailor, Soldier, Spy

Emerging nanotechnologies for manufacturing

Effective Resource Management in Manufacturing Systems: Optimization Algorithms for Production Planning (Springer Series in Advanced Manufacturing)

Advanced Design and Manufacturing Based on STEP (Springer Series in Advanced Manufacturing)

ANEMONA: A Multi-agent Methodology for Holonic Manufacturing Systems (Springer Series in Advanced Manufacturing)

Advanced manufacturing: an ICT and systems perspective

The Tailor of Panama

Tinker, Tailor, Soldier, Spy

Frontiers in Computing Technologies for Manufacturing Applications (Springer Series in Advanced Manufacturing)

Effective Resource Management in Manufacturing Systems: Optimization Algorithms for Production Planning (Springer Series in Advanced Manufacturing)

Condition Monitoring and Control for Intelligent Manufacturing (Springer Series in Advanced Manufacturing)

The Tailor of Panama

The Tailor of Panama

Tinker Tailor Soldier Spy

Tinker, tailor, soldier, spy

Tinker, Tailor, Soldier, Spy

Tinker, Tailor, Soldier, Spy

welded joint design

Welded Design: Theory and Practice

Cumulative damage of welded joints

Particle Control for Semiconductor Manufacturing

Recommend Documents