i
Friction stir welding
ii
Related titles: Advanced welding processes (ISBN 978-1-84569-130-1) This book introduces the range of advanced welding techniques currently in use. It covers gas tungsten arc welding (GTAW), gas metal arc welding (GMAW), high energy–density processes such as laser welding, and narrow gap welding methods. The book reviews general issues such as power sources, filler materials and shielding gases. Particular attention is given to monitoring and process control as well as to automation and robotics. MIG welding guide (ISBN 978-1-85573-947-5) Gas metal arc welding (GMAW) also referred to as MIG (metal inert gas) welding is one of the key processes in industrial manufacturing. MIG welding guide provides comprehensive, easy-to-understand coverage of this widely used process. The reader is presented with a variety of topics from the choice of shielding gases, filler materials, welding equipment and lots of practical advice. The book provides an overview of new developments in various processes such as: flux cored arc welding; new high productive methods; pulsed MIG welding; MIG brazing; robotic welding applications and occupational health and safety. This is essential reading for welding engineers, production engineers, designers and all those involved in industrial manufacturing. Weld cracking in ferrous alloys (ISBN 978-1-84569-300-8) Weld cracks are unacceptable defects that can compromise the integrity of welded structures. Most cracks are the result of solidification, cooling and stresses that develop due to weld shrinkage. Weld cracking can lead to structural failures, which at best will require remedial action and at worst can lead to loss of life. All industries that utilise welding can be affected, including nuclear, aerospace, automotive, shipbuilding and civil engineering. This book covers the processes of weld cracking in different ferrous alloys and for different welding technologies. It also covers methods of testing for weld cracks, avoidance and repair. Details of these and other Woodhead Publishing books can be obtained by: ∑ visiting our web site at www.woodheadpublishing.com ∑ contacting Customer Services (e-mail:
[email protected]; fax: +44 (0) 1223 893694; tel.: +44 (0) 1223 891358 ext. 130; address: Woodhead Publishing Limited, Abington Hall, Granta Park, Great Abington, Cambridge CB21 6AH, 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.
iii
Friction stir welding From basics to applications Edited by Daniela Lohwasser and Zhan Chen
CRC Press Boca Raton Boston New York Washington, DC
Woodhead
publishing limited
Oxford Cambridge New Delhi
iv Published by Woodhead Publishing Limited, Abington Hall, Granta Park, Great Abington, Cambridge CB21 6AH, UK www.woodheadpublishing.com Woodhead Publishing India Private Limited, G-2, Vardaan House, 7/28 Ansari Road, Daryaganj, New Delhi – 110002, India www.woodheadpublishingindia.com Published in North America by CRC Press LLC, 6000 Broken Sound Parkway, NW, Suite 300, Boca Raton, FL 33487, USA First published 2010, Woodhead Publishing Limited and CRC Press LLC © 2010, Woodhead Publishing Limited 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 publishers cannot assume responsibility for the validity of all materials. Neither the authors nor the publishers, 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. Library of Congress Cataloging in Publication Data A catalog record for this book is available from the Library of Congress. Woodhead Publishing ISBN 978-1-84569-450-0 (book) Woodhead Publishing ISBN 978-1-84569-771-6 (e-book) CRC Press ISBN 978-1-4398-0211-3 CRC Press order number: N10049 The publishers’ 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 acidfree and elemental chlorine-free practices. Furthermore, the publishers ensure that the text paper and cover board used have met acceptable environmental accreditation standards. Typeset by Replika Press Pvt Ltd, India Printed by TJ International Limited, Padstow, Cornwall, UK
v
Contents
Contributor contact details
ix
1
Introduction
1
D. Lohwasser, Airbus, Germany and Z. Chen, AUT University, New Zealand
1.1 1.2 1.3 1.4 1.5
Pre-word History of friction stir welding (FSW) Underlying patents Standards for friction stir welding (FSW) References
1 2 7 10 12
Part I General issues 2
The friction stir welding process: an overview
K. J. Colligan, Concurrent Technologies Corporation, USA
2.1
Overview of friction stir welding (FSW) process principles Comparison of friction stir welding (FSW) to other welding processes Welding tools used for friction stir welding (FSW) Parameter effects Materials used with friction stir welding (FSW) Joint geometries Joint preparation Post-weld heat treating References
19 27 28 29 31 35 37 38
3
Material deformation and joint formation in friction stir welding
42
R. Zettler, WTSH, Germany
3.1
Friction stir welding (FSW): a thermo-mechanical joining process
2.2 2.3 2.4 2.5 2.6 2.7 2.8 2.9
15
15
42
vi
Contents
3.2 3.3 3.4 3.5 3.6 3.7
Plastic deformation in relation to material properties Process parameter, temperature and heat loss relationships Material flow and property relationships of the resultant friction stir welded joint Concluding remarks Acknowledgement References
58 66 68 68
4
Friction stir welding equipment
73
T. Zappia, MTS Systems Corporation, USA, C. Smith, Friction Stir Link, USA, K. Colligan, Concurrent Technologies Corporation, USA, H. Ostersehlte, Airbus, Germany and S. W. Kallee, Germany
4.1
4.3 4.4 4.5 4.6 4.7 4.8 4.9 4.10 4.11 4.12
Requirements of friction stir welding (FSW) coming from the process and applications Overview of the machine requirements for friction stir welding (FSW) Friction stir welding machine controller requirements Closed-loop control and friction stir welding (FSW) Control of robotic friction stir welding (FSW) Other controller requirements Other machine requirements Machine requirements summary Part tooling requirements Friction stir welding (FSW) pin tools Machines currently available in the market place References
75 82 84 89 95 97 98 99 107 111 115
5
Industrial applications of friction stir welding
118
S. W. Kallee, Germany
5.1 5.2 5.3 5.4 5.5 5.6 5.7 5.8
Introduction Shipbuilding and offshore Aerospace Railway Automotive Other industry sectors Acknowledgements Conclusions
118 119 130 138 144 156 163 163
6
The future of friction stir welding
164
P. L. Threadgill, formerly with TWI Ltd, UK
6.1 6.2 6.3
Introduction Process aspects of friction stir welding (FSW) Materials aspects of friction stir welding (FSW)
4.2
44 46
73
164 167 172
Contents
vii
6.4 6.5 6.6
Summary Acknowledgements References
179 180 180
7
Inspection and quality control in friction stir welding
183
T. Zappia, MTS Systems Corporation, USA
7.1 7.2 7.3 7.4 7.5
Weld quality requirements definition Online monitoring and statistical process control Offline testing: non-destructive testing Summary References
183 188 196 207 212
Part II Variables in friction stir welding 8
Residual stresses in friction stir welding
S. W. Williams, Cranfield University, UK and A. Steuwer, ESS Scandinavia, Sweden
8.1 8.2
Residual stresses produced by welding Determination of residual stresses in friction stir welding (FSW) Effects of residual stresses produced by welding Mitigation of residual stresses and their effects Residual stresses in friction stir welding (FSW) Active control of residual stresses in friction stir welding (FSW) References
215
245
8.3 8.4 8.5 8.6 8.7
215
219 221 225 231 234 241
9
Effects and defects of friction stir welds
R. Zettler, WTSH, Germany, T. Vugrin, Airbus, Germany and M. Schmücker, German Aerospace Centre, Germany
9.1 9.2 9.3 9.4 9.5 9.6 9.7
Background and introduction Defects from too hot welds Defects from too cold welds Defects from geometrical mistakes Features without significant effect Acknowledgement References and further reading
245 247 255 259 265 272 272
10
Modelling thermal properties in friction stir welding
277
H. N. B. Schmidt, Technical University of Denmark (DTU) and HBS Engineering, Denmark
10.1
Introduction
277
viii
Contents
10.2 10.3 10.4
Analytical model of heat generation Numerical thermal model References and further reading
280 292 312
11
Metallurgy and weld performance in friction stir welding
314
J. F. dos Santos, GKSS Forschungszentrum GmbH, Germany, C. A. W. Olea, Vallourec & Mannesmann Tubes do Brasil, Brazil, R. S. Coelho, Helmholtz-Zentrum Berlin für Materialien und Energie GmbH, Germany, A. Kostka, Max-Planck-Institut für Eisenforschung GmbH, Germany, C. S. Paglia, University of Applied Sciences of Southern Switzerland, Switzerland, T. Ghidini, European Space Agency, The Netherlands and C. D. Donne, EADS Innovation Works, Germany
11.1 11.2 11.3
314 317
11.5
Introduction Metallurgy of friction stir welds in aluminium alloys Corrosion behaviour of friction stir welds in aluminium alloys Mechanical properties of friction stir welds in aluminium alloys References and further reading
Index
411
11.4
360 371 400
ix
Contributor contact details
(*= main contact)
Chapter 1
Chapter 3
Daniela Lohwasser* Airbus Airbus-Allee 1 28199 Bremen Germany
Rudolf Zettler WTSH – Business Development and Technology Transfer Corporation of SchleswigHolstein GITZ, Max-Plank-Strasse 2 Geesthacht 21502 Germany
E-mail:
[email protected]
Zhan Chen Department of Mechanical and Manufacturing Engineering School of Engineering AUT University Auckland New Zealand E-mail:
[email protected]
Chapter 2 Kevin J. Colligan Concurrent Technologies Corporation 15091 Alabama Highway 20, Suite A Madison, AL 35756 USA E-mail:
[email protected]
E-mail:
[email protected] (Formerly with RIFTEC, Germany)
Chapter 4 Timothy Zappia* MTS Systems Corporation 14000 Technology Drive Eden Prairie, MN 55344 USA E-mail:
[email protected]
Chris Smith Friction Stir Link USA
x
Contributor contact details
Kevin J. Colligan, Concurrent Technologies Corporation 15091 Alabama Highway 20, Suite A Madison AL 35756 USA
Chapter 7
E-mail:
[email protected]
E-mail:
[email protected]
Hartmut Ostersehlte Airbus Airbus-Allee 1 28199 Bremen Germany
Chapter 8
E-mail:
[email protected]
Stephan W. Kallee Im Unterdorf 19 63826 Geiselbach Germany E-mail:
[email protected]
Chapter 5 Stephan W. Kallee Im Unterdorf 19 63826 Geiselbach Germany E-mail:
[email protected]
Chapter 6 Philip L. Threadgill formerly with TWI Ltd Granta Park Cambridge CB21 6AL UK E-mail:
[email protected]
Tim Zappia MTS Systems Corporation 14000 Technology Drive Eden Prairie, MN 55344 USA
Stewart W. Williams* Welding Engineering Research Centre Building 46 Cranfield University Cranfield MK43 0AL UK E-mail:
[email protected]
A. Steuwer ESS Scandinavia University of Lund Stora Algatan 4 22350 Lund Sweden
Chapter 9 Rudolf Zettler* WTSH – Business Development and Technology Transfer Corporation of SchleswigHolstein GITZ, Max-Plank-Strasse 2 Geesthacht 21502 Germany E-mail:
[email protected] (Formerly with RIFTEC, Germany)
Contributor contact details
Tamara Vugrin Airbus Airbus-Allee 1 28199 Bremen Germany E-mail:
[email protected]
Martin Schmücker Deutsches Zentrum für Luft- und Raumfahrt (German Aerospace Centre) Institut für Werkstoff-Forschung (Institute of Materials Research) Linder Höhe 51147 Köln Germany E-mail:
[email protected]
Chapter 10 Henrik N. B. Schmidt* Department of Mechanical Engineering Technical University of Denmark Produktionstorvet 425 2800 Kgs.-Lyngby Denmark E-mail:
[email protected]
and HBS Engineering www.hbs-engineering.com E-mail:
[email protected]
xi
Chapter 11 Jorge F. dos Santos* GKSS Forschungszentrum GmbH Institute of Materials Research Materials Mechanics and Joining Solid State Joining Processes D-21502 Geesthacht Germany E-mail:
[email protected]
Cesar Afonso Weis Olea Vallourec & Mannesmann Tubes do Brasil Quality and R&D – Plug Mill Belo Horizinte MG Brazil Rodrigo Santiago Coelho Helmholtz-Zentrum Berlin für Materialien und Energie GmbH D-12489 Berlin Germany Aleksander Kostka Max-Planck-Institut für Eisenforschung GmbH Material Diagnostics and Steel Technology D-40237 Düsseldorf Germany Christian S. Paglia University of Applied Sciences of Southern Switzerland Institute of Materials and Construction Trevano CH-6952 Canobbio Switzerland
xii
Contributor contact details
Tommaso Ghidini ESA (European Space Agency) Product Assurance and Safety Department Materials and Components Technology Division Noordwijk The Netherlands Claudio Dalle Donne EADS Innovation Works Metallic Technologies and Surface Engineering 81663 Munich Germany
1
Introduction
D. Lohwasser, Airbus, Germany and Z. Chen, AUT University, New Zealand
Abstract: This introduction gives a brief overview about the content of this book. The history from the first welding using gas and arc welding followed by resistance and later fusion welding processes (plasma, laser) is briefly described together with the need to develop friction and friction stir welding (FSW) processes. The rapid introduction of FSW is underlined by showing the numbers of publications and patents as well as the evolving standards on this process for ship, train and aerospace applications. Key words: friction stir welding, book introduction, quick history of welding and FSW invention, publications and patents, FSW standards.
1.1
Pre-word
The first time people hear about friction stir welding (FSW) they usually state: “How does it work?”. After a brief introduction describing the process their next reaction is always disbelief: “And that works?”. Without seeing the very unspectacular process themselves – looking, touching, feeling – they cannot believe the extraordinarily good properties described to them. A very impressive way to demonstrate these properties is to give them a bending sample not yet bent and let them perform the 180° bending test themselves. This is a very effective way of getting rid of any remaining doubts. Watching the process is always impressive, although as said earlier, it is quite unspectacular. All one sees is a rotating tool pushing through material. Usually there are no fumes, spattering, red glowing material and the noise is also quite low, making it difficult to believe that such a simple process can produce such good joints (it is similar to inspecting a vacuum cleaner which is not loud enough when it is used – people will not believe that its suction performance is sufficient). So for FSW there is a great need to demonstrate and show the process to the world in order to have a widespread introduction into many industry sectors. This book aims to describe the main process principles followed by a deep analysis of the material flow helping to understand why this process works so well. After that, a description of the equipment requirements followed by equipment examples is given. With today’s knowledge on FSW, numerous applications have evolved which are summarized in an extensive way, followed 1
2
Friction stir welding
by a chapter giving an outlook on the research and development work to come in the near future. In the second part of this book, the complex topic of residual stress in FS welded structures is first discussed. This is followed by a detailed discussion on the effects of defects, which is a topic of high practical importance. To minimize the testing needs, a detailed description of the modelling achievements is given, especially on thermal modelling, which is for both better process understanding and for selecting FSW parameters. In the final chapter, an in-depth discussion on the metallurgy and the evolving mechanical and corrosion properties of FS welded joints are described.
1.2
History of friction stir welding (FSW)
Although joining pieces together can be traced back more than 2000 years, welding emerged as a viable manufacturing process only in the late 1800s (Messler Jr, 2004). This emergence followed a number of important discoveries and inventions: acetylene and oxyacetylene blowpipe leading to gas welding, arc between two carbon electrodes leading to arc welding, and joule heating leading to resistance welding. While gas welding was practised more widely in the early part of 1990s, arc welding has since gradually become a popular and major welding method as many advances have been made in various aspects of arc welding processes. Gas and arc welding processes are fusionwelding, meaning that locally the locations the pieces to be welded first melt together and subsequently solidify. For arc fusion welding, the density of the heat energy is usually low. Low energy density fusion welding processes result in a wide weld (molten) metal zone and wide heat-affected zones, resulting in a high level of solidificationrelated defects, strength reduction and distortion. To address this, in the late 1950s, a more constricted arc (plasma arc) technique was developed and plasma arc welding reduces the adverse effects of low energy density heat sources. Also in the late 1950s, with the invention of laser, which enables a high concentration of heat source and thus deep penetration, laser beam welding was soon developed. Nowadays laser beam welding is used industrially and is capable of making a thick weld with very narrow molten weld zone and heat-affected zones. Thus, quality and properties of laser welds are generally high. Before the invention of FSW, there had been some important technological developments of non-fusion welding processes, which have found some limited industrial uses. A significant process of these is friction welding developed at the time just before laser was invented. During friction welding, the pieces to be welded are compressed together and are made to move relative to each other. Thus frictional heat is generated to soften the material in the joining region. The final step is made by applying increased pressure to the softened material to yield a metallurgical joint without melting the joining
Introduction
3
material. However, the relative movement during the stage of heat generation and material softening can practically only be rotational or linear. Although friction welding operation is simple, the welding geometry is quite restricted and thus its use is also limited. As has been described, the demand for increasing energy density of fusion welding is a major driving force for a number of important innovations and developments in fusion welding. For solid state welding, the thermomechanical principle of friction welding had actually laid an important base for the later invention of FSW. The Welding Institute (TWI) in the UK had for years engaged in various R&D and industrial activities on friction welding and surfacing. Wayne Thomas and his colleagues in TWI had long worked on and developed a number of variants of friction welding. In particular, they developed friction extrusion, friction hydropillar processing and third-body friction joining processes. Over the long period of working on and developing these materials processes, the group in TWI observed and studied a number of important phenomena and accumulated an in-depth working knowledge of those processes. These include the highly plasticized third-body effect and the transportation phenomena of the plasticized material, the adiabatic heating during deformation, and the relationship between the torque and rotation speed when a sufficient amount of the plasticized material is present during processing. With the in-depth understanding and working knowledge of the various friction-based material processes, a welding technique with an effective transportation mechanism for the plasticized material suited for a wide range of geometries of structures to be welded was never far from the minds of Wayne Thomas and his colleagues. One day in 1991, in a flash of inspiration, Wayne Thomas realized that with the use of a rotational probe of a harder material than the workpieces, the workpiece material could be plasticized and an effective transportation mechanism could be provided for the plasticized material to join the workpieces together. This eventual moment of realization, after a long period of gestation, marked the discovery of FSW as we know it nowadays. Based on this discovery, the engineering restriction is quite low for applying the simple friction stir action provided by the probe for welding to a very large range of structures/parts and a wide range of weld geometry. Welding without gross melting is highly significant, as welds can then be readily made free of solidification-related porosities and cracking and with low distortion. Furthermore, there is no need to use protective gases, at least for aluminium alloys, no arc-related emissions and no fumes. Thus FSW is an environmentally friendly process. Furthermore, no filler material is needed during FSW. This gives a further advantage that formation of unwanted phases in the weld microstructure due to the mixing of the filler metal and parent metal, which often differ to a certain extent in alloying,
4
Friction stir welding
can be avoided. The industrial significance of FSW, due to its distinctive advantages, was immediately realized. Wayne Thomas and his colleagues in TWI, as will be described in next section, soon filed the first patent on FSW (Thomas et al., 1991). To date it is with aluminium alloys that FSW is most successfully applied. The reason for the predominant use of FSW on aluminium alloys is a combination of process simplicity in principle and the wide use of aluminium alloys in many major industries. It is especially the case where some aluminium alloys are difficult to fusion weld as, for example, is clearly evident in FSW application made by Boeing for making the Delta 2 rocket tanks. When the variable polarity plasma arc welding process was used, the defect rate was over 90%. FSW allowed them to dramatically reduce their defect rate to nearly zero. Maximum temperature during FSW can reach just below the solidus of the workpiece alloy. For most aluminium alloys, it is significantly less than 660∞C. Thus, H13 tool steel or high-speed tool steel, which is quite inexpensive, is a satisfactory tool material. Thus, FSW of aluminium alloys is relatively straightforward, although FS engineering, particularly for components and structures of high geometry complexity, can be quite challenging. In principle, FSW could be applied for welding of all solid metallic materials. The practical restriction is primarily the integrity issue of the tool during FSW of high temperature materials. For example, during FSW of steels, the local operating temperature generated by both friction and deformation needs to be at 1100–1200°C so that the workpiece material is sufficiently plasticized for stirring and welding. Such high operating temperatures and the necessary forces acting on the tool during FSW create an extraordinary demand on the mechanical properties of the tool material. As will be described in Chapter 4, forces and torque during FSW require the structures/parts to be welded quite rigidly clamped and the tool rigidly positioned. The precision of positioning also needs to be high. Thus, from an engineering point of view, it is reasonable to suggest that the application of FSW is more suited to fabrication environments. As it has turned out, during the last 15 years, FSW has actually enabled many manufacturing processes to become significantly more efficient or enabled new and efficient manufacturing routes to be developed. The first FSW industrial application was a good example of FSW not only being a welding technology but also an enabling technology. After an initial intensive industrial FSW development programme (TWI and Hydro Aluminium) specifically for aluminium shipbuilding, FSW was successfully applied in 1995 to weld aluminium extrusions into large panels with minimum distortion for shipbuilding. The use of these large prefabricated panels without the large amount of fit up work associated with the use of fusion welding had resulted in a significantly more cost-effective route of building fishing
Introduction
5
boats and high-speed ferries. Since then, design and building of various types and sizes of sea vessels using prefabricated panels by FSW has become a normal and cost-effective industry practice. Not long after the first major application of FSW, Hitachi in Japan started the intensive development of applying the technology. One of the major applications has resulted from the use of FSW for welding long aluminium extrusions. The best example of this is the manufacturing of modern traincar bodies by FSW of long aluminium extrusions into double- or singleskin structures. The use of FSW with low distortion eliminates post-weld straightening and filling. Because of this feature of welding long extrusion with very low distortion, by 2000, Hitachi had incorporated FSW in their cost-effective modular design and manufacturing of high integrity high-speed train-car bodies. However, the biggest interest in FSW has actually been from the aerospace industry, starting FSW development work in the early-mid 1990s on structures such as the Delta rocket fuel tank (Boeing), aircraft structures (Airbus) and external fuel tank barrel section (NASA). These and their related organizations, through the development and application of FSW to many aerospace structures over the years, have made a huge forward step in FSW. In late 1990s, the newly formed Eclipse Aviation designed a manufacturing route of making small jets incorporating the intensive use of FSW in the major assembling of Eclipse 500 jets. They have eliminated 70% of their rivets on the aircraft by welding mainly overlap connections in 2024, 7075 or dissimilar 2024 to 7075 alloys. By doing so they have been able to accomplish a production rate of four aircraft per day, which will then lead to significant, reduced lead times for their customers. It is worthwhile to point out that it is quite logical for the aerospace industry to have a strong interest in applying FSW technology. Traditionally, riveting is the dominant joining method in aircraft manufacturing, as the bulk of the high strength aircraft grade aluminium alloys are susceptible to weld defects associated with fusion welding. Since it is a solid state welding process and thus there is no solidification-associated porosity and cracking, all the traditionally unweldable aluminium alloys have become weldable using FSW. Details on this metallurgical advantage using FSW will be explained in Chapter 3 and Chapter 11. It is clear that during the late 1990s, industrial uptake of FSW was gathering a strong pace and an international community of R&D on FSW technology was forming rapidly. By 1998, FSW had started to be a major theme of discussion in a major international conference on welding (Table 1.1). The year after, the first International Symposium on Friction Stir Welding (ISFSW) was organized by TWI and held at Rockwell Science in USA. In this first international FSW conference, 30 presentations were given covering the above-mentioned early industrial applications of FSW
6
Table 1.1 A list of FSW conferences or conferences with FSW as a major theme 1997 1998 1999 2000 2001 2002 2003 2004 2005 2006 2007 2008
Friction stir welding theme in Research Trends in Welding, ASM
∑
International Symposium on Friction Stir Welding, TWI
∑
∑
Friction stir welding theme in AeroMat Conference & Exposition, ASM
∑
∑
∑
∑
∑
∑
∑
∑
∑
∑
∑
∑
Friction Stir Welding and Processing, TMS
∑
∑
∑
∑
∑
Aerospace Friction Stir Welding Symposium, SAE
∑
IIW Pre-Assembly Meeting on Friction Stir Welding, IIW
∑
Friction stir welding conference, AWS
∑
Friction stir processing/welding theme in Thermec Friction stir welding conference, GKSS Workshops
∑
∑
∑
∑
∑
∑
∑
∑
∑
Friction stir welding
Conference
Introduction
7
technology, together with research on process development, structure and property evaluation, and research relating to the thermal-mechanical aspects of the process. Feasibility of FSW of steels was also discussed. This series of ISFSW (Table 1.1) has continued as an annual event from 1999 to 2004 with the exception of 2002 and since 2004 a biennial event has been held rotating in various continents with the most recent one held in Japan in May 2008. The next in this series will be held in Germany in 2010. The strong R&D work on FSW technology starting from the early part of this 2000s can further be seen in another series of international conferences, Friction Stir Welding and Processing (FSW/P) organized by TMS and held in the USA every two years starting from 2001. This series of FSW/P conferences has been run in parallel to ISFSW and now in alternating years. Further to the two major series of FSW conferences, there have also been a number of other FSW conferences (Table 1.1). Overall, in only a short span of time (10 years) more than 30 FSW conferences or conferences with FSW as a major theme have reflected the extraordinary interest in FSW technology in many industries as well as in the wider metal working and welding research communities. In the latest ISFSW, the number of presentations was high (83 oral and 19 poster), covering FSW topics: process control and development, microstructure evolution during FSW and subsequent properties, modelling of heat and material flow, friction stir spot welding, and particularly the recent advancement of FSW of steels and titanium alloys. The trend in the number of scientific research papers published in peer review journals each year has followed a similar trend to the major conference events. Figure 1.1 shows the numbers of published papers on FS-related topics, sourced from Web of Science. The distribution of the total ~ 1000 papers, mostly refereed research papers directly on FSW, has shown an increasing intensity of FSW research in the last few years and this intensive research effort seems to be quite certain to continue for the near future. Detailed overviews on some of the research topics are given in various chapters of this book. An assessment on the current state of FSW used commercially is given in Chapter 6 where new development and future outlook of the technology are described. The technological development of FSW thus far may also be illustrated by examining the patents filed starting with the first TWI patent in 1991. This will be described in the following section.
1.3
Underlying patents
The original patent for friction stir welding was filed by The Welding Institute (TWI) in Cambridge. It was initially filed in the United Kingdom in December 1991 and was granted as EP 0 615 480. There are two major claims in the patent. Claim 1 states: “A method of joining workpieces
8
Friction stir welding 220 200
Numbers of papers from ISI
180 160 140 120 100 80 60 40 20 0 1992 1993 1994 1995 1996 1997 1998 1999 2000 2001 2002 2003 2004 2005 2006 2007 2008 Year (since the invention of FSW)
1.1 Numbers of FS (welding and processing) papers published in (mostly) refereed research journals, from Web of Science.
defining a joint region there between, the method comprising carrying out the following steps without causing relative bodily movement between the workpieces: causing a probe of material harder than the workpiece material to enter the joint region and opposed portions of the workpieces on either side of the joint region while causing relative cyclic movement between the probe and the workpieces whereby frictional heat is generated to cause the opposed portions to take up a plasticized condition; removing the probe; and allowing the plasticized portions to solidify and join the workpieces together.” In Claim 2 the movement of the probe along the joint line is described. Further to this, many more items are already identified in this original patent such as the materials, interrupted or spot welding, crack repair welding, tool configurations regarding motions (rotation, oscillation, reciprocation), adjustable and self-reacting as well as heating of tools. TWI’s interest has not been to keep this technology only for themselves, but they have made great efforts to spread this technology into the world engineering community by issuing FSW licences to a wide range of organizations including end users, equipment suppliers, academia and R&D institutes. As of April 2008 about 200 licences have been granted. These organizations have further developed the process and its applicability. This becomes obvious when analysing the high number of patent filings. Especially the end users have used the process of filing patents in order to protect the use of FSW in their applications, which makes up about 90% of the filed patents. In Fig. 1.2 the number of patents filed since 1996 can be seen. Most of the patents are filed in Japan as can be seen on the grey bar of the Fig. 1.2.
Introduction
9
2500 Patent cooperation treaty European patent office/GB 2000
Japan patent office United States patent and trademark office
1500
1000
500
0 1996
1997 1998
1999 2000
2001 2002 2003 2004 2005 2006 2007
1.2 Third party patent filings since 1996 (Smith and Lord, 2008). 1200 Failed or waiting Granted
1000 800 600 400 200
57%
21%
United States patent and trademark office
Japan patent office
0
40% European patent office
1.3 Granted patents vs. filed patents (Smith and Lord, 2008).
Of course the number of filed patents does not correspond to the number of granted patents, as can be seen in Fig. 1.3. Hitachi (railcars) has filed more than 45% of these granted patents and is therefore the company whose strong philosophy is to protect their applications with these patents. The next organizations following with a rate of about 10% are Boeing (aerospace), KHI (rail and automotive), Showa (aluminium supplier) and NLM (aluminium supplier). Nearly half the granted patents deal with a specific product or industrial applications. Another quarter deal with tools and equipment according to Smith and Lord (2008), as can be seen in Table 1.2.
10
Friction stir welding Table 1.2 Main topics of granted patents Main focus of patent
Total number
Product/application Process variations Apparatus/equipment Tool shape and configuration Friction stir spot welding (FSSW)
300 101 93 54 33
Not many processes have ever reached as many patents as FSW has reached in only 15 years. This process of FSW has very quickly found its way into industries. For new end users, of course, an important question to ask is – Am I infringing on a patent, if I use it on my application? – But normally the answer should be no, as most of the use of FSW is included in the basic patent of TWI to which a licence is accessible and therefore further applications should not be prevented.
1.4
Standards for friction stir welding (FSW)
As has been briefly described and will be detailed later in this book, FSW produces welds that are free of defects associated with local melting and solidification, which are characteristics of the traditional fusion welding processes. However, quality and reliability of FS defect-free welds and reproducibility of the process can only be obtained following acceptable FS practice. In this regard, FSW is no different from the many traditional welding processes. For the latter, many stringent standards, both national and international, have been written over the last many decades and are followed in practice. However, as the operational principle of FSW differs completely from those of fusion welding processes, the existing welding standards cannot be directly applied to FSW. Take, for instance, the earliest application of FSW – shipbuilding in Europe. For the traditional fusion welding for shipbuilding, welding procedures and examination procedures for assessing quality of welds strictly follow the procedures set by the Classification Society or national or international standards. When FSW was first developed and applied commercially (for shipbuilding), as there were no standards that could be applied directly, classification rules urgently needed to be established for the application of FSW in shipbuilding. For this, in the mid-late 1990s, Lloyd’s Register of Shipping issued a Guideline for Approval of Friction Stir Welding. As an illustrative example, it is now an established practice to FS weld aluminium alloy extrusion panels in a FSW operation site and then the welded panels are transported to shipyards for shipbuilding. In order to meet the product qualification that is certified by Lloyd’s Register of Shipping, the
Introduction
11
existing fusion welding code of BS EN ISO 151614-2:2005 with the aid of Lloyd’s Guideline for Approval of Friction Stir Welding is adapted by FSW operators. The guideline specifies the requirements of welding personnel, welding procedure specification, base materials to be FS welded and various non-destructive and destructive tests. After the initial product qualification is gained, samples from production batches are routinely taken and tested, as specified in the Lloyd’s Guideline. As has been described, the aerospace industry has shown the biggest interest in FSW. However, it is possible that the pace of the actual uptake of FSW for the widespread use in aircraft manufacturing may be significantly slower compared to those in shipbuilding and train-car body making. The reason for this may be due in part to the lack of international standards for FSW and design codes, which may then result in the certification process of any structures made using FSW a lengthy or even a very difficult one. To address the issue of FSW standards, NASA started formulating a Process Specification of Friction Stir Welding at the beginning of 2000. The first version of it was issued in 2002 and was comprehensively reviewed and revised in 2007. Also in 2005, an AWS committee was set up to work on Specification for Friction Stir Welding of Aluminium for Aerospace Applications. This specification for aerospace hardware has just gone through the public review and is due to be formally issued soon. It is worthwhile to look briefly at the experience of Eclipse Aviation, who pioneered the use of FSW from the beginning of manufacturing their small jets in 2000. They started building FSW machines with MTS systems corporation in 2001 and at the same time started working closely with Federal Aviation Administration (FAA) on the certification of FSW technology. By 2002, Eclipse Aviation was awarded the approval of FSW process specification by FAA and at the same time FSW was successfully used for the assembly of the lower cabin of the first Eclipse 500 jet. Eclipse’s experience may suggest that a certification process for aerospace structure manufacturing using FSW should not be perceived as an extraordinarily difficult process. For FSW to be widely applied across the many different industries, particularly with the growing importance of the ISO 9001 process adapted worldwide, the need for an international standard on FSW had become increasingly more prominent. This had been recognized not only by FSW engineers and practitioners worldwide, but also the wider international engineering and welding associations and bodies. Responding to this need, an international effort on preparing an ISO (the International Organization for Standardization) standard for FSW began in 2003. The standard, designed and named ISO/DIS 25239: Friction Stir Welding – Aluminium, has been prepared by a working group, WG B-I within Commission III of the International Institute of Welding (IIW), led by Mr David Bolser of Boeing Company. This group included FSW experts and
12
Friction stir welding
practitioners from many IIW member countries. Over the last 5 years, this group (particularly D. Bolser and D. Miller of the USA, W. Thomas of UK, J. dos Santos of Germany, L. Mohlkert of Sweden, H. Gerard of France, K. Namba and S. Matsuoka of Japan) has worked tirelessly, drafting and revising draft after draft following the many international reviews on the drafts and voting procedures set by the IIW and ISO. This standard has now progressed to a very advanced version and may be soon be submitted to ISO/CS for publication. ISO/DIS 25239 is a comprehensive standard consisting of five parts: Part Part Part Part Part
1 2 3 4 5
– – – – –
Vocabulary; Design of weld joints; Qualification of welding operators; Specification and qualification of welding procedures; Quality and inspection requirements.
In this present book, the nomenclature adopted has closely followed those defined in Part 1 of the ISO standard. On the other hand, in various chapters of this book, many important aspects relevant to industrial applications of FSW are described: the principle of FSW as well as the different weld geometries; FSW machines and the associated tooling, clamping and controlling mechanisms; qualities and properties of welds as well as the relevant inspection and testing techniques. Thus, this present book should serve as a good reference source for understanding the different parts of the ISO standard.
1.5
References
Messler Jr R W, Principles of Welding – Processes, Physics, Chemistry, and Metallurgy, Wiley-VCH, 2004. Smith Iain J and Lord Daniel D R, TWI Ltd – FSW Patents – A Stirring Story, 7th International Symposium on Friction Stir Welding Licensees Meeting, May 2008. Thomas W M, Nicholas E D, Needham J C, Murch M G, Temple-Smith P, and Dawes C J, “Improvements relating to friction welding”, European Patent Specification 0 615 480 B1 1991.
2
The friction stir welding process: an overview K. J. Colligan, Concurrent Technologies Corporation, USA
Abstract: This chapter introduces the basic concepts relevant to the use and study of friction stir welding (FSW), including an overview of the process, a comparison to arc welding processes, a discussion of welding tool design and materials, the effect of process parameters, workpiece materials and joint geometries. References are given to point to early contributions in the various areas of study and to the latest progress in the field. Key words: friction stir welding, joining, solid state joining.
2.1
Overview of friction stir welding (FSW) process principles
Friction stir welding (FSW) produces welds by using a rotating, nonconsumable welding tool to locally soften a workpiece, through heat produced by friction and plastic work, thereby allowing the tool to “stir” the joint surfaces. The dependence on friction and plastic work for the heat source precludes significant melting in the workpiece, avoiding many of the difficulties arising from a change in state, such as changes in gas solubility and volumetric changes, which often plague fusion welding processes. Further, the reduced welding temperature makes possible dramatically lower distortion and residual stresses, enabling improved fatigue performance, new construction techniques, and making possible the welding of very thin and very thick materials. Owing to the typically high forces in the process, FSW is usually practiced as a fully mechanized process, increasing the cost of the equipment compared to arc welding techniques, while reducing the degree of operator skill required. FSW has also been shown to eliminate or dramatically reduce the formation of hazardous fumes and reduces energy consumption during welding, reducing the environmental impact of the joining process. Further, FSW can be used in any orientation without regard to the influence of gravitational effects on the process. These distinctions from conventional arc welding processes make FSW a valuable new manufacturing process with undeniable technical, economic, and environmental benefits. Central to the FSW process is the design of the welding tool, as shown 15
16
Friction stir welding
schematically in Fig. 2.1. Many variations and new features have been added to this basic tool, as will be discussed further below and in Chapter 4. Conventional FSW, as the process was originally conceived, is done with a welding tool consisting of a shoulder, which rides on the surface of the workpiece, and a smaller diameter pin, which nearly penetrates the workpiece. The shoulder essentially performs the role of the “lid on the pot”, which prevents the escape of softened workpiece material as the tool is rotated and forced along the joint. The pin commonly employs thread-shaped features which act to push the surrounding workpiece material downward, assisting in the retention of material within the weld zone. The downward force applied to the tool to maintain the correct plunge depth also results in forcible contact between the shoulder and the workpiece surface, and relative motion from the tool rotation results in significant heat generation from friction at the shoulder interface. In conventional FSW the pin accomplishes the breakup of the original faying surfaces of the joint. For this reason, the pin must penetrate to within 0.5mm of the back of the workpiece to ensure complete penetration of the weld through the workpiece. Features cut into the pin surface, originally demonstrated as downward-pushing screw threads, prevent the formation of pores or voids in the weld. The pin generates heat by both friction and plastic work, and both seizure and sliding contact have been observed and predicted by modeling results. The design of the pin and shoulder has been an area of intense research since the conception of FSW, which has resulted in improvement in throughput, joint strength, and weld quality, and in the range of materials, joint geometries, welding parameters, and workpiece thickness that can be welded. Tilt angle
Plunge force Spindle speed
Heel plunge depth
Welding speed Shoulder diameter Pin diameter
Shoulder PIn
2.1 Conventional FSW tool and key variables.
The friction stir welding process: an overview
17
A transverse section from a typical, conventional FSW joint is shown in Fig. 2.2. The weld is bounded on either side by unaltered, base metal (BM). Although BM near the weld zone does experience elevated temperature during welding, this material exhibits essentially the same properties as the workpiece in the as-received condition. Closer to the weld is the heat-affected zone (HAZ), which is heated sufficiently during welding to alter its properties without plastic deformation of the original grain structure. The alteration of properties in the HAZ may include changes in the strength, ductility, corrosion susceptibility, and toughness of the workpiece, but typically will not include changes in grain size or chemical makeup. Heating in the HAZ is generally high enough in aluminum alloys to result in recovery of cold work and coarsening of precipitates, which is the root cause of changes in properties in this region. The thermomechanically affected zone (TMAZ) encompasses all of the plastically deformed material within the joint region. In this region, the workpiece is sufficiently heated and softened and the process forces are sufficiently high, to result in plastic deformation of the original grain structure. The TMAZ can be further divided into the unrecrystallized TMAZ and the nugget, or recrystallized TMAZ. In aluminum alloys, the unrecrystallized TMAZ may be an important feature in the weld, since it can be of significant size and can represent a region of low microhardness and increased corrosion susceptibility. Further, in aluminum alloys the nugget material is generally composed of fine grain size material and is considered to have experienced severe plastic deformation due to interaction with the welding tool pin and in some cases
HAZ
Base metal
Recrystallized TMAZ
TMAZ
2.2 Typical conventional FSW transverse section in 25.4-mm thick 2195 aluminum-lithium plate.
18
Friction stir welding
may actually mimic the shape of the pin profile. However, in materials that experience thermally induced phase transformation, the TMAZ may consist entirely of recrystallized material, while in other materials the TMAZ may be completely unrecrystallized, without regard to the size or shape of the pin. Conventional FSW is typically carried out by first rigidly fixing the plates to be joined in a welding fixture, as shown in Fig. 2.3. Fixture design is a very important consideration in FSW, which is discussed in detail in Chapter 4. The plates are typically fixed with no gap at the joint line. The process requires that the workpieces be prevented from spreading or lifting during welding, so welding fixtures are typically equipped with features which restrain the workpiece. It is common that FSW fixtures are equipped with a removable anvil insert which can be replaced in the event of inadvertent damage to the anvil from contact with the welding tool pin. Since the anvil insert is very closely coupled to the workpiece at the point of welding in terms of heat transfer, it is important to consider the mass and diffusivity of the anvil insert when designing FSW fixtures. The FSW process can be thought to consist of three phases: the plunge phase, where the weld is initiated; the main phase, where the weld is made; and the termination phase, where the welding tool is withdrawn from the workpiece. The properties of the weld produced are, of course, dependent on the process parameters selected for each phase of the weld, so great care must be taken in establishing these settings. The plunge phase consists of inserting, or “plunging”, the rotating welding tool into the joint. This is typically accomplished by commanding the welding system to drive the tool pin axially into the workpiece at a specific rate or with a specific force. Frictional heating and pressure at the end of the pin induce workpiece material to displace, forming a ring of expelled, plastically deformed material around the pin as the pin enters the workpieces. As the tool is plunged into the joint, heat generated conducts into the surrounding material and the anvil. The plunge phase may be facilitated by drilling a hole at the plunge location, reducing the heat and forces produced. Alternatively, Vertical restraint Lateral restraint
Anvil insert Workpieces
2.3 Conventional FSW fixture requirements.
The friction stir welding process: an overview
19
the welding tool may be plunged into the side of the workpiece, although this approach is less commonly applied. Once the welding tool is plunged into the workpiece, the tool is typically driven laterally along the joint without delay, although in some materials, it may be necessary to dwell at the plunge location for some time in order to allow for the welding tool and workpiece to reach a higher temperature. Once the welding tool begins to travel along the joint, friction and plastic work produce heat to maintain sufficient softening in the workpiece to permit material flow around the pin. Features cut into the pin surface, such as screw threads, flats, and spiral grooves, facilitate this material flow by increasing drag between the pin and the surrounding material in such a way as to prevent the formation of internal voids or fractures. Heat from the welding process conducts within the workpiece, serving to precondition the material in front of the tool, producing softening from recovery of work hardening and overaging in materials such as aluminum. This metallurgical alteration may be slight, such as in when welds are made at very high travel speed, or it may dramatically soften the workpiece. As this preconditioned workpiece material interacts with the features cut into the welding tool, the material is rapidly deformed and heated by quasi-adiabatic friction and plastic work, raising the temperature to near the solidus temperature, but macroscopically not above it. Simultaneously, this material is pulled around the welding tool and deposited behind it in a way that prevents the formation of voids. It should be noted that the main phase of the weld can be thought of as consisting of an initial period, where the temperature distribution within the welding tool and workpiece is being established, a steady-state period, where the temperature distribution is stabilized, and a terminal period, where the temperature distribution is altered by the approaching workpiece boundary. Depending on the welding conditions, the initial and terminal periods may be of different durations, and may or may not practically influence the joint produced. As the welding tool arrives at the end of the joint, forward motion of the tool is typically stopped and the tool is withdrawn from the workpiece, leaving a keyhole at the end of the weld. Alternatively, the tool can be run out the end of the workpiece, producing a tear-out. In either case, the end of the weld is generally not usable and must be trimmed away by sawing or machining. For this reason, it is common to use run-on and run-off tabs at the ends of the joint to facilitate trimming the unusable material.
2.2
Comparison of friction stir welding (FSW) to other welding processes
Comparison of FSW to other welding processes is typically done within the context of justifying the use of the process over other, more conventional
20
Friction stir welding
techniques. Successful application of FSW depends upon a clear understanding of the characteristics of the process, so favorable technical and economic justification can be developed. The details of each welding application will dictate the basis for the justification, so in this section the many considerations for technical and economic justification of FSW implementation will be described.
2.2.1 Technical justification of FSW The unique, favorable characteristics of FSW compared to traditional arc welding methods provide several sources for technical justification for use of the process. One key for successful implementation of the process is a clear understanding of the technical improvements offered by FSW, balanced against the requirements of the process. The main points for technical justification of FSW compared to arc welding processes are: ∑ improved weldability ∑ reduced distortion ∑ reduced residual stress, improved fatigue, corrosion, and stress corrosion cracking performance ∑ improved cosmetic appearance ∑ elimination of undermatched filler metal ∑ improved static strength and ductility ∑ mechanized process ∑ high robustness, few process variables. Each of these technical justifications must be balanced against the unique technical requirements of the process, such as: ∑ mechanized process ∑ special fixture requirements ∑ joint design limitations ∑ keyhole at end of weld. Each of these technical points will be discussed briefly, followed by a discussion of the economic justification issues in the following section. Improved weldability Since FSW is a solid state process, weldability in certain materials can be improved. This is especially the case in certain aluminum alloys. Some aluminum alloys or material forms, such as castings, are difficult or impossible to weld by traditional arc welding processes due to problems with the formation of brittle phases and cracking. For these alloys, weldability alone may be
The friction stir welding process: an overview
21
sufficient to form a justification for the use of FSW over conventional arc welding or other joining techniques, such as mechanical fasteners. Further, FSW makes possible the joining of some dissimilar alloys, which can be of significant benefit in certain applications. Reduced distortion The reduced peak temperature reached in FSW compared to arc welding processes also generally leads to reduced longitudinal and transverse distortion, although FSW weldments are certainly not free of residual stress. The balance of residual stress in FSW can result in essentially flat weldments in materials of virtually any practically weldable thickness, although this is affected by welding tool design, joint design, welding parameters and fixture design. This characteristic of FSW makes possible new methods of construction and can significantly affect the total cost of manufacturing an assembly through reduced fit-up problems and secondary machining operations. The earliest production applications for FSW were based on its use for making products from extruded aluminum shapes, as shown in Table 2.1. The business model for these applications was based on the use of FSW to produce very flat, integrally stiffened assemblies from relatively narrow aluminum extrusions. In this case, FSW was justified based on the fact that had arc welding processes been used, the weldments would have been so distorted as to render them unserviceable.
Table 2.1 Chronology of production applications for FSW, through 2004 Year
Application
1995 *Hollow heat exchangers 1996 *Commercial shipbuilding 1998 Delta II rockets 1999 *Commercial shipbuilding 2000 *Automotive components 2000 Laser system housings 2001 *Motor housings 2001 *Automotive components 2001 *Train bodies 2002 *Automotive components 2003 Aircraft structure 2003 *Commercial shipbuilding 2004 Space shuttle external tanks 2004 Food trays *Denotes welding of extrusions
Company Marine Aluminum, Norway Marine Aluminum, Norway Boeing, US Sapa, Sweden Sapa, Sweden General Tool, US Hydro Aluminum (formerly Marine Aluminum), Norway Showa, Japan Hitachi, Japan Tower Automotive, US Eclipse, US Advanced Joining Technologies, US Lockheed Martin, US RIFTEC, Germany
22
Friction stir welding
Improved fatigue, corrosion, and stress corrosion cracking performance The reduced maximum temperature and residual stress can also lead to improved performance under cyclic loading conditions [1–3]. Typically, joints produced by FSW have fatigue strength that is higher than arc welded fatigue strength, but below base metal strength. FSW joints that are machined after welding have been shown to approach base metal fatigue strength. FSW has been justified on the basis of fatigue performance, in concert with economic factors, over the use of mechanical fasteners, such as in its use on the Eclipse aircraft [4] and in other aerospace applications [1, 3, 5, 6]. The reduced residual stress and peak temperature can also improve general corrosion and stress corrosion cracking problems in aluminum alloys. Although this is not generally sufficient as a driver for justification for the use of the process, it could prove to be an enabling technology in some long service life applications. Improved cosmetic appearance The root side of conventional friction stir welds has been shown to be extremely smooth and flat in a variety of materials and thicknesses. After painting, the root side of the joint can be virtually invisible. This has played a role in justification of the use of the process over other joining processes in commercial shipbuilding, in aircraft manufacture, and in the production of food trays. Elimination of under matched filler metal In some materials there are no available filler metals for arc welding that match or exceed the strength of the surrounding base metal. However, FSW is an autogenous welding process, obviating the need for filler metals. In these materials it has been demonstrated that improved joint strength and/or ductility can be achieved with FSW over arc welding processes. The elimination of filler metal in FSW also leads to cost avoidance by eliminating the need for a wire feeding system and improved joint consistency by eliminating problems with wire feeding that can occur in arc welding processes. Improved static strength and ductility Even in cases where adequate filler metals are available, the higher temperature reached and limited material deposition rates in arc welding can degrade the HAZ sufficiently to reduce the joint strength compared to FSW. It is often the case in thin section aluminum alloys that the joint strength in arc welding and FSW are comparable. However, in thick materials, up to 75 mm
The friction stir welding process: an overview
23
thick, the fact that FSW can be accomplished in a single pass can result in significantly improved joint strength and ductility. In some applications, this may be sufficient to justify the use of FSW over arc welding and mechanical fastening. Mechanized process FSW is typically operated as a completely mechanized process, as is further described below. While this can lead to increased capital cost compared to arc welding and mechanical fastening techniques, the mechanized nature of the process leads to improved joint path control and more consistent quality than manually operated processes. The improvement in joint quality alone may be sufficient for justification of the use of FSW, especially in applications that demand very high joint quality, such as in the construction of launch vehicles and other aerospace structures [7]. While FSW offers many technical bases for justification of its use, there are also unique technical requirements that may offset its technical benefits to some degree. Some of the main technical requirements that must be considered are described below. Mechanized process The fact that FSW is a mechanized process was described above as a positive factor in the justification of the process for production use. At the same time, the process forces generated during FSW are typically too high to permit hand operation. In the case of very thin materials, the forces may be sufficiently low to permit manual operation, but typically the welding tool is so small in these cases that mechanical means are needed to maintain accurate tool path control to consume the joint. Certainly, in thick materials FSW can only be operated as a mechanized process. For example, for 25 mm thick 5083 aluminum plate, it may be necessary to apply a force of 44 kN along the tool’s rotation axis to keep the tool embedded in the workpiece, while simultaneously pushing the tool laterally with 15 kN force and applying a torque of about 360 N-m [8]. The forces for thinner sections are understandably lower; for example for 2.8 mm thick 2198 aluminum, forces of about 8 kN along the tool axis, 650 N in the direction of welding, and about 500 N transverse to the welding direction have been observed [9]. Although the forces can be quite low for thin section aluminum workpieces, the welding tools are also very small, requiring precise joint tracking. As a result, relatively expensive, custom made equipment is often necessary for production use of FSW. Although new, lower cost equipment is currently being developed for some applications, the high capital cost is often a significant barrier to developing
24
Friction stir welding
the business case for the use of FSW, often making it necessary to develop the application as a nearly continuous production operation, maximizing the economic value of goods produced by the process. Special fixture requirements As mentioned earlier, FSW requires that the workpiece be rigidly held in position during welding, to ensure that the joint does not separate under the force of the welding tool and to ensure that the workpiece stays in intimate contact with the anvil during welding, thus achieving a smooth weld. Although the special fixture requirements of FSW do impose an economic burden on the justification of the use of the process, which will be discussed below, fixture requirements also place a practical restriction on the size of workpiece that can be produced. For example, the requirement to restrain the workpiece against the anvil may make it difficult to secure very large and thin workpieces, or the requirement to restrain lateral separation of the joint can be difficult for very thick workpieces. These are surmountable requirements, but ones which must be considered. Joint design limitations Since FSW is an autogenous process, it is impossible to make what is generally considered a fillet weld, where a significant amount of material is added to fill a transition between two workpieces. Although it is possible to form a small fillet during FSW of plates at some angle, this is usually achieved at the expense of material from the joint. Typically, FSW is used to produce butt welds, corner welds, and lap welds, as shown in Fig. 2.4. Weld keyhole FSW is a keyhole welding process, as described previously. As a result, in some applications it is necessary to consider how the welded joint will be started and terminated to result in a serviceable assembly, such as in the construction of cryogenic fuel tanks and in welding marine structures. Typically, the start and stop ends are cut away from the main portion of the assembly and discarded. Alternatively, run-on/run-off tabs may be used to reduce the loss of base metal. For structures such as sealed tanks, it may be possible to use friction tapered plug welding, arc welding, or even a sealed fastener to eliminate the keyhole. Finally, a welding tool with a retractable pin has been used in the past to allow variable weld penetration, making it possible to gradually retract the pin over a length of weld, eliminating the keyhole.
The friction stir welding process: an overview
A. Straight butt joint
25
B. Lap joint
D. T joint C. 90° corner joint
E. Oblique angle joint
2.4 Common FSW joint configurations.
2.2.2 Economic justification The positive and negative aspects of economic justification of the use of FSW over more traditional methods are briefly discussed in this section. Very few published studies have been produced that detail the economic justification of FSW, presumably because every application is unique in its justification and it is difficult to develop a general cost model. In spite of this, some publications have addressed the economic justification of FSW by presenting case studies [10–12]. The main factors in economic justification for FSW are: ∑ processing time/labor ∑ licensing ∑ capital investment ∑ production volume ∑ robustness/low defect rate. Other factors play a lesser role in the economic impact of FSW. These secondary factors include the elimination or reduction of conventional welding
26
Friction stir welding
consumables, such as shielding gas and filler metal, the increased cost of friction stir welding tool, itself a consumable item, the reduced production of fumes, and the reduced energy consumption. In this section the main economic factors in the justification of FSW are described. Processing time/labor There are a number of factors that contribute to the reduced processing time and direct labor that can be expected from the use of FSW. FSW can often be accomplished in a single pass compared to the multiple passes often required for arc welding processes, especially as the workpiece thickness increases. Each pass in arc welding can require interpass cooling time, cleaning, inspection, and repair, and may require rotation of the workpiece to allow welding from the opposite side. Further, the superior cosmetic appearance of friction stir welds can lead to reduced manufacturing costs associated with post-weld processing. Licensing At the time of this writing, FSW is a patented process in many countries [13], and a license fee must generally be paid for its use. The specific terms are restricted from publication by agreement with TWI, the owner of the intellectual property rights to the process. That said, the license fee for the use of the basic process is a significant factor in the economic justification of the use of the process, and should be considered until the time of expiration of the patents. Capital investment FSW is a fully mechanized process, so naturally the process is limited by the capabilities of the equipment used. This includes the welding system that rotates the welding tool and traverses it along the joint, and the welding fixture used to restrain the workpiece during welding. Owing to the high forces generated by the process, the equipment and fixture can represent a significant cost, especially for large or thick section workpieces. Some welding system providers have recognized that the capital investment required for FSW is prohibitive for many applications, and are developing low cost approaches for FSW equipment. As an alternative, there are several FSW service providers that can produce welded products without imposing the burden of capital investment on a single production application.
The friction stir welding process: an overview
27
Production volume Production volume is a factor in the economic justification of FSW in the way that it amplifies savings from labor and processing time, and distributes fixed costs from licensing and capital investment. In some applications, the production volume may make traditional arc welding processes impractical simply due to limitations in the available skilled workforce. The combined effect of technical and economic justification of the process is essential to the successful implementation of FSW in the production environment. This justification has been successfully demonstrated in several industries, as evidenced by the list of early production applications presented above in Table 2.1. The expiration of the FSW patents and the development of lower cost FSW equipment and fixtures will, over time, ease the fixed costs associated with implementation of the process and will improve the business case even further.
2.3
Welding tools used for friction stir welding (FSW)
Many of the advances made in friction stir welding have been enabled by the development of new welding tools. The welding tool design, including both its geometry and the material from which it is made, is critical to the successful use of the process. Welding tool geometry development led to the first sound welds made in aluminum alloys, and this field of study has led to higher weld production speeds, higher workpiece thickness, improved joint properties, new materials, and new welding equipment. Welding tool material development has enabled welding of high melting point materials, such as titanium, steel, and copper, and has improved productivity in aluminum welding. These aspects of friction stir welding are covered in detail in Chapter 4. The original welding tool geometry breakthrough developed by The Welding Institute was the discovery that adding a screw thread profile on the welding tool pin prevented the formation of a void near the end of the pin. As the welding tool pin is forced laterally during the welding process, the natural tendency is for workpiece material to be expelled upward and out past the welding tool shoulder. Even a relatively minor loss of material past the shoulder results in a void in the weld zone, since no material is added during welding. By including a thread on the pin, workpiece material could be forced downward, thus establishing a circulation of workpiece material and preventing void formation. Since that time, new welding tool features have been developed with, for example, the goal of reducing process forces, increasing the robustness
28
Friction stir welding
of the process, or simplifying welding control. Different features are used by different practitioners of FSW, depending on the materials being welded and the process performance goals required. For example, some researchers and FSW production operations still use the original FSW tool design demonstrated by The Welding Institute in the early 1990s. FSW practitioners needing to weld at higher travel speeds or with deeper weld penetration may adopt variations to the original tool design. As a result, there is no accepted “optimum” tool design in use today. Tool steel materials are generally acceptable for the FSW of aluminum alloys. However, much like the situation today with welding tool geometry, even for welding aluminum alloys there is no accepted standard tool material. For applications where aluminum alloys from 6 to 12mm thickness are welded, H13 tool steel is generally adequate. For such applications it is also often possible to use a one-piece welding tool design. However, if high productivity is needed or it is necessary to weld thicker aluminum materials, a more elaborate tool design and material selection may be required. In such cases, the pin might be made from a material that has higher strength at the temperature of welding, such as MP159, while the shoulder might still be made from H13. For applications that require welding other materials, such as titanium, steel, and copper, welding tools might be made from tungstenbased materials, from polycrystalline cubic boron nitride, or from any number of other materials that offer high performance at high temperatures. Welding tool design is an area of active research at this time. As new welding tool materials and geometries are developed, it is likely that the scope of applicability of FSW will continue to expand, as it has done since its inception. As a result, it will be necessary to review the latest literature on this subject at any given time to learn of the current state of the art in FSW and its best mode of application for a given application.
2.4
Parameter effects
While the general principles of the effect of process variables on the friction stir welding process have much in common with other welding processes, the details are completely different, as one might expect. The main process variables in friction stir welding are listed in Table 2.2. These variables all act to determine the outcome of the welding process. The main interest in studying the effect of the process variables lies in understanding the effect of the process on joint properties, including static mechanical properties, fatigue strength, corrosion properties, stress corrosion cracking resistance, and toughness, with the goal of maximizing productivity, performance, and reproducibility. The welding process affects these joint properties primarily through heat generation and dissipation, so primary attention should be given to the effect of the welding process variables on
The friction stir welding process: an overview
29
Table 2.2 Main FSW process variables Tool design variables
Machine variables
Shoulder and pin materials Welding speed Shoulder diameter Spindle speed Pin diameter Plunge force or depth Pin length Tool tilt angle Thread pitch Feature geometry
Other variables Anvil material Anvil size Workpiece size Workpiece properties
heat generation and related outcomes. Other areas of study include the effect of process parameters on material flow, defect formation, process forces, grain size, etc. These effects are explored in detail in chapters that follow. Numerous experimental and computed model studies into the effect of process parameters on heat generation have been published since the process was first introduced [8, 14–29]. This work has demonstrated that, while FSW may seem at first to be a simple process, the fact that the process is a fully coupled thermomechanical process can lead to counter-intuitive relationships between variables and in the resulting material flow. In arc welding the heat input is determined entirely by controllable machine parameters (voltage and current), while in FSW the temperature-dependent workpiece properties are important in determining the heat input, which in turn affects the workpiece conditions. As a result, seemingly unrelated variables can participate in the heat generation process. For example, the anvil size and material can have a dramatic effect on process forces and heat generation in FSW, based on their role in thermally preconditioning the workpiece material in advance of the welding tool. This self-referential relationship contributes to the stability of the welding process, but at the same time it complicates the determination of the effect of process variables on heat generation. Further, the fact that FSW is a fully coupled thermomechanical process means that one must carefully consider the unintended effect on interconnected variables when making a change to the process, such as when moving a process developed on one piece of equipment to a new welding system. Chapter 3 fully describes the effect of process variables on heat generation and material flow.
2.5
Materials used with friction stir welding (FSW)
Friction stir welding has been shown to be effective in joining a number of different materials. Although aluminum was the first material to be friction stir welded, over time it has been shown that many of the benefits demonstrated for aluminum can also be seen in welding other materials, such as steel, titanium, copper, magnesium, and lead. However, welding of high melting point materials is made more difficult by the harsh operating conditions
30
Friction stir welding
placed on the welding tool material. As a result, while many materials can be friction stir welded, performance and economic justification must be developed in order to make practical use of the process. No studies have been performed to define the specific characteristics that a material must have in order to be weldable by FSW, but examination of the characteristics of FSW in aluminum suggests some very general requirements for welding of other materials. For example, it is clear from FSW of aluminum that thermal softening of the workpiece material is necessary for the welding process to commence, so it is reasonable to expect that in other materials the welding process will take place at a temperature that is near the melting point of the workpiece material. This also implies that it is necessary that heat be generated with sufficient intensity to overcome the loss of heat from the welding zone through conduction into the workpiece. Another consideration is the need to achieve heat generation, either by friction, plastic work, or by auxiliary heating, at the full spectrum of temperatures from the initial material temperature up to the welding temperature. Although it is not normally needed for FSW of aluminum, shielding gas may be needed for some materials to prevent reactions with atmospheric gasses. In high melting point metals it appears that the main limitation to weldability is the availability of suitable welding tool materials. The development of new welding tool materials and geometries has made it possible to join materials such as steel and titanium in the laboratory environment and in a limited number of production applications [30]. In FSW of steel it has been shown that the lower welding temperature can lead to very low distortion and unique joint properties [31–36]. FSW of steel is an area of active research, so it is reasonable to expect other production applications to emerge over time. A very attractive application is FSW of steel plate for shipbuilding applications, based primarily on the reduction of welding distortion, but the development of low-cost welding equipment and more robust welding tool materials is required before this application can be exploited. Titanium alloys are generally considered to be weldable by fusion welding processes, but FSW is of interest as a joining method for some specific titanium-based alloys, for some product forms, such as castings, and for some alloy combinations which are more difficult to fusion weld [37–39]. Friction stir welding of titanium has been demonstrated in the laboratory environment and it has been used in the construction of relatively large prototype structures. Although titanium is considered a high melting point material, its low thermal conductivity makes it necessary to reduce the heat input of the tool design, either by minimizing the shoulder diameter or by eliminating shoulder rotation altogether [40, 41]. Friction stir welding of titanium is presently an area of active research, which may ultimately lead to production applications. Friction stir welding of copper has been developed for several years
The friction stir welding process: an overview
31
for the construction of canisters for storing nuclear waste [42–44]. In this application, which requires very low defect rates, FSW was selected over electron beam welding for welding 50 mm thick circumferential joints for attaching canister lids. Although it was initially thought that FSW of thick copper would be adversely affected by the high thermal conductivity, it has been found that by using relatively high spindle speed, sufficient heat intensity can be developed to accomplish sound, high quality welds. Several other materials have been joined using FSW in the laboratory environment, and the evaluation of weldability in new materials is an area of ongoing research [45, 46]. Also, some work has been done recently to demonstrate the use of FSW to join dissimilar materials, such as aluminum to steel [47, 48]. Certainly, it is conceivable that titanium could be welded to aluminum using a similar approach. The development of technical or economic justification is required for moving the FSW of these materials to the production floor.
2.6
Joint geometries
A variety of joint geometries are possible with FSW; however, there are certain limitations and requirements that are unique to the process. Since FSW is an autogenous welding process, some types of joint, such as fillet welds, are fundamentally impossible with FSW, although fillet welds can be simulated by special material or fixture designs, as will be discussed below. A summary of common joint geometries is presented in Fig. 2.5. The figure also shows typical anvil sections for each joint type, since anvil and fixture requirements are closely related to joint design. It is assumed in these diagrams that the anvil sections are rigidly mounted relative to the workpieces. Certainly, other joint geometries are possible. In each of these joint designs and fixture arrangements, it is necessary to: ∑ provide sufficient area for the welding tool shoulder path ∑ provide sufficient containment of softened weld metal ∑ provide sufficient force to prevent motion of the workpieces ∑ provide adequate heat sink to dissipate the heat of welding. The area required for the welding tool shoulder is a function of material thickness and alloy. Typically, for aluminum alloys, the area required for the shoulder is about three to five times the material thickness. Steel and titanium typically would require less shoulder area, since these materials have lower thermal conductivity and therefore require a smaller shoulder diameter. Containment of softened weld metal is necessary along the full length of FSW joints. Machined details, such as drilled holes or pockets, that are very close to a weld joint should be avoided or temporarily plugged during
32
Friction stir welding
A. Straight butt joint
B. Lap joint C. 90° corner joints
D. T joints
E. Oblique angle F. Butt joint with mismatch
2.5 Summary of common joint designs with anvil sections and fixture force requirements.
welding to provide additional heat sink and to prevent softened material from pushing out. One exception to this rule is the case where an anvil is specifically designed with rounded corners to produce a small fillet, as shown in Fig. 2.6. Here, the fillet is kept small so that the workpiece material that pushes into the void does not result in internal voids in the weld. This corner radius acts to reduce the stress concentration in the corner [49]. Another example of an exception to the requirement for softened metal containment is the case where a small amount of clad material is removed from the face of the work
The friction stir welding process: an overview
33
2.6 Anvil corner for producing small fillets. Workpieces
Before welding
Cladding removed, leaving small gap Anvil
Clad layer
After welding
Stir zone
2.7 Removal of cladding on root side of a joint.
adjacent to the joint on the root side in order to prevent the cladding from being ingested into the weld region, as shown in Fig. 2.7. If the material is prepared properly, the force of the welding tool causes workpiece material to extrude into the narrow gap, resulting in a smooth root surface. As indicated in Fig. 2.5, it is necessary to forcibly constrain the workpiece components from motion under the process forces. Test data as to the magnitude of the force required is almost non-existent, so experience and
34
Friction stir welding
trial-and-error approaches are necessary for this aspect of fixture design. An example of the negative effect of inadequate fixture force is “drop-out” that is the result of inadequate vertical force in a butt weld, preventing the workpiece from lifting from the anvil, as shown in Fig. 2.8. This commonly occurs at the start of welds, where the process of plunging the welding tool pin into the workpiece tends to bulge the back of the workpiece and lift the surrounding material off the anvil. Once this has occurred it is very difficult to get the workpiece back into contact with the anvil, since forward motion of the tool tends to push the drop-out in advance of the pin. As a result, drop-out is much better prevented in advance by good fixture design, rather than trying to resolve it during a weld. Joint design also plays a role in establishing an adequate heat sink for a stable welding process. An example of a poor joint design from the standpoint of heat dissipation is shown in Fig. 2.9. In this case, one side of the corner joint is very small, which would result in excessive heat buildup during welding, possibly making it impossible to weld. A better joint design would have left more material on the right side of the joint, to be machined away after welding. A few additional guidelines relative to joint design are necessary. First, it is generally desirable to consume the original faying surfaces to the greatest extent possible, since remnant oxide bands can provide crack initiation sites. This may require multiple weld passes, such as in the case of a severe oblique angle joint, or it may require special welding tool designs to generate adequate Workpiece lifting
Drop-out
Longitudinal view Example of drop-out on right side of weld, 6mm 2094 aluminum FSW, transverse view
2.8 Drop-out in a butt weld produced by inadequate vertical holding force on the workpiece.
The friction stir welding process: an overview
35
Poor heat sink
Better heat sink
2.9 A 90° corner joint with inadequate heat sink.
coverage to consume the joint. Secondly, lap joints can be problematic, since they inherently involve leaving a remnant oxide band that enters the weld. Careful consideration must be given to welding procedure and applied load path in this type of joint in order to have reliable results. Finally, it is important to consider the effect of the joint properties across the weld zone and how it relates to the applied load in service when formulating a joint design. For example, in some applications it may be desirable to replace the oblique angle joint shown in Fig. 2.5 with a machined angle piece and two butt joints, in order to move the welded joints away from the corner region, which may be the area with the highest loading. Friction stir welding is a relatively new welding process, but it has rapidly developed into a viable production process in a number of industries. This has been accomplished through intense research in universities and research institutions around the world, contributing to our understanding of the process and expanding its capabilities through technical innovation. The chapters that follow provide a more detailed description of the important aspects of the process, required in developing a complete understanding of its unique characteristics.
2.7
Joint preparation
An important consideration in the design of a successful welding process is the issue of joint preparation. In friction stir welding, there are a variety of joint preparation methods that are commonly employed, depending on the performance requirements and economics of the application. This section will briefly address the issues related to joint preparation, in terms of economics, performance, and cosmetic appearance. It is possible to weld aluminum alloys using FSW with absolutely no joint preparation, using roughly cut edges. However, this will entrap aluminum
36
Friction stir welding
oxide and other contaminants in the weld zone, and could possibly produce volumetric defects or poorly consolidated materials in the weld if the roughness of the surface results in significant gap in the faying surfaces. The negative consequences of surface contaminants include poor fatigue loading performance, localized low ductility, and volumetric defects produced during post-weld heating, such as during solution heat treatment or arc welding. These negative results can be avoided by machining of all surfaces, followed by etching and solvent cleaning, but this can dramatically impact the cost of the process. As a result, it is important to consider the application when deciding how to prepare the joint region prior to welding. It is common when welding extruded 6xxx aluminum alloys to weld with the as-extruded oxide intact, with only minimal cleaning of the parts. In heat treatable alloys, natural aging of the center of the stirred region tends to protect any residual oxide band from high strain levels, since strain is commonly localized in the HAZ regions, thus preventing fracture from originating at a residual oxide band. This would be considered typical for many products that are subjected to essentially static loading. However, for welding of extruded 5xxx aluminum alloys, it is often necessary to remove the oxide on the faying surfaces, since this alloy is not heat treatable and natural aging is not available to protect the stirred region at the center of the joint. In either case, it is important to subject test materials to appropriate mechanical tests to ensure adequate performance of joints produced using the intended joint preparation procedures. When welding materials produced from plate or sheet, the joint is commonly machined along the faying surfaces. In alloys that have a heavy mill scale on the exterior surfaces, it may also be necessary to abrade away the exterior surface oxide adjacent to the joint prior to welding. Heavy surface oxide can result in excessive wear in welding tool components that come into contact with it during welding. Machining and grinding operations are commonly followed by solvent wiping to ensure that residual machining coolant is removed prior to welding. This is the most common method used in preparing non-extruded materials for FSW today. For fatigue or high-deformation applications, it is generally desirable to be very careful about joint preparation procedures. This is also the case when the FSW joint will be subsequently crossed by an arc weld, such as when stir welded components are subsequently arc welded for final assembly, and for lap welded joints, where residual oxide may represent an initial crack that can propagate into the weld. In addition to the machining/grinding steps described above, these components may be subject to a time limit after joint preparation until welding occurs, and may be required to have localized etching prior to welding. The goal is to reduce surface oxides to the thinnest possible condition prior to welding.
The friction stir welding process: an overview
2.8
37
Post-weld heat treating
There are a number of issues related to the heat treating of welded joints, much of which is beyond the scope of this text. Some aspects of this topic are, however, unique to FSW, which will be discussed briefly here. An important issue is the use of post-weld aging to improve static, corrosion, and stress corrosion cracking performance of joints, particularly in aluminum alloys. At the completion of a stir welding pass, the weld zone contains a variety of material conditions, depending on proximity to the welding tool path, which produces the temperature history that the material experiences. For example, material in the stirred region at the center of the weld often has characteristics of material that has been solution heat treated, bounded on either side by materials that are over-aged, with coarsened precipitates or precipitate-free zones, surrounded by materials that are unaffected by the welding process, only because the temperature did not reach a high enough level to alter the local temper [50, 51]. A number of researchers have shown that the static strength, corrosion, and stress corrosion cracking of stir welded materials can be improved by post-weld aging. However, one must consider the impact of the post-weld aging treatment on the properties of material away from the weld zone, so it is typically found that the “optimum” solution is to weld plates that are in an under-aged condition, followed by post-weld aging to bring the welded materials to a state that offers good corrosion performance, for example, while leaving the balance of the component with adequate mechanical properties [52, 53]. Some aluminum alloys, designed to have a particularly strong aging response, can be welded and given optimal strength by closely following the welding tool with a water quench. The goal here is to entrap the maximum amount of solute material in solution after welding, leaving more solute available for post-weld aging, resulting in an increase in joint strength. By using this approach, static strength in some heat-treatable alloys can approach the strength of the base metal. This technique is possible with FSW of aluminum because it is not necessary to use shielding gas, and the welding process is not adversely affected by some amount of water overspray. However, it is important that the welding process be developed with the water quench applied in a systematic way, so unnecessary variability is not introduced. Another issue in considering the use of post-weld aging relates to the impact of the initial temper on the welding process itself. Since FSW is a fully coupled thermomechanical process, the properties of the workpiece material as it reaches the welding tool may affect the process forces and heat input generated. If the travel speed is very slow, the base metal initial temper may not significantly affect the spindle torque, since there is time for alteration of the initial temper prior to coming into contact with the welding tool. However, when welding thin materials at high travel speed, it is likely
38
Friction stir welding
that the initial temper will influence heat generation, friction coefficient, and contact forces. As a result, the effect of changes to the base metal temper, designed to facilitate post-weld aging response, should be confirmed prior to production welding, which may indicate changes in welding parameters to offset the changes in initial temper.
2.9
References
1. Pacchione, M. and Lohwasser, D., “Friction stir welding application to aircraft primary structures,” Aircraft Design Principles PR0407463, Airbus, 2004. 2. Kumagai, M. and Tanaka, S., “Properties of aluminum wide panels by friction stir welding,” 1st International Symposium on Friction Stir Welding, Thousand Oaks, CA, USA, June 14–16, 1999. 3. Talwar, R., Bolser, D., Lederich, R.J. and Baumann, J., “Friction stir welding of airframe structures,” 2nd International FSW Symposium, Gothenburg, Sweden, June 26–28, 2000. 4. Christner, B., McCoury, J. and Higgins, S., “Development and testing of friction stir welding as a joining method for primary aircraft structure,” 4th International FSW Symposium, Park City, UT, USA, May 14–16, 2003. 5. Lohwasser, D., “Application of friction stir welding for aircraft industry,” 2nd International FSW Symposium, Gothenburg, Sweden, June 26–28, 2000. 6. Shepherd, G.E., “The potential for using solid phase welding to repair cracks that may occur on thin aluminum aircraft wing structures,” 2nd International FSW Symposium, Gothenburg, Sweden, June 26–28, 2000. 7. Jones, C. and Adams, G., “Assembly of a full scale external tank barrel section using friction stir welding,” 1st International Symposium on Friction Stir Welding, Thousand Oaks, CA, USA, June 14–16, 1999. 8. Colligan, K.J., Xu, J. and Pickens, J.R., “Welding tool and process parameter effects in friction stir welding of aluminum alloys,” Friction Stir Welding and Processing II, Jata, K.V., Mahoney, M.W., Lienert, T.J. and Mishra, R.S., editors, TMS, 181–190, 2003. 9. Lohwasser, D., “Friction Stir Welding in A350,” EADS Research Workshop, AIRBUS, November 2005. 10. Mononen, J.T., “Cost comparison of FSW and MIG welded aluminum panels,” 3rd International FSW Symposium, Kobe, Japan, September 27–28, 2001. 11. Kallee, S.W. and Mistry, A., “Friction stir welding in the automotive body in white production,” 1st International Symposium on Friction Stir Welding, Thousand Oaks, CA, USA, June 14–16, 1999. 12. Midling, O.T., Kvale, J.S., and Dahl, O., “Industrialization of the friction stir welding technology in panels production for the maritime sector,” 1st International Symposium on Friction Stir Welding, Thousand Oaks, CA, USA, June 14-16, 1999. 13. Thomas, W.M., Nicholas, E.D., Needham, J.C., Murch, M.G., Temple-Smith, P. and Dawes, C.J., “Improvements related to friction welding,” PCT Patent Application No. PCT/GB92/02230, June 10, 1993. 14. Reynolds, A.P. and Tang, W., “Alloy, tool geometry, and process parameter effects on friction stir welding energies and resultant FSW joint properties,” Friction Stir Welding and Processing, Jata, K.V., Mahoney, M.W., Mishra, R.S., Semiatin, S.L. and Field, D.P., editors, TMS, 15–23, 2001.
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15. Record, J.H., Covington, J.L., Nelson, T.W., Sorensen, C.D. and Webb, B.W., “Fundamental characterization of friction stir welding,” 5th International FSW Symposium, Metz, France, September 14–16, 2004. 16. Colligan, K.J., “Relationships between process variables related to heat generation in friction stir welding of aluminum,” Friction Stir Welding and Processing IV, Mishra, R.S., Mahoney, M.W., Lienert, T.J. and Jata, K.V., eds., TMS, 39–54, 2007. 17. Colligan, K.J. and Mishra, R.S., “A conceptual model for the process variables related to heat generation in friction stir welding of aluminum,” Scripta Materialia, 58, 327–331, 2008. 18. Colligan, K.J., “A proposed conceptual model of the process variables related to heat generation in FSW of aluminum,” 7th International FSW Symposium, Awaji Island, Japan, May 20–22, 2008. 19. Zettler, R., Lomolino, S., dos Santos, J., Donath, T., Beckmann, F., Lippman, T. and Lohwasser, D., “Effect of tool geometry and process parameters on material flow in FSW of an AA 2024-T351 alloy, 5th International FSW Symposium, Metz, France, September 14–16, 2004. 20. Chen, Z.W., Pasang, T. and Qi, Y., “Shear flow and formation of nugget zone during friction stir welding of aluminum alloy 5083-O,” Materials Science and Engineering A, 474, 312–316, 2008. 21. Colligan, K.J., “Material flow behavior during friction stir welding of aluminum,” Welding Journal, 229–237, July, 1999. 22. Colligan, K.J., “Material flow behavior during friction stir welding of aluminum,” 1st International Symposium on Friction Stir Welding, Thousand Oaks, CA, USA, June 14–16, 1999. 23. Colligan, K.J. and Chopra, S.K., “Examination of material flow in thick section friction stir welding of aluminum by a stop-action technique,” 5th International FSW Symposium, Metz, France, September 14–16, 2004. 24. Bendzsak, G.J., North, T.H. and Smith, C.B., “An experimentally validated 3D model for friction stir welding,” 2nd International FSW Symposium, Gothenburg, Sweden, June 26–28, 2000. 25. Colegrove, P., Painter, M., Graham, D. and Miller, T., “3-dimensional flow and thermal modeling of the friction stir welding process,” 2nd International FSW Symposium, Gothenburg, Sweden, June 26–28, 2000. 26. Xu, S., Deng, X., Reynolds, A.P. and Seidel, T.U., “Finite element simulation of material flow in friction stir welding,” Science and Technology of Welding and Joining, Vol. 6, No. 3, 191–193, 2001. 27. Williams, S.W. et al., “Integrated modeling of the FSW process,” 6th International FSW Symposium, Saint Sauveur, Canada, October 10–13, 2006. 28. Askari, A., Silling, S., London, B. and Mahoney, M., “Modeling and analysis of friction stir welding processes,”, Friction Stir Welding and Processing, Jata, K.V., Mahoney, M.W., Mishra, R.S., Semiatin, S.L. and Field, D.P., editors, TMS, 43–54, 2001. 29. Frigaard, O., Grong, O., Bjorneklett, B. and Midling, O.T., “Modeling of the thermal and microstructural fields during friction stir welding of aluminum alloys,” 1st International Symposium on Friction Stir Welding, Thousand Oaks, CA, USA, June 14–16, 1999. 30. Nelson, T., Sorensen, C., Packer, S. and Allen, C., “Qualification of friction stir processing for production applications,” 7th International FSW Symposium, Awaji Island, Japan, May 20–22, 2008.
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Friction stir welding
31. Thomas, W.M., Threadgill, P.L. and Nicholas, E.D., “Feasibility of friction stir welding steel,” Science and Technology of Welding and Joining, Vol. 4, No. 6, 365–372, 1999. 32. Lienert, T.J. and Gould, J.E., “Friction stir welding of mild steel,” 1st International Symposium on Friction Stir Welding, Thousand Oaks, CA, USA, June 14–16, 1999. 33. Posada, M., DeLoach, J., Reynolds, A.P., Skinner, M. and Halpin, J.P., “Friction stir weld evaluation of DH-36 and stainless steel weldments,” Friction Stir Welding and Processing, Jata, K.V., Mahoney, M.W., Mishra, R.S., Semiatin, S.L. and Field, D.P., editors, TMS, 43–54, 2001. 34. Konkol, P., Mathers, J.A., Johnson, R. and Pickens, J.R., “Friction stir welding of HSLA-65 steel for shipbuilding,” 3rd International FSW Symposium, Kobe, Japan, September 27–28, 2001. 35. Sorensen, C.D., Nelson, T.W., Packer, S.M. and Steel, R.J., “Innovative technology applications in FSW of high softening temperature materials,” 5th International FSW Symposium, Metz, France, September 14–16, 2004. 36. Mahoney, M., Steel, R., Nelson, T., Packer, S. and Sorensen, C., “FSW of HSLA65 steel with low/no distortion,” 7th International FSW Symposium, Awaji Island, Japan, May 20–22, 2008. 37. Juhas, M.C., Viswanathan, G.B. and Fraser, H.L., “Microstructural evolution in Ti alloy friction stir welds,” 2nd International FSW Symposium, Gothenburg, Sweden, June 26–28, 2000. 38. Trapp, T., Helder, E. and Subramanian, P.R., “FSW of titanium alloys for aircraft engine components,” Friction Stir Welding and Processing II, Jata, K.V., Mahoney, M.W., Lienert, T.J. and Mishra, R.S., editors, TMS, 173–178, 2003. 39. Jones, R.E. and Loftus, Z., “FSW of 5mm Ti 6Al 4V,” 6th International FSW Symposium, Saint Sauveur, Canada, October 10–13, 2006. 40. Russell, M.J. and Blignault, C., “Recent developments in FSW of Ti alloys,” 6th International FSW Symposium, Saint Sauveur, Canada, October 10–13, 2006. 41. Russell, M.J., Threadgill, P.L. and Horrex, N.L., “Recent developments in the stationary shoulder FSW of Ti alloys,” 7th International FSW Symposium, Awaji Island, Japan, May 20–22, 2008. 42. Andersson, C-G. and Andrews, R.E., “Fabrication of containment canisters for nuclear waste by friction stir welding,” 1st International Symposium on Friction Stir Welding, Thousand Oaks, CA, USA, June 14–16, 1999. 43. Hautala, T. and Tiainen, T., “Friction stir welding of copper,” 6th International Trends in Welding Research Conference Proceedings, Pine Mountain, GA, USA, April 15–19, 2002, ASM International, 324–328, 2003. 44. Cederqvist, L. and Andrews, R.E., “A weld that lasts for 100,000 years: FSW of copper canisters,” 4th International FSW Symposium, Park City, UT, USA, May 14–16, 2003. 45. Johnson, R., “FSW of magnesium alloys,” 4th International FSW Symposium, Park City, UT, USA, May 14–16, 2003. 46. Esparza, J.A., Davis, W.C., Trillo, E.A. and Murr, L.E., “Friction-stir welding of magnesium alloy AZ31B,” Journal of Materials Science Letters, 21, 917–920, 2002. 47. Kimapong, K. and Watanabe, T., “Friction stir welding of aluminum alloy to steel,” Welding Journal, vol. 83 (10), 277–282, 2004. 48. Fukumoto, M., Tsubaki, M., Yasui, T. and Shimoda, Y., “Joining of ADC12 and
The friction stir welding process: an overview
49.
50.
51.
52.
53.
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SS400 by means of friction stir welding,” Quarterly Journal of the Japan Welding Society, 2004, 22 (2), 309–314. Smith, C.B., Hinrichs, J.F., Cerveny, L.M., Anderson, R. and Walker, B., “Fabricated shapes using FSW/forging process,” 7th International FSW Symposium, Awaji Island, Japan, May 20–22, 2008. Rhodes, C.G., Mahoney, M.W., Bingel, W.H., Spurling, R.A. and Bampton, C.C., “Effects of Friction Stir Welding on Microstructure of 7075 Aluminum,” Scripta Materialia, 36 (1), 69–75, 1997. Mahoney, M.W., Rhodes, C.G., Flintoff, J.G., Spurling, R.A. and Bingel, W.H., “Properties of Friction-Stir-Welded 7075 T651 Aluminum,” Metallurgical and Materials Transactions A, volume 29A, 1955–1964, July 1998. Juricic, C., Dalle Donne, C. and Dressler, U., “Effect of Heat Treatments on Mechanical Properties of Friction Stir Welded 6013,” 3rd International FSW Symposium, Kobe, Japan, September 27–28, 2001. Li, Z.X., Arbegast, W.J., Wilson, A.L., Moran, J. and Liu, J., “Post-Weld Aging of Friction Stir Welded Al 7249 Extrusions,” 6th International Trends in Welding Research Conference Proceedings, Pine Mountain, GA, USA, April 15–19, 2002, ASM International, 312–317, 2003.
3
Material deformation and joint formation in friction stir welding
R. Zettler, WTSH, Germany
Abstract: This chapter discusses how the friction stir welding (FSW) environment, including process parameters, welding machine, joining tools and backing bar, i.e. material restraint system, in combination with the material properties that constitute this environment, account for processing temperature, subsequent heat losses and thus the plastic deformation (material flow) behaviour, which results in the friction stir welded joint. Key words: friction stir welding environment, hot working deformation process, material properties, process temperatures, heat losses, material flow and joint formation.
3.1
Friction stir welding (FSW): a thermomechanical joining process
Although FSW is generally considered to be a relatively simple process, this chapter will demonstrate that joint formation is a complex process and not solely the result of an interaction between three principal processing parameters, namely tool rotation speed, weld travel speed and axial force. The reader will very quickly discover, through the following discussion of heat and material flow, that joint formation is subject to a specific thermal environment, which in turn dictates and is responsible for the selection of appropriate processing parameters to meet the metallurgical and environmental constraints imposed during processing. Process temperature plays a significant role in determining a material’s ability to deform. Each material after all has a temperature window which will provide for maximum formability during deformation. In the case of precipitation strengthened, otherwise referred to as age hardening or heat treatable aluminium alloys, the temperature developed during processing is known to influence the precipitate distribution across the weld (Mahoney et al. 1998). For the case of wrought alloys which do not respond to age hardening, increasing process temperature accelerates both climb and cross slip (dislocation glide), which lowers the extent of strain hardening, i.e. for a constant strain the flow stress decreases. Of equal importance to processing temperature are the magnitude and type of force involved to cause deformation. In the case of aluminium and 42
Material deformation and joint formation in friction stir welding
43
its alloys, plastic deformation necessitates a shear stress that exceeds a critical stress (sy), the yield stress of the material. This stress corresponds to the stress that is required to push a dislocation through an arrangement of many other dislocations and is applicable to all deformation processes including extrusion, strip or wire drawing, tube drawing over a mandrel, rolling, forging and FSW. On a macroscopic scale most if not all deformation processes share a common feature in that material is forced to flow through a channel or die of some kind. FSW, however, differs from mainstream deformation processes in that it develops a channel or die, which is totally unique to this particular process. This is because the die consists of both fixed and semi-fixed or floating geometry. Here fixed geometry includes the tool shoulder above each workpiece (upper boundary) and the backing bar below each workpiece (lower boundary). In the case of the semi-fixed or floating boundary, this occurs for the non-softened, i.e. rigid, and non-deforming workpiece material either side of the join line (side boundaries). Furthermore, the tool pin provides an additional internal and fixed boundary or central bridge to the die cavity. This fixed boundary actively participates in moving and forcing thermally softened material to flow, while at the same time helping to facilitate bonding through application of pressure and heat during processing. FSW has been identified and termed a thermo-mechanical (Cederqvist and Reynolds 2001) and hot working deformation process (Arbegast 2003). The process is thus described because it relies both on heat (thermal) and material flow, i.e. plastic deformation, to initiate softening and displacement of a finite volume of workpiece material, i.e. material in the direct vicinity of the joining tool (Guerra et al. 2003). This material is then forced to flow and form the so-called solid-state joint, sometimes referred to as the weld nugget or more consistently identified today as the stir zone. The combination of applied forces and resultant processing temperature derived from the interfacial conditions between the rotating and traversing tool with that of the workpiece material causes macroscopic shape change to take place. This occurs in and directly adjacent to the stir zone via grain boundary sliding aided through recovery and recrystallisation, processes (Jata and Semiatin 2000), and for the case of precipitation hardening and non-heat-treatable alloys at temperatures which ensure the dissolution of hardening particles (Shercliff et al. 2005) combined with acceleration of dislocation glide respectively. It should be noted here that macroscopic shape change cannot be adapted at an atomic level without a change in the material’s crystal structure. Structure, however, is conserved when crystallographic planes are displaced (http:// aluminium.matter.org) and this is always a shear process. The ability to activate rafts of atoms to slip over neighbouring rafts is one of the most useful attributes of all metals, including aluminium and
44
Friction stir welding
its alloys, and is the reason why metals are capable of undergoing large plastic deformation, particularly when aided at temperatures above 0.5 the melt temperature (Tm) of the material being deformed. Hence the term hot working.
3.2
Plastic deformation in relation to material properties
For a given working pressure and temperature there will always be a maximum amount of deformation that can be imparted to the workpiece. This limitation is based on the flow resistance at processing temperature of the material being deformed. Commercial metals and alloys are strengthened by various types of obstacles to dislocation motion. These obstacles include: ∑ other dislocations (via work or strain hardening) ∑ grain boundaries ∑ solute atoms (solution hardening) ∑ precipitated GP zones (precipitation hardening) ∑ dispersed particles (dispersoid hardening). It is fair to say that most if not all engineering materials contain at least a small amount of impurities and or deliberate additions, i.e. alloying elements to inhibit crystallographic slip from occurring. In the case of aluminium, some elements, notably copper (Cu), lithium (Li), magnesium (Mg), manganese (Mn), silicon (Si) and zinc (Zn) dissolve, up to a certain limit, which is temperature dependent in the aluminium crystal lattice, so as to form solid solutions within the aluminium matrix. These foreign (solute) atoms substitute for an aluminium atom, where they have a different size compared to the aluminium atoms. This means that solute atoms whose sizes typically vary in comparison to aluminium atoms introduce a distortion into the aluminium lattice and thus aid in inhibiting dislocation motion. A quantitative estimation of crystal lattice distortion and the effect lattice parameter change have on extrusion force has been investigated for binary and ternary aluminium alloys containing a solid solution (Zakharov 1995). The results demonstrate that with the exception of alloying additions of zinc and silicon, extrusion force increases with increasing lattice parameter change, i.e. increasing distortion. In addition to major alloying elements and their contribution to increasing the resistance of an alloy to deformation, minor elements such as scandium (Sc) and zirconium (Zr) also play an important role, since they form a fine intermetallic precipitate in aluminium that helps inhibit recovery and recrystallisation. Apart from physical barriers to crystallographic slip and thus deformation, the obvious limit on the temperature scale is the solidus temperature (the
Material deformation and joint formation in friction stir welding
45
temperature at which a material begins to melt). Additionally, hot shortness, the tendency which some materials have to separate along grain boundaries when stressed or deformed at temperatures near their melting point, often caused by the presence of a low-melting point constituent contained within a material, e.g. second phase particles, is also a limitation. Increasing the strain rate when deforming a material containing such low melting point constituents not only results in an increase in workpiece temperature but more rapid attainment of the critical temperature (Tcr) at which tearing of the grains begins. Here grain boundary embrittlement is seen to lead to the fracture mode changing from one of tough intergranular to brittle intragranular fracture. This is witnessed in extrusion processes and is one of the major reasons why it is not possible to extrude Al-Zn and Al-Cu solid solution strengthened alloys at the same rate as Al-Si based alloys (Zakharov 1995). In most instances very large deformations are only possible using hot working operations. Here hot working refers to deformation conducted at temperatures above 0.6 times the melt temperature of the material being deformed and for strain rates such that recovery processes occur substantially during deformation (Dieter 1988). The need for such temperatures is because the mechanical working of a metallic material at room temperature will lead to a magnitude of dislocations forming in which slip becomes progressively more difficult. Hence dislocations increase internal stress and strain hardening is said to take place. By further applying cold work to the material, this raises the level of internal stress until fracture eventually occurs. Strain hardening thus provides the reason why very large deformations are not possible unless the strain-hardened and distorted grains produced under cold working conditions are first eliminated, which is essentially achieved through hot working the material. During hot working two metallurgical factors contribute to the annihilation and re-formation of the pre-existing microstructure. These processes include recovery and recrystallisation. Dynamic recovery and recrystallisation are seen as active participants responsible for lowering a material’s flow stress, since strain hardening and distorted grain structures produced during deformation are rapidly eliminated by the formation of new strain free grains (Fonda et al. 2007, Prangnell and Heason 2005, Su et al. 2003, Jata and Semiatin 2000). Difficulties, however, arise when attempting to correlate plastic deformation behaviour with recrystallisation behaviour. This is because the deformability of a metallic material, i.e. its capacity to flow and change shape without fracture, is seen to depend on at least six key and inter-related material variables that very often demonstrate complex and non-linear relationships between them. These variables include: (1) the amount of prior deformation, i.e. mechanical working, (2) temperature, (3) time, (4) initial grain size, (5) chemical composition of the material (here alloying additions act as
46
Friction stir welding
deformation inhibitors) and (6) the amount of recovery or polygonisation prior to the start of recrystallisation. Dieter (1988) describes how these six variables can affect recrystallisation behaviour. 1. The amount of prior deformation – there will be a minimum amount of deformation needed to cause recrystallisation. The smaller the degree of deformation, the higher the required temperature. 2. Temperature – peak temperature is the important consideration. The amount of deformation required to produce equivalent recrystallisation behaviour increases with increased working temperature. 3. Time – increasing annealing time decreases recrystallisation temperature. However, temperature is more important than time. An increase in temperature of approximately 10°C is equivalent to a doubling of the annealing time. 4. Initial grain size – the larger the original grain size the greater the amount of cold work required to produce an equivalent recrystallisation temperature. The final grain size, however, will depend on the degree of deformation and to a lesser extent on the annealing temperature. The greater the degree of deformation and the lower the annealing temperature, the smaller the recrystallised grain size. 5. Composition – recrystallisation temperature decreases with increasing purity of a material. Solid solution alloying additions raise the recrystallisation temperature. 6. Amount of recovery or polygonisation prior to the start of recrystallisation – the stored energy prior to deformation is the driving force for recrystallisation. In addition to the material-related variables a number of quantitative factors are also seen to be involved in governing process temperatures generated during deformation. These factors include: (a) the initial temperature of the forming tool and its thermal conductivity, (b) heat generated due to plastic deformation, (c) heat generated due to friction at the workpiece/tool interface and (d) the magnitude plus direction of heat transferred or lost from the deformation zone to the surrounding environment. Because the temperature at which recrystallisation occurs depends on an interplay between all of the variables listed above (1–6 and a–d) there is to date no absolute or fixed law governing these relationships.
3.3
Process parameter, temperature and heat loss relationships
Clearly in terms of FSW it is the rotating and traversing tool in contact with the workpiece material that initiates material flow, i.e. plastic deformation,
Material deformation and joint formation in friction stir welding
47
while at the same time providing the heat required to reduce the material’s flow stress and so allow for constant volume deformation to take place. It should be noted, however, that heat generation is not just limited to the surfaces of the points of physical contact between the tool and workpiece material. This is verified for FSW when one compares area maps of stir zones as viewed transverse to the weld direction and in relation to the shape and size of the FSW tool used to produce these stir zones. Contrary to popular belief is the fact that the stir zone will actually decrease in size rather than increase with ever increasing tool rotation speed (Colegrove et al. 2007). This phenomenon is also witnessed for the deformation zone developed during friction welding (Vill 1962). The fact that the stir or deformation zone will not continuously increase with ever increasing tool rotation speed can be accounted for in that increasing rotation speed increases processing temperature to the point where workpiece material in the direct vicinity of the joining tool significantly loses flow strength. Thus strengthening, which would otherwise be anticipated to occur due to an increase in strain rate at higher tool rotation speeds, cannot compensate for the effects of thermal softening. Subsequently material slip must occur at the tool/workpiece interface. The idea that material slip is possible during FSW is now a well-recognised phenomenon, which is incorporated into most empirical process expressions, particularly when calculating local heat generation rates, dqhg: such as in equation 3.1.
dqhg = d (wr − V sin q) mp dA
3.1
here d defines the extent of slip, w is the angular tool rotational speed, r is the radial distance (as applied to either the pin or more typically to the tool shoulder) from the tool axis, V is the welding speed, q is the angle between the radial vector, r and the welding direction, m is the friction coefficient and p is the local pressure applied by the joining tool on the deforming/ flowing elemental area dA. From equation 3.1 it can be inferred that when d is 1, no material sticks to the tool and all the heat is generated by friction. In contrast, when d approaches 0, all the heat is generated by plastic deformation. The difficulty encountered when attempting to apply equation 3.1 is that both the degree of slip and a uniform and consistent method for determining the friction coefficient have proven difficult, whether by means of fundamental principles or via representative experiments (Nandan et al. 2008). An alternative approach to using process efficiency factors such as slip or friction is by application of the most fundamental relationship applicable to all plastic deformation process, namely the constant-volume relationship. It should be recalled that all deformation processes impart forces onto the workpiece. The categories involved include: direct compression, indirect
48
Friction stir welding
compression, tension, bending and shearing. Obviously, the applied forces must develop yielding in the material being processed (deformed) but the stresses must be such so as not to create localised fracture. For the case of FSW this means that if deformation is generated by grain boundary sliding, this deformation must always occur at some distance, however small, from the direct tool/workpiece interface. Hence a considerable amount of work has been attempted in the past to correlate the degree of flow stress generated during deformation with process parameters through constitutive and empirical models. Here the flow stress is seen to depend mainly on the strain, strain rate and temperature during processing. Typically the temperature and strain rate behaviour of a material is calculated based on a formulation of the deviatoric flow stress proposed by Sellars and Tegart (1972) and modified by Sheppard and Wright (1979) using the Zener-Hollomon Parameter (Z), which represents the temperature compensated effective strain rate, and is calculated according to
Z = e´ exp(Q/RT)
3.2
here e´ is the effective strain rate, i.e. the deviatoric strain rate, Q is the temperature independent activation energy similar to that for self-diffusion, and R is the gas constant, while T is the absolute temperature. The strain rate is found to correlate with the flow stress and temperature as follows
e´ = A(sinh a se)n exp(–Q/RT)
3.3
here the material constants A, n, a and the apparent activation energy Q are derived by fitting the equation to experimental data. By rearranging equation 3.3 the effective material flow stress becomes
se = 1/a (sinh)–1 [(Z/A)]1/n
3.4
A drawback to such empirical methods is that they require the values of the constants to be determined by means of regression. Subsequently the quantitative assessment of these models can yield a wide range of errors, which have been estimated to be up to 60%, i.e. a range of strain rates from 0.0001 to 100 s–1 (Reddy et al. 2008). As an example Jata and Semiatin (2000) have estimated a typical deformation strain rate of 10 s–1 when FSW 7.6 mm thick Al-Li-Cu alloy. Their estimate was based on measuring grain size and using a correlation between grain size and the Zener-Holloman parameter. Chang et al. (2004) have also estimated effective strain rates ranging from 5–50 s–1 when FSW the magnesium alloy AZ31. These estimates were once again based on relationships between grain size and the Zener-Hollomon parameter. On the other hand, Masaki et al. (2008) estimate effective strain rates in the range 2–3 s–1, while Chao et al. 2001 propose that strain rates can be as high as 1000 s–1.
Material deformation and joint formation in friction stir welding
49
Colegrove (2006) presents a plausible reason for the wide variation in strain rates measured during FSW. He notes that while relationships based on grain size and the Zener-Hollomon parameter hold at low temperatures, this is not always the case near the solidus, where significant softening at the tool/ workpiece interface occurs. In the context of heating resulting from plastic deformation, it should be noted that strain rates are very much dependent on material flow and the velocity gradients generated by contact with and the applied pressure given by the FSW tool to the workpiece material. At the near solidus, however, strain rates rapidly diminish, particularly given conditions in which lubrication and slip can occur. Studies by Khandkar et al. (2006), Nandan et al. (2007) and Schmidt and Hattel (2005) further suggest that a condition of slip at the tool/workpiece interface aids in promoting convective rather than conductive heat transfer between tool and workpiece, and that this ultimately affects the local temperature distribution close to the tool/workpiece interface. Admittedly it is more common to use the term convection or convective heat transfer to describe the flow of heat from a surface to a moving fluid. Mills (1992), however, points out that a fluid, by virtue of its mass and velocity, not only transports momentum but, by virtue of its temperature, transports energy. Thus strictly speaking the transport of energy (heat) by motion of a medium, including a moving solid is and can legitimately be described as convection rather than conduction. Kong and Ashby (1991) recognised for the case of friction welding that heat could be lost by conduction, convection and radiation, whereby dry sliding tests confirmed that conduction dominates. Kong and Ashby (1991), however, also noted that heat was likely to flow into two solids partitioned not only as a result of their geometry and thermal properties, but also based on the nature of the heat flow, whether this was linear, radial, transient, steady state or based on the thermal contact resistance between the materials. Subsequently they acknowledged that heat flow will be dependent upon and thus linked to the nature of the deformation taking place. In terms of FSW it can be concluded that the depth of deformation, i.e. the layer or rather layers of deforming, i.e. flowing material, which move around the tool pin, will be dependent on the temperature and velocity gradients developed between the heat source (tool/workpiece interaction) and the energy losses, which go into the surrounding environment. This gradient will be a function of the material’s chemical composition, which determines its resistance to deformation and subsequent thermo-physical properties such as. solidus temperature and thermal conductivity. These in turn will then dictate process temperature limits and energy requirements in respect to heat losses. It is a well-known fact that aluminium alloys with a high amount of alloying elements in solid solution, i.e. 2xxx, and 7xxx series, possess a
50
Friction stir welding
high resistance to deformation. It can therefore be anticipated that these hard alloys will also produce stir zones that develop the extremities of the deformation layer(s) much closer to that of the limits of the joining tool when compared to alloys with a lower resistance to deformation, i.e. soft and low solute content alloys, particularly when both alloy types are processed under identical processing parameters and environmental conditions, i.e. same FSW machine, clamping system and tooling (Zettler 2008). Such an example can be found in Fig. 3.1. The macrographs of stir zones, Fig. 3.1, produced primarily from tool shoulder contact, since the pin was conical and non-threaded with length approximately 2 mm, demonstrate just how different stir zones can be for bead on plate welds produced across diverse aluminium alloys such as AlCu 2024 (hard), Al-Si 6013 and an AlMgSc alloy 8042 (soft).
2024-T351: 800rpm-400mm/min-10kN
5 mm
6013-T6: 800rpm-400mm/min-10kN
5 mm
AlMgSc: 800rpm-400mm/min-10kN
5 mm
3.1 Comparison of stir zone formation between weld macrographs produced in 4 mm thick 2024-T351, 6013-T6 and an 8042 (AlMgSc) alloy using a 15 mm diameter scroll shoulder in conjunction with a 5 mm diameter at shank conical and non-threaded pin. The pointed pin measured 1.75 mm in length and welding was conducted by the author (Zettler 2008) for a constant tool rotation speed of 800 rpm, a weld travel speed of 400 mm/min and an axial force of 10 kN. The weld zone macrographs demonstrate that the softer and lower solute content alloys, namely the 8042 and 6013 alloys respectively, produce the larger stir zone in comparison to the harder and higher solute content 2024 alloy.
Material deformation and joint formation in friction stir welding
51
Not only does the size and shape of the friction stir weld stir zone vary with regard to the alloy type being friction stir welded but so too do process temperatures. Zettler and Potomati et al. (2006a) noted that for the case of dissimilar aluminium alloy friction stir welds the alloy with the lower resistance to deformation always produces the higher processing temperature, regardless of the location of the alloy within the welded joint, i.e. whether placed in the advancing or retreating side. One of the reasons for higher process temperatures in what can be termed the softer low solute content alloys can be seen to be related to an increasing deformation volume, i.e. stir zone. Hence an increased stir zone gives rise to an increased potential for adiabatic shear and heat generation to occur. Subsequently it can be assumed for such alloys that viscous heat dissipation dominates over friction (rubbing) induced heating. It is well acknowledged that process temperatures during FSW are influenced by the processing parameters, while material properties, not just the workpiece but including that of the tool and clamping/backing system, will affect process temperature profiles through cooling rates, and that these relationships will ultimately determine the mechanical properties of the friction stir welded joint (Colligan and Mishra 2008, Yan et al. 2005). When one considers FSW process temperatures, a unique feature, generally, speaking is the presence of asymmetry in the temperature fields that are generated during processing. This asymmetry can and does occur both in relation to each side of the welded joint, i.e. between advancing and retreating sides, but also in terms of temperatures generated for the through thickness of the welded joint. An example of asymmetry for the latter condition is readily identified when one compares the extent of thermal softening that occurs in and around the welded joint for the case of bobbin tool welds, i.e. simultaneous dual shoulder welds, refer to Fig. 3.2.
Y/mm
HV 0,2
4.0 3.0 2.0 1.0 0.0
78 87 95 104 112 121 130 138 147 155 164
5 mm
–20
–15
–10
–5 X/mm
0
5
10
15
3.2 Macrograph (upper image) and representative interpolation of reduction in hardness across the friction stir welded joint as produced by the author in 5 mm thick 2219-T76 alloy using a fixed gap bobbin tool. Here hardness Vickers measurements demonstrate that softening is much more pronounced in the lower half of the weldment.
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Friction stir welding
The bobbin tool weld, Fig. 3.2, although demonstrating relative symmetry in thermal softening between each side of the joint, clearly shows that softening is considerably more developed in the lower half of the weld. The reason for this to occur can be attributed to the fact that heat removal through the FSW tool is considerably lower for the case of the bottom tool shoulder, since the only means for heat loss in this region of the weld, other than into the workpiece, is into the air, where heat loss is considered to be marginal (Colegrove et al. 2007). As such the heat generated by the lower tool shoulder is reflected back into the workpiece where subsequently greater softening occurs. Both the tool shoulder and tool pin are seen to play a significant role when it comes to heat generation during FSW. Xu et al. (2003), for example, provides useful insight for the design of FSW tools and selection of process parameters when FSW thick section (25 mm and above) workpieces. They note that heat generation based on a tool shoulder: pin contribution ratios can vary from 60%: 40% to 30%: 70% depending on process conditions. Furthermore, increasing the tool advancing speed or heat loss through the backing bar increases the fractional heat due to the pin, increases the peak temperature in the weld zone and reduces the time duration for the material to experience elevated temperatures. Additionally, when welds are compared, having the same weld travel speed but produced with and without the use of an insulated backing bar, the welds for the no insulation case demonstrate higher temperatures in the stir zone, although the temperatures drop much more quickly than is the case when the insulated backing bar is used. In comparison to thick section welds, friction stir welds produced in considerably thinner wall thicknesses (5 mm and below) demonstrate significantly more heat contributed by tool shoulder interaction with the workpiece material. Estimates, for example, place the contribution of the tool shoulder in relation to the tool pin as ranging between 98% : 2% and 86% : 14% (Russell and Shercliff 1999, Schmit et al. 2004) respectively. Furthermore, owing to decreasing material thickness it is argued that a greater proportion of heat is removed by the workpiece and thus heat lost through the tool is regarded as being minimal. The fact that tool shoulder to tool pin heat generation ratios vary when FSW should not surprise the reader, since the stir zone of most if not all friction stir welds consist primarily of two dominant flow region, i.e. the upper shoulder induced and the lower pin induced flow zones. Dong et al. (2001) along with Song and Kovecevic (2003) strongly suggest that friction is dominant in the upper region of the stir zone and that plastic work induced heating is significant and dominant in the lower region. As such the relative amounts of heat generated by these regions in overcoming the resistance of the material to deform will influence not only the peak temperatures experienced but also the temperature isotherms produced during processing.
Material deformation and joint formation in friction stir welding
53
Figure 3.3 highlights the nature of temperature profiles as measured using K-type thermocouples embedded at mid-plate thickness for both sides of a friction stir welded joint. It should be noted here that the temperature profiles in this case pertain to the bobbin tool weld, Fig. 3.2. Although bobbin tool welds have three rather than two contacting surfaces between the tool and workpiece, i.e. two tool shoulders and the tool pin, these welds have similarities with that of FSW using an insulated backing bar, since this avenue of heat loss is negated and thus are a good reflection of heating and cooling rates predominantly for the given workpiece material. In the case of the temperature profiles, Fig. 3.3, two features clearly stand out. The first pertains to just how symmetric process temperatures are between the advancing (As) and retreating (Rs) sides of the joint for this weld, while the second feature highlights the typical rapid rise of temperature and much slower rate of cooling after peak temperatures have been reached. The fact that there is a much steeper rate of heat up as compared to cooling down for 600
@ 0 mm
550 500
T
[email protected] mm
450 Temperature (°C)
T 1-As@10 mm
@ 10 mm
400
@ 12.5 mm
350
@ 15 mm
T 3-As@15 mm T 4-Rs@10 mm
300
T
[email protected] mm
250
T 6-Rs@15 mm
200
Computed temperature @ 0 mm
150 100 50 0 0
20
40
60
80 100 120 140 160 180 200 220 240 260 280 300 Time (sec)
3.3 Temperature profiles as measured by the author using K-type thermocouples embedded in the workpieces when FSW 5 mm thick 2219-T76 alloy using a fixed gap bobbin tool. The temperature profiles demonstrate a rapid increase in temperature as the tool approaches the thermocouple. Once the peak temperature is reached, the rate at which the workpiece cools down can be seen to be much slower than the rate at which the workpiece heats up. Hence it can be concluded that temperatures (isotherms) are more elongated at the rear of the tool in comparison to the front of the tool. Additionally the temperature profiles demonstrate that the higher the processing temperature, the more rapid the initial rate at which the workpiece begins to cool down.
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Friction stir welding
FSW process temperatures means that temperature isotherms surrounding the FSW tool can be regarded as being narrower ahead of the joining tool, but considerably more elongated behind the tool, where both material and subsequently heat are deposited. Peak processing temperatures and cooling rates are an important consideration during all hot working processes. In the case of precipitation hardening alloys the strengthening precipitates present prior to FSW are generally dissolved in the stir zone, while in the heat affected zone (HAZ), precipitates coarsen and an over-aged condition is established (Heinz and Skrotzki 2002). The extent of this over-aging will depend on the cooling rate. In terms of FSW it is well accepted that peak process temperatures increase with increasing tool rotation and or decreasing weld travel speed. Furthermore, cooling rates during FSW significantly increase with increasing weld travel speed. Cooling rates, however, are also seen to increase with increasing weld temperature. This means that both higher welding speeds and tool rotation speeds will result in higher cooling rates. By adjusting welding parameters, such as tool rotation and weld travel speed, it is possible to influence either the peak processing temperature or the rate of cooling for the workpiece. This, in turn, will impact on the resultant microstructural transformations. Care should be taken, however, in the selection of processing parameters, since parameters selected based only on optimal heating and cooling rates are generally unable to facilitate constant volume processing and thus increase the potential to produce flaws in the welded joint. Invariably the major standard for judging the quality of a weld, whether produced by fusion or friction/pressure-based welding processes such as FSW remains one of comparing the strength of the weld in relation to that of the base material properties. Simply put, the effectiveness of the welded joint is based on trial and error experimentation. Many distinct factors, however, can be seen to influence the residual strength of a weld. These include the material’s thermo-physical properties, the method of welding, the design of the joint, the energy input and losses which occur, and then the combined interaction between all of these factors. For fusion welding it is standard practice to calculate heat input as a function of voltage, current, weld travel speed and the relative efficiency of the process. Here the shielded metal arc welding process is rated as having an efficiency of 0.75, gas metal arc welding and submerged arc welding, 0.9, while gas tungsten arc has a value of 0.8 (Weman 2003). It is generally assumed that most if not all of the energy generated during FSW is consumed in the formation of the joint. This assumption, however, remains inconclusive. Additionally, FSW differs significantly from fusion welding in that the flow of heat from the heat source is not unidirectional,
Material deformation and joint formation in friction stir welding
55
i.e. operating or moving or allowing movement in one direction only. FSW, after all, relies on the interaction of the tool with the material to thermally soften and activate material flow. In the case of FSW the tool is not just part of the heat source but is itself a not insignificant avenue for heat loss. This avenue along with all other avenues for heat loss is presented schematically in Fig. 3.4 and incorporated within the flow (loss) of heat equation (denoted by Q = thermal energy), equation 3.5.
Qtotal = Qstir zone + Qtool + Qworkpieces + Qatmosphere
+ Qbacking bar + Qclamps
3.5
Obviously the thermal diffusivity, i.e. the ratio of thermal conductivity to the volumetric heat capacity of the materials, and the subsequent masses involved in each of the heat loss pathways, represented in Fig. 3.4 will and do have an impact on weld formation. If the diffusivity is high, for example, the material cooling rate is high and the HAZ of the joint will be small. Conversely, a lower diffusivity leads to slower cooling and a larger HAZ. Furthermore, the rate of heat transfer will depend on the ability during FSW to maintain constant temperature gradients between the heat source and heat sink(s), i.e. this corresponds to the temperature difference potential or concentration potential that heat energy will flow from one region to another (Zettler 2008). Evidence as to the importance heat loss avenues have in terms of stir zone formation is no better portrayed than through changes to backing bar material when FSW (Zettler 2008). Traditionally the backing bar or anvil material used for classical or single-sided friction stir welds has consisted of a structural steel that not only supports but also limits the potential for diffusion bonding of the aluminium to the anvil during processing. These steels normally have a thermal conductivity of between 10 and 20% that of the aluminium alloy being friction stir welded. Hence the anvil material
Heat loss/Air Heat loss/Tool Heat loss/Parent material Heat loss/ Backing bar
1. Frictional heating 2. Adiabatic heating, i.e. a shear process
3.4 Heat flow (loss) pathways that are acknowledged in the literature to occur during classical, i.e. single-sided, friction stir welds (Zettler 2008).
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Friction stir welding
acts not just to support the workpiece but also as a barrier controlling heat transfer during processing. The temperature and the rate at which heat is lost is an important consideration when it comes to all hot working fabrication practices. For the case of precipitation hardening, i.e. heat treatable aluminium alloys, it is generally desirable, when producing these alloys, to maintain as high a working temperature as possible, i.e. one which approaches but also inhibits undue melting. This is because higher temperatures are seen to help minimise recrystallisation after solution heat treatment, which further benefits strength and stress corrosion resistance. Additionally, it is desirable to achieve adequate homogenisation, though this is not always necessary, since, to eliminate soluble second phase particles, it is the ability to return the soluble elements to solution during solution heat treatment that is important. Good solutionising, however, is more readily achieved if the soluble elements have been in solution at some point during processing. For thinner products, this may not be necessary, since coarse particles may be broken up during fabrication, and so can be more readily returned to solution during an intermediate anneal or during solution heat treatment. A homogenisation or intermediate anneal is, however, necessary in order to precipitate Cr, Zr or Mn from solid solution. This is an important processing step since the size and distribution of these particles determine the final degree of recrystallisation. Here either increased time or temperature can be used to increase the solutionising of the soluble elements (www.Metals.com). Since both temperature and time play a significant role in determining the microstructure and subsequent properties of precipitation hardening alloys, it should come as no surprise that processing conditions, including the backing bar and tool material type used to produce a friction stir weld through their capacity to transfer heat, can and do have considerable influence when it comes to weld formation. Zettler (2008) notes that the relative amounts of heat which can be lost through each of the major heat loss avenues, negating air have the potential to vary with increasing weld length, i.e. over and above 2 m of friction stir weld length. Additionally he demonstrates that the relative amounts of heat and the direction of this loss are in fact influenced by the temperature of the tool and the ability of the tool support structure (spindle) to maintain a constant temperature gradient between the heat source (tool/workpiece interaction) and this avenue of heat transfer. The fact that there exists no standard set of FSW parameters, which can be actively transferred to give optimised mechanical properties for the FSW for any commercial grade aluminium alloy, and that process parameters are themselves not readily transferable across different welding machines, has for a long time indicated that the construction of the FSW machine plays a much more significant role in process parameter selection than has
Material deformation and joint formation in friction stir welding
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previously been accredited. The variability which exists in the literature for FSW parameters is not, as many would like to believe, a consequence of process robustness. It is true to say that FSW is a forgiving process in that for any welding machine there exists a window of parameters which will produce flaw-free welds with good mechanical properties. The fact that FSW parameters, unlike parameters established for other industrial deformation processes, cannot be directly transferred across machines, however, makes real just how dependent the process is on the thermal influence provided by and as a result of machine and clamping construction. Naturally, FSW parameters such as tool rotation speed contribute to how rapidly the tool heats up. Higher tool rotation speeds will result in greater friction heating at the tool/workpiece. Slip at the interface between the tool and workpiece further increases the potential for more heat to be lost vertically through the tool away from the stir zone and not radially into the workpiece. As a consequence process efficiency diminishes, since the thermal energy developed is not completely utilised in development of the stir zone formation. At the same time an increasing tool temperature causes the tool support structure to heat up. The capacity to maintain heat uptake through the tool will therefore be subject to the ability of the tool support, i.e. spindle construction, to provide for constant temperature gradients between the heat source and this avenue of heat loss. The effect of a diminishing temperature gradient between the heat source and heat sinks is not a new phenomenon to the FSW process. It is witnessed at the end of the friction stir weld length, i.e. when coming to the end of the workpieces. Here heat is known to reflect back towards the tool, which not only causes increased thermal softening. This can cause the tool shoulder under axial load to sink uncontrolled into the workpiece and is due to an increasing volume of thermally softened material, such that the workpiece can no longer support the pressure (force over area) placed onto it by the tool shoulder. The phenomenon of increasing deformation volume during FSW is demonstrated by Fig. 3.5 (Zettler 2008). Here FSW was undertaken in 4 mm thick 6013-T6 alloy. The macrographs of the stir zone transverse to the weld travel direction highlight the stir zone at weld begin approximately 100 mm after initial tool plunge and then after approximately 300 mm of weld length respectively. It should be noted that the weld was produced for a tool rotation speed of 3600 rpm and a weld travel speed of 900 mm/min in combination with a ZrO2/Y2O3 coated (insulated) backing bar. The results demonstrate that through the use of the insulated backing combined with the fact that the 6013 alloy has a relatively high thermal conductivity (approximately 160 W/mK) this alloy has the potential to force more heat to flow into the workpiece material and thus influence deformation i.e. increase the size of the stir zone. In addition to an increase in deformation volume Zettler (2008)
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5 mm (a)
5 mm (b)
3.5 Weld macrographs produced in a single bead on plate friction stir weld demonstrating growth of the friction stir weld stir zone in approximately 300 mm of weld length. The welds images (a) and (b) were produced by the author (Zettler 2008) in 4 mm thick 6013-T6 alloy, where image (a) originates 100 mm after initial tool plunge and image (b) is obtained from the same weld approximately 300 mm removed from image (a). Here welding was performed using a ZrO2/ Y2O3 coated (insulated) backing bar for a tool rotation speed of 3600 rpm, a weld travel speed of 900 mm/min and an axial force of 10 kN.
also observed that this increase had a noticeable effect on process stability as measured through monitored process signals including spindle torque, tool position relative to the workpiece surface and processing forces.
3.4
Material flow and property relationships of the resultant friction stir welded joint
Macrographs, micrographs, transmission electron microscopy (TEM), microhardness measurements as well as mechanical destructive testing of the weldment are common probes for the investigation of friction stir welds. All of them, however, only give selected and localised information regarding the outcome of the FSW process and not what the actual mechanism was by which the welds, or more precisely how the stir zone formed. As a consequence a large majority of research conducted into the FSW process has focused
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specifically on visualising the material flow around the FSW tool. It is well documented that plastically deforming material is forced to flow in the direction of tool rotation from in front to the rear of the FSW tool (Guerra et al. 2003, London et al. 2001). What is less well recognised, however, is that there exist two separate material flow paths, and that these originate from each side of the workpieces being joined in friction stir butt welds. Furthermore, the flow paths remain separate even when bonding between workpieces is intimate and complete. Here the work of Zettler and Donath et al. (2006b) having placed a Titanium (Ti) powder measuring 20–60 mm into both the retreating and advancing sides of a 2024-T351 alloy with 4 mm plate thickness demonstrated that dispersion of the powder marker was markedly different when originating from advancing as opposed to the retreating side of the joint. It should be noted here that a powder was used in order to limit the extent to which the foreign marker material could act as a single large singularity within the stir zone of the friction stir welded joint and thus not mirror the deformation behaviour of the parent material. Zettler and Donath et al. (2006b) aided through the use of micro-computer tomographic (mCT) renderings of the tomographic volume data further observed for marker originating from the advancing and then the retreating sides of four separate welds (all performed under identical processing conditions) that the marker was not only dispersed as a consequence of tool pin profile but also as a result of the initial marker implant location within the joint. Here the Ti marker material could be seen to be much more finely disrupted and dispersed when coming from the advancing or shear side of the joint. Additionally shearing and dispersion (location) of the Ti marker was notably different as a consequence of FSW having been performed using a threaded pin with three equally spaced flats, as compared to FSW using the threaded pin without flats, i.e. a continuous and uninterrupted thread. Further flow visualisation experiments performed by Zettler (2008) have highlighted for the case of Ti and Cu powder markers embedded simultaneously in the friction stir welded joint, Fig. 3.6 that marker material originating from the advancing side of the joint remains in direct contact with that of the tool pin. Marker material, however, originating from the retreating side attaches itself to the flow arm from the advancing side external to the tool pin and is dragged around the tool pin as a secondary flow. Bonding is also seen to take place and is completed between the two flow arms in the retreating side of the joint. Here it was observed, and as can be seen in Fig. 3.6 that both flow streams are forced to come together under the influence of the rigid and non-deforming material in the side of the joint. Ample evidence exists in the literature clearly demonstrating the importance of controlling deformation (both heat and material flow) in relation to the retreating or flow side of the friction stir welded joint. For example, dissimilar alloy friction stir welds demonstrate that if the material in the retreating side
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Advancing side (Ti)
Retreating side (Cu)
5 mm (a)
Advancing side (Ti)
Retreating side (Cu)
5 mm (b)
3.6 Micro-computer tomographic renderings of the tomographic volume data for marker material displacement (Ti powder light, Cu powder dark) produced by the author (Zettler 2008) when FSW 4 mm thick 6013-T6 alloy using the stop action technique for the FSW tool. Both image (a) mid plate thickness and image (b) weld root clearly demonstrate two flow streams representative of marker implant location occur as the marker materials flow around the FSW tool pin. Here material from the retreating side remains on the periphery of the flow stream originating from the advancing side.
Material deformation and joint formation in friction stir welding
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remains too rigid during processing, this will restrict material from being transported around the tool pin, and thus explains why successful aluminium to steel friction stir butt welded joints (Uzun et al. 2005) place the harder and higher strength material (steel) in the advancing and not retreating side of the joint. Conversely, when the deformation or stir zone is allowed to develop removed from the join line, i.e. biased in the advancing side or extensive thermal softening in the retreating side of the joint, this has the potential to displace the boundary between material flow streams even further into the retreating side of the joint. The biased tool position or increased distance of the join line from the tool pin effectively decreases the likelihood that sufficient pressure develops such that bonding between workpieces will be incomplete. Thus, under bending, the fracture location for such welds appear typically in the retreating side of the joint, i.e. at the displaced join line or interface between workpieces. The notion that independent material flow streams exist and that these originate from either side of the FSW joint is not new. The banding often referred to as the onion ring like structure contained within the stir zone has been clearly identified as a result of regular layering of the flowing material in relation to tool rotation and weld travel speed (Xu and Deng 2008, Chen et al. 2008, Sutton et al. 2002, Krishnan 2000, Threadgill 1999). Additionally, characteristics related to the banding contained within the stir zone of friction stir welds and their correlation with numerically investigated equivalent strain fields (Xu and Deng 2008), grain size and aspect ratio (Srinivasan 2005) as well as particulate distribution within these bands (Sutton et al. 2002) all point to the fact that there is more than a single deformation layer or flow stream which develops during FSW, and that so-called intimate mixing, i.e. reducing both the scale and intensity of segregation within a flow stream, is not truly applicable to the FSW process. Schmidt and Hattel (2005) attribute tool rotation speed as being primarily responsible for not only contributing to the temperature fields but also the rate at which material flows around the tool pin during processing. It has also been reported (Yan et al. 2007, Colegrove et al. 2007, Peel et al 2006a,b, 2003) that the rotation speed of the FSW tool has substantially greater influence on the microstructure and mechanical properties of friction stir welds than either the influence of weld travel speed or axial force. Intuitively, the complexity of the highly deformed material interfaces which evolve in the stir zone of a friction stir weld should play an important role in determining the mechanical strength of the welded joint, since the band spacing has been demonstrated to correspond to the welding tool advance per revolution and that grain size and particle distribution are subsequently affected within the stir zone of friction stir welds (Yan et al. 2007). Furthermore, it has been identified that several weld flaw types are seen to occur as either flow or geometry related (Arbegast 2008).
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The process of correlating residues of material flow against the mechanical properties of friction stir welds remains complicated, however, since these flow streams, as has been indicated throughout this chapter, depend amongst other things on the temperature, strain and strain rate, on the interactions between material flow regimes, i.e. shoulder, pin and pin root, as well as tool profiles, and thus the degree of vertical as well as horizontal material flow brought about by welding parameters and tool features that cause material to circulate (extrude) around the pin, perhaps even more than once during processing (Reynolds 2008). Material flow patterns have been observed to vary considerably during extrusion, particularly in response to the material/alloy type, the material/ tool interface, i.e. conditions of friction/no friction at both the container wall and the billet, and/or significant friction at the surface of the die and its holder, and in response to the shape of the extrusion. Four types of flow patterns have been documented to occur during extrusion (Aluminium and Aluminium Speciality H/B 1993). A key characteristic and thus consideration associated with the development of all flow types, however, is the formation of dead metal zones. Large dead metal zones imply large velocity gradients in the material being extruded. Large velocity gradients are also synonymous with significant internal friction and localised shearing, i.e. the generation and development of heat within the extrudate. Naturally such situations must be avoided, since an increase in temperature of the alloy to the point where the critical temperature is reached will result in the fracture mode changing from tough intragranular to one which is brittle and intergranular. This is witnessed typically on the surface of the extrudate where tearing generally occurs. Although no direct comparison is possible between the residual patterns of material flow that arise when extruding an aluminium alloy through a die against those resulting from the FSW process, a key principle governing material flow behaviour for both processes is that material flow must remain continuous, and gross flow instabilities should be avoided if constant volume and flaw-free processing is to be achieved. When one considers the residual flow patterns occurring in single alloy friction stir welds it is clear to see from the literature that these vary in accordance with the aluminium alloy (series), tool form and the process parameters used during welding. What is fundamental to all stir zones, however, given the weld (stir zone) has been properly etched, is that flaw-free stir zones demonstrate a systematic and non-chaotic layering of the bands, both in the horizontal transverse and vertical longitudinal directions. Such features are highlighted in Fig. 3.7. There are several key visual features which can be used to relate the quality of the friction stir weld to process parameters. These include the relative sizes of the shoulder induced flow zone to that of the pin induced
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2 mm
3.7 Banding or striations which are seen to occur both in the horizontal transverse and vertical longitudinal directions when FSW all aluminium alloys. The current weld zone macrographs were produced by the author in 4 mm thick 6013-T6 alloy and demonstrate a non-chaotic layering within the stir zone.
flow zone, the appearance of the banding, particularly in the lower half of the advancing side of the stir zone, and the extent of material adhesion to the tool shoulder and pin (free from or coated by workpiece material) after retraction from the weldment. The Welding Institute (TWI) and its group of sponsors (Dawes et al. 1995) very early on in the development of the FSW process demonstrated that clogging of tool profiles such as the pin threads was detrimental to bonding between workpieces. Tool coatings such as chromium nitride and titanium nitride were found to enable higher weld travel speeds and to reduce the extent of clogging. It has subsequently been shown that tool coatings also have the ability to influence process forces (Midling et al. 1999), in that they help to reduce process-generated forces opposing the FSW tool for equivalent levels of heat generation. From the above it can be seen that a relatively minor observation such as the condition of the FSW tool, whether tool profiles are dirty or clogged, can have a significant influence on the quality of a friction stir welds. Additionally, this observation in turn speaks against the notion that optimal processing conditions during the FSW of aluminium alloys occur for conditions where material sticks to the FSW tool, e.g. deformation occurs as a cylindrical or conical volume of material attached (sticks) to the tool during processing. A common problem that appears throughout the literature, particularly when attempting to correlate and interpret relationships between FSW parameters, base material properties and resultant friction stir welds is that this is typically performed for only a small part of the operating conditions
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Friction stir welding
which transpire during processing. Unlike FSW most manufacturing processes identify sensor signals which are then used as a means of quantifying process control and hence stability. This is typically based on experience and process knowledge. Here historical data collected for a process provides a useful first approach in formulating the data set from which both the control/input (or dependent) and measured/output (or independent) variables are selected. Limited information, however, still exists concerning the design of FSW tools, since much of the work in this area is commercially sensitive and controlled through licences. Even less information is available on how these tools should be employed, i.e. process parameter selection for both workpiece thickness and across diverse FSW machines. Information is also limited concerning how tool geometry influences processing forces generated during welding and how this impacts both on weld formation/quality and the requirements placed on the FSW machine. The literature identifies that there are many tool types which can successfully friction stir weld aluminium and its alloys. One cannot, however, simply upsize a FSW tool just because it was found to be capable of joining a thinner gauge (for example, under 8 mm thick) aluminium alloy (Thomas and Gittos 1999) in order to weld and join significantly thicker materials (above 20 mm in thickness). This is because increased weld penetration places different demands (forces) on the welding tool, and these demands may not always change in a linear fashion. Zettler (2008) as well as Colegrove and Shercliff (2006) have both identified that tool pin profiles, such as pin threads and flats on their own, have little influence on the temperatures generated when FSW. What is apparent, however, is that profiles contained on tool pins can and do have a significant effect on the forces developed during FSW, and this impacts on the ability to transfer thermally softened material without generating volumetric flaws within the stir zone. Hattingh et al. (2008) presents a good discussion concerning the influence flute design (e.g., number, depth and taper angle), tool pin diameter and taper plus the pitch of any thread form contained by the pin has on the lateral reaction forces and relative orientation of the peak resultant force during FSW. More importantly, however, their data indicates that these profiles provide such advantages not specifically as a result of the FSW process temperature. Both Colligan and Chopra (2004) and Zettler et al. (2004) have identified that tool pins containing flats or flutes wipe rather than draw the flowing material out as if thinning a strand of wire (Guerra et al. 2003). Furthermore Colegrove and Shercliff (2006) observe that the strain rate applied to material adjacent to the flats of the tool pin is lower than that at the periphery and subsequently the flow stress of the material in this region is also lower. This they argue is the reason why tool pins containing flats
Material deformation and joint formation in friction stir welding
65
are capable of producing lower processing forces and greater stability for material flow. Tapered threaded pins containing flats or flutes can be seen to typically produce stir zones with greater uniformity of shape in that the stir zone tends to more closely mirror the shape of the tool pin. Furthermore, the greater the uniformity of structure, i.e. the closer the size of the shoulder induced stir zone to that of the pin induced stir zone, and the less chaotic the banding within and between each side of the stir zone, the better the quality of bonding for the welded joint. Such an example can be found in Fig. 3.8 for the precipitation hardening aluminium alloy 2024. Here it can be seen that considerable symmetry of structure for the bands exists between each side of the joint. Furthermore there is no evidence of any chaotic flow features contained in the lower left-hand side of the stir zone, i.e. advancing side. The stir zone can also be seen to contain two rather than a single central site around which the bands are concentrated and appear to originate from. Typically such structures occur in response to material below this central point opposing the downward transport of thermally softened material that is augured by action of the tool pin. The lack of an outlet for the vertically
2 mm
3.8 Weld zone macrograph demonstrating the size of the stir zone and nature of the banding which occurs in process optimised in terms of mechanical properties for 4 mm thick friction stir welded 2024-T351 alloy. Here the weld was produced by the author using a 13 mm diameter scroll shoulder in combination with a 5 mm diameter (at shank) tapered and threaded pin with three flats. Welding was performed for a tool rotation speed of 1200 rpm, a weld travel speed of 300 mm/min and an axial force of 8 kN. The stir zone clearly shows a weld free from volumetric defects in which the pin induced flow zone is only slightly smaller in width than the shoulder induced flow zone generated at the workpiece surface. Additionally it can be seen that the banding contained in the stir zone demonstrates no chaotic flow features and is almost symmetric between each side of the joint.
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Friction stir welding
transported material results in the generation of a horizontally circulating flow. It should be noted that in terms of mechanical properties both the yield strength and ultimate tensile strength of the welded joint, Fig. 3.8 approached 96% that achieved in the base material, while the rupture elongation measured almost 70% against the base material. More and more evidence suggests that the key to successfully FSW high strength precipitation hardening aluminium alloys is to maintain the working temperature as high as possible, i.e. as close to the solidus, though undue melting or localised/incipient melting should be avoided (Colegrove and Shercliff, 2006, Yan et al. 2005, Rhodes et al. 1997). This necessitates, however, not just strict control of heating but also of cooling rates. Thus temperatures approaching the solidus will lower the alloy flow stress, induce increased slip between the workpiece and the FSW tool shoulder as appears in Fig. 3.8, without unduly compromising the pre-existing microstructure and hence mechanical properties. The weldable range, in terms of tool rotation and weld travel speed, which can be applied to high strength, precipitation hardenable 2xxx and 7xxx series aluminium alloys is much narrower than that of the 5xxx and 6xxx series aluminium alloys. The trend is typically concentrated more towards lower weld travel and tool rotational speeds. This is furthermore supported by the observed tendency that aluminium alloys with good hot extrudability also have wider weldability ranges. As with the case of extrusion the flow patterns resulting in the stir zones of friction stir welds will vary across the diverse aluminium alloy types, primarily in response to the material’s ability to resist deformation at the processing temperatures generated. Although flow features of stir zones produced when FSW high strength, precipitation hardenable 2xxx and 7xxx series aluminium alloys share similarities evidence from dissimilar alloy joints demonstrate that these stir zones are unique and share none of the features common to their single alloy counterpart welds. Additionally, as with the case of microstructural evolution, the flow features produced using a single tool but across diverse FSW machines will be dependent on the entire set of operating conditions encountered during processing, since these are responsible for cooling rates. As such, flow features contained within the stir zones of friction stir welds cannot be standardised and will never be directly transferable across diverse FSW machines simply because the FSW environment is unique to the particular machine responsible for producing the weld.
3.5
Concluding remarks
In concluding it can be stated that the deformation resistance of an aluminium alloy is the fundamental feature which determines its flow behaviour and
Material deformation and joint formation in friction stir welding
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therefore processing rate. In terms of FSW this will depend on the chemical composition of the alloy, its processing history (temper), on the force (strain rate) and temperature developed during processing, as well as the relative capacity to remove heat by means of the FSW environment. For a given working pressure and temperature there will always be a maximum amount of deformation that can be imparted to the workpiece material. In terms of processing temperature the limit is the solidus temperature where onset of melting begins, or if lower, the temperature where a substantial low-melting point constituent contained within the material, e.g. second phase particles, begin to melt. When this takes place the workpiece material in direct vicinity of the joining tool significantly loses flow strength. Subsequently slip occurs at the tool/workpiece interface, which not only changes the extent of the deformation taking place but also has implications regarding process efficiency, since the heat energy developed is no longer completely utilised or distributed for the formation of the stir zone. Optimised processing conditions necessitate that the material reaches a processing temperature which ensures thermal softening takes place and this also means that a degree of slip must occur between the FSW tool and the workpiece material. Very early on in the development of the FSW process TWI and its group of sponsors realised that tool coatings such as chromium nitride and titanium nitride enabled higher tool travel speeds and reduced tool clogging. In fact one of the clear visual indicators as to the quality of a friction stir weld is to look at the welding tool upon retraction from the workpiece. A dirty tool, whose profiles are significantly clogged, is a sure sign of a dirty weld with poor mechanical strength. This statement, however, should be tempered by the realisation that a completely clean and bright tool may also be the result of milling rather than FSW the aluminium alloy and hence no joint is formed. The key to any successful deformation process is the ability to achieve constant volume processing during deformation. In terms of FSW this simply isn’t just about generating a specific temperature for processing to occur. There are clearly a number of quantitative factors which combine and are involved in order to allow for optimal processing to be achieved. These include: (1) the initial temperature of the joining tool and its thermal conductivity (2) the contribution of heat generated due to plastic deformation (adiabatic shear) (3) heat generated due to friction at the workpiece/tool interface (4) the rate at which this heat is transferred, i.e. lost in terms of magnitude, direction and means (conduction, convection) from the deformation zone into the surrounding environment (5) the ability to maintain constant temperature gradients between the heat
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Friction stir welding
source and heat sink(s), i.e. this corresponds to the temperature difference potential or concentration potential that heat energy will flow from one region to another. Factors 1 to 5 above vary in regard to the material type being friction stir welded but also as a consequence of the FSW environment, i.e. machine spindle construction, clamping and backing bar employed to produce the joint. Hence if one is to make any sense of process property and weld formation behaviour, i.e. temperature and material flow relationships, it is essential that all of these factors be taken into account. A lack of such knowledge has led to enormous speculation concerning the FSW process and is the main reason why to date there exist so many and varied parameters for the FSW of any given aluminium alloy, and why the process operator is held captive to selecting a given set of processing parameters, based primarily on the geometry of the joint, construction of the workpiece restraint system and design of the FSW machine.
3.6
Acknowledgement
The author Rudolf Zettler wishes to acknowledge and thank RIFTEC GmbH, Max-Planck-Straße 2, Geesthacht 21502, Germany (E-mail:
[email protected]) where he was an employee at the time of writing Chapters 3 and 9. Rudolf is currently employed as an innovation consultant and can be contacted at the WTSH – Business Development and Technology Transfer Corporation of Schleswig-Holstein, GITZ, Max-Plank-Strasse 2, Geesthacht 21502, Germany (E-mail:
[email protected]).
3.7
References
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of friction-stir-welded 7075-T651 aluminum, Metallurgical Material Transactions A, 29A, 1998, 1955–1964, doi: 10.1007/s11661-998-0021-5. Masaki K., Sato Y.S., Maeda M. and Kokawa H., Experimental simulation of recrystallized microstrcture in friction stir welded Al alloy using a plane-strain compression test. Scripta Materialia, 58, 355–360, 2008 doi: 10.1016/j.scriptamat.2007.09.056. Midling O.T. and Rovik G., Effect Of Tool Shoulder Material On Heat Input During Friction Stir Welding. The 1st International Symposium on Friction Stir Welding, Thousand Oaks, Cal., USA, June 1999. Mills A.F., 1992, Heat Transfer, Irwin, International Student Edition. Nandan R., Prabu B., De A. and Debroy T., Improving reliability of heat transfer and fluid flow calculations during friction stir welding of dissimilar aluminum alloys, Welding Journal, vol. 86, 10, 2007, 313s–322s. Nandan R., DebRoy T. and Bhadeshia H.K.D.H., Recent advances in friction stir welding – Process, weldment, structure and properties. Progress in Materials Science 53, 2008, 980–1023. Peel M., Steuwer A. and Withers P.J., Microstructure, mechanical properties and residual stresses as a function of welding speed in aluminium 5083 friction stir welds. Acta Materialia 51 2003, 4791–4801, doi: 10.1016/S1359-6454(03)00319-7. Peel M., Steuwer A., Withers P.J., Dickerson T., Shi Q. and Schercliff H.R., Dissimilar friction stir welds in AA5083-AA6082. Part I process parameter effects on thermal history and weld properties. Metallurgical and Materials Transactions A 37/7, 2006a, 2183–2193. Peel M., Steuwer A. and Withers P.J., Dissimilar friction stir welds in AA5083-AA6082. Part II process effects on microstructure. Metallurgical and Materials Transactions A 37/7, 2006b, 2195–2206. Prangnell P.B. and Heason C.P., Grain structure formation during friction stir welding observed by the stop action technique. Acta Materialia, 53, 2005, 3179–3192, doi:10.1016/j.actamat.2005.03.044. Reddy N.S., Lee Y.H., Park C.H. and Lee C.S., Prediction of flow stress in Ti–6Al–4V alloy with an equiaxed a + b microstructure by artificial neural networks. Materials Science and Engineering A 492, 2008, 276–282, doi: 10.1016/j.msea.2008.03.030. Reynolds A.P., Flow visualization and simulation in FSW. Scripta Materialia 58/5, 2008, 338–342, doi: 10.1016/j.scriptamat.2007.10.048. Rhodes C.G., Mahoney M.W., Bingel W.H., Spurling R.A. and Bampton C.C., Effects of friction stir welding on microstructure of 7075 aluminium. Scripta Materialia, 36/1, 1997, 69–75. Russell M.J. and Shercliff H.R., Analytical modelling of microstructure development in friction stir welding. Proceedings of the 1st International Symposium on Friction Stir Welding, Thousand Oaks, Cal., USA June 1999. Schmidt H. and Hattel J., Modelling heat flow around tool probe in friction stir welding, Science and Technology of Welding and Joining, vol. 10, 2, 2005, 176–186. doi:10.1016/j. scriptamat.2007.10.008. Schmit H., Hattel J. and Wert J., An analytical model for the heat generation in friction stir welding. Modelling and Simulation Materials Science and Engineering, 12, 2004, 143–157, doi: 10.1088/0965-0393/12/1/013. Sellars C.M. and Tegart W.J.McG., Hot workability. Int. Met. Rev. 17, 1972, 1–24. Sheppard T. and Wright D., Deformation of flow stress: Part 1 constitutive equation for aluminium alloys at elevated temperatures. Met. Technol. 6, 1979, 215–223. Shercliff H.R., Russell M.J., Taylor A. and Dickerson T.L., Microstructural modeling
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in friction stir welding of 2000 series aluminium alloys, Mecanique & Industries 6, 2005, 25–35, doi: 10.1051/meca2005:004. Song M. and Kovacevic R., A Coupled Heat-Transfer Model for Workpiece and Tool in Friction Stir Welding. 4th International Symposium on FSW, Park City, UT, May 14–16, 2003. Srinivasan P.B., Dietzel W., Zettler R., dos Santos J.F. and Sivan V., Stress corrosion cracking susceptibility of friction stir welded AA7075-AA6056 dissimilar joint, Materials Science and Engineering A 392 (1–2), 2005, 292–300. doi: 10.1016/j. msea.2004.09.065. Su J-Q., Nelson T.W., Mishra R. and Mahoney M., Microstructural investigation of friction stir welded 7050-T651 aluminium. Acta Materialia, 51, 2003, 713–729, doi:10.1016/ S1359-6454(02)00449-4. Sutton M.A., Yang B., Reynolds A.P. and Taylor R., Microstructural studies of friction stir welds in 2024-T3 aluminum, Material Science and Engineering A323, 2002, 160–166. doi:10.1016/S0921-5093(01)01358-2. Thomas W.M. and Gittos M.F., Development of friction stir welding tools for the welding of thick (25 mm) aluminium alloys. TWI Research Report 692/1999, December 1999. Threadgill P.L, Friction stir welding – the state of the art, TWI Research Report 678/1999. Uzun H., Dalle Donne C., Argagnotto A., Ghidini T. and Gambaro C., Friction stir welding of dissimilar Al 6013-T4 to X5CrNil8-10 stainless steel, Materials and design 26, 2005, 41–46. doi:10.1016/j.matdes.2004.04.002. Vill V.I., 1962, Friction Welding of Materials, American Welding Society Inc., NY. Weman K., 2003, Welding Processes Handbook. New York, NY, CRC Press. www.keytometals.com/Article111.htm – Metallurgical factors affecting high strength aluminum alloy production. Knowledge Article from www.Key-to-Metals.com. Xu J., Vaze S.P., Ritter R.J., Colligan K.J. and Pickens J.R., Experimental and numerical study of thermal processes in friction stir welding. ASM Materials Solutions Conference and Exposition International Conference on Joining of Advanced and Specialty Materials VI, 2003. Xu S. and Deng X., A study of texture patterns in friction stir welds. Acta Materialia 56, 2008, 1326–1341, doi: 10.1016/j.actamat.2007.11.016. Yan J.H., Sutton M.A. and Reynolds A.P., Process-structure-property relationships for nugget and heat affected zone regions for AA2524-T351 friction stir welds. Science and Technology of Welding and Joining, 10/6, 2005, 725–736, doi: 10.1179/174329305X68778. Yan J.H., Sutton M.A. and Reynolds A.P., Processing and banding in AA2524 and AA2024 friction stir welding. Science and Technology of Welding and Joining, 12/5, 2007, 390–401, doi: 10.1179/174329307X213639. Zakharov V.V., Scientific aspects of deformability of aluminium alloys during extrusion, Advanced Performance Materials, 2, 1995, 51–66, doi: 10.1007/BF00711651. Zettler R., PhD Thesis, 2008, GKSS Forschungszentrum, Geesthacht, Germany. Zettler R., Lomolino S., dos Santos J.F., Donath T., Beckmann F., Lipman T. and Lohwasser D., A study on material flow in FSW of an AA 2024-T351 and AA6056T4 alloys. 5th International FSW Symposium, Metz, France 2004. Zettler R., Potomati F., dos Santos J.F. and de Alcantara N.G., Temperature evolution and mechanical properties of dissimilar friction stir weldments when joining AA2024 and AA7075 with an AA6056 alloy, Welding in the World, vol. 50, (11/12), 2006a, 107–116.
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Zettler R., Donath T., dos Santos J.F., Beckmann F. and Lohwasser D., Validation of marker material flow in 4 mm thick friction stir welded Al 2024-T351 through computer micro-tomography and dedicated metallographic techniques, Advanced Engineering Materials 8 (6), 2006b, 487–490. doi: 10.1002/adem.200600062.
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T. Zappia, MTS Systems Corporation, USA, C. Smith, Friction Stir Link, USA, K. Colligan, Concurrent Technologies Corporation, USA, H. Ostersehlte, Airbus, Germany and S. W. Kallee, Germany
Abstract: The achievement of high quality and consistent friction stir welds depends primarily upon the following three areas: ∑ Pin tool and weld schedule – the FSW pin tool needs to be designed in such a way as to properly mix the material for the given alloys, part thickness, and, if any are used, sealants. Closely coupled to the pin tool is the weld schedule that defines the critical weld parameters such as weld position, loads, spindle rotation, and travel speeds. ∑ FSW machine – the machine performing the welds must be able to control the critical weld parameters within an established range. ∑ Part tooling – the part to be welded needs to be accurately located and held in position during the welding process. In this chapter, each of these items will be examined to provide a better understanding of what is required by the FSW process and what is currently being used in the FSW marketplace. Key words: FSW equipment requirements, FSW machine requirements, FSW welding modes, fixed-pin, adjustable-pin, retractable-pin, self-reacting, bobbin tool, controller requirements, weld schedules, closed-loop control, robotic control.
4.1
Requirements of friction stir welding (FSW) coming from the process and applications
The physics behind the FSW process translates into certain demands that must be met by the equipment used to perform the welds. To start a weld, the FSW pin tool rotates at a predefined speed and then plunges into the part to be welded. Frictional heat is generated and the part material becomes plasticized and reaches a relative steady-state temperature. Once this plasticized state is reached, the pin tool is moved along a predefined weld path. The material which is displaced from the pin entering and traveling through the part attempts to extrude out of the pin hole, but is kept in place by the pin tool shoulder. This process is illustrated in Fig. 4.1. To determine the requirements for the FSW equipment, one needs to understand the range of performance that is required of the process for the 73
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Sufficient downward force to maintain registered contact
Joint
Welding direction
Torque
Leading edge of the rotating tool
Trailing edge of the rotating tool Retreating side of weld
Fx
Fy
Pin tool shoulder
Fz Pin tool tip Fx Fy Fz
Advancing side of weld
4.1 FSW overview. Figure courtesy of MTS Systems Corporation.
Weld cross-section
Friction stir welding
RPM
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parts to be welded. The four critical process parameters that are controlled by the FSW machine are the pin tool position, orientation, loads, rotation and travel speeds. The part tooling needs to be designed in such a way as to ensure appropriate fit-up, hold-down clamping, and stiffness to ensure that the part is held in place, as well as dissipate the heat being imparted from the process. Additionally, it’s important to take into consideration the end-use of the equipment. Is it a production or research application? What is the level of experience and knowledge required for an operator? What is the anticipated throughput for the system? How much automation is desired? Are there any pre- and post-processes to be integrated into the machine? The answers to these questions will have ramifications on the FSW equipment requirements. The following sections take a look at the requirements that generally define the FSW machine and part tooling.
4.2
Overview of the machine requirements for friction stir welding (FSW)
There are many different types of machines used for FSW. Some are configured for a specific application and others are of a more general configuration that allows them to weld a broader range of parts. The details of how a machine is designed dictate the range of applications that it is capable of welding. For example, production applications generally use FSW machines that are designed specifically for the product’s work envelope (i.e., area where the FSW machine is able to weld) and associated weld functionality and parameters (i.e., weld type, torque, RPM, loads, travel speeds). Emphasis can be put into streamlining the design so that only the absolute required attributes are integrated into the system. For research institutes and research and development groups, FSW machines will generally have diverse functionality built into the design so that a wider range of applications can be welded. At the most basic level, the requirements for an FSW machine need to include the functionality and performance required to generate the desired welds. A good starting point for specifying requirements is to decide on the type of welding that is going to be performed; fixed-pin, adjustable-pin or self-reacting. With fixed-pin welding, shown in Fig. 4.2, the FSW pin tool is one piece consisting of both the shoulder and pin. Fixed-pin welding is the most traditional form of FSW and is the easiest to manage from a machine design and control perspective. Any motion of the weld head spindle translates into an associated change in the position and loads of the shoulder and pin. With adjustable-pin welding, shown in Fig. 4.3, the FSW pin tool consists of two pieces: a pin and a shoulder, which are able to move independently of one another. This method of welding allows for more flexibility in the welding of parts. For example, tapered parts (i.e., parts with varying thickness)
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Friction stir welding Fixed pin
Tool section of friction stir weld
Spindle
Welding direction
FSW pin tool
Pin motion
4.2 FSW fixed-pin mode. Figure courtesy of MTS Systems Corporation.
Adjustable pin Tool section of friction stir weld
Spindle
Welding direction Upper shoulder Adjustable pin Pin motion
4.3 FSW adjustable-pin mode. Figure courtesy of MTS Systems Corporation.
can be welded and the pin ligament (i.e., distance of the pin tip to the back of the weld) can be maintained while the shoulder remains on the surface of the part (see Fig. 4.4). Additionally, adjustable-pin can be used to closeout the pin hole that exists when exiting a fixed-pin from a weld (see Fig. 4.5). From a machine design perspective, adjustable-pin welding requires a
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Spindle
Weld direction
Adjustable pin
4.4 Adjustable-pin maintaining pin penetration. Figure courtesy of MTS Systems Corporation.
4.5 Adjustable-pin closing out pin hole.
somewhat more complex machine design and control scheme. The machine needs to have the ability to move the pin and the shoulder independent of one another, including potentially at different rotation directions and/or speeds. This can be done by independent actuation of the two axes within the weldhead, or it can be achieved by controlling the pin motion within the weldhead and the shoulder by manipulation of the tool Z axis. With self-reacting welding, shown in Fig. 4.6, the FSW pin tool consists of three pieces; an upper shoulder, pin, and a bottom shoulder. From a machine design and control perspective there are really only two pieces to control as the pin and bottom shoulder are attached to one another. The upper and bottom shoulders remain on the top and bottom surfaces of the part as the pin moves through the entire cross-section of the material. Self-reacting welding has the advantages of guaranteeing penetration and allows for the possibility of a reduction in tooling as the bottom shoulder is used to selfreact the upper shoulder loads. Similar to adjustable-pin, two separate axes need to be controlled, one each for the upper and lower shoulders. Beside the welding mode, other fundamental requirements for the FSW
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Friction stir welding Self-reacting pin Tool section of friction stir weld
Spindle Welding direction Upper shoulder Adjustable pin Lower shoulder Pin motion
4.6 FSW self-reacting mode. Figure courtesy of MTS Systems Corporation.
machine include the loads, torque, travel rates, and work envelope. While the work envelope is readily defined by the parts to be welded, the correlation of loads, torque, and travel rates to machine requirements is not as straightforward. The reason is that these requirements are dependent upon the interrelated variables such as alloys to be welded, part geometry, joint type (e.g., butt, lap), pin tool design, weld mode, and weld schedules. Loads and torque are directly impacted by the tool design (e.g., shoulder diameter and pin and shoulder features) and the spindle RPM and travel speeds. A similar quality weld measured by UTS or fatigue strength can be achieved using different pin tool designs and weld schedules which will have different resulting loads and torques on the system. Similarly, the alloy type and weld mode (i.e., fixed-pin, adjustable-pin, or self-reacting) impact the loads, torques, and other machine requirements. Given this multi-variable relationship between the process and resulting machine requirements, machines are typically specified to perform within a range. A general relationship is thicker parts require higher torque, lower RPM, and lower travel speeds. Thinner parts require lower torque, higher RPM, higher travel speeds. Table 4.1 lists some known relationships between these parameters.
4.2.1 Work envelope The FSW machine work envelope can be determined by taking the part geometry, the associated weld seam, and the orientation that the part is to
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Table 4.1 Process-to-machine requirements relationship Weld type Alloy
Thickness
mm
in
APT/FPT DH36 Steel 6.35 0.250 HSLA 65 SS 6.35 0.250 1.20 0.047 4.00 0.157 2024 4.83 0.190 9.53 0.375 5182-O 1.20–2.50 .047–.098 5186 6.00 0.236 6056 4.00 0.157 6056 6.00 0.236 6061 4.83 0.190 6061 6106/6005 4.00 0.157 SRPT AL6XN 6.35 0.250 DH36 6.35 0.250 2024 4.00 0.157 2195 8.13 0.320 2219 8.13 0.320 2519 8.13 0.320 6061 4.00 0.157 8.13 0.320
Spindle RPM
Torque Nm
in-lbs
350 788 300 450 850 2000 600 350 215 2400 400 600 900 600 225 1250
203 102 203 113 10 13 30 113 181 20 45 50 30 54 136 10
1800 900 1800 1000 89 115 266 1000 1600 177 398 443 266 475 1200 89
200 350 300 230 750 230 750 225 230 230 460
228 181 52 198 60 192 68 193 113 51 85
2020 1600 460 1750 529 1700 600 1707 1000 450 753
be held while welded, and then accounting for any spindle extension, the length of the FSW pin tool, and the part tooling and clamping mechanism. The resulting work envelope area is used to determine the number of machine axes and associated strokes for the machine. Two-dimensional applications are normally achieved with two or three translational axes (i.e., X, Y, and Z). For three-dimensional applications, two additional rotational axes are required to provide pitch, roll, or yaw (i.e., A, B, or C axes). These rotary motions can be achieved at the weldhead or through a tilting turntable. In order to maximize the work envelope area and allow for the largest range of applications, it is usually preferred to have the rotational axes integral into the machine near the weldhead rather than from the turntable. Turntables can be a nice solution for dedicated circumferential welds and, depending on the turntables size, specific smaller parts. The turntables size can be a drawback as it can take a very large part of the machine’s work envelope
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and usually have a small work surface that can place significant limits on the dimensions of the workpieces that can be mounted to it. Care needs to be taken when designing3-D systems to make sure that the machine motion won’t have any machine singularities or kinematic accelerations that will prohibit the machine from executing the weld paths. Models of the machine configuration can be made along with simulations of part trajectories to determine whether these situations will be an issue and help determine what can be done to overcome them. Typically axes can be reconfigured, part locations can be modified, and different control techniques can be used to eliminate these problems.
4.2.2 Required sensors Another key part of the system design has to do with the transducer selections and their location on the machine. With servo-controlled systems the accuracy and resolution of the sensors, along with the responsiveness of the drive system, play a major part in determining the accuracy and fidelity of control for the machine. Table 4.2 lists the various axes that usually make up an FSW machine and the types of sensors that are generally used. Whenever possible, high resolution absolute position sensors are preferred due to their accuracy, repeatability, and simplicity to use for the control system. However, relative position sensors can be used, but require a “homing” operation to locate the axis on startup. This “homing” operation, commonly used on milling machines, consists of moving the axis until a known position
Table 4.2 Example sensor-to-axis types Axis
Sensor types
Translational axes position Glass scale. Motor encoders, LVDT Rotational axes position Glass scale, encoders Pin and forge position Glass scale, LVDT, capacitance Translational axes loads Load cells or pressure cells Rotational axes loads Load cells or pressure cells Side loads (i.e., loads I along the weld path Load cells or pressure cells and perpendicular to the weld path – commonly referred to as tool X and Y) Pin and forge load Load cell or pressure cells Spindle rotation Encoder pulse modulation Spindle torque Torque sensor, load cells, or pressure cells Seam-tracking Camera or LVDT Stand-off (measurement of the distance of Glass scale, LVDT, capacitance, the shoulder to the material – used to laser interferometer eliminate the effect of spindle and machine thermal expansion)
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is reached and a discrete sensor is located which signals the controller the exact location of the axis. With all sensors, it is preferred to locate the device in such a way that it is able to precisely and as repeatably as possible map the position, motion, or load that is occurring on the machine.
4.2.3 Required accuracies and repeatability The performance of FSW machines, like CNC machines, is defined in large part by their required accuracies and repeatability. That is, how accurately the machine can move to and along commanded trajectories, and how repeatable the movement is from one time to the next. Accuracy is generally measured by an independent measurement device (e.g., laser tracker, interferometer) and done relative to the world coordinate system (i.e., linear and angular positions of the machine based on predetermined reference points in the work cell area). Many factors can affect machine accuracy. This includes design decisions, manufacturing tolerances, assembly and installation techniques, wear, structural deflections, temperature, and the control system. As Jim Destefani points out in his paper titled ‘Getting a handle – on machine accuracy’1, “There are six possible errors per linear axis: linear positioning, pitch, yaw, vertical straightness, horizontal straightness, and roll. Adding in the orthogonal errors that can occur between the mutually perpendicular axes gives a total of twenty-one possible errors that can occur on a threeaxis machining center.” Machine builders strive to minimize accuracy error stack-up starting with the design. Analysis is done to make sure that the system is sized appropriately for the anticipated loads and torques. Stiffness and strength are designed into the appropriate parts of the machine, with the larger structural parts being designed to minimize deflections and the parts that experience stress being designed for strength. The system drive mechanisms are chosen such that the various components (e.g., guides, rails, bearing, gears, ballscrews, motors) are all appropriately sized with minimal errors resulting from things like resonant modes, backlash, friction, and windup. Once a machine has been designed and manufactured, it needs to be assembled in such a way as to achieve the best possible alignment. Each axis should be measured for straightness, flatness, parallelism, and squareness to orthogonal axes, and appropriate adjustments made until acceptable errors are reached. Not only is this important for achieving accuracy, but it can also have an impact on the life of the drive components as inaccurate alignments can cause high wear and stress to the parts. Once the machine is assembled, the various position sensors are calibrated to match the machine motion. At this point the system is going to be as mechanically and electrically accurate as it can be. Any inaccuracies that have to do with deflection, thermal expansion, etc., will need to be compensated
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by the control system. Based on MTS’ experience the “typical” type of accuracies that can be found on standard milling machine designs is shown in Table 4.3. In specifications, accuracy requirements are usually defined as a percentage of the commanded value while taking into consideration the full-scale range of the machine. For example, if the maximum load on a machine is 30 K lbs, the load accuracy could be specified as +/–1% of the commanded value or 100 lbs, whichever is greater. The 100 lbs would be used as a reference point that takes into consideration that machines are designed to perform over the full range of their performance and are going to be limited in their performance due to the size of that range. The reason for this is that the larger the load requirements, the larger the drive components will need to be along with their larger associated errors. In general, the axes that require the most accuracy are those closest to the welding process such as the FSW pin position and load sensing.
4.3
Friction stir welding machine controller requirements
The other major piece of an FSW machine that needs to be defined is the control system, which includes both the electronics architecture and software application. The requirements associated with a production system versus a research system can be significant. With production systems, more emphasis is typically placed on simplification of system design in order to maximize system up-time, ease of use, and synergy with existing control systems that may already be installed and for which there are trained support staff and spare parts inventory. For those doing research, emphasis tends to be on making sure the system has a wide enough functionality range to support the various areas of research. Depending on the research, they may require enough processor throughput to enable future FSW algorithm developments, ensure high speed data acquisition to study signal content and correlations of sampled data, and expansion capability for any future sensor and control scenarios that their research may uncover. The basic requirements for either scenario, production or research, will include the following abilities.
Table 4.3 Typical machine tool accuracies Axis
Approximated accuracy
X (> 4 meters of stroke) Y (2 meters to 4 meters) Z (< 2 meters)
0.076 mm (0.003 in.) 0.038 mm (0.0015 in.) 0.025 mm (0.001 in.)
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Create and execute a weld schedule The system must allow for the operator to enter weld specific commands for a given weld path trajectory. This generally includes variables or inputs for: ∑ rotation speed ∑ travel speed ∑ welding axial load (if in load control) ∑ welding traverse load (optional and requires traverse load measurement system) ∑ travel or tilt angle (see Fig. 4.7) ∑ work angle (see Fig. 4.7). Additionally, the system must allow specific FSW start and stop commands that typically include the following variables or inputs: ∑ ∑ ∑ ∑ ∑ ∑ ∑
nominal plunge depth (start) nominal plunge load plunge speed (start) dwell time after plunge and before traverse (start) dwell time at end of traverse and before retract (end) retract speed (end) exit path offset (end).
The following five points describe the main controller requirements. Perform coordinated motion control The controller must be able to perform multi-axis control that allows the system to follow the desired weld path while executing the various weld commands. Perform position and/or load control The standard means for controlling the FSW process is by position and/ or load control. Position control is the most straightforward method as the weld engineer only needs to program the machine to follow known position
Travel angle
Travel direction
4.7 Travel and work angle description.
Work angle
Travel direction into page
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trajectories. However, owing to machine and part clamping deflections, along with part tolerance error, thermal compensation, position control will not always reliably achieve the desired welds. Load control has the advantage of being able to compensate for the above mentioned inaccuracies as the control system will move the spindle (pin and/or shoulder) to achieve the desired loads throughout the weld. Side load and torque control are variants on load control as the controller uses these parameters to make necessary corrections to maintain the desired load or torque profile. Provide control necessary to support type of welding The controller needs to be able to provide the kind of control necessary to support the weld modes that are going to be performed (i.e., fixed-pin, adjustable-pin, and self-reacting). If required, these weld mode types can be performed in position and/or load control. Monitor process and system feedback The controller needs to be able to monitor and display process and system feedback and take appropriate action should any parameters move outside of a predefined range. Perform data acquisition Data acquisition of key process parameters and machine variables provides a means of evaluating that a satisfactory quality weld has been achieved by the system. Post weld analysis can be done to make sure that these parameters and variables have been maintained within predefined ranges as a quality assurance measurement (see Chapter 7).
4.4
Closed-loop control and friction stir welding (FSW)
For those who may not understand closed-loop control, Fig. 4.8 is a controller diagram for a simple single-axis system using inner and outer proportionalintegral-derivative (PID) control loops. The outer-loop takes the program command (e.g., position or load) and axis feedback and sums the two together to create an error. This error is corrected according to the outer-loop PID function (see Fig. 4.9) and then output to the inner loop servo controller. The inner-loop provides the lowest level of control for the hydraulic servovalve or electric motor, and generally runs at a much higher rate than the outer control loop. The inner-loop works similar to the outer-loop in that it sums the command sent from the outer-loop with feedback from the device being
Outer loop PID gains
Summing junction
Summing junction Command from programmer
DC error
+
Inner loop PID gains
PID
+ Valve command
–
–
PID
Multi-stage servovalve
Inner loop Servovalve spool LVDT transducer conditioner LVDT
Feedback Transducer conditioner Actuator
4.8 Controller diagram with inner and outer control loops. Figure courtesy of MTS Systems Corporation.
Friction stir welding equipment
Outer loop
Valve error
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Friction stir welding Kpe(t)
P
Setpoint
+
t
S
Error
I
Ki
Ú e(t)dt
S
Process
Output
0
–
D
Where: P I D
Kd
de(t) dt
Kp: Proportional gain, a tuning parameter e: Error = SP – PV t: Time or instantaneous time (the present) Ki: Integral gain, a tuning parameter e: Error = SP – PV t: Time in the past contributing to the integral response Kd: Derivative gain, a tuning parameter e: Error = SP – PV t: Time or instantaneous time (the present)
4.9 PID diagram.2
controlled (e.g., electric motor encoder). Again, this error is scaled according to inner-loop PID tuning gains and output to the drive device. With this fundamental controller understanding, Fig. 4.10 shows a functional diagram of a typical FSW controller for a 3-axis positioning system and weldhead. The machine has a spindle, X, Y, and Z drive system that includes individual axis motor and feedback position transducers. An example of what the internal data flow for the above control architecture might look like is shown in Fig. 4.11. a weld schedule is created that contains the machine motions required throughout the weld. This includes all the motion required for the various stages of the weld like plunge (start of weld) and retract (exiting the weld) and any parameter variations that are to be made during the weld (e.g., change in travel speed or spindle rotation speed). The controller takes the weld schedule and generates the associated specific axes commands that are to be sent to the machine. For 2-D welds this operation can be as simple as a straight line interpolation using appropriate acceleration and deceleration rates. Servo control is performed as described above and the appropriate commands are sent to the drive system (e.g., electric motor or hydraulic servovalves). The control system will monitor the various commands and feedbacks of the system and generate any operator messages, alarms, or shutdown the system according to predefined range settings. Data is also gathered at some defined sampling rate and stored on a hard disk for later evaluation.
User interface
Controller
∑ PAC (programmable automation Spindle X Axis Y Axis Z Axis controller) drive drive drive drive ∑ CNC (computer numeric controller) ∑ Custom motion controller ∑ Robotic controller
4.10 Simple control system architecture. Figure courtesy of MTS Systems Corporation.
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∑ PLC (programmable logic controller)
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Weld program ∑ Gantry commands ∑ FSW commands
Program function generator Individual axes command generation
Commands Gantry axes Spindle
Data file
Control servo ∑ PID ∑ Mode servo Feedback conditioners ∑ Encoder cards ∑ Strain gage conditioners DAQ
Limit detection ∑ Command ∑ Feedback ∑ Discrete inputs
4.11 Simple FSW controller data flow. Figure courtesy of MTS Systems Corporation.
Discrete I/O ∑ Overtravel switches ∑ On/Off switches
Friction stir welding
Output conditioners ∑ D/A modules ∑ S/V driver modules
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Depending upon the end use of the system and the level of complexity in the application, the controller type can be one of many solutions. The types generally range from programmable logic controllers/programmable automation controllers, CNCs, custom motion controllers, and robotic controllers. The main differences between these options include the processor speed, built-in functionality, and data acquisition capabilities. The PLC and PAC (e.g., GE Fanuc, Allen Bradley, Siemens) are the most common type of controller used for industrial low complexity motion control. They are usually modular in design with dedicated modules for specific tasks which can be configured to carry out relatively complex tasks. With the constant improvement in processor speeds, the line between PLC/PACs and CNCs is blurring for good effective axis positioning solutions, but PLC/PACs can fall short when complex motion and interpolation is required. CNC systems (e.g., Siemens 805 and 840D, GE Fanuc Series 3Xi models, Bosch Rexroth Indramotion MTX,) are used on most machine tools that require accurate positioning with complex trajectories that are usually generated with CAD/CAM software. Custom motion control systems are used where high performance positioning is required for complex motion profiles for multiple axes. With FSW, custom motion controllers have an added advantage, albeit at a higher price, of being configured to exactly meet the applications requirements. An example of a significant difference between a CNC used for milling machines and FSW is with the importance of controlling the weld travel speed (i.e., feedrate). Travel speed is a calculated variable that results from the derivative of the axes position. With FSW this is an important variable to control as it has direct effect on the frictional heat that is being put into the part. Slowing down can add unwanted heat and speeding up can move the FSW pin tool into areas where there is not enough heat. Milling machines are able to trade off travel speed and accuracy at certain parts of the machining operation. For example, when going over a tight radius on a finishing cut, it is acceptable to slow down the feedrate to maintain the accuracy through the radius. An FSW controller must maintain the accuracy and the feedrate through the radius. While it is not the purpose of this chapter to do a comparison of the different controllers, we will make the general statement that PLCs and PACs are on the lower end of functionality and performance, while CNCs, custom motion controllers, and robotic controllers are on the high-end. Table 4.4 lists some of the advantages and disadvantages between the controllers.
4.5
Control of robotic friction stir welding (FSW)
To implement an appropriate control scheme on a robotic-based FSW system, several unique control solutions are required. These control solutions are based on the characteristics or aspects of robots versus typical custom build
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Table 4.4 Controller advantage/disadvantage comparison Controller type
Advantages
Disadvantages
Programmable ∑ Low cost logic controller ∑ Well-established install base ∑ (PLC) with associated large number Programable of support capabilities ∑ automation ∑ Industrial hardened. ∑ controller (PAC) ∑ ∑
Relatively low processor speeds Not multi-processing Not well-suited for multidegree of freedom systems Limited expandability Limited number of channels and sampling rate for data acquisition.
CNC ∑ ∑ ∑ ∑ ∑ ∑
Limited processor speeds relative to high-end digital signal processors Limited expandability Difficult to customize for FSW special functionality (e.g.,seam-tracking) Limited number of data acquisition channels and sampling rate for data acquisition.
Moderate cost ∑ Multi-tasking Capable of performing 3-D multi-axis control ∑ CAD/CAM interface for easy ∑ path trajectory generation Well-established install base with associated large ∑ number of support capabilities Industrial hardened.
Robotic ∑ Allows low cost robots to ∑ controllers be used for FSW ∑ Moderate cost for controller ∑ Well accepted in production ∑ environment ∑ Multi-tasking ∑ Teach mode ∑ ∑ Industrial hardened. ∑ ∑
Adaptive techniques must be implemented to account for low robot stiffness Limited processor speeds relative to high-end digital signal processors Limited expandability Difficult to customize for FSW special functionality (e.g., seam-tracking) Limited number of data acquisition channels and sampling rate for data acquisition.
Custom motion ∑ controllers ∑ ∑ ∑ ∑ ∑ ∑
High cost Not well accepted in production environment Limited number of people able to support applications.
Multi-tasking ∑ Capable of performing 3-D ∑ multi-axis control CAD/CAM interface for easy ∑ path trajectory generation Easily expandable Easy to customize for FSW special functionality High number and sampling rate for data acquisition Industrial hardened.
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machines. The aspects of robotic-based machines that must be considered in the development of the control system include: ∑
Traditional industrial robots are relatively less stiff than custom built machines. ∑ Traditional industrial robots are repeatable, but not accurate. ∑ Industrial robots have very limited CAD/CAM capability. Robot paths are traditionally taught by an alternative mechanism. Robots are typically jogged to the desired position using the teach pendant. The operator or programmer then indicates that the current position is the desired position along the path. All of the above present several challenges for the control system that must be managed in any application software for FSW. All of the aspects of robots that are listed above can be managed with the development of application software specific to friction stir welding. The application software must have the following general capability (all of which are described in detail later in this section). ∑ Data structures specific to friction stir welding. ∑ Force control. ∑ Real-time editing of vertical position/welding force and position transverse to welding direction. ∑ Ability to provide offsets to programmed paths. ∑ Ability to automatically program robot orientation. ∑ Allow for a teach mode and an automatic mode. ∑ Ability to store welding data for future data analysis.
4.5.1 Data structures As with all types of FSW controllers, the robotic controller must allow for the normal type of FSW functions that let an operator communicate the desired input variables of the process to the machine in a succinct and easy to understand method. The required input variables on a robotic system are those listed earlier when discussing the creation and execution of a weld schedule. As such, data structures are created to support the start/stop operations and the steady-state welding operations. Both the start/stop and welding data structures allow for definitions of plunge depth, force, speed, rotation and travel speed, axial and traverse forces, and tilt angle, etc. The main exceptions for the robotic controller are that there are two additional commands required for the start/stop process. There is an additional plunge after the start and an exit path offset. Each of these variables is used to manage the relative lack of stiffness of typical industrial robots. A robot’s deflection is generally proportional to the force on the robot. Secondly
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FSW requires a different level of force during the plunge operation than when traversing. This means that the robot must artificially be commanded to plunge further during the traverse. This is overcome by allowing for an additional input (additional plunge after start of traverse). The exit path offset is required for similar means, but when the forces are relieved on the robot at the end of the weld, the robot tends to over-shoot. The exit path offset is an artificial offset that is used to prevent the robot from overshooting, to allow for a clean exit path.
4.5.2 Force control There are several potential means by which to implement force control that are dependent on the machine characteristics. Industrial or articulated arm robots generally lack stiffness compared to many machines. Robots tend to mimic a spring with load applied, which means they tend to deflect linearly with applied load. Thus, a change in load can be applied by changing the programmed position (not actual position) of the robot. If the robot is programmed to plunge deeper, the larger effect will be to increase the downward force rather than the actual position. This allows a strategy where a PID control loop can be implemented to control the desired force level by applying offsets to the programmed position.3, 4 On a robot, the mechanical stiffness varies throughout the work envelope of the robot. For example, the robot has greater stiffness near the base of the robot versus the robot being stretched to the limit of its working envelope. Thus, force control can be an important feature of the software that can allow the operator to quickly manage the relative stiffness or deflection of the robot under the prescribed friction stir welding loading conditions.
4.5.3 Real-time control of force/position Because a robot has relatively low stiffness, the programmed position varies from the actual position. The basic force/position control concept is shown in Fig. 4.12. There is a desired position and orientation (Pd and Rd) that is adjusted to achieve a desired force (fd) based on position and force feedbacks. This type of control provides a challenge for implementing friction stir welding on a robot as it results in a condition where the programmed plunge depth will never result in an actual plunge depth that meets the programmed value. If consistent, this could be learned by the operator, but the variation between the programmed plunge depth and the actual plunge depth varies throughout the work envelope of the robot, since the stiffness of the robot varies throughout its work envelope. To help overcome this problem, it is important that the operator have real-time control of the programmed plunge
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Trajectory planner P
R
Position and orientation control
Motion control
Manipulator/ environment
q q f
Dp fd
Force control
4.12 A force/position hybrid control system chart having a desired position, orientation and force input and a joint position, velocity and measured force output.5
depth to determine what value will provide an acceptable weld result. If the operator can preliminarily enter an anticipated plunge depth and then apply an offset during initial welding trials, the final programmed plunge depth can be determined relatively quickly. The alternative is to perform numerous welding trials with various programmed depths to eventually determine an appropriate programmed plunge depth. This real-time control of plunge depth is accomplished by allowing the operator to press buttons that increase or decrease the plunge depth during welding. This allows the robot to overcome variable and unknown deflection of the robot versus work envelope and weld parameters. The real-time control should allow for vertical as well as transverse position control, since the robot can deflect in both the vertical and horizontal planes due to radial and axial forces generated by the process.
4.5.4 Automatic application of program offsets A robot is programmed by jogging the robot to a position and orientation, followed by depressing a button to store the current position as a programmed position. This poses a challenge for friction stir welding, as the actual required position during welding is such that the FSW tool is partially within the part. When programming the position, the FSW tool is stationary (not rotating) so it is impossible to teach a position with the FSW tool rotating and with the FSW tool pin in the part. The best that can be done is to teach a position with the FSW tool pin in contact or near contact with the weld surface. In this case, the actual desired position is an offset in position that is equivalent to the length of the pin. This offset is relatively simple to apply when the weld is in the horizontal plane. However, it is significantly
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more complex when the weld path changes orientation during welding. This generates a requirement that any FSW application software have the capability to automatically apply program offsets that are independent of orientation of the weld path.
4.5.5 Automatic application and programming of travel and work angles Friction stir welding requires relatively precise control of the travel and work angles. The travel and work angles were previously described in Fig. 4.7. As noted above, robots are programmed by jogging the robot to a desired or known position, followed by pressing a key to store the current position. This poses challenges for accurately programming position, especially the orientation. The orientation is made up of three components: roll, pitch, and yaw. Not only is it difficult to accurately or precisely control the robot in the jogging mode to teach a position, but it is generally impossible to measure the roll, pitch, and yaw angles precisely with respect to a complex surface. Thus, any FSW control software must have means of entering and automatically calculating the appropriate roll, pitch, and yaw angles to allow the correct work and travel angles to be applied.
4.5.6 Teach and automatic modes Industrial robots generally have two modes of operation. The first is a teach mode and the second is an automatic mode. In teach mode, the operator generally has control of the robot through the use of the teach pendant and can have significant input to any operation. In the automatic mode, the robot typically performs pre-programmed paths with little ability to affect the robot in real time. The robot can also operate at maximum capable speed in automatic mode, whereas the speed of the robot is typically limited in teach mode. For friction stir welding, it is desirable for the operator to have some real-time input into the process, especially when teaching and developing initial paths and weld processes. In this teach mode, the operator should have ability to adjust the welding force or vertical position, as well as the transverse position to help manage the lack of stiffness of the robot. In the automatic mode, this real-time editing capability is generally eliminated, since it is not required once the process is setup. Given industrial robots are typically repeatable, but not accurate, any deflections that occur tend to be consistent. Thus, once the process is set up, little intervention should be required.
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4.5.7 Data monitoring and storage Like most FSW machines, data monitoring and storage is important for roboticbased machines. In a production mode, data monitoring can be used to set alarms with respect to actual status of FSW process variables. Variables that are typically monitored include force, torque, position, rotation speed, etc. Alarms can be set to warn and/or abort the process. This can be especially useful in production applications to detect abnormal conditions. The process variables can also be stored. This is especially useful in process qualification or development exercises. However, data storage is limited in production applications, due to the relatively limited data storage capability of robots. Data storage is limited due the limited amount of memory that is typically available with robotic controllers. In addition, robotic controllers are not optimized for data storage, so rates of data storage can be expected to be relatively slow. Typical data sampling and storage rates are less than 20 Hertz.
4.6
Other controller requirements
FSW applications that need 5-axis motion for 3-D and complex contour trajectories and have the requirement for very precise control of forces and motion often require more advanced controller configurations. Figure 4.13 shows the different types of functionality that can be integrated into the controller and what the associated data flow might look like. Included could be the following functionality for 5-axis control with various compensation techniques to improve welding accuracy and performance: ∑
Multi-axis kinematic and coordinate transformations – these transformations allow for 3-D trajectory generation along a weld path while maintaining a defined FSW pin tool orientation angle (usually 0 to 3 degrees). The machine can be controlled in either the world or part coordinates system. ∑ Volumetric compensation – a measurement device (e.g., laser tracker) is used to exactly locate the FSW pin tool throughout the work envelope. An error mapping is made which the controller uses to compensate for positioning errors due to mechanical inaccuracies and gravitational effects (e.g., fully extended heavy Z axis on a horizontal boring mill). ∑ Deflection compensation – deflection measurements are taken at different locations in the work envelope using different loading scenarios. A deflection compensation mapping is made and used by the controller to adjust position commands in order to minimize the effects of load induced deflections. ∑ Seam-tracking – seam-trackers can be used to monitor the location of the weld seam in real time and automatically make cross-seam adjustments to ensure proper pin location relative to the center of the seam.
Control servo ∑ PID ∑ Mode servo
Commands Gantry axes Spindle
DAQ
Data file
Advanced compensation ∑ Volumetric compensation ∑ Deflection compensation ∑ Thermal compensation ∑ Seam-tracking ∑ Surface sensing
Feedback conditioners ∑ Encoder cards ∑ Strain gage conditioners
Limit detection ∑ Command ∑ Feedback ∑ Discrete inputs
Loads Seam offsets Distance to surface Component temperatures
4.13 Complex FSW controller data flow. Figure courtesy of MTS Systems Corporation.
Discrete I/O ∑ Overtravel switches ∑ On/Off switches System and process sensors ∑ Load cell or pressure cells ∑ Seam-tracker ∑ Surface sensor ∑ Temperature sensors
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Program function generator ∑ Kinematic and coordinate transformation ∑ Multi-axis command generation ∑ Volumetric compensation ∑ Deflection compensation
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Weld program ∑ Gantry commands ∑ FSW commands
Output conditioners ∑ D/A modules ∑ S/V driver modules
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∑
∑
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Surface sensing – even with very accurate machines and stiff tooling fixtures, it can still be difficult to know exactly where the part surface is. This is primarily due to machine and tooling deflections and thermal expansion in the weldhead spindle and FSW pin tool. The best way to know exactly where the part surface is relative to the weldhead is to measure it in real time from a reference point as close to the surface as possible where the effects of deflection and thermal expansion can be accounted for. This information can then be used as a reference for adjusting position commands. Thermal compensation – ambient temperature variations, along with process induced heat, can cause thermal expansion of the machine that will lead to positioning and load inaccuracies. If the application requires high precision accuracy and the reduction of this type of error, temperature sensors and thermal compensation tables can be used to adjust position commands appropriately.
4.7
Other machine requirements
Some other machine requirements that are not directly related to the FSW process, but will impact machine design, include the following: ∑
Test and calibration stands – often for production system applications a test and calibration stand is used to verify the equipment is functioning properly prior to welding. The control system will perform an automated sequence that tests the system’s ability to perform certain motions and loading scenarios that verify all system components are working properly. In some demanding aerospace applications, a test panel will be welded and fatigue or tensile tests run to verify performance prior to, after, and even during a production run. ∑ Weldhead to perform milling operations – though FSW weldheads are not specifically manufactured for doing milling operations, they have the built-in functionality to perform many milling operations. Some systems will require that the FSW weldhead be capable of performing such operations as drilling and cutting. ∑ Viewing systems – it is very typical to see system requirements include vision systems that are able to record the welds, preferably from both the front and back side. This provides a means for the operator to view the weld and make process adjustments from the operator station rather than being next to the weld. In aerospace production environments the recordings will be stored and saved with the other monitored data providing a complete record of how a particular part was manufactured. ∑ Safety – machine design standards provide guidelines to ensure that safety factors are built into the machine. Components should be selected to
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∑
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ensure that they are sized properly for the application and the expected life-cycle of the machine, care needs to be taken to design in features that reduce the risk of injury where there is the potential for operator interaction with the machine. Guard rails with e-stop triggers may need to be installed at access points, pressure pads may be used, and guard rails and ladders all need to be designed to follow the local safety requirements. E-stop buttons should be easily accessible and will immediately shutdown the system and are to be designed as electromechanical circuits with no chance of software intervention. Maintenance – is an important area of machine design as many components require scheduled maintenance. The design needs to ensure that there is adequate access to the various components to allow service engineers to monitor wear, provide routine maintenance (e.g., grease bearings), and remove and replace components if necessary.
4.8
Machine requirements summary
The ability to consistently weld quality FSW parts is dependent upon the FSW pin tool and weld schedule, the FSW machine, and the part fixturing and clamping system. These items are largely dependent upon each other. That is, all three must be done well, or there is a good chance that the resulting welds will not be satisfactory. Inherent to the FSW process are certain loads and torques that are the primary drivers to the FSW machine and part tooling equipment. Machine tool builders pay particular attention to making sure that there is enough stiffness in the machine and tooling to react the loads and torques to within the accuracy requirements set forth by the application. Key to an FSW machine is its ability to perform the desired type of weld: fixed-pin, adjustable-pin, and self-reacting. The weld types, along with the work envelope, are used to determine the number of machine axes and configuration. The FSW machine is made up of drive components and sensors that are selected so that the machine is able to move accurately and repeatably within the work envelope over the required range of performance. The FSW control system is closely coupled to the drive components and feedback sensors and allows the operator to create and execute a weld schedule. It also provides for process monitoring and data acquisition of the critical weld parameters. The types of FSW control systems include PLC, PAC, CNC, robotic controller, or customer controller. Each has its strengths and weaknesses and should be selected by taking into consideration factors such as the complexity of the application, the type of work that needs to be performed (i.e., research and development or production), data acquisition requirements, and the expected maintenance and service requirements for the system.
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Part tooling requirements
The requirements for part tooling are relatively straightforward; the tooling must be able to hold the part in a known location and react the forces generated by the process. That is, it needs to be made stiff enough not to deflect from the axial and radial forces being imparted on the part. The tooling fixture needs to have clamping mechanisms that allow the FSW pin tool access to the part for the given weldhead articulations and prohibit the part from sliding lengthwise, bending, or separating due to the torque forces. Also, the thermal conductivity of the weld surface and the clamping system can impact the quality of the weld and the welding parameters. For fixed-pin and adjustable-pin welding, the tooling needs to have a backing surface that directly supports and presses against the back side of the part. If there is not proper fit-up, the FSW pin tool will tend to dive into the material, as there will be a decrease in the welding load, and the system will move into the part as a reaction to the change in force. This will occur whether the system is in either position or load control, but will be more pronounced when in load control as the controller will attempt to move into the part until the desired axial load is achieved or a position limit is triggered. With fixed-pin and adjustable-pin welding it is especially important to beware of the impact that the backing plate can have on the thermal flow of the part. Depending on the backing bar or anvil material, for instance stainless steel, steel or a coated material, the thermal flow could lead to a higher (warmer) or lower (colder) weld parameter set. This effect can be used in special cases if more heat is needed on the root side. The backing bar material should be documented in the weld process specification (WPS). Thermal-couples are sometimes embedded into a channel in the backing bar area to better monitor the back-side temperatures (see Chapter 7 Fig. 8). Table 4.5 lists the thermal conductivity of backing plate material and various part materials. The weld quality also depends on the manufacturing accuracy of the weld table and clamping system. The backing bar or anvil should be on the same level as the weld table so that there are no mismatches between the parts being welded (see Fig. 4.14). The clamping system must guarantee to clamp down the work pieces reliably so that no gap can occur during the welding operation. Furthermore, the welding process is easier to handle if the backing bar or anvil is in an absolute plane (i.e., no waves or variations greater than 0.1 mm), which means that the distance from the backing bar or anvil should be constant to the weld tool Z tool axis. In case of a wavy backing, the FSW machine must be able to compensate for these waves to guarantee a constant pin ligament. Similar to the machine design considerations, the exact hold-down clamping
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Thermal conductivity [W/mK]
Mild steel Stainless steel X33CrS16 (1.2085) RAMAX Bras Copper Aluminium alloys
40–60 15–25 17 24 110–150 380–400 110–235
Seam support
+0.05 0 –0.00
Backing
Weld table
4.14 Backing bar tolerance to weld table. Figure courtesy of AirbusBremen.
force is dependent on the material, pin tool, part geometry, joint type, and weld schedule. Please refer to Chapter 2 where sample loads are provided for thick and thin material. Different mechanical clamping systems are used for friction stir welding depending on the application. The simplest and cheapest way to clamp sheets or plates is to use clamping claws. The advantage of this system is a high clamping force and the disadvantage is the high set-up time to clamp the work pieces, the different thermal conductivity if clamping claws are mounted close to the weld seam, and when clamping wide parts, clamping claws are not easily able to reach along the weld seam. The problem of different heat sinking at the clamping claws can be reduced if pressure bars are used beside the weld region (see Fig. 4.15). Figure 4.16 shows a clamping method for holding down Z-shaped stringers for a lap joint weld. You can see that both the top and flange of the Z-stringer are held in place. A challenge with this type of fixture is to make sure that there is adequate access for the FSW pin tool. For serial production it can be desirable to design a special hydraulic or pneumatic fixture so that the set-up time can be reduced (see Fig. 4.17). These fixtures are expensive though and generally only reasonable in production situations.
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Work piece Pressure bar Clamping claw
4.15 Conventional clamping with pressure bar and clamping claws. Figure courtesy of Airbus-Bremen.
4.16 Lap joint Z-stringer clamping. Figure courtesy of MTS Systems Corporation.
An example of how clamping can be done for self-reacting welds is shown in Fig. 4.18. Only the outside area of the part is clamped and there is an opening on the back side of the fixture to allow for the lower shoulder. A good alternative compared to the mechanical clamping systems is vacuum clamping. The set-up time for the vacuum clamping is low and a high rate of parts can be produced quickly and the vacuum plate can be mounted on
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4.17 Hydraulic clamping system. Clamping claws
0
Back-side channel for lower shoulder
Lower shoulder
HEX
2.875
13
10
2.00
4.18 Self-reacting clamping. Figure courtesy of MTS Systems Corporation.
a weld table, or it can be designed to be used as the weld table. Not only flat, but also 3D vacuum clamping systems are available. In general, the 3D vacuum clamping tables can only be used for a specific weld application. The fixture costs, especially for 3D systems, are higher than for standard mechanical clamping systems. A standard vacuum clamping system, shown in Fig. 4.19, consists of: ∑ backing bar (no vacuum clamping) ∑ vacuum plate with vacuum fields ∑ vacuum pump with valves to control different vacuum fields ∑ round rubber sealing ∑ tapped holes for mounting of additional clamping and mechanical fit-up and reaction points ∑ support system for wide sheets. A variable vacuum clamping system should provide independent vacuum areas and a grid of grooves for the sealing to be more flexible in clamping
Vacuum connection
Field 3
Field 2
Field 1
Field 5
Field 6
Field 7
Field 8
4.19 Typical vacuum clamping configuration. Figure courtesy of Airbus-Bremen.
∑ Tapped holes for additional clamping ∑ Wide grid for sealing ∑ Small grid for sealing ∑ Backing bar
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Field 4
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different part geometries. An advantage of several clamping fields is to clamp easily compliant, flexible work pieces. In this case, the sealing has to be larger to reduce the space between the work piece and the sealing in order to generate the vacuum. Tapped holes can be used to mount some physical stops along the work piece in case of insufficient vacuum clamping forces, or to clamp the run-in and run-out with additional clamping claws. In Fig. 4.20, a cylindrical clamping fixture is shown that was used for test panel welding of the Airbus A340-500 HGW. The vacuum fields have to guarantee a sufficient clamping force perpendicular and parallel to the weld direction of the work piece. There are nine vacuum areas on each side of the table. Each area is bordered with a 10 mm rubber sealing to clamp the panel by vacuum. On one side of the fixture are inserted metal plates to adjust the weld position of the panel, and on each side of the fixture there are five adjustable support arms that can be used to support larger parts. Vacuum clamping systems offer some distinct advantages. They are very flexible systems that are easy to use and allow for clamping of different part sizes, both large and small width and lengths. The thermal flow from the FSW process is constant over the whole backing bar, compared to conventional clamping, which leads to good weld quality. However, vacuum clamping forces are not always sufficient for thick plates, which in turn can require that conventional clamping methods are used. When using CAD/CAM packages, the path trajectories are done in the part coordinate system. Once the part is placed in the tooling fixture, the controller generally uses the tooling fixture as a locating reference for transforming the
Support for sheets
Vacuum fields
Backing bar
4.20 Cylindrical vacuum clamping fixture. Figure courtesy of AirbusBremen.
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trajectory from the part coordinate system into the machine world coordinate system. Therefore it is important to have well-defined locating points, usually three separate points, on the fixture that allow the controller to accurately know the location and orientation of the part. One other type of clamping that should be mentioned involves the use of clamping rollers located just ahead or behind the pin tool to press the weld parts against the backing surface. The following figures show different variations to this concept. Figure 4.21 shows hydraulically driven clamping rollers on the Esab SuperStirTM system. In Fig. 4.22, MTS’ surface sensor and roller system is shown. Dual or single rollers are preloaded using mechanical springs and the roller positions are measured and used to locate the part surface relative to the FSW pin tool. Figure 4.23 shows an invention by TWI where a mechanical position system, uses one or two rollers beside or in front of the pin tool. These rollers hold the work piece in place and also work to ensure that the pin tool does not plunge too deep into the work piece. An interesting clamping scheme can be seen in Fig. 4.24. This is the fixture that is used to weld curved gore panels for the upper portion of the liquid hydrogen tank used on the NASA Constellation Ares I upper stage (see Fig. 4.24). This fixture, and the robotic weld system used to do the welding, is installed at the NASA Marshall Space Flight Systems Center in Huntsville Alabama. The fixture allows for panel trimming, fixed-pin welding, and selfreacting welding. The hold-down clamps are pneumatic actuated clamping fingers that are grouped along the weld seam. Each clamping finger has a
4.21 Compliant roller on Esab’s SuperStirTM machine to locally press the sheets on the backing bar. Figure courtesy of ESAB.
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4.22 Compliant roller and surface sensor on MTS weldhead presses the sheet to the backing plate while measuring the location of the surface relative to the pin tool. Figure courtesy of MTS Systems Corporation.
4.23 Conventional concept of rollers beside the FSW tool to maintain the tool heel plunge depth. Figure courtesy of Permission TWI Ltd.
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4.24 NASA gore panel clamping. Figure courtesy of NASA.
predefined load range for clamping and can be fine-tune adjusted by using spacer washers. Originally this application was developed for arc welding and was then adapted for friction stir welding. The pneumatic fingers are not fully sufficient to fix the work pieces in place. Therefore a partial penetration weld is done which lowers the clamping forces significantly so that the pneumatic fingers sustain the piece. For the final weld the tack weld reacts the welding side forces.
4.10
Friction stir welding (FSW) pin tools
Since its original development, friction stir welding advances have often been driven by the development of new welding tools and new welding equipment. The development of welding tools, which is covered in this section, has two aspects; the development of welding tool designs and the development of welding tool materials. Welding tool designs can be divided into two main classes: conventional and bobbin. Conventional FSW tools approach the work piece from one side, typically while the work piece is restrained on an anvil. This type of tool only partially penetrates the work piece, leaving a very thin gap between the end of the pin and the anvil. Material which passes through this gap is plasticized to a sufficient degree to ensure complete consumption of the original joint, when the pin ligament is maintained.
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In contrast to the conventional FSW tool, bobbin FSW tools consist of two shoulders which are connected by the pin, all of which typically rotate in unison. The two shoulders act to contain the softened weld metal from either side, as shown in Fig. 4.25, while generating heat from friction and plastic work. The action of the pin is primarily to distort the faying surfaces and generate additional heat to sustain the process while providing the mechanical connection between the shoulders. Bobbin FSW tools can allow the production of welds without imparting force normal to the work piece plane, and they eliminate the possibility of incomplete penetration of the weld to the anvil side of the work piece. There are two further subdivisions of bobbin tool designs, as shown in Fig. 4.26. Although the notion of a bobbin tool was first described in the original patent filing for FSW,6 it was never practically demonstrated in its original form. The original bobbin tool consisted of a pair of smooth, bare shoulders that tapered outward, presumably for the purpose of ensuring contact with the workpiece surfaces in the event of random or planned variation in the workpiece thickness. A workable bobbin FSW tool was later developed in which flat shoulders were supplemented with spiral grooves or scrolls which acted to pull workpiece material toward the pin.7–13 The shoulders were forcibly actuated relative to each other in order to provide a variable gap between the shoulders, which facilitated the development of controllable compressive force between the shoulders, even in the case where work piece thickness varies along the length of the weld. This type of tool requires a mechanism to actuate the welding tool’s pin in order to achieve the variable shoulder gap. This variable shoulder gap bobbin tool, known as a self-reacting tool, has been demonstrated to produce sound welds in a wide variety of material thicknesses and alloy selections. Later, a practical fixed-
Pin connects shoulders (obscured from view)
Spindle-side shoulder
Work piece
Back-side shoulder
4.25 Bobbin FSW tool.
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Fixed-gap, tapered bobbin tool
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Variable-gap bobbin tool
4.26 Variable- and fixed-gap bobbin FSW tools.
gap bobbin tool was demonstrated, eliminating the need for a mechanism to actuate the welding tool pin. This tool, known as a tapered shoulder tool, has also been demonstrated to produce welds in a variety of materials.14,15 The tapered shoulder tool takes advantage of a protruding shoulder profile combined with spiral shoulder scrolls to result in a tool that has variable shoulder penetration and variable effective shoulder width. Variations of this tool employ non-linear shoulder profiles. Welding tool features have been introduced over time to produce a number of desirable effects, and new tool design features are continually being introduced. Table 4.6 shows a summary of some of the key FSW tool design features that have been developed to date, along with appropriate references. Various combinations of these design features make possible a wide variety of conventional and bobbin FSW tools. In addition to welding tool design features, welding tool ‘motions’ have been developed as a means of achieving desirable effects in weld formation. These have primarily been introduced by TWI, and include variations such as ‘Skew Stir’ (tool axis inclined relative to axis of rotation), dual-rotation FSW (pin and shoulder rotate at different speeds or in different direction), ‘Re-Stir’ (tool rotation direction periodically reverses), ‘Com-Stir’ (compound motion of tool rotation and tool axis orbit), and tandem FSW (two tool operating in tandem).16 These tool motions have been shown to give different material flow patterns during welding, yielding improved weld formation, for example, to improve lap weld performance. New welding tool materials have been introduced in order to increase the strength of the welding tool’s pin, to manage thermal conduction, 17 and to permit welding of high melting point materials. Although aluminum alloys can be friction stir welded using inexpensive tool steel, such as H13, in order
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Table 4.6 Summary of key welding tool design features Feature
Intended effect
Threads on pin
Compression of weld zone against anvil
Flats or other re-entrant features Flat pin tip Frustum pin profile
New mode of plastic work, thicker section welding, higher heat input Improved TMAZ penetration, higher penetration ligaments – better robustness Reduced lateral forces, thicker section welding
Flare pin profile
Wider root profile
Shoulder scrolls
Elimination of tool tilt requirement, containment of softened work piece material
Tapered shoulder
Variable shoulder contact width, variable shoulder penetration
Examples
to maximize production rate and tool life it is necessary to use materials that have high strength at the temperature of welding, such as MP159. Copper is commonly welded using Nimonic 105 and Densimet for the pin and shoulder materials, respectively.18–20 Titanium is generally welded using refractory metals, such as lanthanated tungsten or tungsten rhenium tools.21 Steel is also welded with tungsten based tools, or with tools made from polycrystalline cubic boron nitride (PCBN).22,23 In addition, tool coatings are occasionally used to improve the wear or chemical resistance of tools.24 Certainly, welding tool design and material selection are important considerations in developing a successful friction stir welding process. Careful
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consideration must be given to tool cost, useful life, and the limitations that the tool strength might place on the welding speed. The development of new welding tool designs and materials has been an active area of research in the past, leading to expansion of the utility of the process, and this trend will likely continue in the future. As a result, attention to the latest advancements in welding tool construction is important for optimizing performance in any FSW application.
4.11
Machines currently available in the market place
In this section we will review some of the different machine configurations that are currently in the FSW market place. Robotic Weld Tool (Fig. 4.27) Customer: National Aeronautic Space Agency Application: Welding of curved gore panels for the upper portion of the liquid hydrogen tank used on the NASA Constellation Ares I upper stage. Configuration: Horizontal boring mill Aero system (Fig. 4.28) Customer: Eclipse Aviation Corporation and FHI Application: Welding of stringers and frames to aircraft outer skin. Used for main cabin, wings, aft fuselage, and engine mount beam sections. Configuration: Bridge gantry Process/production development system (PDS) (Fig. 4.29) Customer: Numerous research and development groups Application: Wide range of applications varying from 2-dimensional to 3-dimensional thin gage to thick plate (0.75 mm to 30 mm) Configuration: Bridge gantry Robotic welding system (RoboStirTM) (Fig. 4.30) Customers: Multiple, including Friction Stir Link Application: General production use from lower to higher volumes, moderate thickness materials, moderate to lower melting point alloys Configuration: Serial/articulate arm robot
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Ares I elements Encapsulated service Module (ESM) panels Instrument unit Orion CEV
Interstage First stage
Liquid oxygen/liquid hydrogen tank gore panels
4.27 Robotic weld tool. Figure courtesy of National Aeronautic Space Agency.
Low cost FSW system (Fig. 4.31) Manufacturer: Nova Tech Engineering, Lynnwood, WA, USA Customer: United States Navy Applications: fabrication of stiffened panels from aluminum extrusions, initially for use on construction of the Littoral Combat Ship (LCS) Configuration: C-Frame, traveling work piece Encapsulation of nuclear fuel waste for long term repository (Fig. 4.32) Manufacturer: ESAB AB, Welding Equipment Customer: Swedish Nuclear Fuel and Waste Management Co (SKB)
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4.28 Aero System. Figure courtesy of Eclipse Aviation Corporation.
4.29 Process/production development system (PDS). Figure courtesy of MTS Systems Corporation.
Applications: R & D manufacturing and sealing of copper canisters for nuclear fuel waste for long term repository Configuration: ESAB customized SuperStirTM equipment
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4.30 RoboStirTM. Figure courtesy of Friction Stirlink.
4.31 Low cost FSW system. Figure courtesy of Nova Tech Engineering.
Panel production plant (Fig. 4.33) Manufacturer: ESAB AB, Welding Equipment Customer: Marine Aluminium a.s. Applications: Manufacturing of aluminium panel with max size 16 ¥ 20 meters, in thickness 1.8 to 12 mm Configuration: ESAB customized SuperStirTM equipment
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4.32 R-D FSW. Figure courtesy of ESAB.
4.33 Marine Aluminum FSW. Figure courtesy of ESAB.
4.12
References
1. Getting a handle – on machine accuracy; Society of Manufacturing Engineers, July 2001. 2. http://en.wikipedia.org/wiki/PID_controller#General
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3. Smith, C. TWI FSW Symposium, Gothenburg, Sweden, June 2000. 4. Cooke, G. ‘Robotic Friction Stir Welding’, Industrial Robot, Vol 33(1) 2004, pp. 55–63. 5. Mikael Soron ESAB AB Welding Equipment, Sweden
[email protected] Friction stir welding of high-strength aluminium alloys using an industrial robot system: A feasibility study; 2008 FSW International Symposium, Japan. 6. Thomas, W.M., Nicholas, E.D., Needham, J.C., Murch, M.G., Temple-Smith, P. and Dawes, C.J., “Improvements related to friction welding,” PCT Patent Application No. PCT/GB92/02230, 10 June 1993. 7. Campbell, C.L., Fullen, M.S. and Skinner, M.J., “Welding head,” US Patent 6,199,745, 12 March 2001. 8. Li, T., Hartley, P., Halpin, J., Skinner, M. and Edwards, B., “Characterization of 2195 and 2219 self-reacting friction stir welds,” Aeromat 2002, Orlando, Florida, USA, unpublished reference, 10–13 June 2002. 9. Skinner, M. and Edwards, R.L., “Improvements to the FSW process using the selfreacting technology,” Thermec 2003, Leganes, Madrid, Spain, 7–11 July 2003. 10. Marie, F., “Development of the bobbin tool technique on various aluminum alloys,” 5th International FSW Symposium, Metz, France, 14–16 September 2004. 11. Thomas, W.M., Verhaeghe, Martin, J., Staines, D.G. and Stanhope, C., “Friction stir welding – process variants – an update,” Inalco 2007, Tokyo, Japan, 24–26 October 2007. 12. Neumann, T., Zettler, R., Vilaca, P., dos Santos, J. and Quintino, L., “Analysis of Self-Reacting Friction Stir Welds in a 2024-T351 alloy,” Friction Stir Welding and Processing IV, Mishra, R.S., Mahoney, M.W., Lienert, T.J. and Jata, K.V., eds., TMS, 39–54, 2007. 13. Sylva, G., “Simultaneous opposed FSW of large hollow Al-7249-T6511 extrusions using the self-reacting pin tool process,” 7th International FSW Symposium, Awaji Island, Japan, 20–22 May 2008. 14. Colligan, K.J., “Tapered friction stir welding tool,” US Patent 6,669,075, 30 December 2003. 15. Colligan, K.J. and Pickens, J.R., “Friction stir welding of aluminum using a tapered shoulder tool,” Friction Stir Welding and Processing III, Jata, K.V., Mahoney, M.W., Mishra, R.S. and Lienert, T.J., eds., TMS, 161–170, 2005. 16. Thomas, W., Norris, I.M., Staines, D.G. and Watts, E.R., “Friction stir welding – process developments and variant techniques,” The SME Summit 2005, Oconomowoc, Milwaukee, Wisconsin, USA, 3–4 August 2005. 17. Midling, O.T. and Rorvik, G., “Effect of tool shoulder material on heat input during friction stir welding,” 1st International Symposium on Friction Stir Welding, Thousand Oaks, CA, USA, 14–16 June 1999. 18. Andersson, C-G. and Andrews, R.E., “Fabrication of containment canisters for nuclear waste by friction stir welding,” 1st International Symposium on Friction Stir Welding, Thousand Oaks, CA, USA, 14–16 June 1999. 19. Hautala, T. and Tiainen, T., “Friction stir welding of copper,” 6th International Trends in Welding Research Conference Proceedings, Pine Mountain, GA, USA, 15–19 April 2002, ASM International, 324–328, 2003. 20. Cederqvist, L. and Andrews, R.E., “A weld that lasts for 100 000 years: FSW of copper canisters,” 4th International FSW Symposium, Park City, UT, USA, 14–16 May 2003. 21. Jones, R.E. and Loftus, Z., “FSW of 5mm Ti 6Al 4V,” 6th International FSW Symposium, Saint Sauveur, Canada, 10–13 October 2006.
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22. Sorensen, C.D. and Nelson, T.W., “Progress in polycrystalline cubic boron nitride FSW tooling,” 4th International FSW Symposium, Park City, UT, USA, 14–16 May 2003. 23. Nelson, T.W. and Sorsensen, C.D., “Advances in PBN tooling for friction stirring of high temperature alloys,” 6th International FSW Symposium, Saint Sauveur, Canada, 10–13 October 2006. 24. Subramanian, P.R., Bewlay, B.P., Helder, E.C. and Trapp, T.J., “Apparatus and method for friction stir welding of high strength materials, and articles made therefrom,” US Patent Application 2004/0238599, 2 December 2004.
5
Industrial applications of friction stir welding
S. W. Kallee, Germany
Abstract: Friction stir welding (FSW) is a patented new welding process that has had led to many worldwide applications, predominantly in the fabrication of aluminium components and panels. Trendsetters were the Scandinavian aluminium extruders for the manufacture of hollow aluminium deep freezer panels and for ship decks. The railway rolling stock industry uses FSW for the production of large prefabricated aluminium panels, which are made from aluminium extrusions. In the aerospace industry, large tanks for satellite launch vehicles are being produced by FSW from high-strength aluminium alloys, and several companies manufacture lightweight aluminium airframe structures for commercial and military aircraft. The automotive industry uses FSW now in the high-volume production of components, e.g. light alloy wheels and fuel tanks. Key words: friction stir welding, FSW, friction stir spot welding, FSSW, friction stir processing, FSP, application, industrial, implementation, civil, military, commercial, use.
5.1
Introduction
Since the invention of friction stir welding at TWI in 1991, companies from all parts of the world have implemented the process, predominantly in the fabrication of aluminium components and panels. Trendsetters were the Scandinavian aluminium extruders, who in 1995 were the first to apply the process commercially for the manufacture of hollow aluminium deep freezer panels and for ship decks. In the railway rolling stock industry several companies now exploit the process, e.g. for the production of large prefabricated aluminium panels, which are made from aluminium extrusions. In the aerospace industry, large tanks for satellite launch vehicles are being produced by FSW from high-strength aluminium alloys, and several companies have obtained approval to use FSW for the manufacture of lightweight aluminium airframe structures for commercial and military aircraft. The automotive industry uses FSW now in the high-volume production of components, e.g. light alloy wheels and fuel tanks. The following comments on cost savings have been published by users of the FSW process and speak for themselves: ∑ Hydro Aluminium, a Norwegian extrusion company, reported that at 118
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shipyards using prefabricated FSW panels the “improvement in the aluminium fabrication has resulted in 15% reduction in the man-hour per ton rate”. ∑ Fjellstrand, a Norwegian shipbuilding company, claimed “a total fabrication cost saving of approximately 10% based on improved ship design, streamlined fabrication at the shipyard and by supply of prefabricated FSW panels and structures based on extruded profiles”. They said that using prefabricated FSW panels “has enabled the yard to reduce the production period for a 60 m long aluminium catamaran hull from 10 to 6 months, which means a 40% increase in production capacity and turn-over at the yard”. ∑ The Boeing Company reported that “the FSW specific design of Delta IV and Delta II [satellite launch rockets] achieved 60% cost saving, and reduced the manufacturing time from 23 to 6 days”. ∑ For ‘Slipper’, the US Army’s new cargo interface pallet, “FSW processing reduced the sandwich assembly cost, including raw materials, extruding, and welding, from 61% to only 19% of the total fabrication cost. The Air Force estimates the total cost savings attributed to FSW [for a projected buy of 140 000 Slippers] at $315 million”. ∑ General Dynamics Land Systems, a US defence company, commented: “The use of friction stir welding allowed us to more accurately join different aluminum alloys together, including some lithium-based alloys”. The welding process also provided improved ballistic test results, along with reduced cycle times – up to 400% for 25 mm thick plates compared to the original 2-pass process – and material distortion. The military expects the use of friction stir to provide improved return on investment results within 5 years, with savings upwards of $25.8 million, according to the Office of Naval Research. “We think this is where we have found the most transferable technology for use outside the defense industry.”
5.2
Shipbuilding and offshore
In the shipbuilding and offshore industry several companies use the FSW process for the production of large aluminium panels, which are made from aluminium extrusions. Commercial FSW machines are now available and include complete installations to weld up to 25m length and 14.5m width.
5.2.1 Freezer panels in Sweden and Germany The first commercial application of friction stir welding concerned the manufacture of hollow aluminium panels for deep freezing of fish on fishing boats in November 1996 at Sapa in Finspång (Sweden). These panels are made from friction stir welded aluminium extrusions (Figs 5.1–5.2).
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5.1 Sapa FSW freezer panel for pre-pressing of fish blocks (the first industrial application of FSW in 1996).
5.2 Joint design of Sapa’s freezer panels (weld penetration 4.5 mm, total weld length 16 m).
Minimal distortion and high reproducibility make FSW both technically and economically a very attractive method to produce these stiff panels. Riftec in Geesthacht (Germany) began in 2005 to friction stir weld plates for industrial
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plate freezing plants in batches of up to 700 pieces. These 2.5 ¥ 4.5 m hollow plates weigh 350 kg and contain 50m of friction stir welds per plate.
5.2.2 Ship and oil-rig panels for decks and bulkheads in Scandinavia Pre-fabricated wide aluminium panels for high-speed ferry boats, cruise ships and offshore oil platforms can be produced by friction stir welding and are commercially available. The panels are made by joining extrusions, which can be produced in standard size extrusion presses. Thousands of panels with an overall weld length of several hundreds of kilometres have been produced and delivered by Marine Aluminium in Haugesund (Norway) since 1996 (Figs 5.3–5.5). After welding the panels can be rolled for road transport, as they are stiff only in the longitudinal direction. If they are transported by ship, they can be stacked on top of each other. These panels are also commonly used for making leak-tight helicopter platforms of oil-rigs, to avoid burning aircraft fuel dropping into the living quarters of offshore oil platforms, after a helicopter crash landing. In 2008 SLP Engineering in Lowestoft (UK) ordered a batch of 100t of FSW floor panels from by Marine Aluminium, to make the living quarters for the Valhall redevelopment project in the Norwegian sector of the North Sea
5.3 Marine Aluminium’s prefabricated FSW deck panels for “The World” cruise ship.
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5.4 Fosen Mek’s cruise ship “The World” contains FSW decks of Marine Aluminium.
5.5 FSW module for “Finnmarken” at Marine Aluminium.
under contract by BP Norge AS. The transportable living quarters with 180 single-bed cabins will weigh in total approximately 3100 t (Figs 5.6–5.7). The aluminium extrusions are 5 and 7 mm thick in the FSW joint areas, and the maximum panel size on this contract is approximately 9 ¥ 13 m. Friction stir welded panels are now being used all over the world for highspeed ferries, hovercraft and cruise ships. Large prefabricated aluminium modules can be lifted by crane into ships or oil platforms, to save time during the final assembly. The total length of friction stir welds made by Sapa in Finspång (Sweden) to date is much more than 3000 km. The commercial use
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5.6 FSW of helicopter landing platforms and offshore floor panels at Marine Aluminium.
5.7 Living quarters with Marine Aluminium’s FSW panels for offshore platform Oseberg Sør.
of structures produced by FSW has been supported by component approvals, based on an appropriate welding procedure specification (WPS), for each case.
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5.2.3 Explosively formed hull of an ocean viewer vessel in Australia In Australia, the Department of Mechanical Engineering at the University of Adelaide has developed a portable prototype FSW machine to manufacture a new type of ocean viewer vessel. The machine was used at the Research Foundation Institute (RFI) in Cairns under site conditions of a relatively low-tech shipyard. Six friction stir welds were made in the bow section of the prototype ship using 5 mm thick aluminium alloy 5083-H321 (Figs 5.8–5.9).
5.8 Portable FSW machine at the shipyard of the Research Foundation Institute in Cairns, Australia.
5.9 Friction stir welds on the starboard side of the bow section prior to explosive forming at the Research Foundation Institute in Cairns, Australia.
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The friction stir welded sheets of the bow section were given their final three-dimensional shape after welding by high energy rate forming (HERF). Explosive plates were fixed onto the aluminium sheets, which were positioned in a mould. The mould was filled with water, before detonating the explosives. The resulting structure was then installed in the prototype ocean viewer vessel (Fig. 5.10). This development programme resulted in an innovative and patented prototype ship, which combines the attributes of a fast ferry with those of a semi-submersible reef viewing vessel. The benefits of the new concept include operator flexibility to access different reefs depending on daily sea and wind conditions, so providing a more reliable service. This low impact and environmentally sustainable mode of reef viewing is increasingly being required by international authorities endeavouring to protect their sensitive marine parks.
5.2.4 Honeycomb panels and corrosion resistant panels in Japan and USA FSW applications are being reported in Japan, where the process is being used to produce honeycomb panels and sea water resistant panels. The latter are made from five 250 mm wide 5000 series aluminium extrusions joined by FSW to make a panel of 1250 ¥ 5000 mm. These panels are used for ship cabin walls because of the good flatness of the welded structure, and the excellent mechanical properties of the welds. The Tamano Works of Mitsui Engineering & Shipbuilding (MES) in Japan used FSW, when they built a combined passenger and freight ship with a maximum speed of 42.8 knots. This ship was given the name ‘Super
5.10 Explosively formed bow section of the ocean viewer vessel at the Research Foundation Institute in Cairns, Australia.
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Liner Ogasawara’ and can transport up to 740 persons and 210t of freight (Fig. 5.11). It was successfully tested in up to 2 m high waves. The Nichols Brothers Boat Builders in Freeland, Washington (USA) successfully used FSW aluminium panels for a 55 knots military ship of the X-Craft class, which has been named ‘Sea Fighter’ (Fig. 5.12).
5.11 ‘Super Liner Ogasawara’ with 42.8 knots max speed by MES.
5.12 Bow of US Navy X-Craft ‘Sea Fighter’ by Nichols Bros.
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5.2.5 Deck panels for civil and naval ships in China In 2003, China FSW Centre (CFSWT) in Beijing designed and fabricated its first FSW industrial product-line for a small company in Chang Zhou (Fig. 5.13). This PLC-controlled equipment can weld 2600 ¥ 1100 mm panels from 6 mm thick aluminium extrusions for use in various sectors of the transport industry. China FSW Centre designed and produced the first large FSW machine for wide ship panels in China in 2006 after considering production, weight and transport aspects. This machine can weld aluminium alloy sheets and extrusions from 2 to 6 mm thickness. The largest size panel produced so far was 60 m2 (12 m ¥ 5 m). The CFSWT machine is used for batch production of wide stiffened panels, which are used in high-speed aluminium alloy ships. Since the introduction of FSW panels, a new era started for the Chinese aluminium shipbuilding industry. First of all, batch production of panels by FSW helps to resolve on-site welding problems significantly. This technology has also simplified the design of ships. And most importantly naval architects now have more options, when they design new structures regarding material selection: ∑ FSW can join extruded AA 6082 to corrosion resistant AA 5083. ∑ Forged or stamped parts can be welded to castings. ∑ Thin sheets can be welded to thick sheets or plates. ∑ MIG welds can cross over friction stir welds, e.g. when joining FSW panels to girders.
5.13 FSW of 5 ¥ 12 m large aluminium ship panels at CFSWT.
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∑ Batch production of panels reduces on-site welding activity. ∑ Simplified design. ∑ More options for material selection. Parts of ships can be assembled more accurately, and the precision of ship modules and the final shape of ships can be improved. Nowadays, the concept of using prefabricated FSW panels for shipbuilding is popular at Chinese shipyards in Dalian, Shanghai, Wuhan, Guangxi and Guangzhou. These wide panels have successfully been used in many shipbuilding projects, including ships designed and fabricated in China for export to Vietnam and Micronesia. The Type 022 Houbei Class is the Chinese People’s Liberation Army Navy’s new-generation stealth missile fast attack craft (FAC). The boat features a unique high-speed, wave-piercing catamaran hull with evident radar cross-section reduction design features (Fig. 5.14). A number of Chinese shipyards across the country have been involved in the construction of the aluminium boat, and it has been reported that FSW panels have been used to produce the missile launch system of this very advanced navy vessel in China.
5.2.6 Patrol vessels in New Zealand The first New Zealand fabricators have recently begun to use FSW. In 2004 it was announced that a number of 55-metre Inshore Patrol Vessels would
5.14 ‘Type 022’ new-generation stealth missile fast attack craft with CFSWT’s missile launch containers.
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be procured for use by the Royal Australian Navy and Royal New Zealand Navy, and thus the opportunity existed for both Australia and New Zealand to be involved in building them. The naval architect of these vessels specified that a significant portion of the structure was to be friction stir welded. In mid 2005, the Donovan Group in Whangarei (NZ) implemented FSW for the manufacture of these vessels. The Donovan Group has since then modified a large CNC gantry milling machine to be used as a FSW machine for large-scale production (Figs 5.15–5.16). Tenix Shipbuilding in Whangarei (NZ) uses these panels for assembling the superstructure of these Inshore Patrol Vessels (Fig. 5.17).
5.15 FSW at Donovan Group, Whangarei, New Zealand.
5.16 FSW ship panels at Donovan Group, Whangarei, New Zealand.
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5.17 Friction stir welded superstructure for inshore patrol vessel ‘Rotoiti’ at Tenix Shipbuilding, Whangarei, New Zealand.
5.3
Aerospace
In the US aerospace industry, large tanks for satellite launch vehicles are being fabricated by FSW from high-strength aluminium alloys. The first rocket with a friction stir welded interstage module was successfully launched in August 1999. Recently, the first approval has been obtained for the use of the FSW process in the manufacture of American business jets.
5.3.1 Fuel tanks for spacecraft in the USA An increasing number of fuel tanks for spacecraft are now being produced from difficult-to-weld aluminium alloys. Boeing has applied FSW to the interstage modules of Delta II rockets, and the first of these was launched successfully in August 1999. The Mars Odyssey launch in April 2001 utilized the first pressurized structures. The Mars Odyssey spacecraft lifted off on a Delta II rocket, which demonstrated the strength and quality of longitudinal friction stir welded joints on all three cylindrical tank components. Friction stir welding technology for the Delta IV common booster core tanks increases the weld strength by 30 to 50% and lowers cycle time by nearly 80%. A total of 2100m of defect-free friction stir welds have been produced for Delta II rockets, and 1200 m for the larger Delta IV rocket by July 2001 (Fig. 5.18). The FSW specific design of Delta IV achieved 60% cost saving, and reduced the manufacturing time from 23 to 6 days. The temperature range to which the friction stir welds are submitted during service is –195°C to +183°C. Lockheed Martin Space Systems, Michoud Operations designed and built
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5.18 Boeing’s liquid-oxygen and liquid-hydrogen tanks for the 42 m (125 ft)-long common booster cores.
the Space Shuttle External Tank at the NASA Michoud Assembly Facility in New Orleans (USA). Process development and large-scale demonstrations have used both electro-mechanical and hydraulic equipment. The company also designs and assembles large aluminium and composite structures for aerospace and other applications. NASA is now developing hardware and systems for the Ares I at the Marshall Space Flight Center, Huntsville, Alabama (USA). Beginning in 2015, the rocket will launch the Orion crew capsule to the International Space Station, carrying six astronauts and small pressurized cargo payloads. The Ares I is a single, five-segment, reusable solid rocket booster derived from the Space Shuttle Programme’s reusable solid rocket motor. Ares I may use its 25 t payload capacity to deliver resources and supplies to the space station, or to ‘park’ payloads in orbit for retrieval by other spacecraft bound for the moon or other destinations (Figs 5.19–5.20). Space Exploration Technologies Corporation (SpaceX) in Hawthorne, California (USA) uses FSW for the partially reusable launch vehicles Falcon 1 and Falcon 9. The first stage of Falcon 1 employs a friction stir welded common bulkhead between the liquid oxygen and rocket propellant tanks. It can be transported safely without pressurization, like the heavier Delta II, but gains additional strength when pressurized for flight. The Falcon 9 tank walls and domes are made from friction stir welded aluminium lithium alloy AA2198, and the F9 tank is claimed to be the largest fully friction stir welded
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5.19 Engineers at NASA’s Marshall Space Flight Center successfully demonstrate their first “official” weld with tools for the Ares I upper stage hardware test articles.
5.20 The Vertical Weld Tool at the Marshall Space Flight Center of NASA can assemble tank barrel sections up to 12 m in diameter and 7.6 m length of full-scale Ares I developmental hardware.
structure in the world. The 55 m long vehicle is assembled horizontally and then rolls out to the pad to be placed vertical for launch (Fig. 5.21). Friction stir welding is also being considered for producing Ariane 5 motor thrust frames. A study by Dutch Space has shown that FSW can readily be
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5.21 Circumferential and longitudinal friction stir welds in a Falcon 9 tank of SpaceX.
applied to lap joints in aluminium alloy 7075-T7351. Although the tensile strengths measured in this investigation were lower than those that can be obtained with friction stir welded butt joints, they were at an acceptable level to replace bolted lap joints. For unpressurized structures lap joints offer the significant advantages of generous tolerances at interfaces between components, and ease of assembly.
5.3.2 The first flightworthy parts of cargo aircraft in the USA The toe nails of Boeing’s C-17 cargo ramp were made from friction stir welded aluminium alloy AA7050-T7451 and are the first aircraft parts that have been commercially produced (Fig. 5.22). The company AJT-Inc (Advanced Joining Technologies Inc.) received the production contract in December 2004. Ten toe nails were installed on P136 aircraft and flew away in June 2005. A zero scrap rate had been reported in February 2006, after 100% of the first 135 FSW toe nails had successfully passed non-destructive inspection.
5.3.3 Aluminium panels in aircraft production in the USA The Phantom Works of The Boeing Company are pursuing FSW of thin butt, lap and T-joints and thick butt joints for various aircraft, missile and space applications. There is a strong desire for welding these joint configurations
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5.22 FSW toe nails for military cargo aircraft are supplied by AJT-Inc.
on curvilinear paths thus enabling welding of complex aircraft parts. Boeing has demonstrated curvilinear FSW of a complex aircraft landing gear door by using a patented force actuator. Boeing has also successfully demonstrated FSW of sandwich assemblies by welding thin T-joints for a fighter aircraft fairing, which has been flight tested. The production start of friction stir welded non-structural parts for the Boeing commercial aircraft was scheduled for October 2001. Eclipse Aviation Corporation of Albuquerque, New Mexico (USA) has decided to use FSW to replace traditional riveting and bonding processes. This was the first application of this welding process in commercial aircraft production with the benefit to dramatically lower assembly time and cost (Figs 5.23–5.24). The use of FSW for the Eclipse 500 has eliminated 7000 rivets and fasteners. It replaced associated hole drilling and has also resulted in faster joining speeds: four times faster than manual riveting and twenty times faster than manual riveting. FSW increased the joint strength by up to three times at comparable or better fatigue data. Testing of a full-size FSW barrel simulating the aircraft fuselage showed a fatigue life in excess of 23 aircraft lifetimes. According to the 2006 version of the Eclipse 500 Maintenance Manual, the aircraft was given an initial airframe life of 10 000 hours, 10 000 cycles or 10 years, whichever comes first. Eclipse Aviation Corporation announced in June 2002 that the FAA (US Federal Aviation Administration) has approved the FSW specification created for use in the assembly of the Eclipse 500 jet. The FAA approved the FSW process specification, in conjunction with the receipt of the type certificate, and Eclipse Aviation ramped-up the production of their very light jets using
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5.23 FSW gantry at Eclipse Aviation for welding stringers and spars to aluminium cabin panels.
5.24 Take-off during the first test flight of an Eclipse 500 friction stir welded business jet.
FSW. The FSW process specification describes the procedural requirements, quality assurance provisions, standards for tooling and material preparation necessary for the use of FSW in the assembly of aircraft (Figs 5.25–5.26).
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Eclipse Aviation obtained also certification for private use by the EASA (European Aviation Safety Agency) in November 2008. A few days later, at a time when Eclipse Aviation had delivered 259 friction stir welded Eclipse 500 aircraft, the company filed for Chapter 11 bankruptcy protection, but in February 2009 it was forced into Chapter 7 liquidation under the laws of the United States.
5.25 Eclipse right skin assembly with an oval emergency hatch. FSW results in a smooth and easy to paint surface.
5.26 Eclipse lower panel interior.
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5.3.4 Aircraft fuselages and wings in Europe Airbus has shown interest in friction stir welding since 1998 and has invested in several laboratory machines focusing on thin and thick section welding of aluminium alloys. To date Airbus has matured the process for fuselage and wing applications by validating the mechanical and technological properties of the joints even by successful full-scale testing of several demonstrators, e.g. wing ribs and fuselage barrel sections. Besides validating the robustness and reproducibility of the welds the design improvement and the quality assurance for aircraft application have also been developed. Airbus has investigated many applications, and so far no technical show-stoppers have been identified. The application focus now has to be adapted to different applications from in 1998 due to the change from aluminium to composites in many primary parts of the wing and fuselage. This gives friction stir welding the chance to be further developed especially for dissimilar alloy joining. The first flying application for Airbus will be friction stir welded floor panels in the A400M military transport aircraft. The parts are developed and produced by PFW Aerospace AG according to the Airbus process specification and are currently finishing their part qualification programme. They are made from stiffened extrusions (Fig. 5.27). There will be 102 panels in the cargo hold approximately 1m long and different in width. The friction stir welded solution was selected, because it best fulfils the requirements with the lowest weight.
5.27 Friction stir welded floor panel for Airbus A400M military aircraft.
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Railway
Friction stir welded structures are now revolutionizing the way in which trains and trams are built. In Europe, suppliers to the railway rolling stock industry are exploiting the process for the prefabrication of large panels, which are made from aluminium extrusions. In Japan, complete trains are being assembled from hollow extrusions using the innovative process. Recent investigations into the crashworthiness of aluminium railcars have clearly demonstrated the benefits of using innovative joint and tool designs and optimized procedures for friction stir welding. Modern railway carriages are increasingly produced from longitudinal aluminium extrusions with integrated stiffeners. Using this concept the whole body-shell can be made from either single-wall or hollow doubleskin extrusions. This design approach can enhance the crashworthiness of vehicles because of the absence of transverse welds and the high buckling strength of the panels under longitudinal compression. The railway vehicle industry makes increasing use of friction stir welding. The driver for the uptake of this process is its combination of cost effectiveness and good weld performance.
5.4.1 Aluminium panels for railway rolling stock in Europe The Scandinavian aluminium extruders Sapa and Hydro Marine Aluminium were the first in Europe to commercially apply the friction stir welding process for the manufacture of single-wall aluminium roof panels for rolling stock applications. Sapa in Finspång (Sweden) started production of FSW panels in November 1996. They now have a production rate of approximately 1 000 m of friction stir welds per day, i.e. a yearly production rate of more than 300 000 m. The total length of friction stir welds made by Sapa up to date is more than 3 000 000 m. In addition to the production of large rolling stock panels they also produce smaller components, such as heat sinks for the high-power electronics of electric locomotives, e.g. IGBT coolers for insulated gate bipolar transistors. Sapa supply all the major European train manufacturers (Alstom, Bombardier, CAF, Siemens, etc.) with aluminium extrusions and friction stir welded panels for a number of different projects. Since 1997, Alstom LHB in Salzgitter (Germany) have purchased these prefabricated panels for Copenhagen suburban trains (Figs 5.28–5.29). In March 1999, Alstom LHB engineers considered friction stir welding hollow aluminium profiles for making floor and side panels, but calculated that a three-shift operation would be necessary for achieving return on investment in an acceptable time span. They estimated that the most significant technical
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5.28 Alstom LHB trains for DSB Danish State Railways during production. FSW roof panels for these trains were made at Hydro Marine Aluminium under a contract with Sapa.
5.29 Friction stir welded roof panel produced at Hydro Marine Aluminium for Sapa for delivery to Alstom LHB (Germany).
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and economic benefits could be achieved by applying FSW to aluminium joints of more than 12 mm thickness. This would replace mechanised MIG welding which necessitates the associated activities of pre-heating and grinding of intermediate beads. Additionally, it would lead to improved quality of the welds. Therefore, successful FSW experiments were conducted in up to 23 mm thick aluminium plates, to demonstrate how MIG welds could be replaced in the underframe area of rolling stock. Since early 2001 they have used friction stir welded aluminium side walls and since 2002 FSW floor panels for Munich suburban trains. These panels are made by Sapa (Fig. 5.30). Bombardier Transportation in Derby (UK) have carried out FSW experiments for butt and lap welds and have conducted fatigue tests at TWI. They have stated that one of the major advantages of FSW was the ability to weld larger joints with reduced distortion. However, they concluded that investment in large purpose built FSW machines was at that time difficult to justify – partly due to insufficient volume of work. Using a subcontractor or job shop is a possible solution. Now, they fabricate an increasing number of car bodies from friction stir welded aluminium panels, which are subsequently joined together using MIG welding. The FSW process is used to prefabricate stiff longitudinal extrusions which constitute the car body sidewalls, e.g. for Bombardier’s Electrostar trains (Figs 5.31–5.32). No fewer than 376 friction stir welded vehicles have also been ordered by London Underground for the
5.30 Sapa pre-fabricates curved train side skirt panels by FSW. The cumulative friction stir weld length over the last 12 years has exceeded 3000 km.
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5.31 Friction stir welded side skirt panels are made by Sapa from double-skinned extrusions and used by Bombardier for commuter and underground trains.
5.32 Sapa’s FSW aluminium panels have excellent surface finish after painting at Bombardier due to the low amount of distortion caused by the low-temperature FSW process.
latest Victoria Line upgrade (Fig. 5.33). FSW is likely to be used on future contracts for over 2000 more new trains in the next decade. Hammerer Aluminium Industries GmbH in Ranshofen (Austria) have built and commissioned a gantry-type friction stir welding machine and obtained approval to use this for rolling stock according to DIN 6700. They have
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5.33 Aluminium railcar body for London’s Victoria Line.
produced more than 3000 m of friction stir welds in structural panels. They have welded floor panels and longitudinal girders for prototype aluminium trains of several train manufacturers. An important aspect, which generates increasing interest in FSW, is its potential to contribute to the crashworthiness of aluminium vehicles that could otherwise fail in the heat affected zone along weld seams. This has been observed in European accidents, notably in Eschede in Germany in June 1998 and Ladbroke Grove in Britain in October 1999. In the report on the latter it was recommended that consideration should be given “to the use of alternatives to fusion welding” and “the use of improved grades of aluminium which are less susceptible to fusion weld weakening”. The new designs have been subjected to extensive numerical and physical validations. Computer simulations predict that, with the new designs, overload fracture would occur in the parent material rather than along the weld line. The predictions have been verified by extensive full-scale component tests under high-speed impact at Bombardier’s test facility in France (Fig. 5.34). The newly designed FSW and MIG joints meet the stringent requirements for rail vehicle safety. Modern aluminium trains are designed to have well-defined crash properties, as was impressively demonstrated by the partially friction stir welded Alstom Pendolino train that derailed near Grayrigg in Cumbria (UK) in February 2007 while running at 153 km/h over a defective set of points. According to the Rail Accident Investigation Branch the train exhibited overall a good standard of crashworthiness and this helped to minimize the number of casualties and the extent of their injuries in this high speed derailment. In Europe the state of the art of applying FSW to railway rolling stock can be summarized as follows:
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(a)
(b)
5.34 Computer simulation (a) and (b) specimen after high speed impact test of a new joint design leading to fracture in the parent material away from the weld.
∑ ∑
Steel, stainless steel and aluminium all find favour as materials for rail vehicles. Whatever the material of construction, crashworthiness considerations now play a more and more important part in design. The standards in Europe relating to fabrication are changing and manufacturers will need to be able to demonstrate that their welders,
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welding facilities and systems have been approved to the new standards. ∑ The use of friction stir welding for the fabrication of aluminium rolling in Europe is aided by specialized supply chain companies which provide sub-assemblies to the rail vehicle builders. ∑ Friction stir welding has found commercial application to aluminium rolling stock, and several machine manufacturers can provide suitable welding machines. ∑ Five railway companies (Alstom, CAF, Angel Trains, HSBC Rail and RSSB) have teamed up in the EuroStir® project, to get FSW out of the laboratories and into the industrial manufacturing workshops with the intention to generate additional applications for FSW within the rail industry.
5.4.2 Rolling stock in Japan Hitachi in Japan uses the double-skin design of the car, which is constructed from friction stir welded aluminium extrusions. The main reason for this is the exceptionally low distortion of the FSW process. This contrasts quite markedly with the distortion that can occur when arc welding thin gauge aluminium and eliminates the need for straightening and filling. To date, Hitachi has delivered a range of vehicles for both commuter and express use in Japan (Fig. 5.35), and this has been recognized in Japan by the award of the prestigious Okouchi Award jointly to Hitachi and TWI. Hitachi has now begun to export FSW trains to Europe, e.g. the ‘Olympic Javelin’ trains for the domestic services on the Channel Tunnel Rail Link in the UK (Fig. 5.36). Nippon Sharyo and one other Japanese company have been using friction stir welded panels produced by Sumitomo Light Metal Industries for the floor panels of the new Shinkansen (Figs 5.37–5.38). Nippon Light Metals have also made use of friction stir welding for subway rolling stock. By 1998 they reported that over 3 km of welds had been produced. Kawasaki Heavy Industries (Japan) are using friction stir spot welding to attach stringers to roof panels (Fig. 5.39). They developed a new aluminium car body shell, which is assembled by this method. Main reason for Kawasaki Heavy Industries is the fact that it improves the flatness and visual appearance of the skin panels because of the low heat input.
5.5
Automotive
In 1998, TWI started a study on aluminium tailored blanks for door panels (Fig. 5.40) and demonstrated new concepts on FSW drive shafts and space frames in a confidential group sponsored project involving BMW, Chrysler,
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5.35 Commuter train built by Hitachi with full length friction stir welds of double-skin side and roof panels.
5.36 Hitachi uses FSW for assembling the new Channel Tunnel Rail Link domestic trains in the UK.
EWI, Ford, General Motors, Rover, Tower Automotive and Volvo. As a consequence of the encouraging results of this project, FSW and its variant friction stir spot welding (FSSW) are now being used in the series production of aluminium automotive components at several locations worldwide.
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5.37 Friction stir welded floor panel produced by Sumitomo Light Metal for Shinkansen trains.
5.38 Trainsets with FSW floor panels of Sumitomo Light Metal operate on the Shinkansen in Japan.
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5.39 Kawasaki uses friction stir spot welding for making aluminium roof panels, as shown in Kawasaki Technical Review, No 160 (Jan 2006).
5.40 FSW tailor welded blank produced from 6xxx series aluminium in 1998 at TWI for BMW and Land Rover.
5.5.1 Closures, tanks, suspensions and pistons by international suppliers Ford in Detroit (USA) uses a friction stir welded centre tunnel for the Ford GT sports car. The centre tunnel is a structural part that increases the rigidity of the chassis and is also used as a vapour tight fuel tank (Figs 5.41–5.42). The location of the tank provides good weight distribution and crashworthiness. The mechanical components, including the fuel pumps, level sensors and
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5.41 The friction stir welded aluminium centre tunnel of the Ford GT houses the fuel tank.
5.42 Friction stir welding of the centre tunnel of the Ford GT.
vapour control valves are first mounted on a steel rail. Then, a single-piece tank is blow-moulded around the rail. This ‘ship-in-a-bottle’ design concept maximizes the fuel volume and reduces the number of connections to the fuel system.
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Tower Automotive in Grand Rapids (Michigan, USA) produces aluminium suspension links for Lincoln Town Cars designated as stretched limousines. These have heavy-duty rear axles installed, while the rest of the rear suspension remains unchanged. The suspension link is made from two identical extrusions, friction stir welded simultaneously with two spindles from both sides (Fig. 5.43). This provides excellent fatigue properties. Showa Denko in Oyama City (Japan) joins extruded end-pieces to 20–30 mm diameter tubes for the manufacture of suspension arms. The rubber of the end-pieces of the suspension arms can be vulcanised prior to welding due to the low heat input of the new assembly method (Fig. 5.44).
5.43 Friction stir welded suspension links for Lincoln stretched limousines.
5.44 FSW suspension struts by Showa Denko.
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Riftec in Geesthacht (Germany) provides subcontract production and engineering consultancy, e.g. during the installation of FSW robots in automotive manufacturing lines. One of their automotive-related projects concerned the production of welded test specimens for study of the Berlinbased company Inpro (Fig. 5.45). Riftec is now a full service provider and assists its clients where necessary with product design changes that make it possible to take advantage of the FSW process for series production. Furthermore, Riftec is aided in product development by the fact that it has both 3D capable robotic (5-axis) and bed type vertical (3-axis) FSW machines equipped with complete process monitoring capabilities. Their equipment does not only allow for flexibility, when it comes to multi-dimensional welds, but also provides traceability for in-process and post-process quality control. In a world first industrial application, Riftec has since 2006 been supplying friction stir welded tailored blanks, which are subsequently deep-formed for making the centre closing panel of the Audi R8. These tailor welded blanks with dissimilar thickness of 1.7 and 2.4 mm have a 240 mm long friction stir weld and are used for the series production of the Audi R8. Here Riftec supplies approximately 7000 units or 14 tonnes of these tailor welded blanks per year (Fig. 5.46). This leads to more than 20% material savings, i.e. approximately 1kg weight saving per car. FSW has not only allowed Audi the ability to reduce vehicle weight but also increase efficiency due to a reduction in material and forming cost. Riftec uses a Tricept-9000 parallel kinematics robot for the commercial production of these blanks. This has a Sinumerik 840D control system. With its modified hard and software it has a work envelope of 2700 ¥ 3000 ¥ 900 mm and can weld up to approximately 20 mm thick workpieces.
5.45 Robotic FSSW of automotive parts for attaching skins to spaceframes.
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5.46 Tailor welded blank by Riftec for the Audi R8.
Sapa in Finspång (Sweden) uses a purpose-built FSW machine with two welding heads for welding hollow aluminium extrusions from both sides simultaneously, to produce foldable rear seats of the Volvo V70 station wagon. The machine has a carousel-type loading and unloading station and is automatically loaded by an articulated arm robot (Fig. 5.47). Halla Climate Control in Pyungtaek City (Kyunggi-Do, Korea) uses FSW to assemble small hollow pistons from aluminium castings. These are used in the compressor stage of automotive air conditioning systems (Fig. 5.48). First they machine the two piston parts and align them, then they friction stir weld them, and finally they post-weld machine them to the final dimensions. Mazda in Hiroshima (Japan) uses friction stir spot welding for the rear doors and bonnet of the Mazda RX-8 (Figs 5.49–5.50). The hood of this
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5.47 FSW production of Volvo rear car seats at Sapa.
5.48 Assembly, FSW and machining of pistons at Halla Climate Control.
sports car has an impact-absorbing structure aimed at enhancing pedestrian protection. They use this process, to avoid spatter and to reduce the energy consumption significantly in comparison to resistance spot welding. Friction Stir Link in Waukesha, Wisconsin (USA) is a service supplier focussing on the automotive industry. It provides FSW process development,
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5.49 Robotic FSSW of the rear doors of the Mazda RX-8.
5.50 FSSW rear door outer panel of the Mazda RX-8.
technology transfer, moderate-volume production and robotic friction stir welding system integration services (Fig. 5.51). Sapa produced a friction stir welded prototype engine cradle recently. The cradle is the result of a lightweight study to reduce the weight in the front end of the vehicle. The weight of this substructure is 16 kg, as compared to 23 kg for the steel version. This assembly uses various semi-fabricated products and joining methods. The side members are hydroformed aluminium extrusions. The front cross member is a straight extruded member. The rear
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5.51 CNC controlled FSSW gun on an articulated arm robot at Friction Stir Link.
cross member is built up of a sand cast part and a plate joined to the casting by FSW. The FSW operation was executed in three dimensions. A cost analysis showed the concept to be competitive to other concepts within the framework put forward by the customer. Pierburg in Düsseldorf (Germany) developed a concept for FSW of exhaust gas recirculation (EGR) coolers, a new product which lowers nitrogen oxides (NOx) of diesel engines by feeding cooled exhaust gases back into the combustion chamber. The cooler is made from die cast aluminium to provide a lamellar rib structure. All internal interfaces are welded by FSW instead of using heat-resistant seals and fasteners.
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5.5.2 Trailers in the USA Fontaine Trailer in Jasper (Alabama, USA) produces a revolutionary just 1.22 m high aluminium flat bed trailer. It weighs only 3.5 t, but can transport a maximum pay load of 27 t, thus they called it the Revolution™. It has a FSW aluminium main beam and a FSW floor with integrated side rails. Using FSW adds to the strength of the floor, which allows bigger payloads and reduced fuel consumption (Fig. 5.52).
5.5.3 Wheels in Australia, China and Norway Simmons Wheels in Alexandria (Australia) developed a new method of producing a wheel rim from rolled aluminium 6061-O sheet. From this they form a cylinder with a longitudinal friction stir weld. After cutting this into rim sections they spin form it into the desired rim profile and finally subject this part to heat treatment to the required T6 temper. The company supported UT Alloy Works in Guandong (China) during FSW production ramp-up of light alloy wheels (Fig. 5.53). A new technique of joining two parts of a car wheel has been invented, in which cast or forged centre parts are friction stir welded to rims that are made from wrought alloys. This concept has been industrialized by DanStir in Copenhagen (Denmark) and TWI in Great Abington (UK) on behalf of a leading Norwegian wheel supplier, Fundo Wheels in Høyanger (Norway). FSW allows for an internal cavity, and this reduces the wheel weight by 20–25% providing Fundo Wheels with a business advantage over its competitors, when selling to Volvo, Saab, Audi, VW and BMW.
5.52 FSW flatbed trailer of Fontaine Trailers.
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5.53 Aftermarket three-piece wheel made from FSW and spinformed aluminium cylinders at UT Alloy Works.
5.6
Other industry sectors
Friction stir welding is now being used in various industry sectors outside of the transportation industry. A trend towards non-linear and three-dimensional welds can be observed in these sectors. Friction stir welding is not only being applied to aluminium, but also to materials with higher melting points and dissimilar material joints.
5.6.1 Cathode sheets in Austria AMAG Industrial Services in Ranshofen (Austria) built a large gantry machine and commissioned this in September 2005 for the production of patented AMAG ProCath® aluminium cathode sheets for the zinc electrolysis (Fig. 5.54). The edges of these cathode sheets are hemmed with a patented procedure and provide for perfect stripping of the zinc layer of the cathode sheet metal. The edges are coated with plastic layer in a unique automatic manufacturing plant and connected then by the FSW procedure with the carrying hooks and beams. The low temperatures of FSW preserve the plastic film already applied, the so-called liquid zone protection, which does not to be replaced over the entire life span of the cathode sheet under normal operating conditions. At both ends of the support beam are friction stir welded bimetal contact plates, which provide an electrically highly
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5.54 FSW-Cathode sheet of AMAG.
conductive connection between aluminium and copper. The application of the FSW procedure for this alloy is more economical, more corrosion resistant and environmentally friendlier than conventional procedures as explosive welding. It is also possible to attach connector pieces during refurbishment of recycled cathode sheets by using this method. The gantry machine is now owned and operated by Hammerer Aluminium Industries in Ranshofen (Austria) and is also used to produce a wide range of panels from aluminium extrusions (Fig. 5.55). One of HAI’s first commercial products was a set of façade panels for Vienna airport (Fig. 5.56).
5.6.2 Motor and loudspeaker housings in Scandinavia Several manufacturers use the FSW process for producing housings. Hydro Aluminium in Magnor (Norway) produces FSW housings for electrical motors from four aluminium extrusions. The company PDC Teknik in Helsinge (Denmark) uses FSW for joining two hemispherical pressure die castings to produce loudspeaker housings for Bang & Olufsen.
5.6.3 Heat sinks in Europe Several companies in Europe are licensed by TWI to use the FSW process to produce heat sinks from aluminium extrusions, e.g. Sykatec in Erlangen (Germany), Austerlitz Electronics in Erlangen (Germany), EBG & DAU
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5.55 FSW gantry machine at Hammerer Aluminium Industries.
5.56 Hammerer Aluminium Industries produces a wide range of FSW façade panels.
Kühlerentwicklungs GmbH in Dietenberg (Austria), KMT Group in Kankaanpää (Finland), Profilgruppen Manufacturing in Åseda (Sweden), Euro-Composites® in Echternach (Luxembourg) and Sapa in Finspång (Sweden). These are some of the most cost effective applications of friction stir welding, with large lengths of welds produced every day.
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5.6.4 Heating, ventilating and air conditioning units in Germany When one looks to the future, particularly in relation to new products made from aluminium and the further development of aluminium and its alloys, there is no question that the benefits and innovative use of the FSW process will establish this process as an industrially relevant joining and material working process for years to come. One such example is the application of friction stir welded perforated plates supplied by Riftec for use in an ecologically friendly heating, ventilating and air conditioning unit called ‘Ökolüfter’ (eco fan), which was nominated for the European Aluminium Award 2008 (Fig. 5.57). Additionally, Riftec friction stir welds aluminium alloys which are not produced by the classical ingot route such as dispersion strengthened and spray compaction processed aluminium alloys like DISPAL® of PEAK Werkstoff GmbH. The significance here is that aluminium alloys obtained today by the classical route have generally been optimized by a careful choice of composition, fabrication, thermal and thermo-mechanical treatments. FSW has demonstrated consistently over the last decade its unrivalled ability to join these conventional alloys. It is well accepted, however, that improved mechanical, physical and chemical properties would be possible, if alloys with compositions inaccessible by the classical ingot route could be produced, or
5.57 FSW of heating, ventilation and air conditioning units by Riftec.
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if new processing and or joining techniques could be developed. Friction stir welding not only offers this possibility, but Riftec has effectively demonstrated this point in the case of the DISPAL alloys.
5.6.5 Backing plates for sputter targets in Japan Hitachi Cable Ltd and Hitachi Copper Products Ltd in Tsuchiura City (Japan) use friction stir welding in the mass-production of water cooled copper backing plates, which are the support panels of targets in sputtering machines. To produce these backing plates, meandering grooves are machined into thick copper plates to create a water channel, and non-linear friction stir welds are used to cover these grooves with a copper sheet. The FSW backing plates have an improved dimensional accuracy and surface finish in comparison to electron beam welded plates, because FSW takes place below the melting point. They are used as electrically conducting fixtures for sputter targets in vacuum chambers in the production of large LCD flat screen TV displays, semiconductor devices and CDs/DVDs.
5.6.6 Vacuum vessels in Switzerland, Japan, New Zealand and Germany Several fabricators in Switzerland, Japan, New Zealand and Germany apply the FSW process for making components for vacuum vessels or vacuum valves from aluminium plates, because they can thereby reliably avoid porosity and unwelded root crevices. This is important in the manufacture of vacuum equipment, to achieve a short evacuation time. Riftec GmbH in Geesthacht (Germany) has used FSW since 2004 for dissimilar joints between flat stainless steel sheets and domed aluminium sheets in a vacuum-tight component for X-ray equipment of Siemens Medical Solutions (Fig. 5.58). These are used in the entry screen to radiographic amplifiers of Siemens X-ray equipment. The benefit to Siemens has been a reduction in cost of approximately 20% since implementation of the FSW process. This has been achieved via a significant reduction in defective goods, both as a result of using the FSW technology and the quality assurance programme developed and provided by Riftec for series production.
5.6.7 Panels and components for the food industry in Denmark and Germany Drying trays for freeze drying in the food industry are another common application of FSW both in Denmark and Germany. These relate to the food industry where extensive use of aluminium is made. Aluminium has the advantage of high thermal conductivity, light weight and the ability to
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5.58 FSW of aluminium domes to stainless steel sheets for vacuumtight X-ray equipment by Riftec for Siemens Medical Solutions.
be formed in almost all imaginable shapes and sizes. Of particular concern in the food industry is that any joints containing flaws such as cracks and or pores, even if hair-line, are a hygiene issue and thus a potential health risk. As a consequence FSW was seen as a good solution to solving such healthrelated issues since the solid state nature of the FSW process avoids pore and or crack formation. Simply put, when there is no solidifying weld pool and or entrapment of gases, there is no potential for cracks or pores to form. FSW also has particular benefits, if the parts are anodized after welding. Since 2004, Riftec has successfully provided the food industry with annually approximately 7000 thin-walled and distortion-free food trays, which are absolutely flaw-free (Fig. 5.59). Not only has the welding process been able to meet all hygiene standards but it has allowed the client to reap considerable cost benefits compared to previously used TIG welding alternatives. Since 2008 Riftec has also been supplying Bizerba with friction stir welds for approximately 5000 articles per year made between the carriage and the hand and finger guard of their meat slicing machines. The challenge encountered by Bizerba was that the conventional machine design envisaged
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5.59 Aluminium freeze drying trays by Riftec.
a two component structure between carriage and guard. This meant that in order to clean between surfaces the carriage and guard would have to be dismantled each time after use. Consequently Bizerba approached Riftec to provide a FSW solution, which was successfully implemented.
5.6.8 Copper canisters for nuclear waste in Sweden Svensk Kärnbränslehantering (SKB) in Sweden is planning to encapsulate spent nuclear fuel in copper canisters and to build a final underground repository for these canisters in bedrock at a depth of 500 m. The canisters have a wall thickness of 50 mm, have a diameter of just over one metre and are nearly 5 m long. SKB have chosen to use FSW to encapsulate nuclear waste. The Canister Laboratory of SKB in Oskarsha mm (Sweden) co-operated with TWI to optimize the FSW process specifically for welding 50 mm thick copper, and currently applies the process under production-like conditions for non-radioactive dummy fuel.
5.6.9 Hunting knives in the USA DiamondBlade LLC in Denison, Texas (USA) apply friction stir processing for refining the microstructure of steel blades of hunting knives. They use polycrystalline cubic boron nitride (PCBN) tools by MegaStir for reprocessing
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and forging the cutting edge of knife blades. By doing so, the properties of D2 high-carbon high-chromium tool steel are substantially improved regarding hardness and corrosion resistance. During friction stir processing, the D2 steel becomes plasticized and heats up to the eutectoid temperature. At this temperature, chromium carbides begin to dissolve, until chromium and carbon are quenched into very fine and corrosion resistant martensitic grain structures. After friction stir processing, quenching and tempering, the wear resistant cutting edge has a hardness of 65–68HRC on the Rockwell scale, while the tough spine of the knife blade stays at approx. 45HRC. Friction stir processed blades are hard and tough as well as corrosion and wear resistant. The innovative manufacturing process offers additional benefits in terms of reduced energy input, greater controllability and the elimination of dirty, hazardous and occasionally environmentally problematic heat treatment processes.
5.7
Acknowledgements
The author wishes to thank the companies mentioned for permission to publish photographs and information on their products and manufacturing processes.
5.8 ∑
Conclusions
Friction stir welding is a remarkable new welding method that has grown rapidly into an important industrial process since its invention, by TWI, in 1991. ∑ All transport industries including shipbuilding, offshore, rail, automotive, and aerospace are using this technology. ∑ The main reasons are the cost savings due to high robustness, good repeatability, excellent mechanical properties and low distortion. ∑ Further research and development work is currently underway to assess new FSW joint designs, to establish further mechanical and corrosion data, to specify procedures for the FSW of steel, titanium and other challenging materials and – finally – to develop new applications of this remarkable process.
6
The future of friction stir welding
P. L. Threadgill, formerly of TWI Ltd, UK
Abstract: This chapter is the author’s personal view of the perceived future of friction stir welding. The process has made a very rapid start, and its success continues to grow, particularly in light metals where barriers to further growth are largely dictated by the need for investment in change and equipment, and the acceptance of a radically new process. With high temperature materials, there are significant technical barriers in tool technology, which need to be overcome if the process is not to be restricted to niche applications. Key words: friction stir welding, tools, aluminium, magnesium, steel, titanium, nickel.
6.1
Introduction
The time between invention of friction stir welding and its initial industrial implementation was one of the quickest known for any welding process, although precise data on other processes is not easy to establish. However, friction stir welding was invented in 1991[1], and its first industrial use in was 1995. What is particularly remarkable about this is that the process really did not work very well at the time of its invention. Another remarkable fact of the early development is that there were no suitable machines on which to develop the process. Early trials invariably used milling machines, but these were barely suitable for laboratory work, and without some modification were far from ideal for industrial use. This was recognised at an early stage by ESAB in Sweden, who not only realised that successful implementation of the process would require special machines, but also that friction stir welding could be a competitor to the established arc welding processes on which their equipment and consumable manufacturing were based. History has shown that entering the friction stir welding market was a wise decision, and their lead has been followed by several other companies, so that now fabricators have a good choice of high quality equipment suppliers. The growth of the friction stir welding market is difficult to quantify, certainly in terms of kilometres of weld made, or tonnage of components supplied. Unlike some arc welding processes, there are no external indicators such as tonnage of welding wire sold, and the only real indicator of growth 164
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is the number of licensees of the process. TWI’s licence terms are based on site licensing, and not on quantity of welds produced, and there may be (and often are) several machines on a single site. However, as shown in Fig. 6.1, the growth of licensee numbers has been fairly rapid, although the rate of growth has shown signs of slowing down [2]. This is believed to be due to a number of factors: 1. The obvious major manufacturing companies have already acquired licences, and therefore future growth will be largely in smaller companies. The investment cost for such companies will generally be a higher percentage of their budgets, and so the decision to make the change will be more difficult. 2. The various TWI patents will start to expire from about 2014 onwards, and so some smaller companies may defer the decision to implement the process until after that date, in particular where the cost of the licence fee is seen as a significant barrier. 3. Although becoming established, friction stir welding is not a process which is known or understood by all parts of the engineering community in the way that arc welding is. Therefore, fabricators trying to adopt the process may have to convince their customers and appropriate regulatory authorities of the benefits of the process. 4. Skilled practitioners in the art of friction stir welding are hard to find, and the skills are certainly not yet taught in technical colleges, as the volume of industrial work cannot sustain this investment. 5. The cost of machinery is high, as each machine is designed individually to undertake a certain number of tasks, and therefore the design costs
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6.1 Growth of friction stir welding licensees.
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2001
2000
1999
1998
1997
0
1996
50
1995
Total number of licences
250
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are high, and savings from multiple production of machines are virtually non-existent. Very few machines have seen repeat builds, and standard machines such as the ESAB Legio series and the MTS I-Stir series have tended to end up in universities and other research institutes, where the market is probably approaching saturation. 6. The high cost of bespoke equipment can only be justified in smaller companies if the volume of work is such that the machine can be usefully employed for a high percentage of its available time. 7. For many critical applications, a great deal of procedure development and qualification is required before the process will be accepted. This, of course, takes time and money, although the significant inroads which the process has made into the airframe and railcar markets is proof that this can be done. 8. There has been an overwhelming number of patent applications from many sources covering various aspects of the process, or the applications for which it may be used. The number of known patents and applications stood at over 2100 by late 2008, of which over 600 have been granted, and the numbers continue to increase [2]. Analysis of the patent applications shows that the majority of these originate from large companies, predominantly in Japan. However, there are signs that the number of new patent applications is dropping slightly, but as information on new patent applications is not always easy to find, the extent of the slow down is difficult to quantify. Many of the patent applications are at best trivial, and would be difficult to defend, but nevertheless their presence in large numbers is seen as barrier to implementation of the process, in particular by small companies. The growth of patent applications is shown in Fig. 6.2 [2]. In this chapter, process and materials aspects will be considered, and their possible impact on the future of friction stir welding discussed. The ideas are those of the author, and represent a personal viewpoint at the time of writing. The historical development needs for the process have changed with time as the process has improved. The first significant challenge at TWI in the early 1990s was to repeatedly make a 2-metre long weld, without flaws, in 6mm AA6082-T6. This was achieved, but only after considerable refinement of tool designs and the development of what is now regarded as a primitive control system. This success was followed by development needs related to higher strength alloys, especially the difficult to weld AA2xxx and AA7xxx alloys, and eventually to “thick” materials, which in the mid/late 1990s meant 25mm. Other challenges have been directed towards more difficult materials, particularly thick section copper (which has been achieved), and welding titanium, steels, nickel, etc. These development needs are still being addressed in many places, although significant progress has been made.
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2500 USA 2000
Japan Europe
1500
1000
2008
2007
2006
2005
2004
2003
2002
2001
2000
1999
1998
1997
1996
1995
1994
0
1993
500
1992
Total number of patent applications
PCT
Year
6.2 Growth of patent applications related to friction stir welding.
6.2
Process aspects of friction stir welding (FSW)
6.2.1 Comparison with other processes Growth of friction stir welding will be largely at the expense of other joining processes. In some areas, for example airframe manufacture, the indigenous process (mechanical fastening) is slow and expensive, and so the opportunity to replace it with a mechanised process, which can give good results in high strength alloys, is attractive. In this case, friction stir welding is particularly relevant, as many high strength aerospace alloys are difficult to weld by fusion processes. However, in other industries, replacement of MIG and TIG welding by friction stir welding will never be complete, as these processes offer capabilities with which friction stir welding cannot compete. MIG and TIG welding have the advantage of using a filler (required in MIG, optional in TIG), and this allows the fabrication of joint designs where additional metal deposition is essential, for example in fillet welds, and in butt welds where the fit-up is variable or difficult to control. Thus, in many applications (shipyards, railcars), long straight 6xxx extrusions are welded very economically by friction stir welding (where the absence of a filler in FSW is a distinct advantage), but other attachments, e.g. bulkheads, stiffeners, closing welds, etc., are made by MIG welding. Similarly, longitudinal welds in large tanks are easy to fabricate using friction stir welding, but the dome ends are generally more easily welded by fusion processes as this removes the need for internal support to react the process forces.
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A future prospect for FSW is therefore an optimised integration of complementary welding processes. Although FSW can be adapted to make the closure welds, the technology for this has lagged behind that required for the longitudinal welds. Another major asset of MIG and TIG is that these can be manual processes, and therefore offer improved flexibility, especially when dealing with variable fit-up, difficult access, etc., although there is a risk of more variability in quality than with mechanised welds. Another factor is that the process can be taken to the work, whereas in friction stir welding the process at present has virtually no portability. Although some progress has been made in transportable rather than portable friction stir welding equipment, this is an area that is not yet well developed, although a successful development would open up numerous markets for the process [3]. Good progress has been demonstrated on portable systems with aluminium pipe welding, and shipbuilding. It must also be remembered that other competing processes are also developing. In particular, the last decade has seen enormous progress in the developments of laser technology, and the emerging fibre and disc lasers offer substantial opportunities for welding. Similarly, arc welding has become more sophisticated, as computer technology has underpinned the development of more complex power sources.
6.2.2 FSW process variants Although now out of its infancy, the friction stir process continues to develop at a respectable pace, and this is not expected to decline in the near future. There have been a number of attempts to define derivatives to the process, although most of these are of potential benefit to a limited range of applications, and generally involve far more complex tool and/or equipment designs. For these reasons they have so far failed to make any real impact despite their claimed benefits. For example, Skew stirtm [4] and Com-stirtm [5] would fall into this category. Friction stir processing is a variant on friction stir welding, which has been investigated as a method of improving microstructures in a variety of materials, especially aluminium alloys and naval bronzes [6]. The objective of the process is to refine grain structures to improve properties such as strength or formability, or to improve material quality by removing imperfections, principally removal of porosity and other solidification defects in castings. Friction stir processing of castings can also break up coarse and highly segregated microstructures, as well as intermetallics, etc. This is, of course, an extra operation, and therefore an extra cost, but there is evidence that this may prove to be a good investment. Friction stir processing has demonstrated its usefulness in promoting superplasticity in several aluminium alloys, and it is expected that it will contribute significantly in this specialist market.
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One interesting application is that of using friction stir welding to replace diffusion bonding in the superplastic forming/diffusion bonding (SPF-DB) process [7]. At the moment, SPF-DB is only viable for materials which are easily diffusion bonded, such as titanium, but replacement of the diffusion bonds would allow the technique to be developed for a range of aluminium alloys, especially for those where the friction stir processing significantly improves the ductility of the material. The approach has been demonstrated, and it is clear that further work is needed, but this technique could have a significant future, and its full potential would appear to be considerable. Another major variant is friction stir spot welding (FSSW), which is already in production for aluminium welds in the automotive industry, and is undergoing intense development for very high strength automotive steels [8, 9], which can be more difficult to weld by resistance spot welding. A major advantage of FSSW is the lower energy requirement, as there are no high currents and consequential I2R heating losses. It is firmly expected that this will be an area for major growth, primarily in the automotive industry, and it is a further advantage that the process can be easily adapted for robot use.
6.2.3 FSW tool design There is almost an infinity of tool designs, although most production work uses tools which are based on designs which have been established for many years. New designs tend to differ only in detail, and it is unlikely that further variation on these will lead to significant improvements in the performance of the process or the properties of the welds, although benefits related to specific applications are likely. One significant design variant which is more likely to succeed is the bobbin tool (generally known as the self-reacting tool in the USA), as there need not be any complex engineering involved. The idea is not new, as the basic concept was described in the original TWI patent [1]. There have been a number of approaches to bobbin tool design, and different design approaches have required wide differences in complexity. For example, in some designs the distance between the shoulders can be varied to control the force [10], although such complexity may not always be necessary. The underlying benefit of the bobbin approach is that where access is available from both sides, use of a bobbin tool will reduce the high down force to very close to zero. This in turn has the potential for a significant impact on required machine stiffness, and hence cost, and will open up the market to allow relatively simple modifications of milling machines, as one of the primary weaknesses in milling machines is their inability to apply high downloads without considerable flexing. An additional benefit is that the heat generation is more or less uniform between the top and bottom surfaces, and
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in thicker section materials this appears to reduce distortion, although the same effect could no doubt be obtained by using independent tools on the top and bottom surfaces, welding at the same time. Bobbin tools are also capable of following complex profiles in two and three dimensions, giving great flexibility to this approach. Although data is rather scarce, it appears that the total energy input into a weld is not significantly altered by the use of a bobbin tool [11], and so it is reasonable to assume that the mechanical properties will not be radically altered. A further area where tool design is developing is in tools for thick section welding, and these developments have been made primarily for aluminium and copper alloys, although success has also been reported for magnesium alloys. Aluminium and copper both have high thermal conductivities, but this is generally a problem in thick section as the large bulk of material acts as a very effective heat sink. In conventional tools designed for thinner materials, the bulk of the heat is generated beneath the shoulder, and the high conductivity of the aluminium work piece ensures adequate distribution around the tool. However, above certain thickness (~20–25 mm in aluminium alloys), the large distances over which heat must travel, plus the effective heat sink mean that the redistribution of heat from the shoulder to the tip of the tool is not sufficient. For this reason, single-sided tools for thick section have a relatively small shoulder diameter compared to the length of the probe, so that a much greater percentage of the heat is generated by the probe. An additional point which needs to be considered is the huge bending moment on a long probe, and this is normally overcome by using tapered probes, and also by using tool materials with greater high temperature strength than the normal hot working tool steels employed for aluminium welding. The high energy input also means that extended thermal cycles are inevitable, and this will compromise properties of welds, especially in materials which have been aged or work hardened to high strength conditions. Successful welding requires careful management of the heat generation and removal, which is not a simple task. Fortunately, experience to date has indicated that success can be achieved, and therefore further growth in this area, especially with bobbin tools, is expected to continue. Challenges also exist in tool design (and process control) for welding very thin materials (<1 mm). This has been achieved in a basic fashion for aluminium [12, 13] and magnesium [14] alloys, but it is difficult to machine fine details on such small tools, and in fact this may not be necessary. There are a number of applications, particularly in electronics packaging which may benefit from development of micro-friction welding, but at the moment the technology needs further development. It is understood that Airbus have made significant progress in the friction stir welding of 0.3mm thick aluminium alloy. It is believed that the maximum and minimum thicknesses which can be friction stir welded are more dependent on the
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design of the equipment, the tool and the control system, rather than any materials related issues, but more work would be required to confirm this.
6.2.4 Productivity As with virtually all manufacturing processes, there is a demand for higher productivity, which in welding processes usually means higher welding speeds. However, it is emphasised that welding speed is really only part of the challenge, and what should be addressed is the total joint completion time. In many cases, there is significant time spent in joint preparation, and in loading and unloading materials into the welding machine. Savings here may be more effective than shaving a few seconds off the welding time. A further issue is that welding at high speeds requires closer process tolerances, (and hence more investment and therefore cost in joint preparation and fitup). High speeds also increase stresses and hence increase wear on the tool. In combination, these can lead to an increased incidence of defects, and consequent repairs or scrapped components. The optimum welding speed is therefore not normally the fastest possible speed, and attention to other process aspects may be more productive. There is a limit to welding speed in friction stir welding. The heat generated at the tool must diffuse ahead of the weld to apply some level of preheat. Pushing a tool into cold material will clearly increase stresses, as noted above. Since the linear heat input is approximately inversely proportional to travel speed, welding at higher speeds will require increased rotation speed to generate more heat. However, there is a limit here, as most materials have a maximum shear strain rate which they can endure, and this is determined by the rate of recovery in the highly deformed material. It is common knowledge that an excessive rotation speed under force will cause break-up of the material surface. High linear heat inputs will also cause greater time at elevated temperatures sufficient for microstructural changes to occur, and hence more degradation of properties in heat treated or mechanically hardened alloys. Total process time can also include one or more post weld treatments (e.g., machining flash, QA heat treatment, etc.), and efforts to minimise or automate these will have significant benefits. For any particular material the key to improved travel speeds is clearly in the tool design, and progress will no doubt continue to be made. However, the other restrictions mentioned above are material specific, and it is considered unlikely that significant improvements in travel speed can be made without compromising other aspects of the weld.
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6.2.5 Standards As friction stir welding is still a relatively new process, it can be no surprise that standards for the process did not exist until recently, At the time of writing, an American Welding Society standard (D17.3:200X “Specification for FSW of Al alloys for aerospace applications”) has been published for aerospace alloys, but a more general standard is in the final stages of preparation as an ISO standard (ISO 25239, “Friction stir welding – Aluminium”), and is expected to be published in 2009. Clearly, the appearance of standards by internationally recognised bodies will add significant credibility to the friction stir process, and this again will act as a driver for future growth. In the absence of these standards, friction stir welding has been accepted by various classification societies on a case by case basis, and this trend will no doubt continue.
6.2.6 Process modelling It is not surprising that the emergence of a new manufacturing process is accompanied by numerous efforts to model many aspects of it, and friction stir welding is no exception. It is particularly challenging as, in addition to heat generation and heat flow, complex mass transport phenomena must also be addressed. This is complicated by the fact that the process requires very high strains and strain rates at very high temperatures, which presents challenges to the numerical methods involved, and also a lack data with which to calibrate the models. Despite these problems, very real progress has been made, and our theoretical understanding of the process has improved as a result. Modelling may be a useful tool to develop new tool designs, and it is worthy of note that one concept (the “Trivex” design) emerged directly as a result of modelling work [15]. Modelling offers great prospects in the future for reducing experimental effort in development of welding parameters, tool design, machine design, clamping systems and many other areas, and with the large effort currently being made in modelling it is reasonable to expect a significant return on this investment. Such benefits are already beginning to emerge.
6.3
Materials aspects of friction stir welding (FSW)
6.3.1 Introduction Perhaps the best way to view the future is to look at each alloy type where it might be used, and to consider the factors which limit or support future growth. Alloys have been grouped into two groups, depending on the
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temperature needed to weld them. Further discussion is held on joining dissimilar materials and polymers.
6.3.2 Tool materials for high temperature use There is no doubt that the biggest issue facing the commercial implementation of friction stir welding for high temperature materials is the tool design, and the material from which it is made. There are many papers published which demonstrate quite clearly that good quality welds can be made in materials such as steel, titanium and nickel. However, the problem which keeps recurring is the inadequacy of the tools, and in particular the materials from which they are made. The two principal material types have been based on tungsten rhenium (W-Re) alloys, and polycrystalline cubic boron nitride (PCBN). The advantages and disadvantages of these are summarised below: 1. Source of materials: Both W-Re and PCBN are sourced from single suppliers. In the case of W-Re, material is supplied from which end users machine their own tools. PCBN tools are also currently available from one source only, and the tools are supplied already ground to size and mounted in holders. 2. Cost of tools: The price of Re has risen rapidly in recent years, making W-Re an expensive material. Manufacture of PCBN is a complex process requiring expensive equipment, and therefore the cost of tools is also high. These costs are at present too high to allow FSW of steels for mainstream activities, especially as the life of current tools is generally limited. 3. Quality of materials: There has, without any doubt, been a very significant improvement in the quality of W-Re and PCBN materials over the last five years, as considerable investment has been made into understanding and improving the manufacturing routes and the specific demands placed on the materials. In the case of PCBN, there have also been significant improvements in the design of the tools. This progress is expected to continue, but it is not clear what further investment in time and effort will be required. Other solutions have been proposed, for example other refractory metal alloys (e.g. W-La2O3, Mo-TZM, etc.), cermets based on mixing PCBN and W-Re [16], tools based on iridium alloys [17] and intermetallic alloys, as well as various coatings and surface modifications. Although these and other current research topics may bear fruit, there is no certainty that the results will be economically attractive. Almost all solutions for high temperature tools require the use of very expensive materials whose price is often fluctuating, and users would expect guaranteed tool lifetimes before paying the current high prices. There are, however, two exceptions to this model. The first is
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the use of relatively low cost ceramic tools, based, for example, on silicon nitride, which has shown an acceptable performance in several applications but at a price where the tool is considered as a consumable. Progress has been made in this area on both friction stir spot welding consumables for high strength automotive steels [18], as well as in conventional butt and lap welding of steels and other alloys. For spot welding, the tool should ideally survive for a whole shift, in which at least 3000 welds would be made. The second area is the “niche” market area, where process performance is the primary criterion, and cost is a secondary, although generally still significant, issue. Such applications are usually found, for example, in defence areas, very critical structures (e.g., aerospace and high pressure components) and in some high performance sporting equipment. In these cases, quantities of welds would be relatively small. Other solutions have been investigated, but these developments have generally been protected by patents or patent applications, and so it is reasonable to expect that there will be some financial obligations for potential users. In a process which is so immersed in patents this comes as no surprise, but at the moment there are no materials which have broken through to a point where they are attracting high levels of interest. However, the economic benefits to the suppliers of the perfect tool material are potentially huge, and so further development and research will continue. There has been some interest in the use of hybrid heating systems, for example laser welding or induction heating, to reduce the loading on the tool by preheating the material being welded. These methods are still under development, and will add to the cost and complexity of the process, but this may be a worthwhile price if it opens up the possibility for welding at high temperatures.
6.3.3 High temperature alloys Titanium Titanium is something of an enigma in friction welding, as it is without doubt one of the easiest materials to weld using rotary or linear friction welding, but one of the hardest to weld by friction stir welding, and yet the primary cause is the same, namely its low thermal conductivity. In friction stir welding using conventional approaches, most heat is generated under the shoulder, and relies on good thermal conductivity to be distributed around the weld area. In high conductivity materials such as aluminium, this is not generally an issue (except in thick sections), but in titanium alloys redistribution of heat is a major problem. This is compounded by the relatively narrow temperature range within which titanium alloys can be hot worked, and so friction stir welding requires very even heat distribution
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through the thickness. The accepted approach is to rely on the probe to generate the heat rather than the shoulder, and several alternatives exist, in which either the shoulder is very small [19], or designed to be non-rotating [20], i.e. generating no heat at all. Both methods have claimed great initial success, but at the present time both are in later stages of development. The significant point is that these improved and relatively simple modifications to the process, developed in response to difficult materials, will allow the successful industrial welding of materials of low thermal conductivity in the short term, at least for titanium. Titanium alloys are notorious for their reactivity at high temperatures, even in the solid state. Therefore selection of tool materials has caused difficulties, as titanium will form low melting eutectics with nickel, cobalt and iron. Refractory metals are fortunately virtually inert to titanium, and so most tool development has been based around tungsten alloys. There are certain applications where the use of tungsten is not ideal, and so further research is ongoing to identify alternatives. Steels At the present time, there is no reasonable prospect that the traditional arc welding processes will be displaced from mainstream activities in steel fabrication by friction stir welding. It should not be forgotten that steel is the mostly widely welded alloy, with probably >90% by weight of all welded structures being made of steel. Developments over the last fifty or sixty years or so in steelmaking technology have recognised and responded to the need for good weldability, and other research has led to high quality consumables and process equipment at acceptable prices. Arc processes cannot be surpassed for flexibility of joint design, manual and machine capability, abundance of skilled practitioners at operator and supervisory levels, relatively low costs and a level of experience which cannot be challenged. However, the potential for friction stir welding of steel is high, but for niche applications. The process is being actively developed for various shipbuilding activities (much of this is military), and for other critical applications such as welding of very high strength pipelines, where conventional arc welding consumables are not adequately developed. The potential for lower distortion has also attracted interest from shipyards, and there is now very active research (and some very encouraging results) on the use of friction stir spot welding for ultra-high strength automotive steels. The biggest problem in all of these potential applications, as mentioned already, is the lack of a reliable tool material. It should be noted that steels of virtually all types have been successfully friction stir welded and so the future opportunities for the technology are clearly very large.
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Nickel alloys The total tonnage of nickel alloys used is relatively small, but much of its use is in critical applications, for example jet engines. Weldability has always been an issue, although the weldability of the mainstream alloys, e.g. IN718, is well understood, and alloy developers are aware of the importance of this property. Alloys capable of service in more extreme environments are generally more challenging to weld. At present, friction stir welding has only been demonstrated for solid solution hardened and moderately performing nickel alloys (e.g., IN600, IN625, IN718), and although the results are encouraging, further work is need to optimise the process and fully characterise the welds. Like titanium alloys, nickel alloys normally have very poor thermal conductivity (although it increases at the welding temperature, and so tool design approaches developed for titanium may well be helpful for nickel alloys. There is no doubt that an important niche opportunity could arise for friction stir welding, but this will require further research, and probably significant improvements in tool materials and design.
6.3.4 Low temperature alloys Aluminium alloys At the time of writing more than 99% of all friction stir welds are made in aluminium alloys. This figure is unlikely to reduce by very much in the near future, as new aluminium applications seem to appear as fast as new applications in other alloys. The mainstream activity for friction stir welding of aluminium is in long straight welds, and in this area the economics are so favourable that the process is hard to beat. Even so, adoption of the process has a long way to go, and is little used, for example, in products such as aluminium truck chassis, truck bodies (skip type and tankers), unpressurised or low pressure tanks for storage, processing, etc. There is also a growing market for the process in thicker section materials, in particular for military applications where joining high strength armour has presented a number of challenges to which friction stir welding has responded well. In addition, lightweight transportable bridges and other transportable structures are areas where the process could make a significant impact, and progress is now being made in these areas. The economics of the process look particularly attractive where applied to thick section materials. Although the welding speeds are necessarily rather slow, the joint completion time is much less than for multi-pass arc processes, where interpass grinding, cleaning, inspection, etc., is required. However, the investment in equipment becomes very high for thicker materials, and so machine usage becomes a very significant factor in the economic equation.
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However, the issue of the large thermal cycles and potential for reduced mechanical properties must also be addressed. Friction stir welding has not been widely used to date for pressure vessels, although many aluminium pressure vessels are manufactured. One of the barriers has been that the principle design code for pressure vessels has been the ASME code, which does not at present include friction stir welding among the list of approved processes. This is changing however, as friction stir welding was accepted in a code case in early 2008, and this will lead to eventual adoption of the process in the main body of the code. The process to achieve this is in hand, but may take time. This important development is expected to open up the market initially for lower pressure applications such as aluminium heat exchangers and some process plant, and eventually for vessels for hydrogen storage, where much higher pressures are required. Current designs for such structures require wrapping of aluminium primary vessels with carbon fibres to achieve the desired performance. Magnesium The tonnage of magnesium produced each year is increasing at a faster rate than most other engineering materials, and this is driven primarily by demand from the automotive industry in their attempts to reduce vehicle body weights. However, virtually all of the magnesium alloys which end up in road vehicles are castings, and the requirement for welding by any process is rather limited. This applies to both bodywork and engine/transmission parts. There have been some applications where pressure die castings have been welded, and the results are generally very good, but it is often possible to make the components in a single piece, and from a manufacturing point of view it is clearly preferable to make one large casting than to weld two or more smaller ones together. However, smaller simple castings are easier to make, with lower reject rates, and so an opportunity to integrate friction stir welding into the manufacture of cast components may exist. It is also possible to improve the quality of magnesium castings by friction stir processing, and it is worthy of note that there is no evidence that friction stir welding of magnesium generates any significant fire risk. Copper Copper is a difficult material to weld, due to its high thermal conductivity, but the technology to allow friction stir welding has been well demonstrated, even in single pass welds of 50mm thick. The process is used for the fabrication of heat exchangers, and other components, and this market will presumably grow. There is growing interest in the use of friction stir welding to join copper to aluminium, mostly for electrical conductors, and the feasibility
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of doing this has been demonstrated, even though the technology may not be sufficiently refined for commercial usage. This is, however, an important area, and further development is inevitable.
6.3.5 Dissimilar materials welding One area where friction stir welding has shown enormous potential is in the area of dissimilar metal joining. This can include joining dissimilar alloys from the same family (for example, joining aluminium castings to wrought aluminium products), and joining alloys from different families, for example joining aluminium to steel. In the former case, it is often not difficult to join dissimilar but weldable aluminium alloys by fusion processes, as the fillers are almost always one of two types. However, the usual problems encountered when joining aluminium alloys, for example porosity and other solidification-related issues, must still be surmounted. Aluminium castings, especially pressure die castings, can contain pockets of very highpressure gas which makes fusion welding a challenge. However, there are many examples where friction stir welding has been successful in joining dissimilar aluminium alloys. One feature of the process is that the level of mixing between the two components is often very limited, and so the tendency to form alloys of unpredictable performance is limited. The high forces involved will also heal voids and other imperfections which may be present, particularly in castings. Probably the only potential challenge is to develop where needed a post weld heat treatment which is compatible with both alloys, but this is not likely to be an insurmountable problem. Dissimilar joints have been in commercial usage in critical applications for some time, for example in the Eclipse jet aircraft where 7xxx stringers are lap welded to 2xxx exterior skins [21]. There is also a wealth of experience on joining other dissimilar alloys, for example magnesium alloys, steels, and a lower level of experience on titanium and copper alloys. However, the future looks very attractive for this activity. Joining alloys from different systems is often very difficult. There is always a risk of formation of brittle intermetallic compounds, as well as the formation of low melting point eutectics, and in some alloy systems (for example, aluminium to magnesium) both challenges exist. In addition, widely differences in melting points can be a further complication. Success has been reported in, for example, aluminium to copper for electrical conductors, although further refinement is probably needed before commercial exploitation can occur [22]. The most common area for development has been joining various aluminium alloys to steel, using lap and butt welds [23], and also friction stir spot welding [24, 25]. There are numerous accounts of success from many organisations, and indeed one car maker is already known to use spot welding to join aluminium to steel. The available literature gives good
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cause for expectation of improved weld quality, and more widespread usage of friction stir welding in this area. However, the process appears to require considerable attention to detail, and tolerances on welding parameters are expected to be tight, but this is seen as an area where growth will continue, and where a viable solution to a difficult problem will emerge. Another significant challenge which is being addressed is joining aluminium to titanium. This is of particular importance in current airframe designs where significant amounts of composite are used, as successful development will allow a reduction in the quantity of titanium required. Other materials combinations are likely to prove more even more challenging, for example joining aluminium to magnesium alloys. Success here is very limited, due to the formation of a low melting point eutectic, and the formation of very brittle intermetallics. There is also an issue of galvanic corrosion in any dissimilar joint, and so more traditional joining methods such as mechanical fastening or adhesive bonding are likely to dominate in the future. It is clear that a lot of further investment will be required to reduce this particular combination to commercial reality, and this is not likely to be achieved for some time.
6.3.6 Thermoplastics Thermoplastics (e.g., polypropylene, polyethylene, PEEK, nylon, etc.) can be joined easily by thermal processes, as they soften at relatively low temperatures. Thermosets, such as epoxies, char and decompose as they become hot, and therefore cannot be welded with the same ease and are not considered suitable for friction stir welding. Friction stir welding has been demonstrated for several thermoplastics [26], and encouraging results have been obtained. However, the process has never really captured the imagination of the thermoplastics industry, and it is believed that this is largely due to the economics of the process. As with steels, there are very well-established low cost processes which will not easily be displaced, and so for the foreseeable future friction stir welding will remain in the background.
6.4
Summary
It is hoped that the above thoughts give a reasonable summary of the state of the art for friction stir welding, and for its future. It is expected with reasonable certainty that the process will continue to grow, and the applications and tonnages welded will increase, but the extent of this is impossible to forecast. However, these thoughts are based only on past experience, and history has shown many examples where attempting to predict the future by extrapolation of the past is a dangerous activity. Nevertheless, the following broad conclusions are reached:
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1. Friction stir welding is here to stay, and it will continue to make inroads into light metal fabrication for joints appropriate to mechanised welding, at the expense of MIG and TIG and mechanical fastening. It is, however, inconceivable that these processes will be entirely replaced. Laser welding of aluminium is also making good progress, and will compete for much this market. 2. The technology is not yet fully matured, and it is expected that significant improvements in tool design, tool materials, process control, etc., will continue. It is believed that the full potential of friction stir welding is unknown, but will be dictated by imagination, and limited by investment. Significant development and growth of the process is inevitable. 3. The market for welding high temperature materials exists, but at the moment the technology is not yet ready. It is expected that progress here will be dependent on the development of significantly improved tool materials, especially for steels and nickel alloys. Without such tool materials, it is difficult to see a large market emerging, although specialist applications will no doubt appear. 4. A major barrier to the future progress of the friction stir welding is likely to be the lack of trained practitioners, although the same is probably true for any emerging technology.
6.5
Acknowledgements
The author is indebted to numerous colleagues in the friction stir welding community for discussions which shaped many of the ideas presented here.
6.6
References
1 W M Thomas, E D Nicholas, J C Needham, M G Murch, P Temple-Smith, and C J Dawes: “Improvements relating to friction welding”. Patent US 5,460,317 and EPS 0,616,490, 1991. 2 Smith I J and Lord D R R: “FSW patents – A stirring story”. Proc 7th International Symposium on Friction Stir Welding, Awaji Island, Japan, 20–22 May 2008. 3 Nunn M E: ”System elements and concepts for portable friction stir welding“. TWI members Report 877/2007, July 2007. 4 Thomas W M, Dolby R E and Johnson K I: “Variation on a theme – “Skew Stir” (friction stir welding) technology. Welding and Metal Fabrication, vol. 69, no. 8. Sept. 2001. 5 Thomas W M, Norris I M, Staines D G and Watts E R: “‘Com-Stir’- compound motion for friction stir welding and machining”. TWI Connect, no.124. May–June 2003. 6 Mahoney M, Mishra R S, Nelson T, Flintoff J, Islamgaliev R and Hovansky Y: “High strain rate, thick section superplasticity created via friction stir processing”. In: Friction Stir Welding and Processing. Proceedings, Symposium, Indianapolis, IN, 4–8 Nov. 2001.
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7 Grant G J: “Superplastic forming of Al multi-sheet”. Proceedings, 6th International Symposium on Friction Stir Welding, Saint-Sauveur, Quebec, Canada, 11–14 October 2006. 8 Kyffin W J, Threadgill P L, Lalvani H and Wynne B P: “Recent developments in FSSW of automotive steels” Proceedings of 6th International Symposium on Friction Stir Welding, Saint-Sauveur, Quebec, October 2006. 9 Mironov, R Ohashi, M Fujimoto, Y S Sato and H Kokawa: “Solid-state phase transformation of DP-590 steel during friction spot joining”. Proceedings, 7th International Symposium on Friction Stir Welding, Awaji Island, Japan, 20–22 May 2008 10 Skinner M, Edwards R L, Adams G and Li Z X: “Improvements to the FSW [friction stir welding] process using the self reacting technology”. Proceedings, 4th International Symposium on Friction Stir Welding, Park City, UT, USA, 14–16 May 2003. 11 Threadgill P L, Martin J P, Ahmed M M Z and Wynne B P: “The use of bobbin tools for friction stir processing of aluminium alloys”. invited paper at Thermec 09, Berlin, August 25–29 2009. 12 Teh N J: “Small joints make a big difference”. TWI Connect, no. 143, July–August 2006, p.1 13 Teh N J: “Stirring stuff”. The Engineer On-line, http://www.theengineer.co.uk/ Articles/296139/Stirring+stuff.htm, September 2006. 14 Nishihara T and Nagasaka Y: “Development of micro-FSW [friction stir welding]”. Proceedings, 5th International Symposium on Friction Stir Welding, Metz, France, 14–16 Sept. 2004. 15 Colegrove P A, Shercliff H R and Threadgill P L: “Modelling and development of the Trivex friction stir welding tool”. Proceedings, 4th International Symposium on Friction Stir Welding, Park City UT, May 2003. 16 Steel R, Liu Q, Yao X, Packer S and Leonhard T: “FSW tool material developments for joining high melting temperature materials”. Proceedings, 7th International Symposium on Friction Stir Welding, Awaji Island, Japan, 20–22 May 2008. 17 Fujii H, Kato H, Nakata K and Nogi K: “FSW of high temperature materials (Mo, Ti)”. Proceedings, 6th International Symposium on Friction Stir Welding, SaintSauveur, Quebec, Canada, 11–14 October 2006. 18 Ohashi R, Fujimoto M, Koga S, Ikeda R and Ono M: “Friction spot joining of steel sheets with silicon nitride tool”. Proceedings, 7th International Symposium on Friction Stir Welding, Awaji Island, Japan, 20–22 May 2008. 19 Bernath J J, Krem S and Li T: “FSW of Ti-6Al-4V structural Components”. Proceedings, 6th International Symposium on Friction Stir Welding, Saint-Sauveur, Quebec, Canada, 11–14 October 2006. 20 Russell M J and Blignault C: “Recent developments in FSW of Ti Alloys”. Proceedings, 6th International Symposium on Friction Stir Welding, Saint-Sauveur, Quebec, Canada, 11–14 October 2006. 21 Christner B, McCoury J and Higgins S: “Development and testing of friction stir welding (FSW) as a joining method for primary aircraft structure”. Proceedings, 4th International Symposium on Friction Stir Welding, Park City, UT, USA, 14–16 May 2003. 22 Savolainen K, Mononen J, Saukkonen T and Hänninen H: “A preliminary study on FSW of dissimilar metal joints of Cu and Al”. Proceedings, 6th International Symposium on Friction Stir Welding, Saint-Sauveur, Quebec, Canada, 11–14 October 2006.
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23 A Kostka, R Coelho, J dos Santos, and A R Pyzalla: “Microstructure and properties of Al to steel friction stir overlap welds”. Proceedings, 7th International Symposium on Friction Stir Welding, Awaji Island, Japan, 20–22 May 2008. 24 Mazzaferro J A E, de Souza Rosendo T, Mazzaferro C C P, Ramos F D, Tier M A D, Strohaecker T R and dos Santos J F: “Preliminary Investigation on the mechanical behaviour of aluminium friction spot welds”. Proceedings, 7th International Symposium on Friction Stir Welding, Awaji Island, Japan, 20–22 May 2008. 25 Miyagawa K, Matsumura H, Yasui T, Tsubaki M and Fukumoto M: “The microstructure of interface in spot welding between aluminium alloy and high tensile strength steel by friction stirring”. Proceedings, 7th International Symposium on Friction Stir Welding, Awaji Island, Japan, 20–22 May 2008 26 Nelson T W, Sorensen C D, Johns C, Strand S and Christensen J: “Joining of thermoplastics with friction stir welding”. Proceedings, 2nd International Symposium on Friction Stir Welding, Gothenburg, Sweden, 26–28 June 2000. Session 5, Paper 3.
7
Inspection and quality control in friction stir welding T. Zappia, MTS Systems Corporation, USA
Abstract: The quality of the weld is typically one of the main motivations for choosing FSW as a manufacturing process. Other reasons for using FSW may include the associated end-use efficiencies gained by FSW along with manufacturing costs savings, but what is always of primary importance is that the quality of the weld satisfies the application needs and that the process will be robust and repeatable in production. In Chapter 9, the various types of weld defects are shown. This chapter discusses the typical steps for developing FSW for a given application and some approaches for establishing weld quality requirements, including techniques that can be used to measure results both during welding and post welding. Key words: weld quality requirements, process development and qualification, design-of-experiment, on-line monitoring, statistical process control, off-line monitoring, non-destructive testing, visual inspection, radiography (x-ray), dye-penetrant, phased array eddy current, ultrasonic inspection
7.1
Weld quality requirements definition
The typical process that is followed for implementing FSW is to use a phased approach similar to that depicted in Fig. 7.1. In Phase 1, an assessment is done to verify the feasibility of using FSW for a given application. Welds are made for the desired joint type, material, and part thickness to see if the results are acceptable for the application. These welds are typically done on small coupon samples with simplified clamping techniques used to hold the part in place. The purpose of this step is to gain confidence that the FSW process has enough potential to manufacture the application parts and proceed further with development. If the Phase I assessment is successful, the development moves to Phase II Process development and qualification where more detailed work is done to define and optimize the FSW process. Again, depending upon the application, the amount of work required depends upon the application. On the simpler end, Process development might be limited to verify that a selected pin tool design, weld schedule, and part tooling achieve acceptable weld penetration and material mixture as measured through, for example, 183
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Phase II – Process development/ qualification
Design feasibility study
Weld schedule development
Weld testing and process certification Phase III – Implementation for production
Production equipment selection
7.1 FSW phased implementation approach (courtesy of MTS Systems Corporation).
micrograph cross-section analysis and bend strength testing. On more complicated, high performance, and safety conscious applications such as flight structures for aerospace applications, Phase II will likely require very involved design-of-experiment (DOE) testing to define the process window for the weld parameters associated with the pre-weld part setup, pin tool, and part fixturing. Process development and testing become iterative as the pin tool and weld schedules are refined in order to ensure that the weld quality meets the demanding design requirements in a highly robust way. Tests are generally run on small coupons to establish baseline performance, and then on small- and full-scale test pieces. Depending upon the application, weld quality measurements might include micrograph cross-section analysis, bend tests, and tensile, fatigue, crack growth, fracture toughness, and corrosion testing. Once enough confidence has been established with the FSW process for a given application, and the other production readiness requirements are met (i.e., analysis of various parameters that ensure that the return-on-investment is acceptable), Phase III is entered into and the required FSW equipment is specified for production. To understand if weld quality is being achieved it is first necessary to establish a set of quality requirements by which to measure the welds. These quality requirements depend upon the application and generally need to take into consideration the forces and environment that the part is subject to in its end-use. Examples of acceptable criteria may be in terms of tensile, yield and fatigue strength which can be measured by testing techniques such as bending, tensile, torsion, and fatigue cycling.
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Generally, the forces that a part will be subject to are known by the part designer and are taken into consideration during the part design process. As an example, for aircraft parts manufacturing the design engineers design the part for an expected fatigue life at some predefined stress (e.g., 3 life times at 100 Mpa; R value = 0.05). Often times, however, FSW will be investigated to replace an existing joining technology where the weld quality requirement is not exactly known. It may be that part manufacturing has always been done a particular way and has worked robustly and now the investigators want to replace it with FSW for cost-saving reasons. In this case, it may become necessary to run tests on the existing process to establish the baseline quality parameters by which FSW must be able to at least replicate. In either of the examples above, it is the quality criteria that will be the measuring stick by which to evaluate the FSW process. Meeting or surpassing this quality requirement is the main goal of the Phase II Process development and qualification step. A result of this step is the establishment of weld schedules that can be reliably used to manufacture the parts. The execution of these weld schedules, along with the part material quality, pin tool, part setup and hold down fixtures, ensure the repeatable quality of the weld. As already stated, the rigor by which Phase II Process development and qualification needs to be undertaken depends on the application and overall quality and safety requirements. The various parameters that generally affect weld quality and are of concern during the Process development and qualification phase are shown in Fig. 7.2. These include: ∑ Part material consistency – The material used for process development should be consistent with the end production material. This includes the application of any material coatings. Differences from one lot production Process parameter set Rotation RPM*
Travel speed*
Down force*
Weld path* *Online controlled by machine
Pin tool
Tolerances
Weld quality
Gaps/mismatch
Thickness
Cleanliness Environment
Coatings
Material
7.2 FSW quality-related parameters (courtesy of D. Lohwasser – Airbus).
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of material to another should be understood, especially with thin gauge materials, as variations can affect the weld parameter control window. ∑ Part gap and mismatch – Part tolerance variations are guaranteed with any manufacturing process. The more one pays, the tighter the tolerances can be made. Welding should be done to characterize the impact of tolerance variation to make sure that the desired quality is achievable for the expected tolerance range. ∑ Sealants – In the case where sealants need to be applied (e.g., lap welds for aircraft structure), tests should be run to verify that the sealant can be welded through, or applied post weld. Often corrosion testing techniques such as 30-, 60-, and 90-day exposure tests are used to measure corrosion protection. During process development care should be taken to test for the variations in sealant application that might affect the weld characteristics. ∑ Pin tool – The FSW pin tool design is a key piece of equipment to the FSW process. As discussed earlier, the design of the pin tool is done to maximize the mixing of the joint material for the specific configuration. Pin tool wear needs to be taken into consideration as changes to the pin tool features will impact the quality of the weld for a given set of weld parameters. ∑ Part Tooling – Proper part fit-up and hold-down fixturing is essential to maintaining consistent quality welds. During process development, tests can be run to establish the range for the clamping mechanism (e.g., hold-down pressure, part location). The outcome of the Process development and qualification phase are the weld procedures that are to be followed to achieve repeatable welds. If done properly, this phase will ensure that quality results are built into the weld process, thereby statistically ensuring that the final product welds meet the desired quality requirement parameters. The DOE methodology is considered the most thorough technique for establishing the weld procedures and is used here as a model for establishing the performance window (i.e., the requirements that the production process, including the FSW equipment, must adhere to). There are a number of commercial DOE products available to help organize and evaluate the DOE process. As stated in the NIST Engineering Statistics Handbook, DOE are an efficient procedure for planning experiments so that the data obtained can be analyzed to yield valid and objective conclusions. DOE begins with determining the objectives of an experiment and selecting the process factors for the study. A DOE is the laying out of a detailed experimental plan in advance of doing the experiment. Well-chosen experimental designs maximize the amount of “information” that can be obtained for a given amount of experimental effort.1 The schematic shown in Fig. 7.3 is an example of what the FSW DOE framework process box might look like.
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Discrete factors
Bend strength
Weld path
Tensile strength
Down force
Fatigue strength
Process
Environment
Material preparation
Other Material tolerances
Tool angle Part clamping
Micro structure
Pin tool design
Travel speed
Outputs
Control factors
Operators
Material lot
Machines
Rotation speed
Continuous factors
7.3 DOE FSW process schematic.
The objectives of the DOE generally fall into one of the following categories: ∑ Choosing between alternatives ∑ Selecting the key factors affecting a response ∑ Response surface modeling to: – hit a target – reduce variability – maximize or minimize a response – make a process robust (i.e., the process gets the “right” results even though there are uncontrollable “noise” factors) – seek multiple goals ∑ regression modeling. In the end, the result of the DOE is to have a set of welding procedures that will provide confidence that the resulting weld will have a statistically significant probability of achieving the defined weld quality requirements as long as the weld can be executed within the performance window defined by the DOE. An example is shown in Fig. 7.4 where the process window for forge force is given for one of the weld schedules developed through DOEs by Eclipse Aviation Corporation for their Eclipse 500 jet. To summarize, defining the weld requirements is extremely important to achieving success with FSW. It provides the reference points by which to
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Process capability (Cpk) Values for stringer to skin welds
Upper specification limit
FSW parameter Forge load Travel speed Spindle speed Pin tool lag angle
Cpk 5.0 16.0 11.0 110.0
Lower specification limit
7.4 Eclipse Aviation process window for forge force (courtesy of Eclipse Aviation Corporation).
measure success and creates the boundary for the type and amount of work that is done during the Process development and qualification phase. The end-use of the welded part will dictate the appropriate requirements. There are many different approaches that can be used for doing process development and qualification work. DOEs are typically the most rigorous approach and provide a very disciplined framework by which to manage process development work. The statistical nature of DOEs can provide a high level of confidence that the resulting weld procedures, if followed properly, will meet the part design requirements. DOEs can also have an added advantage of greatly reducing the number of welds that need to be performed during process development and qualification due to the statistical correlations that can be made between parameters that might otherwise not be understood.
7.2
Online monitoring and statistical process control
Once the important work of establishing the process requirements has been achieved, attention can then shift to the also important work of production and executing the welds in a repeatable fashion so as to match the results that were determined during the Process development and qualification phase. As mentioned in Chapter 4, the FSW equipment needs to be capable of controlling and monitoring the process with enough accuracy and repeatability to meet the performance requirements that define the application. Statistical process control techniques can be used to monitor the process and take appropriate
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corrective action such as alerting the operator, stopping the weld, or adjusting control parameters. The online monitoring of the process can include real-time sensing of the FSW process parameters, weld temperature, weld path errors, and visual monitoring of the weld (e.g., excess flash). It can also include more recently developed analytical techniques that find correlations between the monitored data to detect weld attributes that can normally not be sensed in real-time. A description of these techniques is provided below.
7.2.1 FSW process parameters The weld schedules that were produced during the Process development and qualification phase define the process values along the weld path (e.g., FSW pin tool force, tilt angle, RPM, travel speed). The control system generates the appropriate corresponding commands for the various machine axes and monitors the system transducer feedbacks to make sure that the channel of control errors (i.e., command – feedback) are within the allowable range. These commands relate to machine axes positions, speeds, forces, accelerations, etc., and are monitored by transducers such as encoders, linear variable displacement transducers, load, pressure, and torque cell, etc. When errors are detected that fall outside of the defined process window, an appropriate response can be taken such as alerting the operator or stopping the weld (Fig. 7.5). If the FSW Process development and qualification phase has been done properly, and the FSW machine has the appropriate control and monitoring capabilities, the online monitoring of the FSW process parameters can be one of the best ways of ensuring high quality results. In essence, this process is closing the loop on the quality of the weld that was defined and built into the weld as part of the Process development and qualification phase. This type of monitoring is what has been successfully done for production
Control parameters
Control window exceeded – Alert operator – Stop – Other Average value
Time
7.5 Process error band.
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applications such as the Boeing Delta II and IV rockets, Eclipse Aviation’s EAC-500 jet, and Nippon Sharyo’s Shinkansen high-speed train.
7.2.2 Weld temperature Weld temperature can be an indication of the amount of energy that is being imparted into the part and can be used as a signal to monitor weld quality. As stated earlier in Chapter 3, the heat flow through the part directly correlates to the final mechanical properties of the welded part. Precisely measuring the weld temperature is challenging, as it can be difficult to place probes directly at the interface of the part and FSW pin tool. Different sensing techniques include placing a temperature probe inside the FSW pin tool (see Fig. 7.6), infrared radiation thermometry (e.g., pyrometer, IR camera (see Fig. 7.7), and thermocouple placement along the anvil backing plate (see Fig. 7.8). Each technique, along with thermal models, have their advantages and disadvantages, but all can be used to monitor the heat being imparted to the part. The most accurate measurement is thought to be the temperature probe inside, or very near, the pin tool. The other techniques are more complicated to correlate due to temperature flow and unknown and changing emissivity values. Owing to the multivariate interactions of FSW (i.e., load, pin tool,
Thermocouple Shroud for shielding gas Shielding gas inlet
7.6 Temperature probe inside FSW pin tool (courtesy of MegaStir (PCBN pin tool)).
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(a)
(b)
7.7 Infrared radiation thermometry sensing (courtesy of D. Lohwasser – Airbus).
travel speed, material, joint, etc.) and the difficulty in measurement, quality correlations to temperature are not being widely used as a control parameter or definitive measurement of weld quality. However, this remains an area of active research.
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AS RS
Temp. [°C]
200 150 100 50 0
0
20
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Backing bar Thermocouples RS
7.8 Thermocouple placement in backing plate (courtesy of D. Lohwasser – Airbus).
7.2.3 Weld path errors The ability of the FSW machine to maintain the correct position of the FSW pin tool along the weld path is critical to weld quality. During the Process development and qualification phase tests are generally run to see what the allowable offset from the center of the weld seam should be. However, due to machine and tooling deflections, part location offsets, and machine inaccuracies it is possible that there will be pin tool cross-seam positioning errors. These errors can be corrected manually by allowing an operator to “jog” the tool Y axis during the weld and/or by use of a seam-tracking system. Vision and seam-tracking systems can be integrated onto the FSW equipment and be used to monitor and help make appropriate corrections (see Fig. 7.9).
7.2.4 Visual monitoring Using a vision system image similar to that shown in Fig. 7.9, the operator can closely monitor the visual quality of the weld and make appropriate adjustments. For example, excessive flash normally indicates that the FSW pin tool shoulder is too far into the material. It can be corrected by adjusting the tool Z axis, reducing the load, or decreasing pin tool rotation speed (i.e., reducing heat in the part). It should be noted that this kind of manual adjustments to the weld process need to be performed in a very controlled
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7.9 Vision and seam-tracking equipment (courtesy of MTS Systems Corporation).
manner so as not to inadvertently take the process parameters outside of the defined process window.
7.2.5 Analytical sensing A good amount of work has recently been performed in the area of analyzing process data to find correlations to weld defects that have previously been able to be detected only through post-weld non-destructive testing. At the South Dakota School of Mines and Technology, work has been done through the use of neural networks to show the relationship between the oscillation of X and Y axes feedback forces, plasticized material flow, and the detection of wormhole defects.2 The results are shown in Fig. 7.10. Further research showed the correlation between the Y axis tool force (i.e., force normal to weld travel direction) and its derivative Y¢ to the quality of welds using phase space orbital analysis. The orbits in phase space of welds without metallurgical defects tend to remain within a neighborhood surrounding the ideal trajectory while welds containing defects produced more variable orbits.3 An example of the results is shown in Fig. 7.11 and Fig. 7.12. As more information becomes known and the database repository of cause and effect between control parameters and resulting quality are verified, the next step will be to integrate into the FSW control system the ability to make
0.4 0 0.3 2.3 4.3 6.3 8.3 10.3 12.3 14.3 Frequency (Hz) (a)
Spindle frequency bin
0.8 0.4 0 0.3 2.3 4.3 6.3 8.3 10.3 12.3 14.3 Frequency (Hz) (d)
Scaled amplitude
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0.8 0.4 0 0.3 2.3 4.3 6.3 8.3 10.3 12.3 14.3 Frequency (Hz) (b)
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1 mm (c)
Spindle frequency bin
0.8 0.4 0 0.3 2.3 4.3 6.3 8.3 10.3 12.3 14.3 Frequency (Hz) (e)
1 mm (f)
7.10 Representative examples of the manually identified frequency patterns of X force and Y force. The frequency pattern of X force in (a) and the frequency pattern of Y force in (b) correspond to a good weld, whose cross-sectional image is depicted in (c). Likewise, the frequency pattern of X force in (d) and the frequency pattern of Y force in (e) correspond to a bad weld, whose cross-sectional image with a wormhole defect is depicted in (f). Scaled amplitude is the amplitudes divided by the amplitude of the spindle frequency peak. (Courtesy of South Dakota School of Mines and Technology.)
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Scaled amplitude
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195 100 80 60
Delta y (lbf)
40 20 –400
–350
–300
–250
–200
–150
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7.11 Phase space trajectory for the y force vs. the derivative of the y force measured in foot-pounds (lbf). The phase space plot at the top is for a good weld, the plot at the bottom is for a poor quality weld. The trajectory of the spindle frequency (dark circle) is also plotted for reference. (Courtesy of South Dakota School of Mines and Technology.)
real-time adjustments to the process in order to maintain part quality during the welding. It is also important to note that whenever possible it is best to record all the monitoring data so that it can be archived along with any other pertinent information about the part. This has the obvious benefit of providing verification and back-testing from any future problems that might occur from a welded part that are uncovered during its intended use.
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7.12 The location of the resulting wormhole at 40¥ magnification for the bad weld in Fig. 7.9. (Courtesy of South Dakota School of Mines and Technology.)
7.3
Offline testing: non-destructive testing
No matter how much care and effort has gone into the generation of wellthought out and encompassing DOE and the definition of robust and reliable weld schedules during the Process development and qualification phase, it is still reassuring, and sometimes absolutely required, to perform nondestructive testing (NDT) on the welded part. As explained in Chapter 9, even though FSW does not have to contend with many of the error sources that come from fusion welding, there are still errors that can make their way into an FSW weld. Figs 7.13 and 7.14 show examples of the common types of FSW defects. The NDT methods used to inspect FSW welds are fundamentally the same as those used for other types of welds. These include visual inspection, offline monitoring of data, radiography, dye penetrant, ultrasonic, and eddy current. This section will provide a summary of these techniques as they apply to FSW. Many of these techniques were explored by Kinchen and Aldahir in their article “NDE of Friction Stir Welds in Aerospace Applications – The right NDE process for inspecting friction stir welds”4 and are referenced here.
7.3.1 Visual inspection Visual inspection is generally the first and simplest type of inspection. Many surface type errors such as too much flash, galling, and lack of penetration can readily be seen by simply looking at the finished weld. Some examples of visual defects are shown in Fig. 7.15.
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Voids
Flash formaton
Voids
Oxides alignment
Oxides alignment
ZATM
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ZAC
ZAC
10x 2 mm
Thickness reduction Oxides alignment
Kissing-bond (a)
Shoulder
Retreating side
Advancing side 4 Nugget 1
2
3
Types of imperfections: (1) (2) (3) (4)
Lack of penetration (kissing-bonds); Root flaw (weak or intermittent linking); Voids on the advancing side; Second phase particles and oxides alignment under shoulder. (b)
7.13 Typical types of FSW defects (courtesy of Technical University of Lisbon, IST, Secção de Tecnologia Mecânica, Av. Rovisco Pais, 1049-001 Lisbon, Portugal Developments in NDT for Detecting Imperfections in Friction Stir Welds in Aluminium Alloys Telmo Santos, Pedro Vilaça*, Luísa Quintino Technical University of Lisbon, IST, Secção de Tecnologia Mecânica, Av. Rovisco Pais, 1049-001 Lisbon, Portugal * Corresponding author:
[email protected]; Welding in the World, Vol. 52. No. 9/10, IIW-1866-07 (ex-doc.III1426r1-07).)
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(a)
(b)
7.14 Wormhole (tunnel) and oxide layers (kissing bonds) FSW defects: (a) wormholes are long defects that can be meters long and are located in the volume of the weld, generally caused by low rotation speed of the FSW pin tool; (b) oxide layers are located at the weld root and are due to insufficient tool penetration resulting in mechanical contact without the chemical mix of material components. (Courtesy of Olympus10).
7.15 Visual FSW defects showing overheating at startup of Al 7xxx; root side void with lifting; root side void of shelf weld and shoulder void.
Figure 7.16 shows an example of lack-of-penetration defects that become more apparent after etching has been performed. As Kinchen and Aldahir state, “The principal unacceptable root side condition is IJP (insufficient joint penetration). Early on in the friction stir welding program, IJP was considered to be the most critical type of
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DX
LOP
7.16 Lack of penetration defects.
defect. As a result, most NDE testing was conducted with this discontinuity type. Visual examination of the root side of the weld demonstrated IJP discontinuities were detectable when inspected in the post-etched condition. Etching is a post-weld chemical treatment performed most often to prepare mechanically worked surfaces prior to penetrant inspection. In this case, the etching process clearly delineates the weld nugget dynamically recrystallized zone (DXZ) and its surrounding heat-affected zone (HAZ), making the lack of FSW nugget a distinct feature visible to the trained eye. The reason for the successful detection rate is that it is easy to discern the DXZ from the surrounding base metal and HAZ in the post-etched condition. Therefore, visual inspection is a reliable technique to confirm suspected IJP conditions. Fig. 7.16 is a 3¥ magnification view of an IJP defect on the root side of a FSW panel after etching. The metallurgical characteristics of an IJP discontinuity are the determining aspects of the discontinuity and relate directly to the ability of ultrasonics and penetrant inspection techniques to detect IJP. These characteristics are, likewise, directly linked to the weld process itself. Primary factors affecting IJP during welding include heat input or material flow and, most importantly, the depth of the FSW pin tool.”4
7.3.2 Offline monitoring of data The offline monitoring of data is very similar to the online monitoring of data, and both can be done for certain applications (see Table 7.1). The main difference is that with offline monitoring of data alone there will not be any operator or weld engineer intervention should the recorded data fall outside of the allowable process window during a weld. Instead, the data is checked post weld and any anomalies that are found are addressed for
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Table 7.1 Offline monitoring of data Monitoring/inspection Application type Offline monitoring ∑ of data
Advantages
Disadvantages
Correlation to ∑ Simple and ∑ Requires well-defined flaws detected quickly process window that during Process identifies accurately determines development problems range of process and qualification parameters
Radiation
Void Specimen Film
Black area
White area
White area
Black area
After processing
Gray area
7.17 Radiography (x-ray) (courtesy of Engineers Edge – Solutions by Design).
the particular part that corresponds to the data set. This has the effect of keeping the production throughput moving at a known level and any repairs or rework can be done at a later time.
7.3.3 Radiography (x-ray) Radiography, or x-ray, is another common way to inspect welds. It uses the concept of differential absorption of penetrating radiation. Specimens will have differences in density, thickness, shapes, sizes, or absorption characteristics. Unabsorbed radiation that passes through a part is recorded on film, fluorescent screens, or other radiation monitors. Indications of internal and external conditions will appear as variants of black/white/gray contrasts on exposed film, or variants of color on fluorescent screens5 (see Fig. 7.17). Radiography is used widely in the examination of castings and weldments, particularly where there is a critical need to ensure freedom from internal flaws.6 Table 7.2 summarizes the uses, advantages, and disadvantages with radiography. Figure 7.18 shows examples of weld tunnels being detected by x-ray.
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Table 7.2 Radiography (x-ray) Monitoring/ inspection type
Application
Advantages
Disadvantages
Radiography
∑∑ ∑∑ ∑∑ ∑∑ ∑∑ ∑∑ ∑∑ ∑∑
∑∑ Sensitive to finding discontinuities throughout the volume of materials ∑∑ Easily understood permanent record ∑∑ Full volumetric examination ∑∑ Portability
∑∑ Radiation hazard ∑∑ Relatively expensive ∑∑ Long set-up time ∑∑ Necessary access to both sides of specimen ∑∑ Depth of indication not shown ∑∑ High degree of skill required for technique and interpretation ∑∑ Lack of sensitivity to fine cracks and lack of penetration
Inclusions Cracks Porosity Corrosion Debris Lack of fusion Lack of penetration Leak paths
(a)
(b)
7.18 X-Ray example showing large (a) and small tunnels of butt weld (courtesy of D. Lohwasser – Airbus).
7.3.4 Dye penetrant Definition (taken from Wikipedia7): Dye penetrant inspection (DPI), also called liquid penetrant inspection (LPI), is a widely applied and low-cost inspection method used to locate surface-breaking defects in all non-porous
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materials (metals, plastics, or ceramics). LPI is used to detect casting and forging defects, cracks, and leaks in new products, and fatigue cracks on inservice components. DPI is based upon capillary action, where low surface tension fluid penetrates into clean and dry surface-breaking discontinuities. Penetrant may be applied to the test component by dipping, spraying, or brushing. After adequate penetration time has been allowed, the excess penetrant is removed, a developer is applied. The developer helps to draw penetrant out of the flaw where a visible indication becomes visible to the inspector. Inspection is performed under ultraviolet or white light, depending upon the type of dye used – fluorescent or nonfluorescent (visible).7 Table 7.3 summarizes the uses, advantages, and disadvantages with dye penetrants. In tests performed at Lockheed Martin for aerospace applications, it was determined that dye penetrant was found to be unacceptable in the as-welded condition due to poor detection and the excessive background noise produced by the surface, which interfered with the inspection. Single and double etching was required and it was found that a caustic etchant should be used prior to penetrant application so as to remove 0.0004 to 0.0006 of an inch of material in order to improve the ability to detect lack of penetration (see Fig. 7.19).4 The findings from Kinchen and Aldahir were “Inspection of FSW in the etched condition via P135E and P6F4 consistently and successfully detected root-side LOP discontinuities. However, because the sensitivity level of detection for each penetrant solution is different, the results were dissimilar. P135E successfully detected LOP discontinuities that were greater than or equal to 0.064 in. deep, and P6F4 successfully detected LOP discontinuities that were greater than or equal to 0.050 in. deep. Double etching, via caustic etchant solution prior to the application of penetrant, enhanced the detection of LOP in comparison to single etching.”4
Table 7.3 Dye penetrant Monitoring/ inspection type
Application
Advantages
Disadvantages
Dye penetrant
∑∑ ∑∑ ∑∑ ∑∑ ∑∑
∑∑ Inexpensive ∑∑ Sensitive ∑∑ Minimal equipment ∑∑ Application to irregular shapes ∑∑ Versatile ∑∑ Minimal training
∑∑ Non-porous surfaces only ∑∑ Detection of surface flaws only ∑∑ Messy ∑∑ Ventilation requirements
Cracks Porosity Leak paths Seams Laps
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Lack of penetration 0.064”
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Plastic deformation Weak bonding
7.19 Metallurgical cross section of LOP discontinuity (courtesy of Lockheed Martin and American Welding Society).
7.3.5 Ultrasonic inspection Definition (taken from NDT Resource Center8): Ultrasonic testing (UT) uses high frequency sound energy to conduct examinations and make measurements. Ultrasonic inspection can be used for flaw detection/evaluation, dimensional measurements, material characterization, and more. The general inspection principle, a typical pulse/echo inspection configuration is shown in Fig. 7.20. A typical UT inspection system consists of several functional units, such as the pulser/receiver, transducer, and display devices. A pulser/receiver is an electronic device that can produce high voltage electrical pulses. Driven by the pulser, the transducer generates high frequency ultrasonic energy. The sound energy is introduced and propagates through the materials in the form of waves. When there is a discontinuity (such as a crack) in the wave path, part of the energy will be reflected back from the flaw surface. The reflected wave signal is transformed into an electrical signal by the transducer and
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Transducer Initial pulse Back surface echo Crack echo Crack Plate 0
2
4
6
8
10
12
7.20 Ultrasonic general principle.
is displayed on a screen. In the applet below, the reflected signal strength is displayed versus the time from signal generation to when a echo was received. Signal travel time can be directly related to the distance that the signal traveled. From the signal, information about the reflector location, size, orientation and other features can sometimes be gained. Multiple transducers can be incorporated into a single “phased array” inspection probe and multiplexed to increase the speed, coverage and sensitivity of the inspection. Ultrasonic inspection is a very useful and versatile NDT method. Table 7.4 shows a summary of the uses, advantages, and disadvantages of ultrasonic testing. Ultrasonic NDT and FSW: There are two basic types of ultrasonic scans, linear and sector, that can be used for FSW. A linear scan can provide a full coverage of the weld bead with a fast scan of up to 20 KHz which can cover the weld in a one-line pass. A sector scan can create different inspection angles with the same probe and allow inspection of complex shape parts, full volume coverage can also be accomplished with focused beams (see Fig. 7.22).9
7.3.6 Phased array eddy current Definition (taken from NDT Resource Center10): Eddy current inspection is one of several NDT methods that use the principle of “electromagnetism” as the basis for conducting examinations. Eddy currents are created through a process called electromagnetic induction (see Fig. 7.23). When alternating current is applied to the conductor, such as copper wire, a magnetic field
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7.21 Ultrasonic detection of FSW lack of penetration (80 mm long) (courtesy of Olympus).
Table 7.4 Ultrasonic testing Monitoring/ inspection type Ultrasonics
Application
∑∑ Detect lack of penetration (see Fig. 7.21) ∑∑ Detect wormholes ∑∑ Discontinuities in surface and subsurface ∑∑ Thickness measurements
Advantages
∑∑ Fast ∑∑ Only singlesided access is required ∑∑ Full volumetric information ∑∑ Minimal part preparation is required ∑∑ Instantaneous results ∑∑ Detailed images can be produced automatically ∑∑ Permanent record ∑∑ Can be used for thickness measurements
Disadvantages
∑∑ Surface must be accessible and smooth ∑∑ Can have operator dependence ∑∑ Flaw orientation important – linear defects oriented parallel to the sound beam may go undetected ∑∑ Interpretation can be difficult ∑∑ Need for reference standards ∑∑ Difficulty with complex geometries ∑∑ Inability to pass through air – need for couplant
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(a)
(b)
7.22 Ultrasonic testing of weld (a – Linear scan ; b – Sector scan) (courtesy of Olympus).
7.23 Eddy current principle (courtesy of NDT Resource Center).
develops in and around the conductor. This magnetic field expands as the alternating current rises to maximum and collapses as the current is reduced to zero. If another electrical conductor is brought into close proximity to this changing magnetic field, current will be induced in this second conductor. Eddy currents are induced electrical currents that flow in a circular path. They get their name from “eddies” that are formed when a liquid or gas flows in a circular path around obstacles when conditions are right. In the presence of a flaw, the flow of the eddy currents is disturbed, creating a perturbation in the magnetic field at the surface of the part that is detected by the eddy current probe. The frequency of the alternating current used to induce the eddy currents and the electrical conductivity of the material being inspected determines the depth of penetration of the eddy current field and the resulting depth of the inspection. Reducing the frequency of the alternating current increases the depth of penetration. However, even at low frequencies, the depth of penetration of the eddy current field is limited, so eddy current is a surface and near-surface inspection method. It can be a full volumetric inspection in thinner materials. Multiple probes can be integrated into a single inspection head to increase the coverage, sensitivity and speed of the inspection. This use of multiple probes is termed phased array eddy current. In addition to increasing the
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surface covered by using multiple probes, the frequency of the alternating current can be varied to optimize the sensitivy for both surface and subsurface defects. Table 7.5 shows the uses, advantages and disadvantages of phased array eddy current inspections. Eddy current NDT and FSW: Most eddy current probes used for FSW are configured in an array made up of many coils that allow higher sensitivity. Each scan coils a specific area and the coils are optimized for a specific application. The probe is designed so that there is no cross-talk interference between coils and nearly any shape configuration is possible. An example of a phased array configuration is shown in Fig. 7.24.
7.4
Summary
In summary, there are many different techniques that can be used to monitor and inspect weld quality. The particular technique that is selected is driven by the requirements associated with the application. A summary of the various techniques, both online and offline, is given in Table 7.6. Table 7.5 Phased array eddy current Monitoring/ inspection type
Application
Advantages
Disadvantages
Phased array eddy current
∑∑ Cracks, inclusions, dents, and holes ∑∑ Detect lack of penetration (see Fig. 7.25) ∑∑ Detects galling (see Fig. 7.26) ∑∑ Allow bead width sizing (indirect detection of oxide layers – see Fig. 7.27) ∑∑ Coating and material thickness ∑∑ Surface and near surface defects ∑∑ Composition/ conductivity/ permeability ∑∑ Grain size/hardness ∑∑ Dimensions and geometry ∑∑ Alloy sorting
∑∑ Fast ∑∑ Inspection done in one pass ∑∑ Full coverage of weld ∑∑ C-scan imaging for easy interpretation ∑∑ Easy to operate ∑∑ Automation available ∑∑ Permanent record available ∑∑ Specimen contact not necessary
∑∑ Manual surface testing is slow ∑∑ Interpretation may be difficult ∑∑ Depth of penetration is limited ∑∑ Flaw orientation is critical ∑∑ Specimen must be electrically conductive ∑∑ Sensitive to many specimen parameters ∑∑ Surface roughness can produce non-relevant indications
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ECA probe over a flaw
Each coil produces a signal
The amplitude of the signal is color-coded into a C-scan view
7.24 Eddy current phased array configuration (courtesy of Olympus).
7.25 Phased array eddy current detection of FSW lack of penetration (80 mm long) (courtesy of Olympus).
In this chapter an overview was provided on what constitutes FSW weld inspection and quality control. The first step of quality control is in the establishment of the weld quality requirements. These requirements need to be defined so that they can be used as a reference point from which to inspect and measure weld quality. The key phase of implementing FSW which begins to establish the correlation between weld parameters and weld quality
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7.26 Phased array eddy current detection of galling (courtesy of Olympus).
7.27 Phased array eddy current detection of oxydizing layer (kissing bond); (courtesy of Olympus).
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Table 7.6 Online and offline inspection techniques summary Monitoring/ inspection type
Application
Advantages
Disadvantages
Online monitoring (e.g., process parameters, temperature, weld path errors)
∑∑ Correlation to flaws detected during Process development and qualification
∑∑ Simple and quickly identifies problems while a part is being made
∑∑ Requires welldefined process window that accurately determines range of process parameters
Visual monitoring
∑∑ Flash ∑∑ Galling ∑∑ Lack of shoulder or pin penetration ∑∑ Cracking ∑∑ Misalignment ∑∑ Distortion
∑∑ Quickly identifies problems while a part is being made ∑∑ Minimal surface preparation ∑∑ Reduced need for other NDE methods
∑∑ Only able to detect visible surface flaws ∑∑ Observations vary with personnel experience ∑∑ Surface cleaning and preparation ∑∑ Distractions ∑∑ Poor resolution ∑∑ Eye fatigue ∑∑ Good illumination required
Analytical sensing
∑∑ Detection of flaws correlated to signal analysis study
∑∑ Capable of predicting weld flaws while part is being made ∑∑ Reduced need for other NDE methods
∑∑ In early phase of research and much work still needs to be performed.
Offline monitoring of data
∑∑ Correlation to flaws detected during Process development and qualification
∑∑ Simple and quickly identifies problems
∑∑ Requires welldefined process window that accurately determines range of process parameters
Radiography
∑∑ ∑∑ ∑∑ ∑∑ ∑∑ ∑∑ ∑∑
∑∑ Sensitive to finding discontinuities throughout the volume of materials ∑∑ Easily understood permanent record ∑∑ Full volumetric examination ∑∑ Portability
∑∑ ∑∑ ∑∑ ∑∑
Inclusions Cracks Porosity Corrosion Debris Lack of fusion Lack of penetration ∑∑ Leak paths
Radiation hazard Relatively expensive Long set-up time Necessary access to both sides of specimen ∑∑ Depth of indication not shown ∑∑ High degree of skill required for technique and interpretation ∑∑ Lack of sensitivity to fine cracks.
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Table 7.6 Online and offline inspection techniques summary Monitoring/ Inspection Type
Application
Advantages
Disadvantages
Dye penetrant
∑∑ ∑∑ ∑∑ ∑∑ ∑∑
Cracks Porosity Leak paths Seams Laps
∑∑ Inexpensive ∑∑ Sensitive ∑∑ Minimal equipment ∑∑ Application to irregular shapes ∑∑ Versatile ∑∑ Minimal training
∑∑ Non-porous surfaces only ∑∑ Detection of surface flaws only ∑∑ Messy ∑∑ Ventilation requirements
Ultrasonics
∑∑ Detect lack of penetration ∑∑ Detect wormholes ∑∑ Discontinuities in surface and subsurface ∑∑ Thickness measurements
∑∑ Fast ∑∑ Only singlesided access is required ∑∑ Full volumetric information ∑∑ Minimal part preparation is required ∑∑ Instantaneous results ∑∑ Detailed images can be produced automatically ∑∑ Permanent record ∑∑ Can be used for thickness measurements
∑∑ Surface must be accessible and smooth ∑∑ Can have operator dependence ∑∑ Flaw orientation important – linear defects oriented parallel to the sound beam may go undetected ∑∑ Interpretation can be difficult ∑∑ Need for reference standards ∑∑ Difficulty with complex geometries ∑∑ Inability to pass through air – need for couplant
Phased array eddy current
∑∑ Cracks, inclusions, dents, and holes ∑∑ Detect lack of penetration ∑∑ Detects galling ∑∑ Coating and material thickness ∑∑ Surface and near surface defects ∑∑ Composition/ conductivity/ permeability ∑∑ Grain size/ hardness ∑∑ Dimensions and geometry ∑∑ Alloy sorting
∑∑ Fast ∑∑ Inspection done in one pass ∑∑ Allow bead width sizing (indirect detection of oxide layers) ∑∑ Full coverage of weld ∑∑ C-scan imaging for easy interpretation ∑∑ Easy to operate ∑∑ Automation available ∑∑ Permanent record available ∑∑ Specimen contact not necessary
∑∑ Manual surface testing is slow ∑∑ Interpretation may be difficult ∑∑ Depth of penetration is limited ∑∑ Flaw orientation is critical ∑∑ Specimen must be electrically conductive ∑∑ Sensitive to many specimen parameters ∑∑ Surface roughness can produce nonrelevant indications
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is the Process development and qualification phase. The end application will determine the degree of thoroughness that will be involved in this phase. The most widely used and trusted technique is design-of-experiments where a statistical approach is taken to understanding the relation of the various weld parameters to weld quality. From the Process development and qualification phase will come the weld schedules, which define the range of control required for the key FSW control parameters. These parameters, along with other sensed and/or calculated variables, can be monitored online to measure the quality of welds. As with other types of welding and manufacturing processes, FSW can be inspected off-line using traditional NDT techniques such as dye penetrant, radiography, ultrasonics, and phased array eddy current. A summary of the uses, advantages, and disadvantages is provided in Table 7.6.
7.5
References
1 NIST Statistical Handbook: http://www.itl.nist.gov/div898/handbook/pri/section1/ pri11.htm 2 The Use of Neural Network and Discrete Fourier Transform for Real-Time Evaluation of Friction Stir Welding – E. Boldsaikhan, E. M. Corwina A. M. Logar, and W. J. Arbegast South Dakota School of Mines and Technology, Rapid city, SD 57701, USA. 3 Phase space analysis of friction stir weld quality, Enkhsaikhan Boldsaikhan, Edward Corwin, Antonette Logar, Jeff McGough and William Arbegast South Dakota School of Mines and Technology; 510 East Saint Joseph Street; Rapid City, SD, 57702; USA; Friction Stir Welding and Processing IV, ed. Mishra, M.W. Mahoney, T. Lienert & K.V. Jata, TMS, 2007, pp. 101–111 4 NDE of Friction Stir Welds in Aerospace Applications, The right NDE process for inspecting friction stir welds is being investigated by David G. Kinchen and Esma Aldahir – Inspection Trends 2002. 5 Engineering Edge – Solutions by Design: http://www.engineersedge.com/inspection/ radiography.htm 6 An Overview of Nondestructive Evaluation Methods by Marvin Trimm; Practical Failure Analysis – Evaluations, Vol. 3 (3), June 2003. 7 Wikipedia definition: http://en.wikipedia.org/wiki/Dye_penetrant_inspection 8 NDT Resource Center – http://www.ndt-ed.org/EducationResources/CommunityCollege/ Ultrasonics/Introduction/description.htm 9 Eddy current array and ultrasonic phased-array technologies as reliable tools for FSW inspection by André Lamarre, Olympus NDT; Presentation (ppt) in Proceedings of the 6th International FSW Symposium, Saint Sauveur, 2006. 10 NDT Resource Center – http://www.ndt-ed.org/EducationResources/CommunityCollege/ EddyCurrents/Introduction/IntroductiontoET.htm
8
Residual stresses in friction stir welding
S. W. Williams, Cranfield University, UK and A. Steuwer, ESS Scandinavia, Sweden
Abstract: Residual stresses are always produced during welding and often cause problems, either due to distortion or performance degradation. This chapter provides an explanation of the origin of residual stresses during welding, how to measure them and potential problems that may arise due to them. General methods of residual stress mitigation, either through management or control, are discussed along with their applicability to friction stir welding. The effects of the friction stir welding process parameters on the residual stress profile are considered. The chapter concludes with examples of the application of active residual stress control systems using stress engineering methods during friction stir welding. Key words: residual stresses; process effects in welding; residual stress measurement; residual stress control by stress engineering; mechanical and thermal tensioning.
8.1
Residual stresses produced by welding
Residual stresses (RS) are defined as self-equilibrating forces that remain in a material once all external forces have been removed. They arise due to inhomogeneous deformation of thermal, mechanical or micromechanical origin (such as phase transformations) (Withers and Bhadeshia, 2001b). They are introduced in virtually all thermo-mechanical manufacturing steps such as rolling, quenching, welding, forming, cutting, machining and surface treatments. All categories of materials including plastics, glasses and metals, can accumulate residual stresses. We shall restrict our discussion here to commonly used engineering metal alloys which are typically polycrystalline. For these materials, residual stresses can be classified in terms of length scales on which they act over: ∑ Type-I: extend over macroscopic areas and are averaged over several grains. ∑ Type-II: extend between grains or sub-regions of grains and averaged over these areas. ∑ Type-III: act on the inter-atomic level such as around inclusions or dislocation. Whilst all types of RS may be present after welding, generally we are only 215
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concerned with type-I stresses. These stresses lead to the distortion problems often evident after welding. In addition they superimpose externally applied forces (during service) and can therefore have significant influence on the performance of components after welding. The self-equilibrating nature of RS means that their sum perpendicular to any surface in a component is equal to zero. This stress-balancing is frequently employed in their measurement. Because of the significant role of residual stresses produced by welding with respect to distortion and component life they have been the subject of extensive research and published literature (Masubuchi, 1980; Radaj, 1992; Feng, 2005).
8.1.1 Residual stress development in welding To understand how residual stresses arise in welding it is helpful to consider the temperature stress cycle of material affected by the welding process. Figure 8.1 shows the thermal and stress fields around the weld zone. In front of the weld zone there is a large area of compressive stress due to thermal expansion of the warming (and therefore) expanding material which is being constrained by the surrounding cold material. Where the stress exceeds the compressive yield stress, plastic flow occurs. It should be noted that close to the centre of the weld zone the yield stress will be significantly reduced by the high temperature. Behind the weld zone there is a region of tensile stress due to contraction of the cooling material, again being restrained by the surrounding cold material. If the stress level exceeds the yield stress then plastic flow will occur. Residual stresses are normally resolved into three components. These are the longitudinal stress in the direction of welding, transverse stress perpendicular to the direction of welding and normal stress through the thickness of the material. In welding we are generally mainly concerned with the longitudinal stresses, as these are usually much larger than the other stresses. This can be due to a number of reasons but especially the asymmetric shape of the thermal field. The normal stress is usually very small due to the small variation in temperature in this direction, along with lack of restraint. This may not be the case when using FSW to join relatively thick material in a single pass. To envisage how residual stresses arise we can consider the longitudinal stress and temperature history for a point X at the edge of the weld pool or nugget as shown in Fig. 8.1. Figure 8.2 shows schematically the temperature stress history as the weld zone passes by. It should be noted that the exact shape of the temperature yield stress curves shown in Fig. 8.2 will vary significantly for different materials but the overall concept will be similar. As the weld zone approaches, the temperature will rise, along with the compressive stress due to the expanding metal. The resultant strain will be accommodated elastically as shown by line AB. Eventually the compressive
Residual stresses in friction stir welding
Weld pool or stirred zone
217
Tensile stress due to thermal contraction
Welding direction
Point X – on the edge of the Compressive stress due to thermal weld pool expansion
Isotherms TMAX
8.1 Temperature and stress field around a welding process (adapted from Radaj, 1992).
D
A B C Point X Tensile residual stress
Tensile + Tensile yield stress D A
Temperature
Stress Compressive –
C
B Compressive yield stgress
8.2 Temperature stress history ABCD for point X on the edge of the weld pool or nugget as it traverses past the weld zone (adapted from Williams et al., 2008).
stress will exceed the compressive yield stress, and plastic flow will occur as shown by the line BC. As the temperature nears the melting point, the load that can be carried becomes very small (if point X is within the fusion or nugget zone, no significant load can be carried). Upon cooling, the temperature drops and a tensile stress develops due to the contracting material being restrained by the surrounding material. In this case all the tensile stress is accommodated elastically with no plastic flow occurring as shown by the line CD. When the material has cooled back down to room temperature, a tensile residual stress is present as shown by the large arrow in Fig. 8.2. It should be emphasised that the tensile residual stress principally arises due
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to plastic compressive flow on heating. It is not just due to shrinkage forces on cooling as is commonly stated. If there is no plastic compressive flow then there will be no tensile residual stress. From the preceding analysis it can be understood that the tensile residual stress is due to an unequal balance of elastic and plastic flow during the heating and cooling cycles. This is often referred to as thermal or plastic misfit or mismatch. It is clear that if the compressive and tensile plastic flows are equal then no residual stress will arise. This provides an insight on how to control residual stresses in welding – balancing of the compressive and tensile plastic flows. The magnitude of the tensile residual stress at any point depends on the amount of compressive plastic flow that has occurred in the heating cycle. Points further away from the edge of the weld zone will have lower peak temperatures, lower compressive stresses leading to less compressive plastic flow. The result of this will be that the tensile residual stress level will reduce as you move away from the edge of the weld zone. The result is that a typical longitudinal residual stress profile plotted transverse to the direction of welding is as shown in Fig. 8.3. There is a band of large tensile residual stress in the region of the weld. As the sum of the stresses in any one direction is equal to zero, the tensile residual stress in the weld zone is balanced by compressive residual stress further out in the component. It will be noted from Fig. 8.1 that on cooling there is a region of tensile stress behind the weld. As the material is soft in this region, it is possible that this tensile stress will exceed the yield stress and if this occurs then tensile plastic flow will occur. This will compensate for some of the compressive plastic flow that happens during the heating cycle. In this case the level of tensile residual stress will be reduced, with the largest reduction occurring on the weld centreline. This is a common occurrence in FSW with a reduction in the tensile peak in the nugget and the peak values occurring in the TMAZ or HAZ. This profile is commonly referred to as an ‘M’ profile. Longitudinal stress s Tensile +
Transverse distance from the weld centreline
Weld centreline Compressive –
8.3 Typical longitudinal residual profile across the weld centreline.
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219
In summary the large tensile residual stress that occurs in the vicinity of the weld arises from an imbalance in how the thermal strain is accommodated during heating and cooling. During heating some of the strain is accommodated elastically but a significant amount is accommodated plastically. During cooling the thermal strain is nearly all accommodated elastically. This difference between the level of compressive plastic flow on heating and tensile plastic flow on cooling leads to tensile residual stress being produced in the weld zone. This tensile residual stress is balanced by compressive residual stress further away from the weld zone.
8.2
Determination of residual stresses in friction stir welding (FSW)
Residual stresses cannot be measured directly, instead their magnitude and orientation is typically inferred from their effect on other material properties such as shape (strain), conductivity, velocity of sound, etc. (Withers and Bhadeshia, 2001a) The methods for determining RS can be classified into destructive and non-destructive techniques. Each technique, whether destructive or non-destructive, has its own advantages and disadvantages, and we review here the most commonly used techniques in brief, and refer to extensive literature for more details.
8.2.1 Destructive methods The majority of destructive techniques for determining residual stresses in metallic material rely on the change in shape (strain) of a material due to the relaxation of residual stresses. The methods include the slitting (or compliance) technique (Cheng and Finnie, 2006), as well as the contour method (Prime and Gonzales, 2000). Through-depth capabilities of the contour method have been reported using the slitting technique (Lee and Hill, 2007). The contour method was successfully applied to measuring residual stresses in friction stir welds (Prime et al., 2006). The residual stresses are then calculated using basic elasticity theory following certain assumptions about the character of the residual stress field, e.g. uniaxial, biaxial, etc. Owing to their destructive nature, the application of these methods is limited, both in terms of the ability to perform the test on components in service, as well as the range and spatial resolution obtainable. Probably the most frequently applied semi-invasive technique is holedrilling, where the shape change (strain) of the material around a hole due to relaxation is recorded by a strain gauge as a function of drilling depth, assuming an essentially biaxial stress system near the surface. There is a standard procedure for hole-drilling measurements (ASTM E-837, 1999) which specifies hole-depth, depth to diameter ratios, etc. The depth of measurement
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specified for hole drilling rarely exceeds a few mm. Although hole drilling provides relatively good qualitative profiles of stresses as a function of depth, the applicability to accurately map residual stresses, e.g. as distance from weld line in the bulk is rather limited. For example, the density of points that can be investigated using hole-drilling is limited, since the relaxation of residual stresses of one hole can falsify subsequent measurements in the immediate vicinity of a hole. Furthermore, the finite size of the strain gauges employed determines the minimum spacing. Over the last few years, a variant of the technique, called deep-hole drilling residual stress measurement, has been developed (Leggat et al., 1996; George et al., 2002 and George and Smith, 2005). This method is capable of measuring the residual stress up to 750 mm deep in bulk components. Like semi-invasive conventional hole-drilling, it can be applied to components in service but requires additional repairs to be undertaken following the hole-drilling procedure.
8.2.2 Non-destructive methods A variety of non-destructive techniques are employed for residual stress characterisation (Hauk, 1997; Fitzpatrick and Lodini, 2003; Withers, 2006). In general, they estimate the relative distribution and magnitude of residual stresses based on the change in some physical quantity. Typically these are lattice parameter, velocity of sound, conductivity or magnetic properties. Some form of calibration of the technique, or reference standard is required. The lattice parameters are determined using diffraction-based techniques. Using Bragg’s law, small but perceptible changes in the location of diffraction peaks can be used to infer the residual stress. Laboratory X-rays can determine the residual stresses at surfaces but require careful preparation of the surface to avoid artifacts. Synchrotron X-ray and neutron diffraction techniques have the ability to probe the residual stress field in bulk components up to several mm or even cm without any need for sample preparation. The latter methods also offer the highest spatial and strain resolution of all the RS characterisation techniques. To some extent it is possible to achieve depth-scanning capabilities with laboratory X-rays using layering methods. However, this technique requires the partial destruction of the sample using electro or mechanical polishing. In recent years several portable X-ray systems have become commercially available which allow the determination of residual stress in the field. Ultrasonic inspection techniques are based on the change in the propagation velocity of ultrasonic waves in the material caused by residual stress, called the acoustic-elasticity effect. The measured stresses are typically an average over the path length and area given by the ultrasonic transducer, which can be a few mm up to a hundred mm. Non-destructive evaluation techniques based on eddy current analysis have in the past few years found more widespread
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application. In this approach, the residual stress in the near surface region is profiled based on piezo-resistivity, i.e. the effect of stress on the electric conductivity. By analysing the conductivity as a function of frequency the residual stress distribution in the surface layers can be compared to standard samples whose stress distribution is known from X-ray analysis. The Barkhausen Noise Analysis (BNA) technique, also known as the magnetoelastic method, is another non-destructive technique that is frequently employed in the field of non-destructive testing. The basic principle of the technique is the determination of noise that results from domain wall movement upon magnetisation of the ferromagnetic material. The domain wall movements cause a small electric pulse which when amplified can be heard as noise. The noise depends on the general stress state as well as microstructure. The magnetic measurement relies on the calibration of the technique on the base material external to the specimen being investigated. Penetration depths vary between tense of microns and a few mm depending on the frequency and amplitude of the magnetisation field and the magnetic properties of the material. Therefore the application for BNA is mainly surface inspection. For the ultrasonic, eddy-current and magnetic measurement techniques the equipment is fairly compact and portable.
8.3
Effects of residual stresses produced by welding
8.3.1 Distortion The most noticeable and problematic effect of residual stresses arising from welding is distortion of the component. The six primary forms of distortion are: transverse and longitudinal shrinkage, rotational distortion, angular distortion, bending distortion and buckling (Masubuchi 1980). These are shown in Fig. 8.4. Transverse and longitudinal shrinkage plus rotation distortion occur due to the movement of material towards the weld zone. If no material is removed during the welding process (as in FSW) then the weld zone will become thicker. This is because during plastic flow only a small displacement of material occurs. During heating there is significant plastic flow to accommodate the thermal strain. However during cooling most of the thermal strain is accommodated elastically and the net result is material is displaced towards the weld zone. This can be seen in Fig. 8.5 which shows a laser conduction weld made without filler wire in 5 mm thick stainless steel plate. It can be seen that the weld is noticeably thicker than the surrounding plates, even though no material has been added. The result of this will be that the plates are narrower after welding, i.e. transverse shrinkage. Similarly the weld zone will be shorter than the plates, giving longitudinal shrinkage. If the amount of material displaced is different from one end of the weld to
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(a)
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8.4 Common modes of distortion that occur during welding: (a) longitudinal shrinkage, (b) transverse shrinkage, (c) rotational, (d) angular, (e) bending and (f) 1st and 2nd order buckling (adapted from Masubuchi 1980). Top surface of plates
5 mm
8.5 Thickening of the weld zone due to the imbalance of elastic and plastic flow in an autogenous laser conduction weld in stainless steel.
the other, rotational distortion occurs. It can be seen from Fig. 8.5 that the volume of material that is displaced is quite small in relation to the total weld metal volume. This means that the amount of distortion due to this effect is usually quite small unless very thick welds are being made. Angular and bending distortion are due to variations in thermal field, material displacement and therefore longitudinal residual stress through the thickness of the plate. This becomes more of a problem when the plates become thicker and it is more difficult to obtain uniform through thickness conditions. It is frequently a problem in FSW due to large through thickness
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temperature gradients. This is especially the case for thick material or when welding at high speed. Buckling distortion is a direct result of the load that welding residual stresses impart to a structure. The compressive residual stresses induced in the outer regions of the plates produce a load, known as the applied weld load (AWL). When the AWL exceeds the critical buckling load (CBL) of the plates then they will buckle. The size of the CBL depends on the stiffness and geometry of the plates. Generally the thinner and narrower the plates are, the lower the CBL will be. When buckling occurs, the final distortion level and shape is determined by the buckling mode. Note that the 1st buckling mode has the same shape as bending distortion. It is easy to confuse buckling and bending distortion but the causes and methods of dealing with them are quite different. The actual direction of buckling is very unpredictable, often depending on factors such as the direction of pre-existing residual stresses or clamping arrangement. Figure 8.6 shows two examples of distortion in FSW. Figure 8.6a shows lap welds made in 1.6mm aluminium sheet. Very large buckling can be seen with 50 mm of out-of-plane displacement. Figure 8.6b shows a butt weld in 1 m long 3.2 mm thick AA2024 sheet. In this case buckling has occurred (not bending due to the direction of displacement) along with some angular distortion giving a peak out-of-plane displacement of 22 mm. Both these examples illustrate how distortion can be major problem in FSW.
8.3.2 Component performance There can be significant effects on the performance of the component due to the residual stresses produced by the welding process. The residual stresses will add to the in-service stresses, potentially leading to enhanced rates of degradation and premature failure. This is especially the case when
(a)
(b)
8.6 Examples of buckling distortion in (a) a lap weld and (b) a butt weld made using friction stir.
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the component is subject to fatigue (Almen and Black, 1963, Masubuchi, 1980, pp 449–77, Bussu & Irving, 2003, Edwards et al. (2006), Withers, 2007, James et al., 2007) or stress corrosion cracking (SCC) (Masubuchi 1980 pp 478–9). In fatigue there are three effects due to residual stresses that need to be understood: 1 The effect of residual stresses on crack initiation and early growth. 2 The effect of residual stresses on latter stage crack growth. 3 Modification of the residual stresses by the presence and growth of the crack. Unfortunately these effects of the residual stresses are difficult to unravel because the fatigue properties are affected by, and can be dominated by, other factors introduced by the welding process. These include changes in microstructure, surface condition and weld defects. These factors vary across the weld zone along with the residual stress level (Pouget and Reynolds, 2008). It should be emphasised that small fatigue coupons will not contain representative welding residual stresses. There is a minimum sample size needed to include the welding residual stress field. The minimum size depends on the material and welding process but is generally of the order of a few hundred mm (Degramo et al., 1946, Altenkirch et al., 2008). The two stages of fatigue-crack initiation and crack propagation are subject to different processes and are influenced by residual stress fields in different ways. The principal difference is that crack initiation is mostly affected by the local residual stress, whilst crack growth is affected by the mesoscale residual stresses (which it in turn modifies). During crack initiation the action of the local residual stress will be identical to that of a mean stress applied at the crack initiation site. This stress is added to the remote applied mean stress. If the local residual stress is tensile, this will promote crack initiation. Compressive residual stress fields still permit crack initiation to take place but subsequent crack propagation is reduced. Significant crack growth can take place only where there is a tensile stress field to open the crack sufficiently to achieve the minimum threshold for growth. In general, though, crack initiation and early growth are dominated by other factors more significant than residual stress. The later stages of growth are affected by the residual stresses in that they add to the applied mean stress. However, in this case we have to take into account the effects of the crack on the residual stress field. Two effects can expected: stress enhancement at the crack tip and redistribution of the mesoscale residual stresses (similar to measurement of the stress level by hole drilling). Most importantly the effect of the tensile stress can be reduced in the high stress low cycle fatigue regime. This is because if the sum of the applied load and the residual stress (including the stress intensity factor at
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the crack tip) is greater than yield strength then the residual stress level will be reduced (Masubuchi, 1980 pp 449–77). This does not apply in the low stress high cycle regime and the tensile stress can be expected to increase the crack growth rate. In contrast compressive stresses will always reduce the rate, and this can be very significant. It will be appreciated that the effects of residual stresses are complicated, due not only to the varying residual stress levels in the weld alone but also the other factors often present in welds as mentioned previously. However, one can summarise that in general tensile stresses will reduce fatigue life, whilst compressive stresses will enhance life. The residual stress effects can be exacerbated by the fact that in FSW the peak levels of tensile residual stress are usually located in the TMAZ and/or HAZ. This area is often the most susceptible to corrosion so the residual stresses can also cause enhanced rates of stress corrosion cracking (Masubuchi 1980 pp 466–8).
8.4
Mitigation of residual stresses and their effects
8.4.1 Management of residual stresses Methods of management of residual stresses do not attempt to change directly the residual stresses produced by the welding process. Rather the stresses are managed such that their effects are minimised. There are many standard techniques that have been developed for minimising the effects of residual stresses in traditional fusion welding techniques and these have been well documented (Masubuchi, 1980; Radaj, 1992; Feng, 2005). Many of these techniques can be applied to FSW. These techniques are distinguished by when they are applied, before, during or after welding. The first consideration is the design of the component (Masubuchi, 1980, pp 619–26, Radaj, 1992, pp 248–52). The number and length of welds used in a structure should be minimised. For example, this may include the use of intermittent or stitch welds instead of continuous seams. The detail position of the weld should be considered, particularly with respect to the critical buckling loads, which will change as the structure is fabricated and also to the neutral axis. Symmetric weld profiles are very desirable, as this can effectively eliminate bending and angular distortion. In practice in FSW this would mean welding from both top and bottom or better would be to use a bobbin tool which has a shoulder on the top and bottom of the component. Many techniques have been developed that are applied during fabrication. Elastic or plastic pre-bending is a method whereby the parts to be welded are pre-set at an angle so that after welding the plates are pulled flat by the
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angular distortion (Masubuchi, 1980, pp 288–94). This technique is difficult to apply in FSW as the pre-bend angle has to be incorporated in the tooling and the plates will not be normal to the FSW tool. Throughout the welding process it is desirable to make sure the applied weld load (AWL) is less than the local critical buckling load (CBL). As the structure is fabricated both the CBL and the AWL will change. Weld sequencing is a practice whereby welds are carried out in an optimised order to control the AWL and the CBL (Tsai et al., 1999; Kadivar et al., 2000; Mochizuki, 2000). This is done so that the stresses from welds do not add up in an unfavourable way during the weld sequence. This technique includes the order of intermittent welds when they are used instead of a seam weld. The optimised sequence is often found by trial and error but nowadays can be done through numerical modelling (Kim and Brust, 2005, pp 274–87). Another common technique used during fabrication to minimise distortion is to apply restraint during welding (Masubuchi, 1980, pp 313–15). This is done either through tooling or tack welding. It is known that a higher the level of applied restraint can lead to lower levels of resultant distortion; however, this effect is only recently becoming quantified (Kim and Brust, 2005, pp 281, Josserand et al., 2007). However, the reduced distortion will be at the expense of higher residual stress levels. In FSW the parts are necessarily highly restrained so distortion is often low but residual stress levels high. Post-weld techniques generally use a secondary process to add in more stresses in a way that counteracts the residual stresses from the welding process. These processes include thermal straightening and peening (shot or laser). Thermal straightening is normally used to correct buckling distortion in panels (Masubuchi, 1980, pp 322–4). This is achieved by heating material using flames, arcs or induction in a various patterns. This induces material to flow towards the areas that have been heated and effectively stretches the surrounding material and thereby pulls it flat. The problem with this is that quite high temperatures are needed that may not be desirable for heattreatable aluminium alloys. Shot peening is a well-established process for forming aluminium components in the aerospace industry (Friese, 2004) and can be used to eliminate distortion (Masubuchi, 1980, pp 632–3). It works by inducing compressive stresses in the top layer of material and these are used to distort the component to the desired shape. This can be used to correct the weld induced distortion in FSW but it is a very costly process. Any change in welding process or component design requires the development of a new peening strategy. Because the process introduces compressive stresses, it will also improve fatigue performance (Masubuchi, 1980, pp 470–2, Hatamleh, 2008).
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8.4.2 Control of residual stresses and distortion in welding The processes described above may be beneficial to reduce distortion but they do not actually change the residual stresses induced by the welding process. This can be a problem if the residual stresses lead to degraded performance. Also if changes are made to the welding process or the design of the component then the stress management methods often need to be revised. There are several techniques available to control the residual stresses from the welding process and these are collectively referred to as stress engineering (SE). Again these techniques can be distinguished in terms of those that are applied during or after welding. SE techniques are primarily aimed at reducing or eliminating welding residual stresses and associated buckling distortion. This requires a reduction in the AWL which is produced by the longitudinal compressive stresses in the structure. Reduction of the (balancing) longitudinal compressive stress is achieved by minimising the magnitude of the longitudinal tensile residual stress zone around the weld centreline. This can be achieved in two ways. Either the width of the tensile residual stress zone can be made narrower or its peak level can be reduced. The tensile region can be narrowed by optimising the welding parameters or by using lower heat input welding processes such as keyhole laser or electron beam (Colegrove et al., 2009). To reduce the peak level of the tensile stress significantly SE methods need to be used. As discussed previously this requires balancing the levels of compressive plastic flow and tensile plastic flow. There are two basic approaches: 1. Reduction of the amount of compressive yielding directly around the weld (RCY). 2. Induced tensile yielding, either while the weld metal is hot – just behind the weld, or cold – post welding (ITY). Both RCY and ITY are achieved by longitudinal tensile stretching of the material around the weld zone. For RCY this has to be applied in the compressive zone in the front of the weld zone as shown in Fig. 8.1. For the ITY the load needs to be applied in the tensile region just behind the weld zone, also shown in Fig. 8.1. It will be clear that if a tensile load is applied to the whole weldment then this will instigate both RCY and ITY. The effect of the applied tensile load on the temperature stress cycle for point X near the weld zone is illustrated in Figure 8.7. A tensile load L1 is applied as shown, so that at the starting position A, the material is in tension. As the weld zone approaches the increasing stress is accommodated elastically as before, shown by the line AB. However, line AB now intercepts the compressive yield stress curve at a higher temperature but at a lower stress level. The result is that the amount of plastic strain, shown by the line
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Tensile + D A
A Point X
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Compressive yield stress
8.7 Temperature stress cycle for a point close to the weld zone showing how adding a small tensile load L1 during welding reduces the final tensile residual stress peak (adapted from Williams et al., 2008).
BC, is less than in the case with no load applied. As the weld zone moves away the tensile stress is accommodated elastically, as shown by the line CD. After removal of the load L1; the height of tensile residual stress peak is reduced due to the reduced amount of compressive plastic strain that occurred in the temperature stress cycle (BC in Fig. 8.7). If a much larger load L2 is applied then the tensile stress in the weld zone can be replaced by a compressive stress. This is illustrated in Fig. 8.8, where as point X moves through the weld zone, it passes through the temperature stress cycle ABCD. During the heating phase the tensile yield stress curve is intercepted and tensile plastic strain occurs, BC in Fig. 8.8. Upon completion of the cycle ABCD it can be seen the final stress level, D, is lower than the starting stress level A. After removal of the load L2, the result is a compressive stress field in the weld zone. In fact, based on this simple theory one would expect that the magnitude of the residual stress peak would vary linearly with the applied load. With no load a large tensile peak is present. As the applied load increases, the magnitude of the tensile peak reduces until there is no residual stress. Further increases in load produce compressive stresses in the weld zone. Post-weld SE techniques similarly apply a longitudinal tensile load to the weld length. In this case only ITY can be instigated and it allows reduction of the height of the tensile peak produced by the welding process. This can be understood from Fig. 8.9 in which the stages of the application of post-weld SE are illustrated. Initially there is the usual tensile residual stress either side of the weld centreline with balancing compressive residual stresses further out. The room temperature tensile yield stress is indicated. When the SE
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Compressive A B C D Tensile yield residual Tensile + stress Point X A stress D Small load L2 applied
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8.8 Temperature stress cycle for a point close to the weld zone showing how adding a large tensile load L2 during welding induces a compressive residual stress zone around the weld (adapted from Williams et al., 2008). Induced tensile yielding
Tensile +
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8.9 Illustration of the method of post-weld global mechanical tensioning to reduce the longitudinal tensile residual stress by induced tensile yielding.
load is applied, the residual stress curve eventually intercepts the yield stress curve. Further increases in load cause tensile plastic flow and relieving of the tensile peak as shown. Finally after removal of the load, the tensile peak has been much reduced and therefore also the balancing compressive stresses. An alternative approach is to apply a post-weld heat treatment (PWHT) (Masubuchi, 1980, pp 630–1). The PWHT has the effect of lowering the material tensile yield stress below the tensile residual stress and therefore inducing tensile yielding. The result is reduction of the peak level of tensile residual stress. The PWHT may be applied for other reasons such as strengthening of the weld, especially in heat-treatable aluminium alloys.
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The variation in SE techniques is in the manner in which the load is generated. The techniques fall broadly into two classes: ∑ Thermal tensioning (TT) – Heating (TTH) – Cooling (TTC) ∑ Mechanical tensioning (MT) – Global (GMT) – Local (LMT). Thermal tensioning (TT) has been extensively studied (Williams et al., 2001; Gabzydl et al., 2001; Guan, 2005) and is achieved through the use of either hot or cold zones to generate the required loads. The effect of hot and cold zones in TT can be understood from Fig. 8.10. With the orientation shown, the hot zone generates compressive loads in front and behind due to the expanding material. On either side of the hot zone tensile loads are generated due to the stretching effect of the nearby expanding material. In contrast, the stresses generated round the cold zone are the exact opposite, as shown in Fig. 8.10. To use the hot or cold zones it is necessary to place them in specific optimised positions with respect to the weld zone. This is illustrated in Fig. 8.11 for the case of a) ITY and b) RCY using cold and hot zones respectively. Methods of generating the hot zones include simple burners, induction systems and lasers. Liquid CO2 is the preferred method for obtaining cold zones because the cooling power is very large, typically a few kW so that large loads can be generated in a small area. Extensive work has been carried out using liquid CO2 because of the simplicity of applying it. However, a major practical issue has been the problem of applying the CO2 in close proximity to the weld zone due to disruption of either the arc or the molten weld pool. But of course neither of these applies to FSW, so CO2 cooling is very suitable. Weld zone
Welding direction
Stress orientation T C T – Tensile C – Compressive
C
Hot T
C
T
Cold
T
C
8.10 Diagram showing the longitudinal stress fields around hot and cold spots (adapted from Williams et al., 2008).
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T C
Hot
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Welding direction
C
T T
Cold
T
T C
Hot T (a)
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T – Tensile C – Compressive
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Stress orientation (b)
8.11 Optimum positioning of cold and hot zones for (a) reduced compressive yielding and (b) induced tensile yielding (adapted from Williams et al., 2008).
Mechanical tensioning (MT) methods have the advantage of being more controllable and they do not affect the welding process. Global mechanical tensioning (GMT) uses a hydraulic system to stretch the whole component along the weld centreline. This process has the major advantage that it produces highly controlled loads that do not affect the welding process. As a result it has been studied extensively, both experimentally (Staron et al., 2004; Price et al., 2007; Altenkirch et al., 2008; Altenkirch, 2009) and theoretically (Richards et al., 2008b). When applied during welding the application of GMT instigates both RCY and ITY. The main drawback of this method is that large loads and therefore expensive equipment are required, especially for thicker material. Also it is only possible to apply it easily to straight welds. Local mechanical tensioning is achieved using rollers (Williams et al., 2008). When a roller is pushed into the surface of a material, the downward load displaces material sideways. This is similar to the expansion of material sideways in a hot spot and the resulting stresses are the same as shown in Fig. 8.10. Tensile stresses are generated either side of the roller and compressive stresses in front and behind.
8.5
Residual stresses in friction stir welding (FSW)
Residual stresses are a feature of friction stir welding in a similar manner to other welding processes (Sutton et al., 2002). It is a commonly held belief that one of the benefits of friction stir welding is that it is a low distortion process but this is not in fact true. Residual stresses and distortion are as significant in FSW as they are in most welding processes except for keyhole processes. The reason is that FSW is actually quite a high heat input process.
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For example, welding 6 mm thick high strength aluminum with a welding speed of 0.3 m/min and a power input 4 kW corresponds to a line energy (or heat input per unit length) of 800 kJ/m – which is similar to arc welding processes. Parameter optimisation can reduce this heat input somewhat but it still does not approach that of very low distortion processes such as laser welding. The reason that FSW is sometimes considered a low distortion process is the low peak temperature of the process (the maximum is the melting temperature). In fusion welding processes although there are higher peak temperatures, no load can be carried by the molten material. Therefore the residual stresses produced by FSW and arc welding are quite similar.
8.5.1 Process effects The important factor that determines the residual stress profile in all welding processes is the amount and how the thermal strain is accommodated. This is mainly determined by the line energy which is calculated by dividing the power input by the travel speed. Therefore to understand the effect of the processes parameters on residual stresses it is necessary to know their effect on the power input. The main parameters that affect the power input profile are the travel speed and the tool rotation rate. Other process parameters include tool downforce, profile and angle, type and size of backing plate and clamping devices. The heat in FSW is generated by a combination of friction between material being joined and the tool surface (both the shoulder and the pin) and by plastic working of the softened material. The extent of the zone over which plastic work is induced and the intensity vary significantly depending process parameters. This means that the power input profile in FSW is rather complicated. In conventional welding processes the power input and travel speed are independent of each other so that control of the line energy is straightforward. In contrast this is not the case in FSW, with the power input increasing as the travel speed increases (for a fixed rotation rate) (Colegrove et al., 2007). The actual power input to the workpiece depends strongly on the condition, especially temperature, of the material that is being stirred. Similarly, the effect of tool rotation rate is not simple. Initially the power input increases with increasing rotation rate but this eventually reaches a maximum. This corresponds with the temperature approaching the solidus. In this region the material is very soft, with increased slip between the tool and the material. This leads to a reducing rate of increase of heat generation through plastic work. Further increases in rotation rate do not lead to increasing power input. Additionally, the actual power profile changes with welding conditions. With relatively cold conditions the power is deposited over quite a wide zone due the large area of plastic work (Colegrove et al., 2007; Zhang and Zhang, 2009). When the conditions are hotter, the heat generation zone becomes narrower as the material round the tool becomes
Residual stresses in friction stir welding
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much softer. The result is that the power profile becomes narrower even though the total power input is increasing. All of these effects combine such that the effect of process parameters on residual stress profiles in FSW is unlike other welding processes (Steuwer et al., 2006; Peel et al., 2006a, 2006b,). This can mean that the peak residual stress value can actually reduce with increasing line energy rather than increase (Lombard et al., 2009). However, the width of the zone of tensile residual stress will increase, so the overall AWL will also increase. Quite often in FSW there is a significant temperature gradient top to bottom through the material. This is particularly the case for thicker materials or when welding at high travel speeds. This will mean that there will be a large variation of longitudinal tensile residual stress through the thickness as shown in Fig. 8.12 for 20 mm-thick AA7449. At the top surface there is a large tensile residual stress of more than 150 MPa dropping to only 25 MPa on the bottom surface. This can cause a problem by inducing significant bending distortion into the component. Process parameters should be selected to minimise the through thickness temperature profile and therefore minimise the variation of tensile residual stress through the thickness. The other important factor in determining the residual profile is the material type. Figure 8.13 shows the peak residual stress measured for various aluminium alloy types with different thicknesses. Points to note are that the average level is alloy dependent but that there is wide scatter for any particular alloy. This will depend on the precise welding conditions. There is no clear dependence on thickness (measurements were made close to the top surface). It should be noted that some aluminium alloys are age hardenable so that the residual stress profile may change with time (Linton
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8.12 Cross-sectional profile of the longitudinal residual stresses in 18 mm-thick AA7449 showing the large variation of stress through the thickness of the plate corresponding to the large through thickness variation in temperature (adapted from Altenkirch et al., 2008).
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8.13 Peak tensile residual stress values for various alloy types and thicknesses.
and Ripley, 2008). One key attraction of FSW is the ability to join dissimilar materials, and some investigations on residual stress profiles in dissimilar aluminium alloys have been reported in the literature (Steuwer et al., 2006; Prime et al., 2006; Jun et al., 2008), as well as for steels (Reynolds et al., 2003).
8.6
Active control of residual stresses in friction stir welding (FSW)
8.6.1 Thermal tensioning Thermal tensioning (TT) methods are not often used with FSW. This is because TT methods rely on a temperature gradient to be developed to generate the required stresses. FSW is usually used on light alloys such as aluminium and they often have high thermal conductivity which makes the generation of thermal stresses of the required magnitude quite difficult. In addition many aluminum alloys are heat treatable and it is undesirable to have a temperature above about 150°C. However, TT using cryogenic CO2 cooling sprays with FSW has been investigated both experimentally and theoretically (Gabzdyl et al., 2001, Richards et al., 2008a). The equipment used for the investigations is shown in Fig. 8.14. The CO2 was delivered via a simple pair of copper tubes and directed to various positions around the FSW tool. It was applied to 3.2 mm-thick 2024 aluminium panels with dimensions of 300 ¥ 125 mm. The best results were obtained when the sprays were directed as close to and directly behind the FSW tool. The effect on the residual stress profile is shown in Fig. 8.15. There was a significant reduction in the residual longitudinal tensile stress close to the weld centreline. However there was little effect of the main peaks of the residual stress which are located in the TMAZ. The reduction of tensile stress was insufficient to reduce the compressive stresses sufficiently to eliminate buckling distortion in these particular panels. The application of TT using cooling has been studied using FE modeling. The
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8.14 System used to deliver cryogenic CO2 to the rear of a FSW tool.
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With cooling No cooling
Weld centreline
Residual stress (MPa)
75 50 25 0 –25 –50 –75 –100 –125 –40
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8.15 Effect on the longitudinal residual stress profile of thermal tensioning by cooling during FSW of 3.2-mm AA2024 (residual stress measurements courtesy of GKSS).
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results show that with a fully optimised cooling source large changes in residual stresses could be obtained. No studies using TT with heating have been done due to the problems of over heating aluminium alloys.
8.6.2 Global mechanical tensioning Early investigations of mechanical tensioning on arc welds and FSW were found to be beneficial both with respect to distortion as well as residual stress (Price et al., 2007). A system that has been developed for application of global mechanical tensioning for FSW of integrally stiffened panels (ISP) is shown in Fig. 8.16. In the foreground are the hydraulic rams that can produce a load of up to 600 bar. The ISPs are attached to the rams using bolted clamps. The result of applying GMT to FSW of 5 mm-thick AA2199 material is shown in Fig. 8.17. It can be seen that as the load is increased, the peak residual stress value is reduced and eventually changes from tensile
8.16 Mechanical tensioning system developed for FSW of integrally stiffened panels (picture courtesy of Airbus).
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8.17 The effect of global mechanical tensioning load on the tensile residual stress profile (adapted from Altenkirch et al., 2008).
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8.18 Change in out-of-plane distortion and peak longitudinal residual stress as a function of the applied load when using GMT during the FSW of 3.2 mm-AA2024 (adapted from Price et al., 2007).
to compressive. It was postulated in Section 8.1.3 that the peak residual should be linearly dependent on the applied load. Using GMT it is possible to investigate this (Price et al., 2007; Altenkirch et al., 2009; Altenkirch, 2009). Figure 8.18 shows the peak out of plane distortion and longitudinal residual stress values as a function of applied load when applying GMT using the FSW of 1 m-long, 3.2 mm-thick 2024 sheets. Both the distortion
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and residual stress can be seen to show the expected linear dependence on applied load. Altenkirch et al. (2009) have shown this to be the case for a wide variety of materials. This can also be seen when looking at the through thickness residual stress profiles. Figure 8.19 shows the through thickness residual stress profile for 20 mm-thick 7449 material when a GMT load of 34% of yield was applied. This should be compared to Fig. 8.12, which shows the same weld made without the application of GMT. It can be seen that the shift in residual stress value has been nearly uniform throughout the whole thickness of material. There is now a small tensile residual stress of 50 MPa on the top surface and a compressive stress of nearly 150 MPa on the bottom surface. GMT can also be applied post welding but in this case only ITY can be instigated. The effect of applying it post welding is to eliminate the tensile residual stress peak. However, the loads required are much higher than when applying it during welding (Price et al., 2007).
8.6.3 Local mechanical tensioning Local mechanical tensioning is carried out using rollers. A roller tensioning system developed for use with FSW is shown in Fig. 8.20 (Williams et al., 2008; Altenkirch et al., 2009). Roller tensioning can be applied during welding but best results are achieved by applying the load directly on the weld line after welding. For example, Fig. 8.21 shows the change in residual stress by applying rolling to FSW made in 5 mm-2199 alloy. The rolling process effectively eliminates the tensile residual stress with a load in this case of only between 10 and 15 kN. This relatively low load requirement makes this process extremely attractive for FSW as the rolling system can be integrated
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8.19 Cross sectional profile of the longitudinal residual stresses in 20 mm-thick AA7449 showing the effect of applying global mechanical tensioning of 34% of yield (compare to Fig. 8.13) (adapted from Altenkirch et al., 2008).
Residual stresses in friction stir welding
239
8.20 Roller tensioning system being applied to integrally stiffened panels after being joined by FSW (picture courtesy of Airbus UK). 300 250
AA2199 T8 Unc. ±12MPa
as welded 10kN
Longitudinal RS (MPa)
200
20kN
150
30kN 40kN
100 50 0 –50 –100 –150 –200 –40
Retreating side –30
Advancing side
–20 –10 0 10 20 Lateral distance from weld line (mm)
30
40
8.21 Effect of post-weld roller tensioning on the longitudinal residual stresses in FSW in AA2199-T8 (adapted from Altenkirch et al., 2009).
into the FSW machine. The load that needs to be applied in any particular case depends on the material properties, thickness and the roller design. The system shown has been applied to large scale fabrication of stiffened panels made by using FSW to join integrally stiffened panels. The initial panel size was approximately 4 m ¥ 0.4 m and 6 of these were joined together using FSW to make a large panel with a size of 4 ¥ 2.4 m as shown in Fig. 8.22.
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After welding, the panels typically had out-of-plane distortion of more than 20 mm causing difficulty in machining to shape. To eliminate the distortion all five welds were rolled sequentially and the result is shown in Fig. 8.23. The peak distortion is the maximum out-of-plane distortion, whilst the distortion index represents the average distortion measured at several points
8.22 Large stiffened panel made by joining together six integrally stiffened panels using FSW (picture courtesy of Airbus UK). 25 Peak distortion and distortion index (mm)
Peak distortion Distortion index
20
15
1
5
0
0
1
2 3 Roll number
4
5
8.23 Effect of roller tensioning on the distortion of a large panel fabricated by joining together six smaller integrally stiffened panels using FSW (courtesy of Airbus UK).
Residual stresses in friction stir welding
241
around the perimeter of the panel. Each weld was rolled in turn and it can be seen that the distortion reduced as each weld was treated. Eventually the panel was essentially flat. Most importantly the panel was then machined to shape and after milling the panel still remained flat. This is because the weld residual stresses have been reduced rather than managed. This is in contrast to many residual stress management methods where additional stresses are introduced, which often lead to subsequent problems in the manufacturing process. To ensure repeatability the rolling process was applied to three different welded panels. All of the panels responded to the rolling process with effective elimination of distortion in the panels. In summary, residual stresses can be a major concern when applying FSW. This can either be due to distortion of the component or in-service issues. Some limited control of the residual stresses can be obtained through manipulation of the process parameters. Traditional residual stress management methods such as optimised design, weld sequencing and peening can be used with FSW. Stress engineering methods allow direct and full control of the weld residual stresses. Both thermal and mechanical methods have been successfully applied to FSW. One of the most effective and practical ways of reducing residual stresses in FSW is the rolling of welds, especially as the levels of distortion are low after rolling and stay very low after post processes such as milling.
8.7
References
Almen, J. O. & Black, J. P. H. (1963), Residual Stresses and Fatigue in Metals, Toronto, McGraw-Hill. Altenkirch, J. (2009), Stress engineering of Friction Stir Welding: Measurement and Control of Welding Residual Stresses (PhD Thesis). School of Materials. Manchester, University of Manchester. Altenkirch, J., Steuwer, A., Peel, M. J., Richards, D. G. & Withers, P. J. (2008), The effect of tensioning and sectioning on residual stresses in aluminium AA7449 friction stir welds. Materials Science and Engineering A, 488, 16–24. Altenkirch, J., Steuwer, A., Withers, P. J., Williams, S. W., Poad, M. & Wen, S. W. (2009), Residual stress engineering in friction stir welds by roller tensioning. Science and Technology of Welding and Joining, 14 185–192. ASTM (1999), Standard test method for determining residual stresses by the hole-drilling strain-gage method, ASTM Standard E837-99. Philadelphia, PA: American Society for Testing and Materials. Bussu, G. & Irving, P. E. (2003), The role of residual stress and heat affected zone properties on fatigue crack propagation in friction stir weld 2024 T351 aluminium joints, Int. J. Fat. 77–88. Cheng, W. & Finnie, I. (2006), Residual stress measurements and the slitting method, Springer. Colegrove, P. A., Shercliff, H. R. & Zettler, R. (2007), Model for predicting heat generation and temperature in friction stir welding from the material properties’, Science and Technology of Welding & Joining, vol. 12, pp 284–297.
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Colegrove, P. A., Ikeagu, C., Thistlethwaite, A., Williams, S. W., Nagy, T., Suder, W., Steuwer, S. & Pirling, T. (Accepted, 2009), “The welding process impact on residual stress and distortion”, Science and Technology of Welding & Joining. Degramo, E. P., Meriam, J. and Jonassen, F. (1946), The effect of weld length upon the residual stresses of unrestrained butt weld, Welding J. 25(8) pp 475s–486s. Edwards, L., Fitzpatrick, M. E., Irving, P. E., Sinclair, I., Zhang, X. & Yapp, D. (2006), An Integrated approach to the determination and consequences of residual stress on the fatigue performance of welded aircraft structures, ASTM Int. J. JAI, 2547. Feng, Z. (Ed) (2005), Processes and mechanisms of welding residual stress and distortion, Cambridge, Woodhead, Fitzpatrick, M. E. & Lodini, (2003), Analysis of residual stress using neutron and synchrotron, London, Taylor and Francis. Friese, A. (2004), One of the worlds largest shot peening machines installed at Airbus, Metal Finishing News, Vol. 5(May), 18–20. Gabzdyl, J., Cole, M., Williams, S. W. & Price, D. (2001), Control of laser weld distortion by thermal tensioning. 20th ICALEO 2001, Vols 92 & 93, Congress Proceedings, 1236–1245. George, D. & Smith, D. J. (2005), Through thickness measurement of residual stresses in a stainless steel cylinder containing shallow and deep weld repairs, Int. J. of Pressure Vessels and Piping, 82(4), 279–287. George, D., Kinston, E., & Smith, D. J. (2002), Measurement of through-thickness stresses using small holes, The J. of Strain Analysis for Eng. Des., 37(2), 125–139 Guan, Q. (2005), Control of buckling distortions in plates and shells, in Zeng, Processes and mechanisms of welding residual stress and distortion, Cambridge, Woodhead. Hatamleh, O. (2008), The effects of laser peening and shot peening on mechanical properties in friction stir welded 7075-T7351 Aluminum. Journal of Materials Engineering And Performance, 17, 688–694. Hauk, V. (1997), Structural and residual stress analysis by non-destructive methods, Amsterdam, Elsevier Science. James, M. N., Hughes, D. J., Chen, Z., Lombard, H., Hattingh, D. G., Asquith, D., Yates, J. R. & Webster, P. J., Residual stresses and fatigue performance, Engineering Failure Analysis Vol. 14, Issue 2, March 2007, Pages 384–395. Josserand, E., Jullien, J.-F., Nellas, D., Boituout, F. & Deloison, D. (2007), Numerical simulation of welding induces distortions taking into account industrial clamping conditions, Mat. Mod. of Weld Phen. P1073–1092 ISBN:9783902465696.s Jun, T. S., Zhang, S. Y., Golshan, M., Peel, M., Richards, D. & Korsunsky, A. M. (2008), Synchrotron energy-dispersive X-ray diffraction analysis of residual strains around friction welds between dissimilar aluminium and nickel alloys. Stress Evaluation In Materials Using Neutrons And Synchrotron Radiation, 571–572, 407–412. Kadivar, M. H., Jafarpur, K. and Baradaran, H. G. (2000), Optimization of welding sequence with genetic algorithm, Comp. Mech. Vol. 26, 514–519. Kim, D. S. & Brust, F. W. (2005), Mitigating welding residual stress and distortion in Feng Z., Processes and mechanisms of welding residual stress and distortion, Cambridge, Woodhead, 264–294. Lee, M. J. & Hill, M. R. (2007), Intralaboratory repeatability of residual stress determined by the slitting method. Experimental Mechanics, 47, 745–752. Leggat, R. H., Smith, D. J., Smith, S. D. & Farue, F. (1996), Development and experimental validation of the deep hole method for residual stress measurement, J. of Strain Analysis, 31(3), 177–186
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Linton, V. M. & Ripley, M. I. (2008), Influence of time on residual stresses in friction stir welds in age hardenable 7xxx aluminium alloys. Acta Materialia, 56, 4319–4327. Lombard, H., Hattingh, D. G., Steuwer, A. & James, M. N. (2009), Effect of process parameters on the residual stresses in AA5083-H321 friction stir welds. Materials Science and Engineering: A, 501, 119. Masubuchi, K. (1980), Analysis of welded structures, Oxford, Pergammon. Mochizuki, M., Hayashi, M. and Hattori, T. (2000), Residual stress distribution depending on welding sequence in multi-pass welded joint with x-shaped groove, J. of Pressure Vessel Tech., 122, 27–32. Peel, M. J., Steuwer, A. & Withers, P. J. (2006a), Dissimilar friction stir welds in AA5083-AA6082. Part II: Process parameter effects on microstructure. Metallurgical And Materials Transactions A-Physical Metallurgy and Materials Science, 37A, 2195–220. Peel, M. J., Steuwer, A., Withers, P. J., Dickerson, T., Shi, Q. & Shercliff, H. (2006b), Dissimilar friction stir welds in AA5083-AA6082. Part I: Process parameter effects on thermal history and weld properties. Met. And Mat. Trans. A-Physical Metallurgy and Mat. Science, 37A, 2183-2193. Pouget, G. & Reynolds, A. P. (2008), Residual stress and microstructure effects on fatigue crack growth in AA2050 friction stir welds. International Journal of Fatigue, 30, 463-472. Price, D. A., Williams, S. W., Wescott, A., Harrison, C. J. C., Rezai, A., Steuwer, A., Peel, M., Staron, P. and Koçak, M. (2007), Distortion control in welding by mechanical tensioning, Sci. & Tech. of Weld, Vol 12 No 7, 620–633. Prime, M. B. & Gonzales, R. (2000), The contour method: simple 2-D mapping of residual stress. Sixth International Conference on Residual Stresses. Oxford, UK, IOM. Prime, M. B., Gnaupel-Herold, T., Baumann, J. A., Lederich, R. J., Bowden, D. M. & Sebring, R. J. (2006), Residual stress measurements in a thick, dissimilar aluminum alloy friction stir weld. Acta Materialia, 54, 4013–4021. Radaj, D. (1992), Heat effects in Welding, Berlin, Springer-Verlag. Reynolds, A. P., Tang, W., Gnaupel-Herold, T. & Prask, H. (2003), Structure, properties, and residual stress of 304L stainless steel friction stir welds. Scripta Materialia, 48, 1289–1294. Richards, D. G., Prangnell, P. B., Williams, S. W., Withers, P. J. and Morgan, S. (2008a), Simulation of the Effectiveness of Local Cooling as a Technique for Controlling Residual Stresses in Friction Stir Welds, in Proc of the 7th International Symp. on FSW, Kobe, Japan; 2008. Richards, D. G., Prangnell, P. B., Williams, S. W. & Withers, P. J. (2008b), Global mechanical tensioning for the management of residual stresses in welds. Materials Science and Engineering A-Structural Materials Properties Microstructure and Processing, 489, 351–362. Staron, P., Kocak, M., Williams, S. W. and Wescott, A. (2004), Residual stress in friction stir weld Al sheets, Phys. B: Condens. Matter 350 E491–E493. Steuwer, A., Peel, M. & Withers, P. J. (2006), Dissimilar friction stir welds in AA5083AA6082. The effect of process parameters on residual stress. Materials Science and Engineering A, 441, 187–196. Sutton, M. A., Reynolds, A. P., Wang, D. Q. & Hubbard, C. R. (2002), A study of residual stresses and microstructure in 2024-t3 aluminum friction stir butt welds, J. of Eng. Mat. and Tech. – Transactions of The ASME, 124, 215–221. Tsai, C. L., Park, S. C. & Cheng, W. T. (1999), Welding distortion of thin-plate panel structure, The Welding J. 78(5) 157–165.
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Williams, S. W., Scott, G., Gabzdyl, J., Cole, M., Staron, P. & Kocak, M. (2001), In Proc. First Int. WLT-Conf. Lasers In Manufacturing, Munich, Germany, 2001, 188–198. Williams, S. W., Morgan, S. A., Wescott, A. W., Poad, M. and Wen, S. (2008), Stress engineering – Control of residual stresses and distortion in welding, in Proc. Intl. Workshop on Thermal Forming and Welding Distortion IWOTE’08, Bias Verlag, 229–239. Withers, P. J. (2006), Residual stresses: measurement by diffraction, Encyclopaedia of Materials, Science and Technology, 8158–8169 Withers, P. J. (2007), Residual stress and its role in failure, Rep. Prog. Phys. 70 2211. Withers, P. J. & Bhadeshia, H. K. D. H. (2001a), Residual Stress. I. Measurement Techniques. Materials Science And Technology, 17, 355–365. Withers, P. J. & Bhadeshia, H. K. D. H. (2001b), Residual stress. II. Nature and origins. Materials Science and Technology, 17, 366–375. Zhang, Z. & Zhang, H. W. (2009), Numerical studies on controlling of process parameters in friction stir welding. Journal of Materials Processing Technology, 209, 241–270.
9
Effects and defects of friction stir welds
R. Zettler, WTSH, Germany, T. Vugrin, Airbus, Germany and M. Schmücker, German Aerospace Centre, Germany
Abstract: One of the major reasons often cited for using friction stir welding (FSW), particularly when it comes to the joining of light alloys, is the low incidence of flaws that can occur compared with those produced by conventional fusion (arc) welding processes. The FSW process does, however, have its own characteristic flaws. These flaws have been observed to occur in response to different FSW process variables including tool rotation and weld travel speeds, tool plunge depth (axial force), tool tilt angle, tool design, the size of gap between workpieces, thickness mismatch or plate thickness variation, tool offset to the join line, a lack of tool pin length in relation to workpiece thickness, as well as naturally occurring surface oxides remaining entrapped within the stir zone. This chapter discusses the origins of effects and defects, i.e. welding flaws, that can occur when FSW, their implications with regard to residual mechanical strength of the joint, and makes recommendations as how to identify why these occur and how to avoid their occurrence. Key words: friction stir welding, flaw formation, process variables, material thermo-physical properties, recommendations to avoid flaw formation.
9.1
Background and introduction
Defect or flaw formation in this chapter has the same meaning, i.e. a linear or volumetric discontinuity in the weld, which adversely affects the functionality of the joint. For example, a number of flaw types are well established in the fusion welding of aluminium and its alloys. Most notably these flaws include weld porosity, solidification cracking, and heat affected liquation cracking (Leonard and Lockyer 2003). Additionally the heat produced to help facilitate metal parts in the region of the joint to melt and then fuse together (hence the term fusion welding) can lead to micro-segregation of alloying elements, including copper but also of magnesium and manganese and silicon due to solute rejection at the liquid/solid interface (Chong et al. 2003). Furthermore, the heat necessary to melt can cause significant weldment distortion but more so this heat contributes to microstructural change, such that the mechanical properties of the joint are significantly compromised when compared to the parent material(s) (Sato and Kokawa 2003). Friction stir welding (FSW), being a solid state joining process, does not 245
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suffer from problems such as porosity or hot cracking. This is because there is no bulk melting of the parent material in the joint during processing. The fact that FSW is a hot metal working process, however, seen to comprise distinct processing regions (Arbegast 2003), means that it does have the potential for imbalances between these zones to occur, and thus makes it susceptible to flaw formation. Flaw formation when FSW has been identified to occur as a consequence of either imbalances in material flow or geometric factors associated with the position of the tool in relation to the joint (Arbegast 2008). Under optimal processing conditions mass balance, both in terms of material volume and energy, is said to be achieved. This facilitates constant volume processing while ensuring minimal impact on the pre-existing microstructure. In many instances the properties of the friction stir welded joint even approach those of the parent material(s). Typically for hot metal working processes, energy in the form of heat is required to reduce the resistance of a material to deformation. This heat at the same time actively facilitates microstructural change, e.g. recrystallisation, coarsening and or dissolution of strengthening precipitates, grain reorientation and growth. A material’s resistance to both deformation and microstructural change is a function of the material’s chemical composition and since this varies between materials some undergo change at relatively low processing temperatures, while others will resist such change until much higher temperatures are reached. Hence when one speaks of too hot or too cold a processing condition, this is relative to the material and not a specific temperature which can be broadly applied across all materials. In terms of achieving mass balance during FSW this requires the prevention of either insufficient material flow, i.e. too cold a welding condition as witnessed by non-bonding, volumetric or void formation, or too hot a welding condition, giving rise to excessive material flow leading to material expulsion, e.g. flash formation, the collapse of the nugget within the stir zone, and additionally an unwanted degradation of the mechanical properties of the joint. Unlike flow-related flaws, geometric flaws, as the name suggests, are a consequence of incorrect tool placement/adjustment in relation to geometrical features of the joint. Such flaws can come about due to operator error and include incorrect tool pin length in relation to workpiece thickness, known to produce a lack of penetration (LOP) flaw, or incorrect tool position to the joint line such that the tool pin preferentially interacts with one of the two workpiece, i.e. the tool is pushed too far into the advancing side of the joint. This can result in a lack of fusion (LOF) between the workpieces even under hotter processing temperatures, since such thermal softening often facilitates a sideways displacement of the original interface between workpieces, i.e.
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transverse to the weld travel direction without bonding between interfaces taking place. In addition to both FSW flow and geometric flaws there also exists the potential to produce joints that, although comparable in every way with those of flaw free joints, possess traces of the former natural surface layer, e.g. oxides within the weld nugget. The traces often appear as very thin snake-like lines frequently referred to in the literature as root- or weld nugget flaws. Potential FSW flaws as relating to flow, geometric and root- or nugget are presented in Fig. 9.1. All FSW flaw types influence the static and dynamic properties of the friction stir welded joint, both in relation to the base or parent material and when compared against those of a flaw-free welded joint. Figures 9.2 and 9.3 highlight the extent of degradation of properties with respect to LOP and oxide entrapment flaws for the aerospace grade aluminium alloy 6013 in the T6 temper. The data presented in each figure is courtesy of Daniela Lohwasser, Airbus Germany.
9.2
Defects from too hot welds
There are primarily two conditions that define the production of too hot a friction stir weld: 1. Too high a processing temperature, i.e. the generation of processing Surface galling
Surface lack of fill
Excessive flash
Nugget collapse
Tunnel flaw
Material loss at anvil
2 mm
Lack of penetration
50 µm
Lasck of fusion
Oxide entrapment
9.1 Characteristic flaw types in friction stir welds.
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Friction stir welding Tensile strength
Yield strength
450
22.5 396
350 Strength [MPa]
20 351 396
300
17.5
314
309
15
280
250 11.5
200
224
221
203
12.5
224
10
150
Elongation [%]
400
7.5
100
4.6
50
5
4.5
4.4
3.3
2.5
0
0
Base material
Defect free FSW
LOP 0.2
LOP 0.4
Oxides
9.2 The mechanical properties (tensile strength, yield strength and elongation) of friction stir welded aluminium alloy 6013-T6 for LOP and oxide entrapped welding flaws when compared against the parent alloy and a flaw-free friction stir welded joint. 140 IQF IQF at 100 000 cycles [MPa]
120 100 80 60 40 20 0
Base material
Defect free FSW
LOP 0.2
LOP 0.4
Oxides
9.3 Index of weld quality with reference to the fatigue properties of friction stir welded aluminium alloy 6013-T6 for LOP and oxide entrapped welding flaws when compared against the parent alloy and a flaw-free friction stir welded joint.
temperatures in excess of the solidus temperature of the material (temperature at which a substance melts, or begins to change from solid to liquid form). 2. Temperatures approaching the solidus but where heat loss from the direct
Effects and defects of friction stir welds
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deformation zone is sufficiently retarded so as to result in unwanted thermal softening of the workpieces leading to degradation of the mechanical properties of the joint. Several characteristic flaw types are seen to occur as a result of FSW under too hot a processing condition. The flaws that are generated under such processing conditions are visually identifiable through the surface appearance of the welded joint. In extreme cases the surface appears to contain blisters and or surface galling, as portrayed in the image titled “Surface galling”, Fig. 9.1. Furthermore, excessive heat generation can lead to thermal softening in the workpiece material beyond the boundary of the tool shoulder. Subsequently the tool shoulder, rather than actively participating as a means of material containment (and not just heat generation), i.e. upper boundary of the socalled extrusion zone (Arbegast 2003), participates in material expulsion in the form of excessive surface flash formation. In cases where too much thermal softening occurs and where FSW is conducted under active load rather than tool position control, the workpiece material directly below the tool shoulder will arrive at a point where it is no longer able to support the axial load placed upon it. Consequently this leads to a thinning of the workpiece material thickness, since under active load control the FSW machine attempts to maintain a pre-programmed force. When using fixed length pin tools this results in the pin making contact with the backing bar often causing rupture of material in the weld root and ultimately damaging both the pin and backing bar. Such a condition is depicted in the image titled “Excessive flash”, in Fig. 9.1. Not every friction stir welded material can, however, be identified visually as having been processed under “too hot” a welding condition. In the case of aluminium and its alloys the broad spectrum of chemical compositions applied to produce the various types of alloys means that there is substantial variation in thermo-physical properties between the alloys, e.g. solidus temperature and thermal conductivity (the property of material that determines its ability to conduct heat). For example, the higher the alloys solidus temperature and thermal conductivity, the greater the processing temperature that can be tolerated during processing. The higher rate of thermal conductivity ensures more rapid removal of heat from the direct deformation zone. Furthermore the presence or absence of transient local melting within an alloy limits heat generation, since this influences the interfacial conditions between the joining tool and workpiece material (stick or slip) and thereby controls the strain rate that the material experiences during processing (Gerlich et al. 2006, 2007). Consequently this is observed in the shape the weld nugget contained within the stir zone develops during processing. A weld nugget which forms under too hot a welding condition is referred to as having collapsed (Arbegast
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Friction stir welding
2008). Such nuggets have also been characterised through microstructure including melted eutectic formation (Zettler et al. 2006b, Gerlich et al. 2007) and through examination of mechanical properties, e.g. hardness and tensile strength loss of the joint. Colegrove et al. (2007) have demonstrated that the recrystallised region of the stir zone for an Al-Cu-Mg-Mn 2024 alloy can actually decrease rather than increase in size when tool rotation speed is sufficiently increased. This is presented in Fig. 9.4a. Zettler (2008) has also observed the phenomenon of a decrease in the size of the recrystallised region (weld nugget) within the stir zone for friction stir welded Al-Si-Mg-Cu-Mn 6013 alloy. Here tool rotation speeds of 3000 rpm were employed. Additionally heat loss was actively retarded from the stir zone via the use of an insulated backing bar. Such a weld is presented in Fig. 9.5. In Fig. 9.5 it can be observed that the recrystallised region for the 6013 alloy joint like that of the 2024 alloy, Fig. 9.4 comprises a much small portion of the entire stir zone, and in both cases the weld nugget appears distorted by the influence of the much larger thermo-mechanically affected zone (TMAZ) surrounding the nugget.
2 mm
(a) 1600rpm-200 mm/min-8kN
(b) 800rpm-200 mm/min-8kN
2 mm
9.4 Macrographs of stir zone weld nugget formation in an Al-CuMg-Mn 2024 alloy as seen transverse to weld travel direction. (a) hot weld produced using a tool rotation speed of 1600 rpm, a weld travel speed of 200 mm/min and an axial load of 8 kN. (b) optimal processing weld produced using a tool rotation speed of 800 rpm, a weld travel speed of 200 mm/min and an axial load of 8 kN, courtesy of Airbus Germany.
Effects and defects of friction stir welds
(a) 1600rpm-200 mm/min-8kN
(b) 800rpm-200 mm/min-8kN
251
2 mm
2 mm
9.5 Macrographs of stir zone weld nugget formation in an Al-CuMg-Mn 2024 alloy as seen from the top of the weld. (a) hot weld produced using a tool rotation speed of 1600 rpm, a weld travel speed of 200 mm/min and an axial load of 8 kN demonstrating small recrystallised zone restricted to bounds of tool pin with large TMAZ in both sides of the joint. (b) optimal processing weld produced using a tool rotation speed of 800 rpm, a weld travel speed of 200 mm/min and an axial load of 8 kN with much larger recrystallised zone and smaller TMAZ, courtesy of Airbus Germany.
It is the opinion of this author that although thermal softening brought about by very hot processing conditions can lead to slip between the tool pin and the workpiece material, and thus decrease strain rates within the immediate vicinity of the tool pin, the tool shoulder still contributes considerable torsion to material directly below it. This torsion, although not sufficient to activate recrystallisation in the material beneath it, is the reason for the appearance of the TMAZ and also why the weld nugget appears to have collapsed near the surface of the joint. It should be noted that in the cases presented here, including the image in Fig. 9.1 titled “Nugget collapse”, this collapse is always more pronounced in the retreating side of the stir zone indicative of the direction of tool rotation and material flow taking place behind the tool pin. Although there is no way at present of validating actual quantities of
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Friction stir welding
material coming from each side of the joint being transported around the tool pin during processing, (and it has generally been assumed that material flow under hot welding conditions occurs as a single cylindrical body) Zettler et al. (2004) have demonstrated that processing parameters (tool rotation and travel speed) in conjunction with tool pin form can and do have a significant influence on the orientation of the grains within the TMAZ, i.e. vertical displacement, and that this displacement is more pronounced under hot welding conditions. The relevance of the above observation should be clear to those working in the field of corrosion assessment and prevention. This is because vertical re-orientation of underlying elongated and pancake-like grains, particularly for the case of high strength Cu and Zn rich wrought aluminium alloys increases the number of potential intergranular attack sites per unit volume if and when these grains become exposed at the workpiece surfaces. Figure 9.6 highlights how stress corrosion cracking is initiated in the surface area adjacent to the interface where the grain flow is turning. The likelihood that grain re-orientation within the workpiece occurs during FSW is not only seen to increase with increasing process temperature, but also as a result
200 microns
9.6 Micrograph transverse to weld travel direction demonstrating small secondary stress corrosion cracks initiated adjacent to the weld nugget where the grain flow in the TMAZ turns vertically towards the surface, courtesy of Airbus Germany.
Effects and defects of friction stir welds
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of using tool pin forms which strongly encourage vertical displacement of the thermally softened material within the stir zone. This phenomenon is not just associated with friction stir butt welds but has also been observed for the case of overlap (Arbegast 2008) and friction stir spot welds where upward displacement of thermally softened material exacerbates the so-called “hook-in effect”, which leads to substantial weakening of the joint in these locations. Although the reorientation of grains within the workpieces cannot be regarded as a linear or volumetric flaw as defined in Section 9.1, the negative influence this microstructural change has on the friction stir welded joint, i.e. the increase in stress corrosion susceptibility, challenges the original definition of what should be included as a flaw or defect. Perhaps the present definition should also include microstructure? A further challenge to the definition regarding FSW flaws pertains to the formation of the bands, otherwise referred to as onion ring structures produced in the weld nugget of friction stir welds. Krishnan (2002) states that Threadgill (1999) correctly guessed onion ring formation within the friction stir weld nugget was associated with the forward motion of the tool in one revolution. The predominant thought at the time was that although onion rings were notable features in friction stir welded aluminium alloys, they had no practical significance on the properties of the weld nugget. This assumption has more recently proven not to be the case however. Fatigue properties and crack path formation in friction stir welds have typically been analysed in response to processing parameters and their association to welding flaws such as root and nugget flaws (Dickerson and Przydatek 2003, Zhou et al. 2006). More recently, however, fatigue testing has been conducted to relate mechanical performance to processing temperature and weld energy (Lombard et al. 2008). In the case of hot friction stir welds it has been documented that the spacing (width of alternating bands as viewed perpendicular to the direction of welding) increases since more revolutions of the tool and more material transport occur per measure of weld length (Zettler et al. 2004). Fatigue test specimens examined by Lombard et al. (2008) demonstrated that the higher energy (hotter) welds tended to fracture in the weld nugget along the bands. This was also observed in the EMFASIS project (2005), while specimens produced under process-optimised conditions fractured in the parent material. More and more evidence suggests that increasing process temperature can and does have a substantial influence on the formation and subsequent role that the bands play in crack path formation in the weld nugget when placed under a cyclic load. The hotter welds are seen to lead to differences in grain size/shape and inter-metallic particle density within the bands. Furthermore these changes are observed in relation to crack path behaviour within the
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nugget (Sutton et al. 2002, Booth et al. 2007). It has been demonstrated, for example, that under fatigue loading friction stir welded Al-Cu-Mg-Mn 2024 alloy will develop fatigue crack failure paths which correspond to the bands having lower inter-metallic particle density (Sutton et al. 2002 and Booth et al. 2007). Although the actual forming mechanisms of the bands in aluminium alloy friction stir welds remains a hotly debated topic, there is significant evidence to suggest that joint formation and not just spacing between bands is a function of material flow activated by more than just a single shear layer (James et al. 2004, Booth et al. 2007, Zettler 2008). This also explains why the layers within the friction stir weld stir zone etch differently, since these layers experience different thermo-mechanical histories. Furthermore, these differences also present themselves not just as a consequence of temperature and process parameters, but also as a result of tool pin form. This is demonstrated in Fig. 9.7 where both welds were produced under identical processing conditions but using different tool pins.
B
Advancing side
2
C
Retreating side
9.7 Macrographs of the stir zones produced in an Al-Cu-Mg-Mn 2024 alloy under identical processing conditions with the exception of tool pin form when viewed perpendicular to the direction of tool travel (top) and transverse to weld travel below (note macrographs in the top view are seen from the direction as indicated by the arrows). These welds clearly demonstrate not just that the layers within the weld nugget etch differently but that they are significantly different as a result of tool pin form, courtesy of PhD RZ.
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The recommendation to avoid producing too hot a friction stir weld is before welding one should analyse the material’s thermo-physical as well as mechanical properties and then relate these to processing temperature and processing rates. Additionally one should consider the capacity of the FSW environment, e.g. machine, tool spindle structure and material restraint system and its ability to absorb and facilitate heat flow (losses) from the direct deformation zone. In the case of aluminium and its alloys, forgeability and extrudability have been observed to improve formability as a consequence of increasing process temperature. Considerable variation, however, exists in the role temperature plays between the various aluminium alloys. For example, temperature has the greatest effect in reducing flow stress for high silicon content 6xxx series alloys, (which typically possess a relatively high solidus temperature and thermal conductivity compared to 2xxx, 5xxx and 7xxx alloys), but has the least effect on high strength Al-Zn-Mg-Cu 7xxx alloys. In the case of 2xxx and 7xxx series alloys, flow stress is actually observed to increase rather than decrease with increasing temperature (Zakharov 1995). This results in reduced material flow, a greater likelihood for hot tearing along the grains and severe degradation of the pre-existing microstructure. Consequently FSW parameters such as tool rotation and weld travel speeds suitable for processing 2xxx and 7xxx alloys are considerably different (lower) than those which can be applied to 6xxx series alloys.
9.3
Defects from too cold welds
Too cold a friction stir weld can be defined by the processing conditions, e.g. tool rotation and weld travel speed are unable to generate sufficient heat or there exists too rapid a rate of heat loss from the immediate deformation zone to adequately support constant volume processing and/or temperature and time requirements to allow for bonding to occur at the joint. Naturally the processing conditions reflected in temperature are once again in reference to material thermo-physical properties such as solidus temperature, or if lower, the incipient melt temperature of second phase particles contained within the material. Logic would dictate that under too cold a processing condition one should expect a high degree of non-lubricated or dry slip between the tool pin and the workpiece material, since the lower thermal (specific) energy going into the production of the joint implies a higher degree of slip due to increased resistance of the workpiece material to deform. This slip should also manifest itself in reduced spindle motor torque as less material is transported (plastically deformed) by the FSW tool pin. This assumption however, is incorrect. The weld power required to overcome a material’s resistance to deformation under such conditions is in fact higher to the point
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where rupture of the material (machining) or rupture of the tool, as portrayed in Fig. 9.8 takes place. Unlike the case for hot working conditions, where the flow stress of a material is seen to reduce as a result of increasing process temperature, mechanically working a material under too cold a processing temperature will result in work hardening. This causes the density of dislocations in the work-hardened region to increase, making slip (one raft of atoms over a neighbouring raft) more difficult to occur. Macroscopically a material is said to become hard as a result of cold working and, unlike the case of flaw-free welds where tensile fracture has been shown to be dependent on the micro-hardness distribution within these regions, i.e. extent of thermal softening that occurs across the TMAZ and HAZ (Mahoney et al. 1998, von Strombeck et al. 1999, Svensson et al. 2000, James et al. 2004, Liu et al. 2004), welds produced under too cold a welding condition generally fracture (depending on the size of the flaw) through the defect (Liu et al. 2004). Several characteristic flaw types are seen to occur as a result of FSW under too cold a processing condition. The visually discernable flaw types include a lack of surface fill (open voids) and tunnel defects as presented in the images bearing these titles in Fig. 9.1. For the case of the large tunnel defect it is sometimes possible to observe this flaw forming by paying close attention to the surface appearance of the weld. Although in this instance there is no surface breaking void, i.e. “Lack of fill”, Fig. 9.1, there generally exists a region or linear demarcation that is visible separating the tooling marks left by passage of the FSW tool shoulder on the surface of the weld, particularly in the advancing side of the joint. The tunnel or worm hole defect can, however, easily be confirmed by looking directly into the exit hole after retraction of the tool pin from the workpiece. It is well recognised that by reducing the pressure applied to the surface of the joint by the FSW tool shoulder, this will lead to a reduction in processing
5 mm
9.8 Macrograph of the stir zone as viewed transverse to the weld travel direction containing fractured FSW tool pin embedded in AlCu-Mg-Mn 2024 alloy as a consequence of using a too cold welding parameter for this alloy, courtesy of PhD RZ.
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temperature and is often accompanied by swelling of the workpiece through thickness. This swelling has been associated with the formation of volumetric defects, i.e. both surface breaking and subsurface (Zettler et al. 2006a). Figure 9.9, however, demonstrates that there are instances when tunnel defects occur even without swelling of the joint. Most flaws types that develop due to too cold a welding condition, usually produce coarse and visually identifiable flaws. One friction stir weld flaw, however, that is not so easily detected by the naked eye is the so called kissing or cold lap bonding flaw. This flaw type is, however, very much evident when the material is mechanically loaded, for example through bend and or tensile tests. Poor bonding of the joint can often be attributed to inadequate pressure placed onto the workpiece material either due to poor tool pin design, e.g. the use of non-profiled cylindrical pins when possessing materials with a high resistance to deformation, or due to a lack of axial force placed onto the surface of the workpieces by tool shoulder position or inadequate axial load. In order to prevent such cold weld flaws from forming it is again important to provide sufficient temperature and time so that bonding during FSW can take place. A conflict often arises, however, particularly when the goal of the joint is to achieve mechanical properties of the joint that approach those of the parent alloy or rates of productivity that should exceed those of competing joining processes. Here much more intense monitoring of the process and recognition of process instabilities is required. For this reason a great deal of attention has focused on trying to understand the relationships between process monitored signals (machine generated signals) and the quality of the joint.
2 mm
9.9 Macrograph of the stir zone as viewed transverse to the weld travel direction containing a large volumetric flaw in base of the weld nugget for the advancing side of the joint. Here it can be observed that there is no thinning of the joint and hence it can be regarded that the void was formed due to inadequate material flow (too cold a processing condition) to enable constant volume processing to occur, courtesy of Airbus Germany.
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From the literature it becomes apparent that most flow visualisation studies have concentrated their efforts in trying to determine flow trajectories that occur during FSW. Very few studies, however, have attempted to describe the nature of material detachment from the tool once this material has been transported from the front to the back of the tool pin. this point is extremely important, since knowledge of this mechanism, i.e. how often material is rotated around the pin before it is released, what quantities are coming from advancing and retreating sides, will not only give insight into weld consolidation with regard to FSW parameters but also how and why welding flaws such as surface lack of fill, wormhole, or lack of consolidation defects, come about. chen and cui (2008) have shown for the work hardening Al-Mg-Mn 5083 aluminium alloy that a “laminar” or layered flow of thin sheets of material flows down and behind the pin into an area, which they state is vacated by forward motion of the FSW tool as it travels along the joint. the likelihood that a small vacant region at the rear of the tool exists when FSW under too cold a processing temperature is indirectly supported through examination of material flow patterns associated with defect formation and their correlation with fluctuations seen to occur in the magnitude and direction of processing forces (X-, Y-, and Z- axis), as well as spindle motor torque during FSW (Johnson 2001, Reynolds and tang 2001, colligan et al. 2003, James et al. 2004, Arbegast 2005, Colligan 2007, Yan et al. 2007). if one omits for the moment the axial or Z-axis force, then the magnitude and direction of a resultant force (R) through evaluation of Y-weld direction and X-transverse to weld direction forces can be calculated by equations 9.1 and 9.2 respectively. R = (FY )2 + (FX )2
9.1
q = arc tan (FY/FX)
9.2
the reason for differences in the angle q, i.e. orientation of the resultant force R can be explained in terms of changes occurring for interfacial conditions between material ahead of and behind the FSW tool. this is demonstrated schematically in Fig. 9.10. As material contact at the rear of the pin diminishes, the resistive force to rotation Fr2, seen in Fig. 9.10 begins to approach zero, thus reducing the effective force FX opposing sideways movement of the FSW tool pin. It is speculated that as material flow becomes more unstable, i.e. when void formation begins the imbalance between forces opposing tool rotation and travel, both ahead of and behind the FSW tool become greater, causing sideways movement of the tool to increases in magnitude. Arbegast (2005) likewise notes that reactionary forces opposing Z-axis force and spindle motor torque also increase with increasing weld travel speed, i.e. colder welds. the recommendation in response to avoiding
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FWeld results from material resistance opposing transport around pin
Pin
Fx = Fr1 – Fr2
Fr1
w VWeld = Weld travel speed
Fr2 0 Fy = FWeld
9.10 Schematic depicting FSW forces which potentially develop around the FSW tool pin and it is speculated may also be responsible for flaw formation, i.e. volumetric defects occurring when FSW under too cold a welding condition, courtesy of PhD RZ.
the production of FSW flaws due to too cold a weld is again to be aware of processing temperature relationships for a given material, i.e. in reference to the materials thermo-physical properties but also to monitor process generated forces, particularly machine signals which provide data such as the position of the tool in reference to the join line and to the surface of the workpieces. The monitoring of such process signals has the potential to provide an early warning system for the onset of too cold a weld FSW flaws and in time this will help to develop strategies, i.e. intelligent reactionary systems to take measures to correct for these flaws occurring.
9.4
Defects from geometrical mistakes
Geometric flaws as opposed to flow-related flaws are defined in response to producing friction stir welds through improper tool placement/adjustment with regard to the geometrical features of the welded joint. These are therefore machine operator induced errors which lead to flaw formation. Such features can include not adjusting the welding parameters to compensate for workpiece gap, thickness mismatch or plate thickness variation, or welding and allowing for the FSW tool to develop an offset to the join line and or the use of an incorrect tool pin length in relation to workpiece thickness, resulting in a lack of penetration (LOP) defect. In the last two instances, i.e. tool offset and LOP facilitate only partial bonding of the joint, i.e. the bond is not achieved for the entire through thickness of the workpieces, as
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Friction stir welding
highlighted in the images titled “Lack of penetration” and “Lack of fusion” in Fig. 9.1. Christner and Sylva (1996) were one of the first to investigate and report on the effect the presence of a joint gap had on joint formation and subsequent mechanical properties for this type of friction stir weld. Here a 6.4 mm-thick Al-Cu-Mn-Mg 2014-T6 aluminium alloy was friction stir welded with gaps up to 50% the workpiece thickness. In all cases gaps were maintained through insertion of shims, although the length of weld and the distance between shims was not presented. Christner and Sylva (1996) reported that a 2.3 mm or 36% of workpiece thickness gap could be tolerated without significant reduction in joint strength. A joint gap of 3.2 mm or 50% of the workpiece thickness, however, led to incomplete joint consolidation, and a volumetric flaw could be observed to form along the advancing side of the welded joint. Leonard and Lockyer (2003) provided further evidence to support the work of Christner and Sylva (1996) in that they also friction stir welded a 6.4 mm-thick 2014-T6 alloy having a workpiece gap of 2 mm or 33% the workpiece thickness. Their results similarly indicated for a weld length of approximately 260 mm that no discernable flaw formation formed, although mechanical testing was not undertaken to indicate what effect this gap had on joint strength. Although the studies by Christner and Sylva (1996) as well as Leonard and Lockyer (2003) demonstrate that significant gaps between workpieces can be tolerated when FSW, there is evidence to suggest that friction stir weld parameters were selected specifically to assist with joint formation, i.e. hotter welds with lower weld travel speeds were used, since these help to facilitate consolidation of the joint. This is stated because in the same study by Leonard and Lockyer (2003) they demonstrate for the case of momentary plate separation that this led to volumetric flaw formation in an otherwise flaw-free welded joint. One of the early misconceptions concerning FSW has been that temperatureassisted extrusion of the workpiece material alone is capable of consolidating joint formation, and that the interaction of the FSW tool with the material only plays a secondary role, i.e. so as to assist the extrusion process (Seidel and Reynolds 2001). The reality, however, is somewhat different, as can be seen in Fig. 9.11. Here it is possible to see that without tool pin/workpiece interaction, the workpiece material in the gap below the weld nugget does so without necessarily producing a bond in this region of the joint. In this instance the deliberate LOP weld produced by the author, with a 0.5 mm or 12.5% gap in relation to workpiece thickness demonstrates that material from the stir zone is effectively lost from the workpiece surface and extruded into the gap. This loss of material results in a thinning of the workpiece through thickness. Additionally there remains a clearly identifiable join line below the weld nugget.
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2 mm
9.11 Macrograph of the stir zone as viewed transverse to the weld travel direction for a deliberate LOP weld with 0.5 mm workpiece gap in the joint. The weld is characterised by a thinning of the workpieces as a result of material loss (extrusion of workpiece material) into the gap. Further it could be seen that no bonding between workpieces had taken place in this region below the weld nugget, courtesy of Airbus Germany.
Further experiments conducted by this author demonstrate that regardless of whether a cavity exists at the weld root between workpieces or in the backing bar, e.g. a chamfer between workpieces or an equivalent cavity machined into the backing bar, such momentary geometric changes in processing conditions have the potential to result in a region of non-bonding between workpieces for these locations. This is further highlighted in Fig. 9.12. Here a 0.5 mm chamfer (both in terms of width and height) was machined in the underside of the workpieces at their interface so that during FSW material from the stir zone (nugget) would be forced to extrude into this region during processing. Thereafter a root bend test was performed to stress this particular zone of the weld. The result as appear in image (b) of Fig. 9.12 demonstrate that the bend test caused the joint to fracture not just limited to the region of the chamfer but well into the weld nugget. This was not the case, however, for an equivalent weld conducted using the same welding parameters but minus the chamfer, as can be seen in Fig. 9.13. Although poor joint fit up between workpieces should be avoided when FSW, it should also be noted that there are in fact instances when FSW actively benefits from having a small gap between workpieces. In the author’s own experience this holds true when friction stir butt welding thin plate material, i.e. workpieces with a wall thickness of 2 mm and below. Typically having no gap in these instances, especially at the end of the weld length, and where the workpieces are forcibly restrained via clamping gives insufficient tolerance for any thermal expansion (distortion) to take place during processing. As a consequence the pressure exerted by the expanding plates, since material is forced to flow from front to back of the tool pin means that unequal thermal gradients between these two regions cause the plates to come together at the end of the workpiece length. The weld behind the tool naturally prevents
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(a) Prior to bend test
(b) Fracture as result of bend test
5 mm
2 mm
9.12 Macrographs of the stir zone as viewed transverse to the weld travel direction produced in an Al-Li alloy. (a) friction stir welded joint demonstrating that the original 0.5 mm ¥ 0.5 mm chamfer produced for both sides of the workpieces (in the region of the weld root) is no longer visible after welding. (b) the result of a bend test performed on the same specimen image a) highlighting that fracture is not simply limited to the initial region of the chamfer between workpieces, courtesy of Airbus Germany.
this from occurring for this region so plate distortion takes place ahead of the FSW tool. By restraining the plates or not allowing for a gap at the weld end this causes one or both plates to lift vertically away from the backing bar in the near vicinity of the FSW tool. Once this occurs the tool ploughs through the raised material and ruptures the workpieces. Naturally the width of gap that can be tolerated between workpieces when FSW will be influenced by material thickness but additionally material thermo-physical properties such as thermal conductivity will also play an important role. The gap after all is an active barrier to heat transfer between workpieces, i.e. it increases thermal resistivity. This is an even more important issue when it comes to the FSW of dissimilar materials.
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2 mm
9.13 Macrograph of the stir zone after bend testing for the joint produced under identical conditions (including material), as for Fig. 9.12; however, with the absence of a chamfer in the region of the weld root, courtesy of Airbus Germany.
A similar case as with workpiece gap can be made for the position of the FSW tool in relation to the join line. In this instance, however, not only will the position of the workpieces in relation to the joint play an important role but also issues including the stiffness of the FSW machine and the geometry of the profiles at the interface between workpieces. For example, a less stiff machine, such as a robot arm carrying a friction stir weld spindle, will be more prone to deflect from the programmed weld path due to an absence of uniform stiffness. This stiffness varies the further the arm is moved from its central location. And since increasing weld travel speed leads to cooler welds thus increasing force opposing the path of the FSW tool (both in the horizontal and vertical plane) this limits the potential size of the processing window that can be applied to a given material by use of such a machine. Here it can be expected that machine-related instabilities rather than flowinduced process parameter instabilities will lead to earlier flaw formation. The placement of the FSW tool with regard to the join line is also a factor which for less stiff FSW machines such as robotic arms has to be taken into consideration, particularly in relation to increasing weld travel speed. As indicated by Fig. 9.10, Section 9.3 of this chapter there is a tendency when increasing travel speed and producing a lower energy, i.e. cooler welds,
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that the tool attempts to move sideways from the programmed tool path. In such instances it is the findings of this author that joint strength (bonding) benefits by slightly shifting the position of the tool towards the retreating side of the joint line along the length of the weld. Here the natural rotation of the tool and subsequent processing forces bring the tool back towards the joint line during FSW. In the case of dissimilar alloy friction stir welds it is often the case that the position of the FSW tool is biased towards one side of the joint or that workpieces are placed in a given order, either advancing and or retreating side of the joint (McLean et al. 2003, Uzun et al. 2005, Zettler et al. 2006c, Cavaliere and Panella 2008). In the case of dissimilar aluminium alloy welds it has been demonstrated that mechanical strength of the joint is improved by placing the aluminium alloy with the lower strength, higher solidus temperature and thermal conductivity in the advancing side of the joint (Zettler et al. 2006c, Cavaliere and Panella 2008). A very different scenario is presented for the case of dissimilar material friction stir welds. The fact that material from both sides of the joint is forced to flow around the FSW tool pin in the direction of tool rotation makes it impossible to provide a flaw-free bond when the material in the advancing side begins to flow (deform) but the material in the retreating side resists deformation. This is why successful friction stir butt welds between steel and aluminium always have the steel placed in the advancing side of the joint (Uzun et al. 2005). Furthermore, this same principle can be applied to determining tool offset to the joint line. A tool that is placed too far into the advancing side, although it may provide sufficient energy to soften and deform material in this region of the joint, does not automatically mean that this is applicable to material in the retreating side, even when FSW a single alloy combination. The results of such a weld can be seen in Fig. 9.14. In this instance the author produced the weld with the tool pin deliberately positioned further into the advancing side of the joint (right-hand side) than would otherwise have been the case. From the experience of the author the FSW process is more tolerant to placement of the FSW tool into the retreating rather than the advancing side of the joint. Another any equally challenging flaw arising due to a geometrical mistake is incorrect setting of the pin length in relation to the workpiece thickness. This is particularly the case when using a fixed length pin tool where material through thickness varies more than 0.3 mm along the length of the weld. In most applications when FSW aluminium and its alloys up to 6 mm in wall thickness it is the experience of this author that a pin length of up to 0.3 mm smaller than the through thickness of the material being friction stir welded can be tolerated without producing an LOP flaw. When the gap becomes larger, however, the potential for non-bonding below the stir zone
Effects and defects of friction stir welds
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5 mm
9.14 Macrograph of the stir zone as viewed transverse to the weld travel direction where FSW was performed with the tool positioned almost completely in the advancing (right-hand) side of the joint. Here the fracture location is representative of the distorted and somewhat displaced original join line between workpieces and is witness that although the welding parameters were optimal for this alloy, incorrect tool placement to the join line is capable of leading to non-bonding between workpieces, courtesy of Airbus Germany.
increases. Furthermore these non-bonded regions act as crack initiation sites, particularly when the joint is cyclically loaded (EMFASIS). The recommendation to prevent flaws from occurring due to geometrical mistakes is quite simply to put into place weld procedure specifications that an operator has to follow prior to welding. These operating procedures must take into account observation of tool path and any corrections necessary along its length. Similarly, the material through thickness should be measured along the weld length and for both sides of the join line so that the length of the pin stick-out from the tool shoulder is correct and within tolerance for the material. If this is not the case then surfaces including the interface between workpieces should be machined prior to welding so as to avoid workpiece gap.
9.5
Features without significant effect
Aluminium-based material in contact with ambient atmosphere will form a natural surface layer consisting of Al-oxides and -hydroxides. Consequently, when aluminium is friction stir welded, this surface layer is disrupted but also stirred into the weld. However, as these surface layers are very thin, usually well below 1 mm, traces of them are difficult to detect even after careful examination. Cold friction stir welding parameters increase the probability to find traces of the former natural surface layer in the weld, which often appear as very thin snake-like lines frequently referred to in the literature as rootor weld nugget flaws. Here Sato (2005), Okamura (2002), Palm (2003) and Steiger (2005) have investigated root flaws of aluminium friction stir welds
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Friction stir welding
by electron microscopy. Vugrin et al. (2004) investigated systematically the influence of natural surface scales on FSW flaws. In this case, natural aluminium surface layers of varying thickness were created by an oxidation pre-treatment and these were then characterised prior to and after FSW.
9.5.1 Natural surface scales on aluminium and aluminium alloys Owing to its high affinity to oxygen, a fresh aluminium surface rapidly develops an amorphous Al2O3 layer having a thickness of a few nanometres. This dense and continuous surface film prevents further metal oxidation ensuring excellent oxidation resistance for aluminium at temperatures up to 400°C. At higher temperatures, however, the amorphous alumina scale transforms into g-alumina and leads to local flaws and cracks within this layer. Consequently, the corrosion resistance of aluminium is compromised at temperatures above 400°C. Detailed analyses of growth initiation and kinetics of the amorphous Al2O3 surface layers have been published by Cabrera (1948), Wefers (1987), Do (1997), Jeurgens (2002a), Jeurgens (2002b), Cochran (1961), Kirk (1968). If aluminium comes into contact with water vapour, a layer of pseudoboehmite, an alumininum hydroxide, forms on the aluminium surface. Layered hydroxide films are typically formed in humid air. According to Hart (1957) there is a sequence consisting of an amorphous hydroxite layer close to the metal surface followed by a zone of gelatinous pseudoboehmite with boehmite or bayerite crystals in the outer zone. Pseudoboehmite contains various amounts of water, depending on the temperature and humidity of the surrounding atmosphere. Baker (1976) found 27% water in pseudoboehmite layers which were formed in 100°C hot water. Pseudoboehmite may reach a maximum thickness of approx. 1 mm if it is exposed to boiling water for several hours. Detailed descriptions of the natural aluminium-hydroxide layer formed on top of the amorphous Al2O3 and its growth kinetics can be found, for example, in the following publications: Baker (1976), Wefers (1987), Altenpohl (1959, 1960), Ginsberg (1961), Geiculescu (2003), Scotto-Sheriff (1999) and Pereira Antunes (2002).
9.5.2 FSW experiments using pre-oxidized Al sheets with natural surface scales To shed light on the question, as to what extent aluminium surface scales influence microstructure and properties of friction stir welds, Vugrin et al. (2004) investigated a friction stir butt welded 6013-T4 alloy using workpieces which were pre-treated systematically to form relatively thick natural surface layers. In this instance, aluminium plates of 4 mm in thickness were
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mill-cut, cleaned with ethanol and subsequently stored under the following conditions: ∑
one month at room atmosphere (20°C, ca 60% humidity, Fig. 9.15 top) ∑ 1 h exposure to the vapour of boiling water (Fig. 9.15 bottom left) ∑ 48 h exposure to the vapour of boiling water (Fig. 9.15 bottom right). Transmission electron microscopy (TEM) micrographs contained in Fig 9.15 display the corresponding morphologies of the (hydr)oxide layers that were formed. The layer formed at room atmospheric conditions after one month had a thickness of 100–150 nm. The layers formed in the vapour of boiling water after 1 and 48 h had a thickness of 400–500 nm, and 500–800 nm,
Surface (hydr)oxidlayer
6013 T4 Al alloy 200nm
1 month room atmosphere Outer pseudoboehmite layer – feathery appearance
Inner pseudoboehmite layer – structureless appearance 500 nm 1h Water vapour
48h water vapour 200 Nm
6013 T4 Al alloy
9.15 TEM-images of the natural aluminium surface layers formed on 6013 T4-alloy. Top: layer formed after one month at ambient room atmosphere. Bottom left: layer formed after 1h exposure to vapour of boiling water. Bottom right: layer formed after 48h exposure to vapour of boiling water.
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Friction stir welding
respectively. The water vapour treated sample surfaces exhibited an inner zone of gelatinous pseudoboehmite followed by an outer zone possessing a feathery-like appearance. The pre-oxidised aluminium sheets were friction stir welded using laboratory-scale equipment while employing a standard cylindrical and threaded tool pin. The rotational speed used was 2000 rpm and tool travel speed was 1000 mm/min. Ultrasonic testing provided evidence of root and nugget flaws. Welding of sheets having relatively thin surface layers, as those grown at room atmosphere for one month, resulted in only a slight and faint root flaw. Furthermore, these welds were devoid of any evidence of a nugget flaw. In contrast, the thickest hydroxide scale welds observed for this study, grown during exposure to water vapour for 48 h, formed a very pronounced root- and nugget flaw inside the stir zone. In each case all root flaws reached a height of a few 100 mm before they could be seen to bend towards the retreating side of the welds (Fig. 9.16). Regions containing distinct flaws were prepared for subsequent microstructural analyses through scanning electron microscopy (SEM) and TEM. Careful preparation techniques were required to prevent artefacts. SEM sections remained un-etched to ensure the original appearance of voids. TEM foils were prepared by ion beam thinning under very mild conditions. The SEM investigations revealed the structure of the root- and nugget flaws in detail. Figures 9.17 and 9.18 demonstrate that the root flaw corresponds to a string of micron-sized voids. Though the pore series appears as a line in the metallographic sections, it should be observed that the root flaw has a threedimensional form. Hence the voids are arranged in a bend area which turns towards the retreating side of the weld at a height of a few hundred micrometres. The size of the voids is approximately the same in all welds examined in this study. Void density and connectivity, however, could be observed to vary. Obviously, density and connectivity of voids was high in case of thick preexisting surface scales grown under exposure of water vapour. In all welds the density and connectivity of the flaw-voids could be seen to be higher at the beginning of the root flaw, i.e. close to the bottom surface of the plates. This gradually decreased as the flaws progressed inside the weld nugget (as seen in Fig. 9.18). Completely undisturbed and well connected aluminium material could, however, be seen to exist between the voids of the root- and nugget flaws. In the case of pronounced root flaws, i.e. those formed as a consequence of the presence of a thick surface layer, some particles in the voids could be identified via SEM investigation (Fig. 9.17). EDS-measurements conducted on these particles revealed that the elements aluminium and oxygen were present giving a first indication that the particles originated from the former surface layers developed on the workpieces. This was later confirmed by
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rs
4 mm
as
200 µm 1 month room atmosphere rs
as
4 mm 300 µm 1h water vapour rs
4 mm
as
300 µm
48h water vapour
9.16 Optical micrographs of root- and nugget flaws formed by natural aluminium surface layers in 6013 T4-FSWs, cross sections. On the left side an overview of the welds is shown, on the right side the root flaw-region is shown at higher magnification (area marked with a white box). Top: FSW which plates were exposed to ambient room atmosphere for one month prior to welding. Only a root flaw was formed. Middle: FSW which plates were exposed to vapour of boiling water for 1h prior to welding. A root- and a nugget flaw were formed. Bottom: FSW which plates were exposed to vapour of boiling water for 48 h prior to welding. A root- and a nugget flaw were formed.
TEM investigations. TEM micrographs (Figs 9.19 and 9.20) represent rootand nugget flaws of FS-welds pre-oxidised in water vapour for 48 h, or 1 h, respectively. It turned out that particles occurring within the voids of the flaws had the same appearance as the former surface scales. Moreover, the
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200 µm
5 µm
(a)
(b)
9.17 SEM-images of a root flaw in 6013 T4-FSW pre-oxidised for 48h in water vapour, cross section. (a) Overview of the root flaw. The arrows indicate the course of the flaw. (b) Magnification of the area marked with a white box. The arrows indicate particles inside the holes of the root flaw.
(a)
50 µm
End of root flaw
(b)
5 µm (c)
Beginning of root flaw
5 µm
9.18 SEM-images of a root flaw in 6013 T4-FSW, pre-oxidised at room atmosphere over one month, cross section. (a) Overview over the faint root flaw. The arrows indicate the course of the flaw. (b+c) magnification of the areas indicated by the long arrows.
particle thickness corresponded well to the original thickness of the surface layer. Obviously, the surface layer was fragmented into particles. In the root flaw region of the welds, these particles were seen to undergo nearly no deformation by the welding process, Fig. 9.19. In contrast, the surface scale-particles found in the nugget flaw-voids did appear to undergo some deformation, Fig. 9.20. The TEM-investigations confirmed that the welding process had successfully been able to join the workpieces in that metal grains between the flaw-voids appeared sound and free of defects. Though the aluminium surface scales were seen to result sometimes in root- and nugget flaws when friction stir
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500 nm
1 µm (a)
(b)
9.19 TEM-bright field images of a root flaw in 6013 T4 FS-weld preoxidised for 48h in water vapour. The samples were taken close to the bottom of the welded plate at the beginning of the root flaw. (a) Overview of the root flaw. Some surface-scale particles are indicated by arrows, but many more can be observed. (b) Magnification of a root flaw-void filled with many particles.
1 µm (a)
500 nm (b)
9.20 TEM-bright field images of a nugget flaw in 6013 T4 FS-weld pre-oxidised for 1h in water vapour. (a) overview of the nugget flaw. (b) Magnification of the area marked with a black box. Nugget-flawvoids filled with surface-layer particles.
welded, it should be emphasised that these flaws are extremely small in comparison to the millimetre-sized pores typically occurring in aluminium fusion welds. In concluding Okamura (2002), Palm (2003) and Steiger (2005) have suggested that oxidation occurs during friction stir welding. The present
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study, however, yielded strong evidence that (hydr)oxide particles within the weld seam originated from pre-existing surface layers. The recommendation here is that if root- and nugget flaws are to be minimised in friction stir welds, the naturally grown (hydr)oxide layers should be removed as far as possible by a cleaning process in the region of the joint prior to welding. Furthermore the dispersion and disruption of such layers is aided by FSW using comparatively hot welding parameters.
9.6
Acknowledgement
The author Rudolf Zettler wishes to acknowledge and thank RIFTEC GmbH, Max-Planck-Straße 2, Geesthacht 21502, Germany (E-mail:
[email protected]) where he was an employee at the time of writing Chapters 3 and 9. Rudolf is currently employed as an innovation consultant and can be contacted at the WTSH – Business Development and Technology Transfer Corporation of Schleswig-Holstein, GITZ, Max-Plank-Strasse 2, Geesthacht 21502, Germany (E-mail:
[email protected]).
9.7
References and further reading
Altenpohl D.G., (1959), ‘Use of Boehmite Films for Corrosion Protection of Aluminum’, Corrosion – National Association of Corrosion Engineers, 18, 143–153. Altenpohl D.G., (1960), ‘Zur Frage des Auftretens von Böhmit bei der Reaktion zwischen Aluminium und kochendem oder überhitztem Wasser’, Aluminium, 8 (36th year), 438–441. Arbegast J.A., Modeling Friction Stir Joining as a Metalworking Process, Hot Deformation of Aluminum Alloys III, Z. Jin, ed., TMS (The Minerals, Metals, and Materials Society), 2003. Arbegast J.A., Using process forces as a statistical process control tool for friction stir welds, Friction Stir Welding and Processing III, Edited by K.V. Jata et al., TMS (The Minerals, Metals, and Materials Society), 2005. Arbegast J.A., A flow-partitioned deformation zone model for defect formation during friction stir welding, Scripta Materialia 58, 2008, 372–376, doi: 10.1016/j. scriptamat.2007.10.031. Backofen W.A., Deformation Processing, Addison-Wesley Publishing, 1972, pp. 88–115. Baker B.R. and Balser J.D., (1976), ‘Formation of films on aluminum by reaction with water’, Aluminium, 3 (52nd year), 197–200. Booth D.P.P., Starink M.J. and Sinclair I., Analysis of local microstructure and hardness of 13 mm gauge 2024-T351 AA friction stir welds. Materials Science and Technology, 23(3), (2007), 276–284(9); doi: 10.1179/174328407X157290. Cabrera N. and Mott N.F., (1948), ‘Theory of the oxidation of metals’, Reports on Progress in Physics, 12, 163–184. Cavaliere P. and Panella F.: Effect of tool position on the fatigue properties of dissimilar 2024-7075 sheets joined by friction stir welding. Journal of Material Processing Technology 2006, (2008), 249–255; doi: 10.1016/j.jmatprotec.2007.12.036.
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Chen Z.W. and Cui S., On the forming mechanism of banded structures in aluminium alloy friction stir welds, Scripta Materialia 58, 2008, 417–420. doi:10.1016/j. scriptamat.2007.10.026. Chong P.H., Liu Z., Skeldon P. and Thompson G.E., Corrosion behaviour of laser surface melted aluminium alloy in the T6 and T451 tempers. The Journal of Corrosion Science and Engineering, 6 (2003), 12. http://www.jcse.org/volume6/paper12/v6p12.php. Christner B.K. and Sylva G.D.: Friction stir welding development for aerospace applications. Proc. Conf. ICAWT 96, 6–8 November 1996, Columbus OH, USA, 359–368. Cochran C.N. and Sleppy W.C., (1961), ‘Oxidation of High-Purity Aluminum and 5052 Aluminum-Magnesium Alloy at Elevated Temperatures’, Journal of The Electrochemical Society, 108, 322–327. Colegrove P.A., Shercliff H.R. and Threadgill P.L., “Modeling and Development of the Trivex™ Friction Stir Welding Tool”, Proceedings of the 4th International Conference on Friction Stir Welding, The Welding Institute (TWI), Park City, Utah, USA, 14–16 May 2003. Colegrove P.A., Shercliff H.R. and Zettler R., Model for predicting heat generation and temperature in friction stir welding from the material properties, Science and Technology of Welding and Joining, vol. 12, 4, 2007, 284–297. doi: 10.1179/174329307X197539. Colligan K.J., Friction Stir Welding and Processing IV TMS 2007, 39–54. Colligan K.J., Xu J. and Pickens J.R., Proc. Friction Stir Welding and Processsing II, TMS, 2003, 181–190. Dawes C.J., Threadgill P.L., Spurgin E.J.R. and Staines D.G., Development of the new FSW technique for welding aluminium – phase II, TWI Research Report 5651/35/95. Dickerson T.L. and Przydatek J.; Fatigue of friction stir welds in aluminium alloys that contain root flaws. International Journal of Fatigue 25 (2003) 1399–1409: doi:10.1016/ S0142-1123(03)00060-4. Do T., Splinter S.J., Chen C. and Mc Intyre N.S., (1997), ‘The oxidaton kinetics of Mg and Al surfaces studied by AES and XPS’, Surface Science, 387, 192–198. EMFASIS: Effect of Material Flow Patterns on the Properties of Friction Stir Welds in Aluminium Alloys for Aircraft Structures. Patentschaft Project between the GKSS Forschungszentrum and Airbus Deutschland 2002–2005. Forborg B., Halem H. and Marthinsen K., The effect of Sc on the extrudability and recrystallisation resistance of Al-Mn-Zr-alloys. Materials Science Forum 467–470, (2004), 369–374. Gao N., Davin L., Wang S., Cerezo A. and Starink M.J., Mater. Sci. Forum, 396–402, 2002, 923–28. Geiculsescu A.C. and Strange T.F., (2003), ‘A microstructural investigation of low temperature crystallinge alumina films grown on aluminum’, Thin solid films, 426, 160–171. Gerlich A., Avramovic-Cingara G. and North T.H., Stir Zone Microstructure and Strain Rate during Al 7075-T6 Friction Stir Spot Welding Metallurgical and Materials Transactions A, 37A, 2006, 2773–2786. Gerlich A., Yamamoto M. and North T.H., Local melting and cracking in Al 7075-T6 and Al 2024-T3 friction stir spot welds. Science and Technology of Welding and Joining, 12 (6), 2007, 472–480: doi 10.1179/174329307X213873. Ginsberg H. and Wefers K., (1961), ‘Zur Struktur und Morphologie der Hydroxidschichten auf Aluminiumoberflächen’, Aluminium, 1 (37th year), 19–28. Hart R.K., (1957), ‘The Formation of Films on Aluminium immersed in Water’, Transactions of the Faraday Society, 53, 1020–1027
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James M.N., Bradley G.R.H., Lombard H. and Hattingh D.G., The relationship between process mechanisms and crack paths in friction stir welded 5083-H321 and 5383H321 aluminium alloys. Fatigue Fract. Engng. Mater. Struct. 28, 2004, 245–256. doi: 10.1111/j.1460-2695.2004.00830. Jeurgens L.P.H., Sloof W.G., Tichelaar F.D. and Mittemeijer E.J., (2002a), ‘Structure and morphology of aluminium-oxide films formed by thermal oxidation of aluminium’, Thin Solid Films, 418, 89–101. Jeurgens L.P.H., Sloof W.G., Tichelaar F.D. and Mittemeijer E.J., (2002b), ‘Growth kinetics and mechanisms of aluminium-oxide films formed by thermal oxidation of aluminum’ Journal of Applied Physics, 92, 1649–1656. Jo H-H., Cho H., Lee K.-W. and Kim Y.-J., Extrudability improvement and energy consumption in aluminium extrusion process of a 7003 alloy. Journal of Materials Processing Technology 130–131 (2002) 407–410. Johnson R., Proc. 3rd Int. Symposium on Friction Stir Welding, Kobe, Japan, 27–28 September 2001. Kirk C.T. and Huber E.E. Jr., (1968), ‘The oxidation of aluminum films in low-pressure oxygen atmospheres’, Surface Science, 9, 217–245. Krishnan K.N., On the formation of onion rings in friction stir welds, Materials Science and Engineering A327, 2002, 246–251. doi:10.1016/S0921-5093(01)01474-5. Leonard A.J. and Lockyer S.A., Flaws in Friction Stir Welds, The 4th International Symposium on Friction Stir Welding, Park City, Utah, USA, 14–16 May 2003. Liu H., Fujii H., Maeda M. and Nogi K., Tensile fracture location characterizations of friction stir welded joints of different aluminum alloys. J. Mater. Sci. Technol., 20/1, (2004), 103–105. Lombard H., Hattingh D.G., Steuwer A. and James M.N., Optimising FSW process parameters to minimise defects and maximise fatigue life in 5083-H321 aluminium alloy. Engineering Fracture Mechanics 75 (2008) 341–354: doi:10.1016/j. engfracmech.2007.01.026. Mahoney M.W., Rhodes C.G., Flintoff J.G., Spurling R.A. and Bingel W.H., Properties of friction-stir-welded 7075-T651 aluminum, Metallurgical Material Transactions A, 29A, 1998, 1955–1964, doi: 10.1007/s11661-998-0021-5. McLean A.A., Powell G.L.F., Brown I.H. and Linton V.M., Science and Technology of Welding & Joining, 8 (6), 2003, 462–464(3). Midling O.T. and Rovik G., Effect of Tool Shoulder Material on Heat Input During Friction Stir Welding. The 1st International Symposium on Friction Stir Welding, Thousand Oaks, Cal., USA, June 1999. Okamura H., Aota K., Sakamoto M., Ezumi M. and Ideuchi K., (2002), ‘Behaviour of oxides during friction stir welding of aluminium alloy and their effect on its mechanical properties’, Welding International, 16, 4, 266–275. Palm F., Steiger H. and Henneböhle U., (2003), ‘The origin of particle (oxide) traces in friction stir welds’ Proceedings of 4th International Symposium on Friction Stir Welding, Park City, Utah, USA, May 14–16, 2003, ISBN 1-903761-01-8. Payton L.N., Metal cutting theory and friction stir welding tool design. NASA Fellowship Program 2002: http://ntrs.nasa.gov/archive/nasa/casi.ntrs.nasa. gov/20030093619_2003101304.pdf Pereira Antunes M.L., de Souza Santos H. and de Souza Santos P., (2002), ‘Characterization of the aluminium hydroxide microcrystals formed in some alcohol-water solutions’, Materials Chemistry and Physics, 76, 243–249. Reynolds A.P. and Tang W., Friction Stir Welding and Processing TMS, 2001, 15–23.
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Sato Y.S. and Kokawa H., Friction stir welding (FSW) process, Welding International, vol. 17, 11, 2003, 852–855. Sato Y.S., Takauchi H., Park S.H.C. and Kokawa H., (2005), ‘Characteristics of the kissing-bond in friction stir welded Al alloy 1050’, Materials Science and Engineering A, A405, 333–338. Scotto-Sheriff S., Darque-Ceretti E., Plassart G. and Aucouturier M., (1999), ‘Physicochemical characterisation of native air-formed oxide films on Al-Mg alloys at low temperature. Influence of water’, Journal of Materials Science, 34, 5081–5088. Seidel T.U. and Reynolds A.P., Visualisation of the material flow in AA2195 friction stir welds using a marker insert technique. Metallurgical and Materials Trans. A, 32A(11), 2001, 2879–2884. Srinivasan P.B., Dietzel W., Zettler R., dos Santos J.F. and Sivan V., Stress corrosion cracking susceptibility of friction stir welded AA7075-AA6056 dissimilar joint, Materials Science and Engineering A 392 (1–2), 2005, 292–300. doi: 10.1016/j. msea.2004.09.065. Starink M.J., Wang S.C. and Sinclair I., 2005 TMS Annual Meeting held in San Francisco, California, February 13–17, 2005, in ‘Friction Stir Welding and Processing III’, Ed. Kumar V. Jata, M.W. Mahoney and R.S. Mishra TMS, Warrendale, PA, USA, 2005, 233–240. Steiger H., (2005), ‘Grundlagenuntersuchung über den Einfluss von oxidischen Deckschichten auf die Prozess-Sicherheit beim Reibrührschweißen von Al-Werkstoffen’, diploma thesis at the University of Applied Sciences in Munich, Germany, in cooperation with EADS corporate research centre Ottobrunn, Germany, March 2005. Sutiarso S., Lorimer G.W. and Parson N.C., The influence of copper and silicon additions on the extrudability and mechanical properties of a series of aluminium alloys based on AA6111. Materials Science Forum 396–402, (2002), 493–498. Sutton M.A., Yang B., Reynolds A.P. and Taylor R., Microstructural studies of friction stir welds in 2024-T3 aluminum, Material Science and Engineering A323, 2002, 160–166. doi:10.1016/S0921-5093(01)01358-2. Svensson L.E., Karlsson L., Larsson H., Karlsson B., Fazzini M. and Karlsson J., Sci. Technol. Weld. Join. 2000, 5, 285–296. Threadgill P.L., Friction Stir Welding – State of the art, TWI report-678/1999. Uzun H., Dalle Donne C., Argagnotto A., Ghidini T. and Gambaro C., Friction stir welding of dissimilar Al 6013-T4 to X5CrNil8-10 stainless steel, Materials and Design 26, 2005, 41–46. doi:10.1016/j.matdes.2004.04.002. von Strombeck A., dos Santos J.F., Torster F., Laureano P. and Kocak M., in Proc. 1st Int. Friction Stir Welding Symposium, California, USA, 1999. Vugrin T., Schmücker M. and Staniek G., (2004), ‘Root Flaws of Friction Stir Welds – An Electron Microscopy Study’, Friction Stir Welding and Processing III, Proceedings of 2005 TMS Annual Meeting, San Francisco, California, USA, February 13–17. 2005, ISBN 0-87339-584-0. Wefers K. and Misra C., (1987), Oxides and Hydroxides of Aluminum, Alcoa Laboratories. Yan J.H., Sutton M.A. and Reynolds A.P., Processing and banding in AA2524 and AA2024 friction stir welding. Science and Technology of Welding and Joining, 12/5, 2007, 390–401, doi: 10.1179/174329307X213639. Zakharov V.V., Scientific aspects of deformability of aluminium alloys during extrusion, Advanced Performance. Materials, 2, 1995, 51–66, doi: 10.1007/BF00711651. Zettler R., MSc Thesis, 2006, University of Adelaide, Australia.
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Zettler R., PhD Thesis, 2008, GKSS Forschungszentrum, Geesthacht, Germany. Zettler R., Lomolino S., dos Santos J.F., Donath T., Beckmann F., Lipman T. and Lohwasser D., A study on material flow in FSW of an AA 2024-T351 and AA6056-T4 alloys. 5th International FSW Symposium, Metz, France 2004. Zettler R., da Silva A.A.M., Rodrigues S., Blanco A. and dos Santos J.F., Dissimilar Al to Mg friction stir welds. Advanced Engineering Materials, 8 (5), 2006a 415–421: doi: 10.1002/adem.200600030. Zettler R., Donath T., dos Santos J.F., Beckmann F. and Lohwasser D. Validation of marker material flow in 4 mm thick friction stir welded Al 2024-T351 through computer micro-tomography and dedicated metallographic techniques, Advanced Engineering Materials 8 (6), 2006b, 487–490. doi: 10.1002/adem.200600062. Zettler R., Potomati F., dos Santos J.F. and de Alcantara N.G., Temperature evolution and mechanical properties of dissimilar friction stir weldments when joining AA2024 and AA7075 with an AA6056 alloy, Welding in the World, vol. 50 (11/12), 2006c, 107–116. Zhou C., Yang X. and Luan G., Effect of root flaws on the fatigue property of friction stir welds in 2024-T3 aluminum alloys. Materials Science and Engineering A 418 (2006) 155–160: doi:10.1016/j.msea.2005.11.042.
10
Modelling thermal properties in friction stir welding
H. N. B. Schmidt, Technical University of Denmark (DTU) and HBS Engineering, Denmark
Abstract: The thermal modelling chapter is a guide for the development of thermal models for friction stir welding (FSW). The effect of different welding parameters can be analyzed by thermal modelling thereby leading to a better process understanding. Analytical expressions for heat generation are presented in different variants dependent on the level of detail and purpose. Two types of heat generation mechanism are active during FSW: (i) frictional dissipation in the case of sliding and ii) plastic/viscous dissipation in the case of sticking. Each contribution is reflected in the expressions for heat generation. An analytical model for prescribing the material flow in the shear layer is presented, which can be used to account for the convective contribution due to the material flow around the tool. Considerations regarding enmeshment levels and choice of numerical framework (Lagrange/Eulerian) are presented together with a description of the governing expression used in combination with the different numerical solution schemes (static, steady-state, transient). Finally, the temperature-dependent heat source model is presented, which is also denoted the thermo-pseudo-mechanical (TPM) model, having temperature-dependent shear stress as the driver. Key words: friction stir welding, thermal modelling, heat generation, frictional dissipation, plastic dissipation, numerical model, FEM, CFD, TPM.
10.1
Introduction
A thermal model of FSW can be used to get a better process understanding. It can help the scientist or the technical staff in selecting welding parameters prior to performing the actual weld. A thermal model can help to give answers to “what if” scenarios. The heat flow in friction stir welding is complex in the sense that the success of the weld depends on the thermal conditions under which the weld is carried out. The main criterion for a successful weld is that the weld is fully consolidated, i.e. no presence of cavities, channels, voids or root defects. These imperfections in the weld can be related to the material flow around the tool, which itself is dependent on the thermal field. Under certain conditions, the shear layer could collapse leading to some of the above-mentioned failure mechanisms. Another criterion is related to tool 277
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wear/tool life and tool forces, which is reflected by the thermomechanical conditions at the tool/matrix interface. Cold welding conditions could lead to tool breakage as well as excessive tool forces, whereas hot welding conditions could lead to excessive tool wear. Yet another criterion could be the final joint properties which (under assumption of a fully consolidated joint) among others depend on the microstructure evolution in the weld zones. The microstructure evolution during the welding process of typical friction stir welded materials is highly temperature dependent. Especially for the heat treatable aluminium alloys – the thermal history (peak temperature and “holding time”) is important as it alters the hardness distribution (yield strength) and could lead to recrystallization. The non-uniform heat flow through the workpiece introduces thermal expansion and contraction leading to thermally induced plasticity – which introduces residual stresses and distortions in the final workpiece after welding. Finally, machine-specific criteria, e.g. maximum allowed forces/torque, machine stiffness, run-out, spindle deflection, could influence joint properties. A successful weld could be defined by a criterion for the “strength” of the joint, e.g. the weld region should have 80% of strength of the base material. The specific number depends of course on the thickness of the workpiece and material, hence for thin plates in 2024T3 a criterion might be 90%. To fulfill such criteria, the proper welding parameters (rotational and welding speed, plunge force/depth, etc.) should be carefully selected. The tool should be designed to give the desired material flow and heat generation under the constraints given by the welding machine. In the following the process window is defined by sets of welding parameters, e.g. rotational and welding speed, plunge force/depth tool leading to a fully consolidated joint. The process window could be further narrowed by including criteria from microstructure effects, residual stress or machine restrictions. One of the first questions arising is: how does friction stir welding work? The term heat generation is fundamental for characterizing FSW. Unlike other welding processes such as arc welding – the heat generation in FSW is non-controllable, since it is inherently built into the process itself. The heat generation can only be indirectly controlled via the choice of welding parameters. The mechanisms for heat generation are characterized by frictional and plastic dissipation, which in turn are coupled to the contact condition at the interface (degree of sticking/sliding) and the material flow field. Figure 10.1 shows a schematic view of the FSW setup where the identification of the different components is given. Figure 10.2 shows a simplified FSW tool with identification of the three main heat generation contributions from the FSW tool, i.e. the shoulder, tool probe side and probe tip heat generation contributions.
Modelling thermal properties in friction stir welding
279
Extraction Plunge Dwell Tool
Weld
Exit hole
Workpiece
Backing plate
10.1 Schematic of FSW setup with workpiece, tool and backing plate as well as the tool motion during the plunge, dwell, weld and extraction periods.
Qprobe side
Qshoulder
Qprobe tip
10.2 Heat generation contributions from the shoulder, probe side and probe tip.
280
10.2
Friction stir welding
Analytical model of heat generation
It is of great importance to have a phenomenological model that can estimate the amount of heat generation. Such a model is based on some constitutive behaviour for the force equilibrium between tool and workpiece – the most classical is Coulomb’s law of friction. The analytical model of the heat source accounts for some of the process parameters, which in its simple form is the rotational speed. However, classical analytical models do not include the effect of welding speed, which is an extremely important process parameter. Under the assumption of sliding, the plunge force (Z-force) can be used to estimate the pressure under the tool, which then can be used in a Coulomb type of friction law. One of the disadvantages is that some versions of these analytical models need experimentally measured input data, such as the Z-force. In the case of an instrumented weld, the torque for rotating the tool could also be monitored, thereby directly measuring the tool rotational power, which is close to the total heat generation by the tool. The analytical model accounts for the geometrical shape of the tool, and the classical approach is to include the main effect of the shoulder (given by the shoulder radius), the probe sides (given by the probe radius and height under the assumption of a cylindrical shape) and the probe tip (assuming that the probe tip is flat and not rounded, which is often the case in the real tool design). The heat is generated by the relative motion between the tool and the workpiece. There are two contributions to the heat generation: (i) frictional dissipation and (ii) plastic or viscous dissipation. The frictional dissipation is driven by the frictional stress field at the contact interface. But it is important to emphasize that even under a high degree of sliding, some material deformation must be present to accommodate the material flow around the tool probe. The plastic dissipation is driven by the material deformation occurring in the shear layer around the tool. In the following derivation of the heat generation expressions, it is not considered whether the contact condition is sliding or sticking – hence they are written in general form.
10.2.1 Analytical heat source The local heat generation at a surface segment on the tool/matrix interface is given by
q = twr
10.1
where tcontact is the shear stress located at the contact interface between the rotating tool and shear layer/matrix. By integrating the local heat generation over the shoulder contact area the heat generation contribution from the shoulder is estimated to
Modelling thermal properties in friction stir welding
Qshoulder = =
Ú Ú
Rshoulde shoulderr
Ú
R probe probe Rshoulder R probe
2p 0
281
t contact contact w r r dq dr
2 pt contact w r 2 ddrr
10.2
3 3 = 2 p w t conta Rshoulde cont cctt (R sshoulder houlderr – R pprobe ) 3
the heat generation from the probe side is given by
Ú = Ú
Q probe side =
H probe probe 0 H probbee 0
Ú
2p 0
t contact contact w r r dq dz
2 pt contact w r 2 ddz
10.3
2 = 2 p w t contactt R pr pprobe robe obe H probe
assuming that the probe is cylindrical – hence without any non-symmetrical features such as threads or flutes. the heat generation from the probe tip is estimated by
Ú = Ú
Q probe tip =
R probe probe 0 R probe 0
Ú
2p 0
t contact contact w r r dq dr
2 pt contact w r 2 ddr
10.4
3 = 2 p w t contactt R pprobe 3
the total heat generation is the sum of the three contributions i.e. Qtotal = Qshoulder + Qprobe side + Qprobe tip
10.5
which leads to Q = 2 p w t contact (Rs3houlderr + 3R p2robe H probe ) 3
10.6
For tools with a conical shoulder the increased contact area generates additional heat generation by the term tan a where a is the conical shoulder angle. The total heat generation is modified to Q = 2 p w t contact ((Rs3houlderr – R p3robe )(1 + tan a ) 3 3 2 + R pr obe + 3R probe H probe )
10.7
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For getting an impression of the heat generation quantity to be expected in friction stir welding, Table 10.1 shows the heat generation contributions from a typical tool having a shoulder radius of 9 mm, probe radius of 3 mm and probe height of 3 mm. A total heat generation of 1.4 kW is estimated for 3 mm thick panels of typical aluminium alloy where a yield stress of 20 MPa is assumed as descriptive for the material strength at elevated temperatures. Unfortunately, the model cannot directly be used to investigate the influence of the welding speed.
10.2.2 Contact condition The contact condition is first described in literature in Schmidt et al. (2004) as a contact state variable acting as a weighting function between the heat generated by frictional and plastic dissipation. the contact state variable d is also introduced as dimensionless sticking rate (the opposite of the dimensionless sliding rate) for describing the velocity field at the interface between the tool and the shear layer. The contact state variable is defined as
d=
vshearr llayer vshearr llayer = vtool wr
10.8
where d = 0 and d = 1 correspond to pure sliding and full sticking, respectively, whereas 0 < d < 1 describes a contact condition where both sliding and sticking occur, i.e. the partial sliding/sticking condition. the current understanding of the FSW process suggests that a partial sliding/sticking most likely is present, and more importantly, the contact condition is not uniform. the shoulder region has contact condition closer to sliding, whereas the probe contact region has a contact condition closer to sticking relative to that of the shoulder. these are speculations that are not supported experimentally in the current work. Assuming tangential flow only, the velocity component at the contact interface expressed in the cylindrical coordinate system is uq = dw r
10.9
which for the cartesian coordinate system yields Table 10.1 Heat generation contributions based on analytical expressions
Qshoulder Qprobe side Qprobe tip Qtotal
Heat generation (W)
Relative contribution (%)
1230 142 47 1419
87 10 3 100
Modelling thermal properties in friction stir welding
283
ux = – y w d uy = x w d
10.10
here given as a counter-clockwise rotational direction. the contact state variable acts as a weighting function between the heat generation due to plastic and frictional dissipation, i.e. the local heat generation can be defined as q = (d tplastic + (1 – d) tfriction) w r
10.11
where tplastic is the material yield shear stress and tfriction is the frictional shear stress. Assuming that the contact condition is uniform over the entire contact interface, i.e. d is constant, the total heat generation is Qtot = (1 – d)Qfriction + d Qplastic
10.12
which leads to the contact condition dependent total heat generation 3 3 Q = 2 p w ((Rshoulder r – R probe )(1 + tan a ) 3 3 2 + R pr obe + 3R probe H probe )(dt plastic + (1 – d ) t friction )
10.13
the heat generation by frictional dissipation is controlled by a frictional constitutive behaviour. the most common constitutive relation is the coulomb law of friction that simply relates the frictional shear stress to the pressure via a friction coefficient m, hence tfriction = mp
10.14
where p is the normal pressure at the contact interface. Equation (10.13) simplifies to 3 Q = 2 p w Rshoulder mp 3
10.15
Fz 2 and neglecting where Atool = p Rshoulder Atool the probe side heat generation, equation (10.14) can be substituted into equation (10.15) resulting in the well-known heat generation expression
By estimating the pressure by p =
Q = 2 w Rshoulde shoulderr m Fz 3
10.16
The friction coefficient m ranges from 0 to 0.5–0.6 where values around 0.3 are often reported; however, the exact values of m for different welding conditions and tool/workpiece combinations are difficult to measure
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Friction stir welding
experimentally. An additional disadvantage of this expression is that it is based on the experimentally measured value of the plunge force. This has to be measured via a dynamometer or similar load measuring device (strain gauge). The necessity for a dynamometer enables recording the torque required to rotate the tool at the same time. This torque multiplied by the rotational speed gives the total rotational power, Qtotal = Mw. Obtaining these experimental values, makes equation (10.16) useful for estimating the friction coefficient, since Qtotal and Fz are already known. Obtaining/measuring the total heat generation allows for estimating the local heat generation via equation (10.17) (see next section) under the assumption that the power for the linear motion of the tool is negligible. According to equation (10.16) the heat generation should be proportional to the plunge force; however, the experimental welds of 2024T3 in Schmidt et al. (2004) show no such proportionality. Figure 10.3 shows how the plunge force Fz gradually increases whereas the torque, and thereby the rotational power, at the same time decreases. This indicates that if some degree of sliding were present, the frictional dissipation is not proportional with the pressure – meaning that Coulomb friction might not have been descriptive Plunge force
60
Torque
Force [kN]/Torque [Nm]
50
40
30
20
10
0
–10
0
10
20 Time [s]
30
40
50
10.3 Measured plunge force and torque in a 2024T3 FSW. Notice the non-proportionality during the welding period.
Modelling thermal properties in friction stir welding
285
for these welding conditions. Figure 10.4 shows the micrograph of the longitudinal cross section through the exit hole of the FSW in 2024T3. The force data in Fig. 10.3 is from this specific weld (Schmidt et al., 2004). Notice the thin shear layer of approximate 0.5 mm thickness at the tool/ matrix interface. this suggest that the contact condition is different from pure sliding, i.e. more likely closer to sticking, which could explain the lack of pressure dependence. Analytical heat source models are often used as a driver for the thermal models of the heat flow in the workpiece. This means that Q is used as a fitting parameter adjusted until sufficient correlation between measured and simulated temperature profile are obtained.
10.2.3 The local heat generation the local heat generation is the main parameter of interest when developing a thermal model of FSW. Once the distribution of the local heat generation is established it can be prescribed as a boundary condition in a numerical model or used in an analytical model for estimating the total heat generation. Basic equation for the heat generation at the contact surface, i.e. q = tcontactwr, has been addressed in the previous section and most expressions for the total heat generation are based on this under different assumptions. the simplest expression for the local heat source is found by isolating tcontact in equation (10.6) and inserting into equation (10.1). This gives the following expression for the local distribution of the heat generation for known values of Q, q=
3Q r 3 2 p Rshoulder
10.17
Wshoulder
Wprobe side
10.4 Micrograph of the weld region showing the thin shear layer (~0.5 mm) located at the tool/matrix interface.
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Friction stir welding
which has been implemented in numerous models presented in literature, see Shi et al. (2003) and Schmidt and Hattel (2005a). For a model including a flat shoulder, cylindrical probe and flat probe tip, the local heat generation is given by q=
3Q r 2 p ((R Rs3houlder + 3 R 2 H p )
10.18
As an alternative, the heat source can be applied as an uniform volumetric heat source across a shear layer (having the width of wsl), using the following expression q=
3Q r 2 p Rs3houlder wsl
10.19
having units of È W3 ˘ resembling a uniformly distribution though the shear ÎÍm ˚˙ layer thickness.
10.2.4 Frictional and plastic dissipation the heat generation due to frictional dissipation can be described by qfriction = tfriction (1 – d) wr
10.20
which is a surface flux applied at a sliding contact interface. The frictional shear stress can be expressed by a constitutive expression which could be pressure dependent. An example of a pressure dependent friction law can be mentioned the Coulomb law of friction given by equation (10.14), which is commonly used. the local volumetric heat generation due to plastic dissipation is given by qviscous ÈÍ W3 ˘˙ = sij eij Îm ˚
10.21
acting in the shear layer, where sij = d ij – 1 d ij d ii is the deviatoric stress 3 tensor. Furthermore, it can be shown that sij = tij. One of the objectives of the thermal model is to obtain thermal results without making a fully coupled thermo-mechanical/cFd type of model. this means that the material flow field in the shear layer surrounding the tool is not directly known. the local viscous dissipation expressed as a volume flux in equation (10.21) can be “converted” to a surface flux at the contact interface by integrating the volume dissipation through the shear layer width w, i.e.
Modelling thermal properties in friction stir welding
qviscous È W2 ˘ = ÎÍm ˚˙
Ú
w
sij eij ddr
0
287
10.22
Assuming that the flow field concentrates in a thin shear layer surrounding the tool as illustrated in Fig. 10.5. By assuming that the rotational flow is dominant, it follows that only shear components are present, i.e. sij eij = srq erq + sq r eq r + szq e zq + sq z eq r = 2 srq erq + 2 szq e zq
10.23
since sii = 0 under the assumption of rotational flow only, and sij = sji. Assuming uniform temperature distribution through the shear layer (at a given r-position) and a strain rate independent shear stress, allows for a simple integration through the shear layer (having a constant t), which for the probe side gives probe side qviscous =
Ú
wr 0
2 t rq erq ddrr = 2 t rq erq wr
10.24
Shear layer
A
A
Exponential velocity profile
B
B
A
B
Linear velocity profile
10.5 Shear layer surrounding the tool. Lower pictures show the “conversion” of the exponential velocity profile in a real weld to the linear velocity profile used in the analytical shear layer models.
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Friction stir welding
where wr is the shear layer width at the probe sides measured in the r-direction. At the shoulder and probe tip shear layers, the surface flux can be found by shoulder / probe tip qviscous =
Ú
wz 0
2 t zq ezq ddz = 2 t zq e zq wz
10.25
where wz is the shear layer width measured in the z-direction. the strain rate erq can be found by assuming a linear shear layer as illustrated in Fig. 10.5, thus the strain rates for the probe side and shoulder shear layer, are
dw r erq = 1 2 wr
10.26
dw r ezq = 1 2 wz
10.27
respectively, which can be inserted into equations (10.24) and (10.25). The expression for the local plastic heat generation “converted” to a surface flux ÈW ˘ is then ÎÍm 2 ˚˙ qplastic = tplastic dwr
10.28
which can be implemented as a heat source in a numerical model or in an analytical heat source model, despite the flow field not being part of the actual model. Strictly speaking all the assumptions behind this estimate should be recalled when applying such a model. In cases where the velocity field through the shear layer is known, the temperature-dependent heat source can be implemented as a volumetric heat flux, based on the general expression in equation (10.21), which for assuming tangential flow, only, can be simplified to qviscous È W3 ˘ = sij eij = 2t rq erq + 2t zq e zq ÍÎm ˙˚
10.29
For the general case where tangential flow cannot be assumed, the Cartesian coordinate system can be used instead of the cylindrical coordinate system.
10.2.5 Temperature-dependent heat source the analytical expressions for heat generation are based on a uniform shear stress distribution at the contact interface. An implication of this is
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the assumption that the stress is independent on the temperature and the deformation/sliding rates. This is a major limitation of these models – but on the other hand – if the temperature dependence could be accounted for, the solution cannot be found analytically, only numerically. Recently, Schmidt and Hattel (2008a,b) presented a temperature-dependent heat source – the so-called thermal pseudo-mechanical heat source model (TPM). The TPM model is the first model of its kind (being a purely thermal model) to predict welding conditions for welding parameters that are not validated by experiments. the procedure is that once the model is validated for one or two welding parameters it can be used to explore and predict thermal results for other welding conditions. the basic expression for the temperature dependent heat source is qviscous = qplastic = tplastic (T) dwr
10.30
The material flow stress is a function of the strain e, the strain rate e and the temperature T due to the thermal softening of the material. A pure thermal model would not be able to account for more than the temperature dependence – which is for the first order effect of the softening. The temperature dependence of the yield flow stress can be described using a power law type of softening function similar to the one used in the Johnsoncook material law,
t (T ) = t ref re
p Ê Ê T –T ˆ ˆ ref ref Á1 – Á ˜ ˜ ÁË Ë Tsolidus – Tref ref ¯ ˜ ¯
10.31
where Tsolidus is the solidus temperature and p is a power law coefficient changing the temperature dependence (thermal softening). three different temperature-dependent yields stress vs. temperature curves are shown schematically in Fig. 10.6. The linear curve corresponds to having p = 1 and the two single curved plots are for p > 1 and p < 1, respectively, and the double curved plot represents a more complex temperature dependence, e.g. microstructural effect of heat treatable alloy, which is not described by this expression. measured values of the temperature dependent shear stress of friction stir processed 7075-T6 are shown in Fig. 10.7 for comparison, which are evaluated at a reference strain rate of 10–3, see cavaliere and Squillace (2005).
10.2.6 Equilibrium between friction and shear stress A very critical point in the development of the tPm model is the assumption that the shear stress in the region of the shear layer just below the contact interface is in equilibrium with frictional stress acting at the very contact interface. Behind the assumption is that the material flows in the tangential
Friction stir welding p Ê Ê T – Tre ˆ ˆ ref Á ˜ s y (T ) = s (Trref ef ) 1 – Á ˜ Á ËTmelt melt – Tre ref ¯ ˜ Ë ¯
sy
linear p = 1
sy (Tref)
p>1
Heat treatable alloy
p<1
Tmelt
Tref
10.6 Temperature-dependent yield stress described by a power law expression.
250 FSP 7075-T6 at de/dt = 10–3 200 Shear yield stress [MPa]
290
150
100
50
0
0
100
200 300 400 Temperature [°C]
500
600
10.7 Experimentally obtained shear stress data for friction stir processed 7075-T6 evaluated at a strain rate of 10–3.
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direction only (rotating in the same direction as the tool) but not necessarily with the tool velocity, and furthermore the flow field is assumed steady state, i.e. the velocity field in the shear layer does not change seen from the tool position and the acceleration forces are negligible. Figure 10.8 shows a schematic view of the deformation field in the shear layer including velocity profiles and stress components. The layer just at the interface rotates with ticking uqsticking = d w r and the velocity difference (slip rate) between the tool and r Notice that the velocity at the interface shear layer is uqslidingg = (1 – d ) w r.
uqtool = wr Tool shoulder segment Tool shoulder segment
z
t qfriction
· t zviscous = heff gz q q
q uqsl
uqsticking = dwr
uqsliding = (1 – d) wr
Shear layer · gz q
uqsl
uqsliding = (1 – d) wr Tool probe segment · t viscous = heff grq rq
uqsl
uqsl Shear layer uqtool = wr t qfriction
r
uqsticking = dwr
10.8 Velocity profiles, sliding velocities and shear stresses acting on different tool segments. Notice the equilibrium between friction and viscous shear stresses, where the latter are evaluated in the shear layer segment closest to the contact interface.
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Friction stir welding
as well as the slip rate are assumed independent of the welding speed uweld, which could be a critical assumption for high welding speed conditions and high degree of sticking. the following hypothesis states that the presence of the slightest degree of sticking will lead to a shear layer and hence an equilibrium between the frictional and plastic/viscous shear stress must prevail, i.e. tfriction = tplastic = tcontact
10.32
where tplastic = tviscous = heff g for a non-Newtonian viscous behavior. Notice that g should be evaluated just closest to the interface, and subscript rq and zq refer to the cylindrical coordinates at the probe side and shoulder shear layers, respectively. this assumption of equilibrium, makes the heat generations independent of the contact condition, which can be seen by inserting equation (10.32) into equation (10.11) qtotal = dtplastic + (1 – d) tfriction = dtcontact + (1 – d) tcontact = tcontact wr
10.33
which could indicate that the heat generation is contact state independent; however, this holds only for thermal models where the heat generation from plastic and frictional dissipation is modelled as a surface flux having all the assumptions in mind. in general, the contact condition is an important parameter which should be accounted for when modelling. As mentioned before, the contact condition controls the material flow as well as the characteristic of the heat generation (whether generated at the contact interface as a surface flux or inside the shear layer as a volume flux). But in most real FSW the contact conditions are not know in detail and are most likely non-uniform. To overcome this, this simplified contact state independent procedure can be implemented. to summarize, this involve assuming stress equilibrium, moving the viscous stresses to the contact interface and considering the heat source as a surface flux.
10.3
Numerical thermal model
thermal modelling of FSW is the basic modelling discipline since the need for controlling the heat flow is obvious when performing an actual weld. As mentioned previously the heat generation is not adjustable during the welding process – it is inherently built into the process itself and can only be indirectly controlled by adjusting the welding parameters, i.e. welding speed, rotational speed. From a modelling point of view, the thermal results are the driver for residual stress and microstructure models. Furthermore, the pure thermal models themselves are of great interest being a phenomenological
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process model explaining how the FSW works. the analytical expressions of the total heat generation have found their use as input parameters in thermal models. This conflicts in some cases with the purpose of the model itself – namely to estimate the heat generation and temperature fields. Thus, the objective is often to have the heat generation as an output of the model. this calls for performing temperature measurements using thermocouples or other temperature measurement systems. The challenge is then to find the total heat generation Q and the distribution of the local heat generation q by inverse modelling, such that the best fit with the thermal measurements is obtained.
10.3.1 Governing equations For a transient thermal model, the governing equation is the energy equation or heat conduction equation (with convective terms) expressed as
r c p T + ui r c p T, i – (kkT T, i ), i = Q
10.34
where k is the thermal conductivity, r is the density, ui the velocity vector components. the source term Q is a volumetric heat source which is due to viscous dissipation, e.g. in the shear layer surrounding the tool/workpiece interface. If the heat source is applied as a surface flux at the contact interface, the numerical implementation would then be via prescribing the heat flux as a boundary condition. By assuming steady state conditions, the time dependent term r c p T is neglected, thus reducing the energy equation to ui r c p T, i – (kkT T, i ),i = Q
10.35
which is the equation to be solved for when obtaining the steady state temperature field. The combination of an Eulerian framework and a steady state solution is a widely reported model setup in simulation of welding, also in the case of FSW. it should be emphasized that there are no restrictions in combining the Eulerian framework with the transient (time dependent) time scheme. in this case, the plunge and dwell periods, where the heat source is stationary, can be analyzed. the size of the Eulerian domain should be adjusted such that the temperature field is not influenced drastically by the inlet and outlet boundary conditions. Once the welding period begins the convective flow velocity is changed to the traverse welding speed, thereby simulating the linear motion of the tool. In a Lagrange model the mesh is fixed in space, hence there is no material flow through the mesh. A transient thermal model of FSW using the Lagrange framework is governed by the heat conduction equation
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Friction stir welding
r c p T – (kT, i ),i = Q
10.36
For stationary thermal problems T = 0, which leads to the stationary heat balance defined as – (k T, i ), i = Q
10.37
however, this latter expression would not be relevant for modelling FSW, since the tool is moving and the temperature field is transient. In thermal modelling of FSW, equation (10.37) could be used for finding the stationary temperature field in a component, which is moving with the tool/heat source, e.g. FSW machine head or motor. However, with the limitations in computational resources, these effects are often excluded from the calculation domain.
10.3.2 Local thermal model: Eulerian framework A local thermal model can be characterized by having the highly detailed heat source in order to capture the temperature field. Figure 10.9 shows a schematic setup of the local thermal model using the Eulerian framework. Eulerian local model
Stationary heat source
Lagrange global model
Jointline Moving heat source
10.9 Schematic of a local Eulerian model with a stationary heat source and a global Lagrange model with a moving heat source.
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The local heat source distribution is applied as a surface flux at a boundary of the model domain, which is stationary relative to the mesh. The material “flows” through the mesh – or more correctly, the convective contribution due to the material flow is taken into account via the convective term uircpT,i in equations (10.34) and (10.35). Such model framework is called Eulerian – and in some cases also denoted “a moving coordinate system”. By including the effect of having material entering the calculation domain at an inlet boundary and leaving the domain through an outlet boundary, the volumetric size of the enmeshed domain can be reduced. A numerical model is often limited by the number of element contained in the mesh due to limitation in computational resource (CPU and memory). Thus a more dense mesh is allowed in a Eulerian model compared to the corresponding Lagrange model, given that the volumetric size of the former is smallest. The length of the model geometry in the flow direction should be chosen such that the heat source can not “feel” either the inlet or outlet boundary. During simulation of metre long FSW panel, it is obvious, that including the whole geometry as an enmeshed part for studying details around the tool/heat source having size of millimetres, would be a computational challenge. In some cases, it might be impossible to obtain thermal results with sufficient resolution, if the whole domain is meshed. As an alternative, a smaller domain is enmeshed and the effect of the moving heat source is captured via moving the material instead, i.e. the Eulerian framework. The area surrounding the heat source can then be meshed with small elements, thereby ensuring adequate resolution. In many cases, the steady-state time scheme is used in combination with the Eulerian framework, as discussed previously. This efficient model setup allows investigation of heat flow through the tool and backing plate. By modelling the convective contribution of the material flow in the shear layer, highly detailed thermal results can be obtained. Since the heat source is stationary in the Eulerian domain, enmeshment of the tool is readily implementable. Including the tool and especially the tool probe, calls for preventing the material flow through probe. One procedure for this is to implement an analytical shear layer surrounding the tool in which the convective velocity is ramped from the welding speed outside the shear layer to the rotational speed at the tool interface. The convective velocity in the tool shoulder, probe and shaft domains have a rotational speed of wr.
10.3.3 Global thermal model: Lagrange framework The global thermal model using the Lagrangian framework is chosen for analyzing the transient effect of FSW. Such a model setup is the basis for simulating residual stress as well as microstructure models. The global model is characterized by having a heat source starting at one
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end of the workpiece – then moving along the weld path – ending at the other end of the workpiece, which is illustrated in Fig. 10.9. Each location in the model experiences an individual thermal history and the temperature field changes continuously as the moving heat source follows the weld path along the joint line. In the initial period of the simulation, a highly transient temperature field develops. During the plunge and dwell period, the workpiece temperature increases from room temperature to elevated temperatures where thermal softening allows for starting the traverse welding motion. At some point down the weld path, the temperature around the heat source stabilizes – which defines the steady-state region. As the heat source approaches the end of the workpiece, the insulation effect of the workpiece end surface can lead to increased temperatures.
10.3.4 Accounting for the convective heat transfer in the shear layer the simplest version of a thermal model is having a circular boundary at the top surface of workpiece at which a surface heat flux given by equation (10.17) is applied. In an Eulerian framework, the convective heat flow through the mesh is straightforward to model by prescribing it as uweld r c p dT . the next dx level of refinement of the heat source could be including the heat generation contribution from the probe. the probe itself has other thermal properties than the workpiece and it rotates together with the tool shoulder and shaft. The probe heat generation could be applied as a volume flux distributed over the volume of the tool probe. However, if the material is allowed to “flow” through the probe volume, an abrupt change in thermal conditions at the interface between the tool probe and the workpiece will be present. In the real weld, the material flows around the probe and not through it. Bearing this in mind, it could be advantageous find a method of avoiding such discontinuities which moreover conflict with the real heat flow. In the paper (Schmidt and Hattel, 2005a), a procedure for modelling the convective heat transfer within the shear layer is presented. the model includes an analytically defined velocity field that directs the heat flow around the tool probe instead of through it. this way, the convective heat transfer in the shear layer is accounted for. the next challenge is to prescribe the velocity components within the shear layer. the optimal solution would be a velocity field obtained from a fully coupled thermal-flow (CFD) model – however, if such data are available, then the thermal field would already be known – making the purely thermal model irrelevant. By prescribing a velocity field within the shear layer, which (to an appropriate extent) resembles the real flow and at the same time being simply enough to be expressed mathematically,
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the model will response more like the “real weld”, than if the material could flow through the probe volume. the shape of the shear layer can be uniform in thickness or variate along the contact interface. From experimental observations, it is seen that the shear layer is wider at the upper region due to the effect of the shoulder, and more narrow at the root region due to the colder thermal conditions, leading to a somewhat conically shaped shear layer. Figure 10.10 shows a conical shear layer surrounding a conical probe shape indication the dimensioning and dimensionless parameters. the ramping function r (Schmidt and Hattel, 2005a) is defined as a dimensionless parameter across the shear layer in the radial direction, increasing linearly from 0 at the probe/shear layer interface to 1 at the shear layer/matrix interface. The ramping function is defined as
rsl =
r – R probe = wsl
x 2 + y 2 – R probe wsl
10.38
where r is the radial distance to the rotation axis, Rprobe is the probe radius and w is the shear layer width in the radial direction. Using this ramping parameter, the velocity components are defined as u x = – y p d w (1 – rslp ) + uweld rslp
10.39
u y = x p d w (1 – rslp ) upper Rprobe
rsl = 0
rsl = 1
w slupper zsl = 0
Hprobe
zsl = 1 w sllower
lower Rprobe
10.10 Schematic view of conical shear layer used for allowing convective heat transfer around the tool probe.
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Friction stir welding
which gives counter clockwise flow with a tangential velocity of wrp at the tool/shear layer interface and velocity of uweld in the x-direction at the shear layer/matrix interface. the variables xp and yp are the coordinates for the closest point at the shear layer/probe interface relative to the (x, y) coordinates inside the shear layer, corresponding to the intersection point between the probe circle and line from (0, 0) to (x, y). the (xp, xp) coordinates are used in equation (10.39) for the probe side shear layer in order to ensure that the velocity components are defined relative to the innermost part for the shear layer and not the actual (x, y) position. Using (xp, yp) results in a linear velocity field for p = 1 whereas using the (x, y) coordinates leads to a nonlinear velocity field across the shear layer for p = 1, which could be critical for wide shear layers. the (xp, yp) coordinates are defined as xp =
x rp x rp = = r rp + r wsl
y rp y rp yp = = = r rp + r wsl
x rp x 2 + y2 y rp
10.40
x 2 + y2
For shear layers at the tool shoulder and probe tip (i.e., where the shear layer extends in the z-direction), the velocity components can be defined using the (x, y) variables. in the cylindrical coordinate system the shear layer flow field can be defined by the following tangential and radial velocity components uq = w rp (1 – rslp ) + sin q uweld rslp p ur = cos q uweld eld r sl el
10.41
where q is the angular position with zero defined in the welding direction. The convective heat flux inside the shear layer is the given by uxrcpT,x and uyrcpT,y which for normal welding parameters are dominant compared to the convective heat flux outside the shear layer. The convective heat flux in the shear layer results in a asymmetrical temperature field giving the highest temperature at the advancing side. Since width of the conical shear layer varies in z-direction, it is convenient to define the parameter zsl ranging from 0 at the upper region of the shear layer to 1 at the lowest region, i.e.
z sl =
Z probe probe – z H probe
10.42
where Hprobe is the tool probe height and Zprove is the z-coordinate of the top region.
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The shear layer width can then defined as upper lower wsl = wslupper (1 – z sl ) + wsllower z sl
10.43
r lower where Ruppe probe and R probe is the upper and lower shear layer width,
respectively. For a conical tool, the prober radius Rprobe is changing in the z-direction, and can be described by the following expression, lower lower R probe = Rupper probe z sl probe pr obe (1 – z sl ) + R probe
10.44
r lower where Ruppe probe and R probe are the upper and lower radius of the probe,
respectively. this makes the radial ramping function given by equation (10.38) z-dependent, i.e.
rsl (x, y, z ) =
r – R probe (z ) = wsl (z )
x 2 + y 2 – R probe (z ) wsl (z )
10.45
The velocity profile across the shear layer is more likely exponential or parabolic than linear which is resulted from the decrease in shear stress through the shear layer in combination with the temperature gradient across the shear layer and the material tendency to form at shear layer due to the shear thinning non-Newtonian constitutive characteristic of typical friction stir welded materials. To model the non-linearity of the velocity profile, the power law exponent psl is introduced, where psl < 1 gives a non-linear velocity profile such that the velocity decreases for increasing rsl values to a higher extend – resembling the exponential behaviour. It should be underlined, that for this simplified analytical model mass conservation does not prevail, since the full momentum equations are not solved.
10.3.5 Heat transfer to surroundings Even for a thermal model the computational resources are limited. therefore it is necessary to limit the number/extent of the geometries that are enmeshed. During the real welding process, the heat flow to the surroundings is considerable and should somehow be included in the model. the complete system can divided into: (i) the upward direction which includes the tool, tool holder, tool head, bearing, shafts and motor, (ii) the downward direction which includes the backing plate, the machine table and (iii) the side direction which include clamping tools. All of these components influence the the heat flow, and their importance depend on the actual welding setup. From
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Friction stir welding
a modelling point of view, each component has to be evaluated, whether it should be accounted for in the model or not. if the thermal effect of a component is to be included in the model, it can either be enmeshed and given thermal properties, or an equivalent heat transfer coefficient can be applied at the surface interfacing the geometry, which is not included in the model. As an example, the backing plate can either be enmeshed or an equivalent heat transfer coefficient hbp can be applied at the bottom surface of the workpiece. the value of hbp depends on the thermal properties of the backing plate and its geometry and by assuming 1-dimensional heat transfer in the z-direction, only, the equivalent heat transfer can be found from the following expression for the thermal resistance Rbp tbp Rbp = 1 = hbp k
10.46
where tbp is the thickness of the backing plate and k is the thermal conductivity. The equivalent heat transfer coefficient is then hbp = k tbp
10.47
For a backing plate with a thickness of 50 mm in steel having a thermal conductivity of ~ 50 W , the equivalent heat transfer coefficient can be mK Obviously, the thermal capacity is not accounted estimated to ~1000 W m2 K for when excluding the mesh of the backing plate. The heat flow in the x- and y-directions of the backing plate are not simulated when using an equivalent heat transfer coefficient, which could be a major disadvantage. The thermal resistance between to two object can be modelled using an equivalent heat transfer coefficient. As an example, the interface between the workpiece and the backing plate could be important to include in the model, since there is a temperature drop across that could influence the heat flow. if the backing plate is included as an enmeshed geometry, then it is recommended to include the contact resistance in the model. While placing the workpiece on the backing plate, the roughness at the two interfaces in contact results in the presence of small air gaps and microscopic contact segments, which will act as an thermal barrier. the thermal resistance is given by Rwp Æbp =
1 hwpÆbp
10.48
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After clamping the workpiece to the backing plate, it is assumed that the roughness is uniform, hence the thermal resistance can be assumed constant. However, during the plunge and welding periods, pressure from the tool forces in combination with the softening of the material near the process zone can lead to a deformation of the contact layer between the workpiece and backing plate. Such forging effect can result in increased ”compatibility” between the two adjacent surfaces, leading to a reduction in the thermal resistance in the process region. in a thermal model this can be simulated using a heat transfer coefficient that depends on position relative to the tool position, hence h(x, y) for an Eulerian framework and h(x, y, t) in a Lagrangian framework. Values ranging from 100 to 10000 W2 or even m K higher can be applied depending on the conditions, the former is descriptive for a high thermal resistance and the latter corresponds to a low resistance (perfect contact). the thermal resistances can be added, hence the total thermal resistance is RwpÆmt = RwpÆbp + Rbp + RbpÆmt 10.49 where subscripts denotes the geometries, i.e. wp – workpiece, bp – backing plate and mt – machine table. The equivalent heat transfer coefficient is found from RwpÆmt =
1 = 1 + 1 + 1 hwpÆmt hwpÆbp hbp hbpÆmt
10.50
The heat flux from the bottom side of the workpiece can then be modelled as q=
(T Twp bottom – Tmt ttop ) = hwpÆmt (Twp botttoom – Tmmtt ttop ) RwpÆmt
10.51
A thermal resistance between two metals in contact could well attain values of 1e-4, equivalent to a heat transfer of 10000 W2 . Using the equivalent m K W heat transfer coefficient of 1000 2 for a 50 mm steel backing plate with a m K thermal conductivity of approximately 50 W2 , the total thermal resistance m K is RwpÆmt =
1 hwpÆbp
+ 1 + 1 = 1 + 1 + 1 = 0.0012 hbp hbpÆmt 10000 1000 000 10000 10.52
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W underlining m2 K the dominating factor of the backing plate. if the thermal contact is perfect (zero resistance) then the equivalent heat transfer coefficient would be 1000 W , but if the thermal resistance at each interface is 1000 W , then the m2 K m2 K W total heat transfer coefficient would be 333 2 , corresponding to having m K three times the thickness for the backing plate, hence such high resistances must be accounted for. from which the heat transfer coefficient is found to be ~833
10.3.6 Thermal efficiency The term thermal efficiency has been proposed in Schmidt and Hattel (2008a) as a remedy to evaluate how much heat is conducted into the workpiece itself compared to the total heat generation. The thermal efficiency can be expressed both on a global level, describing the heat flow between the parts, and on a local level, describing the heat flow into the workpiece evaluated at a contact region. the heat is generated at the tool/workpiece interface or within the shear layer. The heat can flow either into the workpiece denoted Qworkpiece or into the tool and up through the tool shaft, which is denoted Qtool and can be described as a heat loss. The global thermal efficiency is then given by
h=
Qworkpiece Q – Qttool rk rkpiece ool = total Qtotal Qtotal
10.53
where the total heat generation is found by integrating the local heat generation over the contact area between the tool and workpiece Atool/workpiece, i.e. Qtotal =
Ú
qqd A
10.54
Atool /workpiece rk rkpiece
The heat loss through the tool is found by integrating the heat flux at the surfaces of the tool, from which heat is conducted to the surroundings, hence all other surfaces than the tool/workpiece interface. Alternatively, Qtool can be estimated by integrating a cross section of the tool shaft close to the shoulder. The thermal efficiency is important, when making a thermal model, which does not include the tool. this is often the case in Lagrangian models, or some type of Eulerian models, where only the overall effect of the heat flow through the tool is accounted for. In such case, the local heat flux should be given as
Modelling thermal properties in friction stir welding
q=
3Qwork 3h Qtotal r rkpiece piece r = 3 3 2p Rshoulder 2p Rshou llder der
303
10.55
where values of h ranging from 0.7 to 0.9 are common. An alternative procedure is to define the local heat source using the traditional expression for the heat source, and then subtract the heat flux through the tool. the “real” heat transfer through the tool shaft is by conduction, but under the assumption of 1-dimensional steady state heat flow, it can be modelled by convective heat transfer using an equivalent heat transfer coefficient, i.e. q = tcontact w r – htool (T – T•)
10.56
k and T is the temperature at the far-end of the tool (where • H tool it mounted in the tool holder). where htool =
10.3.7 Thermo-pseudo-mechanical model in the following, the thermo-pseudo-mechanical (tPm) model will be presented. A detailed description of the tPm model is given in the original papers (Schmidt and Hattel, 2008a, b), which must be consulted for additional information regarding the model setup and results. the main feature of the tPm model is the use of the temperature-dependent heat source, described in detail previously in this chapter. this heat source is given by equation (10.30) and requires temperature-dependent yield shear stress data, i.e. t(T) which can be converted from experimentally measured yield stress data using the following expressions
t (T ) =
s (T ) (3)
10.57
When using measured values of the material strength, it is often the case that values for temperature in the high temperature regime are not reported. Therefore, it is necessary to find a method of giving reasonable values for the flow stress at elevated temperature. Based on the assumption that the strength of the material drastically reduces ,when the temperature approaches the solidus temperature, the yield stress is set to zero at temperature at or above the solidus temperature. Yields stress values from the last recorded temperature up to solidus temperature are found by interpolation. Since the yield shear stress is the driver for the local heat generation, this also means that the heat flux is zero if the temperature exceeds the solidus temperature, in which case the heat source “turns it-self off”. As a result, the temperature in a simulation using the tmP model can never exceed the
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temperature where the heat source is cut off, which in this case is the solidus temperature. The “self-stabilizing” effect of the TPM-heat source means that the model will response in changes of thermal boundary conditions, change in material properties and welding conditions (welding and rotational speeds). The discussion of which “cut-off” temperature should be used can be rather academic, but nevertheless very important. The material can be treated as a semi-solid in the temperature range above the solidus temperature and below the liquidus temperature. At a given solid fraction, a reasonable sharp transition in the material flow characteristic (expressed by the viscosity) can be observed. The temperature, at which this transition occurs could be used as a cut-off temperature, being an alternative to the solidus temperature used in the present work. In the following, the TPM heat source model is used to simulate the heat generation and heat flow in an experimental weld of 7075-T6 plates. The welding conditions are listed in Table 10.2 and the thermal properties of both the workpiece (7075-T6) and the backing plate (steel) are given in Table 10.3. The temperature-dependent yield stress data listed in Table 10.4 are converted into shear stress values using equation (10.57). The yield stress data are for 7075-T6 friction stir processed material, evaluated at strain rate of 1e-3 from stress-strain graphs (some loss of accuracy expected), which are reported in Cavaliere and Squillace (2005). The thermal resistance between the workpiece and the backing plate is modelled using a constant
Table 10.2 Welding conditions Welding speed Rotational speed Shoulder radius Probe radius (top) Probe cone angle Workpiece dimensions
40 mm min–1 = 0.66 mm s–1 535 rpm 12.5 mm 5 mm 15° 8 ¥ 100 ¥ 250 mm
Table 10.3 Thermal properties for structural steel and 7075-T6 used in TPM model 7075-T6 (from ASM Handbook Committee, (1990)) Thermal conductivity 130 W (mK)–1 Specific heat capacity 960 J (m3K)–1 Density 2810 kg m–3 Steel (tool and backing plate) (from Comsol 3.3–3.5,www.comsol.com) Thermal conductivity 44.5 W (mK)–1 Specific heat capacity 475 J (m3 K)–1 Density 7850 kg m–3
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Table 10.4 Maximum yield stress for FSP 7075-T6 for e = 10–3 s –1 from Cavaliere Squillace (2005) Temperature (°C)
Yield stress (MPa)
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10.11 Far-field temperature field in workpiece, tool and backing plate.
W , and the convective heat transfer at m2 K the top surface of the workpiece is 10 W2 .the heat loss through the top m K region of the tool shaft is modelled using a heat transfer coefficient of W 10.000 2 . Figure 10.11 shows the far-field temperature field in the tool, m K workpiece and backing plate. Figure 10.12 shows the near-field view of the heat transfer coefficient of 700
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450°C
500°C
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10.12 Detailed view of temperature field in the mid-plane of the workpiece, the conical shear layer and tool probe. The streamlines show the “velocity field” used in the convective heat transfer term. The velocity profile is ramped between the welding speed at the outer region of the shear layer and the rotational speed at the tool/ matrix interface.
temperature and the “material flow”, which are prescribed analytically for controlling the convective heat transfer around the tool probe. In this case, a power law coefficient psl of 0.005 is used, thereby simulating the exponential characteristic of the velocity profile. The temperature-dependent heat source calls for an iterative solution scheme for the non-linear thermal problem, since the local heat generation is part of the thermal solution. In a traditional thermal model, the local heat generation is by q = twr is prescribing at the tool/workpiece interface. This is here denoted as an analytical heat source, since it is defined prior to solving the numerical equation system. Using the TPM-heat source, the model parameters (such as heat transfer coefficient) are adjusted until the best correlation between the experimental and simulated temperature profiles is obtained. Figure 10.13 shows a comparison between the temperature profiles at 15, 30, 45 and 60 mm from the joint-line, and by integration the local heat generation at the tool/ workpiece contact interface, the total heat generation is found to 1.9 kW. For evaluation the main difference between the TPM-heat source and the analytical heat source, the latter is implemented in a second model, having the same
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400 y y y y y y y y
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= = = = = = = =
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– – – – – – – –
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10.13 Comparison between the thermal-pseudo-mechanical model and experimental results for the far-field temperature profiles at y-values of 15, 30, 45 and 60 mm.
properties except for the applied heat flux. The shear stress is adjusted to t = 7.2 MPa such, that the integration of the analytical heat source equals the total heat generation found in the TPM-model. Consequently, the far-field result of the two models are nearly identical; however, the temperature field and the local heat generation distribution are different. Figure 10.14 shows the temperature along the intersection between the joint-line and tool/workpiece interface. In the analytical heat source model, the temperature exceeds the solidus temperature reaching a temperature of around 541°C. However, the TPM-model gives temperatures that are below the solidus temperature, more precisely 520°C. As mentioned before, the peak temperature is limited by the self-stabilizing effect, where the heat generation “turns itself off”, when exceeding the cut-off temperature. The temperature field is more uniformly distributed, since colder regions will response with higher shear stress/heat generation and hotter regions will response with lower shear stress/heat generation – balancing out temperature gradients. Figure 10.15 shows the distribution of the local heat generation along the x-axis. A gradual (exponential) increase is observed at larger distance from the centre for the temperature-dependent heat source. Notice, how the heat generation distribution is asymmetrical, having highest values towards the leading side under the shoulder, due to the colder material “flowing” toward
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Thermal-pseudo-mechanical Analytically prescribed heat source
Tsolidus = 532°C
540
Temperature [°C]
520 500 480
Leading shoulder
460 Leading probe conical side 440
Probe tip Trailing probe conical side
420 Trailing shoulder 400
–0.01
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10.14 Temperature evaluated along intersection between tool and centreline for the thermal-pseudo-mechanical model and the analytically prescribed heat source model with temperatures between 475–520°C and 459–541°C, respectively. Maximum difference is ~21°C.
the tool with the welding speed. For comparison, the analytical prescribed heat source gives a linear V-shaped profile due to the proportionality with wr. In Table 10.5 the different heat generation contributions and thermal efficiency are summarized. These specific values are representative for this specific welding conditions and tool geometry used for welding the 8 mm thick aluminium plates used in this experimental setup. It should be mentioned, that for welding of thicker workpieces, e.g. 25 mm or larger, the tool is designed with a smaller shoulder diameter relative to the probe diameter, compared to the tools used for thinner workpieces. As a consequence, the fraction of heat generated from the probe will then be higher. Once the model is “calibrated” to the experimental setup, the influence of changing welding parameters is analysed. In the present case the TPM-model is used to simulate the effect of varying the welding speed in the range from 1 to 10 mm/s and the rotational speed from 100 to 1000 rpm, i.e. obtaining a process window. Figure 10.16 shows a contour plot of the maximum temperature in the workpiece. The maximum temperatures do not exceed the solidus temperature (532°C); however the peak temperature reaches a plateau 10–15°C below Tsolidus for combinations of low welding speed and high rotational speed due to the self-stabilizing effect in the heat source.
Modelling thermal properties in friction stir welding 12
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¥106 Thermal-pseudo-mechanical Analytically prescribed heat source
Heat generation [w/m2]
10
Leading shoulder Leading probe conical side
18
Probe tip 6 Trailing probe conical side Trailing shoulder
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10.15 Local heat generation evaluated along the intersection between the tool and the centreline for the temperature-dependent and analytically prescribed heat source, receptively. The total heat generation used in the expression for the analytically prescribed heat source model is found from the thermal-pseudo-mechanical model. Table 10.5 Heat generation contributions calculated from the TPM model using the temperature-dependent heat source
(W)
(%)
Total heat generation From shoulder From probe side From probe tip Global heat flow into tool Global heat flow into workpiece
1909 1581 296 32 236 1676
100 83 16 1 12 88
Figure 10.17 shows a contour plot of the total heat generation ranging from 1600–4000 W. The grid of 10-by-10 simulations takes approximately 1 hour on a standard PC, making this a good alternative to the more computationally expensive fully coupled thermomechanical/CFD models. Figure 10.18 shows the maximum temperatures as function of rotational speed for a constant welding speed of 1 mm/s. Notice how the maximum temperatures stabilize for higher rotational speeds – being limited by the melting temperature of
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10.16 Contour plot of the maximum temperature as a function of rotational speed and welding speed.
532°C. Figure 10.19 shows the maximum temperatures as function of welding speed for a constant rotational speed of 100 rpm. The maximum temperatures decrease for increasing welding speeds. This is in good agreement with experimental observations. For this reason, welds with high ratio between the welding speed and rotational speed are denoted cold welds. When developing thermal models of friction stir welding, many different decisions have to be taken by the modeller. Should the Eulerian or the Lagrangian framework be selected for a given thermal problem? What about the time scheme – is the thermal problem time dependent or steady state? How does the level of detail in the heat source description influence the thermal results? How can temperature and heat generation be evaluated for welding conditions not supported by experimental measurements? Using the temperature-dependent heat source model could be an alternative to making a fully coupled thermo-mechanical/CFD model, still having its limitation in mind. Hopefully, this chapter will support the modeller in making some of these decisions during the developing of thermal models of friction stir welding.
Modelling thermal properties in friction stir welding 10 300
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10.17 Contour plot of the heat generation as a function of rotational speed and welding speed. 530 520 510
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500 490 480 470 460 450 440 430 100
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10.18 Maximum temperature as a function of rotational speed (constant welding speed of 1 mm/s). Notice how the temperature stabilizes for higher rotational speeds – being limited by the melting temperature of 532°C.
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10.19 Maximum temperature as a function of welding speed (for constant rotational speed of 100 rpm). The maximum temperatures decrease when increasing the welding speed which corresponds well with experimental observations.
10.4
References and further reading
ASM Handbook Committee, Metals Handbook, Vol. 2 – Properties and Selection: Nonferrous Alloys and Special-Purpose Materials, ASM International 10th Ed. (1990). P. Cavaliere, A. Squillace, J. Materials Characterization 55 (2005) 2, 136–142. Y.J. Chao, X. Qi, J. Materials Processing and Manufacturing Science, 7 (1998) 10, 215–233. Y.J. Chao, X. Qi, W. Tang, J. Manufacturing Eng. 125 (2003) 1, 138–145. C. Chen, R. Kovacevic, J. Mach. Tools Manuf. 43 (2003) 13, 1319–1326. P. Colegrove, Proc. 2nd Int. Symp. on Friction Stir Welding, TWI, Sweden, 2000. P.A. Colegrove, H.R. Shercliff, J. Sci. Tech. Weld. Join. 9 (2004) 6, 352–361. P.A. Colegrove, H.R. Shercliff, R. Settler, J. Sci. Tech. Weld. Join. 12 (2007) 4, 284–297. Comsol 3.3-3.5, www.comsol.com. Z. Feng, X-L. Wang, S. David, P. Sklad, Proc. 5th Int. Friction Stir Welding Symp., Metz, France, TWI, 2004. R.W. Fonda, S.G. Lambrakos, J. Sci. Technol. Weld. Joining 7 (2002) 3, 177–181. Ø. Frigaard, Ø. Grong, B. Bjørneklett, O.T. Midling, Proc. 1st Int. Symp. on Friction Stir Welding (Thousand Oaks, CA, USA) 1999. Ø. Frigaard, Ø. Grong, O.T. Midling, J. Metall and Mater. Trans. A 32 (2001) 5, 1189–1200. A. Gerlich, M. Yamamoto, T.H.North, J. Sci.Tech. Weld. Join. 12 (2007) 6, 472–480.
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P. Heurtier, M.J. Jones, C. Desrayaud, H.J. Driver, F. Montheillet, D. Allehaux, J. Mat. Proc. Tech. 171 (2006) 348–357. M.Z.H. Khandkar, J.A. Khan, J. Mater. Process. Manuf. Sci. 10 (2001) 91–105. M.Z.H. Khandkar, J.A. Khan, A.P. Reynolds, Sci. Techn. Weld. Joining 8 (2003) 3, 165–174. M.J. Russell, H.R. Shercliff, Proc. 1st Int. Symp. on Friction Stir Welding (Thousand Oaks, CA, USA), 1999. H. Schmidt, Modelling the Thermomechanical Conditions in Friction Stir Welding, Technical University of Denmark, 2004. H. Schmidt, J. Hattel, J. Sci. Tech. Weld. Join. 10 (2005a) 2, 176–186. H. Schmidt, J. Hattel, J. Modelling Simul. Mater. Sci. Eng. 13 (2005b) 77–93. H. Schmidt, J. Hattel, J. Wert, J. Modelling Simul. Mater. Sci. Eng. 12 (2004) 143– 157. H.B. Schmidt, J.H. Hattel, Scripta Materialia 58 (2008a), pp. 332–337. H.B. Schmidt, J.H. Hattel, 7th Int. Symposium Friction Stir Welding, Awaji, Japan, TWI (2008b). H. Schmidt, J. Hattel, European Comsol conference, Hanover (2008c). Q. Shi, T. Dickerson, H. Shercliff, Proc. 4th Int. Symp. on Friction Stir Welding (Park City, UT, USA) 2003. M. Song, R. Kovacevic, J. Mach. Tools Manuf., 43 (2003) 6, 605–615. C.C. Tutum, H. Schmidt, J. Hattel. M.P. Bendsøe, Proc. 7th World Congress on Structural and Multidisciplinary Optimization, Soul, Korea 8 (2007) pp. 2639–2646.
11
Metallurgy and weld performance in friction stir welding
J. F. dos Santos, GKSS Forschungszentrum GmbH, Germany, C. A. W. Olea, Vallourec & Mannesmann Tubes do Brasil, Brazil, R. S. Coelho, Helmholtz-Zentrum Berlin für Materialien und Energie GmbH, Germany, A. Kostka, Max-Planck-Institut für Eisenforschung GmbH, Germany, C. S. Paglia, University of Applied Sciences of Southern Switzerland, Switzerland, T. Ghidini, European Space Agency, The Netherlands and C. D. Donne, EADS Innovation Works, Germany
Abstract: The present chapter covers the microstructural development of friction stir welds (FSW), their corrosion behaviour and mechanical properties. The section on microstructural aspects of friction stir welds presents a general description of the different weld zones and their main characteristics. The sub-structural evolution of an AlMgSc alloy is described in detail due to its industrial significance. For the same reason, the metallurgy of dissimilar welds is described based on two case studies: Al-high strength steel and Al-Mg joints. The section on corrosion starts with a general description of corrosion phenomena in Al alloys. This is followed by a specific description of the corrosion behaviour of friction stir welds and their corrosion fatigue properties. This section is concluded with an analysis of the influence of the base material temper on the corrosion resistance of the friction stir welded joints. The chapter is concluded with a comprehensive analysis of the mechanical behaviour of friction stir welded joints focused on fatigue and fatigue crack propagation. In both cases relevant factors controlling joint performance are discussed including thickness and notch effects, loading condition as well as environmental and residual stress effects. In general terms, this chapter addresses high performance applications of friction stir welding such as those found in the transportation industry and other load carrying engineering structures. However the information provided is certainly also applicable to less stringent industrial applications of the process. Key words: friction stir welding, microstructure, sub-structure, AlMgSc alloy, dissimilar joints, Al-steel, Al-mg, corrosion, corrosion-fatigue, mechanical properties, fatigue, fatigue crack propagation.
11.1
Introduction
The complexity of the thermomechanical phenomena in FSW has, since the inception of the process in 1991, not ceased to fascinate scientists. The 314
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phenomena around joint formation and microstructural development has received particular attention in the literature since the first comprehensive overview published by Threadgill in 1997. Similarly, the effects of the gradient microstructure across friction stir welds on the local and global properties of the joint have been consistently reported in the literature. Figure 11.1 presents a general overview of the indexed publications on FSW since 1999 divided into the three knowledge areas: microstructure, properties and corrosion. The presented information has been obtained from a simple search at http://www.sciencedirect.com with very basic filtering and does not claim absolute qualitative correctness but rather describes general trends. Figure 11.1 shows an increase of approximately 1000% in publications in FSW on the selected knowledge area in the last nine years. Although it is almost impossible to strictly separate “properties” from “microstructure”, the trends clearly indicate the amount of scientific effort in these areas. Joint formation and microstructure development of friction stir welds in almost all relevant engineering Al alloys is extensively covered in the literature. Although some questions on the kinetics of precipitation phenomena in the SZ, TMAZ and HAZ as well as on recrystallisation mechanisms in the SZ are still being studied, a general understanding of the metallurgical phenomena in FSW in Al alloys is available. The metallurgy of FSW of other materials has, in general, not been systematically pursued or is presently gaining maturity, since most of the scientific work is currently focused on process-related aspects. Hence, this chapter includes an introductory section summarising the present state-of-art on the metallurgy of FSW in Al alloys. This is followed by an in-depth analysis of the metallurgical phenomena in a AlMgSc alloy which has been specifically developed for complex structural demands such as in aircraft applications. A further section on Al-based dissimilar welds completes this section. The scientific efforts on corrosion resistance of FSW alloys have not consistently grown in the last ten years as indicated in Fig. 11.1. Thus a more generic approach to this topic has been adopted in this chapter. Figure 11.1 indicates that the knowledge area “properties” is the most frequently addressed in the literature. As expected, the emphasis of the studies lies mostly on hardness, tensile strength and fatigue. In this chapter the section on properties is intended to complement the information available in the literature focusing therefore on fatigue, fatigue crack propagation and fracture toughness of friction stir welds in Al alloys.
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11.1 General overview of the indexed publications on FSW since 1999.
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Metallurgy of friction stir welds in aluminium alloys
11.2.1 Introduction The FSW process has emerged as a promising solid-state process with the potential to join low melting point materials, particularly aluminium alloys. The most attractive reasons for joining aluminium alloys through this process are the avoidance of solidification defects formed by conventional fusion weld processes, the low distortion and residual stresses. Consequently, friction stir welding is effective and has been systematically developed for joining aluminium alloys, especially those usually considered non-weldable or difficult to weld such as 2xxx and 7xxx series alloys (Hassan et al. 2003a; Genevois et al. 2005). During the process, thermo-mechanical conditions vary across the weld seam, resulting in different microstructures. Work to date has concentrated on single-pass welds in material thicknesses from 1.6 mm to 10 mm but it has been shown that a thickness of up to 50 mm can be joined with one pass, and up to 75 mm with two passes, one in each side (Dawes et al. 1999, Mishra and Ma 2005). Solid-state stirring leads to fine equiaxed grains in the weld centre, resulting from dynamic recrystallisation (stir zone – SZ), followed by a typically recovered microstructure around the stir zone, which is affected by lower levels of deformation and temperature in comparison to the weld centre (thermo-mechanical affected zone – TMAZ) and a region affected only by temperature, with a microstructure similar to the base material (heat affected zone – HAZ). The stir zone often displays a well-defined onion ring type of inner structure, consisting of concentric ovals; however, in some alloys this feature is not visible. It depends on the alloy used and the process conditions. The diameter of the stir zone is typically slightly larger than that of the tool pin and significantly less than the shoulder diameter. The stir zone also typically extends to the bottom of the weld, but not always. The prediction of the SZ shape is directly related to tool design and welding parameters, besides the hot strength of the material being welded. The macrostructure outside the stir zone generally presents an area adjacent to the SZ where severe plastic deformation of the material occurs, such that the elongated grain structure previously present in the base material of the majority of rolled plates can be rotated by as much as 90°. Local mechanical properties are related to these welded regions. For example, in several previous studies the hardness profiles across the joint have been taken as an indication of changes in microstructure and respective mechanical aspects (Threadgill 1997). Normally the mechanical properties of precipitate-strengthened aluminium alloys tend to deteriorate in the weld zones and, generally, they depend on the dissolution and/or coarsening of strengthening second phase
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particles (Venkateswarlu et al. 2004). The macrographs in Fig. 11.3 shows a typical joint segment, representative of the plane transverse to the weld travel direction of aluminium FSW joints with the respective identification of the weld zones. A reduction in hardness in the weld zones is usual in friction stir welds of artificially aged aluminium 6xxx series (up to 50%, depending on the energy input and temper condition of the alloy). The microstructure of FSW aluminium alloys, in the heat affected zone (HAZ) usually shows no apparent difference from that of the base material. In age-hardened alloys, the hardness level tends to decrease in the HAZ, indicating clearly that the thermal cycle during the welding process has a significant influence, either by overaging, or by decreasing dislocation density, probably by both in fully aged alloys. In the thermomechanically-affected zone (TMAZ), through optical microscopy it is perceived that FSW causes bending of the pancake grains, and occasionally some local recrystallisation. The heat also accelerates the aging and annealing processes, although the latter will compete with some degree of work hardening. Typical hardness profiles for age-hardened alloys have shown that the hardness usually reaches a minimum in this region. Work hardening is the dominant effect close to the stir zone, while overaging/annealing is the most important effect towards the HAZ. It is also possible that local areas may reach a sufficiently high temperature to allow partial precipitate dissolution, particularly in the TMAZ (Mishra and Ma 2005; Olea et al. 2006; Threadgill 1997). There is usually considerable evidence for partial recrystallisation close to the interface with the stir zone. It should be pointed out that this interface is generally very sharply defined, especially on the side where the direction of the tool rotation is the same as the direction of travel (advancing side). The interface on the side where the direction of the tool rotation is contrary to the direction of travel (retreating side) usually shows a wider gradient of microstructure. The microstructure of the SZ is clearly equiaxed and very fine, varying the grain size according to the alloy and FSW process parameters. The typically low resolution of optical microscopy is usually not enough to investigate these fine grains, being in general a few microns or less. Hardness levels are usually below the base material level in age-hardened and mechanically hardened alloys (Threadgill 1997). Typical interfaces between TMAZ and SZ for the advancing and retreating sides of friction stir welded aluminium alloys are illustrated in Fig. 11.4. More thorough investigations of the individual weld regions using electron microscopy have shown the particular effects of the weld energy input on the substructure of several aluminium alloys. Usually a considerable reduction in dislocation density in the SZ is reported, compared to the other regions, as a consequence of the recrystallisation phenomenon (Mishra and Ma 2005; Olea et al. 2006; Genevois et al. 2005). Different kinds of precipitates have
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been observed along the different FSW zones, from dispersoids to very thin strengthening precipitates, occurring intragranular and/or intergranular, depending essentially on the alloying elements present. The coarsening and dissolution of the precipitates along the TMAZ and SZ have been reported for several age-hardened alloys (Threadgill 1997). The different second phase particles formed in different strengthening precipitate alloys vary in chemical composition, size, morphology and distribution, depending on the aluminium alloy and process parameter employed to produce the joint. The different precipitation features for different precipitate strengthened alloys are described in detail in the following sections. The stir zone and HAZ in solution heat-treated and artificially aged aluminium alloys (i.e., 2xxx and 6xxx series) can recover to strength levels close to the fully heat-treated base material, by post-weld heat treatment (Dawes et al. 1999).
11.2.2 An analysis of the welding metallurgy of an AlMgSc alloy and al-based dissimilar joints AlMgSc alloy Effect of scandium addition Excellent mechanical properties at room temperature are attributed to aluminium alloyed with scandium. This is due to the presence of nanometresized Al3Sc precipitates which are effective in pinning mobile dislocations and stabilising a fine grain structure (Toropova et al. 1998). The solubility of scandium in solid aluminium is about 0.22%, 0.15% and 0.05% at 645°C, 600°C and 500°C, respectively. The Al3Sc phase has about 35.7% Sc and is in equilibrium with the aluminium solid solution, having an FCC lattice with the following parameters: a = 0.4106 nm, space group Pm3m, 4 atoms per unit cell. The scandium tends to form supersaturated solid solutions with aluminium, even at relatively low cooling rates. The Al3Sc precipitates are usually stable against coarsening, up to temperatures of 350°C and the decomposition temperature sometimes can reach 650°C, depending on the alloying elements (Toropova et al. 1998; Marquis and Seidman 2001). Precipitates formed in solid state are coherent with the aluminium lattice, normally have a spherical appearance and the volume fraction of precipitates is small due to the low solubility of scandium in aluminium. Therefore, Sc containing particles are generally known to increase the recrystallisation resistance, and consequently grain refinement. The strengthening mechanisms that act in Al-Mg-Sc alloys are mainly precipitation and structural (solid solution) hardening due to the Al3Sc phase, solid-solution hardening due to magnesium and precipitation hardening with respect to the Al3Mg2 phase in alloys containing more than 8% Mg.
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The decreased difference in lattice parameters makes it possible to preserve the coherence of Al3Sc precipitates at higher temperatures and for longer exposures. When compared to commercial age-hardening 2xxx and 6xxx series alloys containing Cu, Mg and Si, where precipitates usually coarsen rapidly above 250°C, it is reasonable to assume that Al3Sc precipitates are much more stable. Thus, aluminium alloys containing scandium precipitation strengthening, can be used at temperatures significantly higher than conventional precipitatehardened aluminium alloys (Seidman et al. 2002). The homogeneous structure and fine grain size promoted by scandium addition improves formability, yield strength and toughness. The material applications include defence (especially in Soviet Union), sporting goods, transportation, and the aerospace industry (Venkateswarlu et al. 2004). The excellent control of the grain structure in thermomechanical treatments produced by Al3Sc precipitates and significant particle strengthening are indications of good performance in welding applications. Thermal history during FSW This presents results of temperature measurements carried out in-situ during the welding of Al-Mg-Sc joints, obtained from thermocouples placed in the workpiece (TMAZ/HAZ) and in the backing bar. Thermal cycles have been obtained from different samples welded with different energy inputs. Temperature measurements obtained from the backing bar are considered as an indication of the temperatures in the SZ. Minor differences in the maximum temperature in the backing bar can be attributed to slight misplacement of the plates. A possible shift of the weld centre line can also be considered, but may not be significant. The thermal history measured for the Al-Mg-Sc joints produced with constant rotation speed (700 rpm), but different welding transverse speeds, in low (525 mm/min), intermediate (350 mm/min) and high (175 mm/min) heat input (HI), are shown in Table 11.1. The increase in the temperature values recorded by the thermocouples agrees with the increase in the heat input produced by the variation of the FSW Table 11.1 Thermal history of Al-Mg-Sc FSW joints using low, intermediate and high heat input, respectively Energy input Weld zone Max. temp. Heating rate Cooling rate Dwell time
Low HI HAZ/TMAZ 250°C 36.6°C/s 2.85°C/s 6 s
SZ 295°C 106°C/s 3.14°C/s 2.5 s
Intermediate HI HAZ/TMAZ 276°C 30.7°C/s 3°C/s 8 s
SZ 309°C 55.8°C/s 3.32°C/s 5 s
High HI HAZ/TMAZ 327°C 12.4°C/s 2.8°C/s 24 s
SZ 372°C 26.3°C/s 2.4°C/s 13 s
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parameters. A decrease of the heating rate can be verified using intermediate HI in comparison with low HI, but the cooling rate does not significantly change. If compared to the low HI joint, the delay in the heat-up to peak temperature of the intermediate HI joint, associated with an inferior welding speed, could be linked to the higher peak temperatures and consequently more homogeneous stirring. A significant variation of mechanical and metallurgical behaviour would not be expected, since no large variations of the peak temperatures and cooling rates occurred. Furthermore, the high stability of the compounds present in the Al-Mg-Sc alloy would not be likely to be susceptible to slight variations in the thermo-cycle (Seidman et al. 2002). For the high HI condition, the increase in the peak temperature values obtained using the thermocouples also agrees with the increase in the heat input produced by the variation of the FSW parameters, as well as with experiments reported in the literature (Reynolds and Tang 2001). Once again, it verified a reduction of the heating rate, followed by an increase in cooling time, but no significant variation of the cooling rate. The heating rate can be directly associated to the variation in welding speed, and also relative to the higher peak temperatures overcome and homogeneous stirring. For this FSW condition, a variation of mechanical and metallurgical behaviour could be expected, since the peak temperatures and thermo-cycles increase reasonably, in comparison with cold and intermediate parameters. However, the high stability of the compounds present in the Al-Mg-Sc alloy could prevail at these peak temperature and thermo-cycles, proving of little consequence for the mechanical and metallurgical properties or even retaining the original properties of the base material (Toropova et al. 1998; Seidman et al. 2002). The influence of the thermal cycles on the Al3Sc stability is represented by the graphs in Fig. 11.2 for the SZ and TMAZ. Even the highest thermal cycle in the SZ is far away from dissolution temperature, but the coarsening temperature has shown to be reached. However, no detectable variation of the precipitate size and distribution is expected and suggests that the thermal cycle (high HI) is too fast for significant coarsening of Al3Sc precipitates to occur. In the TMAZ, all welding conditions have shown to be below the coarsening temperature, indicating stability of the Al3Sc precipitates in these temperature fields. Microstructure evolution As mentioned in Section 11.2.1 Introduction, a typical FSW joint is commonly divided into zones, which are affected by different thermal cycles and degrees of deformation. The microstructure of the Al-Mg-Sc base material consists of deformed elongated grains, while the centre of the welds (SZ), which undergoes the highest amount of deformation and thermal cycle, is
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Friction stir welding 700
Dissolution of Al3Sc
600
Temperature (°C)
500 400
Coarsening of Al3Sc
300
Low HI TMAZ
200
In termediate HI TMAZ High HI TMAZ
100 0 0
20
40
60 80 Time (s) (a)
700
100
120
140
Dissolution of Al3Sc
600
Temperature (°C)
500 400
Coarsening of Al3Sc
300
Low HI SZ Intermediate HI SZ
200
High HI SZ
100 0 0
20
40
60 80 Time (s) (b)
100
120
140
11.2 Correlation between precipitation stability and thermal cycles obtained in (a) TMAZ and (b) SZ of Al-Mg-Sc joints welded with three different energy inputs.
characterised by a fine equiaxed recrystallised grain structure. In contrast, surrounding the SZ, the recovered grains of the TMAZ, which have a highly distorted structure, are characterised by a high degree of deformation and rotation of about 90° of the base material pancake grains. Macrographs and micrographs in Fig. 11.3 show a typical joint segment, representative of the plane transverse to the weld travel direction of Al-MgSc FSW joints. Here the geometry of the weld seam is easily revealed in the macrograph, as well as the plastic flow tendency in the stirred zone, indicating
Metallurgy and weld performance in friction stir welding
HAZ
Z
SZ
TMA
HAZ
TMAZ
BM
323
5 mm (a)
200 µm (b)
500 µm (c)
11.3 Transversal section of the weld seam of the Al-Mg-Sc FSW joint by optical microscopy (a) macrograph showing overview of the joint with respective weld zones, (b) the microstructure of the HAZ/ BM and in (c) shear bands at the interface, delimiting SZ and TMAZ regions.
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Friction stir welding
full penetration and no significant defects. The grain boundary details are not properly revealed using optical microscopy, even using polarised light, as can be seen in the micrographs detailing the base material and interfaces of TMAZ and SZ (Fig. 11.3 a and b). However, it is possible to distinguish the different grain morphologies, coming from the rolled base material (BM) in the missoriented grains in the TMAZ, surrounding the very fine recrystallised grains in the SZ. Such similar microstructural aspects of cold rolled BM have been reported in the literature for sheets annealed at high temperatures (300–400°C), demonstrating a completely non-recrystallised structure (Filatov et al. 2000). Additionally, observing the TMAZ elucidated in Fig. 11.3 (c), particular features of the grains with large through-thickness shear band formation are shown. The bands form initially at approximately ± 35° to the rolling plane and divide the plate into rhomboidal prisms with long axes parallel to the transverse direction of the plate. Very close to the SZ, the region is particularly characterised by shear bands rotated surrounding the interface between TMAZ and SZ. According to the FSW thermal cycle measurements, Genevois et al. (2005), the HAZ experiences peak temperatures from 200°C (bottom) to 250°C (top) measured for the low HI joints, reaching from 300°C (bottom) to 350°C (top) in the high HI joints, which, especially for the high HI, would influence the strengthening features in most aluminium alloys. In contrast, very close to the SZ, the temperature is estimated to reach between 350 and 450°C and, combined with deformation, a rotation of elongated and partially recrystallised grains around the SZ is produced. A large plastic deformation is introduced in the SZ during FSW and temperatures are expected to reach more than 500°C, giving place to a microstructure entirely recrystallised with grain size in the order of about 1mm. It is not possible to acquire quantitative results about grain size using optical microscopy, due to the very refined structure formed in the SZ and the poor grain boundary contrast in all zones of the joint. Substructural evolution TEM investigation of thin foils obtained from different welded zones and base material of Al-Mg-Sc joints is the most adequate technique to observe the grain morphology and its substructure. It allows a deep analysis of the variation of the structural features in the friction stir welded joints. In this section, the investigation of the Al-Mg-Sc alloy joined with different energy inputs is presented in a logical sequence, in order to provide a better understanding to the reader. Initially, the grain structure aspects across the FSW zones and BM are presented in order to reveal the grain structure evolution, which is difficult to elucidate either by scanning electron microscopy or by optical microscopy. Subsequently, the deformed microstructure and
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the mechanisms of recrystallisation and recovery are presented in terms of dislocation features. Next, the precipitation features are presented for the BM and FSW zones with further correlation with mechanical testing performance and thermal cycles produced by the different energy inputs applied. Finally, strengthening predictions are exploited, in order to understand the material strengthening attributed to the precipitation phenomenon. Figure 11.4 shows micrographs representing the different grain structures observed in the BM, TMAZ and SZ of the Al-Mg-Sc joints. In the micrograph presented in Fig. 11.4 (a), a fine pancake grain structure, as a result of the rolling process during the manufacture of the plates, is clearly observed for the Al-Mg-Sc alloy BM. The grain size refined by the Sc addition is significantly smaller when compared to other structural aluminium alloys. The TEM analysis established an average grain length of 2.6 mm and average width of 0.5mm. The effect of scandium in BM grain refinement is to inhibit the recrystallisation process during thermo-mechanical processing (Toropova et al. 1998; Lathabai and Lloyd 2002; Davydov et al. 2000). Figure 11.4 (b) shows a grain morphology combination, which characterises the TMAZ. TEM investigation of this zone allowed observation of the detailed grain structure, which includes a pancake grain structure remaining from the BM, distorted and rotated, combined with some equiaxed grains originating from a mixture of recovered and partially recrystallised structure, resulting from the plastic flow and thermal cycle during FSW. The SZ of the Al-Mg-Sc joints is characterised by only a very fine equiaxed grain structure, as presented in Fig. 11.4 (c). In this zone, dynamic recrystallisation controls the resulting grain structure, having reached the highest temperatures and deformation levels during the FSW process. The average grain diameter in the SZ of the Al-Mg-Sc, established from TEM micrographs for the different FSW energy inputs, are about 2 mm for high HI (175 mm/min), 1,4 mm for intermediate HI (350 mm/min) and roughly 1mm for the low HI (525 mm/min). It is expected that the presence of stable Al-Sc precipitates would inhibit the motion of thermally activated subgrain boundaries under the elevated temperatures promoted during the FSW process, as occur in the SZ (Riddle and Sanders 2000). The recrystallised grain structure supplies a superior elongation to the SZ without significantly affecting its strength and hardness values. In the Al-Mg-Sc FSW joints, a relevant dislocation density presenting singular aspects can be observed along the base material and subsequent weld zones. Typical dislocation features of the different weld zones and base material of Al-Mg-Sc alloy are represented in the micrographs of Fig. 11.5. TEM examinations of welds made with different energy inputs usually show no notable variation in dislocation features. However, the highest dislocation densities are expected to be introduced using lower energy input, since the resulting peak temperatures produced are lower, hindering the dislocation annihilation and rearrangement. The rolled BM present a considerable
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Friction stir welding
2 µm (a)
1 µm (b)
11.4 Bright-field TEM micrographs of the grain structure present in the base material and different FSW regions of Al-Mg-Sc alloy (a) BM, (b) TMAZ and (c) SZ.
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1 µm (c)
11.4 Continued
dislocation density with no order at all, except some particular areas where dislocations presented groups, as indicated in Fig. 11.5 (a). The dislocation pinning by second phase particles can be also evidenced in the BM and weld zones, as indicated in Fig. 11.5 (a). In the TMAZ (Fig. 11.5 (b)), dislocation features show a similar character to the BM. However, the dislocations here indicate more interaction with one another and with second phase particles. The dislocation features most often found in TMAZ include dislocation pileups, forming subgrain boundaries, dislocation tangles (network structure) in some regions, and looping around second phase particles. Dislocation density can be found at a lower level for grains in the SZ, as shown in Fig. 11.5 (c). Pile-up of dislocations and interaction with second phase particles were evidenced, mainly for the higher weld energy inputs, where the higher temperatures favour dislocation mobility. It should be considered, that even being in the SZ, where deformation during FSW is highest, the elevated temperatures and dynamic recrystallisation effects in the SZ usually result in a final microstructure presenting inferior dislocation density levels compared to the BM and recovered zones (Genevois et al. 2005). Such an effect is to be expected, because of the dislocation annihilation resulting from the recrystallisation process, and the atoms, diffusion is helped by the high peak temperature in this zone. However, in FSW, a continuous recrystallisation process (dynamic) occurs, accompanied by incessant deformation, introducing
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Dislocation grouping
Dislocation pinning
100 nm (a)
200 nm (b)
11.5 Bright-field micrographs from Al-Mg-Sc alloy, showing typical dislocation features in (a) BM in [111]Al (g = 202), (b) TMAZ in [001]Al (g = 200) and (c) SZ close to [011]Al zone axis.
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500 nm (c)
11.5 Continued
and constantly increasing the dislocation density at the same time as new grains, free of deformation, are formed. Additionally, certain particular substructural features are encountered at the different weld zones. Figure 11.6 (a) shows a bright field micrograph of BM, showing pancake-shaped grains characterised by contrast of parallel bands crossing the grain in transverse direction, as indicated by the arrows, which are too large to be associated to stacking faults. Therefore, the high stacking fault energy, which is typical in Al alloys, promotes dislocation climb and cross-slip occurrence, which is characteristic of the dynamic recovery phenomenon (Lathabai and Lloyd 2002). The presence of dislocation walls, providing evidence of the initial stages of subgrain formation, can be found across the weld zones and base material as a product of recovery stages, in particular in the TMAZ, as illustrated by the micrograph of Fig. 11.6 (b). A high density of nearly spherical, most likely Al3Sc precipitates, is evidenced for all weld zones and the BM of the FSW joints. The main precipitation characteristics of the BM are presented in Fig. 11.7. Figure 11.7 (a) shows round type precipitates in nanometre scale, disposed along a-Al elongated BM grain, in conjunction with scarcely any dislocations in contrast. The presence of Al3Sc precipitates can be confirmed by selected area diffraction pattern (SADP) analysis, elucidating the superlattice spots,
330
Friction stir welding Parallel bands
100 nm (a)
Sub-grain formation
100 nm
200 nm (b)
11.6 Bright field micrographs of Al-Mg-Sc alloy showing particular substructural aspects (a) parallel bands in BM pancake grain, and (b) initial stages of subgrain formation in TMAZ of WP4 condition.
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200 nm (a)
200 220
020
[100] (b)
11.7 TEM analysis presenting precipitate features in the BM (a) micrograph showing precipitates appearance and distribution along pancake grains, (b) SADP in [001]Al zone axis, (c) EDS peaks from precipitates presented in the micrograph.
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Friction stir welding 100
AI
Atomic %
80 60 40 20 Sc 0 0
2
4
6 8 Energy (keV)
10
12
14
11.7 Continued
due to the presence of Al3Sc ordered precipitates in [001] zone axis of BM, indicating that an orientation relationship exists with a cubic a-Al matrix, as shown in Fig. 11.7 (b). From these observations can be assumed that precipitates are coherent with an Al matrix, if the size of precipitates and lattice constant mismatch with the matrix are not too large, as is the case with Al3Sc. The intermetallic Al3Sc precipitates have the same FCC crystalline lattice as a-Al, with a lattice parameter of 0.4106 nm, differing only 1.3% from the Al matrix (Toropova et al. 1998; Blake and Hopkins 1985; Milman et al. 2000). The presence of Sc-rich particles can be observed by EDS analysis, as shown in the peaks of the diagram presented in Fig. 11.7 (c). Sc-rich precipitates presenting apparently similar size and morphology can be also observed in TEM analysis of the SZ, as represented in Fig. 11.8. In the micrograph of Fig. 11.8 (a), dislocations ordered and pinned by the precipitates are shown, which infers an effective role of the Al3Sc precipitates in order to inhibit recrystallisation and also acting as precipitate strengthener. According to Seidman and co-workers (2002), coherency strain (Ashby-Brown contrast) in bright-field TEM images is one of the indications that precipitates are coherent with the a-Al matrix. The strain-field contrast (SFC), due to the coherent interfaces is also evidenced in the present work for most of the bright-field micrographs and it can be characterised by a lack of contrast, usually crossing the precipitate centre, as can be seen better in some precipitates indicated by the arrows as “PptSFC” in Fig. 11.8 (a). Some authors also refer to this phenomenon as coffee-bean-like stress contrast of coherency (Marquis and Seidman 2001; Lenczowski et al. 2000; Yin et al. 2000). Marquis and Seidman (2001) reported loss of coherency by precipitates over 40 nm diameter, which are typically characterised by interfacial dislocations around the coarse precipitate. Similar contrast
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333
Ppt^
PptSFC
200 nm (a)
200
200
220
[011] (b)
11.8 TEM analysis presenting precipitate features in the SZ (a) micrograph showing precipitate characteristics and interactions with dislocations in equiaxed grain, (b) SADP in [011]Al zone axis, (c) EDS peaks from precipitates presented in the micrograph.
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Friction stir welding 100
Al
Atomic X
80
60
40 Sc
20
0 0
2
4
6 8 Energy (keV) (c)
10
12
14
11.8 Continued
of dislocations is evidenced in Figure 11.8 (a), indicated by Ppt^. SADP spots in the [011] zone axis, presented in Fig. 11.8 (b), also evidenced the presence of Al3Sc precipitates, which agrees with investigations by Blake et al. (1985), who verified parallel orientation relationship for rod-type primary Al3Sc precipitates with a-Al in the [011] direction. Moreover, Kendig and Miracle (2002) reported that the presence of non-primary particles precipitated from supersaturated solid solution was confirmed for superlattice spots of ordered Al(ScZr) precipitates in the [011] zone axis, having a cube-on-cube orientation relationship with an a-Al matrix. Additionally, a good relation of coherency between second phase particles and a-Al matrix was also encountered in these studies of Al-Mg-Sc-Zr alloys. According to their (Kendig and Miracle 2002) estimation of precipitate average size diameter and precipitates disposition from TEM micrographs and strengthening predictions, it indicated a significant enhancement of strength due Al3(Sc,Zr) precipitates. Scandium peaks verified in EDS analysis of precipitates in the SZ are shown in Fig. 11.8 (c). Such investigations show a high stability of Al3Sc particles, which even under a high thermo-cycle remained present, dispersed in the a-Al matrix apparently without severe coarsening and dissolution effects. However, studies of Huneau et al. (2005) suggest that the original Al3Sc particles might have been dissolved during FSW and nucleated again just after recrystallisation, due to the low solubility of Sc in Al. Arguments of such nature are not sustained by the results reported here, when considering the precipitate morphology and disposition features, allied to short time at peak temperatures, to dissolve the very stable Al3Sc precipitates.
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b
The Orowan’s dispersion hardening is suggested to be taking place because of the dislocations looping around the precipitates, which are frequently observed, particularly for high heat input joints, in the TMAZ. The apparent Orowan’s mechanism in the TMAZ of an Al-Mg-Sc joint is presented in Fig. 11.9, indicating the dislocation motion through the Burgers vector. As mentioned before, precipitates containing Sc can be found across the BM, HAZ, TMAZ and SZ of mostly FSW joints. Hence, precipitation features in the BM and the different FSW zones are represented by the micrographs shown in Fig. 11.10. The micrographs obtained from the FSW zones were selected from high heat input condition, which are considered the most critical in terms of high peak temperature and slow cooling rates, with consequent influence on the precipitation features. The mostly round type Al3Sc precipitates are elucidated in the elongated structure of grains and subgrains of BM, as showed in Fig. 11.10 (a). High contrast of Al3Sc precipitates with the same characteristics can be also encountered, often almost fully covering the recovered grain structure of the TMAZ, as showed in Fig. 11.10 (b). These precipitates are also observed in the equiaxed recrystallised grain structure of the SZ, showing no significant variation in size, morphology and distribution, as illustrated in Fig. 11.10 (c). It should be considered that even the high temperature levels attained in TMAZ (≥ 327°C) and SZ (≥ 372°C) were not enough to dissolve or to coarsen the very stable Al3Sc precipitates. The TEM analysis of Al-Mg-Sc joints across BM and FSW zones reported a high density of thin precipitates, randomly distributed along the a-Al matrix, which implies great performance in terms of precipitation strengthening.
1000 nm
11.9 Orowan’s mechanism occurring in TMAZ of hot parameter (WP4 joint), in [011]Al zone axis, g = 200.
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200nm (a)
200nm (b)
2mm
200nm (c)
11.10 TEM bright-field micrographs, representing Al3Sc precipitates in different FSW zones and BM indicated in the macrograph of a AlMg-Sc joint (a) BM, (b) TMAZ and (c) SZ.
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In order to understand the effects of the FSW input parameters on the resulting microstructure and their corresponding influence on mechanical properties of the Al-Mg-Sc joints, some aspects have to be considered. It has been verified that even microstructural variation (evidenced by different microscopy analysis techniques) produced by high deformation and thermal cycle levels does not have significant influence on the hardness values (Olea 2008). The presence of Al3Sc precipitates observed by TEM analysis in all FSW zones, showing no significant variation in size, morphology and distribution, indicates that precipitation hardening is controlling the mechanical behaviour of the Al-Mg-Sc FSW joints. The effect on local mechanical properties as observed in microflat tensile tests (Olea 2008) from different weld zones and BM also agree with this conclusion. In this case, a significant variation of elongation was found for samples from the SZ. However, this behaviour can be attributed to the equiaxed grain microstructure, resulting from the FSW process, which increases ductility and is characteristic of superplasticity resulting from the recrystallisation phenomenon. It is also noted that the increase in ductility can be controlled by means of welding process parameters. It can be assumed that the smaller equiaxed grain size in the SZ reduces the superplasticity effect from recrystallisation. Evaluating Al3Sc precipitate aspects from BM, the presence of a large population of precipitates is observed, in a range from 6 to 21 nm diameter, especially a major proportion of precipitates with a diameter range between 9 and 15 nm. In general, the weld zones presented relative minor variations in Al3Sc precipitates size and population. Compared to the BM, the high HI condition exhibit an increase in size for the major amount of precipitates to a range of 12–18 nm in TMAZ and SZ, as demonstrated in Fig. 11.11. In this case, the higher temperatures and slower cooling rates produced by the low weld speed seems to affect the precipitates from both weld zones in similar proportion. For the lowest HI condition, it is also verified that there is no precipitate coarsening due to the thermal cycles. The major amount of precipitates is found in a size range of 6–15 nm for TMAZ and between 9 and 15 nm for SZ. The lower energy input weld tends to produce no influence on precipitate coarsening, even in the SZ, where peak temperatures are higher. Several researchers (Toropova et al. 1998; Marquis and Seidman 2001; Seidman et al. 2002; Venkateswarlu et al. 2004) have studied Sc added alloys and reported the high stability of supersaturated solid solutions and coherent precipitates formed from its decomposition. Marquis and Seidman (2001) reported the very stable behaviour of Al3Sc precipitates with respect to coarsening, even for long aging times. It was described that Al3Sc precipitates, with sizes in a range of 4–10 nm from Al-0.5 wt% Sc alloy aged above 400°C, results in coarsening of precipitates, which tends to lose coherency when their size exceeds about 20–30 nm. It also agrees with studies from Milman and
Friction stir welding 8 7
Base material
Distribution [%]
6 5 4 3 2 1
60–63
54–57
48–51
42–45
36–39
30–33
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Ppt diameter range [nm]
8 7 Distribution [%]
6
High HI – TMAZ
5 4 3 2 1
60–63
54–57
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30–33
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0–3
0
Ppt diameter range [nm]
8 7 6 Distribution [%]
High HI – stir zone
5 4 3 2 1
66–69
60–63
54–57
48–51
42–45
36–39
30–33
24–27
18–21
12–15
6–9
0–3
0
338
Ppt diameter range [nm]
11.11 Precipitate (Ppt) size and distribution of the Al-Mg-Sc FSW joints using three different input parameters (high HI – WP4, intermediate HI – WP2 and low HI – WP1.33) along TMAZ, SZ and BM.
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8 7
Distribution [%]
6
Low HI – TMAZ
5 4 3 2 1
60–63
54–57
48–51
42–45
36–39
30–33
24–27
18–21
6–9
12–15
–1
0–3
0
Ppt diameter range [nm]
8
Distribution [%]
7 6
Low HI – stir zone
5 4 3 2 1
60–63
54–57
48–51
42–45
36–39
30–33
24–27
18–21
6–9
12–15
0–3
0
Ppt diameter range [nm]
11.11 Continued
co-workers (2000), which showed precipitates with sizes from 3 to 15 nm produce a very strong hardening effect, also growing and losing coherency at temperatures over 400°C. During the FSW process, the critical temperatures are attained in the SZ, and temperatures very close to this zone are around 372°C, as is the case of the high HI condition shown previously. In the SZ, temperatures can be expected up to 200°C higher than these (over 550°C). The high peak temperatures in the SZ and practically no variation in the precipitates size and distribution imply that the thermal cycle is too fast to provoke any instability in the Al3Sc precipitates. It is important to emphasise that the cooling rates could have been different at the SZ, where thermal records could have been affected by changes in position, in relation to the core of the material heating during FSW. The results presented in this section, however, show that the obtained thermal cycles were not sufficient to produce a significant precipitate dissolution and coarsening effect, contrary to the studies by Huneau
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Friction stir welding
et al. (2005). The low solubility of Sc in Al is suggested as responsible for the insignificant size variation of the Al3Sc precipitates, therefore with all the Sc content precipitated as Al3Sc, the low Sc solute a-Al matrix cannot contribute directly to precipitate coarsening. These considerations lead to the assumption that Al3Sc precipitates, which remain in a size range of 6–18 nm (the great majority) are primarily responsible in ensuring the strengthening of the FSW zones in the studied Al-4.3wt% Mg-0.27wt% Sc alloy. Figure 11.11 shows the precipitate size range, related to their respective distribution for Al-Mg-Sc FSW joints in TMAZ and SZ of the three different weld energy inputs, in comparison with values measured in BM. The contributions of microstructural features in Al-Mg-Sc alloys can be basically classified in three different ways: solid solution, grain boundary and precipitate strengthening. The solid solution strengthening is assumed as being basically produced by Mg that is present in 4.3wt% in this alloy. It is known that Mg can also be present as Mg2Al3 particles, but EDS analysis have shown the presence of scarcely coarse particles presenting Mg peaks, which according to size and distribution would not contribute to the alloy strengthening. However it is believed that the majority of Mg content remained in solution, even in weld zones. As commented before in Section 2.4, Sc has very low solubility in Al (about 0.05% at 500°C) and due to the temperatures attained in the PWHT and FSW process, the amount of Sc in solution in Al-Mg-Sc joints is expected to be very low, producing no significant contribution to the alloy strengthening by solid solution (Kendig and Miracle 2002). The grain boundary strengthening contribution in this case can be estimated using the Hall-Petch model (Toropova et al. 1998), which considers the intrinsic resistance of the lattice to dislocation motion added to the dependence of the grain size. In the case of optical microscopy, recrystallised grains in SZ are too small and in the SEM backscattered images, the grain boundary is usually not sufficiently revealed to generate any reliable results. In TEM micrographs, the grain structure can be revealed, as shown in Fig. 11.4, but the number of grains is generally insufficient to generate accurate quantitative results. However, equiaxed grains in a range from 2 to 1 mm has been already evidenced, which indicates a very fine grain structure and so capable of producing a significant contribution to strengthening. The strengthening contribution attributed to fine Al3Sc precipitates is expected to act through two different strengthening mechanisms: anti-phase boundary (APB, particle shearing) or the Orowan (dislocation looping) mechanism. These strengthening mechanisms consider the resistance to dislocation motion produced by the material due the presence of fine precipitates. In Al-Mg-Sc alloy, fine Al3Sc precipitates are characterised by an ordered structure, and the APB mechanism describes the resistance of these precipitates to dislocation motion by formation of an anti-phase
Metallurgy and weld performance in friction stir welding
341
boundary, which results from shearing of these precipitates. The determination of dislocation motion by looping of dislocations around the precipitate, given by the Orowan model, is also applicable in this case and it scarcely depends on inter-particle spacing. For small-sized precipitates, the shearing of precipitates by dislocations is energetically more favourable than looping around the precipitates. However, as the precipitate size increases, shearing becomes difficult and it results in increased strengthening promoted by the precipitates. Upon reaching a critical precipitate size, the easiest way for a dislocation movement during interaction with a precipitate is looping around the precipitate. For a fixed volume fraction of precipitates, if precipitate size increases, the strengthening tends to decrease, but the looping around the precipitate becomes easier. It is worth noting that the optimum precipitate size in terms of alloy strengthening is found at the transition from the shearing to the looping mechanism (Seidman et al. 2002; Kendig and Miracle 2002). Kendig and Miracle (2002) reported precipitate sizes for optimum strengthening for 1.5, 2.5 and 3.5% volumetric fractions as between 20 and 25 nm, as illustrated for the region close to the peaks, where the curves intersect in the diagram of Fig. 11.12. It agrees with studies by Marquis and Seidman (2001), who indicate that precipitates remain coherent in a diameter field of 20–30 nm, and, as previously mentioned, the coherency loss is detected by the presence of interfacial dislocations, when the precipitate diameter reaches about 40 nm. Other authors suggest that Al3Sc precipitates 400
Strengthening increment [MPa]
APB vol% = 1.5 350
APB vol% = 2.5
300
APB vol% = 3.5 Orowan vol% = 1.5 Orowan vol% = 2.5
250
Orowan vol% = 3.5
200 150 100 50 0 0
10
20
30 40 50 60 70 Precipitate diameter [nm]
80
90
100
11.12 APB and Orowan strengthening mechanisms due to resistance to dislocation movement, predicted for 1.5, 2.5 and 3.5% volume fraction (Kendig and Miracle 2002).
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Friction stir welding
cause very strong hardening in a size range from 3 to 15 nm diameter and tend to coarsen slowly with increasing temperature (Seidman et al. 2002; Riddle and Sanders 2000; Milman et al. 2000). Measurements of Al3Sc precipitates volume fraction from TEM micrographs for base material and welded zones has been found in a range from 3 to 4.5%. The highest volumetric fraction values are found in welded zones, showing no significant changes, but decreasing in the BM. By observing values of Al3Sc precipitate size, illustrated in the diagrams of Fig. 11.11, it can be seen that the majority of precipitate sizes are in the range 6–21 nm diameter for base material and different weld input parameters. According to the predicted diagram of Fig. 11.12, for an alloy Al-0.27%wt Sc, the precipitates are in a range (indicated by lines in the diagram of Fig. 11.12) where the APB mechanism is the main contribution responsible for strengthening. Precipitates coarser than the critical size are also evidenced in the diagrams of Fig. 11.11, hence indicating that Orowan mechanism is also acting, but at a minor level, as verified by the TEM analysis (Fig. 11.9). From the considerations mentioned above, it can be affirmed that even with both strengthening mechanisms acting, for the size range of Al3Sc precipitates, the precipitate shearing mechanism is predominantly responsible for the material strengthening. Summary In the Al-Mg-Sc alloy, a high distribution of Al3Sc precipitates with spherical morphology is observed along the TMAZ, SZ and BM for different weld energy inputs. The reduced coarsening effect in the TMAZ and SZ is attributed to the low scandium solubility in aluminium in addition to low diffusion coefficients and fast thermal cycle during FSW. The temperatures attained during the FSW process of Al-Mg-Sc alloy as presented in this section were not sufficient to promote solute diffusion in order to produce dissolution and significant coarsening of the very stable Al3Sc precipitates. Precipitates have been shown to control the mechanical properties and they retained a good stability level along the different weld zones. Predictions of strengthening based on the established values of volume fraction and precipitate diameter, indicate that the APB mechanism (precipitate shearing) is mainly responsible for the strengthening contribution of Al3Sc in Al-Mg-Sc alloys friction stir welded. There is a wide operating window for the production of satisfactory joints concerning the mechanical properties (tensile and hardness properties) of friction stir welded Al-Mg-Sc joints. The precipitation stability indicates that Al-Mg-Sc is a robust alloy for FSW applications, showing a wide range of process parameters to be exploited in order to produce high quality joints with the highest process efficiency.
Metallurgy and weld performance in friction stir welding
343
11.2.3 Al-based dissimilar welds Motivation In recent years, transportation systems increased the use of light-weight materials, such as aluminium alloys (Al alloy), magnesium alloy (Mg alloy), and high strength steels (HSS) for seeking energy saving and pollution reduction. The combination of different materials into a hybrid material system, e.g. car bodies and aircrafts, is the driving force for the development and improvement of joining techniques. However, it still presents a challenge, from the point of view of both scientific interest and industrial application. The differences between the melting temperatures and physical and mechanical properties of two metals lead to complexities in weld pool shape, solidification microstructure and segregation patterns. Hence, different welding energy sources have been developed and improved to further widen the field of application of these materials. In this context, friction stir welding (FSW) appears to be a suitable and promising welding process to join dissimilar materials resulting in the microstructure improvement, since no metal melting is involved (solid-state process). In general, FSW has been found to produce a low concentration of defects associated with cooling from the liquid phase, e.g. porosity, solidification crack, etc. Investigation on the microstructure formation of friction stir dissimilar welding having Al alloy joined to Mg alloy and to HSS is the topic of this section. General aspects concerning joining techniques for dissimilar welding involving those materials and their problems will first be explained. This is followed by an extensive presentation with an emphasis on microstructure examination explaining the microstructure formation, presence of intermetallic phases on the interface region and evaluations of some mechanical properties. The analysis encompasses optical microscopy (OM), scanning electron microscopy (SEM) and electron backscattered diffraction (EBSD), transmission electron microscopy (TEM), tensile tests and shear tests. General aspects The introduction of Al alloy parts into a steel car body necessitates the development of reliable and cost-efficient joining methods. The dissimilar welds between Al alloy and steel, due to the extremely low solubility of Fe in Al, promotes the formation of Al-rich FexAly intermetallic phases (Kubaschewski 1982) known to be brittle and thus detrimental for the mechanical properties of the joints. Mg and its alloy, being the lightest available construction metals, are increasingly applied in replacement of Al alloys and in combination with it in specific structural applications (Avedesian and Baker 1999; Sanders et
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Friction stir welding
al. 1999; Leong et al. 1998). The use of such light-weight metals in hybrid structure cars and other engineering applications shows that the intermetallic layer formed near interface between the metals significantly degraded the joining strength (Borrisultthekul et al. 2005). Force-fit and form-fit joints are obvious technical solutions for assembling dissimilar materials, but they are often disadvantageous with respect to joint geometry, optical appearance of the joints and in particular regarding assembly costs. Early attempts at fusion welding steel to Al alloys relied on preventing the formation of brittle FexAly intermetallic phases by using a buffer layer (Hartwig 1981; Bel’chuk and Proizvod 1963; Bach et al. 2003). More recently sound fusion welds between steel and Al alloys have been achieved by minimising the heat input during welding and thus limiting the thickness of the intermetallic phase seam to a few micrometres (Kreimeyer et al. 2004; Mathieu et al. 2007; Peyre et al. 2007; Agudo et al. 2007). Another option to limit the heat input into the joint is the use of pressure welding techniques (Hartwig 1981; Iwamoto et al. 1975; Rathod and Kutsuna 2004; Lee et al. 2003; Fukumoto et al. 1999, 2000; Kobayashi et al. 2003; Thomas et al. 1992). Solid-state welding processes naturally avoid solidification and, thus, have particular advantages for joining dissimilar materials with very different melting intervals or materials prone to hot cracking (Lima et al. 2003; Sato et al. 1999). Among pressure welding methods, friction welding has gained increasing application within recent years, which is mostly a result of the development of the FSW technique (Thomas et al. 1991; Dawes and Thomas 1996; Rhodes et al. 1997b; Mishra and Ma 2005). Friction stir dissimilar welding Similar FSW of Al alloy and Mg alloys have received a lot of interest in science and technology and have been intensively studied (Lima et al. 2003; Sato et al. 1999; Rhodes et al. 1997b; Mishra and Ma 2005; Li et al. 1999; Reynolds 2003; Zettler et al. 2006a,b, 2005. Tan and Tan 2003; Watanabe et al. 2001,1999; Kim et al. 2001; Somehawa et al. 2005; Woo et al. 2006; Olea et al. 2007). Recently, friction stir dissimilar welds of Al alloys to Mg alloys have successfully been produced (Zettler et al. 2006a; Khodir and Shibayanagi 2007; Kostka et al. 2009; Hirano et al. 2003; Park et al. 2002; Rodrigues et al. 2006). However, it has been shown that the formation of brittle AlxMgy intermetallic phases generally leads to poor mechanical properties of the joint. Compared to similar FSW of Al alloy and Mg alloys, FSW of steel imposes much higher demands on the tool material. The feasibility of FSW of steels was demonstrated as early as 1999 (Thomas et al. 1999), but appeared not to be cost effective at that time, due to strong tool wear. In recent years,
Metallurgy and weld performance in friction stir welding
345
different concepts for tool materials increased the interest in FSW of steels as well as of steel-based dissimilar joints (Lienert et al. 2003; Reynolds et al. 2003; Sterling et al. 2003; Sato et al. 2007). It has been shown that friction stir dissimilar welds of Al alloys to steels can be produced (Fukumoto et al. 2000; Chen and Kovacevic 2004; Uzun et al. 2005; Lee et al. 2006; Coelho et al. 2008a,b) and that the comparatively low heat input in friction welding and FSW reduces or even suppresses the formation of brittle intermetallic phases (Sundarasan and Murti 1993; Fukumoto et al. 1998,1997). Neither, Al-Mg or Al-steel friction stir dissimilar joints have been fully characterised and understood in terms of microstructure formation. Here, some examples on successful joints Al-Mg and Al-steel are presented and investigated in terms of microstructure and intermetallic phases formation and mechanical properties. Examples Al Alloys to HSS As already mentioned in the section before, friction stir dissimilar joints between Al alloy and steel required a lot of effort on tool development and welding setup. The tool material is essential for protecting it against tool wear and the welding setup to reduce the formation of brittle Al-rich FexAly intermetallic phases. Here, successful dissimilar joints between Al alloy and HSS were produced in two different configurations: butt-joint and overlap. All main welding parameters are presented in Table 11.2. The welding setups applied here ensures that the pin intrudes mainly into the Al alloy side or very slightly into the HSS. Therefore the tool wear Table 11.2 Weld parameter details
Butt-joint
Overlap
Materials Al alloy HSS
AA6181-T4 DP600
AA6181-T4 HC340LA
Welding parameters Travel speed (mm/s) Rotation speed (rpm) Down-force (kN)
8.0 1600 5.0
6.0 1600 5.5
Tool properties Pin diameter (mm) Shoulder diameter (mm) Pin length (mm) Pin offset Dy (mm) Tool material
5.0 13 1.35 1.0 WRe25
5.0 13 1.6 – UHB Marax ESR (1.6358)
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Friction stir welding
and the amount of steel particles usually detached and transported into the softened Al alloy can be controlled. These hard steel particles (fine grained microstructure) would act as stress concentration and directly influenced the strength of the joint (easy crack propagation). Butt-joint configuration: AA6181-T4 Al alloy to DP600 HSS (Coelho et al. 2008b) Figure 11.13 show the welding setup applied in this study. The deformation produced by the shoulder and by the Al alloy flow is responsible for the HSS deformation and ensures the bonding observed at the interface by shear strain and frictional heating between the two materials. A typical microstructure of the weld cross section analysed here is presented in Fig. 11.14. The results show a good quality weld containing neither pores nor cracks. There were a small number of HSS particles (arrows) that had been detached and transported into the Al alloy. Neither the surface of the weld nor its cross-section showed any macroscopic defects. The welding parameters chosen (tool offset position) meant that severe plastic deformation occurred mainly in the Al alloy since the pin barely moves into the HSS. Consequently, no evidence of mixing between the HSS and the Al alloy was observed. The thermo-mechanically affected zone (TMAZ) on Axial force
Retreating side Al alloy 1.5 mm
Advancing side HSS 1.5 mm
11.13 Schematic illustration of the friction stir dissimilar joint setup used to join the Al alloy to HSS in a but-joint configuration.
Al alloy
250 µm
11.14 Cross-section view of the different welded zones.
HSS
Metallurgy and weld performance in friction stir welding
347
the HSS side was no wider than 300 mm (Fig. 11.15a). A complex interface is formed between the HSS and the Al alloy characterised by fine equiaxed a-Fe (ferrite) grains and small numbers of thin strips of intermetallic FexAly (a)
Al alloy
HSS
100 µm
(b) Al alloy
HSS
1 µm
(c) Al alloy
Fe2Al5
HSS 100 nm
11.15 SEM micrographs showing the weld interface (a) and a close view of the bonding between both materials (b). TEM investigations conducted on the interface (c) as well on the detached HSS particles (d).
348
Friction stir welding (d) Al alloy
HSS
1 µm
11.15 Continued
compounds (Fig. 11.15b). A close analyses by TEM of this regions revels the presence of Fe2Al5 as the main present intermetallic compound (Fig. 11.15c). The shear strain and the frictional heating support the metallurgical phenomena (chemical reaction and diffusion between Al alloy and the HSS) leading to the bonding observed at the interface. The microstructure analyses of the detached HSS particles show the same characteristics observed on the interface between both materials – fine a-Fe grains and thin strips of intermetallics (Fig. 11.15d). A tensile test was conducted in order to assess the mechanical properties of the joint. The results, presented in Fig. 11.16a, show that the joint stressstrain curve follows the Al alloy curve, which suggests that the test was performed mainly on the Al alloy side. The maximum tensile strength does not reach the yield strength of the HSS and the plastic deformation occurs only in the Al alloy side. The strong grain size gradient shown on Fig. 11.16b originates from the fact that each region of the joint (base material (BM), heat affected zone (HAZ), thermo-mechanically affected zone (TMAZ) and stir zone (SZ)) was exposed to different thermo-mechanical treatment: the gradual accumulation of strain accompanied by the frictional heat lead to the nucleation and growth of new grains during ongoing deformation process (non-homogeneous dynamic recrystallisation). While accumulated plastic deformation increases the mechanical strength of the TMAZ and SZ, the HAZ which grains reveal relatively small defect densities (concentration of dislocations and sub-grains) will control the mechanical behaviour of the joint as the weakest elements of the structure. Figure 11.16c shows that the fracture occurs on the interface Al alloy BM-HAZ-TMAZ. The results suggest that the small amount of Fe2Al5 intermetallic phase observed on the interface Al alloy/HSS promotes the bonding between both materials and do not directly influence the mechanical properties of the joint.
Metallurgy and weld performance in friction stir welding
349
The non-smooth interface observed in Fig. 11.15a and 11.15b appear to play an important role for the joint acting as a mechanical interlock between both components. Overlap configuration: AA6181-T4 Al alloy to HC340LA HSS (Coelho et al. 2008a) The joints produced in overlap configuration follows the idea of the welding setup presented for the butt-joint configuration. The schematic welding setup is presented in Fig. 11.17. The tool material, the pin length and the welding parameters (Table 11.2) were carefully chosen in order to avoid the wear between the HSS and the tool and to control the amount of detached HSS particles transported into the Al alloy. As already mention 700
Stress [MPa]
600 500
HSS Joint
400
Al alloy
300 200 100 0
0
2
4
6
8 10 12 14 16 18 20 22 24 26 28 Strain [%] (a)
(b) 5 µm
70 µm
BM-H face Inter
AZ-T
MAZ
50 µm
11.16 Tensile stress-strain curves of both base materials involved (Al alloy and HSS) and of the joint produced by FSW (a). Macrograph showing the grain size gradient in the critical area before deformation (b). Fractography analysis conducted on the top area of the failed specimen (c).
350
Friction stir welding (c) Interface BMHAZ-TMAZ
Al alloy SZ
2 mm
HSS
Joint interface
11.16 Continued Axial force
Retreating side
Advancing side
Al alloy 1.5 mm HSS 1.5 mm
11.17 Schematic illustration of the friction stir dissimilar joint of Al alloy to HSS in a single overlap configuration applied in this study.
(in previous example) these HSS particles could act as stress concentration and directly influence the strength of the joint. OM was applied for an overview investigation on the cross section as well on the top surface of the joints. A macroscopic examination of the top side of the overlap joint reveals a good surface quality of the stirred zone and neighbouring areas within the Al alloy. The cross section analysis, presented in Fig. 11.18, reveals neither pores nor crack formation. The conducted investigation also reveals the detached HSS particles transported into the Al alloy (arrows). These particles appear in a small amount but in large sizes (of about ~0.1mm).
Metallurgy and weld performance in friction stir welding Retreating side
351
Advancing side
Al alloy
HSS 1 mm
11.18 Cross-section macrostructure of the friction stir dissimilar joint of Al alloy to HSS in a single overlap configuration applied in this study. Detached HSS particles highlighted by arrows.
SEM micrographs taking from the mixing region highlighted in Fig. 11.18 reveal a complex interface between both materials characterised by strong grain size gradient towards the HSS BM and by fine equiaxed a-Fe grains mixed with thin strips of intermetallic FexAly compounds in the interface (Fig. 11.19a and 11.19b). The intermetallic compounds are present along the entire mixing between the HSS and the Al alloy. The microstructure analysis of the detached particles reveals the same characteristics observed on the interface between both materials. A very similar layered microstructure was described in FSW Al-steel lap-joints (Elrefaey et al. 2005). In FSW, it is generally agreed that thermal energy is generated by contact friction between the shoulder and the base material and through adiabatic heating (i.e., shear processes) around the pin. In the overlap joints examined here, the pin intrudes only very slightly into the HSS. Softened Al alloy is transported towards and pressed against the exposed HSS region rubbed by the pin. High local temperatures and high axial pressures promote an accelerated diffusion process leading to metallic bonding between the joint partners. A part of the softened Al alloy is pushed upwards (counter-flow), due to the continuous downward flow imposed by the rotating threaded pin. As a result, some of the detached steel particles are moved away from the bonding line. The chemical reaction and diffusion between the Al alloy transported into the steel result in the formation of the thin stripes with intermetallic compounds. Seams that this thin intermetallic strips appear to be crucial to the bonding adding the mechanical interlocking between both materials. Figure 11.20a shows the interface between the intermetallic compound and the steel at high magnification in TEM Bragg contrast. The ferrite grains (here in dark contrast) are surrounded by bands (layers) of fine grains of the FexAly intermetallic phases. The grain size of the formed intermetallic compound is always smaller than 50 nm. Selected area diffraction patterns, taken from carefully chosen regions, were used for phase identification. Such measured values correspond to the interplanar distances of a-Fe and Fe2Al5 respectively (Fig. 11.20b and 11.20c).
352
Friction stir welding
40 µm (a)
2 µm (b)
11.19 Microstructure investigations on the mixing area between both materials: the strong grain size gradient (a) and the fine ferrite grains together with thin strips of intermetallics compounds.
In terms of the binary Al-Fe equilibrium diagram (Fig. 11.20d) in the Al-rich corner, the three phases FeAl3 (q-phase), Fe2Al5 (h-phase) and FeAl2 (x-phase) might be expected to be present in the intermetallic phase layer. All these phases in the Al-rich corner do not differ strongly in their thermal stability and their chemical composition. Thus, in joints where layers of intermetallic phases between steel and aluminium are as thin as those encountered here, only TEM diffraction patterns allow reliable phase identification. The TEM investigations revealed that may not all possible FexAly phases are present, since only the Fe2Al5 was successfully detected (Fig. 11.20c). Figure 11.21a shows results of the overlap shear test of the welded joints.
Metallurgy and weld performance in friction stir welding
353
Owing to the non-uniform mixing of FSW joints, the contact area of both materials could not be determined precisely along the entire sheet. For this reason, the overlap shear test results are presented as force (N) versus displacement (mm) curves. The mechanical properties of the samples tested in the as-welded condition compared to the mechanical properties of friction stir welded overlap Al alloy joints reveal an efficiency of about 73% of the dissimilar Al alloy to HSS friction stir dissimilar welds. Fracture of the dissimilar friction stir
a – Fe
Fe2Al5
200 nm (a)
1.17 Å 2.14 Å
2.02 Å
3.79 Å
2.43 Å 1.46 Å
1.43 Å 0.76 Å Fe2Al5
a-Fe (b)
(c)
11.20 TEM micrograph showing formation of intermetallic layers in steel (a); selected area diffraction patterns from steel (b) and intermetallic phase (c). The Binary alloy phase diagrams Al-Fe [1] (d).
354
Friction stir welding 1600 1536°C 1400
1392°C
1200
1.95 g
Temperature, °C
1310°C a
1215°C 1092°C
1171°C 1157°C
e
1022°C
1000
a2(h)
911°C Tc
800
a¢2
a2(I)
612°C 552°C
600
700°C
660°C 652°C 99.1
Fe3AI
400 200
L
FeAl2 1092°C 1164°C Fe2Al5 FeAl3
K1 0
10
K2 20
30
40 50 60 70 Atomic aluminium, % (d)
80
90
100
11.20 Continued
welds always occurred on the retreating side of the Al alloy SZ. The force displacement curves obtained on different friction stir dissimilar welds specimens produced with identical process parameters are very similar, indicating a good reproducibility of the process. SEM analysis of fractured samples reveals that failure started in the mixing SZ on the retreating side and propagated into the Al alloy SZ following the detached HSS particles (Fig. 11.21b). These fine grained HSS inclusions are surrounded by the intermetallic compound and acted as stress concentrations upon load. Therefore the cracks propagate following the inclusions. The amount of steel particles incorporated into the Al-SZ presumably determines the mechanical properties of the entire joint. Al Alloy to Mg Alloy In this sub-section we will focus on the details of microstructure evolution and intermetallic phases formation in friction stir dissimilar joint between Al alloy (AA6040) and Mg alloy (AZ31). Figure 11.22 shows the welding setup applied in this study. The process parameters were as follows: rotation speed: 1400 rpm, travel speed: 3.75 mm/s, axial force: 3.5 kN. Figure 11.23 shows a typical microstructure of the joint transverse crosssection. The upper part of the microstructure is characterised by heavy plastic deformation of the joined components resulting in their thorough mixing (white arrows). In contrast, the lower part of the joint shows a much smoother interface and less intermixing (black arrows). The substantial deformation
Metallurgy and weld performance in friction stir welding
355
3500 3000
Force [N]
2500 2000 1500 1000 500 0 0.0
0.5
1.0 1.5 Deformation [mm] (a)
Retreating side
2.0
2.5
Advancing side
1 mm (b)
11.21 Overlap shear test curves of the different Al alloy to HSS friction stir dissimilar weld samples (a). Fractography analysis revealing the crack propagation following the steel detached particles (b). Axial force
Retreating side Al alloy 2 mm
Advancing side Mg alloy 2 mm
11.22 Schematic illustration of the friction stir dissimilar joint setup used to join the Al alloy to Mg alloy in a butt-joint configuration.
356
Friction stir welding
Al alloy
Mg alloy
250 µm
11.23 Montage of backscattered electron micrographs of the joint transverse cross-section. The arrows highlight the (shoulder flow arm) of the friction stir zone (advancing side of the tool is on the right in the AZ31 alloy).
heating and pressure promoted by the shoulder movement during the welding process, confines the materials within the weld region and endorse the differences in plastic deformation comparing upper and lower joint cross section. Microstructure analysis applying EBSD technique was conducted to compare the BM of Al alloy and Mg alloy with the respective SZ microstructures. The strong plastic deformation in the vicinity of the pin causes dynamic recrystallisation (Dougherty et al. 2003; Cabibbo et al. 2007; Su et al. 2005, 2003; Murr et al. 1997) which result in strong subgrain development and variances in crystallographic orientation of both joined components. The result presented in Fig. 11.24 confirms it. High local temperatures and high axial pressures generated by contact friction between the shoulder and the BM promote chemical reactions and accelerate diffusion, leading to metallic bonding between the partners. A closer look on the interface area reveals that intermetallic compounds occurred along the entire interface between the two alloys (Fig. 11.25). These intermetallics intervene between the unaltered alloys on either side. The material flow imposed by rotating pin was so strong that some of Mg alloy particles moved away from the bonding line (see Fig. 11.25a). The chemical reaction and diffusion between the transported Mg alloy and the Al alloy result in the formation of thin stripes of intermetallic compounds. The intermetallic compounds surrounding the particle appear to prevent its complete dissolution by depressing further diffusion during FSW. Further microstructure analyses of the interface between the two alloys revealed the specific morphology of the intermetallic layer: the interface between the Mg alloy and the intermetallic layer appears smooth and flat, while the interface to the Al alloy is rougher. The thin stripes of intermetallic phase and Mg alloy particles penetrate up to few mm into the Al alloy (Fig. 11.25 and Fig. 11.26a). The thickness of the intermetallic layer is not constant and lies between 0.5 and 1mm (Fig. 11.25b).
Metallurgy and weld performance in friction stir welding (a)
(b)
357
Al alloy SZ
Al alloy BM
50 µm
100 µm
(c)
(d)
Mg alloy BM
25 µm
Al alloy TMAZ
Mg alloy SZ
25 µm
11.24 EBSD image quality maps comparing the BM microstructure of both materials with the respective microstructure formation in the SZ: Al alloy (a) and (b), Mg alloy (c) and (d).
The interface region of the friction stir dissimilar joints is characterised by the presence of strong chemical gradients (Lee et al. 2006) and the reaction products (e.g, intermetallic phases formation) very often do not follow the known equilibrium phase diagrams. Therefore the phase identification based on the energy dispersive-X-ray (EDS) quantitative analysis is unreliable and the intermetallic phases can only be identified by TEM diffraction techniques or X-ray diffraction (XRD) patterns.
358
Friction stir welding Mg alloy
Al alloy
30 µm (a)
Mg alloy
Al alloy 10 µm (b)
11.25 Secondary electron micrograph of the Al alloy and Mg alloy interface: Mg alloy inclusion incorporated into the Al alloy surrounded by intermetallic phase formation (a) and detailed view of the intermetallic compound formed (b).
Figure 11.26 shows in detail the reaction products that occurred at the interface region. The main dominant compound was identified as fine grained Al12Mg17 phase (Fig. 11.26a). Further, small inclusions of a different structure were found in the Al alloy in close proximity to the interface (Fig. 11.26b). These Al3Mg2 inclusions reveal typical microstructure for strongly deformed nano-size grained materials: specific blurred ring electron diffraction pattern and absence of the typical TEM contrast during tilting experiments, which did not allow individual crystallites to be observed.
Metallurgy and weld performance in friction stir welding
359
Al12Mg17 Al alloy Mg alloy
0.5 µm (a)
Al alloy Al12Mg17
0.5 µm
Al3Mg2 (b)
11.26 TEM micrographs of the interface region: fine grained Al12Mg17 intermetallic compound separates the Al alloy from Mg alloy (a) and small nano-size grained inclusions of Al3Mg2 phase adjacent to the Al12Mg17 (b).
It is obvious that the interfacial region of dissimilar joints produced by FSW undergoes complex deformation associated with high heating and high cooling rates. Fine grained microstructures are then expected and in the weld studied here new phases appear as a result of solid-state reactions. The
360
Friction stir welding
mechanical energy input can result in an enhancement of the thermodynamic driving force and the kinetics of mass transport. Yashan et al. (1987) Yashan et al. (1987) suggested that diffusion is enhanced during plastic deformation with a high strain rate when studying inertia friction welding of 1100 Al alloy and 316 stainless steel. Additionally, the high energy accumulated in the defect-rich microstructure produced during FSW may drive phase transformation reactions that do not correspond to the equilibrium phase diagrams. All this suggests that the intermetallic phase formation and further grain size refinement during FSW is a very complex process involving both mechanical alloying and solid-state reaction processes that are associated with the solid state joining of dissimilar materials.
11.3
Corrosion behaviour of friction stir welds in aluminium alloys
The general aspects Corrosion is an electrochemical reaction between a metallic material and its environment. During the corrosion process, electronic charges are transferred in aqueous solutions. The corrosion electrochemical reactions consist of an anodic part, where electrons are liberated from the metal (oxidation), and a cathodic part, where electrons are consumed (reduction). The metallic materials exhibit different forms of corrosion such as uniform corrosion, pitting corrosion, galvanic corrosion, environmentally induced cracking, hydrogen damage, intergranular corrosion, dealloying and erosion corrosion (Jones 1996). Aluminum alloys exhibit a relatively stable protective oxide film when exposed in atmosphere. Nevertheless, aluminum and its alloys are subjected to localised corrosion phenomena when in contact with aggressive solutions, in particular with chloride ions. Localised corrosion features are also present when aluminum alloys are friction stir welded. In this latter case, the main forms of corrosion are pitting, i.e. localised attack on a metal surface, intergranular corrosion, i.e preferential attack along the grain boundaries (Fig. 11.27) and environmentally induced cracking, i.e. brittle fracture in the presence of a load and a corrosive environment (Fig. 11.28). Pitting initiates at an anode (equation 11.1) in the presence of chloride ions and is a common form of localised corrosion in friction stir welds. Within the pit, the metal ions Al3+ are hydrolysed (Buchler et al. 2000; Alodan and Smyrl 1998) (equation 11.2), and a reduction of pH is observed. Chloride ions migrate within the pits to compensate the positive charge, aggressive acid environments are formed and the repassivation is prevented. Al Æ Al3+ + 3e-
11.1
Al3+ + H2O Æ AlOH2+ + H+
11.2
Metallurgy and weld performance in friction stir welding Pitting corrosion
361
Intergranular corrosion
40 µ
2 µ
11.27 The main forms of localised corrosion within the heat affected regions of high strength aluminum alloys friction stir welds (for instance 7075-T651). Environmentally induced cracking Fracture surface
Cracks
Cracks Limit of the fracture surface
500 µ
11.28 The main forms of corrosion: environmentally induced cracking in the heat-thermo-mechanically affected zone of a 7050-T7451 high strength aluminum alloy friction stir weld tested in a 3.5 wt% NaCl solution with a constant extension rate device (Mahoney et al., 1998).
The high strength aluminum alloys exhibit a sensitisation of the microstructure when friction stir welded. The sensitisation of the microstructure, present on both sides of the recrystallised weld nugget, is caused by the temperature increase during friction stir welding (McClure et al. 1998). Consequently, the precipitate-free zone along the grain boundaries and the hardening precipitates are subjected to widening and coarsening (Lumsden et al. 1999; Jariyaboon et al. 2006a; Hassan et al. 2003b; Jata et al. 2000; Rhodes et al. 1997a). This causes an increased corrosion susceptibility. The corrosion takes place predominantly within the heat affected zone (HAZ) of the welds (Hannour et al. 2000b; Ambat et al. 2003; Lumsden et al. 1999).
362
Friction stir welding
Increased corrosion susceptibility is also observed along the nugget interface, within the thermo-mechanically affected zone (TMAZ) (Lumsden et al. 2003; Wadeson et al. 2006; Biallas et al. 1999) or within the nugget (Connolly et al. 2004; Jariyaboon 2006). The decrease in the corrosion resistance of aluminum alloy friction stir welds is generally correlated with a decrease in the breakdown potential of the weld regions, i.e. the nugget and the TMAZ/ HAZ regions (Lumsden et al. 1999, 2003; Jariyaboon et al. 2006b). In some cases, however, the weld region of some friction stir welded aluminum alloys may also exhibit, improved corrosion resistance as compared to the base metal (Hu and Meletis 2006; Zucchi et al. 2001; Squillace et al. 2004; Paglia and Buchheit 2006). High strength aluminum alloys friction stir welds generally exhibit an increased susceptibility to pitting within the heat affected zones of the welds (Fig. 11.29) (ASTM G110 1997). The intergranular corrosion, as it is, for instance, for AA 7075-T651 and 7050-T7451 friction stir welds, is also an important form of corrosion in friction stir welds (Hannour et al. 2000a; Lumsden et al. 1999, 2003; Biallas et al. 1999b; Jariyaboon 2006; Paglia et al. 2003a). This type of corrosion generally correlates with the sensitisation of the microstructure and with the copper depletion along the grain boundaries. The intergranular attack may take place for the aluminum alloys either along the precipitate-free zones (Posada et al. 1997; Yasuda et al. 1990; Galvele and De Michele 1970; Scully et al. 2003) or due to the anodic dissolution of the grain boundary phases (Maitra and English 1981; Andreatta et al. 2004). The potential measurements carried out with an AFM device on a friction stir weld exhibit an interesting inhomogeneous distribution of the Voltapotential between the grain boundary regions and the grain interior of the weld zone. This may indicate a galvanic interaction between these two regions that may contribute to the intergranular corrosion (Paglia and Buchheit 2002). The corrosion susceptibility is often asymmetrical with Parent metal
100 µ
Heat affected zone Advancing side
Nugget
100 µ
100 µ
11.29 Localised corrosion in a 7075-T651 friction stir weld after exposure of the samples to a sodium chloride – hydrogen peroxide solution (Paglia et al., 2003a). Note the increased pitting within the heat affected zone.
Metallurgy and weld performance in friction stir welding
363
respect to the weld centre line. For instance, an asymmetric environmental cracking susceptibility is observed for AA 7075-T651, which correlates with a different Cu depletion along the grain boundaries (Paglia et al. 2003a). The role of copper in controlling the corrosion is also observed by welding dissimilar alloys such as 2024/6065, IS237/7040 and 7449/7040. It appears that, alloys with lower copper content suffer from anodic attack (Mercado et al. 2004; Gerard and Ehrström 2004). The nugget also exhibits intergranular corrosion, but generally not resulting from a sensitisation of the grain boundaries. Nevertheless, the small size of the equiaxed grains, limits the propagation of the corrosion attack (Jariyaboon 2006). The intermetallic particles also play an important role in the corrosion of FSW and are often re-distributed with a higher volume density in the retreating side of the weld (Yang et al. 2004). The intermetallic particles such as Al2Cu are more noble than the aluminum matrix and act as cathode. On the contrary, the intermetallic particles present, for instance in AA 7075 FSW, can be more active than the Al-matrix (Fig. 11.30), and act as an anode (Birbilis and Buchheit 2005). Furthermore, the intermetallic particles often act as initiation sites for the intergranular corrosion. The intergranular attack proceeds along the grain boundary and at a later stage within the grain interior involving the entire grain (Paglia et al. 2003a). The welding parameters largely affect the susceptibility to corrosion (Jariyaboon et al. 2006a; Jariyaboon et al. 2006b). Generally, with low rotation speed welds, the corrosion is generally located within the nugget and within the heat affected zones. With higher rotation speeds, the corrosion is mainly located in the heat affected zones (Jariyaboon et al. 2006a). Depending on
2 µ
10 µ (a)
(b)
11.30 Anodic dissolution of an intermetallic particle in a 7075-T651 friction stir weld after exposure to a 3.5 wt% NaCl solution. b: Note the dissolution morphology with depth after a vertical cross section of the particle carried out with a focused ion beam device (FIB).
364
Friction stir welding
the temperature distribution and the asymmetry of the heat transfer between retreating and advancing side of the weld (Cho et al. 2005), the precipitates may exhibit an overaging, as it is for the heat affected zones, or a dissolution and re-precipitation, as it is for the thermo-mechanically affected zones and the nugget in the case of “hot welds” (Jariyaboon et al. 2006b). Thus, the presence, the size and the population density of the precipitates is controlled by the welding parameters and from the temperature distribution across the weld. This largely controls the location and the extent of the corrosion susceptibility across the weld. An increase in the corrosion resistance can be achieved by applying different methods after or during friction stir welding. Low plasticity burnishing (Jayaraman et al. 2003), laser surface treatments (Davenport et al. 2005), cooling during welding (Rockwell 2004), modification of the welding technique (Thomas et al. 2005), and post-weld heat treatments (Li et al. 2002) are the main techniques used to decrease corrosion susceptibility. Corrosion under fatigue It is known that high strength aluminum alloys are generally susceptible to environmentally induced cracking (Hatch 1984). Interestingly, overaged tempers to the T7 conditions increase environmentally induced cracking resistance, but decrease strength (Speidel 1975; Reboul et al. 1992; Polmear 1989), as compared to the peak aged T6 conditions. Nevertheless, for samples tested under fatigue in a 3.5. wt.% NaCl solution, this increase in cracking resistance is not clearly evident. This is particularly observed for aluminum alloy 7050 overaged to the T73 condition. In fact, this latter alloy exhibits an increased susceptibility to the NaCl solution in the overaged condition, as compared to the T6 temper. On the contrary, in stress corrosion cracking tests, the samples tempered to the T73 condition exhibit an increase in the cracking resistance, as compared to the T6 temper. This appear to be caused by the reduction of the intergranular component and the increase in the transangular component in the fatigue tests as compared to the stress corrosion cracking tests (Lin and Yang 1998). The corrosion fatigue of the high strength aluminum alloys tempered to the T7 condition, for instance aluminum alloy 7050-T7451 weld unaffected parent metal, appears to be generally low. In fact, samples exposed to the 3.5 wt.% NaCl solution exhibit a significant decrease in the cycles to rupture as compared to the samples tested in air (Fig. 11.31). The formation of pits followed by a slight intergranular corrosion appears to be the main factor in controlling the crack formation during corrosion fatigue tests. Pits can act as a main corrosion-fatigue crack initiation factor also for aluminum alloy 7050-T73 (Lin and Yang 1998) or control the corrosion-fatigue live of aluminum alloy 7075-T6 (Sankaran et al. 2001). A higher extent of
Metallurgy and weld performance in friction stir welding Weld/pm transverse air Weld solution PM solution Nugget exposed TMAZ exposed
240
220 Stress [MPa]
365
200
180
160 0
5 ¥ 105
1 ¥ 106 Cycles
1.5 ¥ 106
2 ¥ 106
11.31 Diagram stress cycles for a 7050-T7451 weld and the parent metal (PM) tested in air and in a 3.5 wt% NaCl solution (transverse orientation). Loading frequency 10 Hz; R = 0.1. The tests were also carried out by separately exposing the nugget and the thermomechanically affected zone (TMAZ) to the NaCl solution by using a masking lacquer.
intergranular corrosion is observed for the weld unaffected parent metal tested at lower yield strength. This can be clearly recognised by samples with long exposition times to the 3.5 wt.% NaCl solution. A general increase in intergranular corrosion is observed for the friction stir welded samples as compared to the weld unaffected parent metal. This is again particularly observed for the samples tested at lower stresses. Despite the fact that in mechanical testing the fracture generally takes place within the soft heat affected zones (Jata et al. 2000), in the corrosion–fatigue tests (ASTM E 466-96 2002), the fracture is mainly located along the nugget– thermo-mechanically affected zones (TMAZ). A similar fracture location is also observed during slow strain rate tests (Paglia et al. 2003a). This fact appears not to depend on the stress level applied. The thermo-mechanically affected zones and the nugget both exhibit a similar corrosion–fatigue behaviour (Fig. 11.31). Interestingly, the thermo-mechanically affected zones and the regions of the nugget next to the TMAZ can be subjected to a sensitisation of the grain boundaries (Fig. 11.32) similar to that observed within the heat affected zones (Jata et al. 2000). The partial sensitisation of the grain boundaries is responsible for the high intergranular corrosion susceptibility and the partial intergranular fracture mode observed within the crack initiation zones. And yet, the extent of the intergranular corrosion generally increases with decreasing stress level (Figs 11.33, 11.34). The presence of cracks along the grain boundaries also confirms the intergranular
366
Friction stir welding
500 nm
400 nm
(a)
(b)
11.32 Microstructure of the thermo-mechanically affected zone (a) and the nugget (b) of 7050-T7451 FSW.
A
B C
NaCl + corrosion products D 1 mm
10 µ
20 µ
(b)
(a)
100 µ (c)
100 µ (d)
11.33 Fracture surface of the weld tested in a 3.5 wt% NaCl solution (Rhodes et al., 1997b) (stress 237 MPa, 90% yield stress). A: striations with a slight corrosion and cracks; B: dimple zones; C: featureless transangular fracture; D: striation – flat transangular zone.
corrosion susceptibility of the nugget–thermo-mechanically affected zones for instance for 7050-T7451 friction stir weld. Interestingly, by testing the samples above 60% of the yield strength, the difference in the corrosion– fatigue life between the weld unaffected parent metal and the weld appears to decrease (Fig. 11.31). In this case, the rupture takes place below a limit value of cycles (for instance 2.3 ¥ 105 cycles for the condition tested) for
Metallurgy and weld performance in friction stir welding
367
B A C 20 µ
500 µ (a)
20 µ (b)
2 µ (c)
11.34 Fracture surface of the weld tested in a 3.5 wt% NaCl solution (stress 158 MPa, 60% yield stress). A: intergranular corrosion; B: step-like transangular fracture; C: flat – transangular and ductile dimples; dashed line: intergranular corrosion – overload transangular boundary region.
the parent metal and the weld. This fact appears to indicate a tendency to a similar corrosion-fatigue behaviour of the weld unaffected parent metal and the weld at high stress levels. The heat treatment of the base alloy The development of heat caused by friction stir welding, as mentioned earlier, influences, the precipitate distribution, i.e. the corrosion susceptibility, across the welds. This is particularly observed for high strength aluminum alloys (e.g. 7075-T6 and 7050-T651), where a coarsening of the precipitates takes place within the heat affected zones. Within the nugget the precipitates are fully or partially re-dissolved (Rhodes et al. 1997a; Su et al. 2003). Therefore, the heat affected zones or the nugget–thermo-mechanically affected zone boundary appear to be the most corrosion susceptible weld microzones for high strength aluminum alloys friction stir welds (Lumsden et al. 1999; Paglia et al. 2003a; Hannour et al. 2000b). The overaging of high strength aluminum alloys to the T7 tempers reduces the stress corrosion cracking susceptibility of the unwelded base alloys. Nevertheless, the additional heat experienced by the heat affected zones of the T7 tempered base alloys during friction
368
Friction stir welding
stir welding, causes a sensitisation of the microstructure and decreases the corrosion resistance (Lumsden et al. 1999; Hannour et al. 2000b). This is in particular observed for overaged aluminum alloy 7050-T7451 friction stir welds (Jata et al. 2000). Interestingly, aluminum alloy 7075 plates friction stir welded in the annealed condition (ASM 1991; Paglia et al. 2007), indicate a slight overaging of the heat affected zones and exhibit an increase in the mechanical properties as well as in the corrosion resistance as compared to 7075-T7451 welded plates (Paglia et al. 2007). In this concern, it is interesting to outline the mechanical and the corrosion properties of a high strength aluminum alloy welded in the annealed condition and tested with a constant extension rate device (ASTM G129 2000). This may help to understand how small temperature excursions across the welded regions and the temper of the base unwelded alloy may largely affect the corrosion properties. The friction stir welding of high strength aluminum alloys T6 peak aged (Lumsden et al. 1999; Polmear 1989; Mahoney et al. 1998) or T7 overaged (Paglia et al. 2003a; Polmear 1989), causes a further overaging of the weld zones adjacent to the nugget and increases the environmental cracking susceptibility. This can be generally correlated with the sensitisation of the microstructure (Lumdsen et al. 1999; Paglia et al. 2003a; Mahoney et al. 1998). The 7050-O alloy, i.e tempered in the annealed condition, parent metal and the friction stir weld tested in the transverse orientation exhibit a high environmental cracking susceptibility with ductility ratios (strain to failure eenvironment/strain to failure eair) ranging from 0.44 to 0.29 (Table 11.3). The extent of the intergranular corrosion on the fracture surfaces confirms the Table 11.3 Ultimate tensile strength, strain to failure and ductility ratio data for a 7050-O weld tested in the transverse and in the longitudinal orientation within the weld microzones (ASTM G129 2000). Has: heat affected zone advancing side; hts: heat affected zone trailing side Weld zone Ambient UTS [Mpa]
Strain to failure [%]
Ductility ratio [7 env./7 air]
Parent metal air solution (transverse)
410 ± 0.01 350 ± 0.01
14.2 ± 0.01 4.7 ± 0.01
0.33
Weld (transverse) air solution
358 ± 8.5 242 ± 14.2
3.4 ± 0.2 3.9 ± 1.4
0.29
Parent metal air solution (longitudinal)
441.5 ± 4.9 418 ± 2.8
12.8 ± 2.4 5.7 ± 1.3
0.44
Nugget air solution (longitudinal)
382 ± 31.0 364 ± 43.8
6.9 ± 4.8 3.9 ± 1.4
0.56
Has (longitudinal) air solution
388 ± 25.5 379 ± 18.4
16.4 ± 1.8 12.4 ± 0.9
0.76
Hts (longitudinal) air solution
374 ± 31.2 372.5 ± 45.9
14.0 ± 3.9 12.8 ± 1.2
0.91
Metallurgy and weld performance in friction stir welding
369
corrosion susceptibility of the weld unaffected parent metal (Fig. 11.35). A relatively large intergranular corrosion susceptibility is also observed within the nugget, where fracture occurs occasionally along the nugget deformation flow contours (Fig. 11.36). On the other hand, the heat affected zones also exhibit a corrosion susceptibility, but to a lesser extent as compared to the parent metal and the nugget (Figure 11.37). In fact, the ductility ratios of the heat affected zones in the advancing and retreating side of the weld exhibit values of 0.76 and 0.91 respectively (Table 11.3). A similar trend is observed for aluminum alloy 7075-O in the annealed condition, where the heat affected zones exhibit a relatively higher ductility ratio as compared to the other weld microzones and the parent metal. However, in this latter case, the nugget exhibits a higher corrosion susceptibility (ductility ratio 0.28) as compared to 7050-O alloy friction stir weld (ductility ratio 0.56) (Paglia et al. 2007). The slightly lower environmental susceptibility of the heat affected zones may be caused by the heat produced by friction stir welding (Frigaard et al. 2001). This additional heat experienced by the base alloy
500 µ
100 µ
11.35 Fracture surface for a 7050-O parent metal tested in a 3.5 wt% NaCl solution in the transverse orientation. Intergranular corrosion and cavities along the grain boundaries.
500 µ
200 µ
11.36 Fracture surface for a 7050-O nugget tested in a 3.5 wt% NaCl solution in the longitudinal orientation. Note the successive presence of the onion rings (arrows).
370
Friction stir welding
500 µ
20 µ
11.37 Fracture surface for a 7050-O heat affected zone in the advancing side (Has) tested in a 3.5 wt% NaCl solution in the longitudinal orientation.
friction stir welded in the annealed condition may cause a slight overaging of the heat affected zones, as it is for the 7075-O friction stir weld (Paglia et al. 2007). However, this slight overaging of the microstructure, which might take place on the heat affected zones, causing the formation of precipitates with appropriate dimensions and causing a decrease in the environmental susceptibility, is not sufficient to increase the ultimate tensile strength as compared to the weld unaffected parent metal and the nugget (Table 11.3). Nevertheless, the strain to failure of the heat affected zones is similar to that of the parent metal. The 7050-O parent metal exhibits an increased susceptibility to environmental cracking as compared to 7050-T7451 (Paglia et al. 2003a). Nevertheless, in the welded condition and from the point of view of the environmental cracking susceptibility of the weldment, the lowest value of ductility ratio for aluminum alloy 7050-T7451 friction stir weld is ca. 0.2, while the lowest ductility ratio for aluminum alloy 7050-O friction stir weld is 0.29, (both tests carried out in the transverse orientation in a 3.5 wt.% NaCl solution). Thus, concerning the environmental cracking susceptibility of the weld, 7050-O and 7050-T7451 exhibit a similar behaviour. In this concern, the corrosion susceptible weld microzone for 7050 aluminum alloy in the annealed condition is located along the nugget–thermo-mechanically affected zone boundary, which also represent a critical zone with respect to the fracture location in air (Fig. 11.38). The overaging of 7050 base aluminum alloy plates from the O-temper to the T7451 temper in order to increase the environmental cracking resistance of the friction stir welded condition does not appear significant. The heat affected zones of the high strength aluminium alloys friction stir weld in the annealed condition exhibit a slight increase in environmental cracking resistance and a reduced intergranular corrosion susceptibility as compared to the alloys tempered in the T6 and T7 conditions. Thus, once an alloy is friction stir welded, the properties of the welded
Metallurgy and weld performance in friction stir welding
371
Fracture
TMAZ 0.5 cm
Nugget TMAZ
11.38 Weld cross-section: fracture along the nugget–thermomechanically affected zone interface for the samples tested in air and in a 3.5 wt% NaCl solution with a constant extension rate testing device. Extension rate of 2.5 ¥ 10–5 mm/s (Mahoney et al. 1998).
material may change significantly from the properties of the base unwelded material. In this concern, the increase in the mechanical properties gained with aging and the increase in the corrosion resistance obtained with overaging of the base alloys (Speidel 1975; Reboul et al. 1992; Polmear 1989), can be lost once the material is friction stir welded. Therefore, an accurate choice of the base alloy temper is needed prior to welding in order to obtain a corrosion-resistant weldment. Overaged tempers, that are usually corrosion resistant, do not always appear to be an appropriate choice if the alloy is friction stir welded. The thermal excursion occurring during the welding, causes a “thermal treatment”, i.e. sensitisation of the microstructure, with a resulting formation of zones of increased corrosion susceptibility. This is particularly observed when the base alloy is overaged prior to welding.
11.4
Mechanical properties of friction stir welds in aluminium alloys
11.4.1 Introduction Since during the friction stir welding (FSW) process the temperatures are well below the melting point, problems associated with the liquid/solid phase transformation are avoided. This allows high quality joining of materials that have traditionally been troublesome to weld conventionally without distortion, cracks or voids such as high strength aerospace aluminium alloys, like 2024 or 7475. Besides attractive mechanical properties, especially in fatigue and strength, friction stir welded integral structures are claimed to offer cost and
372
Friction stir welding
weight savings (Pettit et al. 2000). Therefore FSW was recently identified by leading aircraft and spacecraft manufacturers as “key technology” for fuselage and wing as well as for propulsion tanks and launchers manufacturing (Talwar et al. 2000; Thompson 2003). “Inspections or other procedures must be established as necessary to prevent catastrophic failure, and must be included in the Airworthiness Limitations Sections” (JAA-FAA 1998). While for space systems: “A fracture control programme shall require that design be based on fracture control principles and procedures when the initiation or propagation of cracks in structural items during the service life can result in a catastrophic or critical hazard” (ECSS-E-30-01A 1999). The recent FAR and ESA regulations for damage tolerance and fatigue evaluations of aircraft and spacecraft applications require fatigue and fatigue crack propagation assessment. If friction stir welding is considered for primary aircraft and spacecraft structures, consistent fatigue and fracture mechanics testing of the joint has to be performed and fatigue and fatigue crack growth predictions methodologies have to be adopted for this specific joint type. An overview on the small coupon fatigue and crack growth properties of friction stir welded butt joints is given. The investigations were carried out on 2024 (Wanhill 1990), 2024A (Warner et al. 1999), 6013 (Cieslak 1987) and 6056 (Blanc and Mankowski 1998) aluminum alloys in the thickness range of 1.6 mm to 6 mm. AA 6013 was welded also in T-joint form (Erbslöh et al. 2003) and fatigue tested. Different effects influencing the fatigue and the fatigue crack propagation lives were identified and investigated and are presented in the present summary. Weld surface, notch and environmental effects on fatigue life as well as residual stress, variable amplitude loading and the environmental effects on the fatigue crack propagation behaviour of the FSW joints are assessed. The butt and T-joints were produced on modified milling machines at DLR and EADS IW on the basis of the TWI patent (Thomas et al. 1992). For most of the welds conventional tools with a cylindrical threaded pin were used. The only exception were the welds of 6013 and clad 2024A material. Welding of 6013 was carried out at a very early stage of FSW development. Therefore a tool without thread was used. The upper clad layer of the 6mm thick 2024A-T3 sheet was removed in the welding area before welding, whereas the weld root layer remained intact. To prevent excessive stirring and therefore disruption of the lower clad layer (as described in Lederich et al. (2001)) a modified tool was used for the welds. This tool allowed production of butt joints with a closed pure aluminum layer. The welding direction was always parallel to the rolling direction of the sheets. Additional heat treatments after FSW were carried out only in the case of 6013, which was welded in the soft T4 condition and heat treated to T6 (4h, 190°C) after welding. Most of the fatigue test results were obtained from welded specimens
Metallurgy and weld performance in friction stir welding
373
which were loaded perpendicular to the welding direction. The denomination “as welded” refers to specimens which were machined and polished only on the edges. The material flash and the typical rippled structure caused by the rotating shoulder were removed in the case of “polished” specimens by manual polishing. The thickness reduction of polished specimens was in the order of 0.1 mm. Extensive thickness reduction by milling of the upper and lower surface of the specimens is referred to as “skimming”. S-N fatigue testing was carried out under controlled lab environment in resonance machines. Depending on thickness of the specimens the frequencies were in the range of 45 to 80 Hz. The fatigue crack propagation tests were carried out at room temperature and in laboratory air on a computer controlled servo-hydraulic testing machine following ASTM E 647 (2008). Crack propagation was always monitored through the potential drop technique, (Bachmann et al. n.d., 1992).
11.4.2 S–N fatigue Weld surface effects Figure 11.39 shows the results of fatigue tests for the base material 6013-T6 and for FSW joints in the as welded surface condition and after polishing h1
Crack
h2
As welded FSW specimen 300
40
60 12
6013–T6 T B = 4 mm R = 0.1 62 Hz laboratory air
30 104
200
Stress amplitude sa [MPa]
50
Notched base material specimen Base material FSW as welded FSW polished surfaces Base material, step h = 0.3 mm Base material, step h = 0.4 mm Base material, step h = 0.5 mm
200
100 90 80 70 60
h
105 106 Cycles to failure [N]
107
11.39 Fatigue strength of “as welded” and polished FSW specimens compared to the strength of smooth and notched base material specimens.
374
Friction stir welding
the weld face side. The fatigue performance of specimens in the as welded condition was inferior to that of the base material. Cracks nucleated from an overlapped region at the advancing side of the weld face, caused by the notch effect of this material flash. For these specimens fatigue strength of about 40 MPa was measured at 107 cycles, indicating a degradation of 50% compared to the base material. If the weld face side of the FSW joints was polished, the fatigue behaviour improved. In this case the reduction of fatigue strength is comparable to the reduction of static strength (–20%). Bussu and Irving (2001) obtained similar results for 6 mm thick 2024-T3 joints. The considerable detrimental effect on S-N fatigue life of the overlap weld toe is confirmed by the base material specimens with machined steps or notches, which were very similar to the FSW weld toe. As indicated in Fig. 11.39 the fatigue strength of these specimens with steps of 0.3 to 0.5 mm of height lies in the range of the “as welded” FSW specimens. In the high stress/low cycle regime the reduction in fatigue strength of polished specimens was related to internal crack nucleating cavities, as confirmed by fractography (Biallas et al. 2000a). The same behaviour was observed in 2024-T3 aluminum alloy (Ghidini et al. 2004a, b). The critical stress level can be estimated by a simple fracture mechanical model requiring the effective (lower bond) base material threshold DKth,eff and a stress intensity factor solution fitting to the size, shape and position of the pores in the weld as input parameters (Biallas et al. 2000a, b). The fatigue life can be also predicted once these parameters are implemented in fracture mechanics software as initial crack size (Ghidini et al. 2004a, b). Thickness effects Generally the ratio of FSW static strength to base material strength decreases with increasing sheet or plate thickness. A similar effect is found for the fatigue strength, as indicated in Fig. 11.40. Welds of 1.6 mm thick 2024-T3 display nearly parent material ultimate strength and fatigue strength, whereas a 20% drop in fatigue strength and a 7% drop in static strength is observed for 4 mm thick material. Both welds were tested in the “as welded condition”. Results of 6 mm thick specimens (Bussu and Irving 2001) are also included in Fig. 11.40. The decrease in fatigue strength compared to the parent material values is comparable to the drop of the 4 mm thick welds. It should be noted, however, that the 6 mm thick FSW specimens were machined on both surfaces before testing. Therefore not only the tool marks on the upper surface but also the problematic root surface region were removed. Decreasing weld properties with increasing thickness are closely related to the frictional heat generation during FSW on the upper surface by the tool shoulder. With increasing thickness the temperature gradient between the
Metallurgy and weld performance in friction stir welding Thickn. [mm] UTSweld/UTSbase 1.6 0.98 4 0.93 6 0.88
100 90 80 70 60 50 40 30 104
60
Stress amplitude sa [MPa]
200
2024–T3 T Kt = 1 R = 0.1 45 Hz laboratory air
B = 1.6 mm B B B B B
= = = = =
200
300
375
12
4 mm base material 6 mm 1.6 mm FSW as welded 4 mm FSW as welded 6 mm FSW skimmed [Bussu ICAF 2001] 105 106 Cycles to failure [N]
107
11.40 Thickness effect on S-N curves of friction stir welded 2024-T3. The values in the table indicate the ratio of weld to parent material ultimate strength.
upper and lower surface increases, resulting in an overheating of the upper surface and a too cool root region. As a consequence of surface overheating, a material flash is generated and in the cold root region incomplete joining and cavities are often found. Notch effects The results of Fig. 11.41 were obtained from flat specimens with a central notch (notch factor 2.34 (Pilkey 1997)). The parent material and FSW samples were machined on both surfaces from an original thickness of 6 mm down to a thickness of 3 mm. The notch was positioned in the centre of the nugget. During the fatigue test, cracks were forced to nucleate in the fine grained nugget material and to grow in the weld direction, because this was the area with the highest stresses. The low cycle fatigue strength of the nugget material was comparable to the base material strength. A slight increase of the FSW fatigue strength was observed in the high cycle regime. This behaviour is attributed to the good fatigue properties of the nugget microstructure: the static strength of the nugget material is comparable to base material, since the hardness curve of the 2024A-T3 weld was very similar to the one shown in Fig. 11.51 (failure locus at x = 0). Moreover a different “initial discontinuity state IDS” (Merati 2003a) is present in the nugget, since the constituent particles are crushed during the stirring process (Biallas et al.
376
Friction stir welding 2024A-T3 T B = 3mm
10 6 3
Net section Stress amplitude sa [MPa]
s6.35
0
20
Milled surfaces Kt = 2.34 R = 0.1 80 Hz laboratory air
Base material FSW 104
105 106 Cycles to failure [N]
107
11.41 Fatigue strength of skimmed, notched specimens (log-log scales).
1999a). As shown by Merati (2003a) the amount, shape and distribution of these particles is controlling the fatigue crack nucleation in bare 2024-T3. Clad material One of the major drawbacks of friction stir welding is the corrosion susceptibility of 2xxx and 7xxx alloys in the weld area and HAZ (Paglia et al. 2003b; Connolly et al. 2004). Therefore, a friction stir welding procedure for clad materials was developed at DLR in collaboration with EADS CRC F. The technique allowed the production of joints with a closed pure aluminum layer on the root surface as presented in (Dalle et al. 2003). The ratio of FSW to parent material ultimate strength was 0.86. In comparison to bare FSW data, a higher amount of scatter was observed in the clad material fatigue strength, Fig. 11.42. Moreover the difference between “as welded” and polished specimens was less evident than in the bare material. This behaviour is attributed to the nucleation of fatigue cracks from different discontinuities in the weld nugget, Fig. 11.43. Most of the “as welded” specimens failed because of fatigue cracks starting from the material flash on the advancing side of the weld. Once the tool marks were polished away, the fatigue cracks started from clad inclusions in the lower side of the weld nugget. Therefore removing of the rippled area did not yield such a distinguished improvement of fatigue strength as observed in Fig. 11.40. Some “as welded” and polished specimens exhibited very low fatigue
Metallurgy and weld performance in friction stir welding
377
24
Kt = 1.03 R = 0.1 80 Hz laboratory air
50 130
Stress amplitude sa [MPa]
2024A–T3 T clad B = 6 mm
R100
–20% Base material (clad) FSW as welded FSW upper surface polished
104
105 106 Cycles to failure [N]
107
11.42 S-N fatigue curves of the clad 2024A-T3 base material and FSW joints.
6 mm
6 mm
adv.
retr.
6 mm
Failure from the material lip (“as welded” specimens)
Root failure from oxide inclusions (polished and “as welded” specimens)
adv.
Failure from clad inclusions (polished specimens)
11.43 Failure location of friction stir welded joints of clad 2024A-T3.
strength and cracks which nucleated form the weld nugget root in the symmetry line of the joint. EDX analysis of the fracture surface revealed oxide inclusions in this area, which were responsible for crack nucleation. Apparently the reduced stirring action, which was necessary to maintain a closed clad layer, was sometimes not mixing up the root material enough to
378
Friction stir welding
disrupt completely the original oxide coatings present on the facing surfaces of the butt joint before welding. Dissimilar materials
4 mm
In general, while FSW welding dissimilar materials, the material more difficult to weld generally implied the welding parameters, (Cavaliere et al. 2009; Krietensen et al. 2004). Since the more difficult alloy had to be joined with higher heat input, i.e. lower welding speeds, the welding speed chosen for the dissimilar joint was too slow for the “easy to weld alloy” of the bi-material combination. In present case of 2024A-T3 / 6056-T6 dissimilar alloys this resulted in an overheating of the “easy to weld” 6056-T6. Obviously the 4 mm thick sheets of both alloys had to be joined in the final heat treatment and no post-weld heat treatments were applied. The overheated HAZ of 6056, which can be easily identified by the strong hardness drop in Fig. 11.44, was the weakest link during the static tensile tests. A drop of 25% compared to the 6056-T6 base material ultimate strength was observed. The trends of S-N curves obtained from polished and “as welded” specimens
180
160
X Retreating: 2024A–T3
Advancing: 6056A–T6
2024A–T3
140 HV1
6056–T6 120
100
80 –25 –20 –15 –10
–5
0 5 ¥ [mm]
10
15
20
25
11.44 Cross-section and hardness curve of a 2024A-T3/6056-T6 dissimilar joint (not in scale).
Metallurgy and weld performance in friction stir welding
379
are similar to the trends discussed so far for similar FSW joints, Fig. 11.45. The broken line indicates in Fig. 11.45 a 20% drop of the 6065-T6 base material fit. Especially at higher stresses the data of polished specimens is above this line, i.e. the drop in fatigue strength was less than 20%. At very high fatigue loads, plastic collapse in the weak 6056 HAZ triggered specimen failure, whereas at lower loads fatigue cracks nucleated in the weld nugget. Unpolished specimens show a higher decrease of fatigue life in comparison to the 6056-T6 base material (approximately 30%). The fatigue cracks nucleated from the small material lip at the edge of the rippled zone on the 6056 side of the weld. The joining of dissimilar Al 6013-T4 alloy and X5CrNi18-10 stainless steel was also successfully carried out using friction stir welding and the fatigue properties were characterised (Uzun et al. 2005). Figure 11.46 shows the cross-section of the dissimilar friction stir weld, AA 6013 T4 to stainless steel. The base aluminium alloy contains elongated grains having diameter range of 80–120 mm in the rolling direction. The weld nugget presents a composite structure of stainless steel particles (pulled away by forge of tool pin from stainless steel surface) reinforced AA 6013. Stainless steel particles have an irregular shape and inhomogeneous distribution within the weld nugget. The S–N curves for Al 6013-T4/stainless steel FSW joints and Al 6013-T6 alloy are shown in Fig. 11.47. Fatigue lifetimes of Al
2024A-T3/6056-T6 B = 4 mm 20
Stress amplitude sa
6056-T6 base material
FSW polished 105 106 Cycles to failure [N]
200
–20%
FSW as welded
104
30
Kt = 1 R = 0.1 84 Hz laboratory air 107
11.45 Fatigue strength of dissimilar FSW joints compared to a fit through 6056-T6 parent material data.
Friction stir welding
Stainless steel HAZ
TMAZ
Weld nugget
TMAZ
HAZ
e a
b
c
f
Al 6013
g
d
4 mm
380
h Advancing side
Retreating side
11.46 Cross-section of the friction stir welded AA 6013-T4 alloy to X5CrNi18-10 stainless steel. 210 FSW Al6013-T4/X5CrNi18-10 Al6013-T6 base mterial [13]
Stress amplitude (MPa)
190 170 B = 4 mm R = 0.1 62 Hz
150 130 110 90
Al6013-T6 base material
70 50
FSW Al6013-T4/X5CrNi18-10 103
104
105 106 Number of cycles to failure
107
108
11.47 S-N curves for AA 6013-T4/X5CrNi18-10 FSW joint and AA 6013-T4 base material (Uzun et al. 2005).
6013-T4 stainless steel joints were found to be approximately 30% lower than that of the Al 6013-T6 alloy base metal. The friction stir welded Al 6013-T4 to stainless steel specimens failed from cracks that had initiated on the root site of the welding zone where the lack of stirred material regions was found, thus resulting in premature fatigue failures. T-joints T-joints were produced by plunging the pin of the FSW tool through the 4 mm thick 6013-T4 cover plate and into the 4 mm thick 6013-T4 and 2024-T3 stringer, (Erbslöh et al. 2003; Ghidini 2006a). A considerable difference is
Metallurgy and weld performance in friction stir welding
381
observed between the two types of welds if comparing the hardness profiles, Fig. 11.48. In the case of the AA 6013-T6 skin welded with the AA 6013-T6 stringer, the hardness profile is similar to that of the AA 6013-T6 butt joints, (Ghidini 2006a; Dalle et al. 2000). Parent plates, dash bands in the diagrams, hardness values were reached at more than 10 mm from the centre of the weld. In the case of the AA 6013-T6 skin welded to the AA 2024-T3 stringer the two minima are lower (almost 60) and the base material hardness values are not reached even 20 mm from the weld nugget. Also in the centre of the weld, the hardness is considerably lower than base material hardness. However the observed effect does not impact the fatigue life of the joints as demonstrated in Fig. 11.49. No difference is observed in the S-N curves of the two welds also when comparing the results obtained with the 12.5 mm 180 160
Base material
HV1
140 120 100 80 AA 6013–T6AA6013–T6 T–FSW 60 –25 –20 –15 –10 –5 0 5 10 15 20 25 Position [mm] 160 140
AA 6013–T6/AA6013–T6 T–FSW Base material
HV1
120 100 80 60 –25 –20 –15 –10 –5 0 5 10 15 20 25 Position [mm]
11.48 Hardness profile of the T friction stir welded joints. In the diagram on the left is reported the hardness profile of the AA 6013T6 skin welded to the AA 6013-T6 stringer, while on the right the hardness profile refers to the AA 6013-T6 skin welded to the AA 2024-T3 stringer.
382
20
51
51
R2
80
65.22
42 200
150
AA 6013–T6 FSW T–joint t = 4 mm R = 0.1
100
110
R2 125
60 Hz laboratory air 50
104
105 Cycles
106
107
11.49 Comparison between the fatigue lives of the 12.5 mm and the 25 mm width FSW T-Joints specimens.
0 28.34
19
110
110 250
110
Maximum stress [MPa]
300
40
4
Friction stir welding
Experimental data specimen 12.5 mm width AA 6013/AA 2024 Experimental data specimen 25 mm width AA 6013/AA 2024
350
Metallurgy and weld performance in friction stir welding
383
width samples with the 25 mm, suggesting that the very low transversal residual stresses didn’t influence the fatigue response of the joints in both sample geometries. Environmental effects A susceptibility of FSW joints to localised corrosion has been proved and correlates to the increase in microstructural heterogeneity associated with local heating. However, in the case of friction stir welds there is little work done (Pao et al. 2001; Sankaran et al. 2002). Previous studies (Paglia et al. 2003b; Meletis et al. 2003; Connolly et al. 2004; Biallas et al. 1999b; Alfaro et al. 2004) have indicated that the heat affected zone (HAZ) and the weld nugget are susceptible to corrosion. Prior to fatigue testing, 2024-T3 un-notched pristine and FSW test samples were corroded with an alternate immersion technique in accordance with ASTM G44-99 standard, (2003) for 100, 250 and 1000 hours. All specimens were prepared by manual polishing and by covering the edges with a protective mask (TURCO 5580G), leaving just the centre of the specimen exposed to the corrosion attack, (Ghidini 2006a; Alfaro et al. 2004). The diagram in Fig. 11.50 shows the effect of corrosion time on the fatigue life of 2024-T3 friction stir welded joints and base material. Also a comparison between base and FSW non-corroded material is given. Corrosion has a detrimental effect reducing the fatigue strength for both types of specimens. The FSW joint show a similar drop in fatigue strength as the base material; this drop is virtually independent from the corrosion time, since the test results of 100 h, 250 h and 1000h pre-corrosion specimens follow in the same scatter band. Most of the un-corroded FSW-specimens failed either in the unaffected base metal (62%) or at the advancing side of the weld nugget (12.5%). Generally, the origin of the failure was found to be a cluster of the second phases present in the microstructure of the 2024-T3 aluminium alloy. The base material has a shorter fatigue life as the FSW material at least on the threshold region: At first glance, this could be contradictory to the normal fatigue behaviour of FSW structures, (Kristensen et al. 2004; Dalle et al. 2004a, b) a preliminary explanation of the experimental observation could be that the surfaces of the sheets have been milled before to produce the weld. This could result in a smoother surface, with smaller clusters, and therefore in a better polishing procedure which leads to better fatigue properties. This effect disappears when considering the corroded specimens where practically no difference exists between the pristine and the FSW specimens. In fact the fracture mechanism is now governed by the corrosion pits, whose dimensions are bigger than the clusters and the constituent particles (Ghidini 2006a; Alfaro et al. 2005).
384
Base material 100h corroded Base material 250h corroded Base material 1000h corroded
400
FSW material 100h corroded FSW material 250h corroded FSW material 1000h corroded
300
FSW weld 12.5
Maximum stress [MPa]
350
250
30 200
220
150
AA 2024–T3 t = 4 mm R = 0.1
100
60 Hz laboratory air
50
104
Pre-corroded area 105
Cycles
106
107
11.50 Comparison between not corroded and 100 h, 250 h, 1000 h corroded polished base and FSW material.
Friction stir welding
Base material not-corroded FSW material not-corroded
Metallurgy and weld performance in friction stir welding
385
Variable amplitude loading The hardness profile of AA 2024-T3 FSW butt joint and a cross-section of the joint are presented in Fig. 11.51, where also the base material hardness (dash bands in the diagram) is present. The diagram shows the same general trend observed in previous works (Biallas et al. 1999a; Dalle et al. 2000). Variable amplitude fatigue tests were performed in centre hole specimens machined taking the hardness profile into account (Ghidini et al. 2009). Samples were manufactured with the hole located 5 mm out of the weld centre in correspondence of the hardness minimum. The centre hole FSW specimens were tested under a standardised flight simulation load history, FALSTAFF (1976). FALSTAFF is a standard load sequence considered representative for the load time history in the lower wing skin near the wing root of a fighter aircraft. The experimental results are plotted in a diagram showing the different levels of smax plotted against the number of flights at failure, Fig. 11.52.
TMAZ I
HAZ
1 2 3 4
= = = =
DXZ TMAZ HAZ Base material
TMAZ II
4 160
DXZ
3
2
1
2
3
4
Base material
140
HV1
BM
120
100 AA 2024–T3 FSW 80 –25 –20 –15 –10 –5 0 5 10 15 Position [mm]
20 25
11.51 Hardness profile of the AA 2024-T3 friction stir welded material.
386
Falstaff
Maximum stress [MPa]
400 5
350 12.5 12.5
300
250
200
FSW weld
ø5
110
AA 2024–T3 FSW t = 4 mm 30 Hz laboratory air 102
103
Flights
104
105
11.52 Experimental results of the centre hole FSW specimen under FALSTAFF spectrum loading.
110
Friction stir welding
450
Metallurgy and weld performance in friction stir welding
387
11.4.3 Fatigue crack propagation (FCP) Residual stress effect In Fig. 11.53a, the da/dN-DK curves of longitudinal weld specimens (closed symbols) are compared to the base material curves (open symbols). At high mean stress levels or R-ratios there was no difference in the FCP behaviour of parent material and welded joint. At low R-ratios and loads (R = 0.1, DK<15 MPa÷m) there was, however, an apparent improvement in the welded material properties. It was demonstrated (Dalle et al. 2000; Ghidini 2000; John and Jata 2001) that this effect was almost entirely caused by the residual stresses, which were not considered in the evaluation of the welded specimen data. On the basis of the Ktr-approach as described in (Ghidini 2006a) and the measurement of stress intensity factor due to residual stresses it was possible to eliminate the residual stress effects from the test data. The da/ dN-DKtr curves of the base material and the friction stir welds then fall in a common scatter band, i. e. the FCP properties of base material and FSW joints were very similar, Fig. 11.53b. A similar result was found by Bussu and Irving (2003), who eliminated the residual stresses by pre-straining the welded specimens before testing. The transverse residual stresses, which were responsible for the effects in Fig. 11.53a, were very low (around 40 MPa). More pronounced effects were expected for crack growth perpendicular to the welding direction, since the longitudinal stresses are usually much higher than the transverse residual stresses (Ghidini et al. 2003). Figure 11.54 shows tests carried out with transverse welded (TW) C(T)-specimens at constant nominal DK-values. In parent material specimens this procedure would have led to constant crack propagation rates. The fatigue crack propagation results of the two TW specimens were normalised with the respective constant fatigue crack propagation values of the parent material at the same nominal loading. The predicted curve was calculated assuming that the real or true resistance of the HAZ and weld were equal to the base material properties and that the changes in da/dN were only due to residual stresses (see details in Dalle et al. (2001)). The favourable agreement between experimental and predicted values indicated that acceleration, deceleration and final acceleration of the crack crossing the weld was mainly caused by residual stresses and not by varying resistance to FCP of the different weld microstructures. This was further confirmed with overloaded base material samples (Ghidini et al. 2006b). Environmental effects The effect of a corrosive environment on the fatigue crack propagation rate of friction stir welded AA 6013-T6 samples has been investigated, (Röhn 1999). Fatigue crack growth tests have been performed while CT specimens
10–2 = = = =
0.1, 0.1, 0.7, 0.7,
R = 0.1, FSW R = 0.7, FSW
base material FSW base material FSW
Friction stir welding
R R R R
10–3
10–4
2024–T3 C(T) specimen W = 50 mm B = 4 mm
10–5
da/dN [mm/cycle]
10–3
da/dN [mm/cycle]
388
10–2
Base material 10–4 2024–T3 C(T) specimen W = 50 mm B = 4 mm
10–5
30 Hz laboratory air
30 Hz laboratory air 10–6
10–6 2
4
6
8 10 20 DK [MPa√m] (a)
40
2
4
6 8 10 DKtr [MPa√m] (b)
20
40
11.53 Fatigue crack propagation in the weld nugget parallel to the weld: (a) nominal FCP curves, (b) true FCP curves.
Metallurgy and weld performance in friction stir welding
389
FSW shoulder width 1.75 1.50 (da/dN)/(da/dN)base material
6013–T6 C(T) TW W = 50mm B = 4mm R = 0.16 DK = 12.7 MPa√m
Spec. 1 Spec. 2 Prediction
1.25
W = 50 mm
Y 1.00 0.75
31
0.50 0.25 0.00 0.2
0.3
0.4
0.5
0.6 0.7 y/ W; a/ W
0.8
0.9
1.0
11.54 Fatigue crack propagation rates of transverse welded specimens under constant DK and R loading compared to the prediction using base material Ktr data and the residual stress distribution in the specimens.
da/dN [mm/cycle]
0.01
1E–3
Thickness R–6 mm W = 50 mm
AA 6013–T6 R = 0.1 3.5% NaCl
W
Frequency = 10 Hs
a
1E–4 FSW FSW FSW FSW
1E–5
sample sample sample sample
#1 #2 #1 #2
1E–6 1
10 DK [MPa √m]
100
11.55 Comparison between base and FSW AA 6013-T6 tested in 3.5% NaCl solution at R = 0.1.
were immersed in a 3.5% NaCl water solution. A comparison between base and FSW materials is given in Fig. 11.55 and in Fig. 11.56 for R ratios of 0.1 and 0.8 respectively. The residual stress effect shown in Fig. 11.53 is still evident also in corrosive environments. Fig. 11.57 and Fig. 11.58 show the comparison of the base material samples in air and in NaCl for R = 0.1 and R = 0.8 loading ratios respectively.
390
Friction stir welding 0.01
da/dN [mm/cycle]
1E–3
AA 6013–T6 R = 0.8 3.5% NaCl Frequency = 10 Hs
1E–4
FSW sample #1
1E–5
FSW sample #2 FSW sample #1
1E–6 1
10 DK [MPa √m]
100
11.56 Comparison between base and FSW AA 6013-T6 tested in 3.5% NaCl solution at R = 0.8. AA 6013–T6 R = 0.1 FSW material
da/dN [mm/cycle]
1E–3
Frequency = 10 Hs
1E–4
1E–5 Tested in air Tested in 3.5% NaCl 1E–6 1
10 DK [MPa√m]
100
11.57 Comparison between base and AA 6013-T6 base material tested in 3.5% NaCl solution and in air at R = 0.1.
A detrimental effect of the corrosive environment in the fatigue crack propagation rates has been demonstrated for both load histories. The crack propagation appears to be much faster in the presence of the corrosive medium than in air. Apparently, the corrosion products acting at the crack tip are locally reducing the fracture toughness of the material consistently with hydrogen embrittlement mechanisms (Piascik and Gangloff 1992). A different behaviour has been observed for the FSW joints at R = 0.1, Fig. 11.59. In this case the fatigue crack growth rate of the FSW samples tested in NaCl solution is identical to the one measured in air. A possible
Metallurgy and weld performance in friction stir welding
391
AA 6013–T6 R = 0.8 Base material
da/dN [mm/cycle]
1E–3
Frequency = 10 Hs
1E–4
1E–5 Tested in air Tested in 3.5% NaCl 1E–6 1
10 DK [MPa√m]
100
11.58 Comparison between base and AA 6013-T6 base material tested in 3.5% NaCl solution and in air at R = 0.8. 0.01
da/dN [mm/cycle]
1E–3
AA 6013–T6 R = 0.1 Base material Frequency = 10 Hs
1E–4
1E–5 Tested in air Tested in 3.5% NaCl 1E–6 1
10 DK [MPa√m]
100
11.59 Comparison between base and AA 6013-T6 FSW material tested in 3.5% NaCl solution and in air at R = 0.1.
explanation for the observed phenomenon could be the compressive residual stresses present in the welded joint. The residual stresses could have generated a closure of the crack mouth reducing the exposure to the corrosive medium and therefore reducing the detrimental effect of the corrosion products acting at the crack tip.
392
Friction stir welding
The effect disappears at higher R ratio, R = 0.8, where the effect of the residual stresses is not relevant anymore, Fig. 11.60. Here, as for the base material, Fig. 11.57 and Fig. 11.58, the corrosive environment shows a severe impact in the crack propagation rate if compared to the samples tested in air. Variable amplitude loading AA 2024-T3 base and FSW material specimens with a crack starter notch of a total nominal length of 10 mm have been manufactured. In the base material samples the notch was placed in the middle, while in the FSW coupons the notch was located 5 mm out of the weld centre line, in the most critical part of the joint where the hardness minimum is located (Ghidini 2006a), see Fig. 11.51. Variable amplitude loading fatigue crack propagation tests have been carried out both in the base as well as in the FSW specimens. The spectrum were covering single tensile peaks overloads every 50 cycles, spectrum 1, tensile over- and underload every 100 cycles and tensile and compressive over- and underloads every 39 cycles, spectrum 3. Spectrum 4 was the FALSTAFF load history. All the sequences have been applied iteratively till specimen failure. Details of the test program and of the spectra are given in (Ghidini 2006a, c, 2005a). The experimental results have been plotted in da/ dN-a Fig. 11.61 and Fig. 11.62 (where also constant amplitude loading tests at R = 0.1 are presented for reference purposes) as well as a-N diagrams, Fig. 11.63. Figure 11.61 and Fig. 11.62 show that both the base material AA 6013–T6 R = 0.8 FSW material
da/dN [mm/cycle]
1E–3
Frequency = 10 Hs
1E–4
1E–5 Tested in air Tested in 3.5% NaCl 1E–6 1
10 DK [MPa√m]
100
11.60 Comparison between base and AA 6013-T6 FSW material tested in 3.5% NaCl solution and in air at R = 0.8.
Metallurgy and weld performance in friction stir welding
da/dN [mm/cycle]
0.1
0.01
80 8
MT specimen W = 160 mm B = 4 mm
393 4
142 100 60 –37 142 100 60 32
FSW joint 1E–3
100 60 20
1E–4
5
360
R = 0.1 Spectrum 1 Spectrum 2 Spectrum 3 FALSTAFF
Crack starter
1E–5 0
10
20
30
40 50 a [mm]
60
70
160
80
11.61 Comparison between constant amplitude, variable amplitude and flight simulation loading for base material specimens. 0.1
MT specimen FSW W = 160 mm B = 4 mm
80 8
39 39
da/dN [mm/cycle]
0.01
100
140 100 60 –37 140 100 60 32 100 60 20
1E–3
R = 0.1 Spectrum 1 Spectrum 2 Spectrum 3 FALSTAFF
1E–4
1E–5 0
10
20
30
40 a [mm]
50
60
70
80
11.62 Comparison between constant amplitude, variable amplitude and flight simulation loading for friction stir welded specimens.
as well as the FSW curves have the same trend: the crack growth rates for the load types 2-4 and FALSTAFF were in all cases lower than in constant amplitude tests. However a real comparison is not possible since the constant amplitude loading test has a R = 0.1 load ratio, while the spectrum 1 is a R = 0.33 and spectrum 2 and 4 are R = 0.6 ratio of the base loading. Referring to the variable amplitude loading spectra, the spectrum 1 (with one high cycle followed by 100 small cycles) showed the lowest overall crack growth
40
a [mm]
a [mm]
40 30
Spectrum 1
10
100 0
200000
70
400000 Cycles
600000 40
40000
80000 Cycles
120000
140 100 60 20 160000
Base material FSW
35 30 a [mm]
a [mm]
60 0
50 40 30
25 20 15
20 Spectrum 3
10 0
Spectrum 2
10
Base material FSW
60
30 20
100 60 20
Friction stir welding
50
20
Base material FSW
60
50
394
Base material FSW
60
0
20000
40000 60000 Cycles
80000
140 100 60 –37 100000
Falstaff
10 5 0
0
1000 2000 Number of flights
3000
11.63 Overview of the complete experimental results: base material and FSW experimental observations are compared for every single loading spectrum.
Metallurgy and weld performance in friction stir welding
395
rates. FALSTAFF is the only spectrum which showed a different behaviour in the base material (lowest rate) and in the FSW, where also the scatter is higher compared to the pristine coupons. Figure 11.63 shows the complete program results as comparison between the base and the friction stir welded material for every single loading spectrum. All the comparisons have been done between one specimen of base material and two of friction stir welded material in order to also have information about the scatter in the welded structures. All the comparisons are presented with the respective experimental crack length. As a general trend the FSW specimens showed a faster fatigue crack growth rate as was expected by considering the residual stresses distribution (Ghidini 2006a), where a considerable portion of the FSW samples showed tension residual stresses (even if of intensity, with a maximum measured value of 15 MPa).
11.4.4 Fatigue and fatigue crack propagation predictions Significant work has been devoted to modelling the fatigue life as well as the fatigue crack propagation of FSW joints using fracture mechanics principles (Ghidini et al. 2003, 2004a, b, c, 2005a, b, 2006a, b, c, 2009; Alfaro et al. 2004, 2005), where details of the predictions assumptions are given. A simple engineering approach has been developed and may provide a practical and reliable basis for the fatigue and fatigue crack propagation analysis of FSW structures under in-service loading conditions and in the presence of corrosion attack, by using widespread fracture mechanics software (NASAESA 2000; Harter 2002). S-N fatigue The fatigue lives of base and FSW welded material have been successfully predicted and the results are presented in Fig. 11.64. The following assumptions have been made: cracks formed immediately at the inclusions, particles for both base and FSW samples and fatigue lives would only comprise crack propagation. The residual stresses were neglected due to the very small dimensions of the samples and the same material data were used for base and for FSW material predictions, (Ghidini 2006a). Same assumptions have been used in the prediction of the centre hole FSW specimens under FALSTAFF loading, combined with the use of the Willenborg retardation model (Willenborg et al. 1971) with a shut off overload ratio of RSO = 3 (Ghidini and Dalle Donne 2009) (Fig. 11.65). When predicting the life of the FSW 6013 T-joint samples, in-house obtained fatigue crack propagation data and b values calculated with FRANC2D/L FEM code (Swenson and James n.d.), were implemented in the calculations, since both, the material and the T geometry were not available in the fracture
300 250 AA 2024–T3 t = 4mm R = 0.1
200 150 100
104
Cycles
400
106
350 300 250 AA 2024–T3 t = 4mm R = 0.1
150 100
104
AA 2024–T3 t = 4mm R = 0.5
200 150
60 Hs Laboratory air 104
Cycles
106
107
Cycles
106
107
Experimental data NASGRO 3.0 AFGROW
350 300 250 200
AA 2024–T3 t = 4mm R = 0.1
150
60 Hs Laboratory air
50 105
105
400
100
60 Hs Laboratory air
50
250
107
Experimental data NASGRO 3.0 AFGROW
200
300
50 105
Maximum stress [MPa]
Experimental data NASGRO 3.0 AFGROW
350
100
60 Hs Laboratory air
50
Maximum stress [MPa]
Maximum stress [MPa]
350
400
Friction stir welding
Maximum stress [MPa]
Experimental data NASGRO 3.0 AFGROW
396
400
104
105 Cycles
106
107
11.64 Comparison between base and FSW material experimental results and computer simulations.
Metallurgy and weld performance in friction stir welding 450
397
Experimental data afgrow Willenborg RSO = 3
Maximum stress [MPa]
400
350
300 AA 2024–T3 FSW t = 4 mm falstaff
250
60 Hs Laboratory air
200
102
103
Flights
104
105
11.65 Comparison between FSW open hole specimens experimental results under FALSTAFF and computer simulations.
Experimental data afgrow
Maximum stress [MPa]
350 300 250 200 AA 6013–T6 FSW T-joint t = 4 mm R = 0.1
150 100
60 Hs Laboratory air
50
104
105 Cycles
106
107
11.66 Comparison between T-joint R = 0.1 experimental results and computer simulation.
mechanics software used for the predictions (Ghidini 2006a). Calculations were in good agreement with the experimental results, Fig. 11.66. The fatigue lives of pre-corroded base and FSW samples showed very good agreement with the experimental results once the following assumptions have been made (Ghidini 2006a): the weld material is treated as a base material (Fig. 11.67). Pitting and inter-granular corrosion are treated as a
398
Friction stir welding 250h Corroded experimental data NASGRO 3.0 AFGROW
AA 2024–T3 t = 4 mm R = 0.1
350 300 250
60 Hs Laboratory air
200 150 100
104
105 106 Cycles
Maximum stress [MPa]
Maximum stress [MPa]
100h Corroded experimental data NASGRO 3.0 AFGROW
300 250
150 100 104
107
300 250
60 Hs Laboratory air
200 150 100
104
105 106 Cycles
Maximum stress [MPa]
Maximum stress [MPa]
AA 2024–T3 t = 4 mm R = 0.1
300 250
60 Hs Laboratory air
200 150 100
104
105 106 Cycles
107
Maximum stress [MPa]
Maximum stress [MPa]
AA 2024–T3 FSW t = 4 mm R = 0.1
107
AA 2024–T3 FSW t = 4 mm R = 0.1
350 300 250
60 Hs Laboratory air
200 150 100 50
107
105 106 Cycles
100h Corroded experimental data NASGRO 3.0 AFGROW
250h Corroded experimental data NASGRO 3.0 AFGROW 350
60 Hs Laboratory air
200
1000h Corroded experimental data NASGRO 3.0 AFGROW 350
AA 2024–T3 t = 4 mm R = 0.1
350
104
105 106 Cycles
107
1000h Corroded experimental data NASGRO 3.0 AFGROW AA 2024–T3 FSW t = 4 mm R = 0.1
350 300 250
60 Hs Laboratory air
200 150 100 50
104
105
106 Cycles
107
11.67 Comparison between base and FSW pre-corroded material experimental results and computer simulations.
Metallurgy and weld performance in friction stir welding
399
single corrosion damage source and the model surface crack comprehends this damage. The several corrosion damaged areas of the specimen surface are simulated with a single semi-elliptical surface crack having the dimensions of the deepest and the widest corrosion damage area. Fatigue crack propagation The fatigue crack propagation of FSW joints has also been successfully predicted under constant amplitude loading and under flight histories (Ghidini et al. 2006a, c, 2003). The two most important effects affecting fatigue crack propagation under complex and in-service loading conditions of welded structures have been identified and firstly separately simulated: the residual stresses (Ghidini et al. 2006b) and the retardation and acceleration effects (Ghidini 2006a). The single solutions of the two different problems have been then superimposed to predict the fatigue crack propagation of bigger friction stir welded specimens under more complex loading spectra and under flight simulation. The simple measurement and inclusion of the residual stresses and the use of the retardation models of the software packages improved the quality of the predictions considerably. Particularly for loading spectra with a predominant influence of overloads the Willenborg model with a shutoff overload ratio RSO = 3 gave the best results, while for loading spectra with a predominant influence of under-loads the best predictions have been obtained without using any retardation model. In Fig. 11.68 and Fig. 11.69 1000000
Base material Spectrum 1
Cyclespredicted
100000
Spectrum 2 Spectrum 3 Constant amplitude Non-conservative
10000
Conservative Falstaff (number of flights)
1000 1000
10000 100000 CyclesExperimental
1000000
11.68 Predicted fatigue life of the base material compared to the actual one.
400
Friction stir welding 1000000
FSW material
Spectrum 3
Cyclespredicted
100000
Spectrum 1 Spectrum 2
Constant amplitude Non-conservative 10000 Conservative Falstaff (number of flights) 1000 1000
10000 100000 CyclesExperimental
1000000
11.69 Predicted fatigue life of the friction stir material compared to the actual one.
experimental results are compared with the predictions both for base and FSW material.
11.5
References and further reading
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Index
AA 2024, 65 AA 6013-T6, 57 AA 7050-O, 369, 370 fracture along nugget–TMAZ interface, 371 ultimate tensile strength, strain to failure and ductility ratio data, 368 weld fracture surface HAZ in advancing side tested in NaCl solution, 370 nugget tested in NaCl solution, 369 parent metal tested in NaCl solution, 369 AA6082-T6, 166 Aero system, 111, 113 aerospace friction stir welds in Falcon 9 tank of SpaceX, 133 FSW applications, 130–7 aircraft fuselages and wings in Europe, 137 aluminium panels in aircraft production in USA, 133–6 first flightworthy parts of cargo aircraft in USA, 133 fuel tanks for spacecraft in USA, 130–3 FSW toe nails for military cargo aircraft, 134 NASA’s first official weld with tools for Ares I upper stage hardware test articles, 132 Vertical Weld Tool at Marshall Space Flight Center of NASA, 132 age hardening, 42 Airbus, 5 A340-500 HGW, 104 A400M military aircraft FSW floor panel, 137
Airwothiness Limitations Sections, 372 Al-based dissimilar welds, 343–60 Al alloy to HSS, 345–54 cross-section macrostructure of single overlap configuration, 351 different welded zones cross-section view, 346 microstructure investigations on mixing area between materials, 352 overlap shear test curves, 355 schematic illustration of single overlap configuration, 350 tensile stress-strain curves, 349–50 weld interface and close view of bonding between both materials, 347–8 weld parameter details, 345 welded joints overlap shear test results, 353–4 welding setup, 346 Al alloy to Mg alloy, 354–60 Al alloy and Mg alloy interface, 358 BM vs SZ microstructure of Al alloy and Mg alloy, 357 joint transverse cross-section microstructure, 356 reaction products at interface region, 359 welding setup, 355 friction stir dissimilar welding, 344–5 general aspects, 343–4 motivation, 343 Al-Cu-Mg-Mn 2024, 254 Al-Cu-Mg-Mn 2024 alloy identical processing conditions in stir zones produced, 254 stir zone weld nugget formation from top of the weld, 251 transverse to weld direction, 250
411
412
Index
Al-Cu-Mn-Mg 2014, 250 Al-Cu-Mn-Mg 2014-T6, 260 Al-Li-Cu alloy, 48 Al-Mg-Mn 5083, 258 Al-Si-Mg-Cu-Mn 6013 alloy, 250 AlMgSc alloy and al-based dissimilar joints welding metallurgy analysis, 319–42 effect of scandium addition, 319–20 microstructure evolution, 321–4 substructural evolution, 324–42 summary, 342 thermal history during FSW, 320–1 anti-phase boundary and Orowan strengthening mechanisms, 341 correlation between precipitation stability and thermal cycles, 322 different grain structures observed in BM, TMAZ and SZ, 326–7 FSW joint thermal history using low, intermediate and high heat input, 320 weld seam, 323 Orowan’s mechanism occurring in TMAZ of hot parameter, 335 particular substructural aspects, 330 precipitate features base material, 331–2 BM and different FSW zones, 336 stir zone, 333–4 precipitate size and distribution of FSW joint, 338–9 typical dislocation features, 328–9 Alstom Pendolino train, 142 aluminium, 29, 266 aluminium alloy, 3, 4, 176–7, 343 corrosion behaviour of friction stir welds, 360–71 base alloy heat treatment, 367–71 corrosion under fatigue, 364–7 general aspects, 360–4 mechanical properties of friction stir welds, 371–400 fatigue and fatigue crack propagation predictions, 395–400 fatigue crack propagation, 387–95 S–N fatigue, 373–86 metallurgy and weld performance of friction stir welds, 314–400 metallurgy of friction stir welds, 317–60 Al-based dissimilar welds, 343–60
AlMgSc alloy and al-based dissimilar joints, 319–42 see also specific alloy AlxMgy intermetallic phases, 344 AMAG ProCath aluminium cathode, 156 American Welding Society standard, 172 analytical model, 280–92 contact condition, 282–5 plunge force and torque in a 2024T3 FSW, 284 thin shear layer in weld region, 285 equilibrium between friction and shear stress, 289, 291–2 velocity profiles, 291 frictional and plastic dissipation, 286–8 field concentrates in thin shear layer, 287 heat source, 280–2 heat generation contributions, 282 local heat generation, 285–6 temperature-dependent heat source, 288–9 shear stress data for friction stir processed 7075-T6, 290 yield stress vs temperature curves, 290 analytical sensing, 193–5 location of resulting wormhole for bad weld, 196 manually identified frequency patterns of X and Y force, 194 phase space trajectory for y force vs derivative of the y force, 195 anti-phase boundary, 340 applied weld load, 223 Ares I, 131, 132 Ariane 5 motor thrust frames, 132 as welded, 373 Ashby-Brown contrast, 332 ASTM E 466-96 2002, 365 ASTM E 647, 373 ASTM E-837, 219–20 ASTM G10 1997, 362 ASTM G129 2000, 368 ASTM G44-99 standard, 383 Audi R8, 150 automotive closures, tanks, suspensions and pistons by international suppliers, 147–54 automotive parts for attaching skins to space-frames, 150 CNC controlled FSSW gun on
Index articulated arm robot at Friction Stir Link, 154 FSW suspension struts by Showa Denko, 149 pistons assembly, FSW and machining at Halla Climate Control, 152 tailor welded blank by Riftec for Audi R8, 151 Volvo rear car seats FSW production at Sapa, 152 welded suspension links for Lincoln stretched limousines, 149 FSW applications, 144–5, 147–55 FSW flatbed trailer of Fontaine Trailers, 155 FSW tailor welded blank for BMW and Land Rover, 147 light alloy wheels FSW production, 156 trailers in USA, 155 wheels in Australia, China and Norway, 155 AZ31, 48, 354 Barkhausen Noise Analysis, 221 base material, 324, 348 AA6013-T4/X5CrNi18-10 joint and AA 6013-T4, 380 and different FSW zones, 336 BM vs SZ microstructure of Al alloy and Mg alloy, 357 constant amplitude, variable amplitude and flight simulation loading, 393 curves of clad 2024A-T3 base material and FSW joints, 377 different grain structures observed in BM, TMAZ and SZ, 326–7 precipitate features in AlMgSc alloy, 331–2 vs AA 6013-T6 base material tested in 3.5% NaCl solution and in air at R = 0.1, 390 tested in 3.5% NaCl solution and in air at R = 0.8, 391 vs actual one predicted fatigue life, 399 vs FSW material for every single loading spectrum, 394 Bizerba, 161 BM, see base material bobbin tool, 51, 169 Boeing Company, 4, 5, 9, 119 Boeing Delta II rocket, 190 Boeing Delta IV rocket, 190
413
Boeing’s C-17 cargo ramp, 133 liquid-oxygen and liquid-hydrogen tanks for Common Booster Cores, 131 Bosch Rexroth Indramotion MTX, 89 BS EN ISO 151614-2:2005, 11 Burgers vector, 335 CAD/CAM packages, 104 ceramic tools, 174 cermets, 173 char, 179 chromium nitride, 63 clamping, see specific clamping CNC, see computer numeric controller cobalt, 175 coffee-bean-like stress contrast of coherency, 332 coherency strain, 332 cold lap bonding flaw, 257 Com-stir, 168 compliance technique, see slitting technique computer numeric controller, 81, 87, 89, 90, 98, 154 constant-volume relationship, 47 contour method, 219 convection, 49 convective heat transfer, 49 conventional clamping, 101 copper, 29, 30–1, 44, 110, 177–8 corrosion behaviour, of friction stir welds, 360–71 base alloy heat treatment, 367–71 corrosion under fatigue, 364–7 environmentally-induced cracking in heat-thermo-mechanically affected zone, 361 fracture surface of weld tested in NaCl solution stress 158 MPa, 60% yield stress, 367 stress 237 MPa, 90% yield stress, 366 general aspects, 360–4 intermetallic particle anodic dissolution after NaCl solution exposure, 363 localised corrosion after exposure to sodium chloridehydrogen peroxide solution, 362 within heat affected regions, 361 stress cycles for weld and parent metal tested in air and in NaCl solution, 365
414
Index
corrosion–fatigue tests, 365 Coulomb’s law, 280, 283 crack propagation, 224 critical buckling load, 223 critical temperature, 45 cylindrical vacuum clamping fixture, 104 data offline monitoring, 199–200 uses, advantages and disadvantages, 200 dead metal zones, 62 decompose, 179 Delta II rockets, 4, 130 Densimet, 110 design-of-experiment methodology, 184, 186, 196 FSW process schematic, 187 objectives, 187 DiamondBlade LLC, 162 diffraction-based techniques, 220 DIN 6700, 141 DISPAL, 159, 160 distortion, 221–3 buckling distortion, 223 common modes during welding, 222 weld zone thickening, 222 DLR, 372, 376 Donovan Group friction stir welding, 129 FSW ship panels at, 129 dual-rotation FSW, 109 dye penetrant inspection, 201–2 metallurgical cross section of LOP discontinuity, 203 uses, advantages and disadvantages, 202 dynamically recrystrallised zone, 199 EADS CRC F, 376 EADS IW, 372 Eclipse Aviation Corporation, 5, 11, 22, 187 Eclipse 500, 134, 187, 190 first test flight, 135 gantry for welding stringers and spars to aluminium cabin panels, 135 jet aircraft, 178 lower panel interior, 136 process window for forge force, 188 right skin assembly with oval emergency hatch, 136 ECSS-E30-01A 1999, 372 EDX analysis, 377 electromagnetic induction, 204
electromagnetism, 204 electron backscattered diffraction, 343 EMFASIS, 253 energy dispersive X-ray, 357 EP 0 615 480, 8 epoxies, 179 equipment, for friction stir welding, 73–115 closed-loop control and FSW, 84–9 controller advantage/disadvantage comparison, 90 controller diagram with inner and outer control loops, 85 PID diagram, 86 simple control system architecture, 87 simple FSW controller data flow, 88 force/position hybrid control system chart, 93 key welding tool design features, 110 machine requirements adjustable-pin closing out pin hole, 77 adjustable-pin maintaining pin penetration, 77 adjustable-pin mode, 76 fixed-pin mode, 76 FSW self-reacting mode, 78 overview, 75–84 process-to-machine requirements relationship, 79 sensor-to-axis types, 80 machine requirements overview controller requirements, 82–4 required accuracies and repeatability, 81–2 required sensors, 80–1 typical machine tool accuracies, 82 work envelope, 78–80 machine requirements summary, 98 machines currently available in market, 111–15 Aero system, 111, 113 low cost FSW system, 112, 114 Marine Aluminium FSW, 115 nuclear fuel waste encapsulation for long term repository, 112–13 panel production plant, 114 process/production development system, 111, 113 R-D FSW, 115 Robotic Weld Tool, 111, 112
Index Robotic welding system, 111, 114 other controller requirements, 95–7 complex FSW controller data flow, 96 deflection compensation, 95 multi-axis kinematic and coordinate transformations, 95 seam-tracking, 95 surface sensing, 97 thermal compensation, 97 volumetric compensation, 95 other machine requirements, 97–8 maintenance, 98 safety, 97–8 test and calibration stands, 97 viewing systems, 97 weldhead to perform milling operations, 97 part tooling requirements, 99–107 backing bar tolerance to weld table, 100 compliant roller on Esab’s SuperStir machine, 105 conventional clamping with pressure bar and clamping claws, 101 conventional concept of rollers beside FSW tool, 106 cylindrical vacuum clamping fixture, 104 hydraulic clamping system, 102 lap joint Z-stringer clamping, 101 NASA gore panel clamping, 107 roller and surface sensor system on MTS, 106 self-reacting clamping, 102 thermal conductivity for backing bars, 100 typical vacuum clamping configuration, 103 pin tools, 107–11 bobbin FSW tool, 108 variable- and fixed-gap bobbin FSW tools, 109 process and applications requirements, 73–5 robotic FSW control, 89, 91–5 data monitoring and storage, 95 data structures, 91–2 force control, 92 force/position real-time control, 92–3 program offsets automatic application, 93–4
415
teach and automatic modes, 94 travel and work angles automatic application and programming, 94 Esab customised SuperStir system, 113, 114 ESAB Legio series, 164, 166 Esab SuperStir system, 105 Eulerian framework, 294–5 Euro-Composites, 158 European Aviation Safety Agency, 136 EuroStir project, 144 excessive flash, 249 Falcon 1, 131 Falcon 9, 131 FALSTAFF loading, 385, 392, 393, 395 fatigue crack propagation, 387–95, 399–400 and fatigue predictions, 395–400 base vs AA 6013-T6 base material tested in 3.5% NaCl solution and in air at R = 0.1, 390 tested in 3.5% NaCl solution and in air at R = 0.8, 391 base vs AA 6013-T6 FSW material tested in 3.5% NaCl solution, 389 tested in 3.5% NaCl solution and in air at R = 0.1, 391 tested in 3.5% NaCl solution and in air at R = 0.8, 392 tested in 3.5% NaCl solution at R = 0.1, 389 tested in 3.5% NaCl solution at R = 0.8, 390 constant amplitude, variable amplitude and flight simulation loading for base material specimens, 393 for friction stir welded specimens, 393 environmental effects, 387, 389–92 predicted fatigue life base material vs actual one, 399 friction stir material vs actual one, 400 residual stress effect, 387 FCP in weld nugget parallel to the weld, 388 tests carried out with transverse welded specimens, 389 variable amplitude loading, 392–5 base material vs FSW material for every single loading spectrum, 394 fatigue test, 253
416
Index
fatigue-crack initiation, 224 FCP, see fatigue crack propagation Fe2Al5 intermetallic phase, 348 Federal Aviation Administration, 11 FexAly intermetallic phases, 344, 345, 347, 351 Fjellstrand, 119 Ford GT centre tunnel welding, 148 friction stir welded aluminium centre tunnel, 148 sports car, 147 FRANC2D/L FEM code, 395 friction stir processing, 168 Friction Stir Spot Welding, 145, 169 CNC controlled FSSW gun on articulated arm robot at Friction Stir Link, 154 Kawasaki uses FSSW for aluminium roof panel production, 147 Mazda RX-8 rear door outer panel FSSW, 153 Mazda RX-8 rear doors robotic FSSW, 153 friction stir welding, 164–80 90° corner joint with inadequate heat sink, 35 anvil corner for producing small fillets, 33 bobbin tool weld, 51 chronology of production applications, 21 common FSW joint configurations, 25 common joint designs with anvil sections and fixture force requirements, 32 controller requirements, 82–4 controller advantage/disadvantage comparison, 90 create and execute a weld schedule, 83 monitor process and system feedback, 84 perform coordinated motion control, 83 perform data acquisition, 84 perform position and/or load control, 83–4 provide control necessary to support type of welding, 84 travel and work angle description, 83 conventional FSW fixture requirements, 18
tool and key variables, 16 transverse section in 25.4-mm thick 2195 aluminium-lithium plate, 17 defects from too cold welds, 255–9 FSW forces developing around the FSW tool pin, 259 stir zone containing fractured FSW tool pin in Al-Cu-Mg-Mn 2024 alloy, 256 stir zone containing large volumetric flaw, 257 defects from too geometrical mistakes, 259–65 Al-Li alloy stir zone, 262 LOP weld from the transverse view of the stir zone, 261 nonbonding of workpieces due to incorrect tool placement, 265 stir zone after bend testing, 263 defects from too hot welds, 247–55 identical processing conditions in stir zones produced in an Al-CuMg-Mn 2024 alloy, 254 stir zone weld nugget formation in an Al-Cu-Mg-Mn 2024 alloy transverse to weld direction, 250 stir zone weld nugget formation in an Al-Cu-Mg-Mn 2024 from top of the weld, 251 stress corrosion cracking initiation, 252 drop-out in butt weld, 34 economic justification, 25–7 capital investment, 26 licensing, 26 processing time/labour, 26 production volume, 27 effects and defects, 245–72 flaw types, 247 mechanical properties of FSW aluminium alloy, 248 weld quality index, 248 equipment, 73–115 closed-loop control and FSW, 84–9 machine requirements overview, 75–84 machine requirements summary, 98 machines currently available in market, 111–15 other controller requirements, 95–7 other machine requirements, 97–8 overview, 74 part tooling requirements, 99–107
Index
pin tools, 107–11 process and applications requirements, 73–5 robotic FSW control, 89, 91–5 features without significant effect, 265–72 natural aluminium surface layers formed on 6013 T4-alloy, 267 natural surface scales on aluminium and aluminium alloys, 266 pre-oxidised Al sheets with natural surface scales, 266–72 root- and nugget flaws formed by natural aluminium surface layers in 6013 T4-FSWs, 269 SEM image of root flaw in 6013 T4-FSW, preoxidised at room atmosphere, 270 SEM image of root flaw in 6013 T4-FSW, preoxidised in water vapour, 270 TEM image of nugget flaw in 6013 T4 FS-weld preoxidised in water vapour, 271 TEM image of root flaw in 6013 T4 FS-weld preoxidised in water vapour, 271 flow streams representative of marker implant location, 60 FS welds in aluminium alloys metallurgy and weld performance, 314–400 corrosion behaviour, 360–71 mechanical properties, 371–400 metallurgy, 317–60 growth of licenses, 165 heat flow pathways, 55 indexed publications on FSW since 1999, 316 industrial applications, 118–63 aerospace, 130–7 automotive, 144–5, 147–55 railway, 138–44 shipbuilding and offshore, 119–29 inspection and quality control, 183–212 introduction, 1–12 history, 2–7 pre-word, 1–2 standards, 10–12 list of FSW conferences, 6 low cost system, 112, 114 main process variables, 29 material aspects, 172–9 dissimilar materials welding, 178–9
417
high temperature alloys, 174–6 low temperature alloys, 176–8 themoplastics, 179 tools and materials for high temperature use, 173–4 material deformation and joint formation, 42–68 material flow and property relationships of resultant friction stir welded joint, 58–66 plastic deformation in relation to material properties, 44–6 process parameter, temperature and heat loss relationships, 46–58 thermo-mechanical joining process, 42–4 numbers of published papers on FSrelated topics, 8 offline testing: non-destructive testing, 196–207 data offline monitoring, 199–200 dye penetrant, 201–2 phased array eddy current, 204–7 radiography, 200 typical types of FSW defects, 197 ultrasonic inspection, 203–4 visual inspection, 196–9 wormhole and oxide layers FSW defects, 198 online and offline inspection techniques summary, 210–11 online monitoring and statistical process control, 188–95 analytical sensing, 193–5 process error band, 189 process parameters, 189–90 visual monitoring, 192–3 weld path errors, 192 weld temperature, 190–1 other industry sectors, 156–63 aluminium freeze drying trays by Riftec, 162 backing plates for sputter targets in Japan, 160 cathode sheets in Austria, 156–7 copper canisters for nuclear waste in Sweden, 162 FSW-Cathode sheet of AMAG, 157 heat sinks in Europe, 157–8 heating, ventilating and air conditioning units in Germany, 159–60 hunting knives in USA, 162–3
418
Index
motor and loudspeaker housings in Scandinavia, 157 panels and components for food industry in Denmark and Germany, 160–2 vacuum vessels in Switzerland, Japan, New Zealand and Germany, 160 vacuum-tight X-ray equipment by Riftec for Siemens Medical Solutions, 161 phased implementation approach, 184 phases, 18–19 main phase, 19 plunge phase, 18–19 terminal phase, 19 process aspects, 167–72 comparison with other processes, 167–8 process modelling, 172 process variants, 168–9 productivity, 171 standards, 172 tool design, 169–71 process development and qualification, 185–6 part gap and mismatch, 186 part material consistency, 185–6 part tooling, 186 pin tool, 186 sealants, 186 process overview, 15–38 joint geometries, 31–5 joint preparation, 35–6 materials used, 29–31 parameter effects, 28–9 post-weld heat treating, 37–8 principles, 15–19 vs other welding processes, 19–27 welding tools, 27–8 quality-related parameters, 185 related growth of patent applications, 167 removal of cladding on root side of a point, 33 residual stress, 215–41 active control, 234–41 effects, 221–5 mitigation of stress and its effects, 225–31 stress determination, 219–21 stress development, 216–19 technical justification, 20–4
elimination of under matched filler metal, 22 improved cosmetic appearance, 22 improved fatigue, corrosion and stress corrosion cracking performance, 22 improved static strength and ductility, 22–3 improved weldability, 20–1 joint design limitations, 24 mechanised process, 23–4 reduced distortion, 21 special fixture requirements, 24 weld keyhole, 24 temperature profiles as measured using K-type thermocouples, 53 thermal modelling, 277–312 analytical model of heat generation, 280–92 FSW setup with workpiece, 279 numerical thermal model, 292–312 three main heat generation contributions, 279 underlying patents, 7–10 granted vs filed patents, 9 main topics of granted patents, 10 third party patent filing since 1996, 9 weld quality requirements definition, 183–8 frictional dissipation, 286 FSSW, see Friction Stir Spot Welding FSW, see friction stir welding galvanic corrosion, 179 GE Fanuc Series 3Xi models, 89 General Dynamics Land Systems, 119 global mechanical tensioning, 229, 231, 236–8 cross sectional profile of longitudinal residual stresses, 238 effect of load on tensile residual stress profile, 237 integrally stiffened panels, 236 plane distortion and peak longitudinal residual stress as a function of applied load, 237 H13 tool steel, 4, 28, 109 Hall-Petch model, 340 Hammerer Aluminium Industries FSW gantry machine, 158 produces wide range of FSW façade paned, 158
Index HAZ, see heat affected zone heat affected zone, 17, 54, 199, 317, 318, 348, 361, 367, 378, 383 heat generation, 278 analytical model, 280–92 contact condition, 282–5 equilibrium between friction and shear stress, 289, 291–2 frictional and plastic dissipation, 286–8 heat source, 280–2 local heat generation, 285–6 temperature-dependent heat source, 288–9 heat treatable aluminium alloys, 42 high strength steels, 343 weld to Al alloy, 345–54 high temperature alloys for FWS, 174–6 nickel alloys, 176 steels, 175 titanium, 174–5 high-speed tool steel, 4 Hitachi, 5, 9 hole-drilling, 219–20 hook-in effect, 253 hot working, 45 deformation process, 43 HSS, see high strength steels hydraulic clamping system, 102 Hydro Aluminium, 4, 118–19 Inpro, 150 insufficient joint penetration, 198 iron, 175 ISO 25239, 172 ISO/DIS 25239, 11 JAA-FAA 1998, 372 Johnson-Cook material law, 289 joint formation and material deformation in FSW, 42–68 material flow and property relationships of resultant friction stir welded joint, 58–66 plastic deformation in relation to material properties, 44–6 process parameter, temperature and heat loss relationships, 46–58 thermo-mechanical joining process, 42–4 K-type thermocouples, 53
419
KHI, 9 lack of fusion, 246 lack of penetration, 246 phased array eddy current detection, 208 ultrasonic testing detection, 205 visual inspection detection, 199 Lagrange framework, 295–6 lap joint Z-stringer clamping, 101 laser welding, 180 layering methods, 220 lead, 29 Lincoln Town Cars, 149 liquid penetrant inspection, see dye penetrant inspection lithium, 44 Lloyd’s Guideline for Approval of Friction Stir Welding, 10 local mechanical tensioning, 238–41 effect of post-weld roller tensioning on longitudinal residual stresses, 239 effect of roller tensioning on distortion of fabricated panel, 240 integrally stiffened panels, 240 roller tensioning system, 239 LOF, see lack of fusion LOP, see lack of penetration low temperature alloys for FWS, 176–8 aluminium alloys, 176–7 copper, 177–8 magnesium, 177 magnesium, 29, 44, 177 magnesium alloys, 343 magnetoelastic method, see Barkhausen Noise Analysis manganese, 44 Marine Aluminium friction stir welding, 115 FSW module for Finnmarken, 122 living quarters with Marine Aluminium’s FSW panels for offshore platform Oseberg, 123 prefabricated FSW deck panels, 121 material deformation affecting quantitative factors, 67–8 and joint formation in FSW, 42–68 material flow and property relationships of resultant friction stir welded joint, 58–66
420
Index
plastic deformation in relation to material properties, 44–6 process parameter, temperature and heat loss relationships, 46–58 thermo-mechanical joining process, 42–4 Mazda RX-8, 151 rear door outer panel FSSW, 153 rear doors robotic FSSW, 153 mechanical tensioning, 231 MP159, 28, 110 MTS I - Stir series, 166 MTS systems corporation, 11
function of rotational speed, 311 maximum temperatures as a function of welding speed, 312 maximum yield stress, 305 near-field view of temperature field, 306 temperature evaluation, 308 thermal properties, 304 vs experimental results for far-field temperature, 307 welding conditions, 304 nylon, 179
NASA, 5, 11 Constellation Ares I, 105 first official weld with tools for Ares I upper stage hardware test articles, 132 gore panel clamping, 107 Marshall Space Flight Systems Center, 105 NASA-ESA 2000, 395 Vertical Weld Tool at Marshall Space Flight Center o, 132 nickel, 175, 180 nickel alloys, 176 Nimonic 105, 110 Nippon Sharyo, 190 NLM, 9 nugget collapse, 251 nugget–thermo-mechanically affected zone, 365, 366, 367, 370 numerical modelling, 292–312 convective heat transfer, 296–9 conical shear layer, 297 Eulerian framework, 294–5 stationary heat source and global Lagrange model, 294 governing equations, 293–4 heat transfer to surroundings, 299–302 Lagrange framework, 295–6 thermal efficiency, 302–3 thermo-pseudo-mechanical model, 303–12 calculations of heat generation contributions, 309 far-field temperature field, 305 heat generation contour plot, 311 local heat generation distribution, 309 maximum temperature contour plot, 310 maximum temperatures as a
Ökolüfter, 159 optical microscopy, 343 Orowan mechanism, 340, 342 Orowan model, 341 Orowan’s dispersion hardening, 335 P135E, 202 P64F4, 202 phased array eddy current, 204–7 configuration, 208 detection of FSW lack of penetration, 208 eddy current principle, 206 galling detection, 209 oxidising layer detection, 209 uses, advantages and disadvantages, 207 PID, see proportional-integral-derivative pin tools, 107–11 bobbin FSW tool, 108 variable- and fixed-gap bobbin FSW tools, 109 plastic dissipation, 286 plastic pre-bending, 225–6 polycrystalline cubic boron nitride, 110, 162, 173 polyether ether ketone, 179 polyethylene, 179 polypropylene, 179 pressure welding techniques, 344 Process Specification of Friction Stir Welding, 11 process/production development system, 111, 113 programmable automation controller, 87, 89, 90, 98 programmable logic controller, 87, 89, 90, 98 proportional-integral-derivative, 84, 86
Index
421
pseudoboehmite, 266
residual stress and distortion control, 227–31 stress development, 216–19 residual profile across the weld centreline, 218 temperature and stress field, 217 temperature stress history, 217 Revolution, 155 Riftec, 160 aluminium freeze drying trays, 162 tailor welded blank for Audi R8, 151 vacuum-tight X-ray equipment for Siemens Medical Solutions, 161 RoboStir, 111, 114 Robotic Weld Tool, 111, 112 Robotic welding system. see RoboStir
radiography, 200 large and small tunnels of butt weld, 201 schematics, 200 uses, advantages and disadvantages, 201 railway aluminium panels for railway rolling stock in Europe, 138–44 FSW applications, 138–44 rolling stock in Japan, 144 commuter train built by Hitachi, 145 floor panel produced by Sumimoto Light Metal for Shinkansen train, 146 Kawasaki uses FSSW for aluminium roof panel production, 147 new Channel Tunnel Rail Link domestic trains, 145 trainsets with floor panels of Sumimoto Light Metal, 146 Re-Stir, 109 recovery, 45 recrystallisation, 45 affecting variables, 46 amount of prior deformation, 46 amount of recovery or polygonisation, 46 composition, 46 initial grain size, 46 temperature, 46 time, 46 residual stress active control, 234–41 global mechanical tensioning, 236–8 local mechanical tensioning, 238–41 thermal tensioning, 234–6 determination of residual stress, 219–21 destructive methods, 219–20 non-destructive methods, 220–1 effects, 221–5 distortion, 221–3 friction stir welding, 215–41 cross-sectional profile of longitudinal stresses, 233 process effects, 232–4 tensile residual stress values, 234 mitigation of stress and its effects, 225–31 management of stress, 225–6
S–N fatigue, 373–86 clad material, 376–8 2024A-T3 base and FSW joints S-N fatigue curves, 377 failure location of clad 2024A-T3 friction stir welded joints, 377 curves of clad 2024A-T3 base material and FSW joints, 377 dissimilar materials, 378–80 AA6013-T4/X5CrNi18-10 joint and AA 6013-T4 base material, 380 cross-section and hardness curve of dissimilar joint, 378 fatigue strength of dissimilar FSW joints, 379 friction stir welded AA 6013-T4 alloy to X5CrNi18-10 stainless steel, 380 environmental effects, 383–4 experimental results vs computer simulations base and FSW pre-corroded material, 398 FSW open hole specimens under FALSTAFF, 397 prediction of base and FSW welded material fatigue lives, 396 T-jont R = 0.1, 397 fatigue and FCP predictions, 395–9 fatigue strength of as welded and polished FSW specimens, 373 of skimmed, notched specimens, 376 not corroded vs corroded polished base and FSW material, 384 notch effects, 375–6
422
Index
T-joints, 380–3 fatigue lives of 12.5mm vs 25mm width FSW T-Joints, 382 hardness profile of T friction stir welded joints, 381 thickness effects, 374–5 friction stir welded 2024-T3, 375 variable amplitude loading, 385–6 centre hole FSW specimen under FALSTAFF spectrum loading experimental results, 386 hardness profile of AA 2024-T3 friction stir welded material, 385 weld surface effects, 373–4 Sapa aluminium panels for railway rolling stock, 138–44 Alstom LHB trains for DSB Danish State Railways during production, 139 aluminium railcar body for London’s Victoria Lane, 142 friction welded roof panel for delivery to Alstom LHB, 139 high speed impact test of new joint design, 143 pre-fabricates curved train side skirt panels by FSW, 140 side skirt panels from doubleskinned extrusions, 141 surface finish after painting, 141 FSW freezer panel for fish blocks pre-pressing, 120 joint design, 120 scandium, 44 scanning electron microscopy, 268, 270, 343 image of root flaw in 6013 T4-FSW, preoxidised at room atmosphere, 270 SEM image of root flaw in 6013 T4-FSW, preoxidised in water vapour, 270 Sea Fighter, 126 selected area diffraction pattern, 329, 351 self-reacting clamping, 102 self-reacting tool, 169 SEM, see scanning electron microscopy servo-hydraulic testing machine, 373 Shinkansen high-speed train, 190 ship-in-a-bottle design concept, 148 shipbuilding and offshore 5 ¥ 12m large aluminium ship panels, 127
Donovan Group friction stir welding, 129 FSW ship panels at, 129 explosively formed bow section of ocean viewer vessel, 125 Fosen Mek’s cruise ship “The World,” 122 friction stir welded superstructure for inshore patrol vessel Rotoiti, 130 friction stir welds on starboard side of bow section, 124 FSW applications, 119–29 explosively formed hull of an ocean viewer vessel in Australia, 124–5 freezer panels in Sweden and Germany, 119–21 honey comb panels and corrosion resistant panels in Japan and USA, 125–8 patrol vessels in New Zealand, 128–9 ship and oil-rig panels for decks and bulkheads in Scandinavia, 121–3 FSW module for Finnmarken at Marine Aluminium, 122 helicopter landing platforms and offshore floor panels, 123 living quarters with Marine Aluminium’s FSW panels for offshore platform Oseberg, 123 Marine Aluminium’s prefabricated FSW deck panels, 121 portable FSW machine at Research Foundation Institute shipyard, 124 Super Liner Ogasawara, 126 Type 022 new-generation stealth missile fast attack craft, 128 US Navy X-Craft Sea Fighter bow, 126 shot peening, 226 Showa, 9 Siemens 805, 89 Siemens 840D, 89 silicon, 44 silicon nitride, 174 Sinumerik 840D control system, 150 Skew Stir, 109, 168 skimming, 373 Slipper, 119 slitting technique, 219 solid-state joint, 43 solidus temperature, 44–5, 248 Space Shuttle External Tank, 131 SpaceX, 131
Index spot welding, 178 steel, 29, 30, 175, 180 stir zone, 43, 317–18, 348 after bend testing, 263 Al-Li alloy stir zone, 262 BM vs SZ microstructure of Al alloy and Mg alloy, 357 containing fractured FSW tool pin in Al-Cu-Mg-Mn 2024 alloy, 256 containing large volumetric flaw, 257 different grain structures observed in BM, TMAZ and SZ, 326–7 formation comparison between weld macrographs, 50 identical processing conditions produced in an Al-Cu-Mg-Mn 2024 alloy, 254 LOP weld from the transverse view of the stir zone, 261 phenomenon of increasing deformation volume during FSW, 58 systematic and non-chaotic band layering on flaw-free SZ, 63 SZ size and nature of banding, 65 weld nugget formation in an Al-Cu-Mg-Mn 2024 alloy transverse to weld direction, 250 in an Al-Cu-Mg-Mn 2024 from top of the weld, 251 strain-field contrast, 332 stress engineering, 227 Super Liner Ogasawara, 125–6 superplastic forming/diffusion bonding, 169 SuperStir, 113, 114 surface galling, 249 Svensk Kärnbränslehantering, 162 SZ, see stir zone tandem FSW, 109 TEM, see transmission electron microscopy temperature stress cycle, 227–8 large tensile load, 229 small tensile load, 228 The Welding Institute, 3, 7–8, 27, 28, 63 thermal efficiency, 302–3 thermal modelling analytical model of heat generation, 280–92 contact condition, 282–5 equilibrium between friction and shear stress, 289, 291–2
423
frictional and plastic dissipation, 286–8 heat source, 280–2 local heat generation, 285–6 temperature-dependent heat source, 288–9 friction stir welding, 277–312 numerical modelling, 292–312 convective heat transfer, 296–9 Eulerian framework, 294–5 governing equations, 293–4 heat transfer to surroundings, 299–302 Lagrange framework, 295–6 thermal efficiency, 302–3 thermo-pseudo-mechanical model, 303–12 thermal softening, 249 thermal straightening, 226 thermal tensioning, 230, 234–6 cryogenic CO2 delivery system, 235 effect of cooling on longitudinal residual stress profile, 235 longitudinal stress fields, 230 optimum positioning of cold and hot zones, 231 thermo-mechanical joining process, 42–4 principal processing parameters axial force, 42 tool rotation speed, 42, 54 weld travel speed, 42, 54 thermo-mechanically affected zone, 17, 250, 317, 318, 346, 348, 362, 364, 365 and nugget of 7050-T7451 FSW microstructure, 366 different grain structures observed in BM, TMAZ and SZ, 326–7 Orowan’s mechanism occurring in TMAZ of hot parameter, 335 weld fracture along nugget–TMAZ interface, 371 thermo-pseudo-mechanical model, 289, 303–12 calculations of heat generation contributions, 309 far-field temperature field, 305 heat generation contour plot, 311 local heat generation distribution, 309 maximum temperature contour plot, 310 maximum temperatures as a function of rotational speed, 311
424
Index
maximum temperatures as a function of welding speed, 312 maximum yield stress, 305 near-field view of temperature field, 306 temperature evaluation, 308 thermal properties, 304 vs experimental results for far-field temperature, 307 welding conditions, 304 thermoplastics, 179 thermosets, 179 titanium, 29, 30, 110, 169, 174–5 titanium nitride, 63 TMAZ, see thermo-mechanically affected zone tool coatings, 63 tooling/tack welding, 226 transmission electron microscopy, 58, 267–71, 343 image of nugget flaw in 6013 T4 FS-weld preoxidised in water vapour, 271 image of root flaw in 6013 T4 FS-weld preoxidised in water vapour, 271 Tricept-9000 parallel kinematics robot, 150 Trivex design, 172 tungsten, 110 tungsten rhenium alloys, 173 tungsten rhenium tools, 110 TURCO 5580G, 383 Type 022 Houbei Class, 128 ultrasonic inspection techniques, 203–4, 220–1 detection of FSW lack of penetration, 205 general principle, 204 types linear, 204 sector, 204
uses, advantages and disadvantages, 205 weld testing, 206
vacuum clamping systems, 101, 104 typical configuration, 103 Vertical Weld Tool, 132 Victoria Line, 141 visual inspection, 196, 198–9 FSW defects, 198 lack of penetration defects, 199 Voltapotential, 362 Volvo V70 station wagon, 151 weld path errors, 192 vision and seam-tracking equipment, 193 process specification, 99 quality requirements definition, 183–8 temperature, 190–1 infrared radiation thermometry sensing, 191 temperature probe inside FSW pin tool, 190 thermocouple placement in backing plate, 192 weld nugget, 43 FCP parallel to the weld, 388 formation in an Al-Cu-Mg-Mn 2024 alloy transverse to weld direction, 250 formation in an Al-Cu-Mg-Mn 2024 from top of the weld, 251 weld sequencing, 226 Willenborg model, 395, 399 X-ray diffraction, 357 Zener-Hollomon Parameter, 48 zinc, 44 zirconium, 44
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