Hydroforming for advanced manufacturing
WPNL2204
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WPNL2204
Hydroforming
for advanced
manufacturing
Edited by
Muammer Koç
Woodhead Publishing and Maney Publishing
on behalf of
The Institute of Materials, Minerals & Mining
WPNL2204
CRC Press
Boca Raton Boston New York Washington, DC
Cambridge England
WPNL2204
Woodhead Publishing Limited and Maney Publishing Limited on behalf of
The Institute of Materials, Minerals & Mining
Woodhead Publishing Limited, Abington Hall, Granta Park, Great Abington
Cambridge CB21 6AH, England
www.woodheadpublishing.com
Published in North America by CRC Press LLC, 6000 Broken Sound Parkway,
NW, Suite 300, Boca Raton, FL 33487, USA
First published 2008, Woodhead Publishing Limited and CRC Press LLC
© 2008, Woodhead Publishing Limited
The authors have asserted their moral rights.
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WPNL2204
Contents
1
1.1
1.2
1.3
1.4
1.5
1.6
1.7
Contributor contact details
Preface
xi
xv
Introduction and state of the art of hydroforming
M. Koç and O. N. Cora, Virginia Commonwealth
University, USA
Introduction
Hydroforming systems, equipment, tooling and controls
Materials, formability, forming limits and advantages
Tribology in hydroforming – friction, wear,
lubricants, coatings and testing methods
Computer simulations for tube hydroforming
Developments in hydroforming and concluding remarks
References
1
Part I: Principles of hydroforming
2
2.1
2.2
2.3
2.4
2.5
2.6
2.7
2.8
Hydroforming systems, equipment, controls
and tooling
D. Gearing and D. Mevissen, Interlaken Technology
Corporation, USA
Introduction
Presses
Pressure intensification systems
Controls
Tooling
Future trends
Sources of further information and advice
References
1
10
15
17
21
24
28
31
33
33
33
39
43
45
50
51
51
v
WPNL2204
vi
Contents
3
Deformation mechanism and fundamentals
of hydroforming
C. Hartl, Cologne University of Applied Sciences, Germany
Introduction
Stress and strain relationships in tube hydroforming
Determination of forming limits
Forming loads and process control
Preceding forming operations
References
3.1
3.2
3.3
3.4
3.5
3.6
4
4.1
4.2
4.3
4.4
4.5
4.6
5
5.1
5.2
5.3
5.4
5.5
5.6
5.7
5.8
6
6.1
6.2
6.3
6.4
6.5
6.6
6.7
Materials and their characterization for hydroforming
C. Hartl, Cologne University of Applied Sciences, Germany
Introduction
Steel materials
Aluminium and magnesium alloys
Formability testing
Future trends
References
Formability analysis for tubular
hydroformed parts
D. E. Green, University of Windsor, Canada
Introduction
Tube formability
Measuring tube formability
Tube-forming limits
Formability analysis for numerical simulations
Formability analysis in the plant
Conclusions and future trends
References
Design and modelling of parts, process and
tooling in tube hydroforming
M. Strano, Università degli Studi di Cassino, Italy
Introduction to tube hydroforming design
Technological classifications of tube
hydroforming processes
Hydroformability of tubular parts
Guidelines for process design
Finite element analysis strategies for process design
Designing a new hydroforming process: a simple example
References
WPNL2204
52
52
54
62
66
69
75
77
77
78
81
84
87
90
93
93
94
99
108
114
115
116
117
121
121
123
127
131
136
139
142
Contents
7
7.1
7.2
7.3
7.4
7.5
7.6
7.7
Tribological aspects in hydroforming
G. Ngaile, North Carolina State University, USA
Introduction
Parameters that influence friction, lubrication,
and wear
Lubrication mechanisms
Development and evaluation of
hydroforming lubricants
Impact of numerical modelling in
hydroforming tribology
Concluding remarks
References
Part II: Hydroforming techniques and
their applications
8
8.1
8.2
8.3
8.4
8.5
8.6
8.7
9
9.1
9.2
9.3
9.4
9.5
10
10.1
10.2
10.3
10.4
Pre-forming: tube rotary draw bending and
pre-flattening/crushing in hydroforming
G. Khodayari, Vari-Form, Canada
Introduction
Concept of rotary draw bending process
Material behavior in rotary draw bending process
Pre-flattening/crushing
Part design and tube formability
Conclusion
References
Hydroforming: hydropiercing, end-cutting,
and welding
L. M. Smith, Oakland University, USA
Introduction
Hydropiercing
End-cutting and saw-cutting
Welding
References
vii
144
144
148
154
163
172
177
177
179
181
181
182
183
193
197
200
200
202
202
202
209
212
215
Hydroforming sheet metal forming components
216
K. Siegert and S. Wagner, University of Stuttgart, Germany
Introduction
216
Hydroforming processes
216
Dies and presses for hydromechanical deep drawing
234
References
237
WPNL2204
viii
Contents
11
Bending and hydroforming of aluminum
and magnesium alloy tubes
A. A. Luo and A. K. Sachdev, General Motors
Research & Development Center, USA
Introduction
Aluminum and magnesium alloy tubes
Aluminum tube bending and hydroforming
Magnesium tube bending
Forming at elevated temperatures
Automotive applications and future trends
Acknowledgments
References
11.1
11.2
11.3
11.4
11.5
11.6
11.7
11.8
12
12.1
12.2
12.3
12.4
12.5
12.6
13
13.1
13.2
13.3
13.4
13.5
13.6
13.7
13.8
13.9
13.10
13.11
14
14.1
14.2
14.3
14.4
14.5
238
238
239
248
255
257
260
263
263
Low-pressure tube hydroforming
G. Morphy, Excella Technologies Inc., Canada
Introduction
Low-pressure hydroforming
Part characteristics
Cross-section pre-forming
Conclusions
Reference
267
Comparative analysis of hydroforming techniques
G. Morphy, Excella Technologies Inc., Canada
Introduction
So many options; how to choose?
Roll forming
Stampings and assemblies
Tube forming
Commonly held misconceptions
Low-pressure hydroforming
High-pressure hydroforming
Other comparative process factors
Conclusions
References
287
Fluid cell pressing in the aerospace industry
M. Bergkvist, Avure Technologies AB, Sweden
Introduction
Evolution of the technology
How fluid cell pressing works
Recent developments
Essentials of ductile materials
315
WPNL2204
267
268
271
277
285
286
287
288
289
289
290
298
299
302
308
313
314
315
315
317
319
323
Contents
14.6
14.7
14.8
14.9
Suitable part applications
Tools (form blocks)
Part manufacture
Conclusions
15
Hydroforming and its role in lightweighting
automobiles
G. Morphy, Excella Technologies Inc., Canada
Introduction
What makes it so difficult to lose weight?
How to lose weight
How tube hydroforming can help you to lose it
Weight loss limitations and how to address them
Conclusions
References
15.1
15.2
15.3
15.4
15.5
15.6
15.7
16
16.1
16.2
16.3
16.4
16.5
16.6
16.7
16.8
ix
325
326
330
333
335
335
336
337
338
348
351
351
Warm hydroforming of lightweight materials
M. Koç, Virginia Commonwealth University, USA
Introduction: motivation for lightweight vehicles
Lightweight materials: advantages and disadvantages
in manufacturing
Forming technologies for lightweight materials
Warm hydroforming: state-of-the-art review
Comparison of warm and cold hydroforming:
a numerical study
Process design and control in warm hydroforming
Characterization of materials for warm
hydroforming conditions
References
352
Index
384
WPNL2204
352
352
353
354
363
367
373
380
WPNL2204
Contributor contact details
(* = main contact)
Chapter 1
Chapter 3
Muammer Koç* and
Omer N. Cora
Mechanical Engineering Dept.
Virginia Commonwealth
University
Richmond
VA 23284-3015
USA
Email: mkoc@vcu.edu
Prof. Dr. Christoph Hartl
Institute of Production
Faculty of Automotive Systems
Engineering and Production
Engineering
Cologne University of Applied
Sciences
Betzdorfer Str. 2
50679 Cologne
Germany
Email: christoph.hartl@fh-koeln.de
Chapter 2
David Gearing* and Dennis
Mevissen
Interlaken Technology
Corporation
8175 Century Boulevard
Chaska
MN 55318
USA
Email: dgearing@interlaken.
com
Chapter 4
Prof. Dr. Christoph Hartl
Institute of Production
Faculty of Automotive Systems
Engineering and Production
Engineering
Cologne University of Applied
Sciences
Betzdorfer Str. 2
xi
WPNL2204
xii
Contributor contact details
50679 Cologne
Germany
Email: christoph.hartl@fh-koeln.de
Canada
Email: Gkhodayari@Vari-Form.com
Chapter 9
Chapter 5
Dr. Lorenzo M. Smith
118 Dodge Hall
Department of Mechanical
Engineering
Oakland University
Rochester
MI 48309
USA
Email: L8smith@oakland.edu
Dr. Daniel E. Green
Department of Mechanical,
Automotive & Materials
Engineering
University of Windsor
401 Sunset Avenue
Windsor
Ontario N9B 3P4
Canada
Email: dgreen@uwindsor.ca
Chapter 10
Chapter 6
Matteo Strano
Department of Industrial
Engineering
Università degli Studi di Cassino
via di Biasio 43
03043 Cassino (FR)
Italy
Email: m.strano@unicas.it
Chapter 7
Prof. Gracious Ngaile
Department of Mechanical and
Aerospace Engineering
North Carolina State University
Raleigh
NC 27695
USA
Email: gracious_ngaile@ncsu.edu
Prof. Klaus Siegert and
Dr. Stefan Wagner
University of Stuttgart
Institut für Umformtechnik (IFU)
Institute for Metal Forming
Technology
Holzgartenstrasse 17
70174 Stuttgart
Germany
Email: klaus.siegert@ifu.unistuttgart.de and stefan.wagner
@ifu.uni-stuttgart.de
Chapter 11
Alan A. Luo* and Anil K. Sachdev
General Motors Research &
Development Center
Warren
MI 48090-9055
USA
Email: alan.luo@gm.com
Chapter 8
Dr. Ghafoor Khodayari
233 Lothian Avenue
Strathroy
Ontario N7G 3J3
Chapter 12
Gary Morphy
Excella Technologies Inc.
82 Rife Avenue
WPNL2204
Contributor contact details
Cambridge
Ontario N3C 2G7
Canada
Email: garym@excellatechnologies.
com
Chapter 13
Gary Morphy
Excella Technologies Inc.
82 Rife Avenue
Cambridge
Ontario N3C 2G7
Canada
Email: garym@excellatechnologies.
com
Chapter 14
Mikael Bergkvist
Avure Technologies AB
Quintusvagen 2
72166, Vasteras
Sweden
xiii
Email: Mikael.Bergkvist@avure.
se
Chapter 15
Gary Morphy
Excella Technologies Inc.
82 Rife Avenue
Cambridge
Ontario N3C 2G7
Canada
Email: garym@excellatechnologies.
com
Chapter 16
Muammer Koç
Mechanical Engineering Dept.
Virginia Commonwealth
University
Richmond
VA 23284-3015
USA
Email: mkoc@vcu.edu
WPNL2204
WPNL2204
Preface
Throughout the last decade, increasing competition and environmental
regulations has forced the transportation manufacturing industry towards
producing low-mass vehicles to achieve fuel savings, reduced emissions and
safe structures. In order to accomplish this goal in a cost-effective manner,
manufacturers have to both develop or use new and lightweight materials,
alloys or composites; and develop and improve new manufacturing techniques that can convert these materials into lightweight functional structures robustly, cost-effectively and with consistent quality. The hydroforming
process, which has been long used to fabricate either intricate and small
fittings for sanitary applications, or large but rather simple parts for aerospace applications, was a relatively new technology for the automotive
industry in the early 1990s. It was taken as an new tool that could enable
part consolidation, high strength-to-weight ratios, tight tolerances, better
rigidity, less post-process operations, easy assembly, cost effective parts and
tooling. With many other advantages and opportunities of design flexibility,
there has been a continuously increasing interest in the hydroforming
process particularly within the automotive and aircraft industries. As
expected, the demand on the development of a knowledge base on this
technology is also increasing so that mass production of new and more parts
can become a reality.
As far as this author could survey and to his best knowledge, more than
one thousand technical papers have been written and published on various
aspects of this technology. Researchers, scientists, engineers from all over
the world, both from academia and industry, researched the deformation
mechanics, material behavior, tribology, tool design, equipment and system
design, process control, etc. of this technique. Analytical and numerical
models were developed and tested for a variety of parts of interest.
Yet, there is no textbook or handbook that serves as a single source of
knowledge in this field of advanced manufacturing. One or two existing
attempts were developed to the desired full extent but they do not suffice
for students, engineers and researchers who are new to the topic.
xv
WPNL2204
xvi
Preface
This handbook on hydroforming was prepared with the aim of it becoming a main reference source of knowledge for decades to come. It is a
product of a truly international and interdisciplinary collaborative work of
many contributors. It presents all aspects of this technology comprehensively in a concise and direct manner. It comprises 16 chapters, and every
chapter takes its topic from the very basics towards the latest and highest
level of know-how. Chapters are organized in a similar manner to take the
readers from fundamentals to the analysis of advanced and complex issues
as well as to the latest trends in this field of research and development.
We hope that readers of all kinds of background, need and interest will
find it useful and take it as a first step towards learning, researching, advancing and applying this technology into whatever their respective application
would be.
I would like to acknowledge and thank all of my collaborating friends
and authors for their contributions by sharing their knowledge and insight
with us. Special thanks go to the publishing team members who relentlessly,
patiently and with a great professionalism steered us towards completing
all chapters in a timely manner.
Muammer Koç
Richmond, Virginia, USA
WPNL2204
1
Introduction and state of
the art of hydroforming
M. KOÇ and O. N. CORA,
Virginia Commonwealth University, USA
1.1
Introduction
This chapter is intended to provide an introduction into the technology of
hydroforming as a briefing for the upcoming chapters of this book. In it
are summarized the fundamentals of hydroforming technology, its developmental background, hydoforming systems including equipment, tools and
controls, as well as its role in the production of lightweight structures and
vehicles. This chapter also provides introductory information on materials
used in hydroforming and their formability issues in addition to information
on hydroforming tribology, pre-forming issues such as effect of loading, path
and process control. Computer simulation techniques and innovations in
hydroforming are discussed towards the end of the chapter. Wherever possible, references to the existing published studies and to the corresponding
chapters in this handbook are given to direct the readers to the right place
for further reading and in-depth understanding.
1.1.1 Definition and examples of hydroforming
Hydroforming is a material-forming process that uses a pressurized fluid
(liquid or gas) in place of hard tooling (punch, die, mold, inserts, etc.) either
to plastically deform or to aid in deforming a given blank material (sheet
or tube) into a desired shape as depicted in Fig. 1.1. With this technique,
more complex shapes with increased strength and low cost can be manufactured as compared with stamping, forging or casting processes. The
cost advantage usually stems from the fact that fabrication steps in hydroforming are significantly reduced, usually to a single step. In stamping, for
example, multiple steps such as blanking, drawing, restriking, trimming,
welding, etc. are needed to finalize a part whereas a sheet blank can be
drawn into the final complex shape (as shown in Fig. 1.2 and 1.3 as examples) in a single step. Most of the time, additional post-processing steps such
as hole piercing or trimming may also be incorporated in this step.
1
WPNL2204
2
Hydroforming for advanced manufacturing
1
2
4
5
Faxial
3
6
Faxial
P
1.1 Steps in a typical hydroforming process shown on a small tubular
part (courtesy of Siempelkamp Pressen Systems).
a
b
1.2 Example hydroformed (sheet) parts: a 2007 GM Pontiac Solstice
GXP has several hydroformed (warm) parts, b fuel tank comprising of
two halves can be hydroformed in a single step.
a
b
1.3 Example hydroformed (tubular) parts for various automotive
applications: a hydroformed steel camshaft (BMW 3.0 L DOHC I-6
Ward’s 2005) offers 28 to 50% mass saving, b 2-piece roll bar for a
convertible car (courtesy of Schuler Inc.).
In addition to various applications in the aerospace industry, such as
panels, fuselage parts and casings and in the appliance industry, such as fittings, joints, knobs and handles, hydroforming has been increasingly used
in the automotive industry since 1990s. Various parts for the automotive,
appliance and plumbing industries are produced by hydroforming technology; they can be summarized as follows (Fig. 1.4).
WPNL2204
Introduction and state of the art of hydroforming
3
1.4 Several hydroforming applications in an automobile
(courtesy of BMW, Tower, Schuler Press, DCX, Vari-form).
Exhaust system parts
Usually made of stainless steel because of the required structural, thermal
and corrosion properties, these include exhaust parts, engine tubes, catalytic
converters, pressure tubes, tail pipes, connectors and manifolds. Exhaust
parts: BMW 3 series; Mercedes Benz E-class, Mitsubishi Carisma 1.6.
Chassis parts
The common material used is low- to medium-carbon steels and aluminum
for structural and cost-related reasons: frame rails, engine sub-frames
(cradles), roof rails, and bows, instrument panels, rear axle frames and radiator frames. Engine cradle/subframe: Audi 100, Ford Mondeo and Windstar;
Opel Astra and Omega; GM 1997 Malibu Cutlas, 1995 Aura and Riviera;
1997 Park Avenue and Pontiac Aztek, 2006 GM Corvette Z06, 2005 Honda
Acura RL, Audi A6, Saab 9–3 Convertible; Front Rail and End Parts:
Porsche Solstice, 1999 GMC full-size pickups, 2004 Ford F-150 pickup, Volvo
850; dashboard members: Audi A4 and A6; side rail: 1999 Corvette, 2006
GM Chevrolet Corvette Z06; roof rail: GM 1999 Buick Park Avenue; rear
suspension: Mercedes Benz S-Class, BMW 5 and 7 series; Rear Axle: BMW
5 series (aluminum 5xxx series); Radiator support: Dodge Dakota and
Ram.
WPNL2204
4
Hydroforming for advanced manufacturing
Engine and power train components
Suspension cross members, hollow camshafts, drive shafts and gear shafts.
Camshafts: 2005 BMW 130i, 2007 BMW X5, 2008 BMW 5 Series.
Body and safety parts
Windshield headers, A/B/C pillars, space frame components, seat frames
and shock absorber housings. Roll-over protector bar: Porsche Boxter;
Bumper beam: Porsche Solstice; A-pillar: Volvo C70; Fuel Tank: VW Golf.
The benefits of hydroforming technology are known as weight and cost
savings through part consolidation and reduced post-forming processes
such as welding and piercing. Hydroforming technology has some weaknesses in terms of process cycle times. But, as the hydraulic system and press
designs are continuously developed, the cycle time is also reduced to acceptable and competitive levels.
Another vital premise of the hydroforming process is its enabling features in fabrication of lightweight structures and parts using lightweight
materials. Approximately 80% of the total energy consumption throughout
the life cycle of an automobile occurs during the utilization period. Hence,
use of lightweight structures is accepted as a prominent and long-term solution to minimize energy consumption and the adverse impacts of transportation on the environment to truly achieve sustainable mobility (Mildenberger
and Khare, 2000). Even in automobiles with efficient and clean powergeneration systems and alternative fuels, the lightweight structures would
further increase the fuel consumption efficiency and reduce emissions (i.e.,
primary and secondary benefits). Lightweight structures can be realised by
(a) using lightweight materials such as aluminum, magnesium, high-strength
steel, titanium, metal–matrix composites (MMC) and polymer composites,
(b) the development of low-cost, robust conversion processes that enable
use effective of these materials (i.e., innovative manufacturing processes).
It was reported that fuel economy improvements of around 6–8% can be
realised for every 10% weight reduction in a vehicle. Apart from its proven
capability to manufacture complex parts cost-effectively, hydroforming is
seen as a process which enables deformation of lightweight materials into
desired shapes with fewer problems than stamping. Aluminum, magnesium
and high-strength steels are the most appropriate materials in lightweight
structures; however, their formability is very low and very sensitive to production speeds. At this point, hydroforming offers opportunities to manufacturers to form these materials by setting the fluid pressure inside the
die precisely by increasing the degree of formability. When combined with
selective heating strategies, hydroforming (warm) even goes further and
offers 100–300% increases in forming limits (Doege et al., 2002).
WPNL2204
Introduction and state of the art of hydroforming
5
1.1.2 History, types and classification of hydroforming
The history of using fluid to form metals dates back more than 100 years.
Early applications were in the forming of boilers and musical instruments.
However, the fundamentals of hydroforming were established in the 1940s
(Grey, 1939; Dohmann, 1991 and Koç, 2001). In the 1950s, alternative manufacturing processes were proposed, such as superplastic forming, explosive
forming and rubber forming, to increase the formability of aluminum and
other lightweight materials. The first patented hydroforming application
was obtained by Milton Garvin of the Schaible Company of Cincinnati,
Ohio, for producing kitchen spouts in the 1950s. Until the 1990s, making
copper plumbing Ts was the most common application. Since the 1990s,
hydroforming has made an impactful comeback due to the advancements
in computer controls, hydraulic systems and recently developed process and
part design guidelines, and various forged or stamped structural parts have
been replaced by parts formed with tube hydroforming technology (THF)
in many North American vehicles. Substantial weight and cost savings were
realized with hydroformed steel parts because of the part consolidation,
less post-forming processes (i.e., joining such as welding and piercing) and
initial thinner material thickness opportunities (Dohmann, 1991; Koç, 2001;
Murray, 1996 and Morphy, 1997).
The hydroforming process, in general, can be divided into two major
categories (a) sheet hydroforming and (b) tube hydroforming as depicted
in Fig. 1.5 (Schmoeckel, 1999).
Sheet hydroforming
In the sheet hydroforming (SHD) process, sheet blank is formed by hydraulic pressure inside the die cavity as illustrated in Fig. 1.6. This technique
allows a much deeper draw, which is necessary for manufacturing panels
with complex curves. Sheet hydroforming can be classified into two parts:
Hydroforming
Sheet hydroforming (SHF)
Hydromechanical
Deep drawing (HMD)
Tube hydroforming (SHF)
High-pressure
sheet hydroforming
Single blank
Double blank
1.5 Classification of hydroforming technology (Schmoeckel, 1999).
WPNL2204
6
Hydroforming for advanced manufacturing
Punch
Internal
pressure
Blankholder
force
Blank
Pi
Die
Die
P
Internal i
pressure
Upper Die
Blanks
Pressure
Lower die
a
b
1.6 a General description of simple sheet hydroforming operation,
b Sheet hydroformed front hood (courtesy of Schuler Inc., 1998)
MRF
BHRF
BHRF
Pc
RF
Main ram
Blankholder ram
Ram
Punch
Blankholder
(upper binder)
Blank
Draw ring
(lower binder)
Hydraulic cylinder
Punch
Blankholder
(upper binder)
Blank
Draw ring
(lower binder)
Pc
Counter pressure pot
Pc
Counter pressure pot
Pc
Press table
Press table
a
b
1.7 Hydromechanical deep drawing: a single action, b double action
(Siegert, 2000).
hydromechanical deep drawing (HMD) and high-pressure sheet hydroforming with single and multiple blanks.
Hydromechanical deep drawing
A typical layout of a HMD process is shown in Fig. 1.7. This type of sheet
hydroforming is similar to conventional deep drawing except the application of the counter pressure on the other side of punch. The hydromechanical deep drawing has no lower die, but applies hydraulic pressure when the
punch forces the blank downwards. The counter pressure is controlled by
a servo or proportional valve. The hydraulic pressure improves drawing
ratio and corner filling. Nakamura et al. conducted HMD experiments and
showed that higher limiting drawing ratios could be achieved with counter
pressure (Nakamura, 1995, 1997, 1998 and 2002) and the blank attached to
WPNL2204
Introduction and state of the art of hydroforming
7
the punch is not stretched during the forming process (Siegert, 2000; Kleiner,
2003).
High-pressure sheet hydroforming
The sheet hydroforming process was developed as a combination of deep
drawing and hydroforming. Figure 1.8 shows a high-pressure sheet hydroforming for a single blank. A blank in a blank holder is placed inside a die
and pressure is applied to fill the die cavity. The advantage of this process
is the possibility of deep drawing with controlled metal flow into the cavity.
For the double blank hydroforming seen in Fig. 1.9, hydraulic fluid is
pumped between the blanks after they have been formed by conventional
deep drawing. The cavity of the die is filled by hydraulic pressure and
contoured by the upper and lower dies. The position of the upper punch
also could be changed. In terms of accuracy, it was shown that highpressure sheet hydroforming (SHF) can achieve better shape accuracy than
PF
BHRF
BHRF
PF
BHRF P
Blankholder ram
Punch
Blankholder
(upper binder)
Blank
Draw ring
(lower binder)
Cavity
Press table
c
BHRF
Pc
1.8 High-pressure sheet hydroforming (single blank) (Siegert, 2000).
Ram
Punch
Punch
Blank
holder
Ram
Blank
holder
Hydromedia
Lower die
Lower die
Press table
Press table
1.9 High-pressure sheet hydroforming (double blank) (Siegert, 2000):
a deep drawing of the double blanks, b converting and calibrating by
hydroforming.
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Hydroforming for advanced manufacturing
conventional deep drawing (Homberg, 2000 and Kleiner, 2003). Another
benefit of SHF is that it needs only ‘one’ die or punch for the forming
process. Therefore, the time and cost of designing and manufacturing dies
is reduced. SHF is very profitable for various production types especially
low–medium volume production. There are other types of hydroforming
processes which are slightly different versions of the above, such as liquid
impact forming and flexforming.
Tube hydroforming
Tube hydroforming (THF) has been known by many other names depending on when and where it was used or investigated. Bulge forming of tubes
(BFTs) and liquid bulge forming (LBF) were two early terms, for instance.
Hydraulic (or hydrostatic) pressure forming (HPF) was another name used
for a while by some investigators. Internal high-pressure forming (IHPF)
was mostly used by German manufacturers and researchers. In some periods,
it was even called as ‘unconventional tee forming’. THF is a materialforming process whereby tubes (straight or pre-bent) are formed into
complex shapes with a die cavity using simultaneous application of internal
pressure and axial compressive forces from both or either ends. The internal
pressure is usually obtained by various means such as pumping hydraulic
and/or viscous medium or squeezing intermediate viscoelastic elements
such as elastomers and polyurethane. Process parameters for a typical tube
hydroforming operation are depicted in Fig. 1.10. Some examples of parts
manufactured by THF are illustrated in Figs. 1.11 and 1.12.
In summary, hydroforming has been proven to be a successful forming
technology replacing conventional stamping and forging processes with
promises of cost savings in terms of elimination of die sets, reduction of
assembly operations via part consolidation, tight dimensional tolerances,
Fq
Re
Rc
Fa
Pi
a
b
c
1.10 a Configuration of a typical tube hydroforming process, acting
loading elements, and geometrical features of importance: Fa axial
force, Fq counter force, Pi internal pressure, Rc corner radius, Re fillet
radius; b sample T-joint part; c exhaust pipe.
WPNL2204
Introduction and state of the art of hydroforming
a
9
b
1.11 a Example parts: hydroformed A pillar in Volvo C 70 (Shah,
2007), b rollover bar protection system in Porsche Boxter.
1.12 Hydroformed engine cradle: the initial straight tubular blank (1),
bent (2), pre-formed (3), hydroformed and pierced (4) (courtesy of
Schuler Inc.).
and complex part formability. Hydroforming of lightweight alloys (sheet,
tube and extrusions) presents challenges as well as opportunities as it promises further reductions in vehicle weight, an increase in the part complexity
and variety, and cost reductions compared with, for instance, stamping
of lightweight materials. On the other hand, since the application of tube
hydroforming technology into mass production is relatively new compared
with other metal-forming processes such as stamping and forging, the existing knowledge base, design rules, and experience for design of parts, process
and tooling are limited. Hence, the application of this technology to new
parts and areas requires extensive development and trial efforts. As a result,
this leads to high development cost, which decreases the competitiveness
of the tube hydroforming process compared with other processes. There are
still some issues that need to be addressed by the research community.
These can be summarized as follows:
WPNL2204
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Hydroforming for advanced manufacturing
(a) Sealing is an problem particularly for the sheet hydroforming.
Improved sealing technologies need to be developed for successful,
robust and consistent production.
(b) Coordination of internal pressure and blank holding force (BHF) (or
axial feeding in tube hydroforming) and their optimal applications is
still being performed based on trial- and-error efforts either by computer or on the plant floor. The efficient optimization of these leading
parameters still remains to be addressed.
(c) Lubricants, coatings and the determination of friction coefficients are
other problems that are usually dealt in a case-by-case mode. Comprehensive methods or computational tools are yet to be developed.
(d) The effect of bending and pre-forming on the hydroforming and the
final part quality is not fully understood yet. Trial-and-error efforts
have been widely used costing much lead time and man power.
1.2
Hydroforming systems, equipment,
tooling and controls
Design, control and maintenance of the tube hydroforming system is of
special importance since high hydraulic pressure levels, very large tooling
and equipment with high capital cost are involved to ensure continuous
mass production of complex shaped parts. The system needed for a typical
hydroforming consists of the following (Fig. 1.13):
• Clamping devices for closing and holding the dies: presses (hydraulic)
• Tooling: dies, inserts, etc.
• Pressure system; pumps, intensifier, valves, sensors/transducers, controls,
• Hydraulic cylinders and punches: for sealing the tube and move the
material
• Process control systems; computers, data acquisition, transducers, etc.
• Hydraulic conditioners: coolers, filters, additives, etc.
1.2.1 Presses or clamping devices
In contrast with other forming operations, in the hydroforming process,
hydraulic presses are typically used to open and close the die and to provide
enough clamping load during the forming period to prevent elastic deflections and die separation. In some cases, equipment specially designed to
provide the necessary high tonnage clamping force is used instead of regular
presses. The necessary tonnage of the press (or clamping device) is dependent on the required closing force. It is, in turn, a function of the maximum
internal pressure used during forming, part size (i.e. diameter, length and
thickness), and material. Large components with thick walls (i.e. chassis
WPNL2204
Introduction and state of the art of hydroforming
11
Deformation
zone
Tool/die
Product
Tube/workpiece
Equipment/press
Tool/workpiece
interface
a
Process control
of horizontal cylinder
Die functions
Pre-filling
(quickfill)
Axial piston pump
for oil-hydraulic
die functions
Emulsion
fine filtration
Process control
of the internal
pressure
Emulsion tank
and emulsion
coarse filtration with
oil skimmer
Pressure
filling
Clear tank
Intermediate tank
b
1.13 Hydroforming system elements: a overall system elements from
workpiece to the equipment and its environment, b schematic of a
typical hydroforming system including part, tools (die, punches,
cylinders), controls, hydraulics, pumps, and cleaning tank.
components), and intricate regions (i.e. small corner radii) need high closing
forces in the region of 7000–8000 ton (Fig. 1.14). At present, presses up to
10 000-ton capacity are in operation at several plants in the world. Existing
hydraulic presses with appropriate closing forces and bed sizes can be utilized for hydroforming process with some necessary additions and changes
in the system.
Clamping devices, other than regular hydraulic press systems, are
being designed and tested for hydroforming purposes based on auxiliary
WPNL2204
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Hydroforming for advanced manufacturing
a
b
Pm
Pm
P
c
P
d
e
1.14 a Hydroforming production line consisting of several presses:
b close-up of a press opening showing the removal of a hydroformed
part out of the tooling after a try-out, c mechanical (bayonet) locking
system to increase the clamping capability cost effectively (courtesy
of Schuler Hydroforming), d hydroforming press with double ram
(courtesy of SPS Inc.), e Sheet hydroforming press; clamping capacity
is increased via pre-stressing by winding of wires around the dies
(courtesy of Avure Tech.).
WPNL2204
Introduction and state of the art of hydroforming
13
mechanical locking for increased and cost-effective clamping forces (Fig.
1.14c) and/or additional hydraulic circuits for rapid cycle times. The purpose
of developing special clamping devices is to increase capabilities on process
control, obtain better dimensional accuracy via high clamping load, access
larger bed size, reduce cycle time, increase flexibility for different parts,
and reduce investments. In such a design, the ram with the upper die half
is actuated up and down through a small cylinder, which would provide
rapid motion and cost less. As the ram closes, the die is at its bottom position and two opposite and horizontally positioned cylinders are actuated
to lock the ram at its required location. Moreover, several other small and
short-stroke cylinders at the bottom of the press bed are moved up to
further increase the clamping load capability. Such designs would not
only be cost effective in terms of initial capital investment, but also would
provide rapid stroking, which consequently contributes to reducing the
production cost. In principle, a tube hydroforming press or machine must
have the following features:
•
•
•
•
•
•
Appropriate die closing force;
Appropriate bed size to hold the dies;
Adjustable/movable axial punches with computer controlled
positioning;
Adjustable/movable rams for counter forces with free and position
control;
Optional: automatic work-piece handling;
High pressure (2000 to 5000 bar/200–500 MPa) and fluid-pumping capability with tight control.
1.2.2 Tooling
Hydroforming tooling consists of die holders, dies, inserts, punches, sealing
systems and, sometimes, counter punches or movable inserts. Due to
the high-pressure values involved in THF process, strong tooling systems
are required to minimize die deflection and part tolerance deviations.
Hence, tool steel such as D2 is used for inserts whereas 1045 steel is
used for the dies. Inserts are usually hardened and polished to achieve
smooth surface finish to reduce friction and die wear. The design of
part positioning and parting lines is of utmost importance to achieve
the necessary reduction in closing force and guarantee the formability
of the part. For structural parts, diagonal positioning is one way of
balancing the die deflection between vertical and horizontal directions
of the part. Because of the need for confidentiality in this high demanding technology, limited information regarding tooling design is released
to the public as it is associated with other aspects of the technology.
WPNL2204
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Hydroforming for advanced manufacturing
Upper die
Guide pins
(for initial location)
Hydroform
contour
Pressure plate
Pressure plates
(adjustable)
Ejectors
Guide plates
(adjustable)
Sealing face
Wear resistant
inserts
Upper/lower
locators
Lower die
a
b
1.15 a Schematic view of a typical hydroforming die; b Nonsegmented hydroform die (courtesy of Siempelkamp Pressen
Systeme-SPS).
a
b
1.16 a Hydroforming tooling for an engine cradle (courtesy of
Siempelkamp Pressen Systeme-SPS); b hydroforming tooling for a
side rail of an SUV chassis. Details show the close-up of a hole
piercing tooling sub-component.
Hence, common guidelines known for forging and stamping technologies
are applied in combination after necessary improvements and trials. The
section view of a typical hydroforming die is depicted in Fig. 1.15a. A
non-segmented hydroforming die and its basic parts can be seen in Fig.
1.15b. Hydroforming tooling examples for an engine cradle and side rail
are presented in Fig. 1.16.
The general features and main requirements for hydroforming tooling
are as follows:
• High strength against stresses due to large internal pressure and axial
loading,
• Good surface finish to minimize friction and increase formability,
• Flexibility by interchangeable inserts,
WPNL2204
Introduction and state of the art of hydroforming
•
•
15
Good guiding systems,
Balanced design to minimize the closing force requirements.
1.2.3 Pressure system
The pressure system (pump, intensifier and control and relief valves, coolers,
etc.) should be designed and selected so as to provide the required pressure
levels for a wide range of parts to obtain flexibility in the system invested.
Figure 1.13 depicts a schematic view of a typical pressure system in a hydroforming unit. The applied pressure should have a range from 2000 bar
(30 ksi/200 MPa) up to 10 000 bar (150 ksi/1000 MPa) depending on the parts
in consideration. In many current industrial applications, the use of pressures up to 3000 bar (45 ksi/300 MPa) are sufficient. The flow rate can reach
up to 50 l min−1 in order to allow short cycle times. In order to increase the
production rate, multiple intensifiers are used to shorten the pressurizing
period and compensate for time losses when rapid pressure increases are
required by any part or process design.
1.2.4 Hydraulic cylinders and punches
The axial punches are necessary (a) to seal the end of the tube to avoid
pressure losses and (b) to feed material into expansion regions. They
should feed the material into the deformation zone in a controlled way,
and in synchronization with internal pressure, i.e. pressure versus time
and axial force versus time should be controlled and coordinated. Counter
punches are sometimes used on bulged or protrusion sections to avoid
premature fracture by providing a controlled material flow. Axial cylinders are expected to generate forces of up to 7000 kN (700 ton) while
counter cylinder limits extend up to 2000 kN (200 ton). The smaller size
also allows close control of the punch position. Various punch tip designs
for effective sealing during hydroforming have been developed.
1.3
Materials, formability, forming limits
and advantages
The overall success of hydroforming product heavily depends on the proper
choice, quality and consistency of the incoming materials whether tubular
or sheet blanks. Material properties such as composition, weld type, yield
and tensile strength, ductility and anisotropy must be determined for the
incoming blanks whether they are in sheet or tubular forms. Monitoring
and controlling of rolling, welding and annealing processes should be
conducted carefully to produce blanks with the desired properties. The
WPNL2204
16
Hydroforming for advanced manufacturing
following are the required characteristics of raw materials for quality hydroforming applications:
• High and uniform elongation,
• High strain hardening exponent,
• Low anisotropy,
• Close mechanical and surface properties of weld line to the base
material (in case of tubular materials),
• Good surface quality, free of scratches,
• Close dimensional tolerances (thickness, diameter and shape),
• Burr free ends,
• Tube edges perpendicular to the longitudinal axis.
According to the requirements above, all alloys that are used in deep
drawing, stamping or extrusion are suitable for hydroforming. Various
testing methods have been used to determine the quality of incoming blanks
(sheet or tube). These tests can be listed as follows: (a) tensile test, (b)
expansion test, (c) cone test and (d) bulge test. From such tests, yield and
ultimate tensile strength (YS, UTS), elastic modulus (E), strength coefficient
(K), strain hardening coefficient (n) and anisotropy (r) values can be
obtained. However, in general, it is observed that hydraulic bulge tests (as
illustrated in Fig. 1.17) result in better representation of material behavior
when compared with tensile tests although they are more convenient and
available to conduct.
Thus far, for various structural body components (such as chassis parts
and side rails) low carbon/mild steel grades (such as SAE/AISI 1008, 1010)
are used in tube hydroforming applications. Stainless steel 304 type grades
are used for exhaust manifold/pipe applications. Several grades of aluminum alloys are used in high-scale automotive applications due to their
rθ
j
q
rz
sθ
t
True stress (MPa)
1600
sz
1200
s = K (e0 + e)n
800
SS304
d0 = 57.15 mm
t0 = 0.6 mm
400
0
a
K = 1357 MPa
n = 0.645
e0 = 0.06
0
0.2
0.4
0.6
True strain
b
0.8
1.17 Hydraulic bulge test principles for a tubular blank and sample
flow stress curve for 304 Stainless Steel tubes (Koç et al., 2001).
WPNL2204
1
Introduction and state of the art of hydroforming
17
lightweighting premises (i.e., 5000 series Al alloys for engine cradles, rear
axle parts, roof rails, etc. Al 6000 series alloys for several body panels).
Hydroforming of advanced high strength steels (DP series) is also gaining
interest for safety components such as side impact bars.
1.4
Tribology in hydroforming – friction, wear,
lubricants, coatings and testing methods
Compared with stamping operations, tribological conditions in hydroforming are more severe due to high surface/contact stresses under high internal
pressure and distributed plastic deformation throughout the workpiece.
Contact stress levels are definitely higher in hydroforming than in stamping,
but their levels are not as high as in forging processes. As a result, surface
conditions, coatings, lubricants and die life problems should be considered
very carefully a part of the design process. Tribological conditions in hydroforming, as they are in any typical metal forming process, are influenced
mainly by the following factors: (a) surface conditions of tube and die, (b)
contact area and associated state of stress, (c) surface pressure, (d) sliding
velocity, (e) tube and die materials and their mechanical properties, (f)
contact temperature, (g) die coatings, (h) positioning of the parting line, and
(j) lubricant. More detailed information on tribology of tube hydroforming
will be given in Chapter 5.
Structural frame parts with particularly long and with varying crosssections require substantial axial feeding in order to form into die cavities
without much expense or excessive thinning. Substantial cross-sectional
changes from round-like to rectangular shapes demand minimum resistance against corner forming and material movement. Friction issues
for such cases become critical. Selection of an appropriate lubricant and
die coating is essential to overcome sliding friction, prevent sticking and
galling to reduce tool wear, axial forces and excessive thinning. In general,
three main friction regimes are identified in a typical hydroforming process
at part-to-tooling interface depending on the differences in metal flow,
sliding velocity and state of stress. These friction regimes and consequent
friction coefficients continuously vary with location and time and can be
described as follows (Fig. 1.18): (Prier and Schmoeckel, 1999; Koç and
Altan, 2001):
(a) Friction in the guided zone where the tube and die surfaces are in
contact under pressure and straight axial compression. Axial movement of material is very rapid compared with expansion (circumferential movement). The material movement rate may vary between
50 and 100 mm s−1. Medium surface pressure, high sliding velocity, high
axial pressure, limited expansion of the surface;
WPNL2204
18
Hydroforming for advanced manufacturing
Fa
3
7
ΔL
1
8
4
1 Die
2 Initial tube
3 Punch
4 Final tube
5 Counter punch
6 Fixed tube ends
7 Guided zone
8 Transition zone
9 Expansion zone
10 Urethane pads
9
5
2
Pi
to
6
10
1.18 Schematic of a basic tooling design for friction testing, and
various friction zones during a typical tube hydroforming process,
similar zones can be identified in a sheet hydroforming case.
(b)
(c)
Friction in the transition zone where curved part and die surfaces are
in contact under pressure and a tri-axial state of stress exists. The rate
of movement of material is slow compared with the guided zone.
Surface expansion or reduction, sliding velocity smaller than that of
the guide zone, but still appreciable, stresses somewhere between axial
pressure and tensile hoop stress, tensile stresses in the tube are in hoop
direction;
Friction in expansion and calibration zones where axial feeding is
negligible and a bi-axial state of stress exists. Material movement in
the circumferential direction is dominant compared with negligible
axial movement. Tensile stresses are prevalent (axial and hoop direction), sliding velocity is small, surface enlargement is large.
Until recent years, there was no reported testing method or equipment
developed to measure or evaluate friction in the tube hydroforming
process. However, the effect of friction and different lubricants on formability and extent of protrusion height was mentioned on many occasions
starting in the 1970s [Limb et al., 1973]. In order to investigate the influence of the above parameters in different zones of friction, Schmoeckel
et al. used an experimental setup where a straight tube is expanded
under internal pressure and pushed to investigate the friction conditions
in only the guide zone (Prier and Schmoeckel, 1999; Schmoeckel et al.,
1997). Simultaneously, Dohmann et al. developed another tooling, which
WPNL2204
Introduction and state of the art of hydroforming
19
would permit investigation of friction in all zones (Dohmann, 1997).
Other researchers conducted pin-on-disk or twist tests to rank the performance of various lubricants suggested for hydroforming applications
(Dalton, 1999). As a result, all parameters affecting friction conditions
should be improved for an overall success in hydroforming. For instance,
a good hydroforming lubricant should be selected based on the following
criteria:
•
•
•
•
•
•
Lubricity to reduce sliding friction between tooling and tube surface
Durability under high pressure values up to 6–15 ksi (40–100 MPa) at
the tube-to-tooling interface to prevent sticking and galling
Minimum abrasivity to reduce tool wear
Compatibility with pressurizing medium and environmental
requirements
Ease of application and removal (washable)
Cost
Lubricants are widely used in metal forming operations to (a) separate
work-piece and die surfaces, (b) reduce interface friction, (c) help material flow to achieve complete cavity filling, (d) obtain parts with required
thickness specifications (reduce thinning), (e) prolong die life by reducing wear and contact stresses. In hydroforming, boundary lubrication
governs the friction conditions. As the internal pressure increase, the
area of contact at the interface also increases and sticking friction may
become dominant. Therefore, under varying pressure conditions during
hydroforming process, various contact conditions may govern. In turn,
various friction laws may be used to model the friction conditions. For
low-pressure levels at the initial stages of hydroforming, Coulomb friction may be used. According to the Coulomb friction law, the tangential
(frictional) stress (t) is proportional to the normal stress (sn) at the
interface. The proportionality constant is called the friction coefficient
(m). If the contact pressure is close to the flow stress of the tube material, then, the Coulomb friction model no longer applies, and the shear
stress model has to be used. According to the shear stress model, the
tangential (frictional) stress (t) at the interface is proportional to the
flow stress (sv). In this case, the proportionality constant is called the
friction factor (f).
Topography of tool and tube also plays an important role in the tribological mechanism of the hydroforming process. In particular, it is necessary to
understand the effect of varying surface roughness of a part undergoing
heavy cold working with changing state of stress in hydroforming. At the
early stages of the process, there are peaks and valleys at the contact
surface. Hence, friction conditions are severe as lubricant is trapped in this
rugged surface structure, and this may not help in separating the die and
WPNL2204
20
Hydroforming for advanced manufacturing
part surfaces. As pressure increases, the part is deformed into the given
shape, and asperities begin to disappear. As a result, friction condition
becomes less hostile in terms of surface. The friction coefficient at lowpressure levels is found to be higher than the friction coefficients at high
pressures. Thus, utilization of intentionally textured tube surfaces can be an
additional approach to lubrication for some hydroform parts. Hydroforming of aluminum may bring additional challenges as surfaces of aluminum
alloys are covered with a thin and hard oxide layer. Breakage of this layer
due to heavy cold working during deformation of the part surface would
expose additional and unexpected surfaces to the contact mechanism. Since
these additional surfaces are not lubricated properly and sufficiently, they
may cause harsh contact conditions resulting in excessive thinning and early
fracture of the part.
Depending on the composition of the lubricant, they fall into the following categories: (a) dry lubricants (solid lubricants), (b) wet lubricants (solutions and emulsions as well as synthetics), (c) pastes, soaps and waxes. Each
group has their own advantages and disadvantages in terms of performance,
application, removal, compatibility, and cost.
Dry lubricants are usually found to be more effective in terms of performance for reducing friction and increasing tool life. Their application is
easy and consistent with proper instrumentation. Their compatibility with
pressurizing fluid is very good when they are dried appropriately. However,
their removal requires special washing fluids. They are found to be more
expensive than wet lubricants when drying time, application and removal
process and their original costs are added. On the other hand, wet lubricants are cost effective, easy to remove, most of the time are compatible
with pressure fluids, but do not perform well as dry lubricant do. Hence,
a compromise must be made depending on the part complexity and quality
requirements.
For a given set of die and tube materials, and surface and loading conditions, selection of an appropriate lubricant is essential to overcome sliding
friction, prevent sticking and galling, reduce tool wear, axial forces and
excessive thinning to produce a sound and acceptable hydroform part.
Selection of production lubricants also requires financial and environmental justifications. Manufacturers need to verify and test these findings to
compare with required dimensional specifications of the actual parts under
production conditions before they start mass production. Wall thickness,
flatness and radius specifications need to be verified with specified values
determined for NVH and crash requirements of a vehicle. Figure 1.19
illustrates the effect of lubricant selection on the critical thinning of a
structural rail part; lubricant #2 performs best for both initial tube thickness
cases and results in less thinning at the critical expansion regions on this
part.
WPNL2204
Thickness (mm)
Introduction and state of the art of hydroforming
to = 4 mm, Section B
to = 3 mm, Section B
4.00
3.80
3.60
3.40
3.20
3.00
2.80
2.60
2.40
2.20
2.00
1.80
1.60
Variation = 3 s
0
Thickness (mm)
21
1
2
3
Lubricants
4
5
4
5
Left End
Region B
Region C Region D
Right End
to = 4 mm, Section D
to = 3 mm, Section D
4.00
3.80
3.60
3.40
3.20
3.00
2.80
2.60
2.40
2.20
2.00
1.80
1.60
Variation = 3 s
0
1
2
3
Lubricants
1.19 Comparison of minimum thickness measurements in regions B
and D for respective lubricants (Koç, 2003).
1.5
Computer simulations for tube hydroforming
In hydroforming, analyzing thinning, thickening, strain and stress distribution on a deformed tube can predict feasibility of forming for a specific part.
The effects of different parameters can be investigated by varying important dimensions or loading conditions on common part types. Hence, generic
rules can be established for future problems. Optimized loading can be
obtained by using controlled analysis techniques.
Many hydroforming operations starting with a tubular blank require a
pre-formed tube in order to (a) fit the tube into the hydroforming die cavity
and (b) reach the desired shape at the end of the process. Thinning and
thickening of tubes particularly during the bending operation may greatly
affect the success of the hydroforming process as thinned sections may not
be able to withstand internal pressure during expansion, and consequently
burst, whereas excessive thickening may lead to wrinkles on the bent tube,
and these may require high pressure for straightening. In order to analyse
the entire tube hydroforming process, it is necessary to carry the results of
bending and crushing analysis into the hydroforming stage. Use of Finite
Element Analysis (FEA) is so far the only way of achieving this. Appropriately selected FEA software would carry the strain history gained during
pre-forming directly into the hydroforming stage just as in actual forming
of complex parts. Figure 1.20 illustrates computer modeling and analysis of
a structural rail part from bending of the tube until its hydroforming into
the final shape.
WPNL2204
22
Hydroforming for advanced manufacturing
Upper die
FEA
4.00
Experiment
3.50
Thickness (mm)
Designed bent tube
3.00
2.50
D
2.00
1.50
1.00
Section 3
C
Lower die
A
E
0.50
0.00
0
Piston
100
200
300
400
Circumferencial distance (mm)
500
1.20 Computer modeling and analysis of the hydroforming process
through all stages of bending, performing and final calibration (Koç,
2002).
Furthermore, the success of hydroforming a sound part without any
defects and with required thickness specifications is dependent on the selection of an appropriate loading path (i.e., pressure versus time and axial force
versus time and/or blank holding force versus time) for an already selected
set of material, lubrication, and part and tooling design. Proper coordination of internal pressure and axial feeding is the key issue as these process
parameters have to be applied synchronously (Koç, 2001; Asnafi et al., 2000).
In practice, trial and error procedures have been used extensively during
prototype try-outs of any new product to determine loading conditions after
material and lubrication selections are finalized. In addition to try-outs and
simple calculation methods for critical parameters, lately several researchers (Doege, 1998; Altan et al., 1999; Yang 2001 and Strano et al., 2004) proposed and implemented an adaptive FEA simulation technique to determine
the loading paths for any given set of part type, shape, material and lubrication. Such simulation techniques are still in the development stage and have
only resulted in success for simple shaped hydroforming parts. The goal is
to determine ‘good’ loading profiles, i.e. internal pressure vs. time and axial
feeding vs. time, necessary to hydroform a good part free of any defects (i.e.
wrinkling, bursting and thinning) and without any trial and error simulations and prototype efforts. This technique employs, in addition to an FEA
code such as PAM-STAMP or LS-DYNA, (a) a software module that scans
the simulated part at every time increment (dt) to detect defects such as
wrinkling and thinning based on certain criterion, (b) an algorithm that
predicts the loading path in the following time step (dt+1) depending on the
detected defect trend, if any, and previous status of loading path. At every
time increment of the FE simulation, the program reads the necessary
current data from the simulation model. If any node has a negative velocity
vector (towards the centerline of the tube), it is detected as the beginning
WPNL2204
Introduction and state of the art of hydroforming
23
of wrinkling, Fig. 1.21. Consequently, as a correctional measure, the program
increases the internal pressure while simultaneously stopping the axial
feeding of the material. If wrinkles have been removed, the internal pressure is kept constant while starting feeding material again. It is noted that
using geometric criteria, such as velocity or displacement, is a very simple
way to detect wrinkling (Altan and Koç, 1999).
FEA of a simple T-shaped hydroformed part was utilized to investigate
the effect of loading path on the attainable bulge height and thinning limitations. Figure 1.22a illustrates the two different loading paths applied. These
loading paths were obtained with the calculation of yielding and maximum
pressure values using simple analytical methods described in (Koç, 2003).
Die
t0
t0
t1
t2
t3
Tube
t4
Axial feeding
Internal pressure
t1
Pressure (Pi)
ΔPi
Axial Feed (Da)
Piy
2ΔDa
Piy
Wrinkle
t2
ΔDa
ΔDa
ΔDa
Piy
t3
Piy + ΔPi
Wrinkle
t4
Piy + ΔPi
2ΔDa
Time
1.21 Illustration of adaptive simulation technique to determine
appropriate loading path for hydroforming process (Altan and Koç,
1999).
Pressure (Case 2)
Pressure (Case 1)
Axial punch movement (Case 2)
Axial punch movement (Case 1)
120
Internal pressure (MPa)
90
80
80
70
60
60
50
40
40
30
20
30% thinning
20
H
Bulge height, H (mm)
100
Axial punch movement (mm)
100
H
35
30
25
20
15
10
5
0
Case 1
Case 2
b
10
0
0
0
2
4
6
Time (s)
8
10
12
a
1.22 Effect of loading path design on the final part properties such as
attainable bulge height and thinning.
WPNL2204
24
Hydroforming for advanced manufacturing
In case 2, internal pressure is slightly lower during the initial stages (up to
4 s) while axial feeding is kept higher compared with case 1. As seen in Fig.
1.22b, loading case 2 provided a bulge height of 32 mm with a maximum
thinning value of 30% while case 1 resulted in 25 mm with same amount of
thinning. This indicates that feeding more at the initial stages where pressure is low contributes more to the compensation of thinning at the expansion zones. Hence, generally speaking, such modes of loading path would
be favorable in hydroforming.
1.6
Developments in hydroforming and
concluding remarks
Recent innovations are aimed to improve competitiveness of hydroforming
technology by reducing initial investment cost, increasing production rate,
and material utilization, consolidating more parts into single parts, and
finding ways to eliminate drawbacks such as excessive thinning. New press
or clamping device concepts are under development to reduce the amount
of initial capital investment as well as to increase the productivity by having
rapid strokes. Even some hydroforming systems without a press or clamping
device are discussed and seem feasible only for low production rates.
In order to increase the material utilization and avoid excessive thinning,
the following innovations are being tested and used nowadays (a) tapered
(conical) tubes for long structural parts having substantial expansion degrees
between two ends, (b) tailor-welded tubes for minimizing thinning at high
expansion zones which are usually at the middle sections of a long part for
which other innovations can not be utilized practically, (c) double tubing is
used to increase the strength of the final part while minimizing the weight.
Particularly used for front rails where extra care has to be taken for excellent crash properties, (d) multiple tubing seems to be an innovative way of
producing whole assemblies at once, which is an excellent way of consolidating more parts into one. Tubes of different preformed shapes are connected to each other, and placed into a hydroforming die altogether. Upon
completion of hydroforming, all parts of an assembly are manufactured and
assembled.
Companies and institutes are looking into every chance and opportunity
to make cost effective production with lighter and stronger products. For
instance, consolidation of lubrication into tube making is considered to
be one way of increasing the production rate. The application of various
welding types, such as gas metal arc welding, laser welding, electron beam
welding, is investigated to search better material properties. Tube making
(forming) cells is in consideration instead of conventional tube rolling mills
in some justifiable cases.
WPNL2204
Introduction and state of the art of hydroforming
25
In spite of the various advantages of hydroforming, forming of lightweight materials into complex shaped parts is still a challenge because of
many unknowns in, for instance, the effect of material behavior under
hydrostatic loading, surface topography and its effect on overall formability,
the necessity for effective lubricants, the low formability of lightweight
alloys such as aluminum and magnesium at room temperature, etc.
As an alternative to cold hydroforming, in order to further extend the
forming degrees of lightweight materials, a hybrid warm hydroforming
process is being investigated by many researchers. Similar ideas have been
discussed within the forming community since as early as the beginning of
1990s. However, written literature suggesting or investigating this possibility
goes back to only early 2000s (Nakamura, 1997; Vollertsen, 1999; Groche
et al., 2002; Lee, et al., 2002). Warm hydroforming simply can be applied
in two manners: (1) bulging of heated blank(s) into a die cavity via fluid
pressure, (2) deep-drawing of a blank against a hydraulic force (hydromechanical forming) as illustrated in Fig. 1.23.
HEATforming is a recently developed tube hydroforming technique
introduced by Schuler. In this process, a heated tube is placed in a heated
die; then the die is closed; and the tube sealed at the ends by sealing cylinders. The tube is subsequently expanded against the die cavity wall by
internal pressure – here provided by a gaseous medium. The process may
also be supported by continued axial feeding of the tube, similar to conventional hydroforming. The tube material and the die can be adjusted to
various temperature zones for control of the material flow. The flow of the
material in the die is further aided by specially developed lubricants. Figure
1.24 depicts the steps for HEATforming.
The very important premise of the warm hydroforming process is to
increase the formability of lightweight materials beyond limits that are
achievable in conventional cold forming processes because of (a) reduction
in the friction between workpiece and tooling elements, and (b) decrease
Pi
Pi
a
b
Pi
c
1.23 a Basic elements of warm hydroforming process, b hydroforming
of double blanks, c warm hydroforming against fluid pressure, i.e.
hydromechanical forming.
WPNL2204
26
Hydroforming for advanced manufacturing
Thermocouple
Axial force
Thermocouple
530°C
560°C
530°C
560°C
Axial force
Pressure gas
Pressure gas
a
c
Thermocouple
Thermocouple
Axial force
Axial force
Pressure gas
Pressure gas
b
d
1.24 Procedure for HEATforming technique a the preheated tubular
blank is placed in the preheated tool and both ends are then sealed;
b the tubular blank is pressurized and material is fed in; c the tube is
formed by inner pressure and simultaneous feeding of material; d the
tube is calibrated under high pressure (courtesy of Schuler Inc.).
in the flow stress of material at elevated temperatures. Thus, material flows
into the expansion areas or intricate regions takes place with very low
forming loads while some sections of the blank in contact with fluid medium
cools rapidly to increase the forming limit via strain hardening. Consequently, reduced forming loads will result in small forming equipment
requirements with low capital equipment investment savings. An increased
formability would result in consolidation of multiple parts leading to reductions in joining/assembly operations contributing to both cost savings and
increased integrity. Figure 1.25 summarizes the advantages of warm hydroforming in terms of the achievable limiting drawing ratio (LDR) in a simple
hydro-drawing case (Groche, 2002).
On the other hand, additional scientific challenges are introduced with
warm hydroforming process. These challenges include: (a) appropriate
design and control of optimal temperature distribution on tooling elements
(i.e., die, blank holder/punches) and blank material, (b) prediction and
compensation of consequent residual stresses and distortions; (c) prediction
of optimal and synchronous loading paths (e.g. pressure vs. time, temperature vs. time, etc.), (d) understanding and modeling of the complex surface
interactions, friction and effective of use of lubrication at elevated temperatures, (e) effect of warm temperature conditions on the material properties,
formability, and failure modes. In addition, practical issues such as (a)
handling and containment of warm blank material, pressurizing fluid
WPNL2204
Blank holder
Fluid bead
Cooled
punch
T
T
q·
Cooled
drawing
die radius
q·
P
T
Heated
die area
P
Counter pressure (MPa)
Introduction and state of the art of hydroforming
25
Phase I
Phase II
20
10
hwall
5
rzk
0
50
100
150
Punch travel (mm)
Counter pressure pot
Limiting drawing ratio
a
Phase III
rst
15
0
27
b
3.0
2.5
2.0
0
50
100
150
200
250
300
Flange temperature (°C)
DD:= Deep-drawing HM:= Hydromechanical deep-drawing
DD: AIMg4.5Mn
punch diameter: 100mm
DD: AIMg0.4Si1.2
die radius:
7 mm
HM: AIMg4.5Mn
punch radius:
7 mm
HM: AIMg0.4Si1.2
punch velocity:
5 mm/s
c
1.25 Tooling and LDR of warm hydromechanical drawing (Groche,
2002), a Tooling, b Pressure control strategy, c Comparison of LDR
values in warm and cold cases.
medium, lubricants, and (b) cleaning and post-processing of formed parts
become development challenges.
Other developments in hydroforming technology include its hybrids in
combination with other processes such as electromagnetic force forming
(EMF) or electrohydraulic forming.
In summary, hydroforming has been proven to be a successful forming
technology replacing conventional stamping and forging processes with
premises of cost savings in terms of elimination of die sets, reduction of
assembly operations via part consolidation, tight dimensional tolerances,
and complex part formability. Hydroforming of lightweight alloys (sheet,
tube and extrusions) presents challenges as well as opportunities as it promises further reductions in vehicle weight, increase in the part complexity
and variety, and cost reductions compared with, for instance, stamping of
lightweight materials. For a robust process, all elements of the hydroforming
operation should be controlled and optimized. Such elements would be
loading path, material, lubrication, preforming operations, surface conditions of the tooling inserts and equipment performance (press, axial cylinders, pressure intensifiers, filtering, cooling units, etc.)
WPNL2204
28
Hydroforming for advanced manufacturing
1.7
References
altan, t.; koç, m.; aue-u-lan, y. and tibari, k. (1999), ‘Formability and design issues
in tube hydroforming’, Proc. Int. Conference on Hydroforming, Stuttgart, Germany,
October 11–12, 1999.
asnafi, n.; nilsson, t. and lassl, g. (2000), ‘Automotive tube bending and tubular
hydroforming with extruded aluminum profiles’, SAE Paper 2000-01-2770, Int.
Body Engineering Conf., Detroit, MI, Oct. 3–5, 2000.
dalton, g. (1999), ‘The role of lubricants in hydroforming’, Automotive Tube Conference, April 26–27, 1999, Detroit, MI, USA.
doege, e.; kösters, r. and ropers, c. (1998), ‘Determination of optimized control
parameters for internal high pressure forming processes with the FEM’, Proc. of
the 6th International Conference on Sheet Metal, SheMet’98, (Eds.) by Kals, Geiger,
et al., University of Twente, Twente, Netherlands, 6–8 April 1998, pp. 119–128.
doege, e.; kurz, g.; walter, g. and meyer, t. (2002), ‘Umformen von Magnesiumfeinblechen mit temperierten Werkzeugen’, EFB-Forschungsbericht, 195.
dohmann, f. and bieling, p. (1991), ‘Theoretical basis and applications of high pressure forming’, Bleche Rohre Profile, 38/5, 379–385.
dohmann, f. (1997), ‘Tribology in internal high pressure forming’, Blech Rohre
Profile, 36–39 (in German).
groche, p.; huber, r. and schmoeckel, d. (2002), ‘Hydromechanical deep drawing of
aluminum alloys at elevated temperatures’, Annals of the CIRP, 51, 2002, (1),
215–218.
homberg, w. (2000), Untersuchungen zur prozessfuhrung und zum fertigungssystem
bei der hochdruck-blech-umformung, PhD thesis, Dortmund University,
Germany.
kleiner, m. (2003), ‘Manufacturing of lightweight components by metal forming’,
CIRP Annals – Manufacturing Technology, 52, 2, 2003, 521–542.
koç, m. aue-u-lan, y. and altan, t. (2001) ‘On the characteristics of tubular materials
for hydroforming – experimentation and analysis’, International Journal of
Machine Tools and Manufacture, 41, 761–772.
koç, m. (2001), ‘Use of FEA in design of part, process and tooling for tube hydroforming technology’, SAE Paper 2001-01-3090 International Body Engineering
Conference, October 16–18, 2001, Detroit, USA.
koç, m. and altan, t. (2001), ‘Overall review of the tube hydroforming technology’,
Journal of Material Processing and Technology, 108, 3, 384–393.
koç, m. (2002), ‘Computer simulations for the tube hydroforming process’, Int.
Conference on Responsive Manufacturing (ICRM 2002), June 26–29, 2002,
Gaziantep, Turkey.
koç, m. (2003a), ‘Tribological issues in the tube hydroforming process – selection of
a lubricant for robust process conditions for an automotive structural frame part’,
ASME Journal of Manufacturing Science and Engineering, 125, 3, 484–492, August
2003.
koç, m. (2003b), ‘Investigation of the effect of loading path and variation in material
properties on robustness of the tube hydroforming process’, Journal of Materials
Processing and Technology, 133, 276–281.
lee, s.; chen, y. h. and wang, j. y. (2002), ‘Isothermal sheet formability of Mg alloy
AZ31 and AZ61’, Journal of Materials Processing and Technology, 124, 19–24.
WPNL2204
Introduction and state of the art of hydroforming
29
limb, m. e.; chakrabarty, j.; garber, s. and mellor, p. b. (1973), ‘The forming of
axisymmetric and asymmetric components from tube’, Proceedings of the 14th Int’l
MTDR Conference, pp. 799–805.
mildenberger, u. and khare, a. (2000), ‘Planning for an environment-friendly car’,
Technovation, 20, 205–214.
grey, j. e.; devereaux, a. p. and parker, w. n. (1939), ‘Apparatus for making wrought
metal T’s’. US Patent 2203868, June 1939.
morphy, g. (1997), ‘Tubular hydroforming: ability and flexibility of pressure sequencing’, Tube and Pipe Association Conference on Hydroforming, November 1997,
Chicago, USA, pp. 199–213.
murray, m. (1996), ‘Advancements using sequenced forming pressures’, Innovations
in Hydroforming Technology, TPA Int’l, September 1996, Nashville, TN, USA.
nakamura, k. (1995), ‘Warm deep drawability with hydraulic counter pressure of
1050 Al sheets’, Journal of Japan Institute of Light Metals, 47, 6, 323–328.
nakamura, k. (1997), ‘Warm deep drawability with hydraulic counter pressure of
1050 Al sheets’, Journal of Japan Institute of Light Metals, 47, 6, 323–328.
nakamura, k. (1998), ‘Effect of hydraulic counter pressure on truncated conical shell
deep drawing of aluminum alloy 5182-O sheet’, Journal of Japan Institute of Light
Metals, 48, 11, 576–580.
nakamura, k. (2002), ‘Control of the deterioration of deep drawing limit by hydraulic counter pressure for 1050-H24 aluminum sheets’, Journal of Japan Institute of
Light Metals, 52, 5, 221–225.
prier, m. and schmoeckel, d. (1999), ‘Tribology of internal high pressure forming’,
PROCEEDINGS of International Conference on Hydroforming, Stuttgart,
Germany, October 12–13.
schmoeckel, d.; hielscher, c.; huber, r. and prier, m. (1997), ‘Internal high pressure
forming at PtU’, PtU der Technischen Hochschule Darmstadt, Germany (in
German).
schmoeckel, d.; hielscher, c.; huber, r. and geiger, m. (1999), ‘Metal forming of
tubes and sheets with liquid and other flexible media’, Ann. CIRP, 48, 2,
497–513.
shah, s. (2007), ‘Auto/steel partnership lightweight front end structure – hydroform
solution and cost analysis’, Great Designs in Steel Seminar, March 7, 2007, Livonia,
Michigan, USA.
siegert, k. (2000), ‘Recent developments in hydroforming technology’, Journal of
Materials Processing Technology, 98, 251–258.
strano, m.; jirathearanat, s.; shr, s. and altan, t. (2004), ‘Virtual process development in tube hydroforming’, Journal of Materials Processing Technology, 146(1),
14 February 2004, 130–136.
vollertsen, f. (1999), ‘Warm forming with liquid pressure’, Personal communication
and presentation.
yang, j. b.; jeon, b. h. and oh, s. i. (2001), ‘Design sensitivity analysis and optimization
of hydroforming process’, Journal of Material Processing and Technology, 113,
666–672.
WPNL2204
WPNL2204
Part I
Principles of hydroforming
WPNL2204
WPNL2204
2
Hydroforming systems, equipment,
controls and tooling
DAVID GEARING and DENNIS MEVISSEN,
Interlaken Technology Corporation, USA
2.1
Introduction
In this chapter, the main components utilized in various high-pressure
hydroforming systems are described. Due to the sustained, high clamping
load required to hold the tool cavities closed during the forming process,
hydraulic presses are typically employed. The high pressures necessary for
the hydroforming process are generated using a pressure intensification
system. Special plumbing and control valves are required to distribute the
high-pressure fluid from the intensification system to the tooling. A control
system is utilized to sequence the press closure, the part sealing and feeding,
the pressure intensification, and any other process requirements. The tooling
provides the shape of the finished part and must be designed to allow part
removal after forming.
2.2
Presses
A hydraulic press used for hydroforming serves two functions. Their primary
function is to hold the die halves together against the separating force
generated by the internal pressure used to form the part. Their secondary
function is to separate the die halves to facilitate changing the parts. A
major advantage of utilizing a hydraulic press in this application is the
ability to apply a variable range of clamping force anywhere in the stroke
of the press. Depending on the process and the tooling utilized, there are a
wide variety of frame styles available to fit the application.
2.2.1 Typical hydraulic press styles used for hydroforming
Four-column and straight-sided comprise the two most common styles of
presses used for hydroforming. These two styles provide the most efficient
means for reacting the clamp load while still allowing access to all four sides
of the tooling area. For all practical purposes, the operation of these two
33
WPNL2204
34
Hydroforming for advanced manufacturing
presses is identical. They are typically down-acting and single-action. A
down-acting press would have the clamp actuator(s) located in the upper
fixed plate pressing downward on the tooling die and reacted against the
lower fixed plate. A single-acting press only has one axis of motion which
is utilized to open and close the press and to apply the required clamping
load (see Fig. 2.1). A double-acting press would have another actuator,
acting in the same direction as the first but operating as a separate axis,
mounted to the same fixed plate (see Fig. 2.2). Where a single-acting press
is typically used for tube hydroforming, the double-acting presses are
usually used for sheet hydroforming. With a double-acting press, the primary
axis is utilized to apply a clamping load to the perimeter of the part. The
secondary axis is used if the part requires a punching operation in conjunction with the hydroforming process.
Four-column and straight-sided presses are both composed of a fixed
lower plate, sometimes referred to as a bed plate or bolster, a moving upper
clamp plate, an upper crosshead, an actuator, a hydraulic power supply, and
an electrical control system (see Fig. 2.3). What differentiates the two press
styles is that the four-column presses connect the fixed lower plate and the
upper crosshead with columns, sometimes referred to as tie-rods, which
transmit the clamping load and also to provide guidance to the upper clamp
plate. A straight-sided press connects the fixed lower plate and the upper
crosshead with heavy plates attached to the sides. The guidance for the
Clamp actuator
Upper crosshead
Moving upper
clamp plate
Fixed lower plate
2.1 Down-acting, single-action press.
WPNL2204
Hydroforming systems: equipment, controls and tooling
35
Punch actuator
Clamp actuator
Upper crosshead
Moving upper
clamp plate
Fixed lower plate
2.2 Down-acting, double-action press.
2.3 Four-column, down-acting, single-action press (source: Interlaken
Technology).
WPNL2204
36
Hydroforming for advanced manufacturing
upper clamp plate is provided through bearings attached to the sides of the
press. Clamping loads applied to the top of the tool are reacted through the
actuator, then through the upper crosshead, then down the columns or
straight sides, through the lower fixed plate, and finally to the bottom of the
tool. All of the components located in the load train, including the die cavities, must be capable of supporting the full clamp tonnage to prevent any
damage from an accidental overloading. In order to avoid undesirable
deflections in the tooling and related hardware, the press components must
be designed to keep deflections from loading to a minimum. The stiffness
of a press component, such as the lower fixed plate, is a function of its width
and its section depth. As the column or side spacing width is increased, then
the section depth must be increased to maintain the same stiffness at a given
clamp tonnage. When the section depth of the lower fixed plate becomes
too great, the press can be recessed into a pit in the floor until the working
area is at a comfortable height for the convenience of the operator.
2.2.2 Hydraulic systems
The hydraulic power supply (HPS) provides hydraulic pressure for the
clamp actuator, the pressure intensification system, and the feed actuators.
A separate hydraulic pump is typically employed for pressurizing the
clamping actuator(s). This is done to assure that the clamp is capable of
maintaining the full clamp tonnage regardless of the oil pressure and flow
requirements of the intensifier and feed actuators. Hydraulic manifolds are
utilized to distribute oil to the control valves and also to the system safety
valves. The hydraulic pump and reservoir for the press clamping system is
typically located on top of the press to reduce the length of the hydraulic
lines and to reduce the amount of floor space required. Depending on the
application, the hydraulic pump and reservoir for the rest of the system
can be located either on the top of the press or on the floor. The typical
system operating pressure is 3000 psi (206 bar). However, newer systems
are beginning to operate at pressures up to 5000 psi (345 bar) to reduce
the size of the actuators, the hoses or piping, and the pump. Filtration and
cooling are required to extend the life of the hydraulic oil, the pumps, the
actuators, and the control components. Cooling is required to maintain a
safe oil temperature and it is accomplished using either a water-over-oil
or an air-over-oil type of heat exchanger. Filtration is required to prevent
contamination in the oil from damaging components and affecting system
performance.
Several factors need to be considered when determining the size of
the HPS which include the clamp actuator(s) volume, the rate at which the
clamp actuator(s) is required to travel, and the time allowed to raise the
pressure of the oil to reach the required clamping tonnage. If only one
WPNL2204
Hydroforming systems: equipment, controls and tooling
37
hydraulic pump is used to operate the entire system, then feed actuator and
pressure intensifier actuator volumes and travel rate requirements need to
be considered as well. The bore size of the clamping actuator is directly
proportional to the rated clamping load and inversely proportional to the
system operating pressure. As the clamping load increases, the required
bore diameter increases and, if for a given load the operating pressure
increases, then the required bore diameter decreases. Since the oil volume
in the actuator is a function of the bore diameter and the stroke, increasing
the operating pressure will reduce the amount of oil required from the
pump to operate the system. This will help to reduce the size of the pump
required. The required pump flow rate is also directly proportional to the
speed at which the press opens and closes. The faster the actuators move,
the higher the oil flow rate requirement. All of these concerns also affect
the sizing of the hydraulic control valves, safety valves, and associated
plumbing.
Various positioning methods are available when flow requirements
become impractical for traditional arrangements. One common method is
to employ the use of a smaller set of actuators, referred to as kicker cylinders, to open and close the press. Since the only force required for them to
generate is to offset the weight of the upper clamp plate, the upper tools,
and the large clamp actuator rods, their bore requirement is much smaller
than that necessary to produce the required clamp tonnage. Because of this,
the oil flow rate required for clamp transit is greatly reduced. Loads required
for any pre-forming or part stripping should also be considered. To allow
the large clamp actuator(s) to move rapidly up and down, a simple, large
flow control valve is utilized to pass oil between the reservoir and the actuator. When the press is closed on the die and the clamping tonnage is
required, the large flow control valve shuts and the necessary pressure is
generated in the actuator. Depending on the application, gravity can also
be taken advantage of to reduce or eliminate the kicker cylinders. Another
method is through the use of mechanisms for positioning and locking the
clamp plate. This approach takes advantage of a short stroke, high tonnage,
clamping actuator(s) to provide the working tonnage while a long stroke,
low tonnage, transit actuator(s) opens and closes the press. The main disadvantage of a mechanical lock style press is that the dimension between
the lower fixed plate and the upper clamp plate, at full tonnage, is typically
fixed or minimally adjustable, thereby limiting the tooling variations to a
particular working height.
2.2.3 Press sizing
Part geometry and tooling requirements will determine the minimum size
of the fixed lower plate in the press. The ease of installation and removal
WPNL2204
38
Hydroforming for advanced manufacturing
of the largest tooling assembly will establish the minimum column spacing
requirement. With the clamp in the maximum closed position and no
tooling installed, the shut height of the press is defined as the distance
between the top surface of the lower fixed plate and the bottom surface
of the upper clamp plate. Ideally, the shut height would be less than the
height of the shortest tooling assembly, but if this is not possible then
spacers can be used if necessary. The working stroke of the press is defined
as the travel of the upper clamp plate from fully closed to fully opened
and should provide enough tooling separation to accommodate the
installation of the part blank and the removal of the finished part. A good
practice is to include an extra stroke to permit in-press tooling inspection
and maintenance.
The total press tonnage is determined by the part geometry at the
maximum hydroforming pressure. A simple calculation can be used to find
the minimum tonnage required to clamp the tooling together (equation
2.1):
F=
pA
2000
[2.1]
where F = the minimum clamping force required (ton), p = the maximum
forming pressure (psi) and A = the final plan area of the finished part, as
defined at the tool parting line (in2).
To allow for geometrical variations and pressure fluctuations, it is a
common practice to factor in a 30 to 50% safety margin. If there are any
additional process functions acting in the direction of the clamp axis such
as hydro-piercing or cam forming, their loads should also be accounted
for.
Small details in the finished part geometry, the thickness, and the type of
material all determine the maximum forming pressure. Typically, an initial
pressure is applied to stretch the material to the die wall followed up with
a higher pressure to yield the part to the die wall, which is referred to as
calibrating the part. Determining the pressure required for calibration is a
somewhat subjective process and is usually a combination of experience,
finite element analysis (FEA), and experimentation. Real factors that can
also affect the calibration pressure include, but are not limited to, material
work hardening, material feeding, die friction, irregular part geometry, and
small corner radii. Since the part will be supported by the tool over much
of its area, other than in the corners with small radii or other similar features, the required calibration pressure is often much higher than the burst
pressure of the tube. A simple formula can provide an accurate estimate of
the pressure required to burst a tube (equation 2.2):
pb =
(UTS)t
r
[2.2]
WPNL2204
Hydroforming systems: equipment, controls and tooling
39
where pb = the burst pressure of the tube (psi), UTS = the ultimate tensile
strength of the tube material (psi), t = the wall thickness of the tube (in)
and r = the radius of the inside diameter of the tube (in).
Equation 2.2 can also be used to calculate the minimum pressure required
to begin forming the tube by substituting the yield stress of the material for
the ultimate tensile stress (equation 2.3):
pmin =
(YS)t
r
[2.3]
where pmin = minimum pressure to begin forming (psi), YS = the yield
strength of the tube material (psi), t = the wall thickness of the tube (in)
and r = the radius of the inside diameter of the tube (in).
Equation 2.3 can also be used to calculate the pressure required to form
the smaller corner radii in the part by using the theoretical inside radius of
the finished part. As the corner radius decreases, the pressure necessary to
form the material into that detail increases. This theoretical maximum pressure should be utilized in equation 2.1 to determine the clamp tonnage
require to hold the die shut.
2.3
Pressure intensification systems
High-pressure hydroforming systems generally operate within the pressure
range of 15 000 psi (1034 bar) to 60 000 psi (4137 bar). Forming pressures
can exceed 60 000 psi, but are generally avoided unless necessary because
these pressures greatly increase the complexity and costs, while decreasing
the service life, of the high-pressure intensifier, the control valves, the distribution lines, and the fittings.
2.3.1 Pressure intensifiers
Generating pressures of this magnitude, while maintaining the ability to
accurately sequence the pressure profile with feeding and other processes,
requires the use of a pressure intensifier. In the lower pressure ranges,
providing the process is not too sophisticated, a high-pressure pump
can be substituted to generate the forming pressure. For more accurate
pressure control, a standard hydraulic actuator is used to intensify the
water pressure in order to utilize conventional servo hydraulic control
valves which provide precise control at normal hydraulic operating
pressures using servo control technology. The same theories that make the
hydraulic actuator operate, allow it to generate the higher pressures through
intensification (see Figure 2.4). Figure 2.5 shows a typical high-pressure
hydroforming water control schematic. These same principles can be used
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Hydroforming for advanced manufacturing
P1
A1
A2
P2
P1A1 = P2A2
2.4 Pressure intensifier theory of operation.
Shut-off
valve
Motor
Check
valve
Tank
Tool
Docking rod
Fill pump
Heat
exchanger
Fill
valve
Pressure intensifier
Filter
Relief
valve
Docking rod
Table drain
Optional
vent
valve
Pressure
sensor
Motor
Sump tank
Sump pump
2.5 Typical fluid hydroforming schematic.
for any fluid required for the process. Figure 2.6 shows a typical gas hydroforming control schematic.
By servo control of the hydraulic actuator, the intensifier output pressure
can be ramped up or down in conjunction with the end feeding or other
process variables. The system is able to control the forming pressure based
on the feedback from a high-pressure sensor connected to the intensifier
output. A linear position transducer is used for controlling the position of
the intensifier actuator, which allows for precise control of the volume of
forming fluid injected into the part. This feature also allows the intensifier
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Hydroforming systems: equipment, controls and tooling
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Charge
Shut-off valve Servovalve
valve
Tool
Tank
Docking rod
Docking rod
Optional
vent valve
Atmosphere
Pressure intensifier
Fill
valve
Hot zone
Relief
valve
Atmosphere
Atmosphere
Pressure
sensor
Pressure
sensor
2.6 Typical gas hydroforming schematic.
to draw in, or reload itself with, more forming fluid, at a controlled rate,
between pressure sequences.
Two primary factors are utilized when calculating the required intensifier volume. One is the total change in the part volume between the initial
blank and the finished part. This should be based on the part with the
greatest volume change from start to finish. The other depends on the
compressibility of hydroforming fluid. In most applications, liquids are
considered to be incompressible, but at the pressures that are used for
hydroforming there is a significant amount of compression in the fluid.
Air trapped in the part can have a large affect on the fluid compressibility,
which would require a larger volume intensifier. The change in fluid volume,
due to compressibility, is directly proportional to the amount of fluid used
in the part, the lines, and the intensifier. As the part volume increases, so
must the intensifier volume capacity. Because of this, the part with the
greatest change in volume, for its required change in pressure, should also
be used. These two factors will provide an accurate estimate of the total
volume necessary to form the part. Since some leakage typically occurs
during the initial sealing stage, and sometimes during the forming stage,
it is recommended to size the intensifier volume to have at least 50%
more forming fluid available as a margin of safety. This may not be feasible, or cost effective, for parts requiring large volumes. As this requires
more insight into the complete system layout, hydroforming equipment
suppliers are able to more accurately size the pressure system for the
specific application.
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Hydroforming for advanced manufacturing
2.3.2 High-pressure control valves and plumbing
There are only a handful of qualified suppliers of equipment that will
operate at pressures in excess of 5000 psi (345 bar), thereby making the
control valves and related plumbing become more difficult to procure. Due
to the high operating pressures, the port sizes in these components are
generally too small for the higher flow requirements of a production application. Because of this, hydroforming equipment suppliers are forced to
build their own special control valves and related plumbing whenever standard components are not available.
The high-pressure lines are typically constructed of stainless steel and are
available in several operating pressure ranges up to 150 000 psi (10 342 bar).
Special fittings and connection blocks, which are matched to the operating
pressure range, are required to connect the high pressure lines to the various
components used in the system. The standard practice is to utilize rigid line
throughout except where moving connections require the flexibility of a
high pressure hose. High pressure hose is available in several operating
pressure ranges up to 60 000 psi (4137 bar). At pressures exceeding 60 000 psi
(4137 bar), the fatigue life of the high pressure lines, hoses, and fittings is
significantly reduced.
High-pressure fluid control valves, which allow process fluid control at
these extreme pressures, are typically produced by hydroforming equipment manufactures. Because of the forces involved and its availability, the
valves are generally hydraulically driven and sequenced by the control
system. Simple check valves can not be used if the process requires that the
pressure be dropped during forming process. Punching into the part during
sheet hydroforming or end feeding during tube hydroforming often requires
that the pressure be relieved which requires high-pressure rated relief
valves or back flow into the intensifier.
2.3.3 Hydroforming fluid handling
The hydroforming fluid handling system usually contains multiple circuits,
depending on the size of the part to be formed and the complexity of the
system. On a larger system, a low-pressure, high-flow pump would be used
to quickly fill the part and also to circulate the fluid through the filtration
and cooling components. An additional high-pressure, low-volume pump
would be utilized to fill the intensifier and also to increase the initial pressure within the system. This minimizes the effect of air trapped in the part,
which helps to reduce the required size of the intensifier volume.
Water mixed with either an oil or synthetic-based emulsion additive is
typically used for forming. The additives provide corrosion inhibitors,
anti-foam/bacterial agents, and also a source of lubrication for the moving
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43
components in the high-pressure portion of the system. An emulsion mixture
of 3 to 10% is typically desired.
After the forming process is complete, the fluid is drained from the part
to be reclaimed. If oil-based lubricants or other non-compatible fluids were
introduced to the emulsion, either through the forming process or during
reclamation, they will need to be removed through either skimming or
filtration.
2.4
Controls
The press control system provides complete supervision and control for all
functions of the press. Three general classifications of control complexity
cover the range from a basic open-loop style, to a programmable logic
controller (PLC) style, or all the way to an advanced closed-loop design.
Because of the need for complex synchronization of multiple control channels utilized in hydroforming systems, it is usually recommended to use a
closed-loop control system to help achieve consistent results.
An example of open-loop control on the clamp actuator in a press
would be to use a simple directional valve to open and close the clamp plate
(see Fig. 2.7). When the clamp is required to close and build tonnage, the
directional valve is shifted into the closed position thereby causing the
actuator to move in that direction. Because there is no source of feedback,
the actuator would continue closing until the tool was clamped. At that
Command
Controller
Hydraulic oil
Directional
valve
Clamp actuator
2.7 Open-loop clamp control schematic.
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Hydroforming for advanced manufacturing
point, the tonnage would be generated to a predetermined level. This level
could be indicated by the use of a sensor or gauge, but not used for either
feedback or control. Proportional control valves, either mechanically or
electronically controlled, are required to determine the rate of travel while
tonnage adjustments are accomplished using either mechanical or electronically controlled relief valves. When using this method of control, the
actual rates and tonnage are only approximate because there are no adjustments made to these valves based on actual readings.
Closed-loop control systems offer tremendous amounts of flexibility, but
they also require more components. An example of closed-loop control on
the clamp actuator in a press would be to use a servo valve, sensors for
indication and feedback of the clamp position and load, and a control
system capable of reading the sensor inputs and providing the required
driving commands back to the servo valve (see Fig. 2.8). Because the servo
valve offers the ability to control either the flow rate or the pressure, the
press can be operated in either position or load control, as required. Position feedback is provided by a digital position transducer connecting the
moving clamp plate to the fixed portion of the press, which allows the controller to maintain the absolute position of the clamp actuator at all times.
The applied tonnage is converted from a sensor reading of the applied
hydraulic pressure within the clamp actuator. In position control, the closed-
Command
– –
+
Controller
Pressure sensor
Hydraulic oil
Servovalve
Clamp actuator
Position sensor
2.8 Closed-loop clamp control schematic.
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Hydroforming systems: equipment, controls and tooling
45
loop control system positions the clamp plate at a desired rate by sending
an electronic signal to the servo valve to open the valve in a preferred
direction. The controller monitors the position sensor for the actual position
and rate while making the necessary adjustments to the control valve
several thousand times per second. In load control, the same adjustments
are made except that the controller would monitor the load sensor instead
of the position sensor before determining the required valve adjustments.
The closed-loop control technique is utilized for the hydraulic actuator(s)
used in the pressure intensification system and the feed actuators in the
tooling of a typical tube hydroforming system. Tube hydroforming systems
require precise synchronization of the feed actuators to the pressure profile
to take advantage of the material flow properties. As the part is being
formed, it typically undergoes a volume change, thinning of the wall section,
work hardening of the material, and friction along the cavity surface of the
tool. These are a few of the variables that commonly occur, but they rarely
exhibit any type of linear behavior during the hydroforming process. This
suggests that the pressure and feed profiles should also be non-linear to
enhance the process. To increase the margin for error, which allows for
additional process and material variables, the program should be capable
of controlling the profiles in a flexible manner.
All of the functions monitored by the controller are displayed on the
operator interface, which is also where the system is programmed. Having
the ability to capture data within the control system allows the process
variables, and also the actual process runs, to be stored and later interrogated. Statistical process control (SPC) can either be performed actively
or offline. As a visual means of monitoring the actuator positions and loads,
the forming pressures, and any other desired functions that the controller
has access to, there is typically a digital oscilloscope (scope) displayed on
the operator interface screen. The scope provides a real time graphical
representation of the process, which is a valuable tool for validating the
process as well as troubleshooting. For example, the ability to monitor the
intensifier stroke during the pressure profile provides information about
the volume requirements during the forming process. Because the part is
well hidden during the forming process, this visual representation provides
precious insight into what is actually happening within the system.
2.5
Tooling
2.5.1 Sheet hydroforming tool styles
There are several basic approaches to sheet hydroforming tooling. Figure
2.9 shows a single-acting die with the lower half of the tool, located beneath
the blank, acting as the pressure pot and the upper half of the tool taking
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Hydroforming for advanced manufacturing
Die
Part
Die
Fluid
2.9 Sheet hydroforming die without bladder. Final part shape in upper
die cavity.
Die
Bladder
Fluid
Die
Part
2.10 Sheet hydroforming die with bladder in upper die. Final part
shape in lower cavity.
the form of the finished part. Adding draft to the tooling cavity is essential
to facilitate removal of the finished part from the tool. The pressure pot is
sealed when the part is clamped together by the two tooling halves. Parts
created using the single-acting method are constrained by the amount of
stretch that the material can provide. Depending on the application, the
pressure pot can be moved to the top half of the tooling with the addition
of an elastomeric bladder (see Fig. 2.10). Bladders also allow the blank to
be placed inside of the tool perimeter, while remaining unclamped during
the process, which allows the material to flow as required to form the part.
Since the bladder material is subject to wear, an intermediate slip-sheet of
another flexible material is usually required to extend its life.
As the demand for forming more complex shapes increases, the use of
traditional binder control on the perimeter of the tools is necessary control
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Hydroforming systems: equipment, controls and tooling
Fluid
Die
Part
Punch
47
Die
2.11 Sheet hydroforming die with punch axis in upper die cavity.
Punch provides part shape.
the material flow into the part with greater precision. Even with this
approach, the requirement to maintain a seal for the hydroforming fluid
still exists (see Fig. 2.10). It would appear that the bladder approach would
be better utilized in this application, but the interface between the elastomeric bladder and the hard tooling can present many design challenges.
Adding a punch axis to the tool allows the sheet to be drawn in a more
conventional method, while using the hydroforming fluid to aid in the
forming process and finish out the part (see Fig. 2.11). However, this method
requires an additional axis of control for the punch actuator and a more
complicated tool design.
2.5.2 Tube hydroforming tooling styles
Tube hydroforming tooling typically consists of an upper die half, a lower
die half, docking rods for the ends, and any other cams or punches required
in the forming process (see Fig. 2.12). The docking rods, which are removable tips that are designed to facilitate various tube shapes, material thicknesses, and sealing configurations, are forced into the ends of the starting
tube by hydraulic feed actuators with enough force to seal the hydroforming fluid inside and also to feed additional material into the tool when
necessary. Parts can be filled with either a high-pressure, low-flow pump
through the porting in one of the docking rods or with a low-pressure,
high-flow pump through a larger port in the tooling that is dedicated for
filling. To vent the air out of the part before forming, the opposite docking
rod can either be ported or just brought very close to the part during the
fill cycle.
The tooling split line is determined by a section cut of the finished part
that allows removal of the part from both of the tooling halves. Using the
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Hydroforming for advanced manufacturing
Addendum
Desired part
Part cut-off
Tool
Addendum
Part cut-off
2.12 Tube hydroforming tool with provisions for addenda.
narrowest possible section reduces the amount of force necessary to clamp
the die halves together. Whenever possible, the design of the part will allow
the split line of the tooling to follow a planar cut, making the overall tooling
layout more straight forward (see Fig. 2.12). Often these simple split lines
are not possible and it becomes necessary to add additional material and
bends to the part design to simplify the interface with the docking rods. This
additional material, commonly referred to as an addendum, would ideally
allow the docking rod interface to take either the shape of the starting tube
geometry or at least a more simplified geometric shape to reduce the
machining complexity. Feed actuators are typically structured such that they
lie on the same plane and are coaxial, which allows them to be used for
multiple die cavities. Because of this, the part addendums need to be modified to conform to the feed actuator axis. Compound split lines and integrated feed actuators become more common as the complexity and the
quantity of the parts increase.
As the tooling increases in size and complexity, the use of segmented
inserts around the entire part allows for small geometry changes without
reworking the entire tool. Inserts permit the use of lower cost, softer tool
steels in the larger, non-critical portions and higher strength, harder tool
steels in the high wearing feed areas. When the part geometry creates low
areas in the lower die cavity, which do not drain naturally after the previous
parts are removed, the use of small vent holes may be required to allow the
trapped fluid to escape during the forming process. Fluid trapped between
WPNL2204
Hydroforming systems: equipment, controls and tooling
49
the die cavity and the outer surface of the tube could prevent the part from
forming all the way to the die wall because the fluid on each side of the
tube could reach a pressure balance.
The force capacity required from the feed actuators is a combination of
the force necessary to maintain a seal with the tube and the force, if necessary, to feed material into the tool. To seal a tube during forming, the
minimum force required can be calculated by multiplying the crosssectional area of the inside of the tube by the forming pressure (equation
2.4):
FS = pA
[2.4]
where FS = the minimum sealing force required (lb), p = the forming pressure (psi) and A = the cross-sectional area of the inside of the tube, at the
sealing interface (in2).
Equation 2.4 only calculates the force required to overcome the load
generated by the internal pressure trying to push the docking rod away
from the part. To create a seal at the docking rod interface typically requires
some mechanical deformation of the part, which varies with the type of
material, the wall thickness, and whether the process requires primarily
sealing or a higher amount of material feeding. A simple, angled tip can be
used for primarily sealing operations, where a combination of lead-in angle
and shoulder are required for feeding to reduce the radial pressures generated (see Fig. 2.13).
Many part geometries take advantage of material feeding to maintain, or
even locally build, wall thickness during the forming process. The forces
required for feeding are a function of the yield strength of the tube material,
Docking rod
Die
Deformation
Tube
Deformation
2.13 Tube hydroforming docking rod sealing styles.
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50
Hydroforming for advanced manufacturing
the part geometry, and the friction between the tube and the die cavity. As
the forming pressure increases, the feed force required to seal the part and
overcome friction also increase. These forces are additive and require FEA
for an accurate estimate of the forces to be expected during the process.
The estimated force required from the feed actuator, at any point in the
process, can be determined by adding the solution to equation 2.4, which
represents the force required to overcome the internal pressure generated
by the hydroforming fluid, with the solution to equation 2.5, which represents the force required to yield the tube:
FY = (YS)ATube
[2.5]
where FY = The force required to yield the tube material (lb), ATube = the
cross-sectional area of the tube, material only (in2) and (YS) = the yield
strength of the tube material (psi).
Typically, after a majority of the end feeding is complete and the material
being formed becomes stabilized against the walls of the die cavity, the
forming pressure can be increased until it reaches the calibration pressure.
Beyond this point, the sealing force is typically the major contributor to
the feed actuator load requirement. A conservative approach to use when
sizing the feed actuators is to solve for FS, at the maximum forming pressure, and then add it to FY. Generally, a safety factor of 25 to 50% is added
to this value to allow for modifications to the geometry and material
changes.
2.6
Future trends
As new applications require materials with increased strength and performance, systems will be required to be capable of providing higher
forming pressures. To achieve this will require increased tonnage from the
press and feed actuators and greater pressure capability from the intensification system. Material suppliers will also be encouraged to develop
higher strength materials with higher elongations, as the two do not typically go hand in hand. Innovative press designs, specific to the hydroforming
process, should provide more economical press solutions. While control
systems and electronics are continuing to become more sophisticated, they
are also becoming more affordable and easier to use. The extreme pressures, required for forming more exotic materials, will drive the need for
higher pressure sealing techniques. Because the tool materials will be
pushed to perform at higher stress levels, conventional forming techniques
and the die insert materials will benefit as they continue to advance. As
FEA is becoming a necessity in the manufacturing plant, advances in
computer and software technology help make this tool more accessible
and affordable.
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2.7
51
Sources of further information and advice
www.interlaken.com
www.ercnsm.org
www.cpforming.org
www.hydroforming.net
Singh H (2003), Fundamentals of Hydroforming, USA, Society of Manufacturing Engineers.
2.8
References
altan t (April, 2002), ‘Sheet hydroforming in automotive applications’, Stamping
Journal.
Interlaken Technology Corporation, 8175 Century Boulevard, Chaska, MN 55318,
USA, www.interlaken.com.
morphy g (May, 2002), ‘Tube hydroforming design flexibility. Part I’, The Tube and
Pipe Journal.
singh h (2003), Fundamentals of Hydroforming, USA, Society of Manufacturing
Engineers.
WPNL2204
3
Deformation mechanism and
fundamentals of hydroforming
CHRISTOPH HARTL,
Cologne University of Applied Sciences, Germany
3.1
Introduction
During the past few years, hydroforming technology has been successfully
adopted by several industries, providing mass production predominantly for
automotive components. Its successful adoption is due to the superior
advantages it offers compared with other conventional manufacturing techniques like welding of stamped components. These advantages essentially
come from being able to form hollow, complex-shaped components with
integrated structures starting from a single workpiece, combined with
improvements in stiffness and strength behaviour due to fewer welding
seams, and reduced assembly costs (Hartl, 2005).
Hydroforming processes are metal-forming processes based on the application of pressurized liquid media to generate a three-dimensional workpiece shape. With regard to existing hydroforming processes, a general
distinction should be drawn between the forming of tubular material, e.g.
straight or bent tubes or profiles, and the forming of sheet material, e.g.
single or multiple sheets. Currently, tubular material predominates in the
mass production of hydroformed components. However, advances in process
and press technology increasingly contribute to the wider industrial application of sheet hydroforming, which is of particular interest for the flexible
manufacture of small batch sizes.
The first classification of existing process variants has been developed
by (Dohmann, 1993a), taking into consideration the acting stress state
within the formed workpiece region and specific characteristics of the
expanded geometry. Currently, the classification of engineering standards
is being enhanced and updated with regard to the description of hydroforming processes, e.g. the engineering standard of manufacturing technologies DIN 8580, published by the German Institute for Standardization
(DIN).
In the hydroforming process, for the forming of tubular material, the
initial workpiece is placed into a die cavity which corresponds to the final
52
WPNL2204
Deformation mechanism and fundamentals of hydroforming
Initial tube
Fc
53
Top die
Fa
Fa
pi
Sealing punch
Bottom die
Hydroformed
component
a
Fc
Fa
Fa
pi
Counter punch
Fg
b
3.1 Principle of hydroforming processes: a rotationally symmetrical
component, b T-shaped component.
shape of the component, see Fig. 3.1a. The die, consisting of a top and
bottom half, is closed under the force Fc while the tube is (a) internally
pressurized with the internal pressure pi by a liquid medium to expand the
blank into the die cavity and (b) axially compressed by punches with the
axial load Fa to force material into the die cavity and to seal it. The component is formed under the simultaneously controlled actions of pi and Fa.
Depending on the part and process types, additional mechanical loads
can be applied to the workpiece. As an example, the hydroforming of Tshaped components, as shown in Fig. 3.1b, or Y-shaped parts, requires an
additional counter punch as well as suitable control of its counter force Fg
during the forming processes. This counter punch acts on the end of the
expanded protrusion and is displaced by the workpiece when the force
exerted by this is the same as Fg.
Design and optimization of hydroforming processes and components
require fundamental details to determine the necessary process loads, to
estimate feasibility, and to obtain an improved understanding of what influences the reliability and quality of component manufacturing. This chapter
deals with basic correlations between forming loads and forming results and
presents approved methods for the determination of suitable process
parameters with an emphasis on tube hydroforming.
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Hydroforming for advanced manufacturing
3.2
Stress and strain relationships
in tube hydroforming
For tube hydroforming, various theories and methods exist to determine
the correlation between forming loads, stress states, strains and non-reversible change of shape. Their efficient application depends on the complexity
of the particular problem being investigated. The following theories and
methods are the basis for the majority of existing solutions for modelling
hydroforming processes:
•
The Membrane Theory, a particular case of the Theory of Shells, assumes
thin-walled surface structures with constant stresses over the wall thickness and negligible bending stresses. The use of this theory provides a
statically defined system and enables stresses to be determined directly
from the equilibrium conditions of forces acting upon the structure.
• The Theory of Shells considers, in addition to the normal forces and
shear forces considered in the membrane theory, transverse forces,
bending moments and twisting moments within the surface structure. To
simplify matters, important assumptions are made that normal stresses
perpendicular to the shell surface are ignored, cross-sections remain
planar and no displacement appears between the outer and inner surface
of the structure. As the equations resulting from the equilibrium conditions of forces and moments describe a statically undefined system, the
determination of the stress state needs to take the fact of additional
strains appearing into account. The result is based upon the solution of
differential equations.
• The Continuum Theory of Plasticity enables the correlation between the
stress state and the resulting strain rates of a ductile continuum to be
described. The solution of a particular problem requires the determination of the strain rate distribution, based upon volume constancy of the
formed material, and provides differential equations of the locally acting
stress state by implying equilibrium conditions of the stress state, yield
criteria and material laws.
• The concept of the Finite Element Method is the discretization of an
investigated structure into a suitable number of finite elements identified by nodal points. With regard to the solution of forming problems,
the unknown variables are the displacements of the single nodes of each
finite element. The nodal point values to be determined are nodal forces.
The correlation between unknown variables and corresponding nodal
values, depending on material properties, is described by separate
element equations. The numerical solution of the assembled global
equations provides the basis for determining the distribution of stresses
or strains within the investigated structure. The formation of element
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Deformation mechanism and fundamentals of hydroforming
55
equations is accomplished by the direct approach method, the variational method, the method of weighted residuals or the energy balance
approach.
At present, fundamental investigations regarding the correlation between
applied loads and forming result have been carried out predominantly for
the hydroforming of straight rotationally symmetrical workpiece shapes
and for T-piece components. The methods used to derive suitable solutions
have been the Membrane Theory and the Continuum Theory of Plasticity
in the majority of cases. The further development of commercial programs
based on the Finite Element Method within the past few years now makes
a detailed and efficient analysis of forming processes possible, where the
component shape differs from parts with round cross-sections and a straight
axis (Dohmann, 2004).
For thin-walled rotationally symmetrical components under uniform
distributed axial load and internal pressure, the forming loads cause
stresses within the workpiece wall which can be referred to as a plain
stress state for the sake of simplicity, based on the assumptions of the
Membrane Theory. It can be described by the main stress sθ, acting in
a circumferential direction, and the main stress sϑ, acting in a longitudinal
direction, perpendicular to sθ and tangential to the component surface.
According to the theory being applied, a resulting stress acting perpendicular to the surface is ignored. The stresses sθ and sϑ can be obtained
from the equilibrium conditions for a rotationally symmetrical membrane
element under the distributed loads p1 along the membrane element
surface and p3 perpendicular to its surface, as shown in Fig. 3.2, which
can be written as:
d
( N ϑrθ sin ϑ ) − N θ rϑ cosϑ + p1rϑrθ sin ϑ = 0
dϑ
dq
a
[3.1]
Nϑ adq
p3rϑ dJadq
rϑ
rθ
Nθrϑ dJ
dϑ
NθrϑdJ
p1rϑ dJadq
Nϑ adq +
∂
(Nϑ adq)dJ
∂J
3.2 Equilibrium of forces for a rotationally symmetrical membrane
element.
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Hydroforming for advanced manufacturing
Nϑrθ + Nθrϑ + p3rϑrθ = 0
[3.2]
where Nθ and Nϑ are the resulting linear distributed loads in a circumferential direction and in a longitudinal direction, respectively. The parameters
rθ and rϑ are the radii of curvature and J is the angle within the plane of
the meridian, see Fig. 3.2. With the solution of the equations [3.1] and [3.2],
the resulting stresses sθ in a circumferential direction and sϑ in a longitudinal direction can then be obtained by the division of Nθ and Nϑ, respectively by the wall thickness t of the membrane.
Klaas (1987) derived the conditions for yield initiation of a cylindrical
straight tube under axial force and internal pressure and the resulting stress
state, based on equations [3.1] and [3.2]. According to these, the circumferential stress sq within a thin-walled straight tube, see Fig. 3.3, can be described
as follows:
σθ = pi
d0 − t0
2 t0
[3.3]
with the initial outer tube diameter d0 and the initial tube wall thickness t0.
The stress in the longitudinal direction, here called axial stress sz, can be
written as
σz =
(
1
p
pi (d0 − 2t0 )2 − Fa
4
p (d0 − t0 )t0
)
[3.4]
The forming of metal components implies that the applied forming loads
induce a sufficient stress state to achieve the current yield stress sY of the
formed material. Assuming the von Mises Yield Criterion, when considering a biaxial stress state, with the current effective stress seff of the tube
material
σ eff = σ θ2 + σ ϑ2 − σ θσ ϑ
[3.5]
Fa
sz
t0
Workpiece
pi
d0
sθ
sθ
sz
Sealing
punch
Fa
z
r
3.3 Tube under axial load and internal pressure.
WPNL2204
Deformation mechanism and fundamentals of hydroforming
57
the beginning of plastic yielding can be determined for seff = sY with sz =
sϑ.
The progression of circumferential and longitudinal stress during the
forming process depends on the applied control of Fa and pi, the material
characteristics, friction conditions within the contact surface of tube and
tool, and the resulting change in workpiece shape. According to experimental and theoretical investigations carried out by Klaas (1987), the ratio of
sϑ/sθ within the maximal expanded diameter of hydroformed spherical
workpieces changes from about −1.4 at the start of the process to between
about −0.3 and 0 at the process end.
In general, stresses and strains in forming processes do not directly correspond. According to the equations developed by Levi and von Mises,
only the momentary plastic strain increments can be calculated in relation
to the currently acting stresses. With regard to the plane stress state of a
rotationally symmetrical component shape, where a stress acting perpendicular to the surface is ignored, the following correlations can be defined
from this:
dε θ =
dλ
(2σ θ − σ ϑ )
3
[3.6]
dε ϑ =
dλ
(2σ ϑ − σ θ )
3
[3.7]
dε t =
dλ
(−σ θ − σ ϑ )
3
[3.8]
with the strain increments deθ in a circumferential direction, deϑ in a
longitudinal direction, det across the wall thickness, and an alterable
scalar l. It should be mentioned here that, due to the higher amount
of plastic strains in comparison to elastic strains, the latter are commonly
ignored.
The effective strain increment deeff, which results from these three strain
increments for the plane stress state considered here, can be generally
defined as:
4
(dε θ2 + dε ϑ2 + dε θ dε ϑ )
3
dε eff =
[3.9]
For the determination of strains, assuming constant strain increments and
unchanged principal directions of deformation during the forming process,
the strains can be written as logarithmic strains:
ε θ = ln
r
r0
[3.10]
ε ϑ = −ε θ − ε t
[3.11]
WPNL2204
58
Hydroforming for advanced manufacturing
ε t = ln
t
t0
[3.12]
with initial mean tube radius r0, expanded mean radius r corresponding to
rθ in Fig. 3.2, initial wall thickness t0 and wall thickness of the expanded
workpiece t. The strain eϑ in a longitudinal direction is derived here from
the condition of incompressibility of the material:
eθ + eϑ + et = 0
[3.13]
With the assumptions made above, the effective stress can then be derived
from equation [3.9] as:
ε eff =
4 2
(ε θ + ε ϑ2 + ε θε ϑ )
3
[3.14]
Using appropriate and generally known material models enables
the relationship between effected strains and resulting yield stress to be
described, e.g.:
n
sY = Ce eff
[3.15]
sY = C(e0 + eeff)n
[3.16]
or
where C is a material constant, n is the strain hardening exponent of the
material and e0 is an effective pre-strain in the material.
Examples of investigations into modelling hydroforming processes by the
use of the Membrane Theory and the Levi and von Mises correlations have
been presented by, e.g. Woo (1973), Sauer (1978) and Fuchizawa (1984) for
the free expansion of rotationally symmetrical workpiece shapes with a
straight axis. The main aspect of this research work was to determine material parameters and their respective influences on the forming result.
These investigations produced iterative, computer-based programs, applying diverse simplifications. Whereas Fuchizawa considered an expansion
merely caused by an internal pressure pi, the investigations conducted by
Woo and Sauer included the application of an additional axial force Fa.
However, Woo’s concept implies that there are only tensile stresses within
the tube wall.
Klaas (1987) applied the Membrane Theory, the Theory of Shells and the
Continuum Theory of Plasticity to derive miscellaneous closed solutions for
the determination of relations between loads, stresses, and strains for the
hydroforming of spherical workpieces under internal pressure and axial
force. The theoretical basis developed was used for the analysis of comprehensive experimental investigations. Based on the Membrane Theory,
WPNL2204
Deformation mechanism and fundamentals of hydroforming
59
Schmoekel (1992) developed an algorithm for a pressure-dependent determination of axial stroke and force parameters.
In principle, deviations of the above-mentioned plane stress state due to
bending forces, e.g. within the area of the workpiece clamped at the tube
ends, can be considered by using the Theory of Shells. However, comparisons of analytically determined forming results together with experiments
have shown the inaccuracy of the calculated shape increases with increasing
expansion of the component (Klaas, 1987).
An essential objective when designing hydroforming processes is to
maintain the initial wall thickness t0 along the hydroformed component.
However, a change in wall thickness is unavoidable in the course of the
expansion process. Figure 3.4 shows the wall thickness distribution for an
exemplary result of a rotationally symmetrical expanded component, here
compared with a numerically simulated result obtained by the use of the
Finite Element Method (Dohmann, 1993a). In general, the expansion of
the hydroformed component involves a reduction in wall thickness, which
is decisively influenced by the amount of expansion resulting from the
applied internal pressure pi, and the amount of compressive axial stress
induced by the axial force Fa. The higher the compressive axial stress, the
lower is the reduction in wall thickness. In theory, the wall thickness remains
constant for an expansion where the ratio of axial to tangential stress is
given as sϑ/sθ = −1 (Sauer, 1978). However, the occurrence of certain failure
modes (see section 3.3) limits the applicable amounts of the loads pi and
Fa, and with that the achievable amount of expansion without wall thickness reduction.
Conventional sheet-forming technology takes advantage of the anisotropic behaviour of the sheet material used, specifically to reduce the
decrease in thickness of the formed blank. The influence of the anisotropy
t0 = 5.0 mm
FEA
Experiment
Length (mm)
300
Hydroformed
cross-section
0
4.0 4.5 5.0 5.5
Wall thickness (mm)
3.4 Wall thickness distribution for the example of a hydroformed
rotationally symmetrical workpiece (Dohmann, 1993a).
WPNL2204
60
Hydroforming for advanced manufacturing
of the tube material on the forming result has been theoretically investigated by Fuchizawa (1987) for the hydroforming of tubes without the assistance of an axial force Fa, but with tube ends permitted to move axially.
According to these investigations, with an increasing anisotropic parameter
in a longitudinal direction, thickness reduction decreases and critical
expanding diameter increases.
Besides the research on hydroforming of rotationally symmetrical shapes,
there are miscellaneous investigations for the hydroforming of T-shaped
and Y-shaped components regarding the determination of correlations
between forming loads, and resulting stresses, strains and forming results.
However, due to the non-axisymmetric forming conditions, these correlations are considerably more complex than rotationally symmetrical forming
operations. The wall thickness distribution of an expanded Y-shaped component in Fig. 3.5 (Klaas, 1997) shows that irregular stress states are acting
during the forming process along the formed part, with predominantly axial
compressive stresses within the main part axis and tensile stresses within
the formed protrusion. The forming process of such geometries requires
higher axial loads and correspondingly higher axial stresses than processes
with predominantly free expansion of the workpiece. These higher axial
loads are possible because the tube wall is almost completely in contact
with the surrounding tool during the whole forming process. The support
provided to the tube wall by the tool from the outside and the internal
pressure from inside reduces the risk of failures like wrinkling caused by
increased axial stresses.
To determine the acting stress state, it is necessary to make simplifying
assumptions in order to obtain closed solutions, e.g. the subdivision of the
formed component into sections with an assumed uniform stress distribution within each section, as described for the hydroforming of T-shaped
components by (Chalupczak, 1983). Nowadays, the Finite Element Method
0.9
1.4
1.4
1.9
0.9
1.3
Wall thickness (mm)
1.6
1.7
2.0
2.0 2.1
1.6 2.2
2.2
2.0
Material: AISI 309
d0: 42 mm
t0: 1.2 mm
L0: 210 mm
Length 125 mm
3.5 Wall thickness distribution of a hydroformed Y-shaped exhaust
system component (Klaas, 1997).
WPNL2204
Deformation mechanism and fundamentals of hydroforming
61
Fa
Sealing punch
Workpiece
t
Pi
d0
Hydroforming tool
z
r
–2
sz
sY
sr
sY
–1
0
sθ
sY
pi = 0.5 sY
t/d0 = 0.2
3.6 Schematic end section of a hydroformed components and stress
state within the tube wall under axial load and internal pressure.
provides an economic way to obtain reliable results regarding such complex
geometries, e.g. Altan (2002) and Jirathearanat (2004).
However, the idealization of the end sections of the workpiece as rotationally symmetrical loaded workpiece regions allows the stress state to be
determined by analytical means within this region of the tube wall when
plastic yielding occurs. Figure 3.6 shows the end section of an exemplary
forming situation, where the formed workpiece shape consists, for example,
of a T-shaped element or any other geometry which has to be kept in a
plastic yielding state predominantly by axial loads. The stresses acting at the
workpiece end, under the condition where the yield stress sY of the material
is achieved due to the axial load Fa and the internal pressure pi, can be
written as
sr = −
sθ = −
sz = −
1
2 3
1
2 3
1
2 3
s y ln Ar − pi
[3.17]
s y(ln Ar + 4 /ar ) − pi
[3.18]
s y(ln Ar + 2br /ar ) − pi
[3.19]
where the parameters Ar, ar and br are functions of the radial direction r:
ar − 1 c + 1
ar + 1 c − 1
[3.20]
ar = 3(2r/d0 )4 + 1
[3.21]
Ar =
WPNL2204
62
Hydroforming for advanced manufacturing
br = 3(2r/d0)2 + 1
[3.22]
and the tube geometry-dependent parameter c, with wall thickness t and
outer diameter of the tube end section d, where the following applies:
c = 3(1 − 2t/d0 )4 + 1
[3.23]
These equations have been derived with the aid of the Continuum Theory
of Plasticity and provide the necessary background for determining required
forming loads and loads acting on tool elements (Hartl, 1995). Here the
tube is assumed to be a thick-walled tube with radial stress sr ≠ 0, in contrast
to the assumption of the Membrane Theory which ignores stresses in directions normal to the tube wall. An exemplary situation of stresses acting
within the loaded tube wall, represented in Fig. 3.6, shows a non-uniform
distribution of sθ and sz and sr along the wall thickness. From equation
[3.17], the contact pressure sN between the tube wall and the surrounding
tool can be derived with sN = −sr (r = d0/2):
sN = bsY + pi
[3.24]
with b ≈ 1.8 t/d0 when the ratio of wall thickness to tube diameter t/d0 is
<0.2.
The use of equations [3.17] to [3.24] requires knowledge of the current
wall thickness and yield stress of the formed workpiece. Analogous to the
determination of correlations between the stress state and strains based on
the Membrane theory, in this case, there is also no direct correlation between
these variables. Hence, an iterative procedure would be required to determine load-dependent deformation, involving equilibrium conditions and
material laws.
3.3
Determination of forming limits
The workpiece geometries attainable with hydroforming processes are
limited by the occurrence of failures. These instabilities are predominantly
local necking and bursting of the workpiece wall, local or extensive wrinkling of the workpiece, and buckling of the initial part.
Necking occurs when the formability of the workpiece material is
exceeded in an area causing the hydroformed component to burst. A criterion for the start of necking, with regard to the rotationally symmetric
expansion of straight tubes, was investigated by Sauer (1978), by assuming
a constant ratio of axial to tangential stress sϑ/sθ = const. The findings were
that this form of necking occurs when the logarithmic strain in wall thickness direction et corresponds to the negative value of strain hardening
coefficient et = −n. In general, the requirement sϑ/sθ = const. does not apply
WPNL2204
Deformation mechanism and fundamentals of hydroforming
63
to conventional hydroforming processes. However, fundamental investigations have shown that it can be used for a first estimation.
For the prediction of the internal pressure pb at the moment of bursting
for straight tubes within a state of free expansion, the correlation investigated by Klaas (1987)
pb = s UTS
2 t0
d0 − t0
[3.25]
Variation of
wall thickness
has been shown to be applicable. Here, sUTS is the tensile strength of the
material. The amount of pb should not be exceeded within a hydroforming
process while the tube is not in contact with the surrounding die cavity. The
internal pressure should only be increased higher than pb if large areas of
the expanded tube are aligned to the die cavity.
For the expansion of rotationally symmetrical components, bursting preceded by necking predominantly occurs within the area of largest expansion. An increase in axial force Fa, within certain limits, increases the axial
compression stress and, with this, the feasible expansion diameter until
necking and bursting are induced (Dohmann, 1993a). With regard to the
expansion of T-shaped or Y-shaped components, necking and bursting occur
as a result of an internal pressure being exceeded at the front of the
expanded protrusion. Adapting the control for the counter punch force Fg
can prevent this failure situation up to a certain point. The application of
measures to reduce friction between the hydroformed workpiece and the
tool can also assist failure-free expansion, due to an improved material
flow.
Complex-shaped components with varying cross-sections along the workpiece axis suffer bursting induced by necking predominantly within crosssection corner areas. The influence of friction on the wall thickness
distribution for an exemplary cross-section is represented in Fig. 3.7, determined by a process simulation with the aid of the Finite Element Method
(Dohmann, 1997). It is obvious from this figure that an increased friction
+10%
8
0
–10%
1 2 3
7
6
2
3
5
1
4 5 6 78 1
m = 0.1
m = 0.01
4
Hydroformed
cross-section
3.7 Effect of friction on the wall thickness distribution (Dohmann,
1997).
WPNL2204
64
Hydroforming for advanced manufacturing
coefficient leads to a local decrease in wall thickness at the corner radii of
the hydroformed component and, thus, to premature necking. Hence, measures aimed at friction reduction enhance feasible expansions and improve
process reliability. In addition, due to previous bending of the initial tube,
the wall thickness is decisively reduced within the bent areas and formability is exhausted to a large extent (Dohmann, 2004). As a consequence, these
workpiece areas tend to fail prematurely as a result of necking and bursting.
It is recommended that the corner radii are increased and/or the overall
expanded cross-section circumference is decreased to avoid a recurrence of
these instabilities.
In conjunction with process simulations based on the use of the Finite
Element Method, Forming Limit Curves are now used for the prediction
of the start of necking, e.g. Koc (2004) and Kroeff (2002). Detailed recommendations for the application of Forming Limit Curves for tube hydroforming have been presented by (Levy, 1999). With the objective of taking the
varying stress states during the hydroforming process into consideration,
Forming Limit Stress Diagrams have been investigated by Prahl (2002),
applying modified Forming Limit Curves.
Besides being influenced by the process controls and hydroformed component geometry, necking and bursting are also induced by the properties
of the applied semi-finished product. As an example, welding seams of longitudinally welded tubes or extruded profiles can be starting points of these
failures.
Instabilities due to wrinkling result predominantly from excessive axial
loads. This failure situation limits the amount of axial stress that can be
applied to reduce the decrease in wall thickness during the hydroforming
process. An axisymmetric component displays wrinkling during the course
of its free expansion whereas the wrinkling of T-pieces and Y-shaped components occurs despite the fact that the tube wall is in contact with the
surrounding tool. In each of these cases, an adapted control of axial force
and internal pressure should be applied to avoid this instability. Reliable
methods to predetermine the occurrence of wrinkling only exist for an
axially loaded tube without internal pressure and unsupported by a surrounding tool, as presented in Geckeler (1928). According to this, wrinkling
occurs in this case when the axial stress within the formed tube exceeds the
critical stress:
s crit =
2
tE
r 3(1 − ν ) 1 + E/T
2
[3.26]
where t is the wall thickness, r the mean tube radius, n the Poission ratio, E
the elastic modulus and T the tangential modulus with T = ∂sY /∂eeff.
For the avoidance of wrinkling during the hydroforming of T-shaped
components, Bogojavlenski and Serjakow developed an empirical
WPNL2204
Deformation mechanism and fundamentals of hydroforming
65
correlation to determine the minimum required internal pressure (Zschekel,
1977):
pmin = (0.13 + 1.15t0/da)sY
[3.27]
with the protrusion diameter da. Their recommendation is to apply this
internal pressure independently from the counter punch force Fg and the
tube length.
In addition, wrinkles in the longitudinal direction of the workpiece can
occur when closing the hydroforming tool if incorrect dimensions for the
semi-finished product or for the preformed component geometry have been
selected. The Finite Element Method can be used to optimise the process
chain with the objective of eliminating such failures, a practical example of
which is given by Boehm (2000).
In a similar way to the formation of wrinkles, buckling of the workpiece
occurs under excessive axial loads. The free tube length, unsupported by
surrounding tool surfaces, is an important parameter for this instability. The
application of an adequate control of Fa and pi is required to avoid this
failure case. Based on the determination of buckling with plastic material
behaviour, Dohmann (1993b) and Bieling (1992) presented an iterative
method to determine suitable load paths for the hydroforming of rotationally symmetrical workpieces with maximum compressive stress. The axial
load that resulted in buckling was taken as a limit for the maximum applied
load with the objective of reducing the decrease in wall thickness during
the hydroforming process.
As well as the geometric parameters and material properties of the
workpiece and tool, the occurrence of instabilities depends predominantly
on the process control selected for the forming loads Fa and pi. According to the investigations into hydroforming carried out by Klaas (1987),
Fig. 3.8 shows schematically the range of feasible process controls for
Axial force Fa
Wrinkling / buckling
Working region
Bursting
Yield start
Sealing force Fp
Internal pressure pi
3.8 Range of feasible process controls for the hydroforming of
rotationally symmetrical components, according to Klaas (1987).
WPNL2204
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Hydroforming for advanced manufacturing
a hydroforming process limited by the occurrence of instabilities, the
yield start of the workpiece and the minimum required force to seal
the tube.
Examples of achievable workpiece geometries within the forming limits
being discussed, assuming a comparatively satisfactory formability of the
component material and a reliable and economic production operation,
have been presented by Klaas (1997) and Dohmann (1998). In general,
the amount of feasible expansion is decisively influenced by the failurefree applicable axial load Fa. However, long feeding sections and bent
workpiece areas, where workpiece material has to be transported by
the axial force into the area to be expanded, impede the material flow
due to friction forces and additional bending forces, respectively (Hartl,
1999).
3.4
Forming loads and process control
Control of the forming loads during the hydroforming process should be
appropriate for obtaining the required forming result due to a continuously
achieved yield stress of the material, avoiding failures such as wrinkling,
buckling and bursting. The internal pressure pi and the axial force Fa are
decisive process control parameters for hydroforming processes. Commonly, these process parameters are determined according to the length of
the forming process. Their variation depends predominantly on the material
properties and strain hardening behaviour, on the tube wall thickness, on
the size of intricate sections of the component such as small corner radii,
and on the potential occurrence of instabilities. As an example, Fig. 3.9
shows the sequence of a typical hydroforming process including the sealing
of the tube, the forming with axial feeding and internal pressure, and the
calibration (Jirathearanat, 2004).
Sealing
Forming
Calibration
Relief of loads
1500 bar
Internal pressure
Workpiece shape
Axial stroke (right)
Axial stroke (left)
Counter punch stroke
100 mm
0
0
time (s)
15
3.9 Exemplary process control for hydroforming, according to
Jirathearanat (2004).
WPNL2204
Deformation mechanism and fundamentals of hydroforming
67
The equations [3.3] to [3.5] enable the required process parameters to be
derived to start the forming of the initial workpiece. Plastic yielding of the
tube is ensured by assuming that the effective stress seff, which results from
the combination of axial and circumferential stress, corresponds to the local
yield strength sY of the tube material. From this, the correlation between
internal pressure pi and the resulting axial force Fa, which is suitable for
inducing plastic deformation of the tube, can be derived as follows (Klaas,
1987):
2
⎡
3 2 ⎛ d0 − t0 ⎞
π
d0 − t0 ⎤
2
Fa = π (d0 − t0 )t0 ⎢ σ Y2 −
pi ⎜
[3.28]
⎥ + pi (d0 − 2t0 )
⎟ − pi
⎝
⎠
16
4
4
t
t
0
0
⎢⎣
⎥⎦
This equation describes an elliptic curve which is represented in Fig. 3.8 as
the yield start of the hydroformed workpiece.
The minimum axial force required throughout the overall process to
ensure the sealing of the workpiece ends can be determined as
Fp = pi
π
(d0 − 2t0 )2
4
[3.29]
An axial force Fa less than Fp will lead to a leakage between the hydroformed tube ends and the punches and, thus, to a premature termination of
the process.
At the end of the forming process, the tube wall has to be formed into
corner radii of the die cavity which have not been formed during the main
expansion of the tube. This requires raising the internal pressure up to its
maximum value pk. Various investigations have already dealt with determining the necessary calibration pressure, e.g. with the aid of the Finite Element
Method (Boehm, 1993) or by analytical means (Dudziak, 1996 and Birkert,
1997). A correlation developed by Koc (1999) gives:
pk =
(
2
σ
r
σ Y ln ⎛⎜ c ⎞⎟ 2 − Y
r
−
t
σ
⎝ c ⎠
3
UTS
)
[3.30]
with the yield stress sY, the ultimate tensile strength sUTS of the tube material, the wall thickness t at the formed radius and the outside corner radius
rc of the die cavity which is to be formed. An equation empirically deduced
by Braeutigam (1992), which is suitable for a first estimation to determine
the maximum necessary internal pressure, is written as:
t
pk ≈ 1.2σ UTS 0
rc
[3.31]
Figure 3.10 shows the ratio pk /sY against the ratio of the die cavity radius
rc to the tube wall thickness t0 in principle according to Hartl (1999). It is
obvious from this figure that the smaller the ratio rc /t0, the higher is the
amount of internal pressure necessary to form the radius.
WPNL2204
68
Hydroforming for advanced manufacturing
3
pk
sY
2
1
0
0
1
2
3
4
5
rc
t0
3.10 Ratio of calibration pressure to yield stress versus the ratio of die
cavity radius to wall thickness.
Fa
Sealing punch
Hydroforming
tool
t
Workpiece
pi
d0
z
Fp
Fu Fr
r
3.11 Force components at the end section of a hydroformed
workpiece.
To determine the axial force Fa shown in Fig. 3.11 throughout the hydroforming process up to its conclusion, this force can generally be made up
of three individual forces:
Fa = Fz + Fp + Ff
[3.32]
The force Fz is the axial force component which is initiated in the tube wall
and, together with the action of the internal pressure, maintains the plastic
flow of the tube wall. For forming processes where plastic yielding is predominantly induced by the axial load, e.g. hydroforming of T-shaped components, the determination of the axial force component Fz can be derived
from equations [3.17] to [3.23], based on the correlations of stresses within
the tube end under plastic yielding. The following applies according to Hartl
(1995):
Fz = ζ tσ Y d 20
π
+ pi tπ (d0 − t )
4
[3.33]
with zt ≈ 4.41 t/d0 for ratios of wall thickness to tube diameter, where t/d0 is
below 0.3. To calculate the amount of the axial force component Fz for
the hydroforming of components with predominantly free expansion, the
correlation
WPNL2204
Deformation mechanism and fundamentals of hydroforming
Fz ≈ asUTSt0p(d0 − t0)
69
[3.34]
has shown to be applicable for a first estimation, with a = 1.2 . . . 2.0. Equation [3.29] applies for the sealing force Fp which results from the reaction
force of the internal pressure acting on the front punch face. Because the
tube ends are in contact with the hydroforming tool throughout the forming
process, the frictional force Ff must be overcome in this section. When
Coulomb’s frictional behaviour is taken as a basis, the friction force is given
as:
Ff = msNd0plr
[3.35]
with sN as shown in equation [3.24] and lr as the length of the tube section
moved along the tool under friction.
In general, the maximum values for Fa and pi are applied at the end of
the forming process when the workpiece is calibrated through an increase
in the internal pressure up to the magnitude pk. Similar correlations to
determine the axial force for the hydroforming of rotationally symmetrical
components have been derived by, for example, (Klaas, 1987) and for the
forming of T-shaped parts by, for example, (Chalupzak, 1983). With regard
to the maximal amount of Fa, an upper limit is the occurrence of instabilities
like wrinkling or buckling (see section 3.3).
Throughout the hydroforming operation, it must be ensured that the
hydroforming tool remains closed by the closing force Fc, applied by
the forming press. The following applies for the minimum required
force:
Fc = piAp
[3.36]
The projected component surface Ap, perpendicular to the closing direction, and the maximum amount of applied internal pressure are decisive for
determining the size of Fc.
3.5
Preceding forming operations
The complexity of hydroformed components requires that, in the majority
of cases, additional preceding forming operations should be considered
together with the hydroforming process itself. These operations can involve
the bending and mechanical forming (preforming) of the initial component
to ensure that it is capable of being inserted into the hydroforming die or
to obtain an optimised material distribution. The forming result of these
preceding operations influences the hydroforming process that follows and
the final component quality. Therefore, the properties and forming limits of
the bent and preformed tubes are of interest.
WPNL2204
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Hydroforming for advanced manufacturing
3.5.1 Tube bending
In general, typical bending processes are rotary draw bending for complex
bent components and press bending for less complex shapes with large
bending radii (Hartl, 2005). In the case of a bent tube geometry with a
constant bending radius, Fig. 3.12 shows the bent area schematically with
the most important parameters. The centreline of the bent part axis is specified by the mean bending radius rm and the bending angle ab. The radius ra
describes the outer radius of the bent tube and ri specifies the inner radius.
Close to rm the radius ru is located where no variation in length of the bent
tube occurs. The position of ru changes during the bending process depending on the workpiece properties, the bent geometry and the kind of bending
process (Franz, 1961). As shown schematically in Fig. 3.12, during the
bending process, tensile stresses parallel to the centreline are acting towards
the tube wall within the outer bend area and compressive stresses are acting
within the inner bend area. The tensile stresses cause a reduction in wall
thickness; the maximum reduction is to be found at the outer bending radius
ra, specified by the minimum resulting wall thickness tmin = ta. Due to the
compressive stress state within the inner bend area, the tube wall thickness
increases. The maximum wall thickness caused by this is located along the
radius ri, specified by the maximum wall thickness tmax = ti. Figure 3.13 shows
the variation of wall thickness distribution for the example of bent tubes.
Nowadays, detailed process simulation using the Finite Element method
enables a similar precise prediction of wall thickness distribution and material strain hardening induced by bending processes, e.g. Dohmann (2004).
For a first approximation of the resulting wall thickness at ra and ri, an estimation can be made based on the volume constancy of formed metal material with:
tmin,max ≈ t0
1
d
1± 0
2rm
[3.37]
ta
ra
–s
ri
ra
ti
ab
+s
ri
ru
rm
t0
d0
3.12 Parameters in tube bending.
WPNL2204
Stress state
(schematic)
Variation of wall thickness
Deformation mechanism and fundamentals of hydroforming
20%
0
Aluminium tube
Steel tube
10%
0
71
θ
60° 120° 180° 240° 300° 360°
θ
–10%
Tube diameter: 20 mm
Tube wall thickness: 1.5 mm
Bending radius: 50 mm
Bending angle: 90°
–20%
3.13 Example of wall thickness variation of bent tubes according to
Khodayari (2002).
15
A
B
C
D
10
rm
d0
E
5
F
0
0
50
100
150
200
d0
A
B
C
D
E
F
t0
Without mandrel
Rigid mandrel
Mandrel with one ball
Mandrel with two balls
Mandrel with multiple balls
Bending not feasible
3.14 Limits of rotary draw bending according to Franz (1961).
when applying the simplifying assumptions that the bent tube cross-sections
remain planar and circular and ru remains identical to rm.
Limits for bending processes commonly result from a local necking and
the appearance of a fracture at ra as well as from wrinkles at ri. Based on
experimental investigation, Franz (1961) developed the working diagram
presented in Fig. 3.14 which shows feasible bend geometries depending on
the ratio of outer tube diameter to tube wall thickness d0/t0 and the ratio of
mean bending radius to the tube diameter rm /d0 for different bending strategies. This diagram is applicable to rotary draw bending processes and still
WPNL2204
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Hydroforming for advanced manufacturing
serves as a basis for the design of tube bending processes. Newly developed
experimental and theoretical results presented by Flehmig (2001 and 2002)
point out that this diagram is based on experiments with stainless steel
tubes with an ultimate strength of about 350 Mpa and that variations of
parameters such as material specifications, bending angle, tool design, friction and distortion of the bent cross-section are not considered.
Based on the simplifying assumptions that linear uniaxial stresses alone
are acting during the bending process without the appearance of shear
stresses, that ru remains identical to rm and that the bent tube cross-sections
remain planar and circular, Flehmig derived conditions for rotary draw
bending processes to determine the forming limits depending on the tube
material behaviour. According to this, the limit of wrinkling can be written
as:
σY
( T1 − E1 ) ≥ 6 fd r ⎡⎢⎣( 2kd t ) − ( 6dr ) ⎤⎥⎦
2
2
m m
R 0
0
0
0
m
[3.38]
and the limit of fracture can be calculated by applying the correlation:
σY
r ⎡d
− 88.857 ⎤
( T1 − E1 ) = 24.157
⎢⎣ t
⎥⎦
d
m
0
0
[3.39]
0
with the limit of elasticity sY, the elastic modulus E and the tangential
modulus T. The parameter kR is a correction value which can be estimated
with kR = 2.22 + 8 t0/d0. The parameter fm considers the kind of mandrel used
to support the tube from the inner side. For non-mandrel bending, fm is 1.0;
for rigid mandrels, fm is between 1.38 and 1.49; for mandrel-bending with
one or three balls, fm is 2.56 and 5.06, respectively. According to Flehmig,
the derived equations have shown their applicability in numerical process
simulations with the Finite Element Method and in experiments.
When designing a tube-bending process, the elastic behaviour of the bent
tube has to be taken into account. The elastic stress state, remaining at the
end of the bending process, leads to an elastic deflection of the bent component after it has been relieved from loads, visible from the increase of
the bent angle. This spring-back effect results from the compensation of
elastic tensile stresses at the outer bend area for compressive stresses at the
inner area of the bend. However, a considerable amount of elastic stress
still remains within the bend after this elastic spring back. In general, the
spring back is described by the ratio:
K=
αp
<1
αb
[3.40]
with the bending angle ab applied by the bending process and the angle ap
after elastic spring back. The size of K depends on the properties of the
tube material and its dimensions, the kind of bending process applied and
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Deformation mechanism and fundamentals of hydroforming
73
dr
dq
3.15 Cross-section deflection of bent tubes (schematically).
the bend geometry produced. Large bending radii with correspondingly
high ratios of rm/d0 cause a decrease in K together with an increase in spring
back (Franz, 1961). Also the scattering of K increases when raising the ratio
rm/d0.
Besides variations in wall thickness caused by the bending operation and
elastic spring back, the distortion of the initially round cross-section due to
this forming process is of further importance. Depending on the principle
of the bending process applied, used bending die elements, bent geometry
and tube parameters, the distortion results in a more or less distinctive oval
shape of the bent tube cross-section, as shown in Fig. 3.15. The degree of
distortion can be described by
u=
dq − dr
d0
[3.41]
with the cross-section width dr in the direction of the bending radius and
the cross-section width dq perpendicular to the bending radius. According
to Franz (1961), u increases with a decreasing ratio of wall thickness to tube
diameter t0/d0 as well as with a decreasing ratio of mean bending radius to
tube diameter rm/d0. In addition, improved support of the inner tube wall
by suitable mandrels reduces u.
3.5.2 Preforming
Preforming operations are predominantly used to enable the reliable insertion of the initial workpiece into the hydroforming die. This applies, for
example, in cases of tube diameters d0 which are larger than the die cavity
width w. The preforming operation reduces the tube cross-section towards
the value of w. Typical methods for preforming are shown schematically in
Fig. 3.16. Methods using bevelled dies as shown in Fig. 3.16a are suitable
for the preforming of straight as well as bent tube sections, whereas the
methods using cross slides for local workpiece reduction are limited when
bent tube sections are to be preformed. In certain cases, preforming is also
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Hydroforming for advanced manufacturing
Punch
Preforming die
Tube cross
Section
Fpf
Preformed
cross-section
a
Cross slides
Fpf
Fpf
Preforming die
b
3.16 Principles of preforming.
used to flatten tube sections, e.g. when flattening by closing the hydroforming tool is not reliably feasible.
Problems caused by preforming consist of shape deviations which remain
after the hydroforming process. These are, for example, wrinkles due to an
inappropriate distribution of material or instabilities, excessive small radii
generated by local folding of the tube wall as well as excessive elastic spring
back of flat workpiece sections due to an insufficient degree of forming
within the hydroforming process.
Besides an adequate tool design, the layout of preforming processes also
requires the determination of the forces required to carry out this forming
operation. From practical experience, the following correlation for the preforming of a circular tube cross-section to a rectangular cross-section has
been derived:
Fpf = VsUTSlt20
[3.42]
with sUTS the ultimate tensile strength of the tube material, l the length of
the preformed section and t0 the tube wall thickness. The factor V considers
whether straight or bent tube sections are preformed, as bent tube sections
requires higher preforming forces in comparison to straight sections. Experiments have shown that V is between about 0.11 and 0.32. In addition, an
extra bead in the longitudinal direction of the formed tubes doubles the
preforming force Fpf.
As with the design of bending processes, the Finite Element method can
also be used here for a detailed process simulation, allowing a similar
precise prediction of results for preforming and the successive hydroforming, and enabling process and tool optimization (Dohmann, 2004).
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Deformation mechanism and fundamentals of hydroforming
3.6
75
References
altan (2002), ‘Process simulation for hydroforming components from sheet and
tube’, Proc 3rd Chemnitz Car Body Colloquium, Chemnitz, 117–132.
bieling p (1992), Untersuchungen zum Aufweitstauchen von Rohren zu Hohlwellen,
doctoral dissertation, Paderborn.
birkert a (1997), ‘Abschaetzung der Kalibrierdruecke beim InnenhochdruckUmformen’, Blech Rohre Profile, 9, 102–107.
boehm a (1993), Numerische Simulation von Verfahren der Innenhochdruckumformung unter besonderer Beruecksichtigung des Aufweitens im geschlossenen
Gesenk, doctoral dissertation, Paderborn.
boehm a, hartl c and abbey t (2000), ‘Process and tool technology for hydroforming:
case study and technical and economic considerations’, Proc Int Conf Tube
Hydroforming, Columbus OH.
braeutigam m and rutsch h (1992), ‘Hydroformen – als Ausweg aus der Investitionsklemme’, in VDI Berichte Nr 946, Duesseldorf, VDI.
chalupzak j and sadok l (1983), ‘The problem of forces and stresses in the hydromechanical process of bulge forming of tubes’, Metalurgia I Odlewnictwo – Tqm
9 – Zeszyt 1, 57–66.
dohmann f (1993a), ‘Innenhochdruckumformen’ in Lange K, Umformtechnik,
Berlin, Springer, 252–270.
dohmann f, boehm a, dudziak k-u (1993b), ‘The shaping of hollow shaft-shaped
workpieces by liquid bulge forming’, Adv Tech Plasticity, 447–452.
dohmann f and hartl c (1997), ‘Tube hydroforming – research and practical application’, J Mat Proc Tech, 71, 174–186.
dohmann f and hartl c (1998), ‘Hydroforming components for automotive applications’, The Fabricator, 2, 30–38.
dohmann f and hartl c (2004), ‘Hydroforming applications of coherent FEsimulations to the development of products and processes’, J Mat Proc Tech, 150,
18–24.
dudziak k-u (1996), Prozessmodell zum Innenhochdruckumformen von hohlwellenfoermigen Werkstuecken, Duesseldorf, VDI.
flehmig t, bluehmel k w and kibben m (2001), ‘Investigation to bending boundaries
of circular tube cross sections and presentation of a new bending mandrel’, Proc
Int Conf Hydroforming, Stuttgart, 41–62.
flehmig t, gerlach j, kibben m and birkenstock a (2002), ‘Biegegrenzen von
Kreisrohren und Ermittlung der Verformungsreserven’, Proc 3rd Chemnitz Car
Body Colloquium, Chemnitz, 93–113.
franz w-d (1961), Das Kaltbiegen von Rohren, Berlin, Springer.
fuchizawa s (1984), ‘Influence of strain hardening exponent on the deformation of
thin-walled tube of finite length subjected to hydrostatic internal pressure’, Adv
Tech Plasticity I, 297–302.
fuchizawa s (1987), ‘Influence of plastic anisotropy on deformation of thin-walled
tubes in bulge forming’, Adv Tech Plasticity II, 727–732.
geckler j w (1928), ‘Plastisches Knicken der Wandung von Hohlzylindern und
einige andere Falterscheinungen an Schalen und Blechen, Z f angew Math u Mech,
8(5), 341–351.
hartl c (1995), Ein Beitrag zur Flexibilisierung der Innenhochdruckumformung,
Aachen, Shaker.
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Hydroforming for advanced manufacturing
hartl c (1999), ‘Theoretical fundamentals of hydroforming’, Proc Int Conf Hydroforming, Stuttgart, 23–36.
hartl c (2005), ‘Research and advances in fundamentals and industrial applications
of hydroforming’, J Mat Proc Tech, 167, 383–392.
jirathearanat s, hartl c and altan t (2004), ‘Hydroforming of Y-shapes – product
and process design using FEA simulation and experiments, J Mat Proc Tech, 146,
124–129.
khodayari g (2002), ‘How material influences bending for hydroforming’, The Tube
and Pipe Journal, January.
klaas f (1987), Aufweitstauchen von Rohren durch Innenhochdruckumformen,
Duesseldorf, VDI.
klaas f (1997), ‘Innovations in high-pressure hydroforming’, Proc 2nd Int Conf on
Innovations in Hydroforming Tech, Columbus OH, 1–31.
koc m (1999), Development of guidelines for tube hydroforming, doctoral dissertation, Columbus OH.
koc m (2004), ‘Advances in tube hydroforming – an enabling technology for lowmass vehicle manufacturing – material, lubrication, loading, simulation issues, and
alternatives’, Tsinghua Science and Tech, 9(5), 527–545.
kroeff a, beherns b-a, peters b-m and koch t (2002), ‘Untersuchungen von Stahl
und Bestimmung von Umformcharakteristiken fuer hydraulische Umformprozesse’, Proc 3rd Chemnitz Car Body Colloquium, Chemnitz, 69–80.
levy b s (1999), ‘Recommendations on the use of forming limit curves for hydroforming steel’ Proc Int Conf Hydroforming, Stuttgart, 77–96.
sauer w j, gotera a, robb f and huang p (1978), ‘Free bulge forming of tubes under
internal pressure and axial compression’, Proc 6th North American metal working
research conf, Gainsville, 228–235.
schmoekel, d, hessler, c and engel, b (1992), ‘Pressure control in hydraulic tube
forming’, Annals of the CIRP, 41(1), 311–314.
prahl u, kaluza w, kim, l and bleck w (2002), ‘Ermittlung von Grenzformspannungsdiagrammen (FLSD) zur Werkstoffcharakterisierung fuer die Blechumformung’, 2. Kolloquium Wirkmedien-Blechumformung, Dortmund, 95–100.
woo d m (1973), ‘Tube-Bulging under internal pressure and axial force’, J Eng Mat
Tech, (10), 219–223.
zschekel r (1977), Ein Beitrag zur Entwicklung des hydromechanischen Ausbauchens fuer die Fertigung T-foermiger Rohrverbindungsstuecke, doctoral dissertation, Dresden.
WPNL2204
4
Materials and their characterization
for hydroforming
CHRISTOPH HARTL,
Cologne University of Applied Sciences, Germany
4.1
Introduction
At present, in hydroforming production worldwide, steel alloys and aluminium alloys are the predominant materials used to manufacture semifinished products. Copper and brass alloys are typical materials for hydroformed
products for use in the piping and sanitary industries. In the majority of
cases, the alloys used correspond to materials which are used for conventional cold-forming processes such as deep drawing or massive forming.
However, with the objective of increasing formability for hydroforming
applications, certain modified alloys have been developed especially for this
technology. In addition, the demands made today by the automotive industry to reduce the weight of car components combined with an increase in
stiffness and strength are advancing the development of new alloys and
treatment strategies for steel as well as for aluminium materials. As an
example, an increasing trend of investigating the application of high-strength
steels and high-strength aluminium alloys with improved formability for
hydroformed components can be seen.
In general, all metal materials with sufficient formability are suitable for
semifinished products in hydroforming processes. A fine-grained structure
in combination with large measures of uniform elongation, elongation at
fracture and strain-hardening coefficient are advantageous to the feasible
expansion of the initial workpiece, achievable without the occurrence of
material instabilities. A distinctive work-hardening of the formed material
improves the strength of the final component. However, it also requires an
increase in the internal pressure pk for the final calibration of the hydroforming process (see section 3.4) which raises the pressure-dependent process
loads like tool closing force Fc and axial sealing force Fa.
Besides the material characteristics, the suitability of a semifinished
product for hydroformed parts is also influenced by its manufacturing
process and any preceding forming operations that might have been performed, e.g. preforming and bending. To regain material formability used
77
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Hydroforming for advanced manufacturing
up by any preceding forming operations, annealing processes can be applied
in theory to a large part of both the used metal materials and semifinished
products. However, due to the additional costs arising from this, it is commonly avoided in industrial mass production of hydroformed components.
Additionally, the application of annealing processes to bent or preformed
components can cause deterioration in the dimension qualities of these
workpieces, resulting in unstable conditions when the hydroforming process
is carried out.
This chapter deals with the application of common materials in tube
hydroforming as well as with future trends for new metal materials, hydroforming process strategies and designs of tubular semifinished products.
Emphasis is placed on the use of tubular materials for lightweight design.
Facts that must be considered when using semifinished products made of
low carbon steel, stainless steel, high-strength steels, magnesium and aluminium alloys are presented.
4.2
Steel materials
Industrial production and investigations using prototyping substantiate that
a wide range of steel alloys is suitable for use as semifinished products for
tubular hydroformed components. Steel alloys that have been used or tested
are ductile low-carbon steels, case-hardened steels, heat treatable steels,
ferritic and austenitic stainless steels as well as high-strength and ultra-highstrength steels. Low-carbon steels with yield strengths between 200 MPa
and 430 MPa are common and up to now have been the predominant materials used, particularly in automotive components (Kroeff, 2002). Examples
of these are axle components and engine cradles, as well as side and crossmembers for chassis components (Hartl, 2005). Within the context of one
of the first industrial investigations into the hydroforming of engine shafts,
steels suitable for case-hardening and heat-treatable steels have been shown
to be applicable for this technology (Ebbinghaus, 1991). Enhancements in
case-hardened steel alloys today provide steel tubes with up to 30% elongation at fracture in a soft condition and an ultimate tensile strength up to
1900 MPa in a hardened condition (Peters, 2005). Austenitic stainless steels
can be found in piping applications (Dohmann, 1997), and in the mass production of exhaust system components, e.g. the steel alloy AISI 309 as presented by Mertes (2000). Double-walled hydroformed tubes are also used
for exhaust system components (Hielscher, 2001). As for the widely used
austenitic stainless steel alloy AISI 304, besides being used in a wide range
of applications for hydroformed components like piping elements, handholds and roll bars, miscellaneous investigation results exist regarding
its formability in hydroforming processes, e.g. Eichhorn (1999), Hielscher
(2001), Jirathearanat (2004) and Hartl (2006). Investigations into the hydro-
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Materials and their characterization for hydroforming
79
4.1 Examples of tubular roll-formed products.
forming of side members made of dual-phase steels with ultimate tensile
strengths of up to 600 MPa as well as residual austenite grades with ultimate
tensile strengths of up to 700 MPa have been presented by Flehmig
(2003).
Tubular steel materials used for hydroforming applications are commonly produced as flat sheet material by continuous roll forming and longitudinal welding, as shown in Fig. 4.1. High-frequency welding is most often
used to close the roll-formed tubular cross-section. A circular cross-section
is generated by the roll-forming process for the production of tubes. Using
the appropriate roll-forming tools, profiles can be generated with crosssection shapes other than a circle. However, semifinished products with
predominantly circular cross-sections are currently used for hydroforming
production of steel components. The typical dimensions of hydroformed
steel tubes are outer diameters d0 between about 20 and 140 mm with ratios
of wall thickness to outer diameter t0/d0 between about 0.012 and 0.16.
In general, for tube specifications that are available on the market, a distinction should be made between tubes that are not annealed after being
cold formed by either roll forming or drawing, tubes drawn with a low strain
resulting from a preceding annealing process and tubes annealed after the
final cold-forming operation, as shown in Fig. 4.2. By following specific
drawing processes during tube manufacturing, the final tube diameter and/
or wall thickness as well as the increase in strength due to cold hardening
effects can be adjusted.
Drawn and non-annealed tubes commonly show a reduced formability
in hydroforming processes, depending on the characteristics of the steel
alloy used and the extent of the strain caused by the drawing operation.
Tubes which have been drawn with a low strain resulting from annealing
provide a cold formability within certain limits. The greatest cold formability is to be obtained by the use of tubes which have been annealed after
the final cold forming operation, e.g. roll forming or drawing. According to
(Kroeff, 2002), controlled drawing and subsequent annealing enable the
creation of a weld seam with a fine-grained structure and formability similar
to the initial sheet material.
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Hydroforming for advanced manufacturing
Roll forming
Welding
Annealing
Drawing
Annealing
Lining
Cutting
Final tube
Final tube
Final tube
Final tube
4.2 Typical process chains for steel-tube manufacturing.
The use of roll-formed and welded tubes in hydroforming processes
requires high-quality weld seams to avoid premature bursting of the workpiece. When positioning the weld seam for the final hydroformed component, it is necessary to avoid placing it within areas where excessive tensile
stresses (expansion) act on the component during the hydroforming
process.
As well as the material specifications which make expansion feasible, as
mentioned at the beginning of this chapter, the material formability is also
influenced by the anisotropy of the workpiece material. However, in contrast to sheet-forming technology, the influence of material anisotropy on
the forming behaviour in hydroforming processes is investigated to a comparatively smaller extent. According to Fuchizawa (1987), with an increasing anisotropic parameter in the longitudinal direction, thickness reduction
decreases and the critical expanding diameter increases for tube hydroforming without the assistance of an axial force but with tube ends permitted
to move axially. Tautenhahn (1996) published experimental results of investigations into anisotropy and textures of tubes dependent on the parameters
of the manufacturing process and the tube material. She recommended high
anisotropic parameters in the circumferential direction to lessen the change
in wall thickness during tube-bending processes. Groche (2005) presented
experimental results of bulge tests (see section 4.4) applied to tubes made
from a dual-phase steel where, besides other things, the influence of anisotropic behaviour became apparent. In this work, three kind of tubes manufactured from one identical coil were investigated: a conventional roll-formed
and high-frequency welded tube, a discontinuously formed and laser-welded
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Materials and their characterization for hydroforming
81
tube (see section 4.5) with the longitudinal axis parallel to the direction of
rolling of the initial sheet, and a discontinuously formed and laser-welded
tube with the longitudinal axis perpendicular to the direction of rolling.
According to the results of the bulge tests carried out with these workpieces,
the tubes made perpendicular to the direction of rolling showed a feasible
expansion about 60% higher than the conventional roll-formed tubes and
about 10% higher than the discontinuously welded tubes formed parallel
to the direction of rolling of the initial sheet.
4.3
Aluminium and magnesium alloys
A great variety of aluminium alloys are applied in hydroforming production or have been investigated. Results of fundamental investigations
into rotationally symmetrical hydroforming of aluminium tubes are to be
found, e.g. in Dohmann (1993) and Yuan (2001). Dohmann (1994) presented experimental results for hydroformed aluminium tubes with a displaced component section, similar to a crankshaft geometry. Various
examples of experimental investigations into the hydroforming of automotive space frame components are described, e.g. in Bartley (2000) and
Trubert (2002). A hydroformed aluminium T-piece with rectangular crosssections suitable for space frame nodal points is presented in Dohmann
(1997). Detailed information about hydroforming production with semifinished products made from aluminium alloys is to be found, e.g. in Schulze
(1999), Leppin (2001), Hoffmann (2001), Boehm (2004) and Schuster
(2005).
Currently, work-hardening aluminium 5000 alloys are used when the
priority is for a high amount of formability and corrosion resistance, whereas
precipitation-hardening aluminium 6000 alloys are applied for components
requiring a high strength. A distinction should be drawn between tubes
produced from flat-sheet material by continuous roll forming with longitudinal welding, and tubular semifinished products manufactured as extruded
profiles. In general, aluminium 5000 alloys are used for the production of
roll-formed tubes, whereas aluminium 6000 alloys are produced as extruded
profiles.
In Schulze (1999), the process chain of one of the first applications of an
aluminium 5000 alloy for the mass production hydroforming of rear axle
components is described in detail. For these components, roll-formed tubes
made of alloy 5083 with cross-section dimensions (outer diameter/wall
thickness) of 82 mm/4 mm, 89 mm/4 mm and 95 mm/3.5 mm were used in a
soft annealed condition for hydroforming. A current example of a massproduced hydroformed part made from a roll-formed aluminium 5154 alloy,
for the manufacture of cross-members in automotive body construction, is
given in Boehm (2004). Tubes with an outer diameter of 95 mm, a wall
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Hydroforming for advanced manufacturing
4.3 Examples of extruded aluminium profiles.
thickness of 1.4 mm and a length of 1800 mm are used for the production
of these components. The main process steps consist of bending the initial
tube, preforming, annealing, hydroforming the final shape and trimming the
component ends.
An advantage provided by extruded profiles is their design flexibility for
complex cross-sections with sharp corners, multiple hollows and flanges, as
shown in Fig. 4.3. However, the reduced formability of these semifinished
products has to be taken into consideration when designing each hydroforming component. Also, the comparatively small amount of plastic
deformation increases the effect of elastic spring back. With regard to the
hydroforming process itself, the sealing of complex-shaped profiles causes
more problems than the sealing of circular tubular profiles. Typical examples of extruded profiles for hydroforming applications, providing relatively
good formability, are aluminium 6061, 6063 and 6082 alloys, supplied in a
soft annealed condition that responds to age hardening by natural or artificial means (Bartley, 2000). The mechanical properties of the final component can be adjusted by altering the heat treatment of the hydroformed
extruded aluminium profiles. The use of an aluminium 6014 alloy for the
production of automotive space frame components is presented in Leppin
(2001). According to this, a suitable process chain for the manufacture of
these components consists of the hydroforming of an initial semifinished
product in T4 condition (solution heat-treated, quenched and naturally
aged) and the improvement of mechanical properties after the forming
operation by a T7 heat treatment (solution heat-treated, quenched, precipitation heat treatment and over-aged), see Fig. 4.4. Detailed investigation
results, dealing with such thermomechanical treatments of hydroformed
extruded aluminium profiles, are described in Kunz (1998). In general, aluminium alloys in a soft tempered condition offer good formability. However,
the properties of these alloys may vary with time because of natural aging.
With regard to the strength-improving elements Mg and Cu in aluminium
alloys, the fact that an increase of these elements leads to a deterioration
in welding qualities and corrosion resistance has to be taken into account
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Materials and their characterization for hydroforming
83
Extruded profile AA 6014 T4
Cutting
Bending
Hydroforming
Machining
Cleaning
Heat treatment T7
Surface treatment
Final component
4.4 Process chain for the example of hydroforming automotive body
components made of aluminium according to Leppin (2001).
(Ostermann, 1998). General rules for the design of extruded aluminium
profiles in hydroforming processes are given, e.g. by Hoffmann (2001),
developed within the field of mass production of automotive space frame
components.
Magnesium alloys offer great potential for weight reduction in vehicle
construction due to their high strength to weight ratio. However, the use of
these alloys in forming processes operating at room temperature is restricted
due to their hexagonal atomic structure. An improvement in formability
can be achieved by using increased temperatures above about 200 °C when
additional gliding planes become activated. Against this background, various
investigations have been carried out during the past few years into hydroforming semifinished products made from magnesium alloys using an elevated temperature. Sebastian (2000) mentions investigation results for the
free hydroforming of magnesium extruded tubes where an expansion in
diameter of 65% was achieved using a pressurizing medium with a temperature above 200 °C. Detailed fundamental investigations into the warm
hydroforming of extruded as well as roll-formed and welded magnesium
tubes made from the alloy AZ31 have been presented by Molitor (2006).
In Geiger (2003) and Stich (2004) the feasibility of warm hydroforming
magnesium sheet material is shown by an example of samples made from
the magnesium alloy AZ 31 B.
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Hydroforming for advanced manufacturing
4.4
Formability testing
Hydroforming process design, as well as the monitoring of semifinished
product quality in hydroforming production, requires suitable and reliable
methods to obtain material parameters which characterise the forming
behaviour of the initial workpieces.
Figure 4.5 shows the principles of formability testing methods used today
for hydroforming semifinished products. Currently, for tube hydroforming,
traditional material testing methods are the main ones used. However, the
suitability of these methods is often limited due to the fact that the typical
F
Expanded tube
pi
Mounting device
F
a Tensile test
a
Sealing punch
b Tube bulge test
Tube crosssection
l1
pi
Die cavity
l2
c Strain analysis
F
Expanding
cone
d Corner fill test
F
Penetrator
Tube wall
Tube
e Expansion test
F
Tube cross
section
F
g Tube tensile test
f Hardness test
F
Tube cross
section
Weld seam
h Flattening test
4.5 Principles of testing methods applied to hydroforming
semifinished products: a tensile test, b tube bulge test, c strain
analysis, d corner fill test, e expansion test, f hardness test, g tube
tensile test, h flattening test.
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Materials and their characterization for hydroforming
85
biaxial stress state of hydroforming processes is not, or only approximately,
reproduced.
The tensile test shown in Fig. 4.5a – a standardized uniaxial material test
– is the method most commonly used to characterize the forming behaviour
of the tubular material used. For roll-formed products, a distinction should
be drawn between the application of this test to the initial sheet material
before roll forming and its application to the roll-formed and welded workpieces. With a test to the initial sheet material, the changes in material
properties due to the manufacturing process remain unconsidered. As a
consequence, specimens for tensile tests are predominantly taken from the
finished tubular component in a longitudinal as well as in a tangential direction. However, this requires the specimen to be flattened before testing
which influences the material properties. Material properties obtained by
tensile tests are: yield stress sY, ultimate tensile strength sUTS, uniform strain
eU, elongation at fracture eZ, strain hardening coefficient n, elastic modulus
E and anisotropic parameters r. In addition, the tensile test enables the yield
curves of the tube material to be determined within certain limits.
With the objective of improving the characterization of tubes for hydroforming applications, several investigations have been carried out into
tube expansion tests, working with an inner pressurization of the tested
tube which is clamped at its ends, see Fig. 4.5b. This bulge test enables
the bursting pressure pb, the pressure-dependent expansion diameter d(pi)
and the maximal achievable expansion diameter dr under biaxial tensile
stress state to be determined. In addition, Altan (1999) presented a strategy
using the bulge test to determine the material properties of tubes, assuming volume constancy during forming, a biaxial membrane stress state, a
parabolic expanded geometry, and isotropic, purely plastic material behaviour. According to Groche (2005), the bulge test enables the yield stress
sY of the tube material to be determined by applying the following
relationship:
σ Y = 3 pi
(dm 0 + 2dr )2 − t0 dm 0
4t0 dm 0
[4.1]
with the internal pressure pi, the mean diameter dm0 of the initial tube, the
maximal currently expanded tube diameter dr and the initial tube wall
thickness t0. Equation [4.1] applies for the condition where the tube ends
are fixed during the expansion process.
Detailed investigations into the application of the bulge test for characterizing the forming behaviour of tubes have been carried out by Hielscher
(2001). According to investigation results described by Hielscher (2001) as
well as by Groche (2005), pressurization of the specimen with a controlled
volume flow is recommended as opposed to a pressure-controlled expansion. This is because the required pressure to expand a tube decreases above
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a certain expanded diameter and impedes exact pressure control. When
applying the bulge test, consideration has to be given to the fact that the
ratio of expanded tube length ld to tube diameter d0 influences the required
pressure to expand a tubular specimen if the ratio ld/d0 is below a certain
limit (Klaas, 1987). As an example, Hielscher (2001) determined that, for
the expansion of stainless-steel tubes with an outer diameter of 60 mm and
a wall thickness of 1 mm, a significant change in pressurization was required
for a ld/d0 ratio below about 2. The first commercial devices to conduct bulge
tests are now available on the market (Boehm, 2003).
A common method for strain analyses in hydroformed components consists of applying circular or quadratic grids to the surface of the initial
semifinished product. The measured distortion of the individual grid elements on the hydroformed workpiece enables local strains to be determined. For the example of circular grids in Fig. 4.5c, the sizes of the principal
axes of the ellipsis l1 and l2, resulting from the individual distorted circular
grid, are used to determine the strains:
l
ε 1 = ln 1
a
[4.2]
l
ε 2 = ln 2
a
[4.3]
and
with an initial grid diameter a. By considering the volume constancy of the
tube material, the local strain in the direction of the tube wall thickness can
be estimated by:
et = −e1 − e2
[4.4]
A comparison of the analysed strains with the Forming Limit Curve of
the respective tube material provides an assessment of the hydroforming
process. However, it must be taken into account that this strain analysis,
strictly speaking, is applicable solely for forming processes with constant
ratios of strains during the forming operation. Detailed recommendations
as well as further developments regarding the use of strain analysis in
hydroforming processes are to be found e.g. in Levy (1999), Prahl (2002)
and Green (2003). Today, optical systems are also available to determine
the resulting strains from the distorted grids produced during the forming
process. An example of the application of such a system to tube bulge tests
is presented in Groche (2005).
The corner-fill test, shown in Fig. 4.5d, is a common experimental test to
investigate the influence of various lubricants on the hydroforming process
(Khodayari, 2002). In this test a straight cylindrical tube is internally pressurized inside a die with a square cross-section that is the same size as the
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Materials and their characterization for hydroforming
87
initial outside diameter of the tube. The tube is constrained by the hydroforming die and merely expands into the corners of the die. As the outer
surface of the tube is almost completely in contact with the surrounding
die, lubricants play a significant role in the forming behaviour. Duroux
(2001) enhanced this test by additionally applying it for investigations into
the influence of material specifications on local necking of the tube wall.
Further and predominantly standardized tests apply a mechanical expansion to tube sections or tube ends. An example from this group of testing
methods is the cone test where the tube end is expanded by a conical punch
until fracture occurs, as shown in Fig. 4.5e. These tests are the main methods
for determining formability, especially for comparing different batches of
delivered tubular material. In addition, failures can be detected, for example
at the tube surface or within the weld seam. When applying such test
methods, the fact that variations in friction conditions, or unequal preparedsurface roughness at the tube end face, may influence the initiation of fracture in the expanded tube section must be taken into consideration.
Standardized hardness tests are suitable for obtaining information about
the hardening behaviour of the tube material due to the forming process,
see Fig. 4.5f. Depending on the tube wall thickness and material strength,
Brinell, Vickers or Rockwell test methods are commonly used. These tests
are applied to cut cross-sections in longitudinal as well as perpendicular
directions to the workpiece axis. The measured hardness numbers allow a
conclusion to be drawn about the increase in strength of the formed
material.
Further miscellaneous testing methods exist, partly derived from standardised tests like the tube tensile test shown in Fig. 4.5g or designed for
checking the weld seam quality as shown in Fig. 4.5h. In general, besides
the material parameters and forming characteristics determined with the
testing methods mentioned here, specifications like the dimensions and
tolerances, and surface quality of the semifinished product used are also
important.
4.5
Future trends
The objective of increasing the number of feasible component geometries
by economic hydroforming as well as the aim of meeting further demands
for weight reduction of hydroformed automotive components are substantial driving factors for current research work into this technology being
carried out by research facilities, hydroforming technology suppliers and
users. Besides the development of steel and aluminium alloys mentioned
previously, this research focuses especially on new strategies in manufacturing semifinished products as well as on the use of specifically applied temperatures within the hydroforming process.
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Weld seam
Modified thickness
and/or properties
One-piece
cylindrical tube
Tailored blank
and tailored tube
Patchwork blank
and patchwork tube
One-piece
conical tube
4.6 Examples of laser-welded tailored semifinished products.
Innovative semifinished product designers are investigating the use of
components that provide optimized distribution of wall thickness and/or
material strength. Alternative manufacturing methods for the initial parts
have been developed like laser welding of tailored blanks and tubes, as
shown in Fig. 4.6, which allow the creation of variations in thickness, strength
and shape, in a longitudinal direction as well, e.g. Hartl (2002), Lamprecht
(2003), Flehmig (2003) and Weil (2003). In addition, initial blanks and tubes
made from flexible rolled sheets with different thickness are now available
on the market (Pohl, 2005). Flexible and discontinuous tube-forming processes are used to create the individual tube from tailor-welded blanks or
flexible rolled sheets as described e.g. by Flehmig (2003) and Weil (2003).
For this kind of tube, the edges must be carefully prepared before the laser
welding operation to ensure a reliable hydroforming process.
Investigations carried out by Kreis (2004) looked at the reduction of
process chains for the manufacture of complex hydroformed parts, based
on an integrated manufacturing cell which allows a combination of double
sheet and tube hydroforming to be carried out with clinching, welding and
trimming by lasers.
Several procedures and process variants have been developed regarding
the use of controlled temperature application to enhance formability in
hydroforming processes. A distinction can be drawn between methods that
apply heat energy to the initial part before hydroforming and processes
applying heat during the forming operation. (Vollertsen, 1999) showed
by numerical simulations that local heating of a precipitation-hardened
extruded aluminium profile before hydroforming improves the forming
result. A T-piece with rectangular cross-sections was used as a hydroformed
workpiece example where the area opposite the branch to be formed was
heated locally. The reduction in yield strength caused by the local heating
resulted in an improved material flow during the hydroforming process at
room temperature.
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Materials and their characterization for hydroforming
Elongation at fracture (%)
120
100
89
ALMg3Mn
300°C
80
250°C
60
200°C
40
150°C
20
0.01
0.1
20°C
1
10
Strain rate (1/s)
4.7 Warm tensile tests of aluminium AlMg3Mn according to Keller
(2004).
In particular, it is already known that aluminium alloys and magnesium
alloys shows an improved formability when hydroformed at specific temperatures by so called warm hydroforming. On the one hand, these alloys
offer great potential for weight reduction in vehicle construction due to
their high strength to weight ratio, but on the other hand the use of these
metals is restricted due to their low formability at room temperature in
comparison with conventional steels. As an example, Fig. 4.7 represents
experimental results from warm tensile tests of an aluminium alloy, showing
the increase of attainable elongation at fracture depending on temperature
and strain rate (Keller, 2004). Therefore, more and more investigations are
being carried out into increasing the formability of sheet and tube hydroforming by the use of elevated temperatures, e.g. Geiger (2003), Stich (2004),
Hartl (2005), Schuster (2005), Kren (2005) and Molitor (2006). Pressurizing
media-heated oil as well as gas in combination with heated tools are used.
Oil – compared with gas – offers the advantages of lower compressibility,
higher thermal capacity and higher applicable pressurization, although the
applicable temperature is restricted due to the inflammability of this
medium. Hence, non-flammable gas enables higher temperatures to be
applied, e.g. transmitted by heated tools.
Besides the warm hydroforming of aluminium and magnesium alloys,
investigations into the warm hydroforming of steel are also being carried
out with the objective of increasing material formability (Pfaffmann, 2003).
For this, ceramic tools with an integrated inductive heating element and
pressurization by gas are being used for the semifinished products.
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Within the past few years, methods to generate metal foams from aluminium or zinc alloys as well as strategies to apply these foams to hydroformed components have been developed further, e.g. Neugebauer (1998).
Methods to produce hydroformed components containing metal foams
include using cast or powder-metallurgical-produced foamed-up workpieces
which are inserted into the hydroformed part or using powder-metallurgical-produced parts which are foamed up within the hydroformed component by the application of heat energy (Neugebauer, 1998). The main
advantages of foamed hydroformed components in automotive applications are improved shock absorption, increased stiffness with reduced
weight and efficient sound absorption. The characteristics of the foam can
be adjusted with respect to composition and density.
4.6
References
altan t, koç, m, aue-u-lan y and tibari k (1999), ‘Formability and design issues in
tube hydroforming’, Proc Int Conf Hydroforming, Stuttgart, 105–122.
bartley g and evert b (2000), ‘Hydroforming of aluminum extrusions for automotive applications’, Light Metal Age, Aug, 24–37.
boehm a and prier m (2003), ‘Innenhochdruck-Umformen – ein Verfahren für die
Zukunft’, Proc EFB-Fortbildungspraktikum, 2003, Magdeburg, 171–178.
boehm a (2004), ‘Development and mass production of the BMW 5 and 6 series
lightweight aluminium front end’, Proc North Am Hydroforming Conf, Waterloo,
Ontario.
dohmann f and hartl c (1993), ‘Moeglichkeiten der Innenhochdruckumformung
unter besonderer Beachtung des Formens von Strangpressprofilen. Proc Conf
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dohmann f and hartl c (1994), ‘Liquid-bulge-forming as a flexible production
method’, J Mat Proc Tech, 45, 377–382.
dohmann f and hartl c (1997), ‘Tube hydroforming – research and practical application’, J Mat Proc Tech, 71, 174–186.
duroux o and tondo s (2001), ‘Influence of material properties on necking in corner
fill test’, Proc Int Conf Hydroforming, Stuttgart, 477–484.
ebinghaus a (1991), ‘Praezisions-Werkstuecke in Leichtbausweise hergestellt durch
Innenhochdruck-Umformen’, Metallumformtechnik, 1D, 15–19.
eichhorn a (1999), ‘Innovative developments concerning hydroforming of tubes’,
Proc Int Conf Hydroforming, Stuttgart, 391–406.
flehmig t and schwarz s (2003), ‘Hydroforming complex hollow sections’, Steel
Grips, 1(6), 408–412.
fuchizawa s (1987), ‘Influence of plastic anisotropy on deformation of thin-walled
tubes in bulge forming’, Adv Tech Plasticity, II, 727–732.
geiger m, celeghini m, haldenwanger g and prier m (2003), ‘Sheet and tube
hydroforming at elevated temperature’, Proc Int Conf Hydroforming, Stuttgart,
259–278.
green d e (2003), ‘Experimental determination of tube forming limits’, Proc Int
Conf Hydroforming, Stuttgart, 299–314.
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Materials and their characterization for hydroforming
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groche p and breitenbach g (2005), ‘Influence of tube manufacturing processes on
hydroforming’, Proc Int Conf Hydroforming, Stuttgart, 219–240.
hartl c, luecke h-u and boehm a (2002), ‘Produktivitaetssteigerung beim Hochdruckumformen – Massnahmen und Strategien’. Proc 3rd Chemnitz Car Body
Colloquium, Chemnitz, 169–182.
hartl c (2005), ‘Research and advances in fundamentals and industrial applications
of hydroforming’, J Mat Proc Tech, 167, 383–392.
hartl c, lungerhausen j, biedermann h and conzen j (2006), ‘Study of hydroforming processes for the production of micro-components’, Proc 1st Jubilee Sci
Conf Manufacturing Engineering in Time of Information Society, Gdansk, 137–
140.
hielscher c (2001), ‘Tube testing for the production of complex hydroforming
parts’, Proc Int Conf Hydroforming, Stuttgart, 63–84.
hoffmann a and birkert a (2001), ‘Design guidelines for hydroformed structural
components of aluminium’, Proc Int Conf Hydroforming, Stuttgart, 323–338.
jirathearanat s, hartl c and altan t (2004), ‘Hydroforming of Y-shapes – product
and process design using FEA simulation and experiments’, J Mat Proc Tech, 146,
124–129.
keller s et al. (2004), ‘Warmumformen von Aluminium-Werkstoffen’, Proc of the
EFB-Kolloquium Fertigungsverfahren für moderne Blechwerkstoffe, Fellbach.
khodayari g, reid j and garnett m (2002), ‘Analyzing tubes, lubes, dies, and friction
using tribology to evaluate lubricant-material combinations in hydroforming’, The
Tube and Pipe Journal, 13(6), 72–74.
klaas f (1987), Aufweitstauchen von Rohren durch Innenhochdruckumformen,
Duesseldorf, VDI.
kreis o, celeghini m, geiger m and merklein m (2004), ‘Integrated manufacturing
by hydroforming, laser welding and cutting of complex hollow parts’, Proc 4th
Conf on Laser Assisted Net Shape Engineering, Erlangen, 1165–1174.
kren l a (2005), ‘Warming up to Al and Mg forming’, Metalforming, March, 28–
30.
kroeff a, beherns b-a, peters b-m and koch t (2002), ‘Untersuchungen von Stahl
und Bestimmung von Umformcharakteristiken fuer hydraulische Umformprozesse’, Proc 3rd Chemnitz Car Body Colloquium, Chemnitz, 69–80.
kunz c (1998), Neue Wege bei der Innenhochdruckumformung von stranggepressten
Aluminiumprofilen, doctoral dissertation, Zuerich.
lamprecht k and geiger m (2003), ‘Umformung tiefgezogener tailored und patchwork blanks’, Proc Colloquium Sheet Metal Hydronforming, Stuttgart, 65–74.
leppin a and boucke b (2001), ‘Hydroforming of aluminium extrusions’, Proc Int
Conf Hydroforming, Stuttgart, 311–322.
levy b s (1999), ‘Recommendations on the use of forming limit curves for hydroforming steel’, Proc Int Conf Hydroforming, Stuttgart, 77–96.
mertes p and schroeder m (2000), ‘Anwendung des Innenhochdruckumformens in
der Automobilindustrie’, Proc EndForm2000, Paderborn, 47–66.
molitor m, eichhorn a and bietke d (2006), Innenhochdruckumformen von Magnesiumrohren bei Erwaermung der Ausgangsteile über das Wirkmedium, Hannover, EFB.
neugebauer r, bräunlich h and wagner u (1998), ‘Lightweight products with metal
foam – properties and methods of processing’, Proc Material Research Society
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ostermann f (1998), Anwendungstechnologie Aluminium, Berlin, Springer.
peters b-m and saeuberlich t (2005), ‘Economic production of high-strength tubes
– a possibility to cost reduction of hydroforming processes’, Proc Int Conf Hydroforming, Stuttgart, 89–100.
pfaffmann g and dykstra b (2003), ‘The next generation process for manufacturing
vehical structural components’, Proc Int Conf Recent Development on Tube and
Sheet Hydroforming, Columbus, Ohio.
pohl s and hauger a (2005), ‘Tailor rolled tubes – weight and function optimised
tubes for hydroforming’, Proc Int Conf Hydroforming, Stuttgart, 29–36.
prahl u, kaluza w, kim l and bleck w (2002), ‘Ermittlung von Grenzformspannungsdiagrammen (FLSD) zur Werkstoffcharakterisierung fuer die Blechumformung’, 2. Kolloquium Wirkmedien-Blechumformung, Dortmund, 95–100.
schulze t (1999), ‘Seam-welded aluminium tubes for lightweight components’,
Tube International, 181(11), 497–503.
schuster c, loretz c, klaas f and seifert m (2005), ‘Potentials and limits with
hydroforming of aluminium alloys’, Proc Int Conf Hydroforming, Stuttgart,
113–136.
sebastian w and schumann s (2000), ‘Einsatzpotential von Mg-Profilen im Fahrzeugbau’, Proc EndForm2000, Paderborn, 21–30.
stich a (2004), ‘Hydroforming of Aluminium and Magnesium Sheets at evaluated
temperature for automotive lightweight construction’, in VDI Berichte, No 1833,
Duesseldorf, VDI, 71–78.
tautenhahn h (1996), ‘Anisotropie und Textur von Rohren’, Blech Rohre Profile,
43(5), 234–239.
trubert f (2002), Herstellung von Automobil-Spaceframeteilen aus Aluminium in
Innenhochdruck-Umformtechnik (IHU), doctoral dissertation, Wien.
vollertsen f, prange t and sander m (1999), ‘Hydroforming: needs, developments
and perspectives’, Adv Tech Plasticity, II, 1197–1210.
weil w (2003), ‘From the blank to laser-welded hydroformed tube’, Proc Int Conf
Hydroforming, Stuttgart, 275–286.
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of aluminium tube hydroforming’, Proc Int Conf Hydroforming, Stuttgart,
334–350.
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Formability analysis for tubular
hydroformed parts
D. E. GREEN, University of Windsor, Canada
5.1
Introduction
In the last 20 years, tube hydroforming has emerged as an important manufacturing process in the automotive industry. Indeed, tube hydroforming
offers the potential for both weight savings and cost savings for structural
components such as frame rails, engine cradles, radiator support beams,
instrument panel beams etc. Significant weight savings can be obtained by
replacing a sub-assembly made from stamped and welded parts with a
tubular structure without compromising its bending and torsional stiffness.
Moreover, the production costs for such structural sub-assemblies can also
be decreased when a number of stamped parts can be consolidated into a
single tubular component. Furthermore, improvements in hydroforming
press technology have led to a decrease in cycle time and a corresponding
increase in the volume of hydroformed structural components.
The production of high-quality components is dependent on correctly
assessing the severity of the entire manufacturing process. A typical manufacturing sequence for a tubular component consists of tube bending, tube
crushing or pre-forming, and then hydroforming to its final shape. During
each successive forming operation the tube wall is deformed and the forming
severity increases. In order to consistently produce quality components, it
is essential that the severity of each forming process remain below the
forming limit of the tube.
This chapter presents some important results of research on the analysis
of forming severity as it applies to tubular hydroformed parts. First, an
overview will be given of the formability of metal alloys in general, and
metal tubes in particular. The next section will be devoted to experimentally
measuring the formability of as-received metal tubes. A third section will
focus on forming limits and the various criteria that have been proposed to
predict the onset of failure. Finally, the implementation of formability analysis is discussed for both numerical simulations of tube hydroforming and
for actual hydroformed parts.
93
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5.2
Tube formability
5.2.1 Formability of metal alloys
The ability of metallic materials to undergo significant plastic deformation
is a property called ductility or formability. Many manufacturing processes
take advantage of the formability of metals to form or shape as-received
material into a semi-finished product that has both functionality and aesthetic appeal. Moreover, as metal is plastically deformed, it generally exhibits an accompanying increase in strength: this is known as work hardening.
This ability of metals to work harden is very desirable since they can be
deformed to a certain extent without failure, and the finished product has
greater strength than the material from which it was fabricated.
The formability of metal alloys depends on their chemical composition
as well as their thermo-mechanical processing history. For instance, mild,
low-carbon steel is generally more formable at room temperature than
aluminium or magnesium alloys, and cold worked metals are more formable after they undergo an annealing process. Therefore, it will be possible
to safely manufacture a metallic component with its given design and geometric features, provided the selected alloy possesses a sufficient level of
formability.
When a metal component is cold formed, different regions of the part
are deformed to different levels of strain, depending on its geometry and
its particular forming process. This signifies that each location in the part
work hardens by a different amount. Additionally, in a multi-stage forming
process the amount of work hardening varies not only from one location
to the next but also from one forming stage to the next. In order to ensure
that a selected alloy possesses sufficient formability, it is necessary to evaluate the forming severity of the entire manufacturing process.
5.2.2 Formability of tubes
Hydroformed tubular components are generally formed in a multi-stage
forming process: for instance, a coil of sheet may be roll-formed and welded
into a tubular structure, sized and cut to the desired dimensions, then bent
in a number of places and possibly crushed before being hydroformed to
the final shape. At each stage in this manufacturing process, the tube wall
is progressively work hardened by various amounts. In order to evaluate
the severity of the entire forming process it is necessary to monitor the loss
of formability that takes place at each stage. In this section, each stage of
the production of a typical hydroformed component will be considered with
a view to understanding how the overall formability of the component is
affected.
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Influence of tube making
Tubes that are used to make structural components are sometimes produced in batches. For instance, press-forming consists of bending a flat blank
in a type of press-brake first into a ‘U’-shaped, then into an ‘O’-shaped crosssection. This batch process is used to produce tubular blanks. Tubular blanks
offer many advantages over conventional roll-formed tubes such as the
possibility of higher diameter-to-thickness ratios (D/T < 250) compared
with conventional tubes (D/T < 75), conical tubes and tailor welded tubes.
The production of tubular blanks is also said to cause less plastic deformation and therefore a greater proportion of the formability available in the
flat sheet remains in the tube for subsequent forming operations. However,
batch processes are generally slower and more expensive than continuous
processes, and therefore better suited to low-volume production.
Continuous tube-making processes include roll-forming, extrusion and
seamless tubing. Tubular parts can be extruded with closed cross-sections
of any shape, making it a cost-effective and versatile process. A definite
advantage of extruded cross-sections is that very small corner radii and wall
angles can easily be extruded whereas extremely high internal pressures
would be required during hydroforming to create the same cross-section
from a roll-formed tube. However, one of the main disadvantages of
extruded profiles is the non-uniformity of the wall thickness around the
cross-section and down the length of the extrusion. For this reason, hydroformed extrusions typically require a high-pressure calibration at the end
of the hydroforming cycle to ensure that the part conforms to the required
dimensions and tolerances.
Many components are also made from seamless tubing, however, by far
the most common method of producing large quantities of tubes for hydroforming applications is by roll-forming in a continuous tube-mill: a coil of
sheet – the skelp – is gradually bent in the transverse direction as it moves
through successive sets of forming rolls until the cross-section is practically
closed. The two edges are then welded together either by electrical resistance welding (ERW), high-frequency (HF) induction welding or laser
welding. Following the welding and the removal of the weld flash, tubing is
typically formed through sizing rolls to ensure that the inner and outer
diameters of the finished tube are within specified tolerances.
The weld seam in a tube leads to an obvious non-homogeneity in material
properties, and the formability of the weld seam and its heat-affected zone
is usually lower than that of the tube. Indeed, micro-hardness tests taken
across a typical weld seam show that hardness may decrease slightly in the
heat-affected zone compared with the tube wall, but will increase significantly in the weld zone. However, welding technology is well understood
and generally well monitored during tube production, so that tube failures
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rarely occur on or near the weld line. Nevertheless, it is customary to locate
the weld seam in one of the least deformed regions of a hydroformed
part.
During tube making, there is an inevitable loss of formability as a result
of the roll-forming and sizing operations. However, it is essential to minimize work hardening at this stage in order to maximize the remaining
formability that will be required for subsequent processes such as tube
bending and hydroforming. The Hydroforming Materials and Lubricants
task force of the Auto/Steel Partnership (A/SP) carried out a study
(Zimerman, 2003) in order to assess how much of the available formability
is used up during tube making. This was done by comparing the stress–strain
curves obtained from specimens before and after tube making. Tensile tests
were carried out on specimens taken from 76 mm outer diameter (od) tubes
that were made from various steel grades. The tensile tests were carried out
in both the longitudinal and transverse directions and for different angular
positions around the tube circumference. Further tensile tests were also
conducted on the flat stock from which the tubes were made. The loss of
formability due to tube making was then estimated by shifting the origin
of the true stress-strain curve for the tube until it coincided with the true
stress–strain curve for the sheet and assuming the tube-making strain is
equivalent to the added strain. It was shown that the range of tube-making
strains varied approximately from 0.01 to 0.08 for all materials and orientations investigated. Figure 5.1 illustrates how the tube making strains were
estimated for high-strength low-alloy (HSLA) tubes. It was also shown that
the estimated tube-making strains were consistently greater in the longitudinal direction than in the transverse direction of the tube: for higher
strength steel tubes (HSLA and dual phase) the difference was significant
(1.5 to 3 times higher in the longitudinal direction), whereas for mild steel
grades the difference was only moderate. Finally, tube-making strains were
always greater at the side that is opposite the weld seam.
The method described above only provides an approximate assessment
of the amount of formability that is used to form a flat sheet into a tube.
The actual loss of formability is tube-mill dependent and also varies from
one set-up to the next. Since the details of the tube-making process are not
normally provided by the manufacturer, it is preferable to determine the
available formability directly from the tube rather than attempting to estimate it from the sheet mechanical properties. Tube mechanical properties
can be determined in a number of ways that will be reviewed in a later
section of this chapter, but suffice it to say, tubes produced in a typical tube
mill exhibit significant variations in mechanical properties around the circumference of the tube. Therefore, the amount of work hardening that is
available for subsequent tube forming operations also varies with the
angular position from the weld seam.
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97
600
Tube
True stress (MPa)
500
400
Sheet
Tube + 0.076 shift
300
200
0.076 shift
100
0
0.000
0.025
0.050
0.075
0.100
0.125
0.150
0.175
True strain
5.1 Graphical method to estimate the plastic strains generated
during tube making for 76-mm-diameter HSLA tubes with a
2.0 mm wall thickness at 180° from the weld seam (adapted from
Zimerman, 2003).
Influence of tube bending
One of the most consequential tube-forming operations is tube bending.
Tube bending is generally carried out as a separate process before hydroforming and enables the tube to fit into the hydroforming die. Depending
on the dimensions of the tube, there are a number of industrial tube bending
processes that can be used. Usually, 3- and 4-roll profile bending is used to
bend large parts with complex cross-sections. Guided profile bending consists of moving a pair of rolls along the length of a tube to bend it to the
shape of a rigid guide profile. Again, roll bending consists of moving a single
roll along the length of a tube to bend it around a circular bend die.
However, the most common tube bending process for industrial hydroforming applications is rotary draw bending. This consists of gripping the end of
a tube between a clamp die and a circular bend die and drawing it around
the bend die. A wiper die located right behind the bend die prevents the
tube from wrinkling on the inside of the bend, and a mandrel is usually
placed inside the tube to prevent it from collapsing during the bend. In
some cases, tube bending can be done in the hydroforming die: the tube is
bent into shape as the die closes.
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During any tube-bending operation, the tube wall on the outside of the
bend is subject to tension forces that cause it to stretch and decrease in
thickness. Conversely, the tube wall on the inside of the bend undergoes
compressive forces that cause an increase in wall thickness. In the case of
rotary draw bending, a tensile deformation is also superimposed on the
bending deformation. The tensile strains and the compressive bending
strains cancel each other out at the neutral axis, but as a consequence of
the additional tensile deformation, this region of zero strain actually shifts
away from the tube centreline toward the inside of the bend. Moreover, the
major strain on the outside of the bend is oriented parallel with the axial
direction of the tube, whereas on the inside of the bend, the direction of
major strain coincides with the hoop direction. And for locations around
the tube circumference between these two extremes, the direction of major
strain gradually rotates between its orientation at the extrados and that at
the intrados. It can be seen therefore that every angular position on the wall
of a bent tube follows a different strain path, which leads to a continuously
changing strain distribution.
As a consequence of the deformations generated during tube bending,
the level of work hardening in the tube wall increases with the distance to
the neutral axis. In other words, the region of least work hardening approximately coincides with the tube neutral axis. Stevenson et al. (2003) demonstrated this by pressurizing tubes that had been bent to 90° in a rotary-draw
bender. These researchers found that the increase in internal pressure in
the bent tube invariably led to extremely localized deformation and failure
at the neutral axis.
The severity of the bending operation also determines the magnitude of
the strains in the tube wall. For example, a 2D bend radius (i.e. the centreline bend radius is equal to two times the outer diameter of the tube)
leads to approximately 20% axial strain on the outside of the bend, a 1.5D
bend radius leads to 25% axial strain and a 1D bend radius leads to 33%
axial strain. Since work hardening increases with bending strains, it is important to design tubular parts with bends that are not overly severe so that
sufficient formability is left in the part for subsequent forming operations.
The recommended minimum bend radius is 1.5D.
The largest amount of plastic strain that is accumulated in a typical
tubular hydroformed component during a single forming stage occurs
during the bending operation. In order to minimize the loss of formability,
rotary draw bending is often accompanied by ‘boost assist’: a compressive
force that is generated by the friction between a pressure die and the tube
and that pushes the tube into the forming zone from the rear. Bardelcik
and Worswick (2006) have shown that bending boost does in fact improve
the strain and thickness distributions in the tube. Indeed, the reduction in
wall thickness on the outside of a bent tube is less severe when a higher
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level of boost is applied, and the combination of an improved thickness
distribution and a reduced level of work hardening contribute to delay the
onset of failure in subsequent hydroforming.
Influence of hydroforming
The last influence on tube formability we will consider is the stress distribution during the hydroforming process itself. When a thin-walled tube is
pressurized and expanded in free space, it is reasonable to assume the wall
is subject to a state of plane stress: that is, the through-thickness stress
component can be neglected. However, when the tube wall is in contact
with a hydroforming die, the through-thickness stress component is equivalent to the internal pressure and the plane-stress approximation becomes
invalid.
Hydroforming often requires internal pressures of the same order of
magnitude as the tube yield stress in order to fully form the part in the die.
Consequently, the section of the tube wall that is in contact with the die
cannot easily deform because of the high contact pressure and the friction
forces that restrain its movement. There is therefore little risk of failure in
sections of a hydroformed part that are in contact with the die. But necking
or splitting frequently occurs at the transition between a section of the tube
wall in contact with the die and an adjacent section that is unsupported. In
this transition zone the tube wall generally undergoes severe gradients in
through-thickness stress, in friction conditions and in curvature (e.g. the
radius of curvature increases to infinity in a flat section of the die wall). It
is important, therefore, that a forming analysis account for this particular
stress state that is so prevalent in tubular hydroforming applications.
As can be seen from this section, tubular hydroformed components are
generally subjected to a complex deformation history. An analysis of forming
severity for bent and hydroformed tubes must accurately account for the
non-uniform distribution of work hardening at each stage of the tube
forming process. Failure to do so could lead to faulty part design or a hydroforming production process with a high scrap rate.
5.3
Measuring tube formability
As discussed previously, tube mechanical properties cannot accurately be
determined from the properties of the sheet, since they depend on the type
of tube mill and on its particular set-up. It is therefore important to determine the formability of the as-received tube before its utilization in a
multi-stage tube-forming process such as bending and hydroforming. There
are a number of simple industrial tests that provide a qualitative or comparative measure of tube formability. For instance, a tube can be flattened
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to determine the bendability of the tube wall, particularly in the area of
the weld seam. A tube with good formability can be fully flattened without
fracture occurring in the wall. Another common test is the cone test or flare
test: this test consists of forcing a cone into the end of a tube so as to cause
an expansion of the tube circumference. The cone continues to penetrate
into the tube until the stretched edge begins to split. The relative depth of
penetration of the cone or the relative expansion of the tube wall provides
a measure of tube formability. The ring tensile test is yet another test that
consists of inserting two finger-like bars into the end of a tube and pulling
them apart in a tensile mode. As the tensile force increases, the tube wall
stretches. The maximum load or maximum deformation before fracture
provides a measure of the tube’s stretchability. These tests are routinely
carried out in the production plant to verify the formability of in-coming
tubes. However, they only consider specific modes of deformation and the
data is only useful as it applies to a given tube size. A more universal and
fundamental description of tube formability is required.
5.3.1 The tensile test
One of the simplest ways of obtaining a quantitative measure of tube formability is to carry out tensile tests on specimens taken from the tube wall.
The stress–strain curve and the mechanical properties (yield stress, tensile
stress, uniform and total elongation, strain hardening exponent, anisotropy
coefficient) obtained from a tensile test describe the formability of the tube
material. Specimens can be taken from the tube such that their length is
parallel with the tube axis: these are referred to as longitudinal specimens.
It is common practice to take longitudinal specimens at certain positions
around the circumference of a tube such as 90°, 180° and 270° from the
weld seam. More meticulous research work may require longitudinal specimens to be taken at positions every 45° from the weld seam as is shown in
Fig. 5.2. Tensile specimens are generally machined according to the ASTM
standard; however, when the tube size is small, sub-size ASTM specimens
can also be used. In an in-depth investigation conducted by the A/SP,
Zimerman (2003) showed that tensile tests done on sub-size ASTM tension
specimens and on standard full-size ASTM tension specimens yielded
similar stress–strain behaviour.
As a result of the tube-making process, the wall thickness and the
mechanical properties – and therefore the formability – are not uniform
around the circumference of a tube. In fact, tubing produced on a typical
roll-forming tube mill shows a significant degree of anisotropy. For example,
Fig. 5.3 represents the longitudinal stress–strain curves at various positions
around the circumference of 76 mm-diameter aluminium-killed-drawquality (AKDQ) steel tubes with a 2.0 mm wall thickness. This figure shows
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Formability analysis for tubular hydroformed parts
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45
0
315
90
270
225
135
180
5.2 Schematic illustrating how tensile specimens are taken from
a tube in the longitudinal direction (courtesy of the Auto/Steel
Partnership).
350
True stress ( Mpa)
300
250
45°
90°
135°
180°
200
150
0.00
0.01
0.02
0.03
0.04
0.05
True plastic strain
5.3 True stress – true plastic strain curves for longitudinal specimens
taken from 76-mm-diameter AKDQ tubes (adapted from Zimerman,
2003).
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that the greatest amount of work hardening takes place at 180° from the
weld seam (i.e. at the 6 o’clock position). Furthermore, the distribution of
mechanical properties around the tube circumference shows distinctive and
consistent patterns. Figures 5.4 and 5.5 represent the variations in yield
stress and n-value, respectively, in the longitudinal direction at various positions around the circumference of the same AKDQ tubes. In Fig. 5.4 it can
be seen that the yield stress distribution has the shape of a ‘W’; the greatest
increase in yield stress (other than at the weld seam) takes place at the 180°
position and the least amount of work hardening occurs at the 90° and 270°
locations. Similarly, Fig. 5.5 shows that the variation in n-value around the
tube circumference has the shape of an ‘M’. The same observations were
confirmed for every grade of steel and for each wall thickness considered
in the A/SP study.
Since much of the expansion in a hydroformed tube occurs in the circumferential or hoop direction of the tube, it is also necessary to evaluate the
formability in this direction. Tensile specimens are machined from the tube
in the hoop direction, flattened and then tested in uniaxial tension. In the
A/SP investigation of tube mechanical properties (Zimerman, 2003), the
longitudinal and transverse properties were determined for a range of tubes
made from different grades of steel and having different coatings and wall
thickness. Zimerman (2003) showed that the distribution of transverse
300
Tube – longitudinal direction
Sheet – longitudinal direction
Yield stress (MPa)
280
260
240
220
200
0
45
90
135
180
225
Position (degrees)
270
315
360
5.4 Distribution of the longitudinal yield stress across the sheet and
around 76-mm-diameter AKDQ tubes (adapted from Zimerman, 2003).
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0.24
0.23
n-value
0.22
Tube – longitudinal direction
Sheet – longitudinal direction
0.21
0.20
0.19
0.18
0
45
90
135
180
225
Position (degrees)
270
315
360
5.5 Distribution of the longitudinal strain hardening exponent across
the sheet and around 76-mm-diameter AKDQ tubes (adapted from
Zimerman, 2003).
mechanical properties around the circumference of the tube exhibited the
same characteristic patterns as did the longitudinal mechanical properties:
thus the variation in the transverse yield stress also has the shape of a ‘W’,
and the variation in the transverse n-value has the shape of an ‘M’. Moreover, Zimerman (2003) showed that for most tube mechanical properties
(tensile strength, uniform elongation and total elongation), the longitudinal
and transverse means were essentially the same. However, the mean yield
stress was consistently somewhat higher in the longitudinal direction than
in the transverse direction, particularly for tubes made from higher strength
steel.
5.3.2 The ring-hoop tension test
When transverse tensile specimens are made, plastic unbending of the
specimen occurs during the flattening process and is known to affect the
values of the yield stress in the hoop direction, particularly for small diameter tubes. In order to determine mechanical properties in the hoop direction without introducing unnecessary work hardening by flattening, some
researchers (Wang et al., 2001) developed the ring hoop tension (RHT) test.
The RHT test consists of mounting a ring specimen with a reduced crosssection onto a pair of ‘D’ shaped blocks that are pulled in tension in a uni-
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Teflon
½ l
D blocks
q
½ l
Reduced section
Pin
Ring sample
5.6 Schematic of the test fixture for the ring hoop tension test
(Wang et al., 2001).
versal tensile testing apparatus (Fig. 5.6). Two layers of Teflon sheet coated
with vacuum grease are used to reduce the friction between the ring specimen and the fixture to near zero. A strain extensometer is attached to the
reduced cross-section and measures the strain in the gauge area. As the ‘D’
blocks are pulled apart, the load is transferred to the specimen which in
turn, deforms without any change in curvature. Wang et al. (2001) described
the analysis procedure used to determine the stress–strain behaviour of the
ring specimen.
In the A/SP study on tube mechanical properties, Zimerman (2003) compared the transverse mechanical properties obtained with both flattened
tensile specimens and with RHT test specimens for 76 mm-diameter AKDQ
tubes. In this investigation, the RHT tests were carried out by the same
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Formability analysis for tubular hydroformed parts
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team of researchers (Wang, 2002) that developed the RHT test, without
having prior knowledge of the tensile data obtained from the flattened
transverse specimens. Zimerman (2003) indicated that the stress–strain
curves obtained by both methods were very similar, and that the differences
between mechanical properties obtained by both methods were not statistically significant. However, it should be noted that the RHT tests carried out
for this A/SP investigation only considered 76 mm-diameter tubes with a
2.0 mm wall thickness (D/T = 38). Prior work carried out by Wang et al.
(2001) on 63.5 mm diameter tubes with a 2.36 mm wall thickness (D/T =
26.9) did show that the two types of tests led to significantly different transverse properties, especially in regards to the yield stress and the uniform
elongation.
It would appear therefore that the RHT test is more suitable for determining the transverse mechanical properties of tubes which have a smaller
D/T ratio. Nevertheless, as pointed out by Zimerman (2003), it is likely that
the conventional tensile test carried out with flattened transverse specimens
is adequate for tubes with a larger D/T ratio, particularly when the standard
tensile test is simpler to carry out than the RHT test.
5.3.3 The tube free-expansion (burst) test
Because of its simplicity and widespread availability, many researchers
prefer to determine tube mechanical properties and the flow stress by carrying out tensile tests. However, the tensile test only describes work hardening behaviour up to the onset of tensile instability (i.e. necking). In order
to use these data in numerical simulations of hydroforming, it is usually
necessary to extrapolate the stress–strain curve from a tensile test to larger
values of effective plastic strain. While such an extrapolation can certainly
lead to good predictions of tube bending (Gholipour et al., 2004; Oliveira
et al., 2005) and tube hydroforming operations (Levy et al., 2004), it remains
that a flow curve representing the biaxial behaviour of the entire tube is
more reliable than an extrapolation of a flow curve obtained from a tensile
test.
It was also seen that continuous roll-forming (to focus on the most
common source of hydroformed tubes) leads to non-uniform work hardening around the circumference of a tube with the greatest work hardening
occurring at 180° from the weld seam. By the same token, the wall thickness
is not constant around the tube circumference, with thickening usually
taking place at the 180° position and thinning taking place elsewhere, but
more so at the 90° and 270° positions. Although the non-uniformity in work
hardening seems to be largely compensated by the non-uniformity in wall
thickness, a ‘weak spot’ inevitably develops somewhere on the tube perimeter. In other words, there is a specific angular position relative to the
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weld seam that is most susceptible to bursting when the tube is pressurized.
The exact location of this weak spot is dependent on the type of tube mill
and on the actual set-up of the mill. But if this weak spot doesn’t happen
to be in a location where tensile specimens are typically taken (i.e. 0°, 90°,
180° or 270° from the weld seam), then tube mechanical properties determined from standard tensile tests will not provide an accurate description
of the tube formability.
In order to more easily determine the location of this weak spot,
Zimerman (2003) suggested that a RHT test be carried out on a ring
specimen without a reduced section. And once this critical location is
identified, longitudinal tensile specimens can be taken from the tube at
this same angular position. However, many researchers (Altan and
Ahmetoglu, 2000; Songmene, 2000; Koç et al., 2001; Green, 2003; Kuwabara
et al., 2003; and Chen et al., 2004) consider the tube free-expansion test,
or burst test (the term ‘bulge test’ will not be used in this chapter since
it originally referred to the test in which a flat blank is clamped around
its periphery and pressurized from one side), to be more representative
of actual tube behaviour than the tensile test. In the tube free-expansion
test, a section of tube is gripped and sealed at both ends and pressurized with a water-based fluid. As the internal pressure increases, the tube
expands in the radial direction (Fig. 5.7). The tube invariably fails at its
Fa
ha
Fa
hb
a
b
5.7 Schematic of the tube free-expansion test with a internal pressure
and axial load and b internal pressure only. It can be noticed that the
bulge height hb < ha. (Koç et al., 2001).
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Formability analysis for tubular hydroformed parts
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weakest point and the burst pressure, which is primarily dependent on
material properties and wall thickness, provides a measure of tube formability. The radial expansion of the tube can be measured with one or
more laser transducers or linear variable displacement transducers (LVDT).
The free-expansion test is sometimes used to quickly verify the formability of tubes supplied to a hydroforming production line by measuring
the burst pressure and the maximum radial expansion just before
burst.
Some free-expansion test facilities allow for an axial compressive force
to be simultaneously applied to the tube ends as it expands (Fig. 5.7),
thus simulating end feeding in hydroforming. Axial compression has
been shown to lead to an increase in the radial expansion of the tube
(Songmene, 2000). By changing the ratio of axial force to internal pressure, the mode of deformation in the tube can be varied between planestrain deformation (no axial force) and approximately uniaxial tension
(maximum axial force). The axial and circumferential strains in the bulged
region of the specimen can be measured by using a spherometer and
large-deformation strain gauges. The corresponding principal stresses
(Fig. 5.8) can be determined from the axial force, the internal pressure
and the geometry of the bulged area (see Kuwabara et al., 2005 for
details). The tube free-expansion test has been used to determine the
work hardening behaviour of tubes in the circumferential direction under
a biaxial mode of deformation. The flow curve thus obtained can then
be used as material data input for numerical simulations of hydroforming
applications.
sz
rθ
j
q
rz
sθ
t
5.8 State of stress in the tube wall during a free-expansion test
(Koç et al., 2001).
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5.4
Tube-forming limits
5.4.1 Strain-based forming limits
The concept of a forming limit for sheet metal alloys was pioneered by
Keeler and Backofen (1963) and Goodwin (1968). They determined experimental forming limits by measuring the principal surface strains on sheet
specimens formed to the onset of localized necking. Keeler and Goodwin
also generated forming limit diagrams (FLD) in principal strain space in
which a forming limit curve (FLC) represents the boundary beyond which
there is a risk of necking for a given sheet metal. Therefore, combinations
of principal surface strains that lie below the FLC lead to a safe forming
operation, whereas those that lie above it lead to failure. Figure 5.9 represents a typical FLD for low-carbon sheet steel. Since many tubes used for
hydroforming applications are roll-formed from cold-rolled sheet, it can be
argued that the concept of the FLD equally applies to tubes.
It can be seen from the shape of the FLC in Fig. 5.9, that formability is
not constant for a given material: it rather depends on the mode of deformation, or the particular combination of principal surface strains. For
instance, plane-strain deformation (when the minor strain is zero) is the
deformation mode with the least formability. The lowest point on the FLC
80
Major
strain
(%)
70
Failure zone
60
50
Marginal zone
40
Safe zone
Safe zone
30
20
10
0
-40
-30
-20
-10
0
10
20
Minor strain (%)
5.9 Typical FLD for low-carbon sheet steel.
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Formability analysis for tubular hydroformed parts
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is commonly referred to as the plane-strain intercept and is written in short
form as FLC0. An empirical relationship was developed by Keeler and
Brazier (1977) to determine the position of the FLC in strain space without
having to carry out time-consuming experiments: the position of FLC is
uniquely located by calculating the value of the plane-strain intercept, as
follows:
FLC 0 (%) = (23.3 + 14.13t )
( 0.n21 )
[5.1]
where t is the sheet thickness in mm, n is the terminal work hardening
exponent and FLC0 is the major engineering strain given in percent. This
relationship has been widely used in North America and elsewhere and
has been shown to be both accurate and reliable for low-carbon sheet
steel.
Since the early work of Keeler and Goodwin, researchers around the
world have developed various experimental and analytical techniques to
determine the FLC for a given sheet material. A comprehensive overview
of these techniques is given by Green and Black (2002). But regardless of
the method used to determine the FLC, the sheet metal forming industry
widely uses the FLD to evaluate the forming severity of stamped components, and this has led to significant reductions in tryout time, improved
process robustness and product quality.
In view of its widespread usage to evaluate the quality of stamped or
drawn parts, it is often assumed that the FLD can also be applied to hydroformed tubular components. The A/SP investigated this assumption by
having the FLC determined experimentally for three automotive grades of
steel tubes [AKDQ, HLSA and interstitial-free (IF)]. Each FLC was
obtained by applying different combinations of internal pressure and axial
tension/compression to a series of tubes. Strain paths thus generated covered
a range of deformation modes from biaxial tension to uniaxial tension. Each
test was carefully interrupted at the onset of necking. It was found (Green,
2003; Green and Stoughton, 2004) that the experimental FLC correlated
very closely with those predicted by the Keeler–Brazier relationship [5.1],
provided the tube mechanical properties at the weakest angular position
were used. However, this work only validates the use of the Keeler–Brazier
prediction for straight tubes tested in free-expansion. As will be shown,
there are important reasons why the standard FLC is not generally applicable to tubular hydroformed parts.
5.4.2 Path dependency of strain-based forming limits
Several researchers (Ghosh and Laukonis, 1976; Kleemola and Pelkkikangas, 1977; Arrieux et al., 1982; Gronostajski, 1984; Graf and Hosford, 1993,
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1994) have shown experimentally that a non-linear strain path can change
the shape and location of the FLC in strain space. Ghosh and Laukonis
(1976) investigated strain path dependence by prestraining aluminium
specimens to various levels of strain and in different deformation modes:
a different FLC was obtained for each value of prestrain. Kleemola and
Pelkkikangas (1977) determined the FLD for steel, copper and brass sheets
after both uniaxial and equi-biaxial prestrains and observed that the change
in the FLC depended on the magnitude and type of the prestrain. Arrieux
et al. (1982, 1983) carried out similar experiments with an aluminium alloy.
Graf and Hosford (1993a, 1993b) also reported strain-path effects for 2008
T4 aluminium pre-strained in uniaxial, equi-biaxial and near plane-strain
tension. The result was a different FLD for each value of prestrain as shown
in Fig. 5.10. Graf and Hosford (1993a) also attempted to predict the shifted
FLD using the Marciniak-Kuczynski (M-K) analysis; however, the predicted FLC did not correlate well with experimental data when the prestrain was in equi-biaxial tension.
These experimental observations show that, depending on the loading
history, the actual FLC can be significantly different from the as-received
FLC. As a result of a change in shape and position, combinations of principal strains that are safe from necking can lie above the as-received FLC,
and conversely, failures can take place at strains below the as-received FLC.
Furthermore, during any forming operation, different locations on a part
undergo different strains and forming modes. If the component is manufactured in two or more forming stages, the overall strain path in each location
can be severely non-linear as a result of following one strain path in one
forming stage and a different strain path in the next forming stage. In order
to properly assess the forming severity of the part after the last forming
stage, a different FLC would be required for each location depending on
the specific strain history at that location. This would call for a large number
of FLC for one and the same part. This signifies that the conventional strainbased FLC is inadequate to evaluate the forming severity of parts that were
produced in complex, multi-stage forming processes such as are typical for
hydroformed tubular components.
5.4.3 Process window
In some recent work, Chu et al. (2006) carried out both theoretical and
experimental investigations into the failure limits of straight aluminium
tubes subject to internal pressure and axial compression. The authors point
out that, because of different loading conditions in sheet forming and tube
free-expansion, different plastic instability criteria may lead to significantly
different FLC. In view of the complexity and uncertainty of predicting
accurate FLD for tubes, these authors (Chu and Xu, 2004b; and Chu et al.,
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2006) recommend the use of a process window diagram (PWD) which
defines a safe operating window in terms of the axial force and the internal
pressure. The PWD offers the advantage of predicting both buckling/
wrinkling instability and bursting limits of a straight tube in free-expansion
hydroforming on the basis of theoretical analyses (Chu and Xu, 2004a), and
was shown to be in good agreement with experimental data. However, it is
not clear how the PWD would be established for typical industrial hydroforming applications in which the tube is bent a number of times, possibly
crushed and hydroformed inside a die.
5.4.4 Stress-based forming limits
The issue of path dependency of strain-based forming limits has led some
researchers (Arrieux et al., 1982, 1983; Gronostajski, 1984; Stoughton, 2000;
Ofenheimer et al., 2005; and Butuc et al., 2006) to consider that the
forming limit of sheet materials is dependent on the state of stress, rather
than the state of strain. Stoughton (2000) presented a detailed procedure
for transforming strains into stresses based on classical plasticity theory.
Assuming the material has a given yielding behaviour and known work
hardening behaviour, any state of strain can be uniquely mapped into
principal stress space. Therefore, the particular states of strain that define
the as-received FLC in strain space also define a unique FLC in stress
space: the SFLC. And, in the same manner that the FLC defines the
lower limit of the failure zone in strain space, any state of stress that lies
above the SFLC also presents a risk of failure. Ofenheimer and Kitting
(2006) showed that the prediction of the SFLC from the experimental
FLC for a 5154 aluminium alloy using this transformation procedure
correlates very well with the SFLC that was obtained experimentally by
Yoshida et al. (2005).
According to this theoretical framework proposed by Arrieux et al. (1982)
and generalized by Stoughton (2000), the stress-based forming limit appears
to be independent of strain path. Indeed, Arrieux et al. (1983) showed that
the different forming limit strain data obtained after various types of prestrain transformed into a single curve in stress space. Similarly, Stoughton
(2000) showed that the experimental forming limits obtained by Graf and
Hosford (1993b) after a wide range of non-linear strain paths (see Fig. 5.10)
all collapsed into a narrow band in stress space. Stoughton (2002) and Butuc
et al. (2006) have demonstrated that, while the SFLC itself is sensitive to
the yield criterion and the hardening law that are used in the stress transformation, the SFLC remains independent of strain-path regardless of the
constitutive equations that are used. It has sometimes been argued that
the SFLC only appears to be path-independent because of the shape of the
stress–strain curve that tends to saturate at higher levels of strain. However,
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0.35
0.05 prestrain
0.12 prestrain
0.18 prestrain
As-received
0.04 prestrain
0.125 prestrain
0.18 prestrain
True major strain
0.25
0.15
0.05
–0.05
–0.2
–0.1
0
0.1
0.2
True minor strain
0.3
0.4
0.5
5.10 Influence of strain path on the FLC (adapted from Graf and
Hosford, 1993).
Stoughton (2001) and Butuc et al. (2006) have shown that this is not the
case.
Following upon this theoretical work, several authors have carried out
experiments to validate the postulation of path-independence of the stressbased forming limit criterion. Green and Stoughton (2004) carried out tests
on straight 76.2 mm-diameter AKDQ tubes subjected to axial tensioncompression loading and internal pressure. Tests were interrupted at the
onset of necking and very meticulous thickness and strain measurements
were made to determine the tube FLC. Another series of similar tests were
conducted on the same tubes after they were prestrained and the tube FLC
was once again determined. While the as-received and prestrained tubes
had very different FLC in strain space, it was shown that they had virtually
the same FLC in stress-space.
Further confirmation of the path-independence of the SFLC was provided by Yoshida et al. (2005). Yoshida and co-workers carried out tests on
straight aluminium tubes subjected to combined tension–internal pressure
using a servo-controlled testing facility built by Kuwabara et al. (2003 and
2005). In these tests, a range of stress paths were followed to fracture and
the stress forming limits were measured. Again, this work showed that the
SFLC is practically strain-path independent. This experimental evidence
strongly suggests that the SFLC criterion is ideally suited to evaluating the
forming severity of hydroformed parts that are subject to multi-stage
forming where strain paths are frequently non-linear.
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Although the SFLC criterion has been shown to be far more suitable
to evaluate the forming severity of hydroformed parts than the conventional strain-based FLC, some researchers have recently shown there
are some limitations to the path-independence of the SFLC. Stoughton
(2001) pointed out that the discrepancy in the predicted SFLC of specimens loaded up to higher levels of prestrain (e = 0.18) compared with
those prestrained to lower levels was no doubt due to kinematic hardening effects following the unloading and reloading along the secondary
strain path. Yoshida et al. (2007) also investigated the path-dependency
of the SFLC by simulating forming limits using an M-K (MarciniakKuczynski) analysis with two types of combined loading: one consisting
of two linear stress paths in which unloading is included between the
first and second loading, and one where the strain path is abruptly
changed without unloading. The authors found that the SFLC was practically path independent for the first type of combined loading, whereas
the SFLC was path dependent when no unloading took place. Yoshida
and Kuwabara (2007) also conducted experiments on steel tubes under
combined tension–internal pressure to further investigate the influence
of loading history on the path dependence of SFLC. They demonstrated
that the strain hardening behaviour of the prestrained material, not the
magnitude of the prestrain, determined whether the SFLC was path
dependent or not: indeed kinematic hardening effects during the reloading led to path dependence.
In order to account for the three-dimensional state of stress in the tube
wall during hydroforming, one team of researchers (Simha et al., 2005;
Simha and Worswick, 2006) proposed an extended stress-based forming
limit criterion (XSFLC). By using equations developed by Stoughton (2000)
and by assuming the stress invariants – the equivalent stress and the mean
stress – are representative of the formability limit under three-dimensional
stress states, they transformed the SFLC into a corresponding curve on a
graph of equivalent stress vs. mean stress called the XSFLC. When the state
of stress in the tube crosses the XSFLC, failure is predicted to occur in the
hydroformed part.
Although recent studies indicate some path dependence of the SFLC
under certain conditions, it appears that this occurs primarily for severely
non-proportional loading histories where the initial prestrain is in biaxial
tension and where the material exhibits definite kinematic hardening effects.
Certainly, these experimental observations indicate that further work is
required to establish a failure criterion that is unambiguous even under
complex loading histories. However, for the time being, it would seem that
the stress-based failure criterion – either the two-dimensional SFLC or the
three-dimensional XSFLC – is the best available tool for evaluating the
forming severity of hydroformed components.
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5.5
Formability analysis for numerical simulations
The hydroforming of tubular components is a process that is routinely
simulated using commercial computation programs that are based on the
finite element method. Early in the product-design stage a computer model
of the proposed part and its corresponding tooling are built. A numerical
simulation of the proposed forming process enables the analyst to predict
the stresses and strains in the part as well as its shape and thickness distribution. Since most commercial codes do not incorporate a failure criterion that automatically determines the onset of necking and splitting in the
part, it is common practice to introduce the appropriate strain-based FLC
along with other pertinent material data. This allows the analyst to determine, during the post processing of the results, whether or not the strains
in the part exceed the FLC. When they do, the virtual model is modified
by making changes to the part or tool design, or by changing the processing
conditions; such iterations continue until both the product and the process
design are deemed to be safe. Only then will prototype or production
tooling be cut.
It is certainly beyond the scope of this chapter to discuss the different
types of finite element programs, the various time-integration schemes and
the multitude of numerical parameters that affect the outcome of virtual
process simulations. A basic set of recommendations for numerical simulation of hydroforming applications is given by Green (2004). Nonetheless it
will be mentioned that, in order to decrease computation time, plane-stress
shell elements are generally recommended for tube-forming processes. As
tube wall thickness increases and corner radii decrease, however, greater
accuracy can be obtained by increasing the number of through-thickness
integration points or by using three-dimensional solid elements. Furthermore, the stress and strain gradients through the thickness of the tube wall
need to be considered when carrying out a virtual formability analysis: it is
generally recommended that the minimum through-thickness stresses or
strains at a given location in the wall be compared with the forming limit.
Experience with virtual forming evaluation has shown that the FLC is
not a reliable failure criterion for tubular hydroformed parts. As previously
discussed, the SFLC has been shown to be far less sensitive to strain-path
than the FLC and, for this reason, Stoughton (2000) and Butuc et al. (2006)
have proposed the use of the SFLC to evaluate forming severity in hydroforming process simulations. Since numerical codes calculate the stresses in
the virtual part, it is straightforward to compare the predicted stresses to
the SFLC. An important caution put forth by these authors is that the plasticity model used to transform the as-received FLC into stress space to
determine the SFLC should be the same as the one used in the numerical
code to compute the stresses in the part.
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115
Simha et al. (2007) implemented the XSFLC failure criterion into a finite
element program to predict the onset of necking in hydroformed parts.
These researchers also carried out hydroforming tests on pre-bent 5018
aluminium tubes and DP600 steel tubes. The comparison of predicted and
experimental expanded cross-sections and failure locations showed that the
FLC was not able to accurately predict the onset of necking. However, the
use of the XSFLC criterion (with some additional assumptions) led to predictions of the failure locations that correlated well with experimental
data.
Finally, both Stoughton and Yoon (2005) and Simha et al. (2007) underscore the importance of tracking the entire stress history in certain critical
elements during a numerical simulation of a hydroforming process. Indeed,
these authors showed that a critical state of stress can be reached at some
intermediate stage in the simulation, and subsequently be followed by an
elastic unloading or a change of path leading to a less severe stress state.
It is therefore not sufficient to compare the stress state in the final stage
of the numerical simulation to the SFLC. Rather, virtual formability analysis should consist of comparing the most critical stress states to the SFLC,
even though they may have occurred before the end of the hydroforming
process.
5.6
Formability analysis in the plant
It appears that the most promising approach to evaluating forming severity
of a virtual hydroformed component is to incorporate a stress-based forming
limit criterion (SFLC) into the numerical simulation program and compare
predicted stresses with the SFLC. This approach is very helpful for simulation analysts, but since forming stresses in a part cannot be measured or
recorded (and residual stresses do not provide a history of the stresses
during the forming process), it cannot be adopted in the prototype shop or
the production plant.
Although the SFLC has been shown to be somewhat sensitive to loading
history for certain classes of material and under severe non-proportional
loading, it may nevertheless be a reasonable engineering approximation to
assume the SFLC is path-independent for a majority of hydroforming applications. This assumption would allow shop floor personnel to follow a
simple, user-friendly procedure to evaluate the forming severity of tubular
hydroformed parts that were formed in a multi-stage process with nonproportional strain paths. A possible procedure is described as follows:
•
the first step consists of identifying critical areas in the hydroformed part
where splitting is likely to occur; this can be done with the use of finite
element simulations or with the helpful insight of a press operator.
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•
•
Hydroforming for advanced manufacturing
Secondly, it is proposed that the strain distribution in the part be determined after each forming stage – for instance after the bending operation, after the crushing operation and after hydroforming. The strains
can be measured by electro-etching a circle grid onto the straight tubes
and measuring the distorted circles after each forming stage.
Assuming the SFLC of the tube and the actual strain path in each critical
location are known, Stoughton’s general procedure (2000) can be followed to transform the SFLC back into strain space and obtain a modified FLC that is a function of the particular strain path for a given
location in the part. As a first approximation, the strain path can be
considered linear between each forming stage, and for greater accuracy,
the strain path predicted by a finite element simulation of the forming
process can also be used.
This procedure can rapidly provide a useful FLC against which to evaluate strains in a given location on the part. Since each critical location will
undergo a different strain path, a different FLC will be required for each
location. For most industrial applications, however, the number of locations
that present a risk of fracture ought to be relatively small and therefore this
procedure should be quite user-friendly.
5.7
Conclusions and future trends
This chapter has provided an overview of the issues related to the formability of tubes for hydroforming applications. In particular, the non-uniform
distribution of mechanical properties of as-received tubes was highlighted
as well as the impact of each forming stage on the overall formability of
the component. Tube bending was shown to lead to a significant loss in
formability and also to severe strain gradients in the tube.
Tube formability tests were reviewed and it was seen that uniaxial tension
tests taken from various locations in the tube wall can easily be carried out
to determine the mechanical properties and the work hardening behaviour
of the tube. However, the free expansion test, when available, is generally
considered to provide a better overall description of tube behaviour than
the tensile test.
Various forming limit criteria were presented and it was pointed out that
the conventional forming limit diagram is not suitable to evaluate the
forming severity of hydroformed parts that undergo a multi-stage forming
process or when strain paths are non-linear. Instead, the stress-based
forming limit criterion, although not strictly path independent, is currently
considered to be the most suitable failure criterion for hydroformed applications. Finally, the integration of a stress-based forming limit criterion into
numerical simulations was briefly discussed, and a user-friendly method for
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117
implementing it into the formability analysis of actual hydroformed parts
was also presented.
It was mentioned that there are some instances where the SLFC is dependent on loading history. Future work will no doubt focus on understanding
more fully the mechanical or metallurgical factors that have an influence
on the stress forming limit. Some of these factors include the hardening
behaviour of the material and particularly transient hardening phenomena,
the precise knowledge of the actual loading history, the anisotropy of the
material, etc. In some recent work, Stoughton and Yoon (2005) address
the issue of accurately describing the forming limit for anisotropic materials
subjected to non-proportional loading. They emphasize the need to consider the forming limit as a surface rather than a single curve in stress
space but propose a technique that enables an analyst to view complex
three-dimensional anisotropic forming data in a simple two-dimensional
diagram.
Future research will undoubtedly lead to a more accurate prediction of
tube forming limits and it is expected such criteria will be incorporated into
finite element codes for virtual simulation of hydroforming processes. While
the complexities of advanced failure models will remain largely transparent
to the simulation analyst, the challenge will remain to turn such advanced
models into a simple, user-friendly technique that can be rapidly implemented to assess the forming severity of actual hydroformed parts.
5.8
References
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d’emboutissabilité à différents matériaux pour emboutissage’, Mémoires et Études
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Société Française de Métallurgie, pp. 685–692.
bardelcik a. and worswick m. j. (2006): ‘Numerical investigation into the effects
of bending boost and hydroforming end feed on the hydroforming of DP600
tube’, SAE 2005 Transactions: Journal of Materials and Manufacturing,
pp. 41–52.
butuc m. c., gracio j. j. and barata da rocha a. (2006): ‘An experimental and
theoretical analysis on the application of stress-based forming limit criterion’,
International Journal of Mechanical Sciences, 48, 414–429.
chen k. k., soldaat r. j. and moses r. m. (2004): ‘Free expansion bulge testing
of tubes for automotive hydroform applications’, SAE Technical Paper No.
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Hydroforming for advanced manufacturing
chu e. and xu y. (2004): ‘Hydroforming of aluminum extrusion tubes for automotive
applications: Part I: buckling, wrinkling and bursting analyses of aluminum tubes’,
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chu e. and xu y. (2004): ‘Hydroforming of aluminum extrusion tubes for automotive
applications: Part II: process window diagram’, International Journal of Mechanical Sciences, 46, 285–297.
chu e., xu y., davies r. w. and grant g. j. (2006): ‘Failure predictions for aluminum
tube hydroforming processes’, SAE Technical Paper No. 2006-01-0543.
gholipour j., worswick m. j. and oliveira d. (2004): ‘Application of damage models
in bending and hydroforming of aluminum alloy tube’, SAE Technical Paper
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ghosh a. k. and laukonis j. v. (1976): ‘The influence of strain-path changes on the
formability of sheet steel’, Proceedings of the 9th Biennial Congress of the IDDRG,
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OH, USA.
goodwin g. m. (1968): ‘Application of strain analysis to sheet metal forming in the
press shop’, SAE Technical Paper No. 680093.
graf a. f. and hosford w. f. (1993): ‘Calculations of forming limit diagrams for
changing strain path’, Metallurgical Transactions A, 24, 2497–2501.
graf a. f. and hosford w. f. (1993): ‘Effect of changing strain paths on forming limit
diagram of Al 2008-T4’, Metallurgical Transactions A, 24, 2503–2512.
graf a. f. and hosford w. f. (1994): ‘The influence of strain-path changes on forming
limit diagram of Al 6111-T4’, International Journal of Mechanical Sciences, 10,
897–910.
green d. e. (2003): ‘Experimental determination of tube forming limits’, Proceedings
of the International Conference on Hydroforming, Fellbach, Germany, Oct. 28–29,
2003, edited by Klaus Siegert, DGM, pp. 299–314.
green d. e. (2004): ‘Summary report of the A/SP Hydroforming work’, a report
presented to the Hydroforming Materials and Lubricants task force of the Auto/
Steel Partnership, March 23, 2004, www.a-sp.org.
green d. e. and black k. c. (2002): ‘A visual technique to determine the forming
limit for sheet materials’, SAE Technical Paper No. 2002-01-1062.
green d. e. and stoughton t. b. (2004): ‘Evaluating hydroforming severity
zusing stress–based forming limit diagrams’, Proceedings of the 2nd Annual North
American Hydroforming Conference, sponsored by SME/TPA, Waterloo, ON,
Canada.
gronostajski i. (1984): ‘Sheet metal forming limits for complex strain paths’, Journal
of Mechanical Working Technology, 10, 349–362.
keeler s. p. and backofen w. a. (1963): ‘Plastic instability and fracture in sheets
stretched over rigid punches’, ASM Transactions Quarterly, 56, 25–48.
keeler s. p. and brazier w. g. (1977): ‘Relationship between laboratory material
characterization and press shop formability’, Proceedings of Microalloy 75, Union
Carbide, New York, pp. 447–452.
kleemola h. j. and pelkkikangas m. t. (1977): ‘Effect of predeformation and strain
path on the forming limits of steel copper and brass’, Sheet Metal Industries, 63,
591–599.
koç m., aue-u-lan y. and altan t. (2001): ‘On the characteristics of tubular materials
for hydroforming: experimentation and analysis’, International Journal of
Machines and Tools for Manufacturing, 41(5), 761–772.
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119
kuwabara t., ishiki m., kuroda m. and takahashi s. (2003): ‘Yield locus and workhardening behavior of a thin-walled steel specimen subjected to combined
tension-internal pressure’, Journal de Physique IV, 105, 347–354.
kuwabara t., yoshida k., narihara k. and takahashi s. (2005): ‘Anisotropic plastic
deformation of extruded aluminium alloy tube under axial forces and internal
pressure’, International Journal of Plasticity, 21, 101–117.
levy b. s., vantyne c. j. and stringfield j. m. (2004): ‘Characterizing steel tube for
hydroforming applications’, Journal of Materials Processing Technology, 150(3),
280–289.
ofenheimer a., buchmayr b., kolleck r. and merklein m. (2005): ‘Forming limits in
sheet metal forming for non-proportional loading conditions – experimental and
theoretical approach’, Proceedings of Numisheet 2005, pp. 460–465.
ofenheimer a. and kitting d. (2006): ‘Validation of numerically determined intrinsic
forming limit stresses based on experimental forming limit strain data’, Proceedings of the 12th International Symposium on Plasticity, Halifax, Canada, ed. Khan
A. and Kazmi R., pp. 70–72.
oliveira d. a., worswick m. j. and grantab r. (2005): ‘Effect of lubricant in mandrelrotary draw tube bending of steel and aluminum’, Canadian Metallurgical
Quarterly, 44, 71–78.
simha c. h. m., gholipour j., bardelcik a. and worswick m. j. (2005): ‘Application of
an extended stress-based flow limit curve to predict necking in tubular hydroforming’, Proceedings of Numisheet 2005, Detroit, USA, August 15–19.
simha c. h. m. and worswick m. j. (2006): ‘Stress-based forming limits in hydroforming’, Proceedings of the 12th International Symposium on Plasticity, Halifax,
Canada, July 17–22.
simha c. h. m., gholipour j., bardelcik a. and worswick m. j. (2007): ‘Prediction of
necking in tubular hydroforming using an extended stress-based forming limit
curve’, Transactions of ASME: Journal of Engineering Materials and Technology,
129, 36–47.
songmene v., bauwens c. and moses r. (2000): ‘Free expansion hydroforming of
AKDQ and HSLA steel tubes – experimental determination of circumferential
stress-strain curves’, a report prepared by the Industrial Research & Development Institute (IRDI) for the Hydroforming Materials and Lubricants task force
of the Auto/Steel Partnership.
stevenson r., ng b.-c. and polidoro p. (2004): ‘Failure in internally-pressurized bent
tubes’, Metallurgical and Materials Transactions A, 35(3), 1151–1158.
stoughton t. b. (2000): ‘A general forming limit criterion for sheet metal forming’,
International Journal of Mechanical Sciences, 42, 1–27.
stoughton t. b. (2001): ‘Stress-based forming limits in sheet metal forming’, Journal
of Engineering Materials and Technology, Transactions of the ASME, 123, 417–
422.
stoughton t. b. (2002): ‘The influence of the material model on the stress-based
forming limit criterion’, SAE Technical Paper No. 2002-01-0157.
stoughton t. b. and yoon j. w. (2005): ‘Sheet metal formability analysis for
anisotropic materials under non-proportional loading’, International Journal of
Mechanical Sciences, 47, 1972–2002.
wang h., martin p. and houghland e. (2001): ‘Evaluation of tube materials for tube
hydroforming’, Proceedings of 43rd Mechanical Working and Steel Processing Conference, Charlotte, NC, October 28–31, 2001, p. 251.
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wang h. (2002): ‘Hoop tensile properties of A/SP hydroforming tubes using uniaxial
RHTT’, a report prepared by CANMET for the Hydroforming Materials and
Lubricants task force of the Auto/Steel Partnership, www.a-sp.org.
yoshida k., kuwabara t., narihara k. and takahashi s. (2005): ‘Experimental verification of the path-independence of forming limit stresses’, International Journal
of Forming Processes, 8, 283–298.
yoshida k. and kuwabara t. (2007): ‘Effect of strain hardening behaviour on forming
limit stresses of steel tube subjected to non-proportional loading paths’, International Journal of Plasticity, 23, 1260–1284.
yoshida k., kuwabara t. and kuroda m. (2007): ‘Path-dependence of the forming
limit stresses in a sheet metal’, International Journal of Plasticity, 27, 361–384.
zimerman z. (2003): ‘Analysis of experimental data developed by the A/SP Hydroforming committee’, a report presented to the Hydroforming Materials and
Lubricants task force of the Auto/Steel Partnership, Dec. 29, 2003.
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6
Design and modelling of parts, process and
tooling in tube hydroforming
M . S T R A N O, Università degli Studi di Cassino, Italy
6.1
Introduction to tube hydroforming design
When starting the design of a new tube hydroforming (THF) operation, the
process planner is usually provided with the desired shape and properties
of the final part. These specifications are non-modifiable, unless an efficient
concurrent engineering approach is followed. Although the design of the
final part is predetermined, several alternative processes can be used to
manufacture the same final product. In fact, a very large number of process
conditions and variables must be properly selected (summarised in Fig. 6.1):
the number and type of intermediate steps (pre-bending, preforming or
crushing, intermediate annealing, etc.), the position of the parting lines on
the dies, the location of the tube welding line (unless tubes are seamless),
the initial wall thickness and cross-section (shape and dimensions) of the
tube, the lubrication conditions, the number and position of axial and transverse punches and counterpunches and, finally, the loading paths, i.e. the
curves describing the internal pressure vs. time and the axial feed (either
displacement or force) vs. time. In some cases, the final desired shape is so
simple or conventional that most of the choices are obvious and only the
selection of time-dependent variables, i.e. the loading paths, can be thought
of as a critical choice.
As an example, if the target part is a T-shape (Fig. 6.2), the following
decisions are straightforward: no preforming is required, the parting lines
should be located as shown in Fig. 6.2, the welding line should be placed at
the opposite side of the T branch, i.e. at the bottom of the tube, where no
tensile deformation will occur, the initial tube cross section should be round
with an initial outer diameter slightly smaller than the final outer diameter,
the dies should be very well lubricated, especially in the guiding zone, and
the punches should be placed as shown by the arrows in Fig. 6.2. Some other
variables cannot be designed in such an intuitive way, but must be carefully
planned, namely the initial tube length (and, in some cases, the initial wall
thickness as well), the control curves of the two axial punches and the
121
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Hydroforming for advanced manufacturing
Process
steps
Dies
parting
lines
Loading
paths
Process
design
Axial and
transverse
punches
Lubrication
conditions
Position of
welding
line
Tube
thickness
and
section
6.1 Conditions and variables to be selected when designing a THF
process.
nch
erpu
t
Coun
l
Axia h
punc
l
Axia
h
c
n
u
p
6.2 Hydroforming of a T-shape.
transverse counterpunch, and finally, the internal pressure vs. time curve.
These variables should be designed according to available guidelines and
by following the indications of FEM simulations.
When less obvious geometries are involved (unfortunately this is the
most frequent case), the design process is usually a difficult and lengthy
task, which often requires a number of iterative steps. In the present chapter,
some basic remarks and guidelines are provided in order to facilitate process
planning of THF operations.
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6.2
123
Technological classifications of
tube hydroforming processes
In order to predict the difficulty of the required design activities and to
select the appropriate tools, methods and software support, it is very useful
to identify the critical features of the hydroforming process and to present
a simple classification of them. Three important characteristics of THF
operations are the process sequence, the type of tooling layout and the
amount of required axial feeding. These subjects will be discussed in the
following three subsections and a related classification is proposed in Table
6.1.
6.2.1 Process sequence
The complete process sequence required for complex parts includes bending,
crushing or preforming, hydroforming and calibrating. It may also require
a number of minor intermediate or complementary operations, such as
sawing, washing, annealing, deposition of lubricant, trimming, punching and
postforming. From a design point of view, bending and calibrating are not
particularly critical, since they are directly related to the part geometry: the
calibration pressure depends mainly on the value of the smallest concave
curvature radius; the number and type of bending operations depends on
the number and severity of bends which lie on the tube centreline.
The two operations that usually deserve most of the attention are preforming and hydroforming itself. An important distinction can be made
between processes that require preforming and processes that can be safely
hydroformed without preliminary mechanical operations. The crushing or
preforming phase is necessary when the outer diameter of the initial tube
(or its height in case of a rectangular cross-section) is larger than the die
Table 6.1 Classification of THF processes
Categories
A
B
C
D
According to
the process
sequence
According to
the tooling
layout
According to
the amount
of axial
feeding
With
preforming
Without
preforming
–
–
With
countertools
Without
countertools
–
–
None
dax/li ≈ 0
Small
dax/li ≈
0.01–0.1
Medium
dax /li ≈ 0.1–0.2
Large
dax /li ≈ 0.4
or more
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Upper die
Upper die
Upper die
Tube
Tube
6.3 Crushing operation of a round tube within a polygonal die.
Upper die
Upper die
Pinching
Tube
Tube
Tube
6.4 Crushing of a bumper rail section with potential pinching.
allowance (Fig. 6.3). When crushing is necessary, pinching must be avoided
as the upper and lower dies are being closed (Fig. 6.4). A classification of
hydroforming processes will be proposed in Section 6.4.3.
6.2.2 Tooling layout
Often, when hydroforming a T-shape (Fig. 6.2), a Y-shape or an X-shape
(Fig. 6.5), a counterpunch is generally used to avoid premature bursting of
the tube. The role of the counterpunch is to inhibit free bulging of protrusions, i.e. to provide compression stresses which have several beneficial
effects: maximum thinning is reduced, the shape of the protrusion is more
regular, the maximum useful protrusion height is increased (Fig. 6.6). The
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125
dax
p
dax
Thickness (mm)
6.5 Hydroforming of an X-shape (Kocanda, 2006).
3.2
3
2.8
2.6
2.4
2.2
2
1.8
1.6
1.4
BOTTOM
section
TOP section
With counterpunch
Without counterpunch
0
50
100
150
200
Curvilinear abscissa (mm)
TOP section
Without
counterpunch
With
counterpunch
BOTTOM section
a
b
6.6 Shape and thickness distribution (initial thickness t0 = 2 mm)
during hydroforming of a protrusion a without a counterpunch and
b with a counterpunch (right).
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Hydroforming for advanced manufacturing
complexity of process planning is obviously strictly related to the number
of independent actuators that must be controlled. THF processes can be
divided into different classes depending on whether they need only axial
feeding or also require transverse counterpunches.
6.2.3 Amount of feeding
The role of axial feeding in THF is to provide axial compressive stresses in
the forming zone. In order to obtain a better correlation between the
control of axial punches and the amount of axial stresses in the tube, many
use the axial force Fax vs. time as the driving design parameter, rather than
the axial displacement dax. The main reason is that, in the hydroforming
presses, a minimum value of axial force is required at the punches in order
to avoid leaking. From the standpoint of the press manufacturer, axial force
control is more straightforward. However, in the present chapter, the
parameter dax is preferred because a large number of decisions are often
based on FEM simulations run with shell elements and explicit time integration, which yield more accurate results in terms of strain rather than
stress distributions. As a consequence, FEM results are more reliably
managed and implemented when stroke control is performed rather than
force control.
In some cases, the amount of tube expansion within the die is very
limited, i.e. the difference between the initial tube’s perimeter and the
die’s perimeter is very small. This happens either because an important
crushing operation is accomplished before forming or because the crosssection of the tube does not significantly change across the tube centreline.
In both cases, the amount dax of displacement of the axial punches is very
small or even zero. The displacement dax is meant as the total amount of
axial feeding, obtained by summing the displacement of the right and left
axial punches. When large expansion ratios must be obtained, a comparably large amount of axial feeding is required, unless the materials’ formability is particularly good, such as for stainless steel tubes with large
initial wall thickness. Given two tubes with different initial length Li and
equal amount of dax, the compressive effect will obviously be larger for
the shorter tube. As an example, the tube labelled ‘A’ in Fig. 6.7 is 39%
longer than tube ‘B’. If an FEM simulation is run with the same displacement vs. time and pressure vs. time curve, the resulting maximum compressive axial stress in the bulging area, before calibration, is about 12% larger
for the shorter tube ‘B’. Therefore, the non-dimensional parameter dax/Li
might have a more general and meaningful use than the absolute value
of axial feed dax. From a technological stand point, THF processes can be
grouped according to the required value of the feeding dax/Li, as summarised in Table 6.1. As an example, if a left axial displacement of 21 mm is
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127
Die
TubeA
TubeB
Li = 160 mm
dax/Li = 0.13
Li = 222 mm
dax/Li = 0.09
6.7 Tubes with different initial length Li.
given to tubes ‘A’ and ‘B’ in Fig. 6.7, the resulting dax/Li values are, respectively, 0.09 and 0.13.
6.3
Hydroformability of tubular parts
The hydroformability of THF parts may be defined as the ability of a given
workpiece to undergo the plastic deformation induced by a THF process
without the insurgence of defects, such as excessive localised thinning or
fracture, wrinkling or buckling. The biggest role in determining the formability of THF parts is played by the desired final geometry of the tube. The
ductility of the tube material is an important factor as well. Several attempts
at classification of tubular hydroformed parts can be found in the literature.
Usually, in papers and technical reports on the fundamentals of THF, parts
are classified according to their function as an end product. Geometrical
classifications of THF parts were proposed in Engel et al. (1995) and Koc
(1998).
6.3.1 Shape of cross-sections
One of the most important geometrical parameters for assessing the feasibility of a new THF process is the maximum required expansion of the
initial tube. The expansion ratio may be defined as the percentage increase
of the tube perimeter. If expanding a round tube into a cylindrical bulge,
the expansion ratio is simply the percentage increase of the tube diameter.
However, at the beginning of the design process the shape of the initial tube
is still an unknown variable and, therefore, the real expansion ratio is not
available. Fortunately, a geometrical parameter which is very similar to the
expansion ratio is available: the expansion ratio between the smallest
(Pcsmin) and the largest (Pcsmax) perimeters of the cross-section that can be
found running across the final tube centreline (Fig. 6.8). The maximum
expansion of the cross-section is therefore:
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Hydroforming for advanced manufacturing
Pcsmax
Welding line
Pcsmin
6.8 Cross-sections with variable perimeters Pcs.
CSexp =
Pcsmax − Pcsmin
100
Pcsmin
[6.1]
Clearly, when CSexp is large, finding a feasible process window becomes
more difficult. When CSexp is too large (compared with the material formability) and axial feeding is difficult due to the part geometry, crushing is
the only other option to be considered.
Another relevant geometrical parameter of the cross-section is the tube
initial wall thickness ti. Obviously, it is primarily dependent on the specifications of the final part. However, different processes will generate different
distributions of final thickness tf across the tube surface. For this reason, the
process designer sometimes is free, to a limited extent, to select the correct
value of initial tube wall thickness ti. The material’s formability usually
increases with larger ti, since the tendency to wrinkling is reduced and the
strains at fracture are generally increased. However, the main drawback of
thick tubes is that very large pressure values are required and, as a consequence, the desired final shape is more difficult to obtain, since the smallest
corner radii may not be completely filled. Furthermore, as the example in
Fig. 6.9 shows, for a given die shape, the point of potential fracture over the
tube surface may shift to different locations as the initial tube thickness
changes. For this reason, the fracture risk is not necessarily a monotonously
decreasing function of ti, but there might be an optimal value of initial
thickness which maximises the formability.
6.3.2 Shape of tube centreline
Once the tube initial outer diameter (OD) is fixed, the longer the length of
the tube centreline, the higher are the friction forces. This limits the material
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Parts, process and tooling in tube hydroforming
129
Probable locations of fracture
ti = 2 mm
ti = 0.6 mm
6.9 Locations of maximum fracture risk for tubes hydroformed in the
same die with self feeding conditions (see Section 6.5.1).
q
CLR
ti
OD
6.10 Tube with a 90° bend.
flow from the tube edges towards the centre. With higher friction forces,
axial feed mainly causes thickening at the tube edges, but does not improve
the feasibility of the process. The tube centreline length can be measured
as the final useful tube length Lf.
Feeding potentialities also decrease drastically due to severe bends.
Nomenclature of tube bending is given in Fig. 6.10: CLR is the centreline
radius of the bend and q is the bending angle. A bend is severe either if the
bend angle is q ≥ 90 ° or if it is sharp. Sharp bends are those where the nondimensional bending factor BF is greater than about 8:
BF =
OD2
≥8
tiCLR
[6.2]
If a part has more than two severe bends (as in Fig. 6.11), axial feeding in
the central part of the tube becomes virtually impossible.
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Hydroforming for advanced manufacturing
6.11 Part shape with three sharp bends.
6.12 Part with three bulges on different positions along the centreline.
6.3.3 Bulges and protrusions
The hydroformability decreases as the number of bulges and protrusions
along the centreline increases. A bulge can be defined as a nearly axissymmetrical local expansion of the tube. A protrusion (T- or Y-shape) can
be defined as a non-axis-symmetric local radial expansion. From a formability point of view, a protrusion is more critical than a bulge. Whenever a
part presents a protrusion or a bulge, axial feeding becomes very important,
since the action of the internal pressure alone would not be enough to form
the part without premature fracture or excessive thinning. The higher the
number of local expansions, the higher the amount of feeding required. If
a part has three local expansions on different positions along the centreline
(as in Fig. 6.12), forming the central expansion becomes very difficult and
the part might be unfeasible with a conventional one-step hydroforming
process. If a part has more than three expansions placed on different axial
positions, it is almost certainly unfeasible, unless the expansion ratio or the
protrusion height is very small. In addition to their number, the position of
local protrusions and bulges is very critical as well. Indeed, even a small
number of bulges, if they are distant from the axial punches, may cause
formability problems.
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131
Analytical or empirical models may be used to approximately calculate
the maximum height of protrusions, in terms of axial force and internal
pressure. As stated before, the use of a counterpunch is generally required
to form a protrusion. The maximum useful protrusion height that can be
reached depends on the material formability and on the specific process
geometry and conditions, but only an optimal process may guarantee a
branch height greater than the outer tube diameter OD. As a general but
rough rule, a counterpunch is recommended when the height of the protrusion is greater than 0.4 ÷ 0.5 OD.
6.3.4 Formability of materials
Given a final desired tube geometry, hydroforming may be difficult with a
material with poor formability (such as several heat-treated aluminium
alloys) and very easy with a material with good formability (such as a stainless steel). In terms of quantitative parameters, the two most important
variables are the strain hardening exponent n and the normal anisotropy
coefficient r. As a general rule, the uniformity of the strain distribution
improves with increasing n-value, and the uniformity of the thickness distribution improves with increasing r-value. Common ranges of n- and rvalues are classified in Table 6.2, in order to quickly assess the formability
of a tube.
6.4
Guidelines for process design
Although the detailed planning of THF processes is carried out through
FEA, a preliminary feasibility study is often recommended or even necessary. A few analytical models and empirical rules are available in the literature for quick calculations of necessary axial feeds and pressure levels
(Ahmed, 1997; Birkert et al., 1999; Rimkus et al., 2000; Asnafi et al., 2000;
Hong et al., 1999; Thiruvarudchelvan et al., 1999; Koç, 2000a; Morphy, 1999
and Hillmann et al., 2000). A good presentation of design guidelines
for hydroforming of aluminium tubes can be found in Hoffmann et al.
(2001).
Table 6.2 Classification of formability of tubes according to n and r values
Formability classes
Poor
Average
Good
n value
r value
<0.14
<<1
0.14 ÷ 0.2
⯝1
>0.2
>>1
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General process planning
Rapid feasibility checks
maximum calibration pressure Pcal
maximum expansion ratio CSexp
Design of process sequence and layout
Rough design of pressure and axial feed
Finite element analysis
Feasibility analysis
Process optimisation
6.13 A procedure for hydroforming process planning.
However, these models are useful for simple parts or for individual THF
features such as T-shapes and axisymmetrical bulges. The complete process
planning of a complex THF part requires several steps. An example of possible planning procedure is proposed in Fig. 6.13.
6.4.1 Rapid feasibility evaluation
The rapid feasibility evaluation is an iterative process which requires several
steps, briefly described as follows. First of all, a provisional initial tube crosssection must be defined. A round tube with maximum possible outer diameter OD inscribed within the die should be initially chosen and an initial
wall thickness ti must be assumed. Then, the maximum available pressure
at the press Pmax must be checked and compared with the predicted calibration pressure Pmax ≥ Pcal. In order to precisely calculate the required calibration pressure, the exact values of wall thickness, plastic strain and flow stress
of the material that is going to fill the die corner with the minimum radius
Rmin (i.e. at the end of the process) are required. For this reason, a precise
calculation is very difficult to obtain before running an FEM simulation.
Nevertheless, a rough overestimation of the maximum calibration pressure
(Fig. 6.14) can be given by the following formula:
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Parts, process and tooling in tube hydroforming
133
Stainless
HSS
Mild steel
Al alloy
Annealed AI
400
350
300
250
200
150
100
50
0
0
0.1
0.2
0.3
0.4
Ratio ti/Rmin
0.5
0.6
6.14 Upper bound estimate of calibration pressure.
Pcal =
2 ti K (ε 0 + n)n
3 Rmin − ti / 2
[6.3]
where K, n and e0 are the flow stress parameters of the power hardening
law s– = K(e0 + e–)n and Rmin is the minimum concave corner radius found
on the die. Please note that equation 6.3 is an upper bound value. Therefore,
if this value falls within the condition Pmax ≥ Pcal, then the available press is
suitable for the given process. If the condition Pmax ≥ Pcal is largely violated,
then a decrease of the initial ti should be considered. If the condition Pmax
≥ Pcal does not hold by a small deviation, then a more accurate value of Pcal
should be searched for, before declaring the process not feasible.
The second rapid feasibility check is about the maximum tolerable expansion of the tube before bursting. The value CSexp, obtained by the empirical
equation 6.4 (instead of equation 6.1) can be checked for its maximum possible limit in order to verify whether the part is feasible or not.
⎛ Pcsmax
⎞
CSexp = ⎜
− 1⎟ ≤ CSmax = 5.6522n − 1
⎝ ODπ
⎠
[6.4]
If the calculated CSexp exceeds the limit imposed by CSmax, a new initial
tube cross-section must be defined. A non-round initial cross-section could
be assumed, or preforming/crushing operations might be considered (Fig.
6.3) and the procedure iterated until both checks of equations 6.3 and 6.4
are passed. The maximum possible expansion calculated by equation 6.4
can be reached only when axial feeding is optimally performed. When axial
feeding is not possible, i.e. when the material is at a plane strain condition,
the expansion limit is severely reduced, and can be estimated as:
⎛ Pcsmax
⎞
CSexp = ⎜
− 1⎟ ≤ CSmax = 2.7183n − 1
⎝ ODπ
⎠
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Hydroforming for advanced manufacturing
Maximum expansion CS
134
1.0
0.9
0.8
0.7
0.6
0.5
0.4
0.3
0.2
0.1
0.0
Optimal
axial
feed
No
axial
feed
0 0.05 0.1 0.15 0.2 0.25 0.3 0.35 0.4
n-value
6.15 Maximum admissible expansion CSmax for different feeding
conditions.
Upper die
Lower die
6.16 Bending performed in the die.
In Fig. 6.15, a comparison of admissible expansions for both conditions
of equations 6.4 and 6.5 is shown.
6.4.2 Design of process sequence and layout
The definition of the complete process sequence starts from the design of
the final part centreline. This curve can be seen as the tube spine and might
present one or more bends. In most cases, bends on the tube spine should
be obtained as preliminary operations, before hydroforming. In a few
exceptional situations, when bends are mild and if they all lie on a vertical
plane or if they fall within areas of expansion, they can be directly obtained
by the hydroforming process itself, as shown in Fig. 6.16. When bends are
obtained during crushing and hydroforming, obviously the die’s centreline
and initial tube’s centreline do not overlie. Once the die centreline has been
identified, the die parting surfaces can be drawn. For approximately axially
symmetrical shapes, the parting surface is easily identified as the horizontal
surface which cuts across the centreline itself. For polygonal or asymmetrical shapes of the die, the part should be horizontally rotated until a compromise is reached between two different purposes. In order to facilitate
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Parts, process and tooling in tube hydroforming
FcB =
FcA
135
2
FcA
3
Upper die
Upper die
Lower die
a
Lower die
b
6.17 Different orientations of the die parting surface generate different
expansion and different maximum closing force Fc.
crushing and to use the maximum possible initial tube diameter, the part
should be placed in order to maximise the area of the horizontal projection
of the die outline AD. Unfortunately, this can be done only for small to
medium sized parts. In fact, the maximum closing force Fc of the press is
equal to AD multiplied by the calibration pressure Pcal. A very simple
example will clarify this statement, considering a square cross-section with
3 : 2 ratio between the two sides. Formability of the first cross-section shown
in Fig. 6.17 is better since a larger initial diameter can be easily accommodated within the die. On the contrary, a much smaller initial tube can be
accommodated within the second die, but the maximum closing force
required at the press is 2/3 of the previous one.
Once the orientation and the separation surface of the die have been
designed, the location of the axial punches is straightforward. The length of
the guiding zone should be selected according to the initial length of the
tube spine. If one or more pronounced protrusions are present on the figure,
then counterpunches should be adopted.
6.4.3 Rough design of fluid pressure and axial feed
A successful THF operation is largely dependent on the proper selection
of pressure Pi and axial feed dax vs. time, i.e. the loading curves. In order to
avoid the occurrence of defects or production problems (leaking, wrinkling,
buckling, bursting) the loading paths must be carefully designed and Finite
Element Analysis (FEA) is often used to reduce the cost of the prototyping
phase. Therefore, fast and effective FEA strategies for the selection of
loading paths are strongly needed. Nonetheless, an initial guess of the
loading curves, rapidly calculated before starting the analysis, is very useful.
To this purpose, the bulging pressure can be easily calculated as:
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Hydroforming for advanced manufacturing
General description
Long parts with no bulges
or protrusions and with
sharp bends
Geometrical
features
Expansion ratio
of sharp bends
of protrusions
of bulges
spline length (Lf/OD)
<0.05
>2
0
0
>20
Parts with
Average length parts with
significant
no bulges, protrusions
expansions
or bends
or bulges
<0.1
>2
0
0
>20
<0.1
≤2
0
0
>15
<0.2
≤2
0
0
>15
Parts with
significant
protrusions
>0.2
Any*
0
>0
≈10
>0.2
Any*
≤2
≤1
<5
Average
feeding
Large
feeding
No
≈0.2
No
≈0.4
General description
No axial feeding
Suggested process
design variables
Crushing
Amount of axial feed dax/Li
Optional
0
Yes
0
Small axial feeding
Optional
≈0.05
Yes
≈0.05
6.18 Classification of THF parts and approximate prediction of
required axial feed.
Pbulging =
2ti σ 0
OD − ti
[6.6]
and the calibration pressure can be estimated by equation 6.3. Furthermore,
it may be possible to understand by large approximation the total required
amount of axial feed (dax). Using Fig. 6.18 as a classification tool, parts can
be roughly divided into four technological groups, which require different
amounts of axial feed, non-dimensionally expressed as the ratio between
the feed itself and the initial tube length dax/Li. A vast majority of typical
THF parts belongs to one of the mentioned groups. However, in some particular cases, THF parts cannot be directly classified. As an example, some
parts could not be feasible by ordinary THF operations, i.e. they might
require particular operations, such as an intermediate annealing or the use
of multiple dies. In other cases, long THF parts may have different features
along the spine, so that they could be decomposed into different modules,
each one of these modules belonging to a different technological group. Or
finally, some parts may present different features at the same location along
the spine, e.g. a bulge or protrusion located at a sharp bend.
6.5
Finite element analysis strategies for
process design
In the previous sections, guidelines and suggestions for quick feasibility
evaluation and design of a new hydroforming process have been given.
However, a proper and correct selection of the amount of left and right
axial feeds and of pressure vs. time curves can be determined only after a
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Parts, process and tooling in tube hydroforming
137
complete Finite Element Analysis (FEA) and prototyping. Clearly, a robust
strategy for designing the simulations to be run must be chosen in order to
keep total computational times under control. Different FEA strategies for
different purposes should be used. If the main goal is to obtain a feasible
solution for a process with a small process window, the following approaches
may be used: adaptive selection of loads, or ‘adaptive loading’ and the ‘selffeeding’ approach.
6.5.1 Determination of feasible solutions
The ‘self-feeding’ (SF) approach is a method designed to restrict the search
for the feasible loading path to a proper family of curves. The starting point
of the SF approach is to run an initial FEM simulation with the following
boundary conditions: no friction at the interface die/tube, no boundary
conditions on the tube edges, internal pressure Pi linearly increasing up to
the calibration pressure. As a consequence, the tube will be pulled towards
the bulging area only by the effect of internal pressure. The output of the
initial run is the amount of left and right axial feed as a function of the
internal pressure dax(Pi). This initial simulation provides a minimum required
value of axial feed and it is also useful to understand the natural balance
between left and right feed. The initial run is not supposed to be realistic
or accurate: the self-feeding loading curve is used only as a reference for
further calculations. Then, in the following simulation runs, the amount of
axial feed is progressively increased, with repeating FEM simulations, until
satisfactory results are achieved, i.e. until a feasible solution is found. The
approach is not appropriate in parts where the natural feeding is prohibited
or restricted by the geometry of the die, i.e. where the bulge area is strongly
non axisymmetrical, such as in T-shapes and Y-shapes. However, the
approach provides good results for structural and frame parts. Usually,
loading curves are obtained, that generate an unwrinkled part although
with very large thinning.
The ‘adaptive loading’ (AL) approach is based on the ability to detect/
identify the onset and growth of wrinkles during the process and promptly
react to them. Loading paths can therefore be adjusted, within the same
simulation run, to correct those defects. The ultimate goal is the selection
of a feasible part program with a minimum number of simulations, or even
within a single run. Since it was first proposed (Doege, 1998), interest is
recently growing around the AL approach, as testified by the numerous
examples of adaptive algorithms for THF that can be found in the scientific
literature in the last few years (Strano, 2004; Aydemir, 2005; Holecek, 2005;
Heo, 2006 and Gelin, 2006). The instantaneous selection of process parameters is often performed through the adaptive control theory. In other
words, a vector of control variables u(t), is optimally and adaptively selected
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Hydroforming for advanced manufacturing
Feed dax (mm)
25
20
15
10
5
0
0.02
Expansion
Calibration
0.04
0.06
0.08
0.10
0.12
Pressure (GPa)
0.14
6.19 Loading curves (feed vs. pressure) (Strano, 2001).
according to the values of a vector of state variables x(t). A dynamic model
is necessary to estimate the relation between process parameters u(t) and
state variables x(t) and an objective function is required in order to select
the optimal solution at each time step. As an example, in Fig. 6.19,
the loading curve obtained by Strano (2001) for a 1.5-mm-thick AISI 304
stainless-steel part is shown.
6.5.2 Strategies for process optimization by
finite element analysis
True process optimisation by the FEM generally requires an extensive
computational effort. A very large number of papers and publications may
be found in the literature concerning optimisation of THF processes. For
this reason only a few citations will be given as examples. Two main different strategies may be used: sequential optimisation and optimisation based
on a plan of computer experiments. The first approach is generally used in
combination with gradient-based optimisation techniques and is more
suited to local optimisation, i.e. when a sufficiently good and reliable initial
solution is available. Generally, sequential optimisation techniques are best
suited for the geometric design of tools and dies. However, they can also
be applied for the selection of the loading paths, i.e. for the optimisation of
time-dependent process control variables (velocities, forces, pressures, etc.).
One of the most convincing examples is given in Yang (2001). Another
example is given in Fig. 6.20, where the optimal loading path obtained by
Fann (2003) for a 2.3-mm-thick AISI 304 stainless-steel part is shown.
The second approach (Koç, 2000b) is more massive and generally requires
longer computational times, but it is more suited to global optimisation on
a larger range of the input variables. When a planned batch of simulation
is run, the results are usually expressed in terms of metamodels (Rijpkema,
2001). A metamodel is a mathematical relation between one quantitative
performance of the process and one or more process parameters. The model
which is most commonly employed is the response surface method (RSM),
in combination with DOE (Design of Experiments) techniques. Recently,
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Parts, process and tooling in tube hydroforming
200
150
100
30 mm
f 61 mm
15 mm
5
10
5 mm
f 61 mm
Internal pressure (MPa)
250
139
50
0
0
15
20
25
30
Axial feeding (mm)
6.20 Loading curves (pressure vs. feed) (Fann, 2003).
ss
cro
are ners
u
q
h s d cor
wit
ge ounde
l
u
b
r
tral and
Cen ction
e
s
oss
cr
nd
Rou ection
s
6.21 Example of a tube hydroforming final desired shape.
interpolation models are increasingly being used, such as the ‘kriging’
approximation, in combination with DACE (Design and Analysis of Computer Experiments) techniques.
A combination of planned and sequential optimisation may be used as
well, and this is probably the most efficient way of obtaining an optimal
solution (Kirby, 2005).
6.6
Designing a new hydroforming process:
a simple example
In this section, a very simple hydroforming example is synthetically illustrated, in order to show an implementation of the design guidelines discussed in Sections 6.4 and 6.5. Consider the hydroforming of the shape
shown in Fig. 6.21 with the data given in Table 6.3. The main design steps
and decisions are synthetically described in Table 6.4. The outcome of the
design process can be observed in Fig. 6.22, which shows 1/8 of the hydroforming die and Fig. 6.23 where the loading curves are plotted.
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Hydroforming for advanced manufacturing
Table 6.3 Data used for the analysis of the part in Fig. 6.21
Property
Description
Values
Isotropic hardening
law
Formability of
material
Friction coefficient
Tube initial
thickness
Useful tube length
Shape of final part
s = K(e + e0)n
K = 562 MPa, n = 0.22, e0 = 0.01
See Table 6.2
Good
Coulomb
–
f = 0.05
ti = 2 mm
–
Diameter of round
cross-section
Side of square crosssection
Maximum expansion
Minimum radius
Lf = 200.7 mm
50.3 mm
Shape of final part
Shape of final part
Shape of final part
59.1 mm
CSexp = 42.4%
Rmin = 5 mm
Table 6.4 Sequence of steps and decisions in the design of the THF process for
the part in Fig. 6.21
1 Rapid feasibility evaluation
1.1 General
The maximum expansion ratio of the die is CSexp
considerations
= 42.4% (calculated before assuming the initial
tube shape and diameter) which is not a critical
value when axial feeding is possible, since the
material shows a good formability
Equation 6.3 yields Pcal = 235 MPa, which is
1.2 Maximum
calibration
reasonably reached by any commercial
pressure
hydroforming press
1.3 Maximum
Using equation 6.4, CSexp = 42.7% results lower
expansion ratio
than the CSmax = 46.4%
As a result of steps 1.1, 1.2 and 1.3, the process is feasible.
2 Design of process sequence and tooling layout
2.1 Prebending
The final part has a straight spine, therefore no
pre-bending is required.
2.2 Preforming
The part shows round cross-sections at both ends,
therefore the guiding zones can be round, the
most reasonable initial tube cross-section is
round with outer diameter OD = 50.2 mm and
no crushing or preforming are needed.
2.3 Dies parting line
In order to minimise the closing force Fc, the die
should be split as in Fig. 6.22.
2.4 Length of initial
As a first approximation, the total length of the
tube
guiding zones could be about half the final
length of spine Lf = 200.7, yielding the initial
length Li = 300 mm.
As a result of steps 2.1, 2.2, 2.3 and 2.4, the hydroforming die can be assumed
to be as in Fig. 6.22.
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Parts, process and tooling in tube hydroforming
141
Table 6.4 Cont’d
3 Rough design of fluid pressure and axial feed curves
3.1 Bulging
Using equation 6.6, the provisional bulging
pressure
pressure can be set equal to Pbulging = 16.9 MPa.
3.2 Axial feed
Since the part is symmetric, left and right feed are
equal; using Fig. 6.18, the part belongs to the
fifth column and the estimated amount of the
sum of left and right feeds is dax = 0.2 Li =
60 mm (30 mm for each side)
As a result of steps 1.2, 3.1 and 3.2, an initial guess for the loading curves can
be obtained (see Fig. 6.23)
4 Process design by FEA
4.1 Adaptive
loading
The strategy of adaptive loading (AL) by FEA is
used in this case in order to quickly obtain a
feasible, sub-optimal solution.
As a result of step 4.1, the AL loading curves can be obtained (see Fig. 6.23)
Symmetry
plane
Guiding
zone
Parting surface
and
symmetry plane
Symmetry plane
6.22 Die divisions and layout, for example, THF process.
Internal pressure
(MPa)
250
200
150
100
Initial guess
Curves obtained by AL
50
0
0
10
20
30
40
50
60
70
80
90
Total axial feed (mm)
6.23 Loading curves obtained by the preliminary studies and by
adaptive loading.
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Hydroforming for advanced manufacturing
6.7
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hoffmann a and birkert a (2001), ‘Design Guidelines for Hydroformed Structural
Components of Aluminium. Hydroforming of Tubes’, in Siegert K Extrusions and
Sheet Metals, Stuttgart, Technische Universität, 2.
holecek m and hora p (2005), ‘Numerical self-regulation of time-dependent parameters in tube hydroforming processes’, in Numisheet 2005, A, 601–606.
hong s z, wang g x and zeng z p (1999), ‘Technological analysis and engineering
calculation on hydraulic extrusion-bulge forging of multi-way tube joint’, in Proc.
of the 6th int conf Advanced Technology of Plasticity, II.
kirby d, roy s and kunju r (2005), ‘Optimization of tube hydroforming with consideration of manufacturing effects on structural performance’, in Numisheet
2005, A, 585–590.
koç m (2000), ‘Use of Guidelines, Analytical Methods and FEA During Development of a Hydroform Part – Concept to Production’, in Proceedings of Innovations In Tube Hydroforming Technology.
koc m, allen t, jiratheranat s and altan t (2000), ‘The use of FEA and design of
experiments to establish design guidelines for simple hydroformed parts’, Int J
Mach Tools Manufact, 40, 2249–2266.
kocanda a and sadlowska h (2006), ‘An approach to process limitations in hydroforming of X-joints, as based on formability evaluation’, J Mater Process Technol,
177, 663–667.
morphy g (1999), ‘Product design for hydroformed structural components’, in TPA
4th Annual Automotive Tube Conference Focus on Hydroforming, Dearborn,
Michigan.
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rijpkema j j m, etman l f p and schoofs a j g (2001), ‘Use of Design Sensitivity
Information in Response, Surface and Kriging Metamodels’, Optimization and
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strano m and altan t (2001), ‘FEA Simulation Strategies for Tube Hydroforming’,
5th Aitem Conference, Bari, Italy, II, 453–462.
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thiruvarudchelvan s and lewis w (1999), ‘A note on hydroforming with constant
fluid pressure’. J Mater Process Technol, 88(1/3), 51–56.
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7
Tribological aspects in hydroforming
G R A C I O U S N G A I L E, North Carolina State University, USA
7.1
Introduction
Better understanding of tribological conditions in hydroforming is extremely
important since the tribological factors highly influence part failure or
success of a hydroformed part. Throughout this chapter, tribological aspects
of hydroforming will refer to lubrication at the tool–workpiece interface,
wear of the dies, resulting surface morphology of the hydroformed part and
the interface friction. Emphasis will be given to tube hydroforming.
7.1.1 Die–workpiece interface and material flow
The schematic of tube hydroforming (THF) and sheet hydroforming (SHF)
variants are shown in Fig. 7.1 and 7.2, respectively. As the tube material
flows to the die cavity due to axial feeding and fluid pressure Pi a pressure
Pe is built up at the interface and a shear stress ts is generated due to friction. Depending on part geometry and variation in the fluid pressure, the
interface pressure (Pe) and shear stress will vary locally along the interface.
To obtain a sound THF part, an optimal combination of Pi(t) V(t), material
characteristics(flow stress), and tribological conditions need to be achieved.
Similarly, with SHF, the same variables will be needed; however, in the SHF
instead of V(t) a blank holder force, BHF, is used.
Figures 7.1 and 7.2 depict the interface at a macro-scale level. As discussed
later in Section 7.3, at a micro-scale level, other tribological variables become
critical in achieving optimal conditions. These variables include: surface
roughness of the deforming material and the die, surface roughening behaviour of materials, oxide layers on the deforming materials, and the change
in the surface chemical properties of the die and the deforming material.
7.1.2 Friction laws and applicability in hydroforming
There are two friction laws that are commonly used in metal forming:
Coulomb’s friction law and the shear friction law. In Coulomb’s law the
144
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Pe
ts
Pe
Pe
ts
ts
V1
V2
Pi Pf(t)
ts
Pe
7.1 Die–workpiece interface variables at macro-scale for THF.
BHF
BHF
Fp
V
P
V
P
ts
ts
Pi Pf(t)
a
BHF
BHF
Pi Pf(t)
ts
ts
Pe
Pe
b
7.2 Die–workpiece interface at macro scale for SHF:
a hydromechanical deep drawing and b conventional
sheet hydroforming.
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friction stress t is proportional to the interface pressure Pi and the constant
of proportionality is the friction coefficient.
t = mP
[7.1]
The shear friction law assumes that the friction stress τ is proportional
to the shear strength of the deforming material, k.
τ = mk =
m
3
σ
[7.2]
where s is the flow stress and m is a constant which varies from 0 to 1. The
essence of this law lies in the comparison of the shear strength of the lubricant to that of the deforming material, e.g., if the shear strength of the
lubricant is equal to that of the deforming material, then m = 1.
While Coulomb’s law is more suitable for sliding contact with little deformation, the shear friction law is suitable for sliding objects subjected to bulk
plastic deformation. Modes of deformation in sheet hydroforming are dominated by sliding and bending, hence, Coulomb’s law may be applicable.
Depending on the severity of deformation, tube hydroforming processes
may fall in between the two friction laws. Therefore, care should be exercised as to which friction law should be adopted.
Since there are numerous variables acting at the tool–workpiece interface, the frictional laws discussed above do not accurately prescribe the
actual interface friction in metal forming. As expressed in equations 7.1
and 7.2, Coulomb’s law and the shear law associate shear stress as a function of the interface pressure and shear strength of the deforming material,
respectively. These laws do not account for: (a) surface roughness of the
mating surfaces, (b) material-dependent surface roughening mechanisms,
(c) micro-opening at the interface, (d) actual die–workpiece contact area,
and (e) thermal induced variables. While these variables are usually lumped
together to account for friction behaviour of the lubricant used, a clear
understanding of the implication of these variables on hydroforming is
imperative.
7.1.3 Typical tribological failures in hydroforming
Failures in hydroforming can be categorized into: wrinkling, buckling, bursting, and unacceptable part surface quality. These failures are caused by
inability to achieve an optimal combination of a multitude of variables,
including pressure loading paths, material feed path, material deforming
characteristics, and tribological conditions. Ineffective lubrication that will
cause excessive friction stress above a predetermined limit value will result
in wrinkling, bursting, or bad surface quality. Figure 7.3 shows typical tribological failures in hydroforming. An effective lubricant should therefore be
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147
7.3 Example of tribological failures.
chosen for a specific hydroforming process or for a family of hydroforming
processes with similar characteristics.
7.1.4 Attributes of hydroforming lubricants
A hydroforming lubricant is expected to satisfy a number of different functions. The major functions are listed below.
1
Reduce friction at the tool–workpiece interface: low friction reduces
forces and power requirement for the feed actuators, press ram,
pressure intensifier systems, etc. With low friction, more homogeneous
deformation can be achieved. For a hydroforming process with
narrow process window, a slight increase in friction can result in a nonproducible part.
2 Separation of surfaces and ability to quickly respond to new surface
generation: the lubricant should prevent metal-to-metal contact. As new
surfaces are generated, the lubricant should rapidly respond to formation of a lubricant film [1].
3 Adaptability to varied conditions: in hydroforming, the interface pressure, sliding velocity, surface temperature, and material deformation
modes change in time and space due to typical varying loading paths.
The lubricant should be functional under these conditions.
4 Compatibility with materials, dies, and secondary operations: the lubricant should perform efficiently in the system defined by the die, the new
and old surfaces [1]. The lubricant should not cause chemical or metallurgical changes in the product. Secondary process such as welding,
painting, annealing should not be affected.
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5
Part surface quality and die wear: the lubrication mechanism should
produce an acceptable surface finish on the hydroformed part. The
lubricant should prevent metal pick-up and reduce die wear.
6 Stability and reactivity: the lubricant should not be affected by temperature, oxidation or bacteriological attach. The lubricant should not be
corrosive to the products or the tools. Any chemical reaction or changes
should be beneficial to reducing friction stress [1].
7 Application and removal: the lubricant should be easy to apply and
remove. The lubricant should be environmentally friendly, i.e., it should
not cause skin irritation, and it should not be toxic or produce strong
odour.
8 Cost effectiveness: in addition to the above attributes, the lubricant for
hydroforming should be cost effective. The cost should be weighed in
accordance with other cost savings the lubricant may bring about to the
process.
7.2
Parameters that influence friction, lubrication,
and wear
There is a multitude of variables inherent in a hydroforming process that
can significantly influence friction, lubrication, and wear rate. These variables can be subdivided into four groups, process parameters, tool parameters, workpiece parameters, and lubricant parameters (Fig. 7.4). A change
in any of the parameters associated with the process, the tool, or the workpiece directly influences the performance of the lubricant and similarly the
wearing characteristic of the tooling; some of these are now briefly
discussed.
Interface pressure
A change in the interface pressure results in a change in friction stress
unless the interface is composed of an ideal lubricant with zero coefficient
of friction. The change in pressure also leads to change in the viscosity of
the lubricant, surface contact morphology, lubricant film thickness, etc.
Figure 7.5 shows possible scenarios when interface pressure is increased.
Depending on the prevailing tribological conditions, friction can either
increase or decrease.
Surface expansion/contraction
Deforming material at the tool–workpiece interface can undergo surface
expansion or surface contraction. Surface expansion leads to generation of
virgin surfaces, change in surface morphology, etc. The resulting effect of
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ra
Lu
br i
rs
ete
ca
nt
• Viscosity
es
s
• Temperature
s
ter
• Composition
• Sliding velocity • Film thickness
• Interface pressure • Pressure stability
Pro
c
149
me
ra
pa
pa
Tribological aspects in hydroforming
• Surface expansion • Compatibility with
• Surface contraction secondary operation
• Size; micro/meso/macro • Die material properties
• Surface topography
• Die surface roughness
pa
me
ra
• Die geometry
To
ol
p
• Geometry
am
ar
s
ter
• Material properties • Die surface coatings
ete
Wo
rk
rs
pie
c
e
7.4 Parameters that influence friction lubrication and wear.
Increase
pressure
Decrease
lubricant
pocket
volume
Increase
hydrostatic
pressure
Friction
decreases
Increase
viscosity
Increase
contact area
Increase
hydrodynamic
and hydrostatic
pressure
Friction
increases
Friction
decreases
7.5 Effect of increased interface pressure on friction [2].
surface expansion is thinning of the lubricant film, lubricant film break
down, and may accelerate chemical reactivity of the lubricant within the
new generated surfaces. On the other hand, the surface contraction may
cause lubricant film delamination or change the surface pocket volume size
and morphology.
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Material properties
The flow characteristic of deforming material can be expressed by equation
7.3 where C is the strength coefficient, e effective strain, s flow stress, n
strain hardening coefficient, and m strain rate sensitivity coefficient. Of
particular importance to tribology is the flow stress of the surface or subsurface of the deforming material, which will change with surface strain and
surface strain rate. For higher n and m the surface will strain harden significantly, causing the surface to be hard and ultimately affecting the lubricant
performance. Another factor that is related to material properties is the
surface roughening behaviour. It has been shown that surface roughening
rate is strongly dependent upon number of available slip planes within the
crystal structure [3]. Therefore, the deforming material make-up, FCC, BCC,
HCP and grain size also result in different roughening characteristics, which
may affect the performance of the lubricant.
.
s– = C e– n e– m
[7.3]
Sliding velocity
The sliding velocity at the tool–workpiece interface will have influence on
the lubrication conditions. Due to relative sliding between the die and the
deforming material friction, heat will be generated as expressed by equation
7.4 [4].
T=
Af σ vF Δt
cρVa
[7.4]
where A is the conversion factor between mechanical and thermal energies, f the friction factor, V the sliding velocity, F the surface area at the
die–workpice interface, c the specific heat of deforming material, r the
specific weight of the deforming material, and Va the volume which is
subjected to an increase in temperature. As a result of the increase in
temperature, lubricant rheology will be changed, flow stress of the surface
asperities and sub-surface material will be decreased, and chemical reactivity between the deforming material and the lubricant will increase. Figure
7.6 shows some of the tribological changes that may occur with increase
in sliding velocity.
Figure 7.6 demonstrates that, depending on the prevailing tribological
conditions, friction can either increase or decrease when sliding velocity
increases. From the above discussion we see that numerous tribovariables
interact at the tool–workpiece interface in a complex manner. Thus, development of an effective lubrication system for hydroforming requires careful
interpretation and quantification of the tribovariables and identifying possible active friction zones of the hydroforming process in question.
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151
Increase
velocity
Increase heat
generated
Increase
hydrodynamic
pressure
Friction
decreases
Decrease
material
flow stress
Increase
contact area
Friction
increases
Decrease
viscosity
Decrease
hydrodynamic
and hydrostatic
pressure
Increase
chemical
reactions
Friction
decreases
Friction
increases
7.6 Effects of increased sliding velocity on friction.
Die
Axial force
Axial force
Fluid
inlet
Expansion
Guiding Transition
zone
zone
zone
7.7 Friction zones in THF.
7.2.1 Friction zones and characteristics
Tube hydroforming friction zones
In a typical THF process, three different friction zones are observed:
(a) guiding zone, (b) transition zone and (c) expansion zone (Fig. 7.7) [5,
6]. In the guiding zone, the workpiece and the die are in full contact during
the process and material deformation is limited to wall thickening. However,
the material is pushed into the deformation zone by means of axial cylinders. Most THF processes require the material to be fed towards the expansion zone. Therefore, depending on the geometry of the part, a high relative
sliding velocity (50 to 100 mm/s) between the workpiece and die may be
present [5]. At a macroscopic scale, the dominant state of stress in this zone
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Hydroforming for advanced manufacturing
is such that sr, sa, and sh are all in compression. Because of the compressive
state of the surface, the deforming tube undergoes surface contraction.
On the other hand, the state of stress at the expansion zone is predominantly tensile, sa and sh are in tension mode. In this friction zone, the sliding
velocity reaches a minimum value while the surface expansion and the
surface strains attain maximum values. Because of the difference in the local
state of stress and material deformation modes in the friction zones, the
tribological conditions will differ as well, leading to different lubrication
mechanisms. This poses a challenge in developing effective lubricants/or
lubrication systems that will satisfy all friction zones.
Sheet hydroforming friction zones
Figure 7.8 below shows friction zones (I, II, III, and IV) that can be identified in sheet hydroforming. The resulting state of stress at friction zone I
(flange) as the fluid pressure, Pi, pushes the blank to the die cavity is such
that s1, and s2, are in compression and s3 is in tension. The radial tensile
stresses lead to compressive hoop stress (s2). The hoop stresses will tend to
cause wrinkling if a blank holder force (BHF) is below a certain limit. In
friction zone II, the deforming material is subjected to bending and unbending stress modes. Friction zone III (wall) is subjected to tensile loading
and compression in the thickness direction caused by fluid pressure, Pi. In
s1
s2
s3
Flange
BHF
BHF
s1
Pi Pf(t)
I
II
III
Wall
Draw
pressure
Base
IV
ts
s1
ts
Pe
Pe
s2
s3
7.8 Friction zones in SHF.
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Tribological aspects in hydroforming
153
friction zone IV the material undergoes bending stress modes. Unlike the
deep drawing process where a punch is used, with hydroforming of the
variant shown in Fig. 7.7 the friction stress will be acting on the die
profile.
7.2.2 Effect of loading paths on tribological conditions
Figure 7.9 shows an example of a Y-shape tube hydroforming component
with a typical loading path (pressure vs time and feed vs time). It can be
seen that, at any time t, the local friction stress ts at the tube–die interface
will acquire a different value due to variation in the Pi and feed as a function of process time t. This scenario calls for better understanding of the
lubrication mechanisms under these conditions and determination of a
lubrication system that has the ability to maintain an effective lubricant film
under rapidly varying tribological conditions.
Pe
ts
Pe
Pe
ts
V1
ts
V2
Pi Pf(t)
ts
Pe
fv1
Axial feed
Fluid pressure (Pi)
Pi
fv2
Process time (t)
7.9 Influence of loading path on interface friction.
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7.3
Lubrication mechanisms
There are four basic types of lubrication that govern the frictional conditions in metal-forming [1, 4]. The Stribeck curve shown in Fig. 7.10 illustrates
the onset of these various types of lubrication as a function of the combination of lubricant viscosity h, sliding velocity v and normal pressure p.
Dry
m>0.3
Dry
Boundary
Coefficient of friction
2
Under dry conditions, no lubricant is present at the interface and only
the oxide layers present on the die and workpiece materials may act as
a ‘separating’ layer. In this case, friction is high, and such a situation is
desirable in only a few selected forming operations, such as hot rolling
of plates.
Boundary lubrication is governed by thin films (typically organic)
physically adsorbed or chemically adhered to the metal surface. These
films provide a barrier under conditions of large metal-to-metal contact
where the properties of the bulk lubricant have no effect. As is the case
with dry conditions, friction is high.
Mixed
hydrodynamic
hv
m
Film thickness
1
Boundary lubrication
0.1<m<0.3
Mixed layer lubrication
0.03<m<0.1
Hydrodynamic or full
film lubrication
m<0.03
hv
m
7.10 Stribeck curve showing onset of various lubrication mechanisms
(adapted from reference 1).
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3
Full film lubrication exists when a thick layer of solid lubricant/dry
coating is present between the dies and the workpiece. In this case, the
friction conditions are governed by the shear strength of the lubricant
film. Hydrodynamic conditions exist when a thick layer of liquid lubricant is present between the dies and the workpiece. In this case, the
friction conditions are governed by the viscosity of the lubricant and by
the relative velocity between the die and the workpiece. Low friction is
always encountered under full film lubrication.
4 Mixed-layer lubrication is the most widely encountered situation in
metal-forming. Because of the high pressures and low sliding velocities
encountered in most metal-forming operations, hydrodynamic conditions cannot be maintained. In this case, the peaks of the metal surface
experience boundary lubrication conditions and the valleys of the metal
surface become filled with the liquid lubricant. Thus, many liquid lubricants contain organics that will adsorb on, or chemically react with, the
metal surface in order to help provide a barrier against metal-to-metal
contact. If there is enough lubricant present, the lubricant in the valleys
of the metal surface can act as a hydrostatic medium. In this case, both
the contacting peaks of the metal surface and the hydrostatic pockets
support the normal pressure. Thus, friction is moderate.
7.3.1 Lubrication mechanisms intube hydroforming
There are different classifications of tube hydroforming that have been
adopted [7]. In this chapter, however, tube hydroforming is classified into
two groups: feed-driven THF and expansion-driven THF. This classification
provides clear distinction on which lubricant mechanisms are predominant
in one class of THF over the other. Feed-driven THF refers to a hydroformed part that requires a considerable amount of material feeding such
as Y shapes and T shapes. In contrast, the expansion-driven THF refers to
hydroformed parts that do not require significant feeding of material from
the tube ends, rather the dominant mode of deformation is expansion.
Typical examples are engine cradles, chassis, etc.
7.3.2 Feed-driven tube hydroforming
The boundary conditions at the guiding zone are such that the lubricant
captured at the interface has little room to escape from the interface. The
pressure gradient given in Fig. 7.11 shows that, if a certain amount of trapped
lubricant is to escape, it will only flow towards the transition zone. If favourable process conditions that prevent the lubricant from escaping the guiding
zone are exhibited at the tool–workpiece interface, it is possible to achieve
hydrodynamic lubrication as long as sufficient amount of lubricant (liquid)
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Pe
Fa
Punch
Pi
Die
t
Pe
7.11 Boundary conditions at the guiding zone.
is trapped at the interface. As will be discussed in Section 7.4.2, experiments
carried out by several researchers on lubricant evaluation using guiding
zone tests have shown that very low coefficient of friction values of the
order of m = 0.02 or lower may occur [8]. These values fall under the hydrodynamic lubrication regime. It should be noted, however, that after the
material has passed the guiding zone, containment of lubricant within the
surface pockets becomes difficult.
On the micro-scale, the mechanisms of lubricant transfer through surface
topography as the material fills the die cavity are extremely complex. The
surface topography changes as a function of loading path, type of material,
type of lubricant, etc. Figure 7.12 shows the surface evolution of the SS 304
Y-shaped part. The figure shows different surface topographies along the
feed zones and along the expansion zone.
In order to understand possible lubrication mechanisms at the guiding
zone, we can treat microscopic surfaces as composed of surface asperities,
which form peaks and valleys. The surface valleys can act as lubricant retainers. Figure 7.13 shows two different scenarios under which boundary and
hydrodynamic lubrication can be achieved. Studies in sheet metal forming
have shown that four lubrication mechanisms, namely mixed layer (hydrostatic), boundary, Micro Plasto HydroStatic and Micro Plasto HydroDynamic lubrication (MPHSL and MPHDL, respectively) mechanisms can
exist at the tool–workpiece interface [10–13]. These lubrication mechanisms
are also observed in THF, though under different stress state conditions
from those observed in sheet-metal forming [14, 15].
Figure 7.13a, shows a guiding zone element of the cross-section of a
typical hydroforming operation where a tube is pressurized by internal
pressure P and fed through a die at a velocity V. For boundary lubrication,
the lubricant pressure q will be equal to zero (Fig. 7.13a). The boundary
lubrication mechanism will occur if the normal force acting between
the tube and the die (FN) is carried only by the asperities. The boundary
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157
~100μm
~100μm
7.12 Surface evolution as the tube is fed into the die cavity [9].
lubrication mechanism can be encountered in THF if the interface pressure
is too low to generate a hydrostatic effect. This may be the result of insufficient trapped lubricant or lubricant failure.
The mixed layer (hydrostatic) lubrication mechanism can occur at the
tube–die interface when the internal pressure P is large enough to generate
a hydrostatic effect. Then the total load (FN) is carried by both the asperity
contacts and trapped lubricant, thus friction is reduced as compared with
boundary lubrication because the asperities are only forced to carry a fraction of the total load (Fig. 7.13b). The MPHDL and MPHSL mechanisms
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Y
P0
t
DIE
P0
t
Tube
st
st
P
Δq
X
a
P0
P0
t
DIE
t
q
q
Tube
st
P
Δq
st
X
b
Tool
V
Tube
c
Tool
V
Tube
d
7.13 Lubrication mechanisms in THF–guiding zone: a boundary
lubrication, b hydrostatic/hydrodynamic lubrication, c microplasto
hydrostatic lubrication, d microplasto hydrodynamic lubrication.
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159
can occur as the tube slides through the die in the guiding zone. The
MPHSL mechanism occurs if the lubricant pressure q becomes greater
than the contact pressure P0 at the leading edge of the pocket. In this case,
lubricant escapes from the pocket in the forward direction (same direction
as sliding) until q becomes equal to P0 (Fig. 7.13c). These pressure gradients
can occur as a result of non-uniformly distributed shear stresses. The
MPHDL mechanism occurs if the trapped lubricant permeates to the interface area where boundary lubrication was present as a result of the relative
sliding between the tube and die (Fig. 7.13d). In this case, lubricant escapes
from the pocket in the backward direction (opposite to the direction of
sliding) [13].
MPHSL and MPHDL mechanisms reduce friction by reducing the shear
stress in the asperity ‘contact’ by interposing a liquid film. For example,
MPHDL occurs if the trapped lubricant permeates to the interface area
where boundary lubrication was present. In this case, the lubricant pressure
q becomes equal to P0. The MPHDL reduces friction by reducing the shear
stress in the asperity ‘contact’ by interposing a liquid film. The friction coefficient under this mechanism is given by:
n
n
n
τ i[ re] ⎛⎜ ∑ Ai[ re] − ∑ Δ Ai[ hdp] ⎞⎟
∑
⎝ i =1
⎠
i =1
i =1
μN = n
n
n
n
n
n
Pi[ re] ⎛⎜ ∑ Ai[ re] − ∑ Δ Ai[ hdp] ⎞⎟ + ∑ qi[ hsp] ⎛⎜ ∑ Ai[ hsp] + ∑ Δ Ai[ hdp] ⎞⎟
∑
⎝ i =1
⎠ i =1
⎝ i −1
⎠
i =1
i =1
i =1
[7.6]
where ΔAi[hdp] is the local interface area occupied by lubricant due to penetration of trapped lubricant to real area of contact (hydrodynamic effect).
n
n
i =1
i =1
When ∑ ΔAi[ hdp] ≅ ∑ Ai[ re] , full film hydrodynamic lubrication is achieved.
The friction shear stress at the interface is equal to the shear stress of the
lubricant. The degree of escape depends on lubricant viscosity h (h as a
function pressure and temperature), relative velocity between the tube and
die, and the geometry of the lubricant pocket [16, 17]. In THF, the material
flow and state of stress (tangential stress st, radial stress sr and axial stress
sa) may alter the geometry of the lubricant pocket. This can either speed
up or slow down the occurrence of MPHSL and MPHDL.
7.3.3 Expansion-driven tube hydroforming
Contrary to the guiding zone, the boundary conditions at the expansion
zone are such that the lubricant trapped at the interface can escape from
the interface due to multiple die–workpiece openings as the tube gradually
establishes contact with the die as pressure increases (Fig. 7.14). This,
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Expansion zone
boundary
Die
ΔL
Element A
Pi
7.14 Boundary conditions at the expansion zone.
Z′
Y
s
a
q
q
p
ex
V
be
Tu
P
i
q
DI
E
t
P
o
wa
ll
q
t
P
o
Y′
σ
a
rexp,i + Δr
rexp,i
ΔL
7.15 Element A from Fig. 7.14.
however, will depend on the geometry of the part in question and whether
liquid or dry film lubricant is used.
Figure 7.14 shows a typical expansion zone with die–workpiece contact
and non-contact boundaries. Microscopically, for a tube material element
with lubricant surface micro pockets at the vicinity of the contact and noncontact region, Fig. 7.14 and 7.15, the following is observed: (a) as the tube
element contacts the die some lubricant will be squeezed out from micro
pockets, (b) because of tensile loading acting on the element sa, the lubri-
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~100μm
~100μm
7.16 Surface evolution for full expansion process.
cant micro pocket will enlarge, (c) the relative sliding velocity between the
tube element and the die, Vexp, will drop due to change in Δr, i.e. an increase
in the instantaneous radius of the expanding tube towards the expansion
zone. These factors diminish the chances for MPHS and MPHD lubrication
to occur. Experimental studies have shown that coating-based lubricants
are more effective in the expansion zone than liquid lubricants. Figure 7.16
shows surface evolution for fully expansion process. The surface topography
on the deformed regions does not indicate the presence of micro-pockets
that can trap lubricant.
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7.3.4 Lubrication mechanisms in sheet hydroforming
Figure 7.17 shows the stress and deformation conditions in SHF at three
element locations (Element A – binder, Element B – draw ring radius,
Element C – punch wall). The figure depicts the interaction of a multitude
of process variables as an infinitesimal blank element is drawn from the
binder to the cavity. The element undergoes an instantaneous change of
state of stress as it passes from location A to B to C. Surface asperities,
which are present on both sides of the blank, act as lubricant reservoirs. The
asperities define the lubrication mechanism when subjected to binder
loading (thermal and mechanical) and draw velocity. Typical lubrication
mechanisms exhibited are: (a) boundary, (b) microplasto hydrostatic and
(c) microplasto hydrodynamic. The local coefficient of friction in the binder
region is a function of contact pressure (P), viscosity (u) and sliding velocity.
Element B, located at the draw ring radius, is subject to bulging due to the
FBHF
P
Element A,
binder
sr
sr
P
Element B
A
B
sr
sr
C
P
Pf
sz
Element C,
punch wall
Pf
sz
7.17 Boundary conditions for SHF.
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163
presence of the pressurized fluid. This bulging virtually eliminates the friction force at the draw ring but makes the sheet susceptible to bursting if
the fluid pressure exceeds a critical value. The coefficient of friction at the
punch wall in element C is a function of contact pressure, viscosity and
sliding velocity.
With the hydromechanical deep drawing variant of SHF, right at the start
of the process, the bottom part of the blank rapidly generates contact with
the punch due to fluid pressure P. Thus, the trapped lubricant at any time
t can hardly escape from punch–workpiece interface. This scenario promotes the occurrence of hydrostatic and hydrodynamic lubrication.
7.4
Development and evaluation of
hydroforming lubricants
As discussed in Section 7.3, different lubrication mechanisms can be
observed in hydroforming depending on whether the specific part in consideration belongs to feed-driven THF or expansion-driven THF. An effective lubricant can therefore be developed or chosen by carefully assessing
plausible lubrication mechanisms that may be induced at the tool–
workpiece interface. Generally, metal-forming lubricants are classified into
several groups as shown in Fig. 7.18. Most of the lubricants are compounded
with additives to improve performance.
7.4.1 Lubricant formulation
Since hydroforming is a relatively new field, lubricant developers have
been using experience gained in lubricant formulation from sheet
Metalworking lubricants
Liquid lubricant
Synthetic
Coating and carriers
Conversion coating
Semi-synthetic
Polymer coating
Soluble oil
Glass coating
Straight oil
Bonded coating
7.18 Classification of metal-forming lubricants.
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metal-forming processes. The choice of base stock and additives is, however,
based on major process variables as discussed in Section 7.2 as well as
anticipated lubrication mechanisms that may be favourable for specific
families of hydroforming process such as feed-driven or expansion-driven
hydroforming.
One of the challenges in the development of lubricant for hydroforming
is the development and/or adoption of multiple additives which upon mixing
with base stock will result in an effective compound that meets environmental regulations and toxicity levels, and still maintains desirable lubricant
functional needs [18]. Figure 7.19 shows typical lubricant base materials and
additives used in formulating lubricants for metal forming. As can be seen
from the large number of chemical compounds involved, to develop a costeffective product, the formulation process will require a carefully planned
iterative process. Usually, a new lubricant passes through rigorous screening
using tribo tests before being commercialized.
7.4.2 Lubricant evaluation methods
After lubricant candidates have been formulated, they undergo rigorous
screening for basic and advanced functionalities using bench and simulative
tests. The results from the bench test are an initial performance indicator
and are used to reduce a large number of lubricant candidates down to two
or three [19, 20]. Typical bench tests used by formulators are the twist compression test and the four-ball friction test. Since the bench tests do not
involve significant plastic deformation of the substrate, nor do they resemble hydroforming processes, it is imperative to further test the best lubricant
candidates using simulative tests. To date, few simulative tribo tests that
mimic hydroforming processes have been developed for the guiding zone
and expansion zone [6, 20, 21].
Guiding zone tribo tests
There are four variations of simulative tests developed to date for the
guiding zone. All these tests are based on pressurizing the tubular specimen
to the required pressure level and push the tube through a cylindrical die.
Figure 7.20 shows the sketches for the four variants.
Variant I was originally developed at the University of Darmstadt in collaboration with Schuler [6]. In this test, the normal load is measured via a
load cell connected to the upper half of the die. Assuming Coulomb’s law,
the interface friction coefficient can be determined by equation 7.7. Friction
force Ff and normal load Fc are both measured by load cells connected to
the system.
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165
Base material and carriers
Synthetic esters
- methyl esters
- polyesters
Vegetable oil
- seed oils
Petroleum
distillates
Synthetic
hydrocarbon
Mineral
- paraffinic
- naphthenic
Water
Additives
Lubricity
Emulsion
Boundary, mixed film
- fatty acid scaps, esters
- alcohols, polyamers
Others
Anianic
- sulfonates, sulfates
- carboxylates,
- phosphate
Corrosion inhibitors
- sulfonates, borates,
- carboxylates, phosphate
- alkanolamines
Extreme pressure
Reactive
- chlorine, sulfur,
- phoshorous
Non-ionic
- alcohol ethoxylates,
- amides
Metal passivators
- trioxides, thiazoles,
- thiadiazoles
Alkalinity
- alkanolamines
- metal hydroxide
Inert solid lubricants
- calcium carbonate
- mica, talc, graphite
- molybdenum disulfide
Antimist, thickener
- polymer
Dispersant
Overbased calcium
and sodium sulfonates
Antioxidant
- hindered phenols
- aromatic amines
Antimicrobial
Dyes, fragrances
Couplers
- glycol ethers,
- alcohols
7.19 Formulation of metal-forming lubricant: carriers and additives.
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Ff
Fc
Pe
V
t
Pi
Pe
Die
V
t
Ff
Pi
Die
Variant II
Variant I
V2
V2
Pi
Die
Pe
Die
Pi
t
t
F1
Pe
F2
V1
Variant IV
Variant III
7.20 Variants of guiding zone tribo tests.
μ=
Ff
FC
[7.7]
This test requires careful design of the tooling to ensure that energy flow
in the system components does not result in significant errors in the measurement of the normal load.
Variant II was originally developed at the Engineering Research Centre
for net shape manufacturing at the Ohio State University [14]. In this test,
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167
the vertical shaft that holds the tube is connected to a load cell which measures the frictional force, Ff. Unlike variant I, the normal load in this test is
determined indirectly using the internal pressure of the tube and the
properties of the tube material. The friction coefficient is determined by
equation 7.8.
μ=
cFf
Pi πDi L
[7.8]
where Ff is the friction force, Pi is the internal pressure in the tube, Di is
the internal diameter of the tube, L is the effective length of the tube and
c is a constant which depends on the flow stress of material Pi. This test
requires c to be determined correctly. For higher pressure and tubular
specimens with higher diameter-to-wall thickness ratio, the value of c will
approach 1.
Variant III developed at the University of Paderborn [22] differs from
variant II by virtue of how the friction force is determined. In this test, two
punches are connected to the tube ends. These punches are also connected
to separate load cells to measure the friction force. With this test, it is possible to emulate compression of the tube when the tube is pushed through
a die. The friction force is determined by taking the difference in the friction
load measured by the two load cells. The friction coefficient can be determined by equation (7.9). The test assumes that the internal pressure is equal
to the pressure at the tool–workpiece interface.
μ=
F1 − F2
Pi πDi L
[7.9]
Variant IV was also developed at the University of Paderborn [8]. The
unique feature of this test is that the test does not require friction load
measurements. The coefficient of friction is determined solely based on
geometric change of the specimen and the internal pressure. Assuming that
the internal pressure Pi is equal to the pressure at the die–workpiece interface, the coefficient of friction can be established as follows [8]. This test
requires that the accurate flow stress of the specimen be known.
F1 = F2 + PiπDahm
[7.10]
μ=
F1 − F2
Pi πDa h
[7.11]
μ=
F1 − F2
A σ − A2σ 2
= 1 1
Pi πDa h
Pi πDa h
[7.12]
where s1 and s2 can be expressed in terms of Pi and effective flow stress
A
seff−1 and seff−2, and seff−1 = Ke 1n, ε 1 = 1 , h is the final height of the specimen,
Ao
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and A1 and A2 are annulus areas of the specimen on the side of the moving
punch and non-moving punch, respectively.
Expansion zone tribo tests
Two commonly used tribo tests to evaluate lubricants for the expansion
zone are the corner-fill test which was originally developed at the Industrial
Research and Development Institute (IRDI) [20] and the pear-shaped
expansion test which was developed at the Ohio State University [15].
Figures 7.21 and 7.22 show the corner-fill and pear-shaped expansion tests,
respectively. The premise behind the corner-fill test is that, under good
Pi
Die
Do
7.21 Corner-fill test.
Heater
d
a
ts
Die
P
ts
Tube
7.22 Pear-shaped expansion test.
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169
lubrication conditions, the tube will be inflated to tight corner radii before
fracture, and the wall thickness distribution will be uniform. For higher
friction at the tube–interface, fracture will occur prematurely or significant
wall thinning will be observed near the corner [20].
Whereas in the corner-fill test material flows towards all four corners, the
pear-shaped expansion test confines the material to flow in one direction
only. After expanding the tube to pressure P, d or protrusion height PH can
be determined, where PH is the distance from the bottom of the tube to the
apex of the deformed tube. The delta at fracture df is a function of friction
stress ts, flow stress s, and the internal pressure Pi. The die angle, α can be
optimized via numerical modelling by maximizing the difference between
d μ2 and d μ1. Performance evaluation of the lubricants can be achieved based
on three criteria: wall thinning distribution (the lower the wall thinning, the
better the lubricant); protrusion height Ph (the higher the height of protrusion, the better the lubricant); and bursting pressure (the higher the bursting
pressure, the better the lubricant) [21].
In both the expansion tests, friction coefficient cannot be determined
directly. However, by combining experimental results and FEA simulation
results, the coefficient of friction can be estimated.
7.4.3 Results from tribo tests
Some results of the performance of hydroforming lubricants are presented
in Table 7.1. The four lubricants given in Table 7.1 were evaluated using the
guiding zone and pear expansion tests.
Guiding zone test results
Tubular specimens of 100 mm length, 57 mm diameter, and 2 mm wall thickness were cut from AISI1008 material. Experiments were conducted using
Table 7.1 Lubricants test matrix
Name
Properties/contents
Application condition
Lub A
Polymeric film and blend of nonabrasive, dissimilar materials
Solid lubricant, free from chlorine and
sulfur
Carbon black, graphite butoxyethanol
and water
Thermoplastic polymer, water and
lithium stearate
Wet
Lub B
Lub C
Lub D
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Load cell
Tube
Die
Die insert
7.23 Experimental setup (one die half removed).
0.05
Lub D
0.045
Coefficient of friction (m)
0.04
Lub C
0.035
0.03
Lub A
0.025
0.02
0.015
Lub B
0.01
0.005
0
0
20
40
60
80
100
120
Sliding length (mm)
7.24 Friction coefficients of tested lubricants using guiding zone
tooling [5].
a 1600-kN hydraulic press at a ram speed of 60 mm s−1 (Fig. 7.23). An internal pressure of 60 MPa and a sliding length of 115 mm were used. Figure
7.24 shows the average friction coefficient for the four lubricants tested. The
results show that Lub B exhibited the lowest friction coefficient followed
by Lub A and Lub C. Lub D was the least effective among the four lubri-
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171
cants tested. The same set of lubricants was then tested using the expansion
zone test.
Pear-shaped tube expansion test
Tubular specimens of 250 mm in length, 57 mm in diameter, and 1.6 mm
wall thickness were cut from SS304 and AISI1008 materials. A 70 MPa
hydraulic pump was used to pressurize the tubular specimens. The tubes
were pressurized using the same pressure loading for all lubricants so that
the difference in wall thickness distribution and protrusion height could
be compared.
Figure 7.25 shows the variations of wall thickness distribution of the
formed pear-shaped specimens. The wall thickness was measured by an
ultrasonic measuring device. The wall thickness distribution profiles for
lubricants A and D decrease linearly along regions I and II. At point A, a
wall thickness of t = 1.6 mm is observed, whereas, at point B, a maximum
thinning of t = 1.32 mm is observed for both Lub A and Lub D. The wall
thickness distribution for lubricants C and B and Non-lubricated (No-Lub)
exhibit different thinning trends along regions I, and II. In region I, the wall
thickness decreases gradually, whereas in region II it decreased rapidly to
a maximum thinning (at point B) of 1.25, 1.25, and 1.18 mm for Lub C, Lub
B and No-Lub, respectively. The linear decrease in wall thickness from
region I to II for Lub A and Lub D implies that the tube material flows
easily toward the apex of the pear-shaped die. This is attributed to the low
friction stress at the tool–workpiece interface. Hence, Lub A and Lub D
have lower friction coefficients. Thus, using the difference in the maximum
wall thinning as one of the criteria to rank lubricants, Lub A and Lub D
1.7
Wall thickness (mm)
1.6
1.5
Lub D
1.4
Lub A
Lub C
Lub B
1.3
1.2
No Lub
1.1
A
Region I
Region II
B
1
1
20
40
60
80
100
Curvilinear length (mm)
7.25 Effect of lubricants on wall thickness distribution [21].
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Protrusion height PH (mm)
172
72.5
72
Specimen material SS304
71.5
71
70.5
70
69.5
69
68.5
68
Lub A
Lub D
Lub C
Lubricants
Lub B
No Lub
Increasing interface friction
7.26 Lubricant performance as a function of protrusion height
(PH) [21].
have the highest lubricity level. In order of increasing lubricity, the four
lubricants can be ranked as Lub A, Lub D, Lub C and Lub B.
Another criterion used to rank the lubricants was the difference in the
protrusion heights attained, i.e., the higher the protrusion height, PH, the
better the lubricant. Figure 7.26 shows the protrusion heights attained for
all lubricants. Lub A resulted in a protrusion height of 72 mm followed by
Lub D with a protrusion height of 71.5 mm. These results show a similar
performance trend as that observed in evaluating the wall thickness
distribution.
It can be seen that the performance ranking of the two tests are different.
In the Guiding zone test, Lub B exhibits the best performance and Lub D
the worst. On the contrary, the pear expansion test shows that Lub D is the
second best and Lub B is the worst. This is attributed to the differences in
possible lubrication mechanisms that are dominant in the respective friction
zones as discussed in Section 7.3.
The challenge facing lubricant development for hydroforming is to formulate a lubricant that can perform equally well in both zones. This is of
particular importance for hydroformed parts that require significant feeding
and expansion.
7.5
Impact of numerical modelling
in hydroforming tribology
One of the major factors that can benefit tribo-chemists in lubricant development is to be able to describe quantitatively the evolution of boundary
surfaces of the deforming part and their tribological derivatives. Though
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173
this is an extremely difficult task, the advancement in numerical computational capabilities has paved a way toward better and efficient examination
of the boundary surfaces. A typical metal-forming discretization problem
used in finite element (FE) formulations is represented by the functional,
Π = I1 + I 2 − I 3 = ∫ σδε dV + k ∫ ε Vδε V dV − ∫ Fiδ ui ds
V
V
[7.13]
SF
ol
To
O
th
va er s
rla ys
bl te
es m
.
.
where k is a penalty constant .and e v is the volumetric strain rate, s– , effective stress, e–, effective strain, e– effective strain rate Fi, traction. The I3 term
mathematically represents the boundary interactions. For a realistic
solution to a lubrication problem, the three terms (I1, I2, I3) should not be
considered in isolation of one another. The lubrication problem in metal
forming is a coupled problem in which the tribological aspects are related
to variables such as material properties of the deforming body, surface
topographies, lubricant and process geometry.
Figure 7.27 shows a numerical simulation databank which could be used
in optimizing a manufacturing process by concurrently considering all manufacturing process system components. The system components may include
(a) tool design, (b) process design and control, (c) machine design, and
(d) material and (e) lubricant development. In Fig. 7.27 tribomechanical–
thermal variables that could be obtained via numerical simulation are
shown.
The process-dependent tribomechanical/thermal variables derived from
computer simulations can be used as initial design inputs in the lubricant
gn
si
de
Numerical modelling
hydroforming process
tribo-variables
• Surface expansion/contraction
• Surface strain/hardness
• Surface temperature
• Surface topography
• Surface stress
Lubricant
formulation
M
l
ac
hi
ia
ne
er
at
de
M
si
g
n
Process design
and control
7.27 Numerical simulation databank supplies initial design
information to manufacturing process system components.
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development cycle. The continuum mechanics employed in numerical modelling through FE can accurately capture the deformation of billet subsurface. We can associate the subsurface deformation with changes in the
surface evolution, dislocation pile-up, stress and strain state of the surface,
and surface hardening behaviour. The manner in which these changes occur
predetermines the behaviour of a lubricant. A global optimization of a
manufacturing process should consider enhancement of tribological performance as one of the objective functions. Enhancement of tribological performance can be done through monitoring surface evolution. The importance
of assimilating process-dependent tribomechanical/thermal variables in
lubricant development is discussed below focusing on tube hydroforming
processes.
Case study
Numerical modelling using FEA of a simple THF process was carried out
to establish some tribomechanical variables that may be relevant for lubricant formulation. Figures 7.28 and 7.29 show a typical THF process and the
loading paths, respectively: the interface is subjected to non-linear mechanical and thermal loading as a function of time. Stainless steel SS 304 was
used in the simulations.
Figure 7.30 shows the local surface expansion distribution for different
axial feeds. The simulation shows half a section of a bulged tube through
THF. A high surface expansion gradient is observed from point A to point
C. In location AB the surface expansion is observed while location BC
exhibits surface contraction. Figure 7.31, shows a very high surface strain
gradient along the tube surface. Converting the flow stress at different strain
Die
Pi
tt
mt f(Pit, Fat, ...Lub...)
7.28 THF process.
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Tribological aspects in hydroforming
30
300
25
Axial feed path
200
20
150
15
100
10
50
Axial feed (mm)
Pressure (MPa)
250
5
Pressure path
0
0
0
10
20
Time (s)
30
40
7.29 Loading paths for THF of bulged part.
Local surface expansion (%)
60
Axial feed 5 mm
Axial feed 15 mm
A
50
40
Axial feed 10 mm
Axial feed 20 mm
B
C
30
20
10
0
–10
0
25
50
A
B
75
100
125
150
175
C
–20
7.30 Local surface expansion distribution.
Tube surface strain
0.7
Axial feed: 5 mm
Axial feed: 15 mm
0.6
0.5
Axial feed: 10 mm
Axial feed: 20 mm
A
B
0.4
C
0.3
0.2
0.1
0
0
A
25
50
75
100
125
B
150
C
Curvilinear length (mm)
7.31 Surface strain distribution.
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Hydroforming for advanced manufacturing
1
15.00
4
V-groove volume (mm3)
14.50
17
14.00
13.50
13.00
12.50
12.00
Before forming
10 mm axial feed
20 mm axial feed
11.50
11.00
0
2
4
6
8
10
12
14
16
V-groove No.
7.32 Lubricant pocket volume change (FEA).
levels to surface hardness shows that region AB is four times harder than
region BC. See equation 7.14 where the hardness is assumed to be proportional to the yield of material. Guiding-zone experiments have shown that
friction increased with an increase in surface strain [24]. Similarly, the stress
states in these regions are entirely different. Region BC has compressive
stresses while region AB has tensile stresses.
n
σ AB K ε AB
K 0.60.6
=
=
= 4.44
n
σ BC K ε BC
K 0.050.6
[7.14]
If it is assumed that the atomic opening on the surface is proportional to
the surface tensile loading and inversely proportion to the compressive
loading, then the diffusion of lubricant chemicals may be influenced by the
state of stress, thus changing the lubrication mechanisms. Similarly, the
surface hardness gradient observed will suggest different chemical/ lubricant formulations.
Figure 7.32 shows FEA results where a tubular specimen with ‘v-type’
lubricant-pockets was simulated. It is observed that lubricant pocket
volumes on the surface vary as a function of loading paths (axial feed and
pressure path). The lubricant pocket volume decreased uniformly from
pocket 4 to 17 as the part is being formed. Through FE analysis it can be
demonstrated that different lubrication mechanisms are exhibited at the
guiding zone, transition zone and expansion zone. In the guiding zone, fluid
type lubricants can perform better due to hydrostatic/hydrodynamic effect
caused by a decrease in pocket volume as compared with the expansion
zone (see Section 7.3).
Furthermore, the influence of subsurface deformation on dislocation
mechanisms can be studied and related to the lubricant chemistry. In the
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177
formulation of hydroforming lubricants, these tribovariables could be considered concurrently with variables discussed in Section 7.2.
7.6
Concluding remarks
In this chapter tribological aspects in hydroforming have been discussed
with an emphasis on THF. Numerous tribo variables interact at the tool–
workpiece interface in a complex manner. By classifying THF into feeddriven and expansion-driven categories, it is possible to distinguish the
lubrication mechanisms in THF which may play a vital role in the determination of effective lubrication systems.
Though tribo tests for tube hydroforming developed to date can provide
performance differences of lubricants suitable for different friction zones,
formulating a single lubricant that can perform equally well in all zones
remains a challenge. The use of numerical modelling presents an opportunity to understand in advance the tribo mechanical and thermal variables
for a specific hydroforming process. These variables can be used as path
indicators in the development of new hydroforming lubricants and lubrication systems.
7.7
References
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20. g. dalton, ‘The role of lubricants in hydroforming’, in: Proceedings of the Automotive Tube Conference, Dearborn, MI, 26–27 April, 1999.
21. gracious ngaile, stefan jaeger and taylan altan, ‘Lubrication in tube hydroforming (THF) Part II: Performance evaluation of lubricants using LDH test
and pear-shaped tube expansion test’, Journal of Materials Processing Technology, 146, 2004, 116–123.
22. f. vollersten and m. plancak, (2002) ‘On possibilities for the determination of
the coefficient of friction in hydroforming of tubes’, Journal of Materials Processing Technology, 412–420.
23. gracious ngaile, ‘Application of numerical methods in metal forming tribology’, 8th International Conference on Numerical Methods for Industrial Forming
Processes, July 2004, Columbus OH, USA.
24. p. groche and a. peter, Behavior of lubricants in dependence of the surface
topography within internal high pressure forming, 2003, MAT-INFO WerksstoffInformationosgesellschaft mbH, Humburger Allee 26, D-60486 Frankfurt, pp.
279–297.
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Part II
Hydroforming techniques
and their applications
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8
Pre-forming: tube rotary draw bending and
pre-flattening/crushing in hydroforming
G . K H O D A Y A R I, Vari-Form, Canada
8.1
Introduction
Hydroforming is used in a variety of industries ranging from plumbing to
aerospace. This process has shown a good potential for manufacturing highquality parts with lower production cost in comparison with conventional
stamping and welding processes. Tubes are one of the important source
materials for manufacturing hollow bodies using the hydroforming process.
In the hydroforming process, a high-pressure fluid applied to the inside of
a tube, forces the material outwards to the cavity surface within a closed
die. Internal pressures can also be accompanied with axial end feed to
improve the formability of the material.
Compared to conventional manufacturing methods, hydroformed parts
exhibit higher dimensional accuracy, structural strength, rigidity, dimensional repeatability and part consolidation (Worswick, 2004 and Gholipour,
2003). However, manufacturing of parts with complex structures from
tubular materials, particularly in the automotive industry, often requires one
or more pre-forming operations such as bending and crushing before the
actual hydroforming process. These operations are necessary to ensure the
tube fits properly inside the die cavity without damaging the die or part
during die closure.
Bending and crushing as two important pre-forming operations, however,
consume a significant portion of formability of the tubular material. Furthermore, due to the non-uniform plastic deformation around the circumference of a bent tube, the wall thickness decreases in the outer radius and
increases in the inner radius. Thus, the hydroformability of the bent tube as
well as the thickness uniformity of the end product is restricted. In special
cases such as pressure sequence hydroforming (PSH; a patented process
developed by Vari-Form Inc.), the process is limited predominantly by the
material’s ability to survive the bending process.
In order to reduce vehicle weight (and greenhouse gas emissions) without
compromising structural integrity, there has been a recent push to increase
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the use of high-strength steels and dual-phase materials in automotive
component. Hydroforming of bent tubes from these materials has been
more and more important as a result. Thus, this chapter deals with the
effects and aspects of tube rotary draw bending as well as the crush-forming
process.
8.2
Concept of rotary draw bending process
Bending is one of the most frequently applied pre-forming operations in
the tube hydroforming industry. It is the first necessary forming operation
for almost all automotive structural components and sets the stage for
quality and production down the line. The purpose of bending for hydroforming is to achieve geometry closer to that of the final part to be manufactured by hydroforming, enabling the tube to fit into the die cavity.
A number of different tube bending techniques and concepts are in use
in various industries, but the most common and advantageous of them is
CNC rotary draw bending. Its major benefits are speed, accuracy, repeatability and relatively good control of wall thickness variation due to bending.
Rotary draw bending applies both moment and transverse loads on a
straight tube (Khodayari, 2003a). It can be loaded and unloaded manually
or fed with automated equipment. Figure 8.1 shows the schematic of a
standard rotary draw bender utilizing five major tools, which all interact
with one another via the tube during the bending process. The tools used
Pressure die
x
Boost
a
Tube
Mandrel
Rake angle
Wiper die
Bend die
Clamp die
8.1 Schematic of a tube rotary draw bender.
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are: the bend die, the clamp die, the pressure die, the wiper die and the
mandrel. Mandrel and wiper die are stationary in the forming zone as the
tube is being drawn through this zone. Because they are stationary, they
increase the ‘drag’ on the tube, which further adds to the clamping
requirement.
In the initial state – before starting the bending process – the bend die,
clamp die and pressure die are all oriented parallel to the tube length axis.
The wiper die is positioned at a small angle from the tube axis, which is
usually referred to as the ‘rake angle’. It reduces the friction drag from the
wiper during bending, while assisting in preventing wrinkling on the inside
of the bend. The mandrel consists usually of a cylindrical mandrel shaft with
one, two or more mandrel balls utilizing swivel connectors. The mandrel
balls support the tube from the inside and minimize the ovality (see 8.3.2)
of tube cross-section during the bending process.
Once the tube is slid into position over the mandrel, the clamp die is
closed applying a force on the tube. The bend die and clamp die then begin
to rotate together, pulling the tube over the mandrel, while the pressure die
advances parallel to the mandrel shaft, thereby pushing material into the
bend region and acting against the bending reaction forces. In order to bend
a tube with end-feed (push-assist), an independent actuator can be used to
push on the pressure die (and tube via friction) to feed more material into
the bend region. This pushing action is known as ‘boost’. The axial boost
improves the material flow and leads to less thinning of the tube wall on
the outside of the bend, which is the most critical area when hydroforming
bent tubes (Khodayari, 2003b).
As with any other metal-forming process, there are restricting parameters
in tube rotary draw bending too. For instance, some of the system related
limits and variables are machine capability (interference), machine rigidity
(robustness), tribological conditions, clamping and boost force (Khodayari
2002a; Dwyer, 2002a and 2002b). The most restrictive is often the clamp
length, which is a prime limitation on hydroform part design as well. The
straight length between bends drives the clamp length. The recommended
minimum clamp length is about twice the tube diameter. When too short a
clamp is selected, the leading edge of the clamp will dent the tube and not
bend the specified angle.
8.3
Material behavior in rotary draw bending process
During the bending process, all material layers in the bend area between
the inner and outer radius undergo stress and strain conditions which
change continuously throughout the bending process. No layer is continuously strain-free and no layer is continuously stress-free. The strain-neutral
axis (layer) coincides with the stress-neutral axis only in the straight tube
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Tension
Mandrel side
Forming zone
(stationary)
Neutral axis
B en
do
sion
ress
ion
Com
p
Ten
CLR
nd
Be ide
ins
α
id e
uts
Compression
Clamp side
8.2 Forming zone and the bending terms.
or at the beginning of the bending operation (hence, neutral-axis). With
increasing deformation and decreasing bend radius, the strain- and stressneutral layers differ from each other and at the same time from the position
of the initial neutral axis. However, when going from tension stress (bend
outside) to compression stress (bend inside), there is always one layer (axis)
which is stress-neutral and one layer which is strain-neutral. After bending,
the strain-neutral or non-elongated axis will have exactly the same length
as before bending.
At the very beginning of the bending process, the deformation of tube
remains elastic until the induced compression and tension stresses in the
bend-inside and/or bend-outside (Fig. 8.2) reach the yield stress of the tube
material. Then, plastic flow begins. At this angle position, the forming zone
is still not fully made-up. Therefore, the actual bend radius of tube is larger
than the radius of the bend die.
By further rotation of bend die, the area of plastic deformation grows in
both the tube’s cross-section and along its length axis, while the bend radius
of tube decreases gradually until it coincides with the centreline radius
(CLR) of the bend die. At this point, the forming zone is completely developed and the maximum level of possible plastic deformation in the tube’s
cross-section is achieved. Upon completion of the forming zone, the plastic
deformation propagates only along the tube length axis. During the rest of
the bending process, the tube is quasi-drawn through the forming zone and
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Tension zone
(bend outside)
Neutral axis
before bending
T
d
Displacement
of neutral axis
Neutral axis
after bending
Compression zone
(bend inside)
CLR (R)
C
Bend axis
8.3 Displacement of neutral axis in bend area.
plastically deformed into the bend die. The position of forming zone remains
unchanged throughout the whole bending operation. In this forming phase,
the bend loads and bend reaction forces remain at their maximum level
until the process terminates.
Because of the bending, the outside tension loads and inside compression
loads cause thickness reduction on the outside and thickness increase on
the inside of the bent tube, respectively. Consequently, the location of the
strain-neutral axis (called the neutral axis) moves toward the bend inside.
A new axis below the tube’s centerline (initial neutral axis) becomes the
actual neutral axis (Fig. 8.3). This phenomenon is characterized as the ‘displacement of neutral axis’.
With D as outside diameter of tube and R as centreline radius (CLR) of
the bend die, the displacement of neutral axis d can be determined using
equation 8.1.
d≈
D2
D
=
4 R 4(R/D)
[8.1]
In equation 8.1, the term (R/D) refers to the so-called ‘bend-ratio’, which
is a measure of bend severity. The smaller the bend-ratio, the higher the
bend severity. Therefore, the displacement of neutral axis increases with
bend severity, which means: decreasing the bend radius and/or increasing
the outside diameter of tube. The bend severity also indicates the desired
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clamp length and the length of straight section between adjacent bends. The
higher the bend severity, the longer the desired clamp length. Furthermore,
more severe bends cause higher work hardening and higher loss of material
formability in the bend regions, leaving less formability for subsequent
operations like crushing and hydroforming.
Because of the ‘displacement of neutral axis’ (equation 8.1), the initial
neutral axis experiences an elongation which entails an axial elongation of
the entire tube in the bend area. Considering the displacement given by
equation 8.1 and α as the desired bend angle in radians, the elongation
‘ΔL’ of tube can be determined by equation 8.2.
ΔL = dα ≈
D α
4(R/D)
[8.2]
When designing tubular hydroformed parts, particularly with severe
bends, the tube elongation, due to bending, must be considered as a source
for possible geometrical deviations. These deviations will cause difficulty
when inserting the bent tube into the hydroform die.
An additional effect of axis displacement is the relatively higher thickness strains in the bend outside (tension zone) compared with those in
the bend inside (compression zone), which can be seen on strain measurements around the circumference of bent tubes (see Fig. 8.8). Failure in
the tension zone occurs when grain boundaries can no longer hold the
integrity of the material. This is usually preceded by severe ‘necking
down’ (localized reduction of wall thickness) and is an indication of
exhausted material formability. Failure in the compression zone can be
from one of two modes. The first is a pure compression failure where the
material uniformly thickens and fails by shearing between the grain
boundaries. The second mode more commonly seen, comes from buckling
or wrinkling.
8.3.1 Springback
Springback, which is inevitable in a cold-forming process, is caused by the
elasticity of materials. It is characterized by the tendency of materials
subjected to a cold-forming process to return to the initial state after
unloading. In bendforming, the springback occurs when clamping forces
that hold the tube against the bend die are removed. Apart from some
special cases like spring manufacturing, springback is considered as a main
defect in a metal-forming process and has been studied by several researchers (Asnafi, 2000, Shr, 1999). The overall deviation of a bent part increases
with increased number of the bends. The complete elimination or compensation of the effects of springback can increase production costs
substantially.
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Δα = α2 – α1
α2
α1
α1 = Angle under load
α2 = Angle after unloading
8.4 Angular springback of a bent tube after unloading.
In tube bending, springback occurs in both angular and radial directions.
The angular springback is defined as the difference between the angle a1
before unloading and the measured actual angle a2 after releasing the bent
part from the bender (Fig. 8.4).
The angular springback can be compensated by overbending the target
angle about the springback angle Δa. The desired compensation for springback, however, requires its exact determination or prediction. Although
most of the industrial CNC tube benders are able to keep the target angle
with an accuracy of a hundredth degree, the necessary CNC data to compensate the springback is still obtained from a time-consuming and
material-wasting trial-and-error procedure for any given tube material,
tube geometry or bend parameter setting. The effort increases with the
number of bends in a part and causes major problems when manufacturing
hydroformed assemblies utilizing multiple bent tubes. Therefore, the understanding of springback behavior in the bending process would help to minimize the efforts to its prediction. Figure 8.5 shows the characteristic diagram
of angular springback of tubular materials in the mandrel rotary draw
bending process (Khodayari, 1998).
Analogous to the material behavior in the bending process (Chapter 8.3),
the angular springback shows a non-linear characteristic in the range of
smaller bend angles. In these angle settings, the non-linearity of springback
is mostly caused by the shape and size of the forming zone which indicates
the elastic–plastic transitions from a straight section (start point of forming
zone) to a fully bent section of tube. Since the size of the forming zone
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Spring back (°)
188
Non-linear
springback
ack
ingb
pr
ar s
Line
Bend angle (°)
8.5 Angular springback characteristic bent tube.
(after its completion) does not vary for different angle positions, its relative
impact on total springback is higher for smaller bend angles. Thus, the
prediction and control of springback angle for smaller bends is much more
challenging, specially, for angle positions below about 15°. Beyond this
relatively small angle range, the springback exhibits a linear behavior
versus bend angle. This means, knowing the amount of the springback angle
at two different bend angle positions would enable its prediction for any
other bend angle.
In general, tubular materials with a higher yield/tensile stress and lower
Young’s modulus are expected to have higher springback after bending
(Khodayari, 2002b). Besides the material properties and tube geometry,
there are a large number of other process variables influencing the level of
springback, for instance:
•
•
•
•
•
•
bend radius and bend angle,
mandrel type and number of mandrel balls,
wiper die and rake angle,
lubrication,
level of applied axial force (end feed),
bend speed and robustness of the bender.
Therefore, an exact prediction of springback angle will always be a challenging task for design and manufacturing engineers (Khodayari, 1994).
One of the most recommended methods of minimizing the efforts of springback determination is the establishment of a database which contains the
springback data collected from tests, experiments and actual manufacturing
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Radial springback = R2–R1≈d
I
R2
R1
R1 = Nominal center-line radius of bent part (bent die)
R2 = Actual centre-line radius of the bent part
8.6 Radial springback of a bent tube.
process as a function of material properties, tube geometry and process
parameter setting. Over time, and utilizing a capable and intelligent data
analyzing software, the collected data can be optimized and used as a
powerful engineering tool to predict the springback for new bend combinations. The degree of accuracy of the prediction will increase with the amount
of collected bend data.
As mentioned earlier, the bendforming causes springback in the radial
direction as well. It is caused by both the displacement of the neutral axis
and the elasticity of the tube material. The major portion of the radial
springback is, however, caused by displacement of the neutral axis (equation 8.1), and therefore the material elasticity can be neglected. Because
of radial springback, the actual radius of the bent part (after unloading)
is larger than the nominal bend radius (centreline radius of the bend die;
Fig. 8.6).
In contrast to angular springback, the radial springback cannot be compensated in the bending process. Therefore, in order to manufacture a bent
part with an exact bend radius, it is necessary to modify (correct) the design
of the bend die in such a manner as to ensure that the end part obtains the
desired radius after unloading. This can be realized by reducing the nominal
centre-line radius of the bend die by the amount of the displacement of
the neutral axis d determined by equation 8.1; exact CLR = (nominal
CLR) − d.
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8.3.2 Cross-section ovality
In bending, ovality refers to the flattening of tube cross-section due to the
plastic deformations which occur during the bending process (Fig. 8.7). The
relatively opposed radial forces, resulting from tension loads in the bend
outside and compression loads in the bend inside, deform the cross-section
of tube from an originally circular shape into an oval form. In the rotary
draw bending process (Fig. 8.1), the tube is usually supported on the outside
of the bend by the pressure die, the inside of the bend by the bend die and
inside the tube by mandrel balls which prevent the tube from collapsing
(ovalizing). However, once the tube has left the support region of the
mandrel and circular cavity of the pressure die, the cross-section loses the
necessary support and become oval as a result of the radial components of
the bend reaction forces (Khodayari, 1994). The ovality U can be expressed
as the percentage of cross-sectional deviation of a bent tube from its original circularity (equation 8.3).
U=
D2 − D1
100
D0
[8.3]
where D0 is the initial outside diameter of tube, D1 is the smallest and D2
the largest diameter measured in the bend area of the tube.
Figure 8.7 illustrates the ovalities measured in 5° increments along the
bend arc on aluminum and steel tubes bent at 90° under the same bending
10
Aluminum
8
Ovality u (%)
7
Steel
6
5
Tension
4
D1
3
2
Compression
1
0
10
20
30
D2
40
50
Bend angle (°)
8.7 Cross-section ovality of a bent tube.
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80
90
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conditions and process parameter settings. The higher ovality of the aluminum tube compared with the steel tube indicates the effect of material on
cross-section ovality during the bending process. The ovality has the same
characteristic for both materials; almost constant through the middle area
of the bend with a smooth ramping up and down at the beginning and
end of the bend area, respectively. The fact that the ovality does not disappear at the start and end points of the nominal bend angle (0° to 90°)
indicates that the plastic deformation extends into the straight areas immediately preceding and following the nominal bend area.
Mandrel ball size determines the clearance to the inside surface of the
tube wall. Although tighter mandrel fit minimizes tube ovality, it has to be
in an acceptable range for normal production operations. In general, the
more severe the bend, the tighter this range must be and the higher the
desired clamp load. Too tight a mandrel fit will cause excessive wall thinning
and stretching in the bend area which can result in premature failure. Conversely, excessive clearance between the mandrel and tube inside wall can
result in buckles or wrinkles during the bending process.
8.3.3 Thickness strain distributions in a bent tube
Because of the non-uniform plastic deformations in the bending process,
the wall thickness around the cross-section of the bent tube decreases in
tension zone and increases in the compression zone. The thickness strains
increases with the severity of bend. The wall thinning which occurs reduces
the formability of the bent tubes in hydroforming drastically. Hence, it is
important to investigate the thickness strain distributions after the bending
process in order to evaluate and predict its hydroformability.
Figure 8.8 presents the characteristic of thickness strain distribution measured at 25 point intervals around the circumference of a bent tube. The
start and end points of the measurements were on the weld seam, which
were positioned on the neutral plane during the bending operation.
It can be observed (Fig. 8.8) that the strain curve on the inside of the
bend (left-hand side of the diagram) intersects the strain 0-line at an angle
position of less than 180°, which means that the peripheral portion of the
stretched cross-section zone is larger than the peripheral portion of the
compressed cross-section zone. The distance of the intersection point from
the position of the initial neutral axis at 180° is equivalent to the displacement of the neutral axis given by equation 8.1. It represents the circumferential angle position in tube cross-section for which the wall thickness did
not change as a result of the bend forming. Furthermore, the graph shows
that the absolute value of the maximum thickness strain measured at the
outer point on the tension zone is approximately 5% higher than that measured at the outer point on the compression zone. Since the wall thickness
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0.20
Bend inside (compression zone)
Bend outside (tension zone)
True thickness strain
0.15
0.10
Displacement of neutral axis
0.05
0.00
Position of neutral axis after bend
–0.05
–0.10
Initial position of neutral axis
–0.15
–0.20
0
30
60
90
120
150
180
210
240
270 300
330
360
Angular position from the weld (°)
8.8 Thickness strain distribution around the circumference of a bent
tube and displacement of neutral axis.
has a major influence on formability of bent tubes in the hydroforming
process as well as on thickness uniformity of the final part, it is necessary
that the bent tube has the desired minimum wall thickness and formability
in the areas of critical deformations.
One of the significant process parameters influencing the thickness strains
during the bend forming, particularly on the outside of the bend, is endfeeding (see Fig. 8.1). End feeding is realized by applying an axial force on
the straight end (mandrel-side) of the tube. In order to enable bending with
end-feed, a boost block can be attached to the tail end of the pressure die
for a direct axial push on the end of the tube. The boost (expressed as a
percentage) is defined as the ratio of the pressure die displacement (X) to
the tangential displacement of a point on the centreline radius (CLR) of
the bend die. The bend die rotates by an angle (a), thus the boost is given
by equation 8.4.
(
)
X
Boost = − 1 100
αCLR
[8.4]
Bending without boost refers to the condition when the pressure die
velocity is equal to the tangential velocity of the bend die which means
X = α /CLR.
For the strain curves shown in Fig. 8.9, the maximum end-feed achieved
was 4%. End feeding can only be increased if it does not cause wrinkling
on the inside of the bend and if no inadvertent deformations appear on the
bent tube.
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0.20
True thickness strain
0.15
Tube bent with push-assist
0.10
0.05
0.00
–0.05
Tube bent without push-assist
–0.10
–0.15
~7% strain resuction (relatively)
–0.20
0
30
60
90 120 150 180 210 240 270 300 330 360
Angular position from the weld (°)
8.9 Influence of push-assist on thickness strain distribution.
The axial force improves the material flow and minimizes thickness strain
on the outside of the bend, which is the most critical area when hydroforming bent tubes. Figure 8.9 demonstrates the influence of push-assist on
thickness strain distributions measured around the circumference of DP600
bent tubes with a 3.5 inch (88.9 mm) outside diameter and a 1.5 mm wall
thickness. The tubes were bent at 90° with a bend-ratio of 2.5. The comparison of the measured circumferential strain distributions shows that the axial
end-feed leads to a relative thickness strain reduction of approximately 7%
on the outside of the bend.
Although push-assist improves tube formability in bending and subsequently in hydroforming, the maximum level of applied axial force is,
however, limited by the formation of wrinkles on the inside of the bend
(excessive boost would cause earlier emergence of wrinkles). When
bending tubes for hydroforming, wrinkles in the bend inside must be
avoided since wrinkles typically cannot be hydroformed out of the part.
Manufacturers must also estimate the thinning that occurs in the bend
outside in order to determine whether the resulting tube thickness is
adequate to satisfy minimum thickness requirements of the part after
hydroforming.
8.4
Pre-flattening/crushing
Depending on the complexity of the final product to be manufactured,
pre-flattening or crushforming may be required (Fig. 8.10). This is a targeted re-shaping of a tube’s cross-section over some or all of the part length
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(1) Bent tube
(2) Crushformed tube
(3) Hydroformed and
hydropierced tube
8.10 Forming stages; bending, crushing and hydroforming (courtesy
of Vari-Form).
in order to prevent pinching within the hydroform die during the closing
process. In addition, an appropriate crushforming can improve the tribological condition and lead to better sliding of tube in the die cavity during
axial feed if applied. It also enables parts to be formed with higher thickness uniformity and reduces the residual stresses on the hydroformed parts.
The crushforming, however, may cause additional deformations in the
tubular part, which have to be removed in the subsequent hydroforming
operation.
In order to achieve full plastic deformation in the entire part, the tube
perimeter can be made slightly smaller than the die geometry in each crosssection in order to minimize springback. Furthermore, it is recommended
to orient the tube before clamping in the tube bender so that the weld line
is (as far as the part design allows) in the neutral zone and is not subject to
tension or compression during bendforming. It is also good practice to orientate the weld seam in such a manner that it does not pass through holes
to be pierced in the hydroforming die.
The crushforming operation can be eliminated in hydroforming operations utilizing a start tube with a nominal outside diameter smaller than the
part width. Crushforming is necessary when the cross-section of tube is
larger than the part width. It is important to distinguish between crushform-
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Cross-section
(bend area)
Cross-section
(straight area)
Initial tube
cross-section
Pre-forming unit
prior to hydroforming or
in the hydroform die
Tapered die pre-forming
8.11 Crush-forming processes (courtesy of Vari-Form).
ing in straight sections of the tube and crushforming in the bend area. The
first squeezes the round tubular cross-section perpendicular to the direction
of die travel because the section in the cavity is narrower than the tube OD.
This type of pre-forming can be achieved before the hydroforming process
using a special device consisting of two contoured plates that are pushed
together with a hydraulic cylinder or, if possible, directly in the hydroform
die (Fig. 8.11).
The second is crushing a tube in bend section. Crushing or flattening of
a tube in a bend section can cause significant geometrical deformations in
the tube (e.g. extreme changes in the bend angle; depending on the amount
and direction of the applied crush forces). Hence, in order to keep the tube’s
geometry in an acceptable tolerance range, it is necessary to use a tool that
fully captures or encloses the part (or at least along a portion of its length
which needs to be crushed); see Fig. 8.11. The manufacturing effort and cost
of tools for this type of crushforming increases with the complexity of the
part shape. It is much higher if it captures the whole part from end to
end.
The position of the start (bent) tube, when it is nested in the die cavity,
is one of the most important keys to success in the hydroforming operation. If any part of the tube cannot be contained between the die halves
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Cross-section not crushformed
Cross-section crushformed
Press travel
vector
Forming
vector
Area
of pinch
b Forming vector and press
trabel vector (aligned)
a Forming vector and press
trabel vector (not aligned)
8.12 Controlled crush-forming. Forming vector and press travel vector
a not aligned and b aligned (courtesy of Vari-Form).
as the press closes, the tube will shift and/or likely pinch. In order to avoid
this, the imaginary line between the start tube and the centroids of the
finished part have to be in alignment with the press travel vector (the
path along which the press opens and closes). This means that, during
hydroforming, the press travel vector should run through both the finished part’s section-centroid as well as that of the start tubes for each
representative cross-section of the part (Fig. 8.12). Therefore, during the
design process, the direction of forming vectors for each section of the
part must be checked and aligned with the press travel vector as accurately as possible. This ensures that the tube material distributes itself
evenly in the die cavity without pinching as it closes. Furthermore, maintaining contact between the tube and lower die cavity, at least on two
surfaces, will prevent the start tube from shifting in the die during die
closure.
In those design situations where it is impossible to align the sectioncentroids of the start tube with the direction of press travel vector or seat
the tube in the die cavity due to the finished part geometry or bending
limitations, the tube has to be crushformed in a controlled manner, so that
the necessary condition (mentioned above) is realized. The extent of crushforming is predicted during part design, but developed during the prototyping process.
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In Fig. 8.12a, the start tube sits in the die cavity but its forming vector is
not aligned with the press travel vector and the finished tube centroid. As
the die closes, the tube will not be contained in the die cavity and it will be
pinched between the die halves. In Fig. 8.12b, the start tube has been preflatted (crushed) to an elliptical shape. It too nests in the die cavity, but its
centroid aligns with the direction of the press travel vector and the finished
part cross-section-centroid. As the die closes, the tube will fill the cavity in
a uniform manner with little risk of pinching or buckling.
Although there are several reasons to achieve sharper corners (e.g.
extending flat surfaces for hole placements near the corner, improving the
rigidity of a particular section or achieving tight manufacturing tolerances),
it is best practice to avoid forming sharp corners or wrinkles in pre-forming
operations. Sharper corners are key determinants of the maximum pressure
required to completely hydroform a part. Equation 8.5 demonstrates the
relationship between the smallest inner radius r a tubular part and the
required internal pressure Pi to achieve it by high-pressure hydroforming
process.
Pi ≈
UTSt
r − 0.5t
[8.5]
where t is the tube initial wall thickness and UTS is the ultimate tensile
strength of the tube material. Thus, the smaller the (corner) radius, the
higher the internal pressure required to achieve it.
Contrary to high-pressure hydroforming, in pressure sequence hydroforming (PSH), the tubular material does not experience significant expansion. Thus, the required internal pressure to achieve the same part geometry
is significantly lower. By appropriate design of hydroform die, the PSH
enables manufacturing of hydroformed parts with extreme tight corners
(flanges) and a high degree of complexity that could not be manufactured
by a conventional high-pressure hydroforming process. It is because the
tube cross-section can be formed during the die closure.
8.5
Part design and tube formability
The design of the product determines the simplicity and the success of the
hydroforming process. It dictates the feasibility/complexity, speed and cost
of the manufacturing equipment and process. Knowing the geometry of the
start tube and the shape of the final part (after hydroforming) enables the
designer to estimate the expected maximum deformations/strains and
predict the required minimum material formability in order to withstand
the necessary forming stages to achieve the desired final shape. Therefore,
it is crucial for the designer to understand the significance of material properties such as yield stress, tensile stress, uniform elongation, n and r values
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as well as the geometrical parameters such as cross-section perimeter and
initial/finished part wall thickness.
Although the existing analytical methods developed for tube hydroforming process are very useful tools to determine process limits; the increasing
complexity of hydroform parts, however, desires more and more the application of numerical methods to model the process and simulate the forming
behavior (Koc, 2002). Most commercially available simulation and FEA
software packages, which are excellent design engineering tools, can be used
to model and analyze almost any mechanical forming process including
stamping, bending and hydroforming. By proper interpretation of the
forming process and appropriate definition of boundary conditions, the
software provides qualitative and quantitative information regarding
the strain distributions in critical forming areas as well as predicting locations for possible material failure. However, the selection of a proper tube
material which can withstand all the forming stages such as bending, crushing and hydroforming, is a very crucial step. Currently, the most common
tool used to select a steel material for a given forming process is the forming
limit curve (FLC) which has been developed to evaluate the forming severity of stamped components (Fig. 8.13). The FLC is commonly plotted in
strain space and defines the combinations of principal surface strains beyond
which there is a higher risk of necking or splitting of the sheet material.
70
60
Failure
zone
Major strain (%)
50
40
FLC
30
Marginal
zone
Strains measured along
the outside of a DP600
bent tube
20
Safe zone
10
0
–40
–30
–20
–10
0
10
20
30
40
Minor strain (%)
8.13 Standard forming limit diagram (FLD) and the plotted strains
measured along the outside of a bent tube (tube not failed).
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The forming limit diagram (FLD) has become a widely used tool for
product designers and simulation analysts to optimize the design of steel
metal components, and for tooling and process engineers to design and
build the corresponding stamping dies.
Since tube hydroforming became an established manufacturing process
in the automotive industry, it was plausible for design and process engineers
to resort to the forming limit diagram as a failure criterion to evaluate the
robustness of the hydroforming process as well as the quality of tubular
components. However, the formability of a roll-formed tube differs significantly from that of the sheet material from which it was made (Green,
2003).
Although the current theory for determining thickness and strains during
bending correlates well with the results from experimental investigations,
using the standard forming limit curve would, however, lead to an underutilization of the tube materials. As shown in Fig. 8.13, in relatively severe
bending operations, the strains measured along the outer radius of the bent
tubes can exceed the upper formability limit (FLC) generated from its
original sheet material. Interpretation of the forming limit diagram indicates that the tube material should experience a failure in these regions.
However, this is not the case. Therefore, because of the lack of experimental evidence, the forming limits (or the likelihood of failure) of tubular
materials subjected to a mandrel bending operation are only ‘estimated’.
Hence, it is not known if the forming limit curve for bent tubes has the
same standard shape as that for sheet materials nor has it been clearly
proven that the tube-FLC can be predicted from the tube mechanical
properties as it is commonly done in North America for low-carbon sheet
steels.
The fact that the strains induced by the tube-bending process can exceed
the upper limit (FLC) of the marginal zone without failure, indicates that
the material behavior in bendforming differs significantly from that of the
stamping process. This difference can be caused by: (a) different mechanical
material properties due to the tube manufacturing process (different preforming histories), (b) different stress–strain conditions in bending compared with the stamping process, (c) the circumferential strains in
bendforming (subsequently the axial and thickness strains) are influenced
by the material flow from the compression zone into the tension zone (see
Fig. 8.2). Of equal importance is the effect of the bend die and tools (e.g.
mandrel, pressure die, clamp die, and wiper die) as well as the tribological
conditions in the bending process, which contribute to the differences in
material behaviors between stamping and bending processes. Therefore, the
exact determination of formability limits for tubular steel materials in bendforming will lead to much higher process stability as well as better materials
utilization.
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8.6
Conclusion
The tube rotary draw bending and tube crushing described in this chapter,
are two important pre-forming processes when manufacturing hydroformed
parts utilizing multiple bent tubes. Understanding the behavior of the material, geometrical changes in the bent tube, the influence of process parameters and their interactive influences during these pre-forming processes
will allow designers to perform products and tools design in a better, faster
and more reliable way. Bending with end-feed (boost) leads to an overall
thicker and more uniform tube wall and reduces the work hardening in the
bend region which improves the stability of the end product as well as the
hydroforming process. This would allow a reduction in material usage and,
therefore, lower the weight and cost of the end part.
In contrast to the stamping process (and in some cases, the hydroforming
process), the forming limits of tubular materials for bending have not been
explored widely enough. Despite the large number of material and process
variables (see 8.3.1) which makes it difficult to develop a global and standardized concept to predict the formability of tubular materials for bending
process, further research would improve the state of the art and the possibility to determine the failure criterion and generate the desired forming
limit diagram (FLD) for tube bendability.
8.7
References
n. asnafi (2000) ‘Springback and fracture in V-die air bending of thick stainless steel
sheets’, Material and Design, 21 217–236.
n. dwyer, m. worswick, j. gholipour, c. xia and g. khodayari (2002a) ‘Pre-bending
and subsequent hydroforming of tube: simulation and experiment’, Numisheet
2002, Proceedings of the 5th international conference and workshop on numerical
simulation of 3D sheet forming processes, Jeju Island, Korea, Vol. 1, pp. 447–452.
n. dwyer, m. worswick, j. gholipour and g. khodayari (2002b) Proc. Plasticity ’02,
the ninth international symposium on plasticity and its current applications, edited
by A. Khan, pp. 394–396.
j. gholipour, m. j. worswick, n. dwyer and g. khodayari (2003) ‘Damage development during bending and hydroforming of aluminum alloy tubes’, Proc. Plasticity
2003, 172–174.
d. e. green (2003) ‘Experimental determination of tube forming limits’, in K. Siegert,
Hydroforming of tubes, extrusions and sheet metals, Stuttgart, Germany.
g. khodayari, ‘Untersuchungen zum elastisch-plastischen Biegen von Stahrohrprofielen’, Doctor thesis, Siegen, Germany.
g. khodayari (1998) ‘Theoretische und experimentelle Untersuchungen zum walzrunden von plattierten und nicht-plattierten Grobblechen’, Post Doctoral thesis,
Siegen, Germany.
g. khodayari, j. v. reid and m. garnett (2002a) ‘Analyzing tubes, lubes, dies, and
friction. Using tribology to evaluate lubricant–material combinations in hydroforming’, Tube and Pipe Journal (TPJ).
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201
g. khodayari (2002b) ‘How material influences bending for hydroforming. Effects
on ovality, springback, and wall thickness in tubes’, Tube and Pipe Journal
(TPJ).
g. khodayari (2003a) ‘Hydroforming of pre-bend aluminum tubes’, Int conference
on accuracy in forming technology (ICAFT), Chemnitz, Germany, pp. 149–161.
g. khodayari and m. worswick (2003b) ‘Examining the effects of push assist on the
formability of aluminum tubes’, Tube and Pipe Journal (TPJ).
m. koc and t. altan (2002) ‘Prediction of forming limits and parameters in the tube
hydroforming process’, International Journal of Machine Tools and Manufacturing
42 123–138.
alex shr, gracious ngaile and taylan altan (2000) ‘Evaluation of tube bending
experiments – effect and design parameters in rotary draw bending process’,
THF/ERC/NSM-00-R-08.
m. j. worswick, j. gholipour, d. a. oliveira and g. khodayari (2004) ‘Severity of the
bend and its effect on the subsequent hydroforming process for aluminum alloy
tubes’, American Institute of Physics.
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Hydroforming: hydropiercing,
end-cutting, and welding
L . M . S M I T H, Oakland University, USA
9.1
Introduction
In Chapter 3, the introduction of fundamental concepts of the mechanics
of hydroforming helped establish the background for the discussion of
material property characterization in Chapter 4, and tribological aspects in
Chapter 7. In Chapters 2 and 6, a systematic approach for selecting and
customizing the tooling and controls for the hydroforming process was then
introduced. The topics of pre-bend effects and formability quantification
were addressed next in Chapter 5. In this chapter, some of the key processes
after the pre-processing (pre-bending, pre-forming) and processing phases
(swaging, expansion) are addressed. Three of the processes, which fall within
what may be called the ‘post-processing’ phase, include hydro-piercing, endcutting, and welding.
9.2
Hydropiercing
9.2.1 Preliminary considerations
Tube hydroforming operations have allowed for the replacement of subassemblies with a single structural unit featuring increased stiffness-toweight ratios. Such units however, are typically designed to be integrated
into an assembly of other structural components. The most common mechanisms for joining the constituent units of the assembly include weld joints,
pushpins, self-threading screws, and threaded fasteners. The properties of
the joints within the assembly directly influence its overall stiffness-toweight ratio, cost, and visual appeal. Therefore, proper attention to processes that prepare the hydroformed unit for subsequent operations should
be given.
A common element among many of the fastening mechanisms for tube
hydroformed products is the presence of some type of hole through which
a fastening device may be inserted. Hydropiercing is one method of producing a hole in a tube hydroformed part. There are many other methods,
202
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however. Some of the alternatives to hydropiercing are 1 milling and drilling, 2 conventional punching, 3 laser cutting, 4 plasma cutting, 5 flow drilling,
and 6 post piercing. For a given part, any combination of these methods
may be applied. The most critical factors which should be considered when
choosing a method for creating a hole are as follows:
•
•
•
•
•
•
•
•
•
Repeatability: Over the duration of the production, there should be a
minimal amount of variation in the hole geometry and location.
Accuracy: The intended hole geometry and location should satisfy a
pre-defined tolerance.
Cost: The overall cost associated with tooling, labor, and maintenance
should, of course, be minimized within the constraints of the entire
process.
Hole geometry: The process should accommodate the hole geometry. For
example, an S-shaped hole intended for a Class-A surface (i.e. surface
easily visible to the customer) will probably require a laser cutting
process, while a simple circular hole on a hidden surface will probably
only require a conventional punch process.
Application: The attachments, adaptors, and/or mating components
should influence the type of hole-producing process employed. For
example, if the application for the hole is to receive a wire harness plug,
then there may be little need for a high tolerance hole with no visible
burrs.
Design Stage: The stage of design for the product should also influence
the process selected. For example, early-phase prototypes typically do
not require high-precision machining, while near-production prototypes
may require more precision.
Throughput: The desired throughput time associated with the overall
production of the part should also be taken into account. Because the
time required for hole production varies with respect to the process,
those working on the development of the prototype should consult
closely with design engineers responsible for the production-level manufacturing processes.
Accessibility: The tooling used to create the hole must have relatively
easy access to the hole location. If, for example, a supporting internal
mandrel could not be positioned at the hole location, then the conventional punching method may not be an option.
Slug and chip removal: Handling of the slug (portion or metal removed)
and metal chips after the creation of the hole must always be addressed.
In some cases, it is acceptable to allow the slug to remain inside of the
tube. In other cases, the slug and chips must be removed in order to
avoid unacceptable rattling noises emanating from within the hydroformed parts.
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•
•
Hydroforming for advanced manufacturing
Part structural integrity: The propensity of the hydroformed tube to collapse or severely distort should be known before any hole-producing
operation. For certain cases, finite element analysis may be necessary to
ascertain the structural integrity of the part. In most cases, however,
rules of thumb and judgment based upon experience will suffice.
Die design: The ability of the die to withstand the internal forces imposed
by the expanding tube is also important. Holes may be created immediately after the expansion process while the tube is still in the die, or
after the tube has been removed from the die. In the former case, access
holes/slots must be featured in the die design. Such features can lead to
high stress concentrations in the die, which may compromise its structural integrity.
9.2.2 Hydropiercing processes
Hydropiercing is the process by which a hole is created in the interior
portion of the hydroformed component. The portion of material removed
from the main body is called a slug. Hydropiercing may be divided into
two categories; they are inward hydropiercing, and outward hydropiercing. The schematic in Fig. 9.1 illustrates the inward hydropiercing process
Countersink
A
C
B
D
9.1 Cross-section schematic illustrating inward hydropiercing with
uncoupled slug (A), inward hydropiercing with attached slug
(B), hydropiercing with flanged land for self-threading screws (C),
outward hydropiercing with uncoupled slug (D).
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9.2 Picture of structural member with large and small inward
hydropierced holes (courtesy of Hydrodynamics Technologies, Inc.,
Auburn Hills, MI).
(A). For inward hydropiercing, the material is forced into the interior of
the tube cavity while the tube is under a high internal pressure condition.
The agent for shearing the slug is the applied force of a rigid tool. If
the slug is completely severed from the tube body, then rattling noises
from within the tube may develop (if it is not removed). To circumvent
this possibility, the tooling can be designed such that the slug remains
attached to the tube body (B). Circular slug diameters for the inward
hydropiercing processes can be as small as 6.0 mm. When self-threading
fasteners are required, pilot hole punching may be followed by a flanging
process; this sequence of steps provides a larger surface area for thread
engagement (C).
The outward hydropiercing process is illustrated at point D. For outward
hydroforming, the agent for removing the slug is the force associated with
the internal hydraulic pressure. The slug is allowed to be forced into a die
cavity, where it may be recovered. The outward hydroforming process is
better suited for larger holes or slots (Singh, 2003). In Fig. 9.2, a picture of
a structural member with both a large and small inward hydropierced hole
is shown. In the following sections, some analytical models for hydropiercing are presented.
9.2.3 Mechanics of inward hydropiercing
In Fig. 9.1a, a depressed region in the metal, called a countersink, will
develop for inward hydropiercing. The boundary conditions for the countersink region are shown in Fig. 9.3. For some cases, the presence of a
pronounced countersink is benign. For other cases, the presence of a
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M
M
FT
T
F
T
F
Δ2
P
Δ1
DT
9.3 Two-dimensional section of axisymmetric countersink body during
inward hydropiercing process.
countersink of any magnitude may be prohibitive. Therefore, it is desirable to have access to some analytical tools for the purpose of predicting
the geometry of the countersink as a function of the process variables.
Analytical modeling of the countersink geometry may be approached
from more than one perspective. One approach is to consider an energy
balance (Lu, 2002). The general idea is to consider the work associated
with tearing/shearing, bending, friction, and hydraulic pressure. (In Lu’s
analysis, square tubes with no internal pressure were considered.) For
the energy approach, the total work done by the external force, We, is
given as
We = FTd
[9.1]
where FT = applied tool force (net) and d = tool travel distance.
The internal energy components, Wi, are
Wi = WR + WB + WF
[9.2]
where the subscripts R, B, and F, denote tearing, bending, and friction,
respectively.
For this approach, Lu employed empirically based calibration factors. In
principle, this approach could be employed in order to introduce a closed
form analytical model for the prediction of the countersink geometry. For
this chapter, however, an alternative approach, which is based upon force
equilibrium, is introduced due to its relatively simple approach.
The countersink region for a circular hole punch may be defined by the
dimensions D1 and D2 as shown in Fig. 9.3. The dimension D1 is called the
countersink width, and the dimension D2 is called the countersink depth. In
this section, force equilibrium forms the basis of an analytical model (Choi,
2004) which may be used to predict the countersink geometry.
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In Fig. 9.3, the two-dimensional section of the axisymmetric countersink
body is shown. Dynamics effects are ignored. Due to naturally arising
bending effects, the countersink region features a certain amount of curvature. For this model, membrane behavior of the metal in the countersink
region is assumed, however. Therefore, for this analysis, the angle f is constant, the moment M is defined as zero, and the membrane force T is
assumed to be directed in the x direction. (It is acknowledged that, if f is
constant, then the membrane force T must have some component in the y
direction; the calibration factor in the below equation accounts for this
inconsistency.)
The tool force required to shear off the metal slug is
FS = P π
( ) + πDC tτ .
DT
2
2
T
[9.3]
where t = thickness of the metal, t = shear strength of the metal, DT =
diameter of the punch, P = hydraulic pressure and C = calibration factor to
account for neglecting friction and bending effects.
Choi documents that the sum of the y-direction forces, F, resisting the
applied tool force is
F=P
π
DT2 + ⎡⎣ 4 Δ12 + Δ22 (Δ1 + DT )⎤⎦ cos φ
4
{
}
[9.4]
Equating (9.3) with (9.4) leads to
DT tτ = CP Δ12 + Δ22 (Δ1 + DT )cos φ
and if cos φ =
Δ1
Δ + Δ22
2
1
[9.5]
then,
DT tτ = CP Δ12 + Δ22 (Δ1 + DT )
Δ1
[9.6]
Δ + Δ22
2
1
One way to determine the relationship between D1 and D2 is through
experimental observation. Choi proposes to develop such a relationship
through the experimental evaluation of three different material cases for
which a best fit line may be found as shown below.
D1 = 7.1D2 − 2.8
[9.7]
In the experiments conducted to obtain equation (9.7), a range of hydraulic pressures was considered by Choi. The empirically obtained equation
(9.7) may be applied to equation (9.6) in order to yield the following analytical model for predicting the countersink depth, D2.
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DT tτ = CP (7.1 + 2.8 Δ2 )2 + Δ22 (7.1 + 2.8 Δ2 + DT )
7.1 + 2.8 Δ2 + DT
(7.1 + 2.8 Δ2 + DT )2 + Δ22
[9.8]
Assuming that the material obeys the von Mises yield criterion, the shear
strength is given as
τ=
Y
[9.9]
3
where Y = the uniaxial tension yield strength of the work piece material.
The appropriate calibration factor is found by determining a best match
between theory and experiment. Choi found that, for the specific case of t
= 10 mm and DT = 10 mm, the calibration factor C = 3 produced the best
match. The engineer should be aware that because the calibration factor C
may vary from case to case, a different value may be more applicable for
their particular case. The following results, therefore, should be applied
judiciously.
Application of equation (9.9) to equation (9.8), and using C = 3, leads
to
DT t
Y
3
= 3P (7.1 + 2.8 Δ2 )2 + Δ22 (7.1 + 2.8 Δ2 + DT )
7.1 + 2.8 Δ2 + DT
(7.1 + 2.8 Δ2 + DT )2 + Δ22
[9.10]
Finally, the tool force FT required to shear off the metal slug for the
inward hydropiercing process is
FT = πDT t
Y
3
+
πDT2
P
4
[9.11]
From equations (9.10) and (9.7), the countersink depth D2 and countersink width, D1 can be predicted. Some important points related to the model
are:
• Both countersink depth and countersink width decrease with increasing
internal pressure.
• Both countersink depth and countersink width increase with increasing
tube thickness
• The results are based upon a simplifying assumption, for which bending
and internal shear stress effects are not explicitly included.
• A calibration factor of three was employed, but it need not be the same
for all cases.
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•
209
Caution should be exercised in using the proposed model for cases
featuring material properties, thicknesses, and pressures significantly
different than those considered in its derivation.
9.2.4 Mechanics of outward hydropiercing
Analysis of the mechanics of the outward hydropiercing process is more
straightforward than that of the inward hydropiercing process. As pressure
in increased, the absence of a rigid support behind the prescribed hole
region allows the metal to be pushed outwardly into the die cavity (point
D in Fig. 9.1). Because the expanding metal is surrounded by a rigid support,
problems associated with countersink geometry do not prevail. The following equation arises through force equilibrium considerations:
Pπ
( ) = τπD t
DT
2
2
[9.12]
T
Applying (9.9) to (9.12) leads to
PDT Yt
=
4
3
[9.13]
Equation (9.13) may be used to approximate, for example, the required
pressure to shear off a slug for the outward hydropiercing process.
9.3
End-cutting and saw-cutting
9.3.1 End-cutting
After the hydroforming and hydropiercing processes are completed, the
ends of the part usually require further processing so that the part may be
joined to a larger assembly. Although it is possible to achieve the desired
end location and geometry through the application of appropriate axial end
forces/tools, proper end geometry usually involves some form of cutting.
Three of the main end-cutting processes for tube hydroforming are sawcutting, shearing, and laser-cutting.
9.3.2 Saw-cutting
Figure 9.4 shows a simplified schematic of the saw-cutting process. It is,
perhaps, the most simple and cost-effective method of end-cutting a tube
hydroformed part. The saw cutting process may be carried out manually or
fully automatically. One advantage of saw cutting is that is generates a clean,
burr-free edge, if the tooling is properly maintained. The saw-cutting
operation can only generate straight cuts, though (Fig. 9.5). For this reason,
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9.4 Simplified schematic of saw end-cutting process.
9.5 Picture of structural member with saw-cut end (courtesy of
Hydrodynamics Technologies, Inc., Auburn Hills, MI).
alternative methods (such as shearing and laser-cutting) are selected when
ends featuring curves or corners are required.
9.3.3 Outside shearing
For the outside shearing operation, the blade enters the work piece from
the outside surface of the part as shown in Fig. 9.6. Unlike the saw-cutting
operation, the outside shearing operation can lead to the presence of burrs
if the tooling is not maintained well. The outside shearing operation sometimes can lead to distortion or partial collapse of the cross-section if adequate planning is not made. Some of the variables that influence the
propensity of cross-section collapse are:
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9.6 Simplified schematic of shearing inward end-trim process.
9.7 Simplified schematic of shearing (shimmy) outward end-cutting
process.
•
•
•
•
•
•
Tube thickness
Tube-cross section geometry
Material properties
Tooling geometry
Cutting blade sharpness
Blade location and angle (relative to work piece and tooling)
9.3.4 Inside (shimmy or brehm) shearing
The shimmy trim operation involves a set of shearing tools which are driven
outward as illustrated in Fig. 9.7. The support from the tooling on the
outside of the part provides excellent resistance to cross-section distortion.
The shimmy trim operation is more accommodating to complex end geometries, relative to the outside shearing operation (Goode, 2001). One of the
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9.8 Simplified schematic tube hydroformed part with hole and slot
produced by laser-cutting process.
most prominent limitations of this method, however, is that the trim tooling
must be able to fit into the open end of the tube. Therefore, the tooling for
the application of axial feed and end sealing must be designed under the
expectation that a shimmy trim operation will be employed.
9.3.5 Laser-cutting
Laser-cutting involves the projection of a high-powered, concentrated light
source onto the surface of the metal. The extremely narrow laser beam
allows for a very clean (burr-free) and precisely defined cut. It is far more
versatile than the saw-cutting and shearing operation. The cost, however, is
typically four to five times that of the saw-cutting or shearing operation
(Sing, 2003). The laser head, if secured to a robotic arm, can be controlled
by CNC programming. The design possibilities associated with the lasercutting method are great. Fig. 9.8 shows a schematic of a product with a
hole and slot produced by the laser welding operation.
9.4
Welding
9.4.1 Preliminary comments
Each individual tube hydroformed part features a relatively high stiffnessto-weight ratio. In most cases, however, the individual hydroformed part is
designed to be one of several hydroformed components within a larger
structural assembly. The benefit of high stiffness-to-weight ratios in individual components can be completely negated by ill-functioning joining
mechanisms throughout the overall assembly. Therefore, the joining mechanisms must be carefully considered during the design of each individual
component. There are many types of joining mechanisms for tube hydroformed parts. Some of them are threaded fasteners, push pins, clips, rivets,
and weld nuggets. In this chapter, only welded joints are considered. The
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practice of welding is a mature discipline. Detailed analyses and descriptions of the many welding techniques are thoroughly documented in literature (Kalpakjian, 2003). The purpose of this section is to provide a very
brief overview of the three most common welding techniques for tube
hydroformed parts; they are gas–metal–arc welding (GMAW), spot welding,
and laser welding.
9.4.2 Gas–metal–arc welding
In gas–metal–arc welding (GMAW), sometimes called metal inert gas (MIG)
welding, a filler wire is fed from a coil directly to the weld region on the
parent material. When contact between the filler wire and the parent material is established, an electrical arc is generated. The resulting current flow
produces enough heat to melt the filler wire and surrounding regions of the
parent material. The weld region is engulfed within a cloud of inert gas such
as argon, helium, carbon dioxide, or some type of gas mixture, in order to
prevent oxidation with oxygen and nitrogen in the atmosphere. This process,
which was developed in the 1950s, is suitable for many ferrous and nonferrous materials used in the metal fabricating industry. Because the process
lends itself well to automated, robotic manufacturing systems, it is economical and accommodating to high throughput demands. One of the great
advantages of the GMAW process within the context of tube hydroforming
is the fact that only access from the outside of the tube is required. It should
be noted that the minimum thickness of the parent material is about 0.05 in.
(Longhouse, 2000). An illustration of a bracket joined to the tube hydroformed part through the GMAW process is shown in Fig. 9.9.
GMAW
Spot
welding
Laser
welding
Spot
welding
9.9 Cross-section schematic illustrating attachments joined by the
GMAW, spot weld, and laser weld processes.
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Hydroforming for advanced manufacturing
9.4.3 Spot welding
In spot welding, two metal work pieces are pressed together by two
copper electrodes through which electrical current is conducted (Fig. 9.9).
The resulting current flow generates enough heat to melt the metal pieces,
thereby forming an internal weld zone. Unlike the GMAW process, spot
welding does not require a filler metal rod. In order to help promote a
quality weld joint, an appropriate amount of clamping force is required.
On one hand, if the stiffness of the work piece is too low, then the propensity for improper alignment between the electrode and the metal
surface may arise, resulting in a substandard weld joint. On the other
hand, if the stiffness of the work pieces is too high, then the propensity
for not bringing the two work pieces into sufficiently close proximity may
arise, resulting in a substandard weld joint also. The required electrode
tip force for a typical spot weld for a tube hydroformed part is 500 lb,
approximately (Singh, 2003). As can be seen in Fig. 9.9, an access hole
for the electrode is often required for the spot welding for tube hydroformed parts; this represents a significant limitation for the spot welding
process.
9.4.4 Laser welding
The laser welding process involves the application of a laser beam onto a
local region at the interface of two metal pieces. The laser beam may be
guided by CNC controlled robots (Kreis, 2001). The laser welding process
requires a very precise geometric match between the two parts to be joined.
Maintaining such precision can be relatively expensive. Accordingly, the
laser welding process is not as popular as the GMAW or spot weld process
for tube hydroformed parts.
9.4.5 Distortion due to welding
During pre-bending and expansion of the tube within the die, residual
stresses within the work piece will develop. A well-designed process will
account for the presence of residual stresses, which factor into the sum of
the equilibrating forces within the work piece. Depending on how cooling
is controlled (if at all), the high temperatures introduced by the welding
process may totally change the internal stress distribution in the heat
affected zone. Such changes in the internal stress distribution can lead to
geometric distortion of the work piece, with respect to the desired work
piece final geometry. Some rules of thumb for reducing the effect of weld
temperatures on the final geometry of the tube hydroformed part are given
below:
WPNL2204
Hydroforming: hydropiercing, end-cutting, and welding
215
• Introduce welds in regions where the residual stresses are the least.
• Introduce short-stitch weld lines instead of long continuous welds.
• Apply welds symmetrically.
• Specify a weld sequence such that a build-up of heat in any one weld
region can be prevented.
9.5
References
choi s k, kim w t and moon y h (2004), ‘Analysis of deformation surrounding a hole
produced by tube hydro-piercing’, Proc Inst Mech Eng, J Eng. Manuf 218 B,
1091–1097.
goode m (2001), ‘Trimming hydroformed parts using cam trim dies’, Hydroforming
Journal (A Supplement to TPI – The Tube and Pipe Journal, March 19).
kalpakjian s and schmid s (2003), Manufacturing processes for engineering materials, 4th Ed., Prentice Hall.
kreis o and hein p (2001), ‘Manufacturing system for the integrated hydroforming,
trimming and welding of sheet metal pairs’, J Mat Proc Tech, 115, 49–54.
longhouse b (2000), ‘Hydroformed tubes design considerations for joints and
attachments’, Tube/Pipe Fabricating Conference, May, TPA.
lu g and wang x (2002), ‘On the quasi-static piercing of square metal tubes’, Int J
Mech Sci, 44, 1101–1115.
singh h (2003), Fundamentals of hydroforming, Society of Manufacturing
Engineers.
WPNL2204
10
Hydroforming sheet metal
forming components
K . S I E G E R T and S . W A G N E R,
University of Stuttgart, Germany
10.1
Introduction
Hydroforming of sheet-metal components requires high-capacity presses,
special deep drawing dies and special high-pressure hydraulic components.
To compensate for these disadvantages, new designs for presses and dies as
well as new optimized hydroforming processes have been developed. All
these developments have made hydroforming sheet metal components
more competitive, especially for low-volume production, but also in some
cases for high-volume production. In the automotive industry there is a
need for low-volume production of auto-body components. Aluminium, and
high-strength and ultra-high-strength steels are the first choices of material
for reducing the weight of the cars, in order to reduce the fuel consumption.
In nearly all cases, this accompanies the need to produce high quality, difficult-to-form components at low cost.
10.2
Hydroforming processes
In hydroforming of sheet metals, it is possible to differentiate between:
• Hydraulic stretch forming;
• Hydromechanical deep drawing axisymmetric round products and nonaxisymmetric products;
• Combination of hydraulic stretch forming with hydromechanial deep
drawing; and
• Combination of conventional deep drawing with hydromechanical deep
drawing.
It is useful to compare the hydromechanical deep drawing process with
the conventional deep drawing process in order to discuss the advantages
and disadvantages of hydromechanical deep drawing processes. In addition,
when talking about hydraulic stretch forming and the combination of conventional deep drawing with hydraulic stretch forming, the possibility of
forming double blanks should be discussed.
216
WPNL2204
Hydroforming sheet metal forming components
Fst
Blankholder
with upper
pressure plate
pi
Fst
217
Blankholder
with upper
pressure plate
pi
Blank
Draw Ring
Fst
Fst
Blank
Lower die
(female die)
b
a
Fst
pi
Fst
c
10.1 Hydraulic stretch-forming of single blanks: a bulge test, b with
rigid lower die, c with rigid upper die (source: IFU Stuttgart).
10.2.1 Hydraulic stretch-forming
Figure 10.1 shows the principles of the hydraulic stretch-forming process.
The blank is tightly clamped and sealed around the draw ring opening (Fig.
10.1a) or outside the cavity of the upper or the lower die (Fig. 10.1b and c).
The hydraulic pressure p forms the sheet metal through the draw ring
opening (Fig. 10.1a) or into the cavity of the upper or the lower die (Fig.
10.1b and c).
Stretch-forming a blank through an open draw ring is known as the Bulge
Test, which is used for determining the Limited Dome Height Value (LDH
Value), Flow Curves and Forming Limit Curves (FLC). For the Bulge Test,
round and elliptical draw rings are used. If the inner diameter of the draw
ring DDR in relation to the thickness t of the blank is greater than or equal
to 1000, i.e.
DDR
≥ 1000
t
[10.1]
one can assume that the bulge can be seen as a membrane element
(Gologranc, 1977) (see Fig. 10.2).
With this assumption, one can calculate the main stresses s1 and s2 given
radii d1 and d2, thickness t of the sheet metal at the pole of the bulge,
hydraulic inner pressure pi and hydraulic outer pressure pa (counter
pressure).
WPNL2204
218
Hydroforming for advanced manufacturing
sy
sx
r2
sy
r1
sx
z
x
y
10.2 Membrane element at the pole of the bulge under stress–stress–
load when hydraulic stretch forming after (Gologranc, 1977).
σ1 =
pi − pa
δ2
2t
[10.2]
σ2 =
pi − pa ⎡
δ 22 ⎤
2
δ
2 −
2t ⎢⎣
δ 1 ⎥⎦
[10.3]
According to (Panknin, 1959) one can find the third main stress with
normal pressure pi acting on the inner blank surface and normal pressure
pa acting on the outer surface of the blank under the condition
pi ≥ pa
σ3 =
[10.4]
1
( pi + pa )
2
[10.5]
If there is no outer pressure, pa = 0 (counter pressure), which is normally
the case, and if a round draw ring is used, assuming d1 = d2 = d and s1 = s2 =
s, the pressure p = pi for forming the bulge can be described by the
equation
p=
2t
σ
δ
[10.6]
From Tresca’s flow hypothesis and taking equations (10.2) and (10.6) into
account, it can be said that the flow stress kf (true stress) is:
kf = σ max − σ min
σ max = σ 1 = σ 2
σ min = −σ 3
kf = σ 1 − σ 3
p
p
kf = δ +
2t
2
p δ
kf = ⎡ + 1⎤
⎦⎥
2 ⎣⎢ t
[10.7]
WPNL2204
Hydroforming sheet metal forming components
219
With t as the actual thickness of the sheet metal at the pole of the bulge
and t0 as the exit thickness of the blank, the equivalent strain jg at the
pole is
ϕ g = { ϕ 1 ; ϕ 2 ; ϕ 3 }max
ϕ g = ln
[10.8]
t
t
= ln 0
t0
t
By measuring the pressure, the radius of the bulge and the thickness of
the sheet metal at the pole, it is possible to get flow curves kf = f (jg) with
the hydraulic bulge test.
For non-axisymmetric components, it is possible to describe the process
of hydraulic stretch forming the sheet metal into the cavity of a rigid die
(Fig. 10.1b and c) by an analytical model. It was shown by (Geiger, 2005)
that the contour of the component can be approximated by different elliptical regions.
Banabic (1999) elaborated on an analytical model for elliptical bulging,
taking the anisotropy of the sheet metal into account. Analytical models
like that developed by Banabic have the advantage of a fast calculation for
the pressure needed for hydraulic stretch forming non-axisymmetric sheetmetal components. A more accurate analysis may be obtained by performing Finite Element Method Process Simulations.
The pressure at the end of the stretch-forming process is of great
importance. Here a high-pressure calibration is used to form the sheet
metal into the cavity of the die. The smaller the radii of the component and the higher the strength and the thickness of the sheet metal,
the higher the pressure has to be. Figure 10.3 shows results of experimental investigations carried out for fluid cell presses. It can be seen
that high pressure is needed to form parts with small radii (Johannisson, 2001).
By numerical simulation, it is possible to find stresses, strains, thickness
and geometry of the workpiece related to the pressure that closely approximate experimental results.
Figure 10.4 shows a hydraulic stretch-formed tray and hydraulic stretchformed hump element plates.
Figure 10.5 shows a component (door outer, scale 1 : 2) of a passenger car.
It was stretch formed into a female die. The advantages of this process are
inexpensive dies and the excellent surfaces and geometry of the workpiece.
In addition, it only takes a short time to develop and make up the die. Disadvantages are the low cycle time (∼1 part per minute), the need for highpressure hydraulic equipment and the high capacity of the clamping device,
which might be a press.
Figure 10.6 shows the production of automotive components by hydraulic
stretch forming double blanks welded together before forming. This process
WPNL2204
Ratio of radius to thickness of sheet metal (R/t)
220
Hydroforming for advanced manufacturing
8
R
7
t
6
5
4
Stainless steel
3
2
Mild steel and
Al (AA 2024, solutionised)
Aluminium (AA 2024-T0)
Aluminium (Pure)
1
0
500
1000
1500
2000
Pressure (bar)
10.3 Radius R related to the sheet metal thickness t of hydraulic
stretch-formed components versus the pressure p obtained by
experimental investigations with a fluid cell press (Source:
Johanisson, 2001).
10.4 a Tray from CuZn36-brass and b hump element plate with mirror
finished outsight surface (right) produced by hydraulic stretch forming
(Source: HDE, Germany).
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Hydroforming sheet metal forming components
221
10.5 Hydraulic stretch formed door outer scale 1 : 2 (pressure max. 250
bar, DQ-steel DC05, sheet thickness t = 0.6 mm) and experimental set
up (Source: IFU Stuttgart).
U1 ≅ 2 × U0
U0(100%)
10.6 Hydraulic stretched formed autobody components from double
blanks welded together before forming using internal high-pressure
forming (Source: Thyssen Company, Germany).
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222
Hydroforming for advanced manufacturing
is a hydraulic stretch-forming process using internal high pressure. There
are special patented pressure inlet systems in use.
10.2.2 Hydromechanical deep drawing compared with
conventional deep drawing
It might be useful to compare the hydromechanical deep-drawing process
with the conventional deep-drawing process to show the various advantages
and disadvantages of the hydromechanical deep-drawing process.
Processes
Figure 10.7 shows the principles of the conventional deep-drawing process
and Fig. 10.8 the hydromechanical deep-drawing process in four steps. It
can be seen that with hydromechanical deep drawing as opposed to conventional deep drawing, instead of a rigid female die, hydraulic pressure
forms the sheet against the contour of the punch. This pressure is built up
by compression when the punch is forced downwards. The pressure over
the stroke can be regulated by a pressure relief valve. Because a highpressure pump is not necessary for this, the hydraulic system is called a
passive hydraulic system. After reaching the final forming position, the
punch travels upwards, the produced sheet-metal component can be taken
out and a low-pressure filling pump fills the counter pressure pot until the
beginning of the next cycle is reached.
Forces
For conventional deep drawing, Fig. 10.9 shows the ideal force Fid, and
the three friction forces FFriction/Sheet/BH to overcome the friction between
sheet and blankholder, FFriction/Sheet/DR to overcome the friction between sheet
and draw ring and FFriction/Sheet/DRR to overcome the friction between the sheet
and the round draw ring edge.
In addition, one has to consider the bending force FB for bending the
sheet around the round draw ring edge.
When using drawbeads for controlling the metal flow between the blankholder and the draw ring, one has to consider the drawbead force FDB, which
is built up by bending and friction forces acting on the sheet metal, as a
restraining force when pulling the sheet through the drawbead.
The ideal force Fid is the punch force for the ideal condition where there
are no friction or bending forces. This force cannot be reached in reality.
For axisymmetric parts, the ideal force Fid is needed to plastically form the
sheet metal between the outer edge of the blank Douter and the inner edge
WPNL2204
Hydroforming sheet metal forming components
223
Punch (male die)
Upper binder
(Blank holder)
Blank
Draw ring (lower binder)
Matrix (female die)
a
FBH
FBH
b
FBH
FP
FBH
c
FBH
FP
FBH
d
10.7 Conventional deep drawing: a outgoing situation, inserting the
blank; b closing the binders, introducing the blankholder force FBH;
c deep drawing with the punch force FP and the applied blankholder
force; d forming the component between the punch and the matrice
(female die).
WPNL2204
Punch
Upper binder
Sealing
Blank
Counter pressure pot
Pressure relief and
safety valve
a
FBH
Filling pump
FBH
b
FBH
FP
FBH
c
FBH
FP
FBH
d
10.8 Hydromechanical deep drawing: a filling the counter pressure pot
and inserting the blank; b closing the binders and introducing the
blankholder force; c hydromechanical deep drawing with the punch
force FP, the blankholder force FBH and the counter pressure p;
d reaching the bottom dead center, calibrating the work piece
by forming against the male die with high counter pressure.
WPNL2204
Hydroforming sheet metal forming components
FBH
1 Blankholder
(upper binder)
2 Draw ring
(lower binder)
3 Female die (matrice)
4 Male die (punch)
5 Blank
FPunch
Douter
DDB
DBH
DPunch
1
Douter Outer edge of the blank
Position of the draw bead
DDB
5
DBH
Inner diameter of the blankholder
DDR
DPunch Diameter of the punch
DDR
Inner diameter of the draw ring
rDR
Radius of the inner draw ring edge
rPunch Radius of the punch
h
Draw depth
F1 = Fid1 + FFriction/Sheet/BH + FFriction/Sheet/DR
rDR
rPunch
2
3
FFriction/Sheet/BH
Fid1
225
FFriction/Sheet/DR
F1 + Fid2 + FDB
4
h
F2 = F1 + Fid2 + FDB
F2
FFriction/Sheet/DRR
Fid3
FB
FTotal = FB + FFriction/Sheet/DRR + Fid3
FTotal = F2 + FFriction/Sheet/DRR + Fid3 + FB
FTotal = Fid1 + FFriction/Sheet/BH + FFriction/Sheet/DR + Fid2 + FDB + FFriction/Sheet/DRR + Fid3 + FB
Respecting
Fid = Fid1 + Fid2 + Fid3
One gets
FTotal = Fid + FFriction/Sheet/BH + FFriction/Sheet/DR + FFriction/Sheet/DRR + FB + FDB
10.9 Forces acting on the sheet metal when conventional deep
drawing.
D0
Douter
jgouter
kf
kfinner
kfouter
Dpunch
jginner
jgouter
jginner
jg
10.10 Determining the true stresses of the outer edge of the blank and
of the inner edge of the draw ring.
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226
Hydroforming for advanced manufacturing
of the draw ring DDR. This force can be calculated with the equation
(Panknin, 1959):
⎞
⎛D
Fid = Dpunch πtkfm ln ⎜ outer ⎟
⎝ Dpunch ⎠
[10.9]
The medium true stress (medium flow stress) kfm = 0.5(kfo + kfi), which is
the average of the flow stress at the outer edge of the blank kouter and at the
inner edge of the drawring opening kinner can be determined (Panknin, 1959)
with:
ϕouter = ln
( DD )
0
outer
and
ϕ ginner = ln
2
2
D02 + Dpunch
− Douter
2
dpunch
[10.10]
out of the flow curve kf = f(jg) of a sheet metal (see Fig. 10.10).
Blankholder force
The blankholder pressure pn, which is the normal pressure between the
sheet and the blankholder and between the sheet and the draw ring, should
be high enough at least to avoid wrinkles of type A in the sheet guided
between blankholder and draw ring. For axisymmetric components, according to E. Siebel (Siebel, 1954) one finds
Dpunch ⎤
pn = 0.002 . . . 0.0025 ⎡(β0 − 1)3 + 0.5
σ UTS
⎢⎣
100t0 ⎥⎦
[10.11]
D0
, D0 = Outgoing blank diameter, Dpunch = Punch
Dpunch
diameter, t0 = Outgoing blank thickness and sUTS = Ultimate tensile stress.
With non-axisymmetric sheet-metal components it is necessary to control
the metal flow by adjusting the blankholder force properly and/or by using
drawbeads to control the metal flow. To avoid wrinkles of type A, the blankholder force FBH should take into account the results of equation (10.11),
at least up to the point where
where draw ratio β0 =
π
2
FBH ≥ pn ⋅ (D02 − Dpunch
)
4
[10.12]
To avoid splitting it is necessary to use the equation
σ Fracture =
FFracture
≤ aRσ UTS
t
WPNL2204
[10.13]
Hydroforming sheet metal forming components
227
with aR ≥ 1 as a fracture factor according to Doege (1963), taking into consideration the transfer of forces from the punch to the forming zone between
the binders, by friction between the punch and sheet metal, with FFracture as
fracture force, sFracture as fracture stress, sUTS as the ultimate stress of the
sheet metal, ᐉ as a certain length of the product wall and t as the thickness
of the wall of the component.
Figure 10.9 shows the forces acting on the sheet metal for conventional
deep drawing.
Considering
Fid = Fid1 + Fid2 + Fid3
[10.14]
FFriction/Sheet/BH = FFriction/Sheet/BH1 + FFriction/Sheet/BH2
[10.15]
FFriction/Sheet/DR = FFriction/Sheet/DR1 + FFriction/Sheet/DR2
[10.16]
and
with FB = force for bending the sheet around the inner drawing ring edge,
FDB = force for pulling the sheet metal through the draw beads and FFriction/
Sheet/DR = force to overcome the friction between the sheet metal and the
round draw ring edge, the punch force Fpunch acting on the sheet metal for
conventional deep drawing may be obtained.
Fpunch = Fid + FFriction/Sheet/BH + FFriction/Sheet/DR
+ FDB + FFriction/Sheet/DRR + FB
[10.17]
Figue 10.11 shows the forces acting on the sheet metal for hydromechanical deep drawing. For the hydromechanical deep-drawing process, the situation regarding the stresses and strains in the sheet metal between the outer
blank diameter Douter and the seal (diameter Ds) is the same as with conventional deep drawing for the same blank form and size.
The differences between conventional and hydromechanical deep
drawing are the different stresses, strains and forces inside the sealing
(diameter Ds). Here the hydraulic pressure p (counter pressure) has to be
taken into account.
10.2.3 Forces related to the counter pressure in
hydromechanical deep drawing
When looking at the differences between conventional and hydromechanical deep drawing processes, it has to be borne in mind that there is absolutely no difference between the outer edge of the blank (Douter) and the
sealing (DS).
In the area between Douter and Ds the blankholder pressure is the same for
both processes [see equation (10.11)]. But normally, with hydromechanical
deep drawing, the blank is bigger than with conventional deep drawing.
WPNL2204
FBH
1 Blankholder
(upper binder)
2 Draw ring
(lower binder)
3 Counter pressure pot
4 Male die (punch)
5 Blank
FPunch
Douter
DDB
DBH
Ds
DPunch
1
rDR
rPunch
2
4
Douter Outer edge of the blank
DDB
Position of the draw bead
Inner diameter of the blankholder
DBH
Ds
Position of the sealing
DPunch Diameter of the punch
DDR
Inner diameter of the draw ring
rDR
Radius of the inner draw ring edge
rPunch Radius of the punch
h
Draw depth
pc
Counter pressure
h
pc
5
DDR
3
FFriction/Sheet/BH
Fid1
F1 = Fid1 + FFriction/Sheet/BH + FFriction/Sheet/DR
FFriction/Sheet/DR
F1 + Fid2 + FDB
F2
FFriction/Sheet/DRR
F2 = F1 + Fid2 + FDB
Fid3
FB
FTotal = FB + FFriction/Sheet/DRR + Fid3
FTotal = F2 + FFriction/Sheet/DRR + Fid3 + FB
FTotal = Fid1 + FFriction/Sheet/BH + FFriction/Sheet/DR + Fid2 + FDB + FFriction/Sheet/DRR + Fid3 + FB
Respecting
Fid = Fid1 + Fid2 + Fid3
One gets
FTotal = Fid + FFriction/Sheet/BH + FFriction/Sheet/DR + FFriction/Sheet/DRR + FB + FDB
10.11 Forces acting on the sheet metal when hydromechanical deep
drawing.
DBH
d02
Punch
Blank holder
Douter
Ds
DContact
t
h
DBH
Draw ring
pc
d01
t
DContact
pc
Counter
pressure
pot
s1
s2
10.12 Forming a bulge with hydromechanical deep drawing
components with tapered shaped walls.
WPNL2204
Hydroforming sheet metal components
229
Inside the sealing (D ≤ Ds) for components with tapered shaped
walls, one has to differentiate between three areas (see Fig. 10.12) as
follows.
Area I
In this area, the sheet metal is pressed by the counter pressure pc against
the punch. The outer contour of this area is the contact line Dcontact = f(h)
where the sheet metal comes in contact with the punch (see Fig. 10.10). The
vertical force acting on the punch in this area is
Fpunch/counter pressure =
π
2
pc Dcontact
4
[10.18]
Area II
In this area, the sheet metal is made to bulge upwards by the counter pressure against the travel direction of the punch where there is a critical gap
between DBH and Dcontact with respect to the yield stress sy and the thickness
t of the sheet metal.
This area is between the inner diameter of the blank holder (DBH) and
the contact line (Dcontact = f(h)). The force acting on the bulge is
FBulge/Vertical =
π
2
2
pc (DBH
− Dcontact
)
4
[10.19]
This force acts vertically on the blankholder and on the punch.
Incorporating the stresses s1 and s2, according to equation (10.19), see
Fig. 10.12, gives
σ 1 πDBH t + σ 2 πDcontact t = pc
π 2
2
(DBH − Dcontact
)
4
[10.20]
If s ≈ s1 ≈ s2, then
σ=
pc
(DBH − Dcontact )
4t
[10.21]
As shown by Khandepakar (2007), the bulge force acting on the punch
is
FPunch/Bulge = sDcontactpt
[10.22]
and the bulge force acting on the blankholder is
FBH/Bulge = sDBHpt
[10.23]
WPNL2204
230
Hydroforming for advanced manufacturing
Area III
This area is between the sealing (Ds) and the inner diameter of the blankholder. In this area, the counter pressure acts in the gap between the draw
ring and the sheet metal. The resulting vertical force acting on the blankholder is
FBH/counter pressure =
π
2
pc (Ds2 − DBH
)
4
[10.24]
Due to the counter pressure it can be assumed that, with hydromechanical deep drawing, no bending force is needed to bend the sheet metal
around the round draw ring or, in the case of a bulge, around the blankholder edge, FB ≈ 0.
Furthermore, because there is no contact between the sheet metal and
the inner round draw ring edge, the friction force is zero.
Ffriction/sheet DRR = 0
[10.25]
The following table shows the forces acting on the sheet metal for
conventional and hydromechanical deep drawing. It is obvious that
with hydromechanical deep drawing, there are much greater vertical
forces.
Table 10.1 Comparison between conventional and hydromechanical
deep drawing
Conventional deep drawing
Hydromechanical deep drawing
FPunch = Fid
+ FFriction/sheet/BH
+ FFriction/sheet/Draw ring
+ FFriction/sheet DRR
+ FDrawbead
+ FBending
FPunch = Fid
+ FFriction/sheet/BH
+ FFriction/sheet/Draw ring
+0
+ FDrawbead
+0
+ FPunch/Bulge
+ Fcounter pressure
FBH = pn
π 2
(D a − D 2s )
4
+ FBH/counter pressure
π 2
2
(Da − DBH
)
4
FBH = pn
+ FBH/Bulge
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Hydroforming sheet metal forming components
231
10.2.4 Combination of hydraulic stretch forming and
hydromechanical deep drawing
For flat auto-body parts like bonnets, roofs and wings in particular, a combination of hydraulic stretch forming and hydromechanical deep drawing
should be considered to obtain a higher rigidity and greater resistance to
severe impact.
Using hydrostatic pressure for stretch forming the blank before hydromechanical deep drawing results in a significantly more uniform strain distribution and, because of this, in a higher work hardening than is the case
without stretch forming. The sequence of actions for this process combination is shown in Fig. 10.13. The process combination is called ‘prebulging’
followed by hydromechanical deep drawing (Nakamura, 1987) and ‘Active
Hydromec’ (Steinmetz, 1996).
After stretch forming the panel to approx. 2% in the middle area of
the component, reverse drawing can be accomplished by the punch moving
downwards. In this case the pressure is generated ‘passively’ by penetration of the punch or, in other words, by compression of the hydrostatic
medium. However, with reverse drawing, the danger of wrinkling has to
be taken into consideration. For this reason the sheet metal must be in
contact with the punch in the outer surface area of the formed product.
This should be the case at the end of the hydraulic stretch forming
Fpunch
Punch
FBH
FBH
FBH
Fpunch
FBH
Blank
Blankholder
Draw ring
Seal
pc
pc
Counter
pressure pot
(a)
(b)
10.13 Combination of hydraulic stretch forming followed by reverse
drawing and hydromechanical deep drawing: a Hydraulic stretch
forming (prebulging); b Reverse drawing and hydromechanical deep
drawing.
WPNL2204
232
Hydroforming for advanced manufacturing
process and during the hydraulic deep drawing process in order to prevent
marks on the outer surface area from wrinkles ocurring during reverse
drawing.
10.2.5 Combination of conventional deep drawing and
hydraulic stretch forming
Figure 10.14 shows the possibility of conventional deep drawing up to a
given draw depth followed by subsequent hydraulic stretch forming. Here
the sheet metal can be formed (10.14a) into the cavities of the punch or
(10.14b) into the cavities of the female die. The pressure for this purpose
must be applied (‘actively’) from outside by a pressure generator.
Figure 10.15 shows the possibility of conventional deep drawing double
blanks up to a predefined draw depth followed by hydraulic stretch
forming in such a manner that the lower blank is formed into the cavities
FBH
Fpunch
FBH
Blankholder
(upper binder)
Punch
(male die)
Blank
FBH
Fpunch
FBH
Seal
Draw ring
(lower binder)
pc
Counter
pressure
pot
(a)
FBH
Fpunch
FBH
Blankholder
(upper binder)
FBH
Fpunch
FBH
pc
Punch
Blank
Seal
Draw ring
(lower binder)
Matrice
(Female Die)
(b)
10.14 Combination of deep drawing followed by hydraulic stretch
forming the work piece into a the cavities of the punch or b into the
cavity of a matrice (female die).
WPNL2204
Hydroforming sheet metal forming components
Fpunch
Fpunch
FBH
FBH
233
Punch
FBH
FBH
Blankholder
(upper binder)
Sheet 1
Sheet 2
pinner
pinner
Draw ring
(lower binder)
Matrice
(female die)
(a)
(b)
10.15 Conventional deep drawing double blanks; a not welded
together, followed by b reverse drawing and hydraulic stretch forming
the upper blank into the cavities of the punch and the lower blank into
the cavities of the matrice (female die).
of the female die and the upper blank is formed into the cavities of
the punch. The punch can be pulled back in a controlled manner,
if necessary.
The final stretch forming of the two sheet-metal parts not connected to
each other is accomplished by applying hydraulic pressure between the
blanks using a special pressure inlet.
This process combination greatly expands the possibilities for conventional deep-drawing of non-rotationally symmetrical sheet-metal parts, e.g.
structural body parts of passenger cars such as longitudinal members and
cross-members as well as A- and B-pillars, which is of particular importance
in terms of drawing difficult-to-form sheet-metal parts from high-strength
steel sheets and super-high-strength steel sheets as well as from aluminium
sheets.
The following demonstrates the use of this process for forming a petrol
tank. Figure 10.16a shows the conventional deep-drawn parts produced
from a double blank, not welded to each other, Fig. 10.16b shows the lower
blank, hydraulic stretch formed into the cavity of the female die and Fig.
10.16c shows the upper blank bulged upwards and then hydraulic stretchformed into the cavity of the punch. In this process, the punch is withdrawn
in a controlled manner.
The combination of conventional deep drawing of double blanks followed by hydraulic stretch-forming expands the possibilities for the conventional deep drawing of, e.g. structural body parts of passenger cars
including longitudinal members and cross-members, as well as A- and
B-pillars.
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234
Hydroforming for advanced manufacturing
Conventional deep drawn products
from double blanks, not welded
to each other
(a)
(b)
+
Conventional deep drawn and hydraulic
stretch formed part
Conventional deep drawn,
reverse drawn and hydraulic
stretch formed part
(c)
10.16 Combination of conventional deep drawing and hydraulic
stretch forming of a gasoline tank from double blanks, not welded to
each other (source: IFU Stuttgart).
10.3
Dies and presses for hydromechanical
deep drawing
Special dies for hydromechanical deep drawing have been developed at
IFU. They are designed with a pyramidal-shaped blankholder in contact
with a hydraulic multipoint cushion system integrated into the die (see Fig.
10.17). The blankholder forces are introduced downwards via the cylinders
of the multipoint cushion system into corresponding pyramids. These forces
are transmitted to the base of the blankholder pyramids, each covering
separate areas of the blankholder surface.
This setup makes it possible to control the blankholder pressures locally
over defined areas of the blankholder (see Fig. 10.17). By this means it is
possible to optimise the blankholder pressure, area by area. When optimising the blankholder force on a certain blankholder area, the neighbouring
areas are not affected. This makes it possible to adjust the blankholder
WPNL2204
Hydroforming sheet metal forming components
235
Locally increased
blankholder pressure
Draw ring
Sheet
Blankholder
Fpin,increased
10.17 Segment-elastic blankholder (source: IFU Stuttgart).
pressure over the blankholder without adjusting the gap between blankholder and draw ring by hand.
Figure 10.18 shows the punch and the cylinders of the cushion system
integrated into the die for the hydromechanical deep drawing of a car hood
(scale 1 : 1.25). Integrating the cushion system into the die produces higher
die costs, but the advantage is that a standard hydraulic single-acting press
can be used if the die can be mounted in the press and the capacity of the
press is high enough. When introducing the production of hydromechanical
deep-drawn products, normally only a certain percentage of the press availability is needed. The remaining availability can be used for other forming
processes like conventional deep drawing, stretch forming, cutting and
calibrating.
When using a specially developed press for hydromechanical deep
drawing processes, and only a certain percentage of its availability, e.g. 20%,
is needed, 100% of the press costs still have to be attributed to the products
produced, because this press cannot be used for other forming processes.
Therefore, the use of a hydraulic single-acting press is recommended
when introducing the production of hydromechanical deep drawn products.
After increasing the production of hydromechanical deep drawn parts up
to 100% availability of the press, it makes sense to go over to the use of a
special double-acting hydraulic press with a hydraulic multipoint cushion
system integrated into the blankholder ram.
If no hydraulic single-acting press with sufficient capacity and appropriate die clearance, stroke and table dimensions is available, a new press can
be made. To be competitive this new press should be as inexpensive as possible. A low-cost single-acting hydraulic press was built at the Institute for
Metal Forming Technology (IFU), University Stuttgart, (see Fig. 10.19).
This press has a capacity of 20 000 kN which is provided by 4 × 5000-kN
cylinders acting between two frame plates and the ram. No welding is used
in this press. Both frame plates are solid and screwed together horizontally
by tie rods.
WPNL2204
10.18 Die for hydromechanical deep drawing: a punch and blank
holder cylinders of a die for hydromechanical deep drawing a hood;
b tool with integrated multipoint cushion system, segment-elastic
blank holder and counter pressure pot for hydromechanical deep
drawing hoods mounted in a hydraulic single-acting 20 000-kN press;
c hydromechanical deep drawn hood (source: IFU Stuttgart).
1200
Upper traverse with
distance plate
Frame plate (2×)
Hydraulic cylinder (4×)
8000
Ram
Press table
1500
Distance (2×)
Tie rod (4×)
Base plate
2900
0
154
10.19 Single-acting hydraulic press built up at IFU (source: IFU
Stuttgart).
WPNL2204
Hydroforming sheet metal forming components
10.4
237
References
banabic, d (1999), ‘Closed-form solution for bulging through elliptical dies,’ Proceedings of SheMet’99, Sept 27–28, 1999 pp. 623–628.
doege, e (1963), ‘Untersuchung über die maximal übertragbare Stempelkraft beim
Tiefziehen rotationssymmetrischer zylindrischer Teile.’ Diss. TU-Berlin 1963.
geiger, m (2005) ‘Sheet hydroforming: analytical modelling and experimental verification of complex structures’, production engineering, XII(2), 69–72.
gologranc, f (1977), ‘Theoretische Betrachtungen zur Aufnahme von Fließkurven
im kontinuierlichen hydraulischen Tiefungsversuch’, Blech, Rohre, Profile 4/1977
S.116/121.
johanissson, t (2001), ‘Low Volume Production of Sheet Metal Parts’, in:
Hydroforming Tubes, Extrusions and Sheet Metals, ed. K. Siegert, WerkstoffInformationsgesellschaft mbH, Frankfurt 2001, pp. 159–179.
khandeparkar, t (2007), ‘Investigations on hydromechanial deep drawing sheet
metal components with high counter pressure.’ Dissertation Universität Stuttgart,
2007.
nakamura, k (1987), ‘Sheet metal forming with hydraulic counter pressure in Japan.’
Annals of the CIRP, 36(1), 191–194.
panknin, w (1959), ‘Der hydraulische Tiefungsversuch und die Ermittlung von
Fließkurven.’ Dr.-Ing. Dissertation TH Stuttgart 1959.
siebel, e (1954), ‘Der Niederhalterdruck beim Tiefziehen.’ Z. Stahl und Eisen, 4,
155–158.
siegert, k (1999), ‘Sheet Metal Hydroforming’, in: Hydroforming Tubes, Extrusions
and Sheet Metals. Ed. K. Siegert, Werkstoff-Informationsgesellschaft mbH, Frankfurt 1999, pp. 221–247.
steinmetz, m (1996), ‘Aktives Hydromechanisches Tiefziehen für großflächige
Ziehteile mit erhöhter Formsteifigkeit’. Technologiepräsentation der Fa. SMG,
Waghäusel.
WPNL2204
11
Bending and hydroforming of aluminum and
magnesium alloy tubes
A. A. LUO and A. K. SACHDEV,
General Motors Research & Development Center, USA
11.1
Introduction
Tube hydroforming is a metal forming process that uses pressurized fluids
such as water to make various perimeter shapes from tubes. Recently,
hydroformed aluminum parts such as front and rear subframes are being
used in Europe for vehicle weight reduction [1]. The front and rear subframes of the 2005 Acura RL are also made from hydroformed aluminum
tubular members [2]. In North America, the most recent applications of
hydroformed aluminum tubes are the side rail and roof bow for the 2006
Chevrolet Corvette Z06 frame. The side rail is the largest aluminum hydroformed part in the world.
Both 5xxx series (Al-Mg based) and 6xxx series (Al-Mg-Si based) alloys
are used in automotive structural applications. However, there is limited
literature on the bending and hydroforming of these tubes and their
mechanical properties after hydroforming [3–6]. This information is very
important to automotive design and manufacturing engineers.
Magnesium components are being increasingly used by major automotive
companies including General Motors (GM), Ford, Volkswagen and Toyota
[7–14]. Current major automotive magnesium applications include the
instrument panel beam, transfer case, steering components, and radiator
support. However, the magnesium content in a typical 2005 model family
sedan built in North America was only about 0.3 per cent of the total vehicle
weight [15]. While high-pressure die casting is the dominant process for
current magnesium automotive applications, wrought magnesium alloys
and their manufacturing processes are receiving increasing attention from
academia and industry. As magnesium is expanding from interior components to more critical applications in powertrain, chassis and body areas,
there is a greater need for developing wrought magnesium products including tubes and sheet and their forming processes to provide improved
mechanical and physical properties, crash performance and corrosion
resistance.
238
WPNL2204
Aluminum and magnesium alloy tubes
239
This chapter will provide a summary of recent development in bending
and hydroforming processes for lightweight aluminum and magnesium
tubes for automotive structural applications.
11.2
Aluminum and magnesium alloy tubes
Aluminum tubes are made from wrought alloys by (a) extruding a solid or
hollow billet; or (b) roll-forming and welding of a sheet metal strip. Tubes
are sometimes brought to final dimensions by drawing through dies for
better dimensional control. The properties of aluminum tubes depend on
the alloy composition, mechanical working and heat treatment involved in
tube processing. Magnesium tubes are currently available in the extruded
condition only. In this section, the physical metallurgy of wrought aluminum
and magnesium alloys, used in tube manufacturing for hydroforming applications, will be discussed.
11.2.1 Aluminum extruded tubes
Extrusion is a thermomechanical process that converts a cast billet (in solid
state) into a continuous length of generally uniform section by forcing it to
flow through a die which is shaped to produce the required form of product
[17]. In a modern extrusion plant, cast cylindrical billets are loaded into a
cylindrical container and pushed through a profiled die with a hydraulically
actuated ram. The most common dies used today are called ‘bridge dies’ or
‘port hole dies’ where the billet is broken into about four streams that get
their internal shape as the streams flow over the supporting mandrel whose
end face sits essentially flush with the die exit face that contains the external
profile of the extrusion. The material seams become longitudinally welded
together as the material finally exits the die over the very short ‘bearing
area’, which is a longitudinal section of the die exit where the internal
supporting mandrel and die exit plate finally set the total tube dimensions.
Seamless aluminum tubes can be extruded using a so-called ‘die and
mandrel’ process in which the mandrel becomes a component of the
extrusion press rather than a part of the die. In Table 11.1, several extrusion
alloys that are of interest in automotive applications are listed, and in Table
11.2 the mechanical properties of some of these alloys are summarized.
6xxx Extrusion alloys
The Al–Mg–Si based 6xxx alloys are by far the most widely used alloys
in extruded products. The 6xxx alloys have excellent general corrosion
resistance and provide good strength at low cost. Alloy 6063 has the lowest
strength of this group, but is easiest to extrude (see Fig. 11.1) and is
WPNL2204
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0.6–0.9
0.5–0.9
0.4–0.8
0.2–0.6
0.4
0.25
Si
0.35
0.35
0.7
0.35
0.5
0.4
Fe
0.3
0.3
0.15–0.4
0.1
0.1
0.1
Cu
* Maximum limit, unless a range is shown.
6005
6005A
6061
6063
5086
5454
Alloy
0.5
0.5
0.15
0.1
0.2–0.7
0.5–1
Mn
0.4–0.7
0.4–0.7
0.8–1.2
0.45–0.9
3.5–4.5
2.4–3
Mg
0.3
0.3
0.04–0.35
0.1
0.05–0.25
0.05–0.2
Cr
0.2
0.2
0.25
0.1
0.25
0.25
Zn
0.1
0.1
0.15
0.1
0.15
0.2
Ti
0.05
0.05
0.05
0.05
0.05
0.05
Each
Others
0.15
0.15
0.15
0.15
0.15
0.15
Total
Table 11.1 Chemical composition limits (in wt. %) of aluminum extrusion alloys* [16, 17] Maximum limit, unless a range is shown
Aluminum and magnesium alloy tubes
241
Table 11.2 Typical mechanical properties of aluminum extrusion alloys [16, 17]
Ultimate
tensile
strength
(MPa)
Tensile
yield
strength
(MPa)
Elongation
in 50 mm
(%)
Ultimate
shear
strength
(MPa)
Modulus
of
elasticity
(GPa)
6005-T5
6005A-T5
6061-T1
6061-T4
6061-T6
6063-T4
6063-T5
6063-T6
5086-O
5086-H112
5086-H116
5454-O
5454-H111
5454-H112
305
305
150
240
310
170
185
240
260
270
315
250
260
250
270
270
90
145
275
90
145
215
115
130
230
115
180
125
12
12
20
22
12
22
12
12
22
14
16
22
14
18
200
200
100
165
205
105
115
150
160
–
–
160
160
160
69
69
69
69
69
69
69
69
71
71
71
70
70
70
Extrudability rate (%) relative to Alloy 6063
Alloy and
temper
100
90
80
70
60
50
40
30
20
10
0
6063
6061
5086
7075
Alloy
11.1 Relative extrudability of aluminum alloys [16].
therefore used most often for very complex cross-sections. Alloys 6005,
6005A, and 6061 have the highest strengths.
The formation of magnesium silicide (Mg2Si) makes the 6xxx alloys
heat treatable and capable of achieving medium strength in the T6 condition. The proportion of Mg to Si for equilibrium Mg2Si phase formation is
WPNL2204
242
Hydroforming for advanced manufacturing
1.73 : 1. In a balanced alloy such as 6061, the alloy can be considered as a
pseudo-binary Al–Mg2Si system. The precipitation of Mg2Si particles after
extrusion offers considerable strength to the alloy, and 6061 is commonly
used for medium to heavy sections for heavy-duty applications. Alloy 6063
has lower Mg and Si content and, therefore, offers lower strength. In other
alloys such as 6005 and 6005A, an excess of Si over that required for Mg2Si
formation will further increase the strength but reduce the extrudability.
Minor additions of Mn and Cr can improve toughness of the alloys by
forming fine, stable incoherent dispersoids during homogenization. However,
these additions increase quench sensitivity by providing nucleation sites for
Mg2Si and decrease extrudability. Mn is preferred over Cr because it is far
less detrimental to extrusion speed and surface quality, although it is more
quench sensitive [17].
One unique phenomenon with age-hardening alloys such as the 6xxx
series is their softening behavior upon a retrogression heat treatment which
is a short time exposure (1–60 s) at 200–500 °C. Such a heat treatment has
been applied to 6061-T6 extruded tubes for compression-fit joining [18] and
6111-T4 sheet metal for enhanced workability during hemming and trimming operations [19]. Therefore, retrogression heat treatment presents an
option for enhanced hydroformability of 6xxx extruded tubes.
5xxx Extrusion alloys
The Al–Mg–Mn based 5xxx series alloys represent the most important
commercial non-heat-treatable alloys. The 5xxx alloys provide moderateto-high strengths, excellent welding characteristics, and good corrosion
resistance in marine environments. The 5xxx alloys have extensive use in
aerospace, shipbuilding and forged components industries. Extruded tubes
made of 5086 (containing about 4% Mg) and 5454 (containing about 2.7%
Mg) are often used in marine and aerospace applications. Alloys containing
over 4.5% Mg are rarely used due to their lower formability and susceptible
intergranular and stress corrosion. In addition, the hot-working range of
5xxx alloys is limited due to their strong work-hardening characteristic
which makes the alloys prone to incipient melting at higher extrusion temperatures. Thus, compared with 6xxx alloys, 5xxx alloy tubes have much
lower productivity due their lower extrusion speeds (70–80% lower than
6063 as shown in Fig. 11.1).
The work-hardening effect of 5xxx alloys is primarily due to the large
atomic size difference of Mg in Al (the Mg atom is 12% larger than the Al
atom). Various alloying elements including Mn, Cr, Fe and Si are added to
control the grain structure. It should be noted that reducing grain size can
significantly increase the strength of 5xxx alloy products in the O temper,
while grain size is not a major factor in obtaining high strength in other
WPNL2204
Aluminum and magnesium alloy tubes
243
aluminum alloys [17]. Nevertheless, grain size is controlled in all aluminum
alloys to optimize the formability. Honda has recently developed an Al–Mg
based extrusion alloy with very carefully controlled Cr, Mn and Fe levels
that restrict grain growth and property degradation during subsequent hot
gas forming [2].
11.2.2 Aluminum seam-welded tubes
The production of seam-welded aluminum tubes involves sheet production,
coil slitting, and tube mill processing. Current aluminum sheet processing
consists of direct-chill (DC) casting, and various hot- and cold-rolling
operations. In processing aluminum fin stock sheet, it has been reported
that continuous casting (CC) can substantially lower the operating cost of
aluminum sheet by reducing hot-rolling steps and in-process scrap. An
additional saving may be realized if the final gauge can be produced directly
off the continuous casting and continuous hot-roll line. The reduction in
rolling passes reduces cost but limits the control of the alloy chemistry and
grain size, which can reduce formability and other mechanical properties.
Recently, it has been demonstrated that low-cost tubes made of continuous
cast aluminum alloy 5754 sheet provided acceptable formability upon
bending and hydroforming, as well as good mechanical properties after
hydroforming [20, 21].
The tube mill process, shown in Fig. 11.2 [1], starts with a sheet aluminum
coil slit to the size of the outer diameter of the tube. The slit coil is fed to
the tube mill which consists of several pairs of roll-form dies, a welding
station, a cooling area and a sizing area. In the first step, the aluminum strip
2. Calibration
1. Forming
Forming
Coil edge rolls
clearing
Knife
rolls
Outer
definning Cooling
stage
Squeeze rolls
3. Straightness
Sizing
rolls
4. Quality
control
Saw
“Turks heads”
Sorting
area
Inner definning station
11.2 A typical tube mill for producing seam-welded aluminum tubes
(courtesy of Hydro Aluminum [1]).
WPNL2204
244
Hydroforming for advanced manufacturing
is formed into an open round tube that passes into a welding station where
a high frequency induction coil rapidly heats the edges of the strip almost
to its melting temperature. These edges are then squeezed together under
high pressure to form the tube, which then passes through cooling and sizing
rolls before final cutting and annealing.
5xxx Sheet alloys
5182, 5454 and 5754 are the principal Al–Mg series (5xxx) automotive sheet
alloys. There is also a German-registered alloy, AlMg3.5Mn, extensively
used in Europe for both sheet and tubular applications. Since these alloys
are not precipitation-hardenable, they do not strengthen during the paint
bake cycle. Since any strengthening developed due to work hardening
during forming may be lost during the paint bake cycle, their annealed (O-)
temper properties are generally used as a conservative estimate of their
design strength for most situations [6]. The 5xxx alloys are exceptionally
formable and have very high resistance to corrosion.
6xxx Sheet alloys
6009, 6022 and 6111 are Al–Mg–Si or 6xxx sheet alloys. They are heattreatable, and are typically supplied in the T4 temper. Subsequent aging
during the typical automotive paint bake cycle will cause some strengthening due to precipitation hardening of the Mg2Si phase. Of the three alloys,
6111 has highest strength and superior stretch-forming characteristics, while
6009 has slightly greater formability in bending [22].
Tables 11.3 and 11.4 summarize the automotive sheet alloys and their
mechanical properties, respectively. Both 5xxx and 6xxx alloys can be used
to make tubes for hydroforming, and Hydro Aluminum is presently the only
supplier of such tubes.
11.2.3 Magnesium extruded tubes
Extrusion alloys
Table 11.5 lists the nominal composition and typical room-temperature
tensile properties of extruded magnesium alloy tubes [23–25]. Of the
commercial extrusion alloys, AZ31 is most widely used in non-automotive
applications. With higher aluminum contents, AZ61 and AZ80 offer higher
strength than AZ31 alloy, but have much lower extrudability. The highstrength Zr-containing ZK60 was designed for applications in racing
cars and bicycles, such as wheels and stems [24]. The extrusion speed of
ZK60A tubes is extremely low, rendering it uneconomical for automotive
applications.
WPNL2204
WPNL2204
0.20
0.25
0.40
0.6–1.0
0.8–1.5
0.6–1.1
Si
0.35
0.40
0.40
0.50
0.05–0.20
0.40
Fe
0.15
0.10
0.10
0.15–0.6
0.01–0.11
0.50–0.9
Cu
* Maximum limit, unless a range is shown.
5182
5454
5754
6009
6022
6111
Alloy
0.20–0.50
0.50–1.0
0.50
0.20–0.8
0.02–0.10
0.10–0.45
Mn
4.0–5.0
2.4–3.0
2.6–3.6
0.40–0.8
0.45–0.7
0.50–1.0
Mg
Table 11.3 Chemical composition limits (in wt. %) of aluminum sheet alloys* [22]
0.10
0.05–0.20
0.30
0.10
0.10
0.10
Cr
0.25
0.25
0.20
0.25
0.25
0.25
Zn
0.10
0.20
0.15
0.10
0.25
0.15
Ti
0.05
0.05
0.05
0.05
0.05
0.05
Each
Others
0.15
0.15
0.15
0.15
0.15
0.15
Total
246
Hydroforming for advanced manufacturing
Table 11.4 Typical mechanical properties of aluminum extrusion alloys [22]
Alloy
and
temper
Ultimate
tensile
strength
(MPa)
Tensile
yield
strength
(MPa)
Elongation
in 50 mm
(%)
Ultimate
shear
strength
(MPa)
Modulus
of
elasticity
(GPa)
5182-O
5454-O
5754-O
6009-T4
6009-T62
6022-T4
6022-T62
6111-T4
6111-T41
6111-T62
275
250
220
220
300
255
325
280
270
360
130
115
100
125
260
150
290
150
150
320
24
22
26
25
11
26
12
26
26
11
205
160
130
130
180
155
195
175
160
215
71
70
71
69
69
69
69
69
69
69
Table 11.5 Nominal compositions and typical room-temperature tensile
properties of extruded magnesium alloy tubes [22–24]
Alloy
AZ31
AZ61
AZ80
ZK60
ZK60
AM30
Temper
F
F
T5
F
T5
F
Composition (wt. %)
Tensile properties
Al
Zn
Mn
Zr
Yield
strength,
MPa
Tensile
strength,
MPa
Elongation,
%
3.0
6.5
8.0
–
–
3.0
1.0
1.0
0.6
5.5
5.5
–
0.20
0.15
0.30
–
–
0.40
–
–
–
0.45
0.45
–
165
165
275
240
268
171
245
280
380
325
330
232
12
14
7
13
12
12
AM30 (Mg–3% Al–0.4% Mn), a new extrusion alloy developed by GM
[25], is aimed to provide a good balance of strength, ductility, extrudability
and corrosion resistance. The alloy was designed based on the following:
• Aluminum has the most favorable effect on magnesium of any of the
alloying elements [23]. It improves strength, hardness and corrosion
resistance, but reduces ductility. An aluminum content of about 5–6%
yields the optimum combination of strength and ductility for structural
applications. Increasing aluminum content widens the freezing range
and makes the alloy easier to cast, but more difficult to extrude due to
increased hardness. For example, alloys containing less than 3% Al can
be extruded at higher extrusion speeds compared with high-strength
WPNL2204
Aluminum and magnesium alloy tubes
247
alloys such as AZ61 (Mg–6% Al–1% Zn) and ZK60 (Mg–6% Zn–0.5%
Zr) [26]. To maximize the ductility and extrudability, while maintaining
reasonable strength and castability (for billet casting before extrusion),
an aluminum content of 3% was selected for the new alloy.
• Zinc is next to aluminum in effectiveness as an alloying ingredient to
strengthen magnesium [23]. However, it reduces ductility and increases
hot-shortness of Mg–Al based alloys. Zinc-containing magnesium alloys
are prone to microporosity [27]. Zinc is also reported to have mild to
moderate accelerating effects on corrosion rates of magnesium as determined by alternate immersion in 3% NaCl solution [23]. Therefore,
unlike most commercial magnesium alloys, Zn was not selected in this
alloy.
• Manganese does not have much effect on tensile strength, but it does
slightly increase the yield strength of magnesium alloys. Its most important function is to improve the corrosion resistance of Mg–Al based
alloys by removing iron and other heavy-metal elements into relatively
harmless intermetallic compounds, some of which separate out during
melting. For this purpose, Mn is added at about 0.4% as recommended
by the ASTM Specification B93-94a.
Depending on the extrusion temperature used, the AM30 alloy developed was shown to extrude 20–30% faster without visible defects compared
with the conventional AZ31 alloy, [25]. Figure 11.3 summarizes the relative
Extrudability rate (%) relative to Alloy 6063
100
90
80
70
60
50
40
30
20
10
0
6063
AM30
AZ31B
Alloy
AZ61
AZ80/ZK60
11.3 Relative extrudability of magnesium alloys compared with alloy
6063.
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extrudability of magnesium alloys compared with alloy 6063, based on
information from the literature [23–25] and the authors’ experience.
Extrusion processes
Magnesium alloys can be warm or hot extruded in hydraulic presses to form
bars, tubes, and a wide variety of profiles [23]. While hollow magnesium
extrusions can be made with a mandrel and a drilled or pierced billet, it is
generally preferable to use a bridge die as detailed earlier.
The hydrostatic extrusion process, typically used for copper tubing fabrication, is a much faster extrusion process compared with the conventional
direct extrusion. It was recently reported that seamless magnesium tubes
were extruded using the hydrostatic process at speeds up to 100 m min−1,
due to the absence of friction between the billet and container since the
billet is suspended in hydraulic oil [28]. Although the process is capable of
extrusion ratios up to 700, the outer diameter of tubes produced by this
process is limited to about 45 mm, even with a large 4000-ton press [28].
11.3
Aluminum tube bending and hydroforming
11.3.1 Tube bending
The rotary draw bending process as shown in Fig. 11.4 is generally used for
bending aluminum tubes. In this process, the tube is clamped against the
bend die insert with the clamp die, and a multi-ball steel mandrel is positioned inside the tube. The pressure die holds the collet end of the tube
against the wiper die as the bend die rotates and the tube is drawn
forward.
Friction needs to be better controlled in bending aluminum tubes as
compared with the steel tube bending process. This can be achieved by the
following: (1) a slightly larger tolerance between the tube inner diameter
and the mandrel diameter; and (2) a dry-film lubricant is recommended for
the aluminum tube outer surface in contact with the bend tooling [29].
Figure 11.5 shows an aluminum alloy 6061-T4 tube bent to 1.5D
compared with a steel 1008 tube of the same outer diameter (70 mm) and
centerline bend radius. The D ratio of the centerline bend radius (CLR) to
the outer diameter (OD) of the tube describes the severity of the bend and
determines the maximum strain developed during bending. 2D is generally
required for tubular structural components in automotive applications, and
1.5D is sometimes required in other applications. Figure 11.6 compares the
major strain distribution in the bent steel and aluminum tubes for a 1.5D
bend condition. A grid pattern of 2.5-mm-diameter circles was electrochemically applied to the tubes and the deformed grids were measured
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249
Pressure die
boost cylinder
Pressure die back-up
carriage
Carriage bearing
Pressure die
Clamp die
Multi-ball
mandrel
Cleat
Wiper die
Bend die
Cleated clamp plug
11.4 Schematic of the tooling used for bending.
steel
Aluminum
11.5 Aluminum tube 6061-T4 bent to 2D compared with a steel 1008
tube.
using a CamSys Grid Pattern Analyzer (GPA)-100 strain measurement
system. Figure 11.6 shows that both tubes achieved the 90° bend for a 1.5D
condition with very good quality bends as evidenced by the relatively
smooth strain distribution. The steel tube shows a slightly higher maximum
strain (∼46%) compared with the 42% strain observed in the aluminum
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50
45
Steel
Aluminum
True strain (%)
40
35
30
25
20
15
10
5
0
0
10
20
30
40
50
Bend angle (°)
60
70
80
90
11.6 Strain distribution along the tension side of the bend in 1008
steel and 6061-T4 aluminum tubes (1.5D/90° bend).
40
30°
60°
90°
35
True strain (%)
30
25
20
15
10
5
0
0
10
20
30
40
50
60
Bending angle (°)
70
80
90
100
11.7 Strain evolution during bending of 6061-T4 tubes (2D bend).
tube. These results show that the bendability of a tubular material does not
only depend on the level of maximum strain it can withstand, but also its
ability to spread the maximum strain over a large area along the bend. In
this regard, the aluminum tube shows similar behavior in spreading the
bending strain longitudinally over the tube as the steel tubes. An uneven
strain distribution in bent tubes can lead to tube splitting and/or wrinkles
[29].
The bending process creates a tensile stress zone on the outside radius
and a compressive stress zone on the inside radius. Figure 11.7 shows the
strain evolution on the tension side, plotted longitudinally along the 2D
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40
Major
True strain (%)
30
Minor
20
10
Tension
0°
0
Minor
strain
–10
–20
180°
0
20
40
60
80 100 120 140 160 180
Location (°)
Compression
11.8 Circumferential strain distribution in a 2D/90° bent 6061-T4 tube.
bend of 6061-T4 tubes. The strain builds up gradually to about 22% at a
bend of 30° as shown in Fig. 11.7, and a maximum strain of about 32% is
reached at a bend angle of about 60°. Further bending increases the
maximum strain only slightly (to about 35%), but widens the strained area
along the bend. Extruded 6063 and seam-welded 5754 tubes also showed
similar results in bending strain distribution.
Figure 11.8 shows the strain distribution along the circumference of a 45°
radial section of a 6061-T4 tube bent to the 2D/90° condition (the longitudinal major strain distribution of this tube is shown in Fig. 11.7). It is evident
that the major strain starts from its maximum tension (37%) at the maximum
thinning point to maximum compression (−14%) accompanied by maximum
thickening. It is also interesting to note that the tension-to-compression
transition did not occur at 90°, but rather at about 125°, indicating that there
is an overall tension force pulling the tube during the rotary draw bending
process. On the other hand, the minor strain is very low (slightly compressive) across the circumference of the tube, confirming that the tube is
approximately in the plane-strain condition during bending.
11.3.2 Hydroforming
Hydroforming of pre-bent aluminum tubes can be carried out at different
pressure levels depending on the forming strain required for the product.
Fig. 11.9 shows a 6061-T4 tube formed on a 5000-ton press at a forming
pressure of 62 MPa. The round cross-section (70 mm outer diameter and
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11.9 A Hydroformed 6061-T4 tube (forming pressure of 62 MPa).
40
Major
Minor
True strain (%)
30
Corner B
20
Corner A
10
Tension 0°
A
0
Minor
strain
–10
B
–20
0
20
40
60
80
100
120
140
160
180
Compression
180°
Location (°)
11.10 Circumferential strain distribution in a 6061-T4 tube after
bending (2D/90°) and hydroforming (forming pressure of 62 MPa).
3.5 mm wall) of the tube was converted to a 73 × 49 mm rectangular section.
The average corner radius of this hydroformed part was measured as
14.5 mm.
The strain distribution along the circumference of 6061-T4 tubes in the
90° bend area was also measured after hydroforming. Figure 11.10 shows
the major and minor strain distribution along the circumference sectioned
at ∼45° in the bend. While the major strain does not change significantly
after hydroforming at 62 MPa, except for the slight increase in the corners
(A and B), the overall minor strain increases to ∼+15% due to the circumferential expansion during hydroforming at a pressure of 62 MPa. A closer
examination of Fig. 11.10 indicates that the major strain increase near
Corner A (tension side of the bend) is greater than Corner B (compression
side of the bend). This suggests that any further increases in forming
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pressure may cause undesirable localized thinning in the aluminum tubes.
Forming failure is normally observed near the tension corner at the tangent
area.
Free expansion burst tests are used to evaluate the formability of the
tubes. In a typical burst test, a fixed length of tube is sealed at both ends
and internally pressurized with a water-based solution. The pressure vs. time
is recorded with a computer data acquisition system and strains are
measured with a circle grid pre-etched on the tube surface. The burst test
resembles the hydroforming process in expanding the tube with an internal
pressure, but differs from hydroforming in that the tube is not placed inside
a die that restricts the forming to a desired shape.
Figure 11.11 shows the forming limit diagrams (FLD) obtained from the
burst tests for 6063-T4 (extruded tube), 5754-DC (seam-welded tube made
of direct-chill cast and rolled sheet) and 5754-CC (seam-welded tube
made of continuous cast and rolled sheet) tubes, respectively. The three
materials exhibit slightly different forming limit (necking) strains for the
plane strain condition (FLC0), 24% for 6063-T4, and 25% for both 5754-DC
and 5754-CC tubes. The seam-welded tubes made from two differently
processed sheet materials (5754-DC and 5754-CC) have essentially the
same formability in tube free expansion, indicating similar formability in
hydroforming.
11.3.3 Retrogression heat treatment for
improved formability
Retrogression heat treatment using induction heating and fast waterquenching has been shown to improve the formability and hydroforming
performance of extruded 6063-T4 [30] and 6063-T6 aluminum tubes [31].
For 6063-T4 alloy tubes, heating for 5 s at 350–400 °C was found to produce
50% softening, which allowed larger strains during burst testing and sharper
corner radii during hydroforming, indicating a formability improvement.
At room temperature, the hardness of the retrogression-treated material
increased by 30% in 8 h and reached the original 6063-T4 level in about
215 h [30].
For 6063-T6 tubes, significant softening (60–70%) was possible with a
retrogression heat treatment for 2 s at 450 °C, which improved the burst
strains to the same level as the 6063-T4 material, suggesting equivalent
formability. At room temperature, the retrogression-treated material
subsequently aged back to T4 hardness in 9 days, and with time, continued
to age above the T4 hardness, but not to the original T6 level, even after
6000 h. Additional heat treatment is required to achieve the original T6
hardness [31].
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45
40
Major strain (%)
35
30
25
20
15
Safe
Necked
10
5
0
–12 –9
–6
–3
0
3
6
9
12
Major strain (%)
Minor strain, %
a
50
45
40
35
30
25
20
15
10
5
0
–12 –9
Safe
Necked
–6
–3 0
3
Minor strain, %
6
9
12
b
45
40
35
Major strain (%)
254
30
25
20
15
10
Safe
Necked
5
0
–12 –9
–6
–3 0
3
6
Minor strain (%)
c
9
12
11.11 Experimental forming limit diagrams (FLD) for a 6063-T7,
b 5754-DC; and c 5754-CC tubes.
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11.4
255
Magnesium tube bending
11.4.1 Bending at room temperature
Bending of conventionally extruded magnesium AZ31 and AM30 tubes at
room temperature is generally unsuccessful; a 2D/90° bend could not be
consistently achieved [32]. Figure 11.12 compares the longitudinal cross
section micrographs of an AZ31 tube before and after a 30° bend at room
500um
a
100um
b
11.12 Optical micrographs showing the microstructure of an AZ31
tube: a before bending; and b after bending at room temperature
(fractured at 30°) [32].
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temperature. The micrograph in Fig. 11.12b clearly shows the onset of strain
localization initiating a crack before fracture. There is no evidence of any
grain elongation even near the outermost surface where the strain is the
largest, suggesting that twinning is the predominant deformation mode at
room temperature. One may notice that the grain sizes in the outer layer
on either side of the localized band are different.
11.4.2 Bending at elevated temperatures
A moderate temperature (100–200 °C) bending process was developed at
GM for magnesium tubes, using a Pines rotary draw bending machine with
heated tooling [32]. In this bending process, three parts of the bend tooling:
the bend die, the pressure die, and the mandrel, were heated, and the
temperature for each heating zone was controlled separately. Using the
optimal parameters developed for the moderate temperature bending
process [32], a 2D/90° bend, as shown in Fig. 11.13, can be consistently
achieved with AZ31 and AM30 tubes at about 150 °C. No surface cracks
were detected in these bent tubes.
In the rotary draw bending process, as described in Section 11.3.1, the
tension (outer) side of the tube is thinned while the compression (inner)
side is thickened. Figure 11.14 shows the thinning distribution along the
tension side of magnesium alloy tubes bent to the 2D/90° condition at
150 °C. The smooth thinning distribution in Fig. 11.14 suggests uniform
deformation during bending at elevated temperatures. The improved
bendability of the magnesium alloy tubes is due to the higher ductility and
formability reported in the literature [34, 35].
AM30
AZ31
11.13 AM30 and AZ31 tubes bent (2D/90°) at 150 °C.
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257
25%
Thinning (%)
20%
15%
10%
AM30
AZ31
5%
0%
0
20
40
Bend angle (°)
60
80
11.14 Thinning distribution in magnesium tubes bent at 150 °C.
11.5
Forming at elevated temperatures
Elevated temperatures increase the formability of certain aluminum
alloys (especially the 5xxx alloys) and magnesium alloys. More complex
geometries and greater circumferential expansion can be achieved at
elevated temperatures using pressurized fluids such as oil or gas.
11.5.1 Gas forming of aluminum tubes
A hot gas forming process was developed by Honda to form Al–Mg alloy
tubes with up to 90% circumferential expansion at a fairly high temperature
(510 °C) but a very low forming pressure (about 1% of the hydroforming
pressure) [2]. Recently, a warm gas forming process at a much lower temperature range (150–350 °C) was developed using an Interlaken 5000-kN
press (Fig. 11.15) [36]. This press is equipped with two separate controllers
and pressure intensifiers to perform the tests with either gas or water. For
warm forming, nitrogen is pre-charged into the system with a pre-charge
control valve, and forming under gas pressure or volume control is achieved
with a pressure intensifier. Fast forming cycles of about 10 s can be obtained
with either operating modes. Two end-feed actuators are used to seal and
provide axial feeding during forming. The forming-die holder is heated with
electric resistance elements and a temperature controller is used to measure
and control the temperatures independently in the six zones of the die. The
die cavity has two rectangular zones, a large and a small zone, and die inserts
are used to create different geometries required for tube forming. Complete
expansion involves an average tube expansion of 25 and 50% in the smaller
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11.15 Warm gas forming press at CANMET, Ottawa, Canada [36].
and larger cavity regions, respectively. Figure 11.16b shows aluminum alloy
5754 seam-welded tubes gas formed at 250 °C.
11.5.2 Gas forming of magnesium tubes
The warm gas forming process has been applied to magnesium alloy tubes.
Figure 11.17 compares an AZ31 magnesium alloy tube hydroformed at
room temperature with a similar tube gas formed at 250 °C. While both
tubes failed at an extrusion seam, gas forming could provide a circumferential expansion of 28% compared with only 8% with hydroforming at
room temperature. Further optimization of extrusion and warm forming
process would improve the formability of the magnesium tubes especially
if controlled end feeding can be accomplished. Alternatively, seamless
extrusions can be considered.
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259
a
b
11.16 Aluminum alloy 5754 tubes gas-formed at 250 °C:
a hydroforming at room temperature (8% circumference expansion);
b Warm gas-formed at 250 °C (28% circumference expansion).
a
b
11.17 a Room temperature hydroforming; and b warm gas forming of
magnesium alloy AZ31 tubes: a Lower die; b AZ31 tube formed at
350 °C (80% circumference expansion).
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Seamless extrusions of AZ31 magnesium alloy were the subject of another
study of warm hydroforming (gas forming using nitrogen) [37]. The tooling
for the warm forming process consists of two units, the clamping and the
forming dies. The clamping unit ensures the sealing of the tube and the
supply of the pressure medium. The tube is put into the forming unit, consisting of a main body and a free expansion guided zone equipped with
cooling and heating devices. The guided zone makes it possible that tubes
of different outer diameters can be formed under plane strain condition.
Band heaters are used to heat the forming unit and the sealing punches.
11.6
Automotive applications and future trends
Hydroformed parts generally replace stamped components, achieving
weight savings and part consolidation. In many cases, one hydroformed part
can replace several hat-shaped box sections (stampings) that are welded
together. Some of the advantages and disadvantages of hydroforming
compared with stamping are listed below.
Part consolidation
This is the single most important driver for replacing welded box sections
of steel stampings with hydroformed tubular structures. Compared with a
stamped design, the hydroformed engine cradle for Buick LeSabre reduced
the total part count from 40 to 18 [38]. It was also reported [39] that a 10piece hydroformed radiator support assembly for Dodge Ram replaced a
17-piece stamped assembly.
Mass reduction and improved structural performance
Mass reduction of a hydroformed design is achieved by eliminating the
welding flanges, more efficient section design, and reduced wall thickness
because of improved integrity in a tubular structure. Massing savings of
15–20% are generally reported when converting a stamped and welded
structure to a hydroformed application. Despite reduced mass, a hydroformed tubular structure offers improved stiffness and strength due to the
elimination of massive welds and the use of tailored section design. The
GMT800 (Chevrolet Silverado and GMC Sierra) truck frame, introduced
in 1999, was the first high-volume application of hydroformed front rails
and cross-members. The hydroformed frame doubled the torsional stiffness
compared with a conventional stamped truck frame while achieving a 15%
mass savings and better ride quality. The hydroformed frames are further
improved in GM’s newly introduced Silverado and GMC Sierra trucks
(GMT900), which have an annual production run of 1.3 million units. The
Chevrolet SSR, a low-volume sport truck, features full-length hydroformed
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261
rails. A traditional stamped frame with equivalent strength and rigidity
would weigh roughly 20% more than the hydroformed structure. This
improved strength allows the frame to take on the road inputs and allows
the suspension do its job more precisely, resulting in better ride quality.
Additional mass savings can be achieved when the hydroformed steel
structures are converted to aluminum or magnesium hydroforms. Depending on the design criteria (strength or stiffness), aluminum and magnesium
tubes can save about 20–30% and 40–50% of mass, respectively, compared
with steel tubular structures. For example, the 4.8-m long aluminum frame
rail for the 2006 Corvette, the largest hydroformed part in the world, saved
20% mass from the steel rail it replaced.
Lower tooling costs and fewer processes steps
The hydroforming process uses pressurized fluid as the mating die and
relatively simple dies and fixtures in the forming process, hence, substantially reducing the number of tools and thus tooling cost. In a typical
stamped and welded box section assembly, each stamping involves blanking,
drawing, trimming, flanging and piercing operations, and those stampings
have to be welded together. A hydroforming solution reduces the process
steps because no welding of sections is needed and hole can be punched
during hydroforming.
Higher dimensional stability
There is significantly less springback in hydroformed parts due to the higher
applied pressure compared with a stamping operation. The part consolidation (less welding) and the incorporation of hole-piercing into the hydroforming operation improve the dimensional tolerance and thus reduce
scrap.
Disadvantages
Despite the above advantages, the hydroforming process also has some
drawbacks such as longer cycle times and more expensive capital equipment, which are often offset by part consolidation and lower tooling costs.
The need for both high pressure (for die holding) and long stroke (for
die closing/opening) in hydroforming makes the hydraulic system very
expensive.
11.6.1 Aluminum applications
Hydroformed aluminum tubular components provide very efficient design
solutions for automotive chassis structures. The BMW 7 series is a com-
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11.18 Hydroformed aluminum side rail for 2006 Chevrolet Corvette
Z06.
bination of hydroformed tubes and vacuum die cast members, with a total
mass of only 14.1 kg [40]. The newly developed aluminum front and rear
cradles by Honda for the 2005 Acura RL used warm gas formed tubular
members.
In North America, the most recent applications of hydroformed aluminum tubes are the side rail (Fig. 11.18) and roof bow for the 2006 Chevrolet
Corvette Z06 frame. The use of this lightweight aluminum frame (129.5 kg)
and a magnesium engine cradle (10.3 kg), combined with a powerful LS7
7.0L/500HP V8 engine for Corvette Z06, set a remarkable performance
standard: 0 to 60 mph in less than 4 s and a top speed of more than
190 mph.
Unlike steel which cannot be extruded to complex hollow sections, aluminum alloys can be readily extruded to form complex tubular sections
which can be bent and/or hydroformed. Due to the tooling and process
cost, hydroforming is to be avoided in the design if no geometric section
changes are required along the tube length. This limits the application of
hydroformed aluminum components compared with steel hydroforms.
Nonetheless, hydroformed aluminum tubes are increasingly used to provide
additional mass savings and/or design efficiency and flexibility in automotive and other industries.
11.6.2 Magnesium applications
Hydroforming of magnesium alloy tubes is still at the research and development stage, and there have been no reported commercial applications
of hydroformed magnesium tubes. In Table 11.6 the potential applications
of magnesium extrusions in the automotive interior, and body and
chassis areas are summarized, some of which may involve hydroforming
processes.
Magnesium extrusions are used to make prototype parts such as bumper
beams and most parts of a spaceframe for VW 1-Liter Car [10]. However,
the production of magnesium tubes/extrusions in automotive structures
would require more development to meet all performance and cost targets
as well as a supply base for high-volume automotive production. Recently,
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Table 11.6 Potential automotive application of magnesium extrusions
System
Component
Interior
Instrument panel
Seat components
HVAC (Heating, ventilation and air-conditioning) components
Body
Roof frame
Bumper beam
Radiator support
Shotgun
Frame rail
Chassis
Engine cradle
Subframe
a US–Canada–China collaborative effort has been focused on the development of enabling technologies for magnesium body applications using front
end body structure as a test bed. A magnesium front end would weigh about
30 kg, and provide up to 40 kg of mass saving compared with the equivalent
steel construction [41].
Bending and gas forming processes of magnesium tubes will be further
developed and validated in this project for automotive applications.
11.7
Acknowledgments
The authors would like to acknowledge many colleagues at General Motors,
especially Bob Kubic, Raj Mishra, Reggie Joyner, Susan Hartfield-Wunsch,
Mark Cline and Terry Kent, for their support and contributions. Collaboration with CANMET (Pierre Martin) is also acknowledged.
11.8
References
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Bending for Hydroforming’, unpublished work, General Motors Corporation,
2000.
30. hartfield-wunsch s e and luo a a, ‘ATIM Heat Treatment of Extruded 6063-T4
Aluminum Tubes for Hydroforming’, unpublished work, General Motors
Corporation, 2000.
31. hartfield-wunsch s e, coe t and luo a a, ‘ATIM Heat Treatment of Extruded
6063-T6 Tubes for Hydroforming’, unpublished work, General Motors Corporation, 2003.
32. luo a a, sachdev a k, mishra r k and kubic r c, ‘Bendability and Microstructure
of Magnesium Alloy Tubes at Room and Elevated Temperatures’, in Magnesium
Technology 2005, eds. Neelameggham N R, Kaplan H I, and Powell B R, TMS,
Warrendale, PA, 2005, 145–148.
33. luo a a and sachdev a k, ‘Development of a Moderate Temperature Bending
Process for Magnesium Alloy Extrusions’, presented at the International Conference on Magnesium, Beijing, China, Sept. 20–24, 2004, Materials Science Forum,
2005, 488–489, 477–482.
34. luo a a and sachdev a k, ‘Mechanical Properties and Microstructure of
AZ31 Magnesium Alloy Tubes’, in Magnesium Technology 2004, ed. Luo A A,
The Minerals, Metals and Materials Society (TMS), Warrendale, PA, 2004,
79–85.
35. krajewski p e, ‘Elevated Temperature Behavior of Sheet Magnesium Alloys’, in
Magnesium Technology 2002, ed. Kaplan H I, TMS, Warrendale, PA, 2002,
175–179.
36. martin p, baragar d, boyle k p, luo a a, jonas j j, godet s, neale k w,
‘Elevated Temperature Property Measurements for Warm Forming Aluminium
Alloy Tubes’, in Proceedings of the 2nd International Light Metals Technology
Conference 2005, Ed. Kaufmann H, June 8–10, 2005, St. Wolfgang, Austria.
51–56.
37. ben-artzy a, spinat e, dahan o, siegert k, jager s, mueller k, altan t, ‘High
Internal Pressure Forming of Magnesium Tubes’, in Magnesium Technology
2006, eds. Luo A A, Neelameggham N R and Beals R S, TMS, Warrendale, PA,
2006, 253–258.
38. shah s and bruggemann c, ‘Tube Hydroforming Process Overview and Applications’, presented at Innovations in Hydroforming Technology, Nashville, TN,
September 25, 1996.
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39. mason m, ‘Tube Hydroforming: Advancements Using Sequenced Forming
Pressure’, presented at Innovations in Hydroforming Technology, Nashville, TN,
September 25, 1996.
40. hirsch j, ‘Automotive Trends in Aluminum, The European Perspective: Part
Two’, knowledge article from www.Key-to-Metals.com, 2006.
41. luo a a and mccune r c, ‘Magnesium Front End Projects’, Automotive Lightweighting Materials, FY 2006 Progress Report, United States Department of
Energy, 2007.
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Low-pressure tube hydroforming
G. MORPHY,
Excella Technologies Inc., Canada
12.1
Introduction
There are many tubular parts being hydroformed for today’s automotive
structures that are usually categorized as high-pressure (HPH) and lowpressure hydroforming (LPH). The proportion made using the latter method
is surprisingly large considering that more companies use the HPH process
but this is the case as LPH can provide all of the advantages shown in
Figure 13.3.
The individual aspects of product design are interrelated, but are considered as separately as possible to improve understanding of LPH and
hydroforming in general, how part features work, as well as the design and
economic effect of their inclusion. When designing parts, it is best to work
with an experienced hydroformer knowledgeable about both processes to
help you make the right decisions for the right reasons to ensure the best
result. This chapter provides information to allow design flexibility to be
balanced with minimal weight and cost and gives the reasoning behind it.
In the range of metal-forming options available to make structural
parts there is a technique that is less expensive than hydroforming called
mechanical tubeforming (Figure 13.2). The shortcomings of mechanical
tubeforming led to the development of sequenced-pressure hydroforming
in the mid to late 1980s. This found market acceptance since it achieved
better dimensional stability and a broader range of part feature design
options at a cost that was higher, but acceptable. Other now commonly
recognized benefits of tube hydroforming (see Fig. 13.3) came to be recognized also. A fundamental aim was minimizing the cost increase compared
with mechanical tubeforming. The forming method developed uses much
smaller presses and lower internal forming pressure than that used in other
hydroforming processes. It was devised to use the material and tube quality
that was normal in the mid 1980s, since special specifications would increase
cost. Its greater expense relative to mechanical tubeforming is worthwhile
for those parts where it provides sufficient benefit. This can include many
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hydroformed parts currently made by other methods following the rationale explained in Chapter 13 that the process that delivers the features
needed should be employed.
A review of factors that are important for hydroformed parts and how
this affects the design of LPH parts will provide more insight into this less
understood hydroforming approach. It will show more clearly how the
benefits of LPH can be applied to design parts for the best balance of
essential part functions and cost. Low-pressure hydroforming (LPH),
sequenced-pressure hydroforming (SPH) and HydroDieForming (HDF)
are discussed. Each uses forming pressure in different ways, but applying it
while the hydroforming die is closing is a common trait.
12.2
Low-pressure hydroforming
The simplest hydroforming method uses a die cavity whose periphery equals
the starting blank. Cross-section shapes must be simple as shown in Fig. 12.1
and transitions between them gradual. Proper bend positioning and changing the cross-sectional shape around bends are particular problems. Internal
pressure is too low to yield the material after the die is closed.
Cross-sections that are too complex result in material being crushed at
the split line; this is known as pinching. These restrictions motivated process
development researchers to strive for more capable and economical processes and, therefore, this method is seldom used because of the greater
capabilities gained by later developments. For the rest of the chapter, LPH
will be used as a general term for SPH and HDF. Some explanations are
related to HPH since it is more widely understood.
12.2.1 Sequenced pressure hydroforming
The process description near the beginning of Chapter 13 mostly describes
this type of low-pressure hydroforming, but there are a few more points to
be added. Later sections in this chapter describe how this process relates
to material and tube properties and part design features.
Split line
12.1 Simple hydroformed cross-section.
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The two-step process described in Chapter 13 uses two pressure levels.
The first during the latter part of die closing (Fig. 12.2c) is well below the
pressure that would yield the material. This supports the material by forcing
it to form in an orderly manner, but die design has to be aligned with
forming behavior.
The second higher-pressure level that is applied after the die closes is not
used to form the corners. This is true for HPH and is a very persistent notion
assumed to apply to all hydroform processes. For this process, the crosssection corners are completely formed once the die is closed. Instead, the
increased pressure flattens indentations that may occur on flat surfaces and
provides support for hole punching.
SPH part design normally has a cross-section periphery equal to that of
the start tube blank. Parts also have the same peripheral length from one
end to the other, whereas the cross-section shape can change as needed.
12.2.2 HydroDieFormingTM
As with many technologies, development has continued in low-pressure
hydroforming. The HydroDieFormingTM method results from improvements and advancements that allow greater part and process design flexibility (more features can be formed with less effort and good results), process
stability and lower cost. The process is similar to Fig. 12.2 conceptually, but
differences in tools and processes used provide substantial benefits. The
exact process sequence is adjusted to suit the required part characteristics
and features.
Features or advancements that distinguish HDFTM from other LPH processes are:
1.
2.
3.
4.
5.
6.
7.
Parts can be expanded:
a. At the ends of the parts (i.e. 50%) as is common in HPH, but with
more options;
b. 30–40% (perhaps more) at the center of long parts; previously
unavailable.
Numerous strategies to expand local part cross-sections.
Wall thickness can be more consistent around a given cross-section,
thus avoiding the corner thinning pattern that is commonly associated
with expansion during hydroforming.
Wall thickness can be controlled in expanded areas by several
techniques to put thicker material where it is most beneficial; higher
rigidity, lower weight.
More opportunities to reduce weight and increase rigidity.
More control of material forming during the whole cycle.
More accurate tube location, process and dimensional control.
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Before pressure
die open
Before pressure
die closing
First pressure stage
die closing
Second pressure stage
die closd
12.2 Sequenced pressure hydroforming with two pressure levels.
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9.
10.
11.
12.
13.
14.
15.
16.
17.
271
A refined approach to apply the right pressure at different points in
the cycle coordinated with die movement to increase the number of
part feature options and decrease cost factors.
Fewer die and part design considerations.
Greatly reduced rupture risk by avoiding or reducing corner wall
thinning when expanding.
More processing options to suit different part requirements.
Improved process stability.
Lower cost and less material used, fewer processing steps and shorter
cycle time.
Lower tooling and equipment cost.
Suited for high-volume and lower-volume parts.
More options to eliminate or minimize end scrap.
Extensive combination of mechanical and fluid-forming process
elements.
It is common with conventionally hydroformed parts to have a continuously varying cross-sectional periphery along its length. Some expanded
sections provide a significant structural benefit and others do not. Distinguishing between these and choosing to design a constant periphery part,
except where expansion gives a substantial benefit is a way to expand cost
effectively. These are usually formed differently than normal HPH forming.
12.3
Part characteristics
12.3.1 Material type
Most ductile metal, including copper, brass, steel, titanium, aluminum and
many of their alloys can be hydroformed. Structural tube hydroforming
applications have most commonly used steel alloys such as mild, highstrength low-alloy (HSLA) and galvanneal and 409 stainless steel. Other
options that are gaining acceptance as the industry moves toward much
greater use of high-strength steels (HSS) are dual-phase (DP) and TRIP at
various strength levels.
LPH has the advantage of substantially lower material formability
requirements. This provides a much improved ability to form steel parts
with more complex part features than would otherwise be feasible. Less
formability is needed because the material is reshaped rather than expanded
by stretching. A second reason is that the reshaping is done by compressive
forces generated as the die closes on the fluid-filled tube.
Thus, LPH can be applied to the lower formability materials mentioned
above. Of most probable interest for automotive structural interest are DP,
TRIP steel grades and aluminum, which offer the potential to reduce weight.
The ability of LPH to readily form these materials makes hydroforming a
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viable option. Any past or current applications of this sort should be reassessed from the LPH perspective, since it can change a previously unfeasible
solution into a reasonable one. Several grades of each of these materials
have been formed using LPH with prototype and production tools. Using
them in production is not viewed as being a problem.
Material type should be determined early in the design process to
best suit the function of the part. The choice may be based on economy,
weight reduction, part count reduction, corrosion resistance, load carrying
ability, energy absorption during crash (to name some important factors)
or in any combination. The first condition often creates a preference for
mild steel (SAE 1008/1010 steel). Often in the past, the largest consideration when selecting material has been meeting the forming requirements of the conventional hydroforming process. Such compromises do not
need to be made with LPH and part function-related requirements can be
focused on.
During part development and prototype testing, component performance can be less than anticipated or requirements that dictate using
higher strength steel, a thicker wall, or cross-section redesign, for example,
may change. Industry-wide efforts to reduce part and therefore vehicle
weight have led to increasing use of high-strength steel (may include HSLA,
but more recently DP and TRIP steels). These steels have become relatively
common in stamped sheet metal applications, but not with tube
hydroforming.
The perception may be that the materials are not compatible with tube
hydroforming. This is probably based on the forming mechanics of conventional hydroforming. Aluminum and high-strength steels do not stretch
easily or very much but using LPH changes the situation dramatically.
A number of examples where LPH has been used with lower formability
material includes:
1. An engine cradle that was in production for six years with 45 000 psi
minimum YS HSLA.
2. A single production die has also been used to form several grades of
aluminum.
3. Other hydroforming dies made for mild steel production have made
parts from aluminum, stainless steel, and ultra HSLA [88 000 psi yield
strength (YS)] steel with no process or die changes.
The fact that there is no strain or stretching of the material to make the
process work and therefore no hydroformed wall thinning is key to using
these less formable materials. This is true for the normal LPH design practice of being able to change the cross-section shape, while maintaining a
constant periphery along the whole part length. With HydroDieFormingTM,
cross-section expansion can be provided as well. Special materials are
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usually not needed to accommodate the process, only those that may be
advantageous to the function and cost of the part.
The most economical material or tube is that which is produced to
‘normal’ specifications that were standard in the tubing industry before the
rise to prominence of hydroforming. Where material property requirements
increase, higher costs tend to follow. This is a substantial advantage for LPH,
since often 60% or more of the total part cost is the material.
12.3.2 Properties
Material properties that commonly concern metal-forming engineers are
yield strength S and available elongation E, which is also true for hydroforming. Ultimate tensile strength is of secondary concern. The stress S is
applied to plastically deform the material. The relationship (equation 12.1)
between S, wall thickness T and the smallest cross-section corner radius
(Ri) and the process conditions applies for HPH but not for LPH. Wall
thickness, YS or corner radius may be changed dramatically with no change
in P because of the forming mechanism previously explained.
Pi =
ST
Ri
[12.1]
Therefore, the press size and cost to buy and operate are unchanged.
Elongation needed is significantly less because expansion and material
stretching are not necessary for the process to work and more formability
can be used to form part features. This is particularly pertinent for highstrength steels. The n and R values have little impact on parts made
with LPH.
When using HDFTM and expanding some cross-sections, equation (12.1)
and the material properties apply to the expansion step or stage, but required
pressure, n value and R values are much lower.
12.3.3 Tube diameter and wall thickness
A tube-making company must assess the diameter to wall thickness ratio
to determine the feasibility of producing the tube. As the wall gets thinner,
the ratio increases and it becomes progressively more difficult to control
the tube-making process, since aligning the edges for welding and maintaining weld integrity becomes more challenging. Ratios of D/T (where D is
the outside diameter and T is the wall thickness) exceeding 60 are usually
considered difficult and fewer companies can produce quality tubing. The
cost to produce thin-walled tube results in higher scrap rates and increased
handling damage compared with tubing having thicker walls. For a given
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diameter, this can result in the surprising situation that a thinner wall with
less material could cost more. If D/T < 20, the wall is heavy; 20 < D/T < 50,
medium and D/T > 60 gives a light wall.
Four other considerations to determine the most advantageous wall
thickness are stripout torque on self-threading fastener holes, welding
stability and corrosion. The first can be affected to some degree with punch
style and sizing, but the biggest influence is material thickness. Welding
stability refers to having enough thickness to allow good penetration while
avoiding unacceptably high burn-through rates. As welding technology
improves, this is becoming less of an issue for thin material. In applications
where corrosion may occur in service, resulting weakening should be carefully considered, but prevention is preferred.
The largest diameter, thinnest wall tube will provide a given structural
rigidity with the lowest weight. However, factors mentioned above may
cause the wall thickness to exceed this minimum.
Using thinner walled tubes is less of a problem when using LPH. Since
there is no thinning of the wall during hydroforming, the likelihood of
rupture is avoided. With HDFTM, controlling material forming during the
whole cycle and greater variability of pressure during the forming process
makes it more suited to forming thinner tubes.
While tubing may be produced in any diameter and wall thickness desired,
it is recommended that common diameters be used whenever possible, in
available increments (i.e. 5 mm or ¼ in) increments. Specifying a diameter
between these increments tends to increase cost by having to purchase
special sized tube-mill tooling and each production run is inherently ‘a
special run’. The attraction to do so is the effort to minimize weight and
cost, but a more effective way is to use a thinner wall and/or a larger tube.
During the prototype development phase of a project, using a common tube
size can significantly reduce cost and required lead time for material. In
the unpredictable, fast-response automotive industry, this can be a large
advantage.
12.3.4 Tube making
Mechanical steel tubing is the most common (and lowest cost) tubing used
for hydroformed tube blanks and is used for any known LPH parts. This
contrasts with the relatively frequent need to use more expensive tube types
for HPH parts. Factors that should be considered when selecting tube diameter and wall thickness include required section strength, deflection limits,
available packaging space, tendency of flat surfaces to ‘oil can’ and tube
manufacturing capabilities for high D/T ratios. The first three should be
naturally taken into account by designing to satisfy demands on the part in
the space available, as with any other forming process.
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There are three main tube types: electrical resistance welded (ERW),
welded seam, and drawn over mandrel (DOM) or seamless. The last is
relatively expensive and uncommon for hydroformed parts. ERW tubing is
the most economical and should be used if it satisfies part function and
process requirements. Welded seam tubing refers to that made using other
welding techniques such as laser and TIG. These deliver a more consistent,
stronger weld, but are more expensive since the production rate is slower.
LPH process requirements are satisfied by ERW tube, while the weld
strength and consistency has been seen as inadequate for HPH applications.
This is most commonly addressed by using welded seam tube.
The ERW process rolls the strip into a round shape and just before the
shaping is complete the edges are heated, softened and pushed together.
Some excess metal (flash) is bulged into the inner and outer surface of the
weld and is sliced off during manufacturing to prevent problems with the
bend tooling. Tubing used in LPH applications are produced to the ASTM
513 standard for mild steel or equivalent for galvanneal products. Thus, the
increased costs that accompany more stringent requirements are avoided.
For HPH applications where expansion is required to avoid pinching or
varying the periphery, a special material with higher formability, particularly n value, is commonly required. It is logical to preserve these heightened levels through tube making, bending, and preforming as much as
possible. In many existing tube mills, the relatively short distance over which
the flat strip is rolled into the normal round shape causes a greater reduction in these two properties than is desirable. Special tube-mill design can
improve this situation substantially by using more roll stands and new
rolling tools to develop the cross-section shape more gradually. This requires
major modification to an existing tube mill and/or purchase of new equipment. LPH parts do not require specially made tubing.
ASTM A513 has proven to be inadequate for many HPH applications.
Some factors such as mechanical properties YS, E and n value must be
controlled more carefully. Others include outer and inner surface finish,
length, wall thickness and cleanliness. These heightened standards increase
cost because of higher scrap rates, level of inspection and testing being
required.
12.3.5 Tube bending
Bending is necessary for most automotive structural parts. A number of
bending techniques are used in various industries, but the most common
and advantageous for these parts is CNC rotary draw mandrel bending. The
chief benefits are speed, accuracy, repeatability, minimizing cross-section
reduction around bends and better control of wall thickness variation from
bending (thinning on the outside of the bend, thickening inside).
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Pressure
die
Clamp
die
Clamp
straight
Wiper
die
Bend
die
External tooling
Bend angle
Mandrel
body
Centerline
bend radius
Mandrel
balls
Internal tooling
12.3 Typical rotary draw bend tooling configuration.
Usually the most difficult part of making a thin-walled part is bending it.
With LPH processes, if it can be bent, it can be hydroformed, but bending
sensitivity may lead to using a thicker wall. Product design focuses on bend
radius and angle, and the straight length between the tangent points of
adjacent bends as shown in Fig. 12.3. The first two factors use some of the
material elongation and n value; as the radius decreases and bend angle
increases, the amount of formability used increases.
For HPH, product design feasibility for a given material depends on
having sufficient formability for bending and hydroforming, while there is
more formability for bending with LPH since less is required for hydroforming. Several methods have been developed to handle these higher
HPH requirements.
Annealing returns the material to the high-elongation, lower-YS state
that the material was in before work hardening was caused by sheet
rolling, tube making, bending and preforming. Although the whole part can
be annealed, local (probably induction) annealing is common since it can
focus on the bends, which are worked the most. When annealing, a cooling
period of up to 20 min or more is necessary to prevent quenching the
material when it is exposed to fluid during hydroforming. Annealing must
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be done in an oxygen-free environment to prevent formation of black
scale, which is normally not acceptable. It is quite expensive and while it
has been used for some hydroformed parts, none made with LPH have
required it.
A second approach is to limit bend angle and radius severity to prevent
elongation from dropping too low and breaking the tube. This can avoid
the need for annealing but may require a multi-part assembly where a single
part could do the job. In addition, limitations on the number of useful features that can be designed into the part, such as abrupt indents and local
sharp cross-sectional corners may reduce the part’s utility. To achieve some
of the main benefits of hydroforming some difficult choices may have to be
made. This approach has not been used for LPH parts.
As discussed above, special high formability material can be used to have
a suitable residual formability left after bending and preforming. Although
this increases cost, it may be the best alternative.
12.4
Cross-section pre-forming
Pre-forming the tube cross-section over some or all of the part length may
be required to avoid pinching some of the tube between the die sections.
There are two main methods.
The first squeezes the round perpendicularly to the direction of die travel
(making an oval shape) where a section in the cavity is substantially
narrower than the start tube outside diameter. This has little effect on elongation, using a simple cylinder to compress the section in one direction and
allowing it to form freely otherwise. Cycle time is much shorter than for
hydroforming. All pre-forming for LPH parts is done this way.
Alternatively, the cross-section can be formed close to the desired shape
with some intermediate deformation (not wanted in the final part), then
hydroformed to complete forming and remove these deformations. This is
common for HPH parts where the difference between blank and final part
peripheries is inadequate to prevent pinching along the length of the part.
When the blank is compressed in one direction, it is also constrained in the
others. The forces required to do this are substantially higher and must be
done in a press at a cycle time more similar to the hydroforming operation.
This method has not been used for LPH parts.
12.4.1 Hydroforming
The many advantages of tube hydroforming (Fig. 13.1) make the automotive industry receptive to it. Its extraordinary ability to reshape the
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tube cross-section with a high degree of complexity and repeatability is
compelling.
Several aspects affect the design of the final part and are discussed
separately.
•
•
•
•
•
•
Variable and constant periphery design,
Wall thickness distribution,
Cross-section shape,
Cross-sectional corner radius,
Hole piercing,
Section expansion.
12.4.2 Variable and constant periphery design
Variable
It is commonly understood that all hydroformed parts must be expanded.
This is rooted in the fact that it is true for HPH, since the starting tube is
smaller than the smallest cross-section in the finished part to avoid pinching.
Therefore, all cross-sections are expanded to varying degrees to completely
form the part. It is forming these small radius cross-section corners that
creates the need for high pressure, not expansion itself. Section expansion
also is credited with increasing YS and providing superior dimensional
capability, but thorough supporting data is unavailable. This variable periphery approach and its effects are discussed in Chapter 13.
Stretching an undersized tube to fill the cavity presents several challenges
to successful forming. In many cases, the restrictions necessary to deal
with them can be greater than the freedom of variable periphery design.
Many parts that exhibit varying peripheries do not necessarily need them.
Designing and objectively judging the merits of a design for each forming
technique is rarely done, but is advisable and advantageous.
Constant
It is normal to design LPH parts so that the final part cross-section peripheral lengths equal that of the tubular blank and these cross-section lengths
are equal to each other along the part length.
This approach to part forming and pinching avoidance dramatically
lowers the internal pressure required to form complex parts. Pressure
is normally one-fifth to one-third of that needed for a similar HPH part.
Equation 12.1 does not apply to LPH because of its different forming
mechanism. Combining low internal pressure in the tube while the die
closes with a die cavity periphery that is equal to the periphery of the
blank increases design flexibility with respect to material selection, YS, wall
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thickness, corner sharpness and the extent to which the material can be
stretched. Maintaining a constant periphery throughout the part also
reduces costs by using smaller pressure generating equipment, hydraulic
presses, faster cycle time, economical material and much less power. Wall
thinning does not occur during LPH. Techniques to control it such as end
feeding, lubrication (cleaning) and high formability material are unnecessary. Having the discipline of designing a constant periphery along the part
length allows access to a number of benefits that otherwise are not apparent
and are unappreciated.
HydroDieFormingTM
HDFTM removes these LPH periphery design constraints by providing the
benefits of a variable periphery design while adjusting the part production
process to eliminate or minimize the negative cost impact. Elimination is
preferred to minimize cost, but in some instances a certain level of cost is
acceptable. Expanded sections along the part length can be included in the
design in several ways.
With HDF, many very beneficial fectures can be built into a part for free.
All that is needed is the insight to design it properly.
For economy and process stability, the goal is to devise a part design with
expansion, using the low forming pressure approach that is such an advantage. Where this is not possible, using higher pressure is always an option.
12.4.3 Wall thickness distribution
See Section 13.9.
12.4.4 Cross-section shapes
Hydroforming’s attractiveness is based on its ability to easily reshape
the tube cross-section and continuously vary it along the length of the part.
This combination of form flexibility, strength and lower weight of a tube
and elimination of joints (stress concentrations and rigidity reducers) is
structurally sensible and is a vital tool in the structural designers toolbag.
As vehicle development continues, one overriding factor will remain. Every
vehicle needs an effective, efficient structure to perform the required functions at minimal weight and cost.
Adding structural strength or rigidity often does not require expansion.
Figure 12.4 shows examples of formable shapes with no expansion that can
be combined with more normal rectangular and other shapes with LPH.
This greater ability to dramatically change the shape can be a more advantageous approach.
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b
a
c
d
f
e
12.4 Examples of hydroformable cross-sections.
Reshaping the cross-section is important for a number of reasons of
which maximizing section strength in the most advantageous direction to
withstand applied loads in the most efficient way is one of the most attractive. Others include avoiding interference with surrounding components,
better joint configuration and greater flexibility when designing joints with
brackets or other parts for welding or mechanical fastener attachment.
These shapes are made up of flat and curved surfaces, as well as corners
bent to varying degrees of severity. Many functional requirements and
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shaping flexibility can be provided by LPH. Since LPH reshapes the starting
periphery, some shapes like sharp corners and wall-to-wall sections can be
made that are difficult to achieve by HPH methods. Generally the surface
complexity of LPH parts is higher than other hydroformed parts.
HDF offers several ways to make such parts with the added option of
expanding where needed. HDF uses HPH and LPH design and process
elements. How they are used differs depending on the individual design
requirements. HDF eliminates corner thinning and offers several new ways
to make material thicker on thinner where it is most beneficial.
12.4.5 Cross-section corner radius
Corner sharpness is often expressed as a radius that is a multiple of the wall
thickness (T) in a similar way to the stamping industry. A difference is that
stampers specify the inside radius while, with hydroform parts, the outside
is usually indicated. Figures 12.5 and 12.6 show examples. Sharp corners
(3T) can be achieved throughout a production part or 1-3T in a local
area.
Sharper corners can be an advantage for several reasons. One is extending flat surfaces for hole placement near a corner. Structurally, sharper
corners improve the rigidity of a particular section. A less tangible reason
is that the part looks sharper, better defined and higher quality.
For stampings, a smaller corner radius is accompanied by an increased
likelihood of fracture. The real reason for this is that the part is formed
using tensile forces that stretch the material into place. The same concept
applies to conventional hydroforming. However, LPH forms tubular components using compressive forces. Corners of 2T and less can be formed
12.5 Localized sharp corner cross-section.
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12.6 Localized sharp corner cross-section.
9.5 mm
9.5 mm
9.5 mm
9.5 mm
6.3 mm
2.0 mm
a
12.5 mm
12.5 mm
b
12.7 Cross-sections – different corner radii.
12.8 Hole cross-section with folded back slug.
with no indication of cracking. While corners this sharp would rarely be
needed throughout a part, in local areas it can be an asset.
Figures 12.7 and 12.8, as well as Table 12.1, demonstrate that corner
sharpness can decrease by 30% with no effect. Even a decrease of 70% may
only generate a 10% pressure increase.
12.4.6 Hole piercing
Piercing in an LPH die offers the most economical and dimensionally stable
method of providing holes or openings in a hydroformed component.
Adding piercing to the forming operation increases the hydroform die cost,
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Table 12.1 Cross-section corner radius of Fig. 12.7
Figure
number
Material
grade SAE
Wall thickness
minimum = T (mm)
Internal
pressure
(MPa)
Outside corner
radius
Yield
stress
(MPa)
12.7a
12.7b
1010
1026
2.0
6.3
37.2
41.4
4.75T
1.5T
263
386
but will not significantly affect capital, labour, or operating costs. Combining
forming and piercing into a single operation ensures that the holes will
always be punched in the correct orientation to the part surface and to the
highest degree of accuracy possible, since the part is positioned where it is
formed. Versatility is very high and increasing as techniques improve to
allow greater design flexibility of size and particularly position, both anywhere on the tube surface and relative to each other. Almost all holes
(highest known hole count is 64 in a 2-m-long part) are punched in LPH
production applications. This is normal practice in the stamping industry.
Using HPH reduces the holes that can be punched for several reasons,
including higher pressure requiring larger punch units, interference with die
inserts and the strength of the die to contain higher pressure fluid.
Using the pressurized fluid in the part as the punching backup results in
some different piercing characteristics compared with stamping. Punch life
is usually very long in a hydroform die because there is no die button, as
found in a stamping tool, for the punch to be misaligned with that could
cause punch damage or excessive wear. The surface immediately around a
hole will be indented or rounded slightly toward the inside of the part.
Higher forming pressure reduces the amount of indenting around holes, but
has other drawbacks that are discussed throughout this chapter.
For LPH systems, whenever possible the slug made by punching is left
attached to the tube. This prevents loose slugs from falling into the die
cavity, marking the following parts or rattling in the finished component.
The amount the slug is folded back can be changed depending on how the
punch is made. When the slug interferes with hole function, it may be
removed by a modified punch arrangement.
The hole shown in Fig. 12.8 is an extruded hole, which intentionally has
a flange formed around the hole and can only be formed in the hydroforming die. This type of hole is most commonly used for self-threading fasteners
where indenting acts as a lead in to start the screw.
Figure 12.9 shows a slug being pushed out from the cavity by internal
pressure. Fluid leakage from hydropiercing such a hole may be considerable, but it is not a problem if the tool is designed properly. Large holes can
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Pi
12.9 Internal pressure pierced hole.
be made this way to a high standard and it is a cheaper alternative than
cutting later with laser, plasma or other means mounted on robots.
Maximum flexibility in hydropiercing is important for the highest quality,
most economical product. Many hole shapes are in production including
round, oval, rectangular and hexagonal. Hole size is only limited by what is
reasonable given the part geometry. Holes that cannot be hydropierced
require a secondary or post forming cutting or piercing operation at additional cost and reduced repeatability.
12.4.7 Tube section expansion
Section expansion of hydroformed parts has attracted a lot of attention and
has been closely associated with the need for high forming pressure. This is
not necessarily true. For tube diameters and wall thicknesses that are normal
for automotive structural parts, it is typical that pressure to expand when
the table is round is low. Referring to equation 12.1, a 3 in-diameter mild
steel tube (YS = 240 MPa [35 000 psi]) with wall thickness of 1.5 mm will
only need 10 MPa (1500 psi) to expand. For a given material and wall thickness, it is the reduction in effective radius that dictates elevating the pressure to completely form the part. In other words, forming the corners drives
the need for high pressure.
The reasons for expansion discussed here are to improve structural
strength or satisfy mounting conditions, not the smaller amount needed for
HPH processing. The amount will probably be applied in a local area or at
most probably less than half the part length.
It is commonly understood that expansion in the hydroforming die is
the best and most economical method, but this seems a misconception in
many cases. Expanded sections of 10% and above are mostly possible at
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the component end, or close to it where extra material can be forced in to
compensate for wall thinning. Anything that increases friction, such as
bends, low, wide rectangular shapes, dents or increasing distance from the
end, impede feeding additional material in and limit how far that end
feeding is effective. Expanding in the die requires large end-feeding
cylinders and hardened inserts to withstand the intense abrasive forces that
are generated. Cycle times are longer and special material is commonly
needed to meet the deformation challenges.
Designs with expanded sections extending beyond the first bend may
be best done in the hydroforming die to simplify bending, but this assessment depends on the ‘extra’ things necessary to successfully form the part.
Conditions such as high material formability, the ability to effectively feed
material around the bend, as well as the amount and concentration of
expansion, have to be suitable to make production feasible.
The approach taken with HDF extends the part design possibilities
beyond the non-expansionary limitations associated with LPH and can
reduce or eliminate the increased cost factors that accompany expanding
with the HPH approach. The amount of expansion HDF can achieve is
potentially larger than HPH, without drawbacks such as high pressure and
uneven wall thickness (where end feeding is ineffective).
Additionally, several HDF techniques have been developed to control
wall thickness and put it where it is needed to a degree previously unavailable. Larger sections with thinner or thicker walls is one option. Thicker
walls in localized portions of a cross-section while others are thinner are
another option. Using these techniques to tailor wall thickness during
hydroforming represent opportunities for substantial weight reduction that
are needed for renewed industry efforts to reduce vehicle weight at a
reasonable cost.
12.5
Conclusions
There are many design considerations and interrelated complexities to
consider when designing hydroformed structural components. These include:
material type and properties, tube diameter and wall thickness, tube making,
bending and preforming, as well as hydroforming factors such as variable,
constant and HDF periphery design, wall thickness distribution, crosssection shape, corner radius, hole piercing and section expansion.
It is best to make informed judgments and it is unlikely that efficient,
effective parts can be designed to maximize the benefits of tube hydroforming without knowing as much as possible or working with a company that
does. Keeping all options open to use the technology most beneficial for
any new part is the most likely path to competitiveness in the automotive
industry.
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12.6
Reference
1. morphy g, Product Design for Hydroformed Structural Components, TPA
4th Annual Automotive Tube Conference Focus on Hydroforming, Dearborn,
Michigan April, 1999.
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Comparative analysis of
hydroforming techniques
G. MORPHY,
Excella Technologies Inc., Canada
13.1
Introduction
In the past 20 years, hydroforming of structural parts has gone from development to a mainstream commonly accepted practice to make a wide
variety of shapes in many applications. Automotive structures comprise the
bulk of the hydroformed parts in production today, because of the number
of beneficial attributes and their value.
There are a number of manufacturing options that can be used to produce
automobile parts. Examples of common techniques are shown in Fig. 13.1.
The wisest approach is to choose the technique that will deliver the functional performance needed for the lowest cost. Giving too much importance
to either performance or cost will result in a less efficient solution. An objective view of the different methods is important to choose the best balance.
Putting too much weight on performance is a key reason why hydroforming has come to be seen as an expensive process. Making cost a distant
second consideration can be rationalized for a while, but eventually it
becomes an undeniable issue. A second reason is that insufficient innovation to reduce cost perpetuates the disparity. These factors lead to a plateau
in demand and may even lead to a downtrend in applications where hydroforming is used. Applying hydroforming differently can allow design
engineers to devise more effective and cost-efficient structures for vehicles
and other applications. Efficient processing is required to achieve this.
When faced with such a problem, there are several avenues one can
consider. These are:
1. Trying to do what you have been doing, but aim to do it better;
2. Decide you have tried everything and the current state is as good as it
can get;
3. Try something different;
It should be pointed out that comparisons might seem questionable considering the comparer’s perspective. Working for a company promoting one
287
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Forming
complexity
Stamped and welded
assemblies
2 stampings 4+ stampings
Roll forming
Stamping
Tube forming
Relative process cost
13.1 Comparative cost and forming complexity of metal
forming processes.
process can be seen to affect objectivity. While my experience has been with
LPH but not dependent on it, my preference is solely guided by using the
most effective method to produce parts with best performance, quality,
process stability and lowest cost. This is what most people who consider
using hydroformed parts are interested in.
While this hydroforming development for structural applications was
occurring, hydroforming was already being used to form very high volume
plumbing T-shapes and low-volume parts for aerospace. The HPH process
was adapted from the plumbing T process.
This comparison of different hydroforming methods has the benefit of
more experience and perspective than was possible earlier; for knowledge
is gained if you pay heed to the signs and results to arrive at logical conclusions. Thus, we can plot the most sensible course for the future. People are
always eager to know future trends and the best path to follow. We can
determine what these are if we choose carefully.
Study the past if you would define the future.
Confucius
13.2
So many options; how to choose?
An important thing to remember is that parts do not have an allegiance to
be made in one way or another. It is people who choose, and often technical
decisions are made based on partial information and logic. Sometimes
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Forming
complexity
Tube forming
Hydro die forming
Low pressure
hydroforming
Liquid impact
forming
High-pressure
hydroforming
Mechanical
forming
Hot forming
Bending
Relative process cost
13.2 Comparative cost and forming complexity of tube
forming processes.
these are referred to as business decisions, meaning presumably that a nontechnical line of logic dictates doing what is less technically sensible. This
has been a factor in the development of many manufacturing technologies,
including hydroforming.
Effectively using innovation to get the result that you want while minimizing the disadvantages is the best concept, but is easier said than done.
It is an important principle because it should always be the aim.
Figures 13.1 and 13.2 compare forming complexity and process cost for
a number of processes. What is needed for the part being considered determines the right manufacturing technique. Within each of these generalized
options lie more detailed suboptions, such as Fig. 13.2 which is a more
detailed view of the tube forming category.
13.3
Roll forming
This approach to making parts is inexpensive, but also limited in part design
flexibility. However, the most sensible approach is to take full advantage of
its economy for parts whose functions can be achieved. The cross-section
shape must be consistent along the part length.
Additional features can be added with operations such as bending and
localized forming to deviate from normal uniformity. These operations
increase cost, but, if they achieve the desired function, are still worth it.
13.4
Stampings and assemblies
When comparing methods to make structural parts, the discussion should
be relative to the most common incumbent technology that has been used
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and refined for a number of years. Stamping metal parts into desired
shapes has been used since cars started being made. As referred to in Fig.
13.1, stampings are more expensive than roll forming due to slower
throughput and greater complexity, but it is an inexpensive way to make
non-uniform cross-section parts. As with roll forming, where the required
part functions can be addressed with a stamping, this should be used for
best economy.
Structurally stampings often fall short since the nature of the process
only allows open cross-section parts. Decades ago designers recognized
that this shortcoming could be improved by combining two or more stampings into a closed cross-section assembly or tube-like structure. Tubes have
been long recognized as an efficient load-bearing structural element.
Assembly is usually by means of welding: resistance, MIG, laser or another
method.
Closed rectangular structures are important in vehicle design. No viable
way existed to form tubes into complex shapes at required cost and quality
levels. Adding holes was more costly than in a stamping and some hole types
were unachievable. Over time, designers and engineers recognized a number
of other shortcomings and sought better solutions.
13.5
Tube forming
Tubes can perform required structural functions using less material and may
cost less than stamped and welded structures without some of their drawbacks. Figure 13.2 shows several potential processing options ranging from
simple tube bending through several hydroforming methods to warm or hot
forming.
The structural and material use efficiency of tubes has led to a number
of tube forming techniques that offer differing degrees of forming complexity and capability. Accompanying this is the reality that cost increases as
well. As mentioned above, the guiding principle should be using a process
that performs the required part function, but simultaneously minimizes
cost. Sometimes when the design community is confronted with a new
technology, there is a tendency to think that it is good for more applications
than it really is. This is because a ‘common sense’ of what is most appropriate, logical and how it would work out in the final analysis is still being
developed. Simpler forming processes, with lower capital, part and tool
costs are best to choose where their limited forming capabilities are adequate for the functions that are needed.
Tube hydroforming removes these constraints, and increases the range of
applications whose requirements can be satisfied with tubing. Increased
strength, reduced variability, welding, weight, tooling cost, capital cost, and
part count can be realized simultaneously.
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Capital cost
Tool cost
Piece cost
Assembly cost
Diminishing variability
Number of parts
Number of joints
291
Strength
Design flexibility
Bend, torsion rigidity
Overall part quality
Strength/weight
Crash performance
Integrate fasteners
a
b
13.3 Advantages and disadvantages of tube hydroforming.
More demanding bending and torsional rigidity requirements can be
partly addressed by reshaping part cross-sections for best stiffness in the
plane of in-service force application, a chief strength of hydroforming.
Intrusion by surrounding components can be avoided by abrupt surface
changes. Expanding cross-sections, forming high-strength steel and other
materials, sharper corners, thick-walled tube are among some capabilities.
Hydroforming has become popular because it increases the number of
achievable part design features; reduces weight and cost by providing more
advantageous manufacturing options. Where the sum of these benefits is
large enough, the motivation is created to use hydroforming to make a more
efficient structure. Some are well recognized, while others are less so. In any
case the main advantages and disadvantages are shown in Fig. 13.3.
When hydroforming, as with all industrial processes, part design dictates
what manufacturing steps are needed, and the efficiency with which they
can be produced dictates the cost. To make sensible decisions, a thorough
understanding of how differing versions of this new process work why
certain actions are taken and their cost are essential to achieving the desired
efficiency improvements.
Four forming processes will now be described in terms of how they work
and compared in a number of respects.
13.5.1 Tube hydroforming processes
A natural consequence of forming a tube into a reasonably complex shape
is that it will get squeezed between the die halves or pinch. Four ways to
force it to form are: low-pressure hydroforming (LPH), high-pressure
hydroforming (HPH), liquid impact forming (LIF) and hot forming (HF).
These are the focus, but there are also other process options.
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High-pressure hydroforming
The high-pressure hydroforming (HPH) process avoids pinching by designing the tube periphery to be smaller than the desired finished product. The
process starts with a round tube, Fig. 13.4a and fully closing the die,
Fig. 13.4b before filling it with water, Fig. 13.4c. Often, one of the results
Before pressure
die open
a
Before pressure
die closed
b
Pressure applying
die closed
c
Maximum pressure
die closed
d
13.4 High-pressure hydroforming process steps: a before pressure die
opens; b before pressure die closed; c pressure applying die closed;
d maximum pressure die closed.
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FHPH
Reduced
thickness
Pi
T
FHPH
13.5 Close-up of typical HPH cross-section corner.
is undesired ripples in the tube wall roughly perpendicular to the tube
centerline as shown in Fig. 13.4b and c.
A second more significant characteristic is the difference between the
tube and die cavity peripheries. As a result, the cross-sectional corners
are unfilled when the die is closed and material thickness is the same as
the bent tube blank. Pressure is applied such that material stretches into
previously unfilled areas of the die cavity. Increasing pressure completes
forming as shown in Fig. 13.4d.
The result is uneven wall thickness, often in the pattern shown in Fig.
13.4d and 13.5. Both show that internal pressure stretches the material into
the corner, creating tensile stresses in the tube wall. The wall thins because,
as forming progresses in Fig. 13.4b to c and d, more of the tube wall contacts
the die cavity surface. Pressure needed to stretch the material is high enough
to make it to stick to the die surface and stretching occurs in a decreasing
portion of material, which concentrates wall thinning.
Strategies to reduce wall thinning are employed and work well with some
limitations. Examples are end feeding, lubrication, annealing and high n
value material. End feed is only effective for a limited distance (i.e. 500 mm)
from the tube end. All four options increase piece cost.
Low-pressure hydroforming
The low-pressure hydroforming (LPH) method of preventing pinching is
less intuitive for most people than HPH. It is avoided by forming the tube
filled with low-pressure water as the die closes on it. At any cross-section,
its periphery is essentially equal to the die cavity. For this chapter and
Chapter 12, the general term LPH describes forming processes using this
concept. The different approaches are explained in Chapter 12.
As seen in Fig. 13.6a, the die halves contact the round tube. The normally
round starting shape is forced toward the intended shape as the die closes
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Before pressure
die open
a
Before pressure
die closing
b
First pressure stage
die closing
c
Second pressure stage
die closed
d
13.6 Low-pressure hydroforming process steps: a before pressure die
open; b before pressure die closing; 1st pressure stage die closing;
2nd pressure stage die closed.
in Fig. 13.6b. The tube is filled with water (signified by shading) in Fig. 13.6c
and low pressure is applied while the die continues to close.
What distinguishes this process is that the periphery or circumference of
the starting tube is essentially equal to the desired periphery of the finished
part. Coupling this with low-pressure water to keep the tube wall in contact
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FHDF
R 3T
Pi
T
FHDF
13.7 Close-up of typical HPH cross-section corner.
with the die cavity wall (but still able to slide) effectively harnesses the large
mechanical force generated by the closing press and die. This force resolves
into compressive components that act parallel to the tube wall as shown in
Fig. 13.5. The water resists unwanted deformation inward (i.e. crushing or
ripples in the metal), while the die prevents outward deformation, which
facilitates a controlled reshaping of the cross-section throughout the length
of the part.
This combination of hydraulic and mechanical forces used during the
cycle facilitates complex part forming. Using low pressure water when the
die is closing allows material to slide on the cavity surface and get pushed
into place using the power of the press.
As the die closes, forming is almost complete. Pressure is increased to a
higher level (normally <60 MPa) to produce the final form of the part as
shown in Fig. 13.6d. Holes are punched at this stage since the backing force
provided by the water is greatest. The other reason for this higher pressure
is to flatten the planar areas, not to finish forming the corners. It takes much
less pressure for the former than the latter (Fig. 13.7).
Liquid impact forming
The liquid impact forming (LIF) process provides many of the attributes of
tube hydroforming, but uses a stamping press rather than the hydraulic
press that is normal for tube hydroforming. Faster cycle times (8–15 s) also
contribute to it being a less expensive option than hydroforming. For these
reasons, parts that can be formed with LIF should be. Those that cannot
should be hydroformed.
This process was developed to operate in a stamping press, which removes
the investment required by most companies to hydroform parts with the
less common, slower hydraulic press. Many companies who would or have
considered doing this have stamping presses. The tube is submerged, ends
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are capped, and the tube is squashed in a specific area. The water that is
displaced is utilized to reshape, bulge, or support the tube in a specific area.
At the bottom or any point during the press stroke, the internal pressure
(fully adjustable) can be released and the finished part removed (Fig. 13.8).
This process forms a tube into a variety of straight, twisted or curved shapes
as shown in Fig. 13.9. Depending on the final product shape, it may be necessary to pre-bend the tube to fit into the die. Because the tube is fully filled
and supported with liquid, it can be pierced at the bottom of the press
stroke, reducing the tendency for the metal to ‘cave in’ around the pierced
hole.
a
b
13.8 Liquid-impact forming die sequence. Courtesy of Greenville Tool
& Die, Greenville, Michigan.
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a
b
c
13.9 Liquid-impact formed parts. Courtesy of Greenville Tool & Die,
Greenville, Michigan.
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Hot forming
A fourth processing option, hot forming (HF), involves heating to dramatically increase formability. It reduces the flow stress and the strain hardening
induced by the shape change. The reduced hardening is based on temperature-induced recovery and recrystallization effects. The linear dependence
of the forming pressure on the flow stress results in the ability to form the
parts using significantly lower pressures and hence require lower tonnage
presses. The reduced hardening effect leads to a significant increase in formability of the material. Even materials with very limited elongation at room
temperature, i.e. glass, some high-strength steel and aluminum alloys, exhibit
a marked increase in formability at elevated temperatures.
Using air (nitrogen or argon) as the pressure medium, heated (>350 °C)
metallic hollow blanks, tubes, and extruded profiles can be formed by means
of internal pressure. Material expansion in excess of a 270% increase in
circumference has been demonstrated at production-intended cycle times
of 20 s. As shown in Fig. 13.6, the tool and tube are heated, with the hottest
point in the area of maximum forming. The tool temperature is monitored
and controlled by a thermocouple (approx. 30 °C less on the left compared
with the area of the thermocouple on the right). The desired temperature
profile along the tube is generated by means of inductive heating outside
the tool. The tube ends are kept at a lower temperature to maintain material
strength to aid with sealing and to achieve a controlled amount of axial
feeding without the tube buckling.
Using hot forming, a high level of complexity can be reached that would
not be achievable with room-temperature hydroforming. So this process
increases the range of parts that are formable, including machined pieces
such as this part. The internal surfaces will be smoother than the machined
component as well and deliver better fluid flow. It will win business for parts
where its cost is lower or where part features are better than the current
method. However, when parts can be made with both processes, it makes
sense to hydroform at a lower cost. To see the procedure for Schuler’s
HEATforming technique see Chapter 16.
13.6
Commonly held misconceptions
Process comparisons can be done in several ways. Since the aims are to
improve understanding of the overall subject and to provide the details to
help those interested in hydroforming make more informed decisions, two
approaches are taken. First, the misconceptions people have about hydroforming processes will be discussed based on feedback gathered over a
number of years. Second, and of equal or greater importance, the features
that are important for tube hydroformed parts and the relative ability of each
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process to provide them are discussed. Comparison is mostly limited to LPH
and HPH, since they have been used more extensively and longer.
13.7
Low-pressure hydroforming
13.7.1 Not suitable for making large parts
Tube hydroforming makes use of internal pressure to form metal tubing.
For parts of any size, the force exerted on any part of a die, the press and
other components increases proportionately to internal pressure. For a part
that can be made by either LPH or HPH, using the former will inherently
exert less stress on the tool and equipment. Commonly, the pressure required
for LPH is 1/3 to 1/5 of that needed for HPH. It makes sense then that, by
extension, the benefits that are realized with smaller parts increase in proportion to part area.
It seems likely that this understanding is based on the fact that companies
using LPH made business decisions not to produce large parts like frame
rails. Additionally, LPH as it has been applied since inception has not
included expanded cross-sections in any production parts. Local expansion
is a common advantageous feature for frame rails and cross-members and
is seen as a requirement by many designers and engineers.
New developments in LPH technology, like HydroDieFormingTM
(HDFTM) have removed this limitation. As a result, the range of part design
options for any parts is even greater than has been feasible with LPH or
HPH alone, especially including large ones. When using (HDF), large
expansions potentially anywhere along the part length, and cross-section
shapes that were not considered feasible earlier, are now possible. Other
improvements when using LPH for large parts are the much reduced likelihood of die cracking and wear, no required lubrication, as well as being able
to make longer parts (i.e. frame rails) with lower clamping tonnage.
As will be seen elsewhere in this chapter, there are no other limitations
that prevent making larger parts with LPH, and lower forming pressure
actually facilitates making much larger parts with a given tonnage press. So,
contrary to this contention, LPH is actually advantageous to make large
parts.
13.7.2 Can only form simpler parts
This seems to be based on the idea that pressure is needed to force
the material where you want it to form the required shape. So, when you
have less pressure, forming complexity must be reduced. It seems to be
unrecognized that other process methods can change this assessment
dramatically.
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There are a number of examples of production LPH parts that have
features and cross-sections that would be difficult or not feasible for HPH.
Generally, the lower forming demands of LPH allows bending to be
more aggressive, cross-section corners to be sharper, material to be less
formable and wall-to-wall cross-sections (Fig. 12.4), while maintaining
even wall thicknesses show excellent fatigue durability and good process
stability.
Recent development efforts further improve the part design options
when using the HDF approach. This can be key to the economic sense of
hydroforming a part. Having more options and features that can economically be formed in the part allows designers to perform more functions with
a part or an assembly with little more, similar or even lower cost. It is the
difference between a knife and a Swiss Army knife. Having more design
and processing tools leads to more functional, economical solutions and
being able to address a wider range of needs and functions. Generally,
LPH part complexity is higher today and the options to go even further
with HDF are available now.
13.7.3 Low pressure limits process capabilities
Any manufacturing process has limitations including each of those discussed in this chapter. The fairest comment is that the limitations of each
is different, including HPH and LPH. Where the use of high fluid pressure
with HPH does provide the means to force material to form the shape
required, success depends on the material being able to form without
rupture. Also, a part with a continuously changing cross-section periphery
likely requires high pressure to form the corners.
Alternatively, LPH can make a product with large differences in wall
thickness, material yield strength and cross-section corner radius without
changing forming pressure where for HPH each factor would alter it
proportionately according to equation 12.1.
As for process stability as expressed by uptime and dimensional
variability, it seems anecdotally that the LPH approach provides similar to
better results, apparently because the forming mechanism is less affected
by variations in the incoming material and parameters.
13.7.4 Low pressure cannot expand
It is common to combine straining the material (expanding) and changing
the shape of the tube while forming it, as with the HPH process. Reshaping
by closing the die immediately before expansion substantially reduces
the radius or Ri in equation 12.1. This greatly increases the pressure
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required to expand the material. Simultaneously, material starts to work
harden, further increasing the pressure needed, which is the reason for the
H (high) in HPH.
Instead, by expanding in the round, relatively low pressure is required,
usually 2000–3000 psi for normal tube used for automotive structural applications. This pressure is much lower than normal maximum LPH pressure.
So the question now becomes when during the process expansion happens,
rather than whether it can be done at all. The relationship between pressure
P, yield stress S, material thickness T and smallest corner radius Ri being
formed is shown in equation 13.1.
13.7.5 Low pressure is less able to accommodate part
design changes
This can be looked at during the part program or as one program is ending
and considering what will replace it. For the first case, it varies according to
what is changed. For example:
1.
2.
3.
4.
5.
The part surface location or shape changes or the options are
a Cut more die material;
b Build up and reshape the surface – limited amounts;
c Remake that section of the tool – inserts or new section;
d The situation is similar to what it would be with HPH;
e Where local strain increases substantially, review carefully.
A hole location needs to be changed:
a Move a punch unit – in-die hole punching is common with
LPH;
b Reprogram robot – many holes are cut with laser after hydroforming with HPH – press size or die strength may be inadequate.
Material is being changed to higher strength steel:
a Make adjustments for differing springback – no LPH process
change;
b Increase forming pressure proportional to material yield with
HPH.
Increase wall thickness:
a Forming pressure does not change with LPH, but needs to increase
according to equation 13.1 with HPH.
Make corner sharper in at least one local area:
a Forming pressure does not change with LPH, but needs to increase
according to equation 13.1 with HPH.
Material formability needs to be examined, as well as related topics such
as end feed and lubrication, which latter also affects when there is a model
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change. Substantial differences in wall thickness, material strength and
corner sharpness do not change the LPH forming pressure, making it likely
that the equipment will still be able to make the new version of the part.
Similar changes using the HPH approach can lead to needing a larger press.
So, the solution depends on what the changes are, but often LPH is more
able to adapt to alterations in part design.
Pi =
13.8
ST
Ri
[13.1]
High-pressure hydroforming
13.8.1 Process stability can be challenging/troublesome
For any metal-forming process including hydroforming, material consistency is important. The lack of it can mean having tubes rupture and excessive scrap, having to frequently adjust the process conditions to the material
properties and perhaps even to reject it if a consistent process cannot be
achieved. This is particularly true when expanding the cross-section as is
prevalent with HPH.
13.8.2 Expanding (and work hardening) material creates
higher strength steel (buy low carbon steel;
work harden to create HSS)
This is a concept that has been promoted as a chief advantage of HPH since
its beginning. There are several considerations that make this claim seem
optimistic at best, misleading and its reliability questionable. However,
it is clear that engineers and, maybe more so, purchasing managers in the
increasingly cost-sensitive automotive industry find the concept of getting
something for nothing highly compelling.
The most common form in which this is explained, because it is most
economically attractive, is to buy low carbon 1008 or 1010 steel and with
work hardening increase its strength, approaching that of HSLA. The
former is the cheapest grade of steel, widely available with wide specification ranges. The first two are advantageous when purchasing flexibility is
desired, while the third results in the material properties varying widely
from material lot to lot received over months and years.
There are two points to consider as a result of this. The first is that
straining the material by expanding the cross-section will not reduce this
variable property effect, thus it cannot be ‘made into HSLA’. It will increase
the strain states of different points around the tube, concentrated in the
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13.10 Stress distribution in typical cross-section corner.
cross-section corners, but as long as rupture is avoided that may not matter.
The second is that this variability in the properties of the tubular blank
makes the amount of benefit or increased yield strength that results just
as variable.
A third issue is the unevenness of the effect due to the uneven strain that
is concentrated in the cross-section corners. This is illustrated in Fig. 13.10
where the strain is much higher in these corners. This is balanced by the
small strain occurring on most of the flat surfaces.
It is normal in automotive part design to assume the minimum properties.
For these reasons determining what a reliable gain is and modeling it
accurately must be done carefully. Since this minimum is likely to be low
and require extensive efforts to quantify it, the position of a number of
designers and companies is to not include this benefit in FEA calculations.
Including this factor as it occurs instead of assuming a minimum, also substantially complicates these calculations.
13.8.3 Expansion and high pressure needed to
form more dimensionally stable parts
Springback is the bane of all metal formers. The material does not stay
exactly where the forming process put it, but returns to some degree to the
shape that it was. This characteristic increases as the yield strength of the
material increases. Springback or elastic deformation is eliminated or at
least greatly reduced by plastic deformation.
This has been a problem for all of metal forming history that experts
and manufacturers have had to deal with. Hydroforming is no different,
but it substantially reduces variability compared with a similar part
made by other methods, such as stamp and weld as well as the variability
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introduced by previous processes, such as bending. It is also apparent that
it does not eliminate springback or variability. Using higher pressure offers
a logical avenue for the lowest variability, since it expands the cross-section
along the whole part length.
Later in this chapter more detailed information is provided, but, in
summary, the tolerances that can be achieved in a number of situations
using LPH was established in a 1998 SAE paper. This is still largely true
today, but no similar data are publicly available for HPH parts, thus preventing factual comparison. However, anecdotal opinions from HPH users
indicate that, while HPH may be more accurate in some circumstances, in
others it is not, leaving the general assessment that the two processes deliver
similar tolerances.
It seems that warm/hot forming has a big advantage in this respect, but
this is not recognized because of a lack of production or capability data.
Generally, springback is more of an issue at lower temperatures and a
heated process is likely to be more accurate.
13.8.4 High pressure needed to form more complex parts
While this seems conceptually sensible, the situation is more complex than
this simple statement would indicate. It is true that high pressure fluid has
the greater impetus to force material to go where you want it to (i.e. more
complex geometries) compared with low pressure. The catch is that the
material must not rupture or tear. Theoretically, the fantastic shapes
that could be formed, become more conservative and rounded when they
actually are.
Though counterintuitive for many, experience shows that the relationship
between part shape complexity and the pressure being used can also be
characterized in another way. More complex cross-section shapes have been
made using lower pressures, perhaps since the material has more opportunity to move relative to the die cavity surface. High pressure inherently
forces the material against the wall forcefully enough that sliding is
prevented.
In summary, it is best to work with a hydroformer that is adept with both
approaches to allow use of the most suitable one, combination or innovation to achieve the most advantageous part design.
13.8.5 Greater design flexibility
This general assessment is firmly lodged as an advantage of using highpressure fluid and this is important for design engineers. However, when
you separate the usually interwoven threads of design flexibility into their
logical strands it once again does not agree with the evidence. As is detailed
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in a 1999 SAE paper focused on this subject, there are few design characteristics where using high pressure holds a clear advantage.
The design flexibility categories are:
1. Variable vs. constant periphery design;
2. Material type, properties, and thickness (overall and process caused
variation);
3. Bending;
4. Finished part cross-sectional corner sharpness;
5. Part features (forming severity);
6. Holes;
7. Dimensional stability;
8. Process conditions;
9. Cross-section expansion;
10. Future ability to adjust to design changes.
Of these categories, (9), for cross-section expansion, may improve some
design options by using HPH whereas the others are either improved when
using LPH or are similar for either.
13.8.6 Needed to reduce process steps and cost
Reducing process steps is a good goal, but may come with the presumption
that it reduces cost. It does in some cases, but not in others and it is important to determine which is true in any circumstance. Another factor that
needs to be considered is the potential for limitations, compromise or even
an increase in some other aspect of cost that may be inherent in eliminating
one or more process steps.
This, as is the case with a number of design and processing subjects, is
one where general explanations fall short, since there are a number of
considerations essential to arriving at the right conclusion. To illustrate,
examples are:
1.
Eliminating a tube pre-forming step:
a May increase limitations on parts design;
b Increase process instability leading to unforeseen scrap and other
extra costs.
2. Punching holes in the hydroforming die to avoid the higher cost of
cutting after forming:
a Better hole location accuracy and repeatability, extruded holes
cannot be subsequently formed;
b May be a small increase in cycle time.
3. Eliminating lubrication during hydroforming:
a Reduces cost for applying and removing;
b May decrease part formability or process stability and add cost.
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The final way to determine the best way to proceed depends on relative
assessments and interactions. The magnitude of the advantages and disadvantages of these examples makes the most prudent way to proceed evident.
Employing as much knowledge as possible (internal or external) leads to
the most productive design.
13.8.7 Always better to expand in the hydroforming die
This is accepted knowledge when hydroforming parts. It is also tied with
several other concepts like eliminating a processing step by doing expansion
while hydroforming as well as increasing material strength by straining
material during forming. It is a widely accepted approach to expanding local
sections of a part.
This complex mix of issues is often oversimplified. Other approaches can
be used to accomplish a similar part function. The only debate should be
‘Which is best?’. For most designers this means ‘Which manufacturing
sequence gives the broadest range of product design options’? and/or
‘Which is cheapest?’.
Fully answering three fundamental questions should help:
1.
2.
3.
What is the real cost of each method?
What are the advantages?
What are the consequences and limitations?
Shortcuts taken in this assessment may lead to surprises. An additional
process step clearly adds cost. However, it must be determined how much
and be compared with the alternatives. Expanding before hydroforming has
a number of advantages, when used to best advantage. These include much
improved wall thickness control and material flow over the part length, less
corner thinning, larger expanded sections, ability to use lower formability
(lower cost) material and improved process stability.
Expanding in the hydroforming die saves a pre-forming step, but the
disadvantages can be considerable, both from a product design and cost
perspective. Each of the factors in the previous paragraph, as well as others
are adversely affected. These costs and limitations are often not recognized
or readily accepted without objective consideration of other approaches.
13.8.8 Expansion is free
From the perspective of HPH, where at least a small amount of cross-section
expansion is necessary for the process to work, expanding further for added
rigidity would seem to require nothing more than the elevated pressure
required for the process anyway. From this perspective, larger, structurally
meaningful expanded areas can be done for no additional cost.
However, it is not that simple or clear cut. Additional costs for the HPH
process to work are already incurred, so claiming no extra cost is true, but
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certainly a relativistic point and perhaps misleading. This means extra costs
from a number of different factors are needed to facilitate better forming.
Some are listed above and the extra cost can be substantial. An example is
when higher formability material is needed, the cost will be higher and
material is usually the majority of the part cost. The only question is how
much higher.
13.8.9 High pressure is needed to form
sharp cross-section corners
Using HPH is one way of forming sharp corners, but it has several
drawbacks, including formability limitations. Designers of HPH processes
make it a priority to reduce corner radii to mitigate this effect. The benefits
of this are a proportionate reduction in the forming pressure required;
reduction in the stress concentration that such a feature creates in the
hydroforming tool; and reduction of wall thinning, the likelihood of rupture
and the need for other wall thinning reduction strategies.
Making sharp corners is actually more easily achievable using LPH. The
way the process works allows localized mechanical bending of the metal as
the die closes and since the material is formed compressively, it does not
thin or increase the chance of rupturing it. It also has no effect on the
internal pressure required.
13.8.10 Holes are better cut in the part after forming
This idea seems to have grown from the fact that parts made with HPH
have a relatively small number of holes punched in the hydroforming die,
the rest are cut afterward, commonly with robot mounted lasers. A benefit
is that cutting holes later allow easier adjustment of hole location leading
up to production. Also any engineering changes requiring hole moves are
easy to implement. A downside is that, since holes can be easily moved, a
manufacturer must monitor closely to ensure they do not.
The reasons fewer holes are punched in an HPH die are rooted in the
use of higher pressure fluid. This means larger punch units and cutting large
holes in the die body to mount them. These same holes also introduce stress
concentrations that with pressure cycling can become cracks.
A counterpoint to this logic is found in the stamping industry. It is widely
recognized that, for cost and accuracy reasons, as many holes as possible
are cut during stamping. Designers only deviate when they have to. This
same logic seems applicable to tube hydroforming and is in fact the approach
used with LPH. The use of much lower forming pressures facilitates being
able to make more than 60 holes in one part with no die cracking issues. It
is safe to say that almost as many holes as are likely to be needed can be
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punched in the hydroforming die. Also some holes, the most prominent
being the extruded one, can only be punched in the die. Positioning 30 or
more punch units increases the tool cost, but the cost is similar to an HPH
die. Apparently the cost of punch units compensates for the extensive
number of inserts placed in HPH dies to address wear (not for LPH). The
cost of a system to make many holes after hydroforming can be large
depending on hole count, part size and volumes, among other factors. After
the purchase cost is that of maintaining and operating it.
13.8.11 Parts made with high pressure are lighter
Normal practice for HPH parts is to buy a tube with a diameter slightly
(2–5%) smaller than the minimum periphery of the part to be formed and
the wall thickness required. When the pressure is exerted and the tube is
expanded to fill the cavity, wall thinning occurs where end feeding does not
have any effect. The concept is that a smaller tube fills the space of a larger
one, thus making the part lighter by a similar percentage. Seemingly not
considered in this approach is that usually parts have a minimum material
thickness. When the corner regions get thinner, they would be below the
minimum. This should mean that thicker material would need to be bought
to end up with the minimum in the corners. This would make the HPH
manufactured part heavier.
The approach when using LPH is to buy tube that is the diameter or
periphery and minimum wall thickness required. When the process does
not expand the tubing, the wall thickness stays the same. Also in the scenario above, if the thinner wall in the corners was acceptable, it would seem
to present an opportunity to buy thinner material to start with, still making
the part made with the LPH approach lighter.
13.9
Other comparative process factors
Another avenue of comparison is to look at how the LPH and HPH
processes were applied to make two similar engine cradles. It is a clear and
specific way to contrast differences in these techniques. Since corresponding
data for HF and LIF are not available, this section concentrates on LPH
and HPH. In general, hot forming seems more suited to smaller parts,
judging from the examples so far publicized by companies promoting this
process.
13.9.1 Process flow diagrams
The process steps for LPH and HPH are shown in Figs. 13.11 and 13.12. The
differences of note are a simple preforming fixture vs. a small preforming
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Bending
Hydroforming
and
hole punching
Pre-forming
fixture
Welding
309
End trim
13.11 LPH process operation sequence.
Bending
Pre-forming
press
End
lubrication
Cleaning
End trim
Hydroforming
and
hole punching
Rust
prevention
Drying
Welding
13.12 HPH process operation sequence.
press and the added steps of lubricating, cleaning, rust inhibitor application
and drying. The last three operations were shown separately for clarity, but
they occur in a single piece of equipment. These extra operations, and
dealing with the lubricant residue from cleaning adds cost.
13.9.2 Material and tube specifications
The LPH part uses 300 MPa minimum yield HSLA steel, which was chosen
by the customer based on part function with no process restriction. The
tube is 70 mm in diameter with 2-mm wall thickness. Normal commercial
standards for cut length, cleanliness and seam weld quality are sufficient for
LPH production.
The HPH part uses 240 MPa boron mild steel in a 65-mm-diameter tube
with 2-mm wall thickness. It is apparently dictated by the process and will
cost more than mild steel, but comparison with 300 MPa HSLA is unclear.
More stringent standards of cleanliness, length and weld quality are required
which increases cost.
13.9.3 Bending
The LPH part has nine bends, the greatest of which is ≈140°. The centerline
bend radius is ≈115 mm. This more complex profile was worked out to
accommodate the function of the related components in the most economical way. Two benders are used.
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The HPH part has five bends and the largest is about 100°. The centerline
bend radius is ≈140 mm. The larger bend radius reduces demands on the
material and leaves more formability for expansion during hydroforming.
The simpler bending program means that bending cycle time is substantially
less and also uses two benders to produce this high-volume part.
13.9.4 Pre-forming
The LPH part is placed in a press loading fixture where it is made into an
oval shape through the part legs by using simple blocks sliding perpendicularly to press travel. This is done because the leg width is substantially less
than the starting tube diameter.
The HPH part undergoes a more intense pre-form in a 200-tonne press.
It is done to form the round starting tube cross-section closer to the desired
rectangular shapes of the final part. This is necessary because the periphery
is not expanded in this area and the steel would likely not be contained
within the die cavity, or pinch. Using the LPH technique would make this
type of pre-forming unnecessary.
13.9.5 Hydroforming
Following pre-forming, the HPH tube ends are lubricated with oil before
being loading in the press. Dramatically different forming pressures are
required. The HPH part requires 152 MPa (1500 bar). This pressure over the
surface area of the part is contained by a 3500-tonne hydraulic press, which
requires 630 kW/press to operate. The LPH part is not lubricated and needs
only 48 MPa (475 bar) and a 1100-tonne hydraulic press at 94 kW/press.
The LPH part has 21 holes of varying sizes, the largest being 42 mm in
diameter. The remaining holes are medium and small rounds, as well as
rectangular, which are punched in the hydroforming die with slugs attached.
The HPH part has 22 holes with the largest being a 18 mm × 26 mm slot.
The rest are small round, rectangular, and hexagonal with all slugs attached,
indicating that all were punched in the forming die.
Both parts are then trimmed at an end shearing station and inspected for
hole presence. Additional material is trimmed to accommodate rear mounting bushings. The HPH part requires washing to remove the oil, followed
by application of rust inhibitor and drying to complete the process, while
the LPH part does not.
13.9.6 Wall thickness distribution
Figure 13.13 shows the wall thickness pattern for each of the parts in the
leg area where the HPH part undergoes expansion to avoid pinching as
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2.20
2.10
Thickness
2.00
1.90
LPH
HPH1
HPH2
1.80
1.70
1.60
1.50
1
4
7
10
13
16
19
22
25
28
31
34
13.13 Material thickness distribution for similar LPH and HPH
engine cradles.
described previously. Measurements were taken in a straight section to
avoid the wall thinning and thickening effect of bend areas, which should
give a clear comparison of thinning from hydroforming. They were taken
every 10° for a total 36 readings around the circumference.
Wall thickness for the LPH part is virtually the same as the starting round
tube, since there is no expansion. The section shown has a higher than
normal range of 0.10 mm or up to 5% above the minimum reading.
The two sections on the HPH part vary substantially more (22%) from
highest to lowest readings. The amount of expansion is approximately 5%.
Even though sections 1 and 2 are approximately 40 mm apart there is a
noticeable difference in the thinning pattern. Stretching in two of the crosssectional corners are the reason for the low measurements between 100 and
150° as well as 200 and 250°. The other two corners in these rectangular
sections do not exhibit this pattern, perhaps because the start tube is closer
to that side of the finished section.
Failure or rupture in such sections normally occurs at the tangent points
of the corner radii. It is noteworthy on section 2 at 200 and 250°, and
particularly at 110°, that thickness abruptly drops, suggesting the onset
of necking. Extending into this zone of plastic deformation may reduce
the fatigue life of the product before cracking, depending on stress
application.
Wide variation in wall thickness affects weld quality. Conditions set for
thicker material in flat areas may burn through thinner corners or settings
for corners may cause cold welding in thicker areas. Controlling this is
important since many structural hydroformed parts have other components
welded to them.
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13.9.7 Equipment power and energy consumption
The pressure used for these parts is 48 and 152 MPa for LPH and HPH
parts, respectively. The power and energy required to produce parts is an
important measure of the relative efficiency of the processes and directly
affects cost. The LPH uses an 1100-tonne press whereas the HPH uses a
3500-tonne unit. For the latter, there are two presses, doubling the energy
consumption indicated.
Figure 13.14 shows these factors together to give a clearer comparison.
Energy consumption per part is important and is based on the rated power
of the respective units. These numbers are meaningful relative to each other
but may not represent actual energy consumption. They give an idea of cost
magnitude.
When considering the whole process, the difference shown here is only
part of the story. Energy needed for the HPH equipment that pre-forms,
lubricates, cleans and dries adds to the already large difference. End trim
is similar for both parts and therefore energy consumption is assumed to
be the same.
13.9.8 Cycle time
LPH cycle time was <22 s whereas HPH was 34 s, i.e. >55% longer (Fig.
13.15). These times are for the production of one press for each process.
LPH
LPH
HPH
HPH
0
100 200 300 400 500 600 700
Power (KW)
1
2
3
4
5
6
7
0
Energy consumption/part (KW h/piece)
13.14 Power usage and energy consumption for similar LPH and HPH
engine cradles.
LPH
LPH
HPH
HPH
0
10
20
30
40
0
Cycle time (s)
200 400 600 800 1000 1200
Floor space (m2)
13.15 Cycle time and floor space for similar LPH and HPH
engine cradles.
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The HPH system has two presses for a line production rate of 17 s. This
significantly longer time may be explained by larger presses being inherently slower, as well as the longer time it may take for the HPH sequence
to complete forming, including end feeding. Because of the great similarity
of the parts from the functional perspective, it seems unlikely that this time
difference is driven by some constraint of the part design. This greater cycle
time has made two presses necessary to meet the volume requirements.
13.9.9 Required floor area
The LPH production cell occupies 240 m2, while the HPH equipment needed
to produce a functionally similar part requires 1090 m2 as shown in Fig.
13.15. This is because the equipment required for HPH is not needed (i.e.
no lubrication, cleaning, rust protection application or drying equipment,
second hydroform press or preform press) for the simpler LPH process.
13.10 Conclusions
It is easy to conclude that, because more companies use HPH, it must be
the better approach. Also, when seeking to start making hydroformed parts,
it is what equipment makers offer. However, LPH can also be used for most
hydroformed parts. All the processes discussed are suitable and have made
60 million production parts. The challenge is to identify the most costeffective way to make what you need. It is also worthy of note that processes
and equipment are reasonably interchangeable, meaning that HPH equipment can be used for the LPH process, or elements of each may be combined to gain greater benefit of some of the economies described. However,
doing so with insufficient knowledge can lead to incorrect design choices
and conclusions. Designers are advised to work with those having extensive
knowledge and experience in this area.
It is worth noting the number of total parts produced by LPH in spite of
the few companies doing it. This steady increase in business is telling. Each
process has its strengths and merits. The ability of LPH to make parts that
fulfill most performance and design requirements, accomplish some tasks
that are not reasonable with HPH and its lower cost structure make it a
compelling process to use.
Liquid impact forming is even more cost effective for parts of a suitable
design and size.
Hot forming offers attractive design options for making parts that would
be unreasonable to consider for any of the other processes discussed. It can
combine parts and allows substantial weight and cost reductions for parts
that up to now have had to be made by more expensive methods like
machining. Its greater expense relative to other hydroforming methods is
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justified for this range of parts, since it is cheaper than their current method
of manufacturing. It will also substantially increase future design options as
it finds its place in the market.
It is common to look for one process to replace another but more often
a new process only replaces a subset of the part range that the previous
process was used for. This was the case for hydroforming compared with
stamped and welded assemblies and is true for LIF and HF compared with
LPH and HPH. Also, new forming techniques can extend or expand the
potential market with their abilities, such as has been described with hot
forming. Generally speaking, all the processes discussed are competitive to
some degree and complementary.
When the capability to produce the required part features is objectively
balanced against cost, the right process will become evident. It is a crucial
analysis, since business competitiveness relies on getting it right.
13.11 References
1 Provided Courtesy of Greenville Tool & Die, Greenville, Michigan
Brady Paulson, Greenville Tool & Die, Greenville, Michigan
2 Title: HEATforming: A New Freedom in Forming Tubular Structures
4th Annual North American – Hydroforming Conference & Exhibition – Sept
2006
Harry Singh – Vice President Schuler Hydroforming, Inc., Canton, Michigan
Tube Hydroforming: Efficiency and Effectiveness of Pressure Sequence Hydroforming, IBEC98 Conference and Exposition, Cobo Hall, Detroit, MI, 1998, Gary
Morphy, Van-form
Tube Hydroforming: Dimensional Capability Analysis of a High Volume Automotive structural Component Production Process
SAE International Congress and Exposition, Detroit Michigan, 1998, Gary Morphy,
Van-form
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Fluid cell pressing in the aerospace industry
M. BERGKVIST,
Avure Technologies AB, Sweden
14.1
Introduction
A variant of hydroforming, fluid cell pressing is being used by aerospace
manufacturers such as Boeing, Airbus, Cessna, Bombardier, TAI, Embraer
and others to form a wide variety of sheet metal components with precision.
Although the process was introduced decades ago, recent advances in press
and tooling design have made this high-pressure forming technology more
viable and affordable than ever.
In this chapter the history of the technique and the operational details
of the fluid cell forming process are described. Other topics include:
•
•
•
•
•
Recent developments
Ductile material properties
Suitable part applications
Tooling
Part manufacturing procedures
14.2
Evolution of the technology
With the advent of metal airplanes after World War I, there grew an
urgent need for an economical process to manufacture sheet metal parts.
Mechanical stamping equipment required expensive, multi-piece tooling
which could not be cost justified by the relatively small volume of airframe
parts to be produced. The only other alternative was hand forming, a laborintensive process that resulted in long lead times and inconsistent part
quality.
14.2.1 Rubber pad pressing
In the late 1930s, the rubber pad forming technique was developed and
subsequently adopted by a number of aircraft manufacturers. Instead of
using two rigid tool halves to form a part, the upper tool was replaced with
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a soft rubber pad. Under sufficiently high pressure, the rubber would force
a sheet metal blank to assume the shape of the lower tool surface. This
concept offered the advantages of significantly lower tooling costs and
improved surface finish. The process, however, had two basic limitations:
• Longer forming time than mechanical pressing. Each forming cycle
required pressure buildup, which extended cycle time to several
minutes.
• Insufficient pressure. Standard rubber pad equipment lacked the pressing force to form more complex parts or to achieve uniform pressure
distribution over very high tools. Higher tonnage models could not be
cost justified.
14.2.2 Fluid cell pressing
In the early 1950s, ASEA in Sweden and other companies spearheaded
the development of the fluid cell press. This design replaced the rubber
pad with a thinner, more flexible rubber diaphragm, or ‘bladder’. A strongly
reinforced press frame meant that significantly higher forming pressures
could be attained at a capital cost similar to lower-pressure rubber pad
presses.
14.2.3 Acceptance by aerospace manufacturers
Economically, the fluid cell press made sense to airframe parts producers,
since they dealt with relatively low production volumes and a good deal of
prototyping. The longer cycle times were offset to a great degree by the
ability to form multiple parts in a single cycle.
Beginning in the 1960s, major aerospace original equipment manufacturers OEMs began to incorporate fluid cell pressing into their fabrication and assembly operations. Since then, a growing number of contract
suppliers have also found it economically feasible to invest in this technology. Many companies have retained their rubber pad presses for lowerpressure forming of simple shapes, while employing fluid cell presses for
more complex, critical tolerance components.
Fluid cell pressing has proven to have a number of economic and processing advantages, particularly when compared with mechanical or hydraulic
stamping:
•
Tooling savings. As with rubber pad pressing, only a single, shapedefining punch or die is required, cutting tool costs by 50 to 90% from
the two- or three-piece tool sets used in stamping presses. Tool matching
and alignment procedures are eliminated and, with the uniformly
applied pressure of the diaphragm, stresses in the tool are almost purely
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R
317
8
t
7
100 bar
3600 psi
250 bar
10150 psi
700 bar
14500 psi
1000 bar
20300 psi
1400 bar
29000 psi
2000 bar
6
5
R/t
1450 psi
4
3
Stainless steel
2
Mild steel and
Al (AA 2024, solutionized)
Aluminium (AA 2024, T0)
aluminium (Pure)
1
Parts formed in a
V-shaped cavity due to
different pressure levels
0
500
7,250
1000
14,500
1500
21,700
2000 bar
29,700 psi
Pressure
14.1 Part definition of various materials shown as a function of
pressure and blank thickness under increasing forming pressures.
•
•
•
•
•
•
•
compressional, permitting the use of inexpensive, lower-strength,
easy-to-machine materials such as wood and plastic.
Very high pressures up to 20 000 psi (140 MPa, corresponding to a press
force up to 1000 MN or 100 000 ton) form parts to precise tolerances,
with little or no manual correction (see Fig. 14.1).
Intricate shapes, including undercuts, holes and trimming, can be produced, often in one operation.
Virtually any sheet metal from 0.1 to 16 mm (0.004 to 0.63 in) can be
formed.
Different gauges can be formed on the same tool.
Surface finishes are free from scratches or gall marks.
Lead times are dramatically shortened for both prototypes and production parts.
Complex parts can be designed and produced that would be prohibitively expensive with conventional presses. Parts as long as 4 m can be
formed, allowing designers to combine smaller components into single,
seamless units, saving downstream assembly time and the extra weight
of fasteners.
14.3
How fluid cell pressing works
14.3.1 Press construction
Figure 14.2 illustrates the basic components of a large fluid cell press and
Fig. 14.3 shows side and end cross-sectional views. Shuttling feeding trays
contain the tools and sheet metal blanks. The press frame is a horizontally
positioned cylinder, most typically reinforced with pre-stressed wire winding
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5
6
4
2
8
3
7
1
8
14.2 Basic components of a large fluid cell press: 1 Cylindrical press
frame; 2 Forming tray; 3 Tray station; 4 Automatic tray pad roll-up
device; 5 Hydraulic system; 6 Motor control center/switchgear;
7 Operator control panel; 8 Operator start/emergency stop buttons.
Press frame Pressurized
reinforcement
fluid
Press
Diaphragm Formed part
cylinder
Pressure
fluid inlet
Throw pad
Forming tray
Blank
Tool
14.3 Cross-sectional views of a fluid cell press.
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to safely contain the high hydrostatic pressure. Attached to the frame are
the rubber diaphragm which is pressurized by hydraulic oil, and the wear
pad which serves to protect the diaphragm. Flexible tray pads and throw
pads are used as required to shield the wear pad from heavy local wear and
to enhance forming results.
14.3.2 Pressing cycle
Forming tools with blanks are placed more or less randomly in the feeding
tray; there is no need for fine alignment or fixing in place. The tray is rolled
into the press, and the area above the diaphragm is rapidly filled with oil.
Pressure is increased by high-pressure pumps and the wear pad is forced
downward, forming the metal completely and uniformly around and into
the tool. When the set pressure is reached, the oil is decompressed and
pumped out of the press. The diaphragm and pad return to their rest position, the tray rolls out of the press, formed parts are removed, and new
blanks are loaded for another cycle. Depending on tray capacity and
maximum pressure required, cycles typically run from one to three minutes.
Figure 14.4 depicts the three basic stages of the forming cycle.
14.4
Recent developments
14.4.1 Serving the new supply chain
Cost efficiency is paramount in the globally competitive aerospace industry,
and fluid cell press manufacturers have responded with advancements
in product design and forming technology. Today, a variety of press sizes
and capabilities are available to conform to specific operational and
volume requirements in discrete segments of the airframe parts production
business.
These segments include OEMs along with their tier 1, 2 and 3 suppliers.
As with most companies in industrialized nations, aerospace OEMs are
outsourcing an increasing share of their parts production. Similarly, many
of the original tier 1 suppliers have shifted to more value-added assembly
and integration activities, and are contracting component fabrication to
smaller tier 2 and 3 sources. The broad selection of fluid cell press designs
with tailored functionality creates new competitive opportunities throughout the expanding supply chain.
14.4.2 Tooling technology
In recent years, a great deal has been learned through testing and research
about predicting and controlling the behavior of aerospace aluminum alloys
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a
b
c
14.4 The three stages of fluid cell forming: a blank and tool
placement, b pressurization of the diaphragm, and c depressurization
leaving a formed part. A cutting edge built into the tool trims the part.
and other metals under very high forming pressures. Studies centered on
several materials and a variety of processing characteristics: stretching and
elongation under varying pressures, draw ratios, bending radius, strain hardening, elasticity modulus, etc. The result has been a much greater practical
understanding of the elastic/plastic thresholds and the tensile and yield
strengths of aerospace metals. Among the many findings is the value of
introducing tensile stress during the forming process to better control
stretching uniformity and springback compensation.
This new knowledge has led to significant innovations in fluid cell tooling.
Until now, tool design has been a rather inexact science, often involving the
modification of existing rubber pad tooling. Owners of fluid cell presses
have either built their own form blocks or sought the aid of toolmakers who
were experts in mechanical pressing but had little experience with the fluid
cell process. Through trial and error, a degree of acceptability was achieved,
although manual rework of formed parts was often required.
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New tooling designs now being developed are dramatically changing
the economics of fluid cell parts production (see Figs. 14.5, 14.6 and 14.7).
These tool design innovations result in far more precision in the control of
materials, producing final net size components in one operation with closer
tolerances and little or no manual correction.
14.5 Modified rubber pad tools are seldom suitable for fluid cell
forming. The part exhibits defects such as wrinkling, incorrect flange
lengths, and surface ‘orange peeling’.
14.6 A similar but larger tool than that shown in Fig. 14.5 produced
sufficient circumferential strain to cause a tear at the outer left edge of
the part.
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14.7 A tool designed specifically for fluid cell pressing has reduced
the total processing time for the part shown in Figs. 14.5 and 14.6 by
more than 80%.
14.8 Tail cone attachment ring on empennage for Eclipse 500 very
light jet.
This new precision, easily incorporated into CAD platforms, provides
critical support to the aerospace industry’s ‘part-to-part’ manufacturing
philosophy. Until now, component assembly time has often been excessive
due to misalignment problems caused by failure of the component to meet
key assembly tolerance specifications. In addition, parts are generally made
with pilot holes, requiring re-drilling before attachment. Now, fluid cell
pressing can produce components that fit the first time, with full-size holes,
speeding assembly operations significantly by eliminating time wasted on
manual rework, multiple positioning and re-positioning, and redundant
drilling.
The complex part shown in Fig. 14.8 was originally to be produced by
rubber pad pressing, but early development trials resulted in a great deal
of material gathering, wrinkling, and excessive springback. Many parts had
to be scrapped, and those that were not required substantial rework. Total
manufacturing time was an unacceptable 8 h per part. The problem has been
resolved with a new fluid cell tool (Fig. 14.9) that fully incorporates the new
knowledge of the stress/strain behaviors exhibited by modern aerospace
aluminum alloys. The part is now formed to a tolerance of ± 0.030 in and,
even with a slight amount of final hand dressing, can be produced in less
than 30 min, a 95% reduction in manufacturing time.
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323
Bosses to be removed with magnet
before part is removed from tool
Form tool located on recessed boss
Form tool to be removed
with magnetic lifting chuck
M12 lifting holes
with inserts
Jabroc dam inserted
into base plate
Mild steel upper packer
Mild steel lower packer
14.9 Redesigned fluid cell tool.
These closer tolerance parts can also promote greater use of friction stir
welding, a joining process that is two to three times faster than traditional
riveting, with a subsequent reduction in the weight of the airframe.
14.5
Essentials of ductile materials
14.5.1 Material properties
All metals can be formed to some extent. The ductile properties that are
most relevant to good forming results are:
• Ultimate tensile strength (Rm)
• Yield strength (Rpo2)
• Elongation (Σ)
Table 14.1 lists the values of these and other properties for common
aerospace metals. Forming takes place when materials are stretched beyond
their yield strength and start to permanently deform. The best forming
materials will have a low yield point, high elongation, and high ultimate
tensile strength relative to yield.
14.5.2 Draw ratio
This ratio is defined as the blank diameter divided by the formed diameter
of the part, or D/d (see Fig. 14.10). For most ductile materials, the maximum
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Table 14.1 Properties of common aerospace metals
Material
Rpo2
(N mm2)
Rm
(N mm2)
Elogation
(%)
180
280–350
36–43
180–290
270–410
25
107
Stainless steel
ASTMA 167-81a type 304L
210
490
45
200
Aluminum, heat treatable
alloys (aircraft quality)
AA 2024 O
T3
AA 7075 O
T6
75
345
103
503
185
485
228
572
20
18
17
11
47
120
70
150–185
145
210–250
16
HV 55–77
65
170–215
20
HV 45–60
Pure titanium at
room temperature
400 °C
200
450
220
25
28
Ti 6 Al 4 V at
room temperature
400 °C
850
900
700
10
28
Nickel 7376 (Hastelloy X)
280
700
35
<242
Nickel 7327 (Nimonic 75)
240
650–850
30
166–219
Mild steel, deep drawing
quality
SS 2 1147-32
ASTMA 620-75
DIN ST 14 03
Mild steel
SS 2 1142-32
ASTMA 366-72
DIN ST 12 03 m
Aluminum (quarter hard)
Not heat treatable
SS 4120-22
AA 5052 H32
Aluminum (annealed)
Not heat treatable
SS 4120-0
AA 5052-0
D
d
14.10 Draw ratio of cylindrical cup.
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ratio for pressing straight-sided cylindrical cups in one step is 1.8 to 2.0
(a linear reduction of 45 to 50%). Higher draw ratios can be achieved by
forming in multiple steps. In addition, special deep-drawing fluid cell presses
use a movable punch tool that can produce parts with a ratio of up to 3.0
in a single operation.
14.6
Suitable part applications
14.6.1 Acceptable part geometries
Fluid cell presses can form any part configuration that can be produced
on conventional rigid tools. These include single and compound curves,
contoured parts with stretch or shrink flanges, shallow recesses, undercuts,
and deep-drawn parts. Designs can incorporate forming features such as
internal and external beads, lightening and access holes with or without
flanges, and joggles to provide space for adjoining parts. Simple shapes with
small-radius straight bends are normally produced more economically on
a press brake.
14.6.2 Materials
For aerospace applications, the sheet metals most commonly formed on
the fluid cell press are high-strength aluminum alloys, stainless steel, and
titanium. Other materials can include mild steel and nickel-based alloys
such as Hastelloy®, Inconel®, and Nimonic 75. Thicknesses range from 0.1
to 16 mm (0.004 to 0.63 in). Very thin materials may require special tooling
techniques, while maximum thickness depends very much on the shape and
complexity of the part.
14.6.3 Production quantities
Fluid cell pressing is not, and was never designed to be, a mass production
technique. Smaller fluid cell presses with short cycle times can produce relatively simple parts at the rate of 120 h−1, but speed decreases as parts get
larger and more complex. In most cases, the cost of tooling is the deciding
factor on whether fluid cell or mechanical/hydraulic pressing is more cost
efficient.
Fluid cell form blocks vary in cost depending on production quantity.
For prototyping and very small runs, the tool is made as simple as possible
and part imperfections can be economically hand corrected. For medium
quantities, it usually pays to construct a form block that will minimize hand
work. Multiple duplicates of the tool also improves efficiency. For production of larger quantities of parts, the life of the form block needs to be taken
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into consideration. More durable materials would be necessary to limit tool
wear. It should also be noted that fluid cell pressing has the advantage of
very short set-up times which can be an important consideration with batch
sizes that require frequent tool changes.
14.7
Tools (form blocks)
Single-piece tooling is the primary reason for growing adoption of fluid cell
pressing among aircraft component manufacturers. While the concept of
high-pressure formation is simple and straightforward, the design and construction of form blocks is anything but. As mentioned earlier, a great deal
of research, experimentation and laboratory and in-plant testing has led
to increasingly efficient fluid cell tooling, resulting in a higher-than-ever
percentage of single shot finished parts directly out of the press. This section
can only cover the basics of tooling, but press manufacturers can supply
volumes of design data and technical assistance for all kinds of part-forming
applications.
14.7.1 Tool types
In basic terms, fluid cell tools are either male or female (punch or die).
Many tools incorporate both punch and die forming areas, and are classified
as most appropriate. More specifically, there are four tool types: punch,
punch with blank holder, die and expansion die.
Punch
Also called a hydroblock or male die, the punch (Fig. 14.11) is the simplest,
least expensive, and most widely used type of fluid cell tool, having the same
shape as the internal surface of the pressing. The blank is normally trimmed,
positioned on the tool, and pressed into final shape in one step. Curved
parts with either stretch or shrink flanges are a common application for
punch tools.
14.11 Punch tool with sample parts.
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14.12 Punch with blank holder and sample parts.
14.13 Die (cavity die, female die) with sample parts.
Punch with blank holder
As the blank is drawn over the punch, the outer edge is drawn inward,
and the material inside the forming perimeter is compacted, generating
compressive stresses that may cause wrinkling. This is remedied by placing
a fixed blank holder around the punch (Fig. 14.12). The blank holder
must support the entire unstressed area of the blank and can be made as a
combined unit with the punch or as a separate piece.
Die
The die (or cavity die or female die; Fig. 14.13) consists of a central cavity
surrounded by a flat or curved section upon which the blank rests, thus
eliminating the need for a separate blank holder. The surface of the die
produces the outer surface of the pressed part.
Expansion die
An expansion die (Fig. 14.14) is used when the blank is to be pressed
radially outwards into undercuts in the cavity. In many cases, the blank is
pre-formed to fit more readily into the die. To prevent excessive elongation
of the diaphragm, the space in the blank is filled with soft rubber. The
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14.14 Expansion die with sample parts.
expansion die is made in two or more pieces and is pulled apart after the
pressing cycle to allow removal of the part. The sections of the tool are
typically held together during pressing by the uniform pressure of the diaphragm, but in some cases, special holding fixtures are required.
14.7.2 Tool material
Fluid cell tooling can be made in many materials depending on cost, production volume, blank material and maximum forming pressure. In general,
material hardness determines tool life (number of cycles), and material
strength defines the variety of possible shapes without tool failure at high
pressure. The most commonly used form block materials are the following.
Bismuth and tin eutectic alloys with low melting points (e.g., Cerrotru)
These alloys are expensive but can easily be melted down at 200 °C and
re-used a number of times. Lack of hardness limits their use to prototyping
and short series at low pressures.
Plastics (e.g., polyurethane, epoxy, and other casting resins)
Plastic form blocks are easy to cast and machine, but care must be taken to
eliminate any internal voids which will cause deformation of the tool under
pressure. Plastic tools are generally for low volumes and pressures up to
1000 bar.
Laminated resin-impregnated hardwood
Layered in criss-cross sections <1 mm thick, wood tools can be used at
pressures up 1400 bar. They are not recommended for flanging of thickwalled parts. Parts requiring high surface finish should be positioned
parallel to the lamination.
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Remeltable zinc or aluminum alloys (e.g., Zamac and Kirksite)
Shrinkage-compensated molds are required, which are costly but result
in high casting accuracy. Material cost is relatively low and finished
tools are suitable for medium volumes and forming pressures up to
2000 bar.
High-strength aluminum
Aluminum alloyed with zinc, magnesium or copper combine high strength
with good machinability, making form blocks with long life at pressures up
to 2000 bar. However, these tools may show excessive wear when pressing
certain aluminum parts.
High-strength steel alloys
Hardened or tool steel make excellent tools for high-volume, high-pressure
forming of any blank material up to 16 mm thick.
14.7.3 Tool type selection
General considerations
• For parts requiring quality surface finish and close tolerances, machined
form blocks are preferred over cast tools and the high-surface finish
side of the pressing should normally face the diaphragm. It is even
possible to form painted blanks if the painted surface is facing the
diaphragm.
• Prototype parts can generally be pressed on less expensive tools, where
hand correction is economically tolerable.
• Thick and strong blank materials require high-strength hardened
tools.
• The larger the number of parts to be produced, the more important
is the need for a harder tool surface and a stronger overall tool
design.
Punch applications
Punches are commonly used for many relatively simple shapes, including
those with shallow corners and stretch or shrink flanges. Because they
reduce material thickness less than other tool types, punches are the
preferred tool in deep drawing applications.
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Punch with blank holder applications
The blank holder offers more precise control in punch-formed parts and is
frequently used on visible exterior components where surface finish and
shape accuracy are critical.
Die applications
Dies have a larger free area available for material extension than do
punches, and are thus more suitable for deeper cavity parts with deeper
corners and large bending radii. Dies are typically the choice for complex
shapes.
Expansion die applications
These multi-piece tools can be disassembled in order to remove parts
formed with undercuts. Wherever possible the split lines should be located
at points where marks on the pressing will not be an issue.
14.8
Part manufacture
14.8.1 Tryout
Even with the advent of more precisely designed form blocks, all first-time
pressings must undergo preliminary testing before the production run
begins. It is preferable to perform tryouts on one tool at a time.
Preparation
Basic checks include the availability of necessary blank holding and filler
pieces, rounding off of any sharp edges in the tool outside of the forming
area, air evacuation holes where required, plugs to cover lifting holes in
heavier tools, and cutting of blanks to approximate size.
Tool lubrication
If needed, the surface of the tool is lubricated – generously where sliding
is desirable and sparingly where it is not. Different parts of the tool are
lubricated in different ways. For example, a punch with blank holder should
have low friction at the blank holder but high friction in the punch area.
Grease is commonly used as a lubricant, sometimes in combination with
thin polyethylene film (0.07–0.15 mm) which achieves very low friction
values.
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Blank size and testing
In tools requiring a blank holder, an oversized sheet is normally used.
Pressure should be initially low and increased in stages to assess the flow
of the material. Some parts, particularly those with deep or complex recesses,
may require two forming stages. Rubber filler pieces are normally placed
in the recess in the first stage to limit initial material stretching or excessive
elongation of the diaphragm. Before the second forming stage, the filler
pieces are removed and excess blank material is trimmed off. Any necessary
adjustments in blank shape and tool lubrication are made before the next
pressing, for which a higher initial pressure may be set.
If significant thinning or cracking occurs during the operation, it may
be necessary to reduce the size of the sheet in the blank holder area,
increase lubrication, or reduce the pressure and trim the sheet. In the event
of wrinkling, the reverse of these measures apply.
Extra care is needed to protect the diaphragm during part tryouts.
Damage to blanks before the sheet reaches the bottom of the tool may
cause diaphragm rupture, especially when relatively heavy gauge metal is
used. Precautions include extra layers of wear rubber on top of the blank
in critical areas, increasing pressure in small increments to carefully control
stretching of the sheet, and adding protective rubber after each pressing
stage.
Documentation
It is very important to document all stages of the tryout procedure. A pressing from each stage may be saved, marked with trim lines. Once the entire
manufacturing procedure has been optimized, production may begin.
14.8.2 Trimming and hole cutting during pressing
To reduce the number of secondary operations in the part production
process, form blocks can be designed with integral cutting edges for trimming and hole cutting.
Trimming
In-tool trimming can be economically performed for a production series as
small as 50–100 pieces. The blank is stretched over a hard cutting edge,
which may be countersunk in a softer tool material (see Fig. 14.15). The
amount of forming pressure required for trimming is determined by the
sharpness of the edge and the shape of the cutting area.
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14.15 Dies with trimming edges.
14.16 Hole cutting techniques.
Some part geometries may require intermediate trimming during preforming to prevent breakage. This is typically done at low pressure (50–
200 bar). Final trimming takes place at high pressure after completion of
forming in the cutting area. Tools can be supplied with both intermediate
and final trimming edges, but the distance between them must be sufficient
to prevent the scrap from being formed over the final trimming edge, rather
than being stretched and trimmed off.
Hole cutting
Two basic tooling designs are used for cutting holes (see Fig. 14.16). The
first features a cone below the cutting edge which results in a clean cut and
easy scrap removal. If the blank is subject to sliding over the tool surface
at low pressure, a pre-stressed spring device opens at higher pressure to
make the cut and eject the scrap.
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14.8.3 Series production
The total quantity of each part to be produced has a bearing on production
techniques, tool design, and tray loading procedures. For very short runs
(<25 pieces), the tryout method is adquate. Additional guidelines apply for
larger series production.
Quantities up to 1000 pieces
•
•
•
•
•
•
Blanks should be cut to near final size by laser or, in some cases, by
cutting edges in the form block during the pressing cycle.
The tray is generally loaded with the maximum number of tools possible
to reduce per-unit production costs.
Pressings requiring the same forming pressure are processed in the same
batch.
Spacing of tools varies depending on height and form, but, in all cases,
void areas between and outside tools should be filled with cut rubber
blocks to help ‘level out’ the top of the forming area. This not only
prevents unnecessary stretching of the diaphragm, but also reduces the
volume of oil needed to pressurize the diaphragm, resulting in shorter
cycle times.
Lifting facilities such as overhead cranes with suction cups are required
for placing and removing very large blanks and parts.
Production rates are increased by using multiple duplicates of the form
block for identical parts.
Quantities of more than 1000 pieces
•
•
•
Tools with complete trimming/hole punching capability should be used
to the maximum possible extent.
Special ejectors can simplify the removal of finished parts.
Automated pallet systems using steel plates pre-loaded with tools and
blanks are available to reduce the handling time between pressing
cycles.
14.9
Conclusions
Since 1965, the world’s aircraft industry has purchased some 125 fluid cell
presses, almost all of which are still in operation. Of these, 90% have
forming trays of 1 m × 3 m or larger, with maximum pressures ranging from
1000 to 1400 bar. Most users have had satisfactory forming results at pressures between 690 and 1000 bar.
It should be noted that fluid cell pressing is also gaining acceptance in
the automotive and general sheet metal forming industries. Ford, Volvo,
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Daimler–Chrysler, and several other companies have reported dramatic
savings in time and money in prototyping operations and in lower-volume
parts production for ‘niche’ and specialized commercial vehicles.
Aircraft manufacturing, however, is, and will remain, the largest market
for fluid cell pressing. The technology is constantly evolving as press suppliers, component designers, metallurgical technicians, and plant engineers
seek more cost-efficient ways to produce high-quality sheet metal parts.
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Hydroforming and its role
in lightweighting automobiles
G. MORPHY,
Excella Technologies Inc., Canada
15.1
Introduction
Transportation of people and goods has always been a challenge
facing humanity, progressing through the years from basic walking and
carrying goods to our modern motorized modes that are so ubiquitous that
we take them for granted. For many people, the basics are still what they
have to use most of the time, but everyone wants the ease of the faster
options of using vehicles such as motorcycles, cars, trucks, planes and
others.
The already massive size and level of resource consumption of the
transportation system, face the prospect of ongoing increase, continuation
of the trend that has been evident for many years. Recently, this trend has
accelerated because of the huge number of people in the developing world,
most notably India and China (37% of total world population) with a pentup demand for modern conveniences and transportation. Reducing this
need/want seems unlikely. It is, therefore, to investigate how best and most
efficiently to make our transportation system consume fewer resources.
A key part of the strategy is, and must continue to be, making vehicles more
efficient, thereby consuming fewer resources. This must include designing
and devising ways of manufacturing, maintainance and operation, which,
while attaining a reasonably long working life and efficient means of
disposal, minimize the required effort and cost at each stage.
The focus of attention here is on the manufacturing of automobile parts
and how it affects the operating characteristics of such vehicles. The most
important of these characteristics are:
• Bending and torsional rigidity;
• Crash energy absorption;
• Fuel economy;
• Potential to reduce cost – less material (≥60% total cost) vs. possible
higher process cost.
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It has been recognized since people started using vehicles that a significant part of the energy required to propel a vehicle is due to its weight. This
is the sum of the weight of the vehicle itself and what it is carrying. The first
would be all parts of the vehicle, while the latter is people, luggage or any
other cargo. The goal should be to minimize the first and maximize what
the vehicle is capable of carrying safely. The challenge is how to do so using
the most effective and economical methods.
Vehicle weight is composed of the structural elements that hold the
vehicle together and allow it to perform the functions required of it, as well
as the weight of everything else. The focus of this chapter is on those that
play a structural role including those that are hydroformed today and others
that are not, but where it would be beneficial. Non-structural parts, such as
exhaust components (some are hydroformed) are part of the opportunity
for weight reduction.
Lightweighting is a general term used to describe a number of techniques
aimed at reducing weight. Generally it is the activity of producing lightweight components for improved performance. Extremely clever and
weight-efficient (high strength/weight ratio) examples can be found in
nature such as structures found in bones including notably the human spine.
Its ability to move and bear large loads relative to its small structural weight
is remarkable.
Lightweighting is a theme that has been prevalent in the automotive
industry for many decades and yet the commonly held perception is that a
lot more remains to be done. Two reasons for this are the addition of equipment and features and a relaxing of the weight-reduction discipline. This
leads to the need for a periodic renewal of the discipline of shedding mass
by one or more of a number of different strategies. A third reason is the
continual development of technology that provides new options that are
difficult to synthesize with other elements like function requirements,
material, design practices and meshing it with current design. This is a
complex subject, which is difficult to manage.
15.2
What makes it so difficult to lose weight?
This situation seems similar to personal weight loss in several respects. One
is that there always seems to be a few more pounds to lose and it is easy to
regain what was previously shed. Vigilance to prevent or minimize future
gains and recognizing new reduction opportunities are important, but challenging to maintain.
Engineers and designers are not willfully adding weight to the vehicle.
For automobiles, there are several factors that make the goal of designing
the lightest weight structure to perform the required functions a moving
target. These are:
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•
•
•
•
•
•
337
Changing requirements for vehicle function and expectations of
consumers;
Expansion of the capabilities of current forming or assembly
technologies;
Development of new technologies;
Materials;
Factors of safety (or uncertainty) that are built into the design;
How well informed were many of the design choices made, considering
how many options;
The potentially synergistic effect of losing weight in some parts has
then to be taken into account in the design of the parts that must
carry it.
Each of these categories has more detailed areas that will be discussed with
a focus on how they relate to hydroforming.
15.3
How to lose weight
Lightweighting means different things to different companies, who assign
them different levels of value. It seems reasonable to view lightweighting
as a collection of strategies that make sense to a particular company or
industry. They should also be open to new strategies or technology that may
offer additional advantages. These strategies should be coordinated to
achieve the lightest overall part or assembly. It is often not highlighted, but
is an important point to make, that part weight reduction is conditional on
the parts functional requirements still being met.
Manufacturing options should be explored for their weight-reduction
potential. Relative cost should not be a concern initially, because lightweighting and cost can be at cross purposes. Too often, cost is the first and
by far the largest consideration. However, if the weight reduction benefits
are quantified first, more rational decisions can be made concerning benefit
vs. cost, particularly if the total includes some components that require the
combined effect of two or more of these factors. They include:
1.
2.
3.
4.
5.
6.
Eliminating joints.
Eliminating welding.
Eliminate material added to make joints.
Improving joint design for rigidity.
Secondary stress analysis – deflection, stress reassessment.
Using high strength steel or lightweight materials – aluminum, magnesium, plastics.
7. Maximizing section size; minimize wall thickness.
8. Cross-section shaping to minimize material use – expansion, wall
thinning, loads.
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9. Wall thickness control – tailor welded or tailor rolled blanks; even
more during hydroforming.
10. Reconsidering options where different processes may reduce
weight.
11. Selecting the most effective processing steps – i.e. LPH, HPH, HDF
or a combination.
12. Effectively coordinating use of the applicable strategies to maximize
the benefits.
13. Holistic design strategy – The whole structure should be considered.
Design can be handicapped by several factors. Often the focus is on
many separate efforts focused on different parts and assemblies, while
the whole structure is given less attention. Jumping to conclusions with
insufficient information as a guide.
14. Reconsidering the whole structural system: make two parts into one
to reduce part count; part function multitasking; overview – ensure
right technology for each part; allow structure to be effectively filled
out; and reassess design choices to avoid structural degradation.
Once the potential weight loss ideas are understood, cost can be considered to decide how to most effectively implement them. The decision may
be that the cost is unjustified for some components, but it will be more
informed and made for better reasons.
15.4
How tube hydroforming can help you lose it
Hydroforming is a relatively new technology that has played and will
continue to play an important role in creating lightweight, efficient and
effective parts for vehicles. As such, its potential contribution tends to
increase as further capabilities are developed. The potential for further
development is greater due to its 15–20 year history vs. 100+ years for
stampings.
Tube and sheet hydroforming are two different technologies that can
both play roles in lightweighting. They have a name in common and both
use fluid to function, but differ substantially in most other respects. This
chapter will focus on tubular applications. Additionally, hydroforming tends
to focus most commonly on structural parts, but plays an important role in
exhaust systems and secondary structures, as well as other parts. Sheet
hydroforming can also be used to reduce weight compared to conventional
stamping, but the magnitude tends to be less.
Hydroforming ideally should provide a way to make parts that gain
some advantage from being tubular while retaining favourable features,
capabilities and economics compared to other techniques like stamped and
welded assemblies. In other words it should do for curved sheet metal what
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stamping or roll forming does for flat. So far, the way hydroforming has
been applied is too often focused on making the process work, rather than
achieving as many part features and minimized weight as are needed to
maximize value. Hydroforming capability still needs to improve. Such
advancements are being developed in order to increase the benefits hydroforming can provide for the broadest range of parts and to play a larger
role in lightweighting.
Of the points listed in 15.3, many are directly applicable to tube hydroforming. Some more explanation will to clarify how they can reduce
weight.
15.4.1 Eliminating joints, welding, reducing joint material
The greatest loss of rigidity in an assembled structure is at joints between
parts. While having no joints is preferred, this is impractical in most designs.
Hydroforming facilitates the goal of fewer joints while also satisfying part
requirements such as mounting and attachment points, rigidity, strength and
surface locations. Meeting these objectives determines the complexity of
the part design as designers strive to achieve them as simply and economically as possible.
Joints can be mechanical (bolts, rivets, self threading screws, etc.) for
overall assembly or logistical reasons or for when they need to be disassembled and reassembled during the life of the vehicle. This type tends
to be more flexible, thus requiring attention to minimize movement.
Welded joints are also common; a number of welding processes are used
for permanent attachment. Rigidity of these joints varies and can be found
wanting as well, although they are generally better than mechanical joints.
Characteristics of these are point loading or stress concentrations where
the weld is. Also, because of the heat generated to create fusion, the material is locally annealed. This area of reduced material strength through
which loads transferred between joined parts are passed can also be the site
of other discontinuities resulting from welding such as where (or near
where) burnthroughs occur. They are a problem that is best dealt with using
good design and welding practices with effective quality control.
15.4.2 Improve joint design for rigidity
Where joints are necessary or unavoidable, the best that can be done is to
apply good design practices to maximize their rigidity. This can include
strategies such as positioning welds or mechanical fasteners in shear rather
than bending, and accurate modeling of the joint to reduce areas of high
stress because of inherent point loading.
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15.4.3 Secondary stress analysis – deflection,
thickness and stress reassessment
When the design has been mostly settled, it is expedient to revisit loading
and material behavior to identify further opportunities for weight reduction
by any of the identified techniques. This allows a reassessment of appropriate wall thickness, tube size and local expansion to handle the predictable
stresses and deflections. Potential weight reduction can be surprisingly
high.
This reassessment should be done with the overriding question of ‘What
do we really need?’ with respect to rigidity, material strength, minimum
thickness and other properties of interest. When parts are being modeled
for inclusion in FEA studies, they should ideally have thickness and residual
stress data that represent what they really are throughout the part. This
allows better insight into obtaining sufficient structural strength to perform
the tasks required while not inadvertently overspecifying the amount or
strength of material needed.
This is a key concept since it is common practice to specify a minimum
material thickness for structural parts. This actually functions as a preliminary minimum since forming processes often change it (usually thinning)
locally. These sites are often where potential problems occur. For hydroforming there are three common sources of material thickness change.
They are:
• Bending – decreases thickness on the outside of the bend and increases
it on the inside;
• Cross-section expansion for structural advantage – a complex situation;
thinning is the common effect, can be quite uneven, several measures
(i.e. end feeding) reduce or prevent thinning;
• Cross-section expansion to prevent pinching, which is necessary for the
HPH process.
It is common practice in the automotive industries use of finite element
analysis (FEA) to use the thinnest wall, since detailed thickness and stress
levels over the whole part surface are often not available and difficult to
use. Approximations are used instead, which are assumed not to conceal
‘hot spots’. It is also beneficial to model the part or assembly at the low and
high side of the wall thickness tolerance to assess differences in weight and
performance. From this, it can be judged if it is worth the extra cost to use
tighter tolerance material.
The benefits of intentionally varying the wall thickness to put more
where it is needed and less where it is not are crucial to assess, for weight
reduction opportunities. Options for this are discussed in more detail in the
section below on wall thickness control. In general, ways to control wall
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thickness in specific areas of hydroformed parts have been limited. Often
changes in thickness were a by-product of the process. Clever marketing
has portray such circumstances in a way that made these changes into an
advantage. This has been the case with hydroforming and the idea of making
high-strength steel from low-carbon grades by work hardening (see Section
13.8.2 for a more detailed description).
15.4.4 High-strength steel, aluminum,
magnesium and plastics
These materials are being used more commonly to reduce weight, while
performing the required structural tasks. It should be noted that the
progression of technology and a steady increase in vehicle benchmark
standards has led to the need to improve existing materials or introduce
new ones. For instance, improving crash performance would make it
necessary to:
•
•
•
Increase the amount of material;
Form cross-section shapes that are most beneficial;
Use high strength steel to allow elastic bending further to absorb more
energy during plastic deformation.
The first option is the least desirable due to the weight increase. The other
two steps are preferred and hydroforming improves the range of options
for advantageous shapes while reducing the number of joints. High-strength
steel (HSS) is a challenge to form, but its use is facilitated by use of LPH.
Material selection is a very important aspect of design flexibility when striving to fulfill part functionality requirements including energy absorption
during a crash situation. Maximizing the range of choices is fundamental to
making the most effective, efficient part.
Automotive structural parts are mainly composed of steel and aluminum,
with various alloys of each serving a number of purposes.
The LPH technique, described in Chapters 12 and 13 of cross-section
reshaping by bending widens the range of feasible materials for prototype
and production. This is useful for reasons such as:
• Using inexpensive material where the properties are relatively widely
variable since the process is less affected by it;
• Forming high-strength or stainless-steel alloys since formability is low
or even if it is higher but the work hardening rate is high;
• Forming aluminum where formability limits possible stretching.
The LPH process is able to form complex shapes because of these
low process demands on the material. In all production applications, the
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customer selects the material. Commonly, mild steel is chosen because it
provides the greatest rigidity for the least cost.
A challenge of material development is retaining formability while improving yield strength. Meeting this target has led to the development of
DP and TRIP steels. A cautionary note is that, even though they have
impressively high elongation numbers, the work hardening rate is much
higher than mild steel and therefore must be formed differently or less,
particularly by stretching. Compressive forces generated by LPH forming
facilitate using these materials with a broader range of design features.
Aluminum is another material that generates a lot of interest for structural applications. Its most recognized asset is weight reduction because of
its low density, but it does not ‘like’ to be stretched. Since LPH does not
involve stretching, the design options are more open than with other
approaches. Its weight reduction benefit is reduced by its rigidity being
directly proportional to its density. Most automotive structural parts are
designed for rigidity, so that material may need to be added but hydroforming may prevent it and increase a net weight reduction.
There are numerous examples of high strength, low formability materials
being successfully formed in prototype or production.
15.4.5 Maximize section size, minimize wall thickness
Table 15.1 provides some information that may guide the designer to making
the most effective choice of tube diameter and wall thickness for a given
Table 15.1 Comparing tube weight of differing sizes for a given load
Tube cross-section
Column 1
Column 2
Column 3
Column 4
OD – in. (mm)
Wall thickness – in
(mm)
OD/Wall thickness
ratio
Mass – lb/ft (kg/m)
Weight reduction – %
Allowable deflection
Max. stress due to
load
2.5 (63.5)
0.079 (2.0)
2.25 (57.2)
0.079 (2.0 mm)
2.5 (63.5)
0.056 (1.42)
2.75″ (70.0)
0.041 (1.03)
31.6
28.5
44.6
67.1
2.05 (3.05)
1.84 (2.73)
10.2
At limit
<Limit
1.47 (2.18)
28.7
At limit
<Limit
1.19 (1.77)
42.0
At limit
At limit
<Limit
<Limit
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product. In general, occupying the same cross-sectional space and reducing
the wall thickness to meet the criteria required for the application at hand
makes a more efficient structure.
To show this, column 1 is assumed to form the same cross-section shape
and wall thickness as the part being replaced. Reassessing the stress and
deflection levels provides a secondary weight saving opportunity that may
exceed the first if handled properly. It seems to be a common conception
that minimizing the cross-sectional periphery throughout the part length is
the best way to make the most efficient part. While the concept is true in
some circumstances, often it is not the most effective path to minimizing
weight. Column 2 shows a tube that was reduced 10% in periphery and
weight, where deflection reaches its maximum. In contrast, if the original
diameter is retained and wall thickness is reduced as in column 3 to make
deflection equal the limit, the resulting weight loss is 28.5%. If a little larger
diameter tube (10%) could fit the available space, weight reduction would
rise to 42% as shown in column 4. Further benefit can be gained by reshaping the part for more rigidity in the direction of application of the force.
Other restrictions, such as being able to make good quality tubing,
stable welding of assemblies or sufficient fastener thread strength, that may
prevent full advantage being taken of these potential weight savings are
discussed in 15.5. When this occurs, it makes sense to reduce tube size, but
as can be seen in column 2, weight reduction is smaller. In some instances,
the strength of the part may be too high for best occupant protection. Then
the best method is still using thinner wall or smaller tube, but crash initiators or holes can be designed to create the crash behavior desired.
15.4.6 Cross-section shaping to minimize material use
This is one of the main tools that tube hydroforming can offer a designer
in the battle to reduce weight with a range of options that can provide a
high degree of benefit. Recent developments discussed in 15.4.7 dramatically increase design flexibility and weight reduction opportunities in a
cost-effective manner.
Most automotive structural components are designed to deliver required
rigidity. This means that lightweighting opportunities can be generally categorized as:
• Expansion/cross-section size;
• Wall thickness control;
• Cross section reshaping options;
• Material properties;
• Variability of section strength requirements can mean more is needed
in one place rather than another and this adds unnecessary weight.
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Space constraints and high-strength material use complicates the options
even further.
Expansion and cross-section size
There are three fundamental reasons to expand the cross-sectional periphery of a hydroformed component.
•
•
•
HPH normally expands the tubular blank to prevent squeezing part of
the tube between the die blocks or ‘pinching’. This process is explained
further in Chapter 13.
Greater section strength is another common reason and variable periphery part designs in some cases strive to provide larger sections where
applied loads are higher and thereby improve material use efficiency,
thus facilitating the goal of lightweighting.
Mounting conditions for mating parts.
Before assuming expansion is required, it is valuable to determine if
reshaping the cross-section may be adequate to address locally higher structural requirements. The depth of the part can be much thicker in the plane
of force application in the middle than at the ends. A proportionate decrease
in the perpendicular direction means no expansion, but a much stiffer
section where needed. This is an example where being able to form more
complex surfaces can be a big asset, facilitating a different solution that
is simpler to achieve. The flexibility to change cross-section shape quickly
along the part length is greater than many expect when using LPH
(Fig. 12.4).
Expansion in the hydroforming die (common in HPH parts) is considered
by some to be ‘free’ but the correctness of such an assessment depends on
the assessor’s viewpoint. When using HPH, it seems free because equipment
etc. has to be put in place for pinching avoidance, but compared with LPH
it has a higher cost.
Wall thinning
The thickness of the tube wall is crucial to structural integrity and component performance in a crash. Automotive manufacturers routinely simulate
part performance and crash scenarios on computer systems, and this is
becoming more important all the time.
Wall thickness consistency in different areas of the same part and knowing
how it varies is very important. Normally, a structural part will be bent and
possibly preformed before hydroforming. Bending thins material on the
outside of the bend and thickens it on the inside. The magnitude depends
on the severity of bending and the bending technique used. Some types of
preforming do not change the wall while others do.
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Wall thickness variation around the cross-section is inherent with expanding the tube in the hydroforming die. The corners get thinner where end
feeding is not effective because the material is stretched to form the larger
periphery of the finished part. A newly developed technique can prevent
this thinning effect in many more cases.
All of these variations must be established to improve the accuracy of
FEA work. The minimum values must be used for analysis so that in the
worst case the part meets performance requirements. Failure to do so
results in real part performance that is less than what FEA predicted. This
discrepancy may not cause a problem, but it could be sufficiently serious
for the part to fail. The most serious source of thinning seems to be what
occurs in the hydroforming die when end feeding is not effective, presumably because of its effect on the cross-section corners, which can lead to
cracking after time in service.
Various measures have been developed to reduce wall thinning during
hydroforming, such as end feeding, lubrication, special materials, annealing
and preforming. These are effective depending on the part design and
manufacturing details, but rarely is thinning eliminated throughout the part.
The fact that the thinning effect concentrates in the corners as shown in
Fig. 13.13 increases the importance. Cross-sectional corners are normally
considered to carry a disproportionately large part of the total load compared with flat areas, and give more resistance to deformation in a crash.
15.4.7 Wall thickness control
For a number of years there have been several opportunities to exert some
level of control over the wall thickness of a tube to facilitate lightweighting.
It is important to use these strategies to their fullest extent since they hold
the potential for the largest weight reduction.
Examples are:
•
•
•
Cross-section expansion
䊊
With end feeding – effective near the part end
䊊
Without end feeding – material thinning is uneven and proportional
to expansion
Tailor welded tube
Tailor rolled tube
It has been widely understood in the hydroforming industry that expansion is done in the hydroforming die with the effects mentioned above and
in Chapter 13. Tailor welded or rolled tubes are quite easy to hydroform.
Any challenges are usually due to bending with a mandrel, since the wall
thickness variation will affect the inside surface.
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Recently, additional techniques have been developed that allow substantially greater design flexibility to make parts that were unfeasible or uneconomical and reduce the cost of making most hydroformed parts. They and
some of their benefits are:
1.
More even thickness reduction (where reduction is desired):
a Concentrated wall thinning prevented, as cracking and rupture;
b Lower cost process and material requirements.
2. Larger expansions where needed including:
a More expansion for local section strength or joining to other
parts;
b Substantial expansion in the center of long parts like frame rails.
3. More control of wall thickness in expanded areas than previously
possible:
a Maintain original material thickness;
b Allow material to thin evenly in proportion to expansion to reduce
weight;
c Thickness between these high and low values to balance rigidity,
weight and function;
d Place thicker material at bottom and or top of a part for a more
efficient structure.
4. Local formations (bulges) with controlled wall thinning.
These new technologies offer the opportunity to carve unneeded localized
quantities of material out of parts. These will find applications that make
economic sense as the weight reduction principles are applied. This will take
some effort, since designers need to ensure structural adequacy is maintained. Additionally, future developments will further enhance the ability
to put the right amount of metal where it is needed, in an economically
acceptable manner.
A normal FEA model of a part under stress in the vehicle will show
a pattern of large low stress areas (usually blue), high small stress areas
(usually red) and other stress regions (and colours) in between. It is
normal to concentrate on the red areas to alter design to reduce or
eliminate these ‘hot spots’. Unfortunately this may include adding more
material (often throughout the part), due to a lack of alternatives. It is a big
advantage to add material only where it is needed or reduce the stress
concentration.
The blue areas are in fact understressed and offer the opportunity to
remove material and reduce weight. The stress level in this area will increase.
Following this approach should aim for the goal of making the in-service
stress in the part equal throughout at a chosen safe level. In this manner
all material is equally utilized. Coordinating this with economical processing methods is crucial to advanced weight reduction efforts. This may
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seem like a lot of trouble, but the easy reductions are already done and
like personal weight loss, the next pound is usually tougher to lose than
the last.
15.4.8 Reconsider processing options for parts
where weight may be reduced
A number of parts on vehicles today that are made by different methods
would benefit in one or more ways from being hydroformed. These could
include any of the advantages highlighted in Fig. 13.3. These have either not
been considered in the past or were not changed to hydroforming due to
forming limitations, feasibility or cost. The latter is often the deciding
factor.
The benefits provided by LPH, combined with the recent developments
in the last section, increase the range of product features that can be provided at reasonable cost. Some of these are much lower equipment cost,
lower cycle times and other part cost factors, as well as a more developed
process sequence. In summary, parts that may be lighter, easier to make,
improve functionality or reduce cost should be assessed by those with the
best hydroforming expertise to maximize the benefits.
15.4.9 Hydroforming techniques – different
approaches, different results
As discussed in Chapter 13, different processing alternatives offer a number
of benefits and drawbacks. Two hydroforming methods (LPH and HPH)
dominate the manufacture of automotive structural hydroformed parts.
It is important to distinguish them and recognize which part design
features and processing elements are most advantageous because of their
different effects on wall thickness, material yield strength (YS) and part
design flexibility, which affects how parts can be made more suitable for
lightweighting.
What is necessary to get the best advantage from this technology is
an effectively devised strategy of choosing the process best suited to
deliver minimized weight at a reasonable cost; a probable goal of designers.
To be clear, this does not mean choosing one or another of the alternatives
described in Chapter 13. Instead, it means choosing the elements that
are useful to achieve the desired benefits of minimal weight, maximum
performance, adequate quality, and acceptable cost leading to the right
process.
Often in the automotive industry, cost is treated as the only decisive
criteria; this seems likely to be detrimental to other criteria such as weight
reduction that costs only slightly more.
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15.4.10 Is rigidity or strength more important?
The answer whether rigidity or strength is more important depends on the
design purpose. Taking the right approach can make a difference to what
the part weight ends up being. Generally, for automotive structural parts,
rigidity is the most important factor.
Rigidity and strength are commonly confused with each other, but are
distinctly different. Rigidity of a part is affected by the size and shape of its
cross-section including wall thickness, as well as the material’s Young’s
modulus (modulus of rigidity). All the steel alloys developed over many
decades have essentially the same rigidity, with the exception of stainless
steels, which have a lower modulus because other elements, such as
chromium, nickel and molybdenum make up such a large percentage of
the alloy. It does not change during deformation.
Strength indicated by yield or ultimate tensile strength is a characteristic
that differs substantially from one alloy to another. It can be changed by
deformation or work hardening and heat treatment. High-strength steel will
bend the same amount under a given load as does low-strength steel. The
only difference is that the former will bend further elastically before it
transitions to plastic deformation and stays bent. A well-known example of
HSS are springs. Springback increases as material strength does.
15.4.11 Material strength added by processing
Processing can increase material strength in at least two ways:
•
•
Cold working is the most associated with hydroforming since it is
commonly referred to that expanding the tube elevates yield and tensile
strength;
Heat treatment, such as can be used for steel with added boron.
These are attractive options because the cost of the first is relatively low
if the amount of expansion is not too high and the method economical. The
gains are inherently localized to where deformation happens and in proportion to it. When expansion is too much, cracking and rupture results. The
gains from heat treatment can be dramatic; triple or quadruple the starting
yield strength, but cost is higher and its balance with the benefits must be
examined.
15.5
Weight loss limitations and how to address them
There are a number of process and part performance factors that have been
viewed by designers as limitations of the hydroforming process and furthermore of structural design in general. It now seems apparent that rather than
assume the limits are irrefutable, there are strategies to avoid or at least
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reduce the potential limitations imposed by some of these factors. It is best
to address these in a positive fashion by asking how it could be done rather
than looking for why it can not be done. The following list is not exhaustive,
but provides a number of examples.
15.5.1 Welding burnthroughs
The material must have a minimum thickness, below which burnthrough
defects would occur creating an unacceptably high scrap rate. Limiting wall
thinness reduces the tendency to burnthrough. Improved welding techniques are reducing the thickness needed to avoid this problem.
15.5.2 Weld annealing effect
This is a closely related effect that is part of the nature of welding since the
heat used is sufficiently high to locally anneal the material. However, introducing less heat can also reduce this effect (laser welding), as well as using
a material that retains higher yield strength.
15.5.3 Attachment point stress concentrations
It is common that transfer of stress from one part to another through a
mechanical or welded joint with their relatively small joint cross-section
leads to high local loads. The coincidence of this high stress level with local
annealing means that material must be thick enough to prevent yield in an
annealed state.
Another option is using material whose annealed yield strength is higher,
thus allowing significant thickness reduction.
15.5.4 Insufficient stripout torque for self-threading holes
When punching holes in the part during the hydroforming operation for
use with self-threading fasteners, a key characteristic is the material thickness. Other factors such as hole diameter and material strength have a
smaller effect.
While this is a consideration, there may be options such as using a lower
tightening torque or a different male fastener, or adding a female fastener
to the tube. It is a matter of judging tradeoffs and picking the option with
lower weight for a reasonable cost.
15.5.5 Oil-canning effect
Attachment in the middle of a flat surface applies force to bend the wall.
Thicker walled material is one way to reduce this effect, but not an efficient
one. Several alternatives are:
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• Forming opposing walls together for the strength of two material
thicknesses;
• A connecting element to allow the load to be carried by two walls of a
tube spaced apart;
• To locate the mounting hole closer to a cross-section corner.
15.5.6 Tube-making feasibility
This is a big consideration, because, without good-quality, consistent tubing,
the number of problems that follow can be considerable. A high diameterto-wall thickness ratio (D/t) is highly desirable from a structural design
perspective, providing a higher strength-to-weight ratio. A number of issues
can arise that adversely affect tube weld seam quality. It is becoming more
common that high-strength steel tube is also a requirement, which presents
greater challenges as the strength increases. Combining both can make the
tube-making process unsustainable. This must be addressed with a good
tube maker.
15.5.7 Denting and other incidental damage
This is chiefly a material-handling problem, which happens more as
the material gets thinner. Use of high-strength material is a partial
solution since it takes a larger force. Even a subtle dent in the tube can
make it impossible to get the tube on the bending mandrel. This normally
means the tube is scrapped. Hydroforming will reduce an early dent. If
the damage happens after bending, it is unlikely to cause problems
with further processing. However, denting at any point in the process can
also be a cosmetic or visual concern and cause it to be scrapped at the
customer’s site. Commonly, a wall thickness of 1.5 mm or more has
fewer problems, but 1 mm or less is quite prone to denting. More careful
handling is a preventative measure, but in many industrial settings this may
prove difficult.
15.5.8 Wall thinning
In most hydroformed structural parts there are several reasons for wall
thinning. These are bending, cross-section expansion and hydroform wall
thinning. The amount can be quite substantial and the pattern differs for
each of these causes. If the final wall thickness is less than the specified
minimum, it may be necessary to start with a thicker wall. Finite element
analysis and process simulation can be used to calculate the required starting thickness.
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Conclusions
Hydroforming offers the potential to reduce weight in a number of ways.
Performance and cost improvements are possible depending on the specifics of the situation and how they are handled. It also offers significant
weight-reduction opportunities. These can range from those with no cost
impact to those which entail a substantial increase. Designing and making
the most efficient part is not automatic, easy or obvious. As with any other
technology, there are many ways to misapply it and few ways to do it optimally. The best way to ensure proper application is to learn as much as
possible about different methods to allow logical judgment of the merits of
each approach and/or work with experts in the field to decrease learning
time and increase confidence.
Important principles to bear in mind when designing structures are:
1. Design for structure effectivity first.
2. Make structure design firmer; limit compromises that weaken or add
compensating material. Even though hydroforming can easily accommodate such design features, it is not a structurally good idea.
3. Reducing the number of joints is very important and can be worth some
extra cost.
4. Use the most advantageous cross-section shapes for rigidity or stress
levels where best suited.
5. Some compromises will be necessary to fit with the rest of the
structure.
Losing weight requires a tough, ongoing effort from which relaxing often
leads to a relapse. A disciplined approach combined with determination,
the most knowledgeable and innovative hydroforming and structural design
engineers are the keys to success. Like personal weight loss it is not easy,
but it is necessary.
15.7
References
morphy g., Product Design for Hydroformed Structural Components, TPA 4th
Annual Automotive Tube Conference Focus on Hydroforming, Dearborn
Michigan April, 1999, Gary Morphy, Vari-form.
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Warm hydroforming of lightweight materials
M. KOÇ,
Virginia Commonwealth University, USA
16.1
Introduction: motivation for lightweight vehicles
Lightweight materials such as aluminum, magnesium, advanced high
strength steels (AHSS) and composites have been considered and evaluated to replace automotive components made of low carbon steel to reduce
fuel consumption and hazardous emissions in the transportation vehicles.
The transportation industry is the largest and the most resource-intensive
manufacturing enterprise in the world comprising production of cars, trucks,
aircrafts, locomotives, wagons, and ships. This industry’s approach to protecting the environment and still being able to offer affordable products is
quite diversified including the development of technological solutions to
reduce the environmentally hazardous emissions/wastes during the entire
life cycle of a vehicle. The technological solutions and innovations for environmentally friendly products and for environmentally benign manufacturing can be categorized in three stages (1) design / manufacturing stage, (2)
utilization (actual use of the product) stage, and (3) end-of-life (EOL) stage.
Since approximately 80% of the total energy consumption throughout the
life cycle of an automobile occurs during the utilization period, the use of
lightweight structural parts is accepted as a prominent and long-term solution (Mildenberger, 2000). Even in automobiles with efficient and clean
power generation systems and alternative fuels (such as fuel cells), the
lightweight structures would further increase the fuel consumption efficiency and reduce emissions (i.e., primary and secondary benefits).
16.2
Lightweight materials: advantages and
disadvantages in manufacturing
Lightweight structures can be realized by using lightweight materials such
as aluminum, magnesium, high strength steel, titanium, and composites, and
developing enabling, low-cost and robust manufacturing processes that can
transform these materials into the desired complex functional shapes. For
352
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Warm hydroforming of lightweight materials
353
instance, a viable and cost-effective manufacturing process that leads to thin
wall castings would be an enabling agent to use aluminum alloys for engine
blocks. Similarly, a manufacturing process that produces complex-shaped
parts (i.e. high formability) would enable the use of lightweight materials
to achieve further weight reductions in a vehicle.
Currently, the use of lightweight materials, such as aluminum, magnesium,
and AHSS, is limited to cast and forged underbody parts. Thus, automotive
manufacturers and their suppliers are facing an opportunity and challenge
to further reduce the total weight of the vehicles by using lightweight materials for body parts, too. Schultz presented that 10% reduction of body
weight could improve the fuel efficiency by 6–8% (Schultz, 1999). The mass
reduction by replacing steel with aluminum can be as high as 40–60% and,
with magnesium, it can go up to 60–75%. However, the cost of a part is
expected to increase because of the raw material and production costs. In
the case of aluminum alloys, the cost increases by 30–100% whereas, for
magnesium, it is a 50–150% increase. On the other hand: since 80% of the
total energy consumption throughout the life cycle of an automobile occurs
during the utilization (driving) period, the use of lightweight parts is still
seen as a prominent, long-term and cost-effective solution (Mildenberger,
2000). If we consider the fuel savings during the first 10 years of operation
of a vehicle, assuming that there would be continuous cost reductions
through development of new forming processes and increased use of aluminum and magnesium alloys, competitive and affordable realization of the
lightweight auto body and structures can be expected. Thus, alternative,
novel and cost-effective forming processes are the key issues to be addressed
properly since such lightweight materials have low formability degrees
particularly at room temperature conditions.
However, formability of lightweight materials at room temperature is
much lower than that of mild steel (two-thirds of drawing quality steel for
aluminum alloys). They are more prone to forming defects such as wrinkling
and springback when compared with mild steel (Bolt, 2001). Such drawbacks have pervented an extensive use of lightweight materials and
composites in automobiles. Even though there is a huge potential for further
fuel savings by using lightweight material for various components of
transportation vehicles, the justification of lightweight material structural
parts over steel is very difficult and questionable with conventional production processes like forging and stamping because of the formability
constraints.
16.3
Forming technologies for lightweight materials
The material characteristic of lightweight materials is one of the major
barriers for their wide implementations into automotive body panels. It is
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Hydroforming for advanced manufacturing
well known that aluminum and magnesium alloys are not ductile and
have low formability at room temperature due to their crystal structures.
Therefore, conventional manufacturing processes for lightweight materials
are usually hot forging, extrusion, casting and machining. Recently, rapid
solidification processes such as spray deposition or melt-spinning have also
been investigated. However, they are far away from leading to thin, smoothsurfaced body parts.
As a viable alternative to existing stamping and hydroforming operations,
warm forming of lightweight material has been investigated since the 1960s.
At warm temperatures (100–300 °C), the formability of Al and Mg increases
and up to 300% total elongation can be achieved. Sophisticated post-process
heat treatment needs can also be avoided as opposed to the hot forming
case. Yet, it would be possible to achieve precise part dimensions and surfaces as the warm forming compared with hot forming that ensures close
and reasonably controllable tolerances. Various investigations and attempts
have been made to form lightweight materials at elevated temperatures
since the 1940s (Ayres, 1978 and Doege, 1996). A substantial increase in the
elongation and decrease in the tensile strength with increasing temperature
and decreasing strain rates have been reported for Al–Mg alloys (Ayres,
1978 and Shehata 1993). The effect of temperature (20–300 °C) on the
formability of several Al–Mg alloys over a wide range of strain rate was
presented as a result of uniaxial and biaxial stretch forming experiments
(Naka, 1998, 1999, 2001). Bolt and his colleagues (2001) conducted warm
forming experiments on various Al alloys (1050, 5754-O and 6016-T4)
between 100 and 250 °C using box-shaped and conical-rectangular dies.
Similar studies on Mg alloys were conducted by Takuda (1999, 2002) and
Doege (1996, 1997, 2001). Takuda conducted their experimental work on
AZ31 and AZ91 magnesium alloys as they were commercially available in
different forms and commonly used. Doege, on the other hand, conducted
comprehensive experimental work including aluminum and magnesium
alloys at different temperatures and forming speeds. Flow stress curves of
AZ31B and AZ61B magnesium alloy at 50, 100, 150, 200 and 235 °C at
various strain rates (i.e. 2–0.002 s−1) were obtained. It was observed that, for
AZ31B alloy, maximum LDR was achievable between 175 and 210 °C. All
the results show that formability and LDR (limiting drawing ratio) increase
at elevated temperatures (150–300 °C). High strain rate is shown to increase
the flow stress, and hence decrease the formability and LDR (Fig. 16.1).
16.4
Warm hydroforming: state-of-the-art review
In order to further extend the forming degrees of lightweight materials, a
hybrid warm hydroforming process is being investigated by many researchers. Similar ideas have been discussed within the forming community
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Warm hydroforming of lightweight materials
LDR
3
2
355
A5181-O
A5052-O
1
0
Deep drawing depth at 25°C
27
100 150 200 250
Temperature (°C)
Deep drawing depth at 250°C
16.1 Effect of heating on the formability of Al alloys; limiting drawing
ratio (LDR) increases more than 100% (after Takata, 2000).
since as early as the beginning of the 1990s. However, written literature
suggesting or investigating this possibility goes back only to the early 2000s
(Nakamura, 1997; Vollertsen, 1999; Groche and Schmoeckel, 2001; Lee
et al., 2002). Warm hydroforming simply can be applied in two ways: bulging
of heated blank(s) into a die cavity via fluid pressure and deep-drawing of
a blank against a hydraulic force (hydromechanical forming) as illustrated
in Fig. 16.2.
The very important premise of the warm hydroforming process is to
increase the formability of lightweight materials beyond limits that are
achievable in conventional cold forming processes because of a reduction
in the friction between workpiece and tooling elements and a decrease in
the flow stress of material at elevated temperatures. Thus, material flow into
the expansion areas or intricate regions takes place with very low forming
loads while some sections of the blank in contact with fluid medium cool
rapidly to increase the forming limit via strain hardening. Consequently,
reduced forming loads will result in small forming equipment requirements
with low capital equipment investment savings as illustrated in Fig. 16.3.
Fig. 16.3 also illustrates another benefit of warm hydroforming where
intricate part features can be formed with less pressure. Similarly, increased
formability would result in consolidation of multiple parts leading to reductions in joining/assembly operations contributing to both cost savings and
increased integrity.
On the other hand, additional scientific challenges are introduced with
the warm hydroforming process. These challenges include:
(1)
(2)
appropriate design and control of optimal temperature distribution
on tooling elements (i.e., die, blank holder/punches) and blank
material,
prediction and compensation of consequent residual stresses and
distortions;
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Hydroforming for advanced manufacturing
Part consolidation and
weight reduction
Stamping at room
temperature
Hydroforming
(tube and sheet)
Warm forming
(stamping)
Warm
hydroforming
Increased formability
Increased formability,
part consolidation, weight reduction,
cost effective
a
Pi
b
Pi
Pi
c
16.2 a Basic elements of the warm hydroforming process, b
hydroforming of double blanks, c warm hydroforming against fluid
pressure, i.e., hydromechanical forming.
(3)
prediction of optimal and synchronous loading paths (e.g., pressure vs.
time, temperature vs. time, etc.),
(4) understanding and modeling of the complex surface interactions, friction and effective of use of lubrication at elevated temperatures,
(5) effect of warm temperature conditions on the material properties,
formability, and failure modes.
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Warm hydroforming of lightweight materials
357
Required press tonnage (cold hydroforming)
9.0E+05
Required press tonnage (warm hydroforming)
8.0E+05
Pi
Press tonnage (kN)
7.0E+05
6.0E+05
5.0E+05
4.0E+05
3.0E+05
2.0E+05
1.0E+05
0.0E+05
0
100
200
300
400
500
Internal pressure (MPa)
a
Rc @ T = 200 °C
40
Rc @ T = 25 °C
35
(Pi)max=
2
√3
sf ln
Pi
rc
rc – t
Corner radii rc (mm)
30
25
rc
20
Pi
15
Pi
10
5
0
0
50
100
150
200
250
300
Internal pressure Pi (MPa)
b
16.3 a Increase in formability with warm hydroforming process; b the
same corner radius size can be formed with an internal pressure of
70 MPa instead of 170 MPa with cold hydroforming, leading to lower
equipment requirements.
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Hydroforming for advanced manufacturing
In addition, practical issues such as handling and containment of warm
blank material, pressurizing fluid medium, lubricants, and cleaning and
post-processing of formed parts become development challenges.
Warm hydroforming for tube and sheet have been investigated and tested
by a few groups mainly in Germany. HEATforming is a recently developed
tube hydroforming technique introduced by Schuler. In this process, a
heated tube is placed in a heated die; then the die is closed; and the tube
sealed at the ends by sealing cylinders. The tube is subsequently expanded
against the die cavity wall by internal pressure (Pi) – here provided
by a gaseous medium. The process may also be supported by continued
axial feeding of the tube, similar to conventional hydroforming. The tube
material and the die can be adjusted to various temperature zones for
control of the material flow. The flow of the material in the die is further
aided by specially developed lubricants. Figure 1.24 denotes the steps for
HEATforming.
Siegert and co-workers at Stuttgart University conducted experiments
for high pressure sheet hydroforming and hydromechanical deep drawing
(HMD) at elevated temperatures (250–300 °C) (Siegert, 2003). They
measured material properties for AZ31, a Mg alloy, by bulge and tensile
tests. Similarly to Al alloys, the Mg alloy has high formability at elevated
temperature, and flow stress increase at high strain rates. For high-pressure
sheet hydroforming, they made a license panel pocket (Fig. 16.4a).
Heating cartridges
Upper pressure chamber
Pi
Upper die
Heating
devices
Lock-bead
Blankholder
Punch
Blank
Draw ring
Insulation
Counter
pressure pot
Band-heater
Pc
Female die
Pc
Lower pressure chamber
860
350
55
Tp = 200 °C
a
b
16.4 Tooling and products of warm hydroforming: a sheet
hydroforming tooling and final product (Siegert, 2003), b HMD tooling
and final product (Siegert, 2004).
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Warm hydroforming of lightweight materials
359
Pneumatic bulging was used to form the license panel pocket. The female
die was heated by a band heater outside of a female die. The upper die was
heated by cartridge heaters. The sheet was fixed by lock-bead and high
pressure was applied from upper die. The pocket geometry was 350 by
850 mm. Thickness was 2.3 mm and part depth was 55 mm. For hydromechanical deep drawing (Fig. 16.4b), they used a drawing punch with a dome
at the center of punch. The flange area including both die and blank holder
region was kept at an elevated temperature (240–260 °C). The temperature
of fluid used for hydromechanical deep drawing was controlled by a heat
exchanger and kept at 220–250 °C. the pressure for warm HMD was 7.5 MPa.
They measured equivalent strains at the bottom of cup. At 220–250 °C,
the effect of fluid temperature on equivalent strain was small. They also
investigated the effect of punch temperature on the effective strains at
the bottom of cups. As the temperature of the punch was increased, the
equivalent strain at the cup bottom was also increased.
Groche and co-workers at the Technical University of Darmstadt (2002)
investigated hydromechanical deep drawing of Al alloys at elevated
temperatures (Fig. 16.5). They reported reasonable strategies to increase
drawing ratio. First, the reduction of flow stress of the material in the flange
improves the drawing ratio. Therefore, heating of the flange region is needed.
Blankholder
Fluid bead
Cooled
punch
T
q·
Cooled
drawing
die radius
P
T
T
q·
Heated
die area
P
Counter pressure pot
Counter pressure (MPa)
a
25
Phase I
Phase II
20
Phase III
rst
15
10
hwall
5
0
rZK
0
50
100
150
Punch travel (mm)
b
16.5 a Tooling for warm HMD, b pressure control strategy (Groche
2002).
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Limiting drawing ratio
Secondly, the reduction of friction force at the die corner increases the
drawing ratio. This could be reduced by drawing over fluid bead. Third, an
increase in force transmission was achieved by increasing the friction
between the punch and the workpiece by means of a counter pressure.
For material properties, they also determined the properties by bulge
tests at elevated temperatures. The strain rate was controlled by flow rate
control. The results of the bulge test showed higher flow stress than tensile
tests. The friction conditions were also tested. They conducted the friction
test at different temperatures and contact pressures. In order to find
suitable system and process parameters, thermomechanically coupled
simulation was conducted. Counter pressure was controlled with three
steps. First, no counter pressure was applied at the beginning, because of
excessive stretch. Second, the counter pressure was determined for each
time step by calculating the drawing force and support blank on a fluid
bead. Third, a controlled decrease in the counter pressure control was
attempted. This control was initiated as soon as the friction force between
punch and blank exceeded the difference between the wall load and the
transmissible force in the bottom area (Fig. 16.5b). After the preliminary
finite element analysis (FEA) test, they studied the limiting drawing ratio
of warm deep drawing and warm hydromechanical deep drawing. The
results shows that a higher drawing ratio could be achieved by hydromechanical deep drawing at elevated temperatures (Fig. 16.6).
Geiger and co-workers at the University of Erlangen (2004) used an
optical strain measurement system to measure the mechanical properties
of Al alloys. The test sets were biaxial and bulging test, equipped with heatable tools and fluid. The feeding of a warm medium to the pressure intensifier was effected by a three-tank system. The medium could also be quickly
evacuated from the tool by use of a vacuum suction system. The forming
3.0
2.5
2.0
200
250
300
100
150
Flange temperature (°C)
DD:= Deep drawing HM:= hydromechanical deep drawing
Punch diameter: 100 mm
DD: AIMg4.5Mn
7 mm
Die radius:
DD: AIMg0.4Si1.2
7 mm
Punch radius:
HM: AIMg4.5Mn
5 mm/s
HM: AIMg0.4Si1.2 Punch velocity:
0
50
16.6 LDR of deep drawing and warm HMD (Groche, 2002).
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Warm hydroforming of lightweight materials
361
R ≈ 7 mm
R ≈ 20 mm
a
b
16.7 Sheet hydroforming at room and elevated temperature (Geiger
et al., 2004), a cylindrical cup tool, left at room temperature and right
at 230 °C, b box-shaped part tool, left at room temperature and right at
230 °C.
a
b
16.8 Parts formed by a AA6016 and b AA5182 at elevated temperature
(Geiger et al., 2004).
tools are equipped for heating, insulation and cooling of the punch. Measurement and control of temperature by thermocouples, multiple splash
guards and vapor evacuation and filtering system were equipped. For sheet
hydroforming, cylindrical and box-shaped were conducted for AZ31, the
Mg alloy. While at room temperature, a premature failure occurs in the die
corner region, long before the sheet contacts the tool upper surface, at
temperature around 230 °C, a good part depth and corner filling was
achieved at elevated temperatures (Fig. 16.7). For example, a license plate
support was investigated (Fig. 16.8). This aluminum part is not feasible by
conventional deep drawing, but could be completely formed in the 6400 kN
hydraulic press at pressure around 150 bar.
Tube hydroforming experiments at elevated temperatures were carried
out with simple bulging tests. The tube material was an Al alloy (AlMg3.5 Mn).
The tube was welded and annealed in order to get good formability. Figure
16.9 shows the pressure curve versus the injected volume for the process at
room and elevated temperature. Also, the sequence of the forming process
is represented. At room temperature, the pressure increased very quickly
and failure of the tube occurred. By increasing the tool temperature up to
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Hydroforming for advanced manufacturing
300
Room
temperature
bar
Tool: 275 °C
Medium: 225 °C
Pressure p
200
150
100
50
0
a
0
100 200 300 400 500 600 700 800 900 1000
Volume (ml)
b
16.9 Bulging of tube (Geiger, et al., 2004), a warm bulging test at
220 °C, b tube hydroforming at room and warm temperature.
275 °C and medium temperature to 225 °C, the volume could be achieved
around 600 ml at 100 bar.
In summary, a few research attempts on warm hydroforming are still in
their early stages as far as reported in the literature. They are mostly focused
on demonstrating the feasibility of warm hydroforming of a few selected
alloys such as AA5182, AA6016 and AZ31B. Novel tool designs, process
improvements and understanding the temperature effect on feasibility are
being studied. Limited studies have been carried out to characterize the
material and friction conditions at warm hydroforming condition. On the
other hand, in order to bring these technologies to real plant floor applications from lab-scale experimentation, R&D in the following areas need to
be conducted:
(1)
Comprehensive modeling of material behavior and friction
condition.
(2) Full understanding of temperature distribution, heating strategies and
thermal effects on formability of the lightweight materials. Development of methods or models to enable quick determination of optimal
WPNL2204
Warm hydroforming of lightweight materials
363
temperature distribution and a proper heating scheme for a given part
shape, material, etc. are needed.
(3) Similarly, understanding and modeling of the effect of non-uniform
heating on the formed part quality is needed.
(4) Combinational optimization of Pi and blank holder force (BHF), in
the case of warm sheet hydroforming, is a necessary task. Development of methods of numerical modules is needed in order to reduce
costly and lengthy trial errors in the plant floor and computer.
(5) Improved heating, heating control, insulation and sealing techniques
are necessary for the success of warm hydroforming technology.
16.5
Comparison of warm and cold hydroforming:
a numerical study
The warm hydroforming process enabled higher LDR (limiting drawing
ratio) compared with the warm forming and cold hydroforming processes
(Fig. 16.7). To understand the deformation mechanism and the differences
in these three processes, temperature, stress, strain and thickness, data from
the FEA were compared. The FE model based on Groche’s experiments
(2002) was used for the analyses. The blank diameter was 145 mm for all
cases. The temperature conditions for each process are shown in Fig. 16.11.
The punch temperature of the warm hydroforming and warm forming was
considered as 50 °C to reflect an increase in temperature during the manufacturing process. The fluid temperature and die corner temperature of the
warm hydroforming process were assumed to be the same, because they
were in contact with each other. The flange temperature of the warm hydroforming and warm forming process was assumed to be 250 °C to ease material flow in the flange region.
As depicted in Fig. 16.10, the warm hydroforming process ended successfully with a part depth of 191 mm. But both the warm forming and cold
hydroforming processes resulted in local thinning (35%) at smaller part
depths of 38 mm and 17 mm, respectively, compared with the warm hydroforming case. In brief, the warm hydroforming process resulted in higher
part depths for the following reasons:
(1) High temperature in the flange region: material at the flange region
was heated, and therefore, its flow stress was decreased remarkably
leading to high ductility and easy movement.
(2) Low temperature near the die corner and punch wall region: at the
punch wall region, blank material is in continuous contact with the
punch wall due to hydraulic pressure. Therefore, the temperature level
of the blank at the punch wall and die corner regions is low because
of contact heat transfer with the punch wall (conduction) and with the
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Hydroforming for advanced manufacturing
Warm hydroforming
Warm forming
Cold hydroforming
Tp = 50 °C
Tf = 250 °C
Tdc = 70 °C
N/A
Tb = 20 °C
Tp = 20 °C
Tf = 20 °C
Tdc = 20 °C
Thm = 20 °C
Tb = 20 °C
Tp = 50 °C (punch)
Tf = 250 °C (flange)
Tdc = 70 °C (die corner)
Thm = 70 °C (hyd. medium)
Tb = 20 °C (initial blank)
167 °C
188 °C
203 °C
188 °C
167 °C
203 °C
146 °C
Thinning
146 °C
125 °C
Thinning
104 °C
125 °C
NT11
+2.500e+02
+2.292e+02
+2.083e+02
+1.875e+02
+1.667e+02
+1.458e+02
+1.250e+02
+1.042e+02
+8.333e+01
+6.250e+01
+4.167e+01
+2.083e+01
+0.000e+00
83 °C
Fluid
NT11
+2.500e+02
+2.292e+02
+2.083e+02
+1.875e+02
+1.667e+02
+1.458e+02
+1.250e+02
+1.042e+02
+8.333e+01
+6.250e+01
+4.167e+01
+2.083e+01
+0.000e+00
104 °C
NT11
+2.500e+02
+2.292e+02
+2.083e+02
+1.875e+02
+1.667e+02
+1.458e+02
+1.250e+02
+1.042e+02
+8.333e+01
+6.250e+01
+4.167e+01
+2.083e+01
+0.000e+00
Fluid
83 °C
63 °C
a
S, Mises
(Ave. Crit.: 75%)
+2.838e+08
+2.602e+08
+2.365e+08
+2.129e+08
+1.872e+08
+1.656e+08
+1.419e+08
+1.183e+08
+9.461e+07
+7.096e+07
+4.731e+07
+2.365e+07
+0.000e+00
S, Mises
(Ave. Crit.: 75%)
+2.966e+08
+2.719e+08
+2.472e+08
+2.225e+08
+1.978e+08
+1.730e+08
+1.483e+08
+1.236e+08
+9.888e+07
+7.416e+07
+4.944e+07
+2.472e+07
+0.000e+00
S, Mises
Thinning
(Ave. Crit.: 75%)
+3.602e+08
+3.302e+08
+3.001e+08
+2.701e+08
+2.401e+08
+2.101e+08
+1.801e+08
+1.501e+08
+1.201e+08
+9.004e+07
+6.003e+07
+3.001e+07
+0.000e+00
Thinning
Fluid
Fluid
38.08 mm depth;
successful drawing (191 mm)
38.08 mm depth;
thinning at the die
corner
16.84 mm depth;
thinning at the
punch corner
b
16.10 a temperature and b stress distributions on the blank material
in warm hydroforming, warm forming and cold hydroforming
processes.
hydraulic medium (convection) resulting in an increased material
strength providing additional resistance to drawing force. However, in
the warm forming case, the temperature range between punch wall
and die corner regions is higher than that with warm hydroforming
(167–183 °C vs. 146–169 °C, respectively, as in Fig. 16.10a). Thus, the
blank material does not have enough strength (i.e., low flow stress) to
resist the forming forces, and thinning occurs at a smaller part depth
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Warm hydroforming of lightweight materials
365
value (in this case, at 38 mm, versus 191 mm). The contact of the blank
material with the punch wall was found to have a profound effect on
reducing the temperature level and increasing the strength because
the conduction heat transfer (i.e., between blank and punch wall,
1400 W m−2 K−1) is much higher than the convection heat transfer
(between blank and hydraulic medium, 173 W m−2 K−1).
(3) Friction effect: application of the hydraulic pressure levitates the blank
material over the die corner region leading to a reduction of friction
between blank and die corner. Consequently, this enhances the movement of material from the flange region into the cavity, thus leading
to less thinning. In addition, increased friction between the punch wall
and the blank because of hydraulic pressure and low temperatures
contributes to a reduction in the stress levels in that region giving rise
to stress reduction and less thinning. At the punch face, punch corner
and punch wall regions, the friction force helps the material stick to
the punch and not to move in a radial direction. Therefore, the friction
force shares the radial stress (σ1) and helps to release tension stresses
which causes thinning within the material.
The temperature distribution (Fig. 16.11a and b) shows an advantage of
warm hydroforming. As the forming process proceeds, the temperature of
the part near the punch wall decreases due to the contact between punch
wall and the material. For the warm hydroforming case shown in Fig. 16.11a,
the temperature of a part near the punch wall and face is mostly 69 °C. The
blank material cooled by the punch then sustains high drawing forces with
reduced thinning. For warm forming, because of the high temperature near
the die corner, the material strength is low and cannot withstand the bending
and stretching forces (Fig. 16.11a and b). The equivalent plastic strain
distributions at several forming steps (i.e., of different depth value) are
presented in Fig. 16.11c and d). For the warm hydroforming case, the
equivalent strain was distributed more evenly over the blank material
compared with the warm forming and cold hydroforming. It is apparent
that the maximum strain of the warm hydroforming at a depth of 50 mm
was lower than that of warm forming at a part depth of 38 mm. For the cold
hydroforming process, the strain at the punch face region was very high and
thinning appeared near the punch corner region. The thickness distribution
on the blank is similar to the strain distribution as shown in Fig. 16.11e.
The warm hydroforming process results in an even thinning distribution
minimizing the local thinning occurrence when compared with the other
processes. For the warm forming process, the localized thinning occurs
near the die corner and thinning ratio was very sharp. The thinning for the
cold hydroforming occurs at the punch corner region and thinning was
distributed all over the punch corner.
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Hydroforming for advanced manufacturing
Warm hydroforming
Warm forming
path
Temperature (°C)
A=174
B=156
C=139
D=122
E=104
F=87
G=69
Depth=191 mm
250
250
depth
150
100
50
path
a
A
250
5 mm
10
15
25
38.08
150
100
Path (mid-plane)
Depth=191 mm
depth
1
0.8
A
B
A
C
B
C
D
E
F
G
1.2
5 mm
10
15
25
50
100
150
191
0.6
0.4
0.2
Plastic equivalent strain
1.2
A
D
E
GF
0
0
Plastic equivalent
strain
A=0.687
B=0.591
C=0.494
D=0.397
Depth=38.08 mm
E=0.299
F=0.202
Path (mid-plane)
G=0.105
Plastic equivalent
strain
A=0.444
B=0.381
C=0.317
D=0.254
Depth=16.84 mm
E=0.191
Path (mid-plane)
F=0.127
G=0.064
EF
E
C
D
G
A BC D
1.2
0.4
0.2
1.4
1.3
1.2
1.1
1
0.9
0.8
0.6
thinning
0.4
0.2
0
Node number for the path
depth
1.5
1.4
1.6
1.2
1.1
1
0.9
1.4
5 mm
10
15
16.84
1.2
1.1
1
0.9
0.8
0.8
0.7
0.7
0.6
0.6
Node number for the path
d
1.3
0.7
50 100 150 200 250
depth
1.5
0.8
0
50 100 150 200 250
Node number for the path
5 mm
10
15
25
38.08
1.3
c
5 mm
10
15
16.84
0
1.6
5 mm
10
15
25
50
100
150
191
depth
1
50 100 150 200 250
Thickness (mm)
depth
1.5
b
FG
E
A BCD
0.6
0
Thickness (mm)
1.6
50 100 150 200 250
Node number for path
5 mm
10
15
25
38.08
thinning
0.8
Node number for the path
uniform temp.
Node number for path
depth
1
50 100 150 200 250
100
50 100 150 200 250
0
0
150
0
0
Node number for path
Plastic
equivalent
strain
A=1.009
B=0.866
C=0.724
D=0.582
E=0.439
F=0.297
G=0.155
5 mm
10
15
16.84
50
Plastic equivalent strain
50 100 150 200 250
depth
200
0
0
Plastic equivalent strain
Depth=16.84 mm
50
0
Thickness (mm)
Temperature (°C)
CONSTANT=20
depth
200
Temperature (°C)
200
Temperature (°C)
CB
ED
F
G
5 mm
10
15
25
50
100
150
191
Cold hydroforming
Temperature (°C)
A=223
B=198
C=173
D=148
Depth=38.08 mm
E=123
F=98
path
G=72
A
CB
D
E
F
G
Temperature (°C)
366
0.6
0
50 100 150 200 250
Node number for the path
0
50 100 150 200 250
Node number for the path
16.11 Comparison of a and b temperature, c and d strain, and e
thickness distribution on the blank in warm hydroforming, warm
forming and cold hydroforming processes.
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e
Warm hydroforming of lightweight materials
Warm hydroforming
Warm forming
367
Cold hydroforming
d
c
thinning
d
d
c
b
0.2
15 mm
50 mm
150 mm
191 mm
d
b
0
–1.2 –1 –0.8–0.6 –0.4–0.2 0
Minor strain
1
1
0.6
0.4
thnning
c
a
0.2
0.8
0.6
c
0.4
b
a
0.2
15 mm
38.08 mm d
0
–1.2 –1 –0.8–0.6 –0.4–0.2 0 0.2
Minor strain
Major strain
Thins
1
Thickens
0.8
b
a
a
b
Major strain
Major strain
a
c
0.8
0.6
0.4
c
b
0.2
a
16.84 mm d
0
–1.2 –1 –0.8 –0.6 –0.4 –0.2 0 0.2
Minor strain
16.12 Comparison of minor–major strain distribution on the blank at
each stage of warm hydroforming, warm forming and cold
hydroforming.
Finally, major and minor strain distributions at different stages of the
forming processes are depicted in Fig. 16.12. For the cold hydroforming
process, the blank material near the punch face was elongated biaxially (a).
Peak strain appeared at the punch corner region (b). Then, the strain
decreased in the die corner and flange region because the drawing force
could not be delivered owing to the thinning (c,d). At the edge of the flange
region, minor and major strains were almost zero (d). For the warm forming
process, the punch face region (between a and b) was hardly stretched compared with the cold hydroforming process. Near the die corner region (c),
magnitudes of major and minor strains were increased and thinning
occurred. At point (c), excessive thinning (35%) occurred, and failure was
assumed. The magnitude of major and minor strains at the flange edge did
not change much. But a slight thickening of the blank was observed. In the
case of the warm hydroforming process, the strain distribution was quite
similar to that in the warm forming until part depth was 50 mm. As the
forming process proceeded, the minor–major strain curves became smooth
and also dropped causing thickening at the flange region. In Fig. 16.12, it
can be seen that the warm hydroforming process utilizes the blank material
more efficiently.
16.6
Process design and control in
warm hydroforming
There are many variables that should be controlled for a successful warm
hydroforming process (Fig. 16.13). These variables can be categorized into
WPNL2204
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Hydroforming for advanced manufacturing
Controllable variables
Response characteristics
on deformation
Die temperature
Punch temperature
Hyd. medium and die
corner temperature
Slow response (~h)
Spatial control
Temp.=f(x,y,z)
Control strategy
Benefits
Hybrid approach with DOE and
adaptive isothermal FEA
Optimal temp. distribution
Hydraulic pressure
Blankholder force
Punch speed
Fast response (~0.01 s)
Temporal control
Pressure, BHF, Speed
=f(x,y,z,t)
Rapid process
development
Less trial and error
Less dimensional
variation
Adaptive FEA
with fuzzy control algorithm
37=2187 ...
Optimal loading profile
Blankholder force
Punch
temperature
Die temperature
p
Pi
BHF
Hydraulic
temperature
Hydraulic pressure
t
Punch speed
V
t
t
16.13 Process control strategy for warm hydroforming.
two groups. The first group includes the variables that have a slow response
on the deformation characteristics of materials. An example of such a
variable is the temperature of the tooling and hydraulic medium, which in
turn affects the temperature distribution in the blank material. Generally,
the tooling and the hydraulic medium have a large thermal capacity which
results in very slow responses, in ranges of hours. Therefore, during the
forming cycle, the temperature levels of the tooling and the hydraulic
medium cannot be controlled on a time basis which is usually in the range
of 10–30 s depending on the part size. To ensure increased part formability,
an appropriate way would be to control the temperature levels spatially
with selective heating or cooling devices. The second group consists of
variables that have a rapid response on the part forming characteristics such
as blankholder force, hydraulic pressure and punch speed. These variables
can be controlled temporally on the foundation of the pre-determined
optimal temperature distribution (Fig. 16.13). The second group of process
variables is also optimized through a fuzzy control algorithm (Choi and
Koç, 2007).
16.6.1 Temperature distribution design and control
It was shown that certain temperature conditions in the warm hydroforming
process increase the forming limits of the lightweight materials. For the
regular warm deep drawing process, it is well known that the cold punch
temperature and the warm flange temperature increase the formability in
WPNL2204
Warm hydroforming of lightweight materials
350
Flange region
1 , 3, 4
Flow stress kf (MPa)
2
Punch face
region
a
Flow stress kf (MPa)
300
Material: AZ31B, s0= 1.0 mm
T = 50 °C
300
Punch wall 2
region
T = 100 °C
2
250
T = 150 °C
200
T = 200 °C
T = 235 °C
150
1
100
.
j = 0.002 s–1
50
0
0
0.05
0.1
0.15
0.2
0.25
0.3
Logarithmic strain j
4
200
150
0.35
0.4
b
Material: AZ31B (s0= 1.0 mm), T = 200 °C
250
369
.
j = 2.0 s–1
.
j = 0.2 s–1.
j = 0.02 s.–1
j = 0.002 s–1
3
100
50
0
0
0.05
0.1
0.15
0.2
0.25
0.03
0.35
0.4
Logarithmic strain j
c
16.14 Determination of proper temperature levels and strain rates
for a part made from lightweight materials: a a simple cup part,
b temperature-dependent flow curves (Doege, 2001), c strain-ratedependent flow curves (Doege, 2001).
a simple deep drawing case. The temperature condition has a large influence
on the material properties (flow stress and strain) of lightweight materials
as shown is Fig. 16.14. For instance, it was reported (Jager, 2002) that, for
the Mg alloy AZ31, flow stress decreases by 50% at elevated temperature
levels ( Fig. 16.14c) compared with room temperature (Fig. 16.14d).
Moreover, at the elevated temperatures, the flow stress and strain become
highly sensitive to the strain rate (Fig. 16.15c), while the sensitivity is not
observed at the room temperature conditions. Even though the effect of
the strain rate is less than that of the temperature, it affects the flow stress
and the strain (total elongation) considerably at the elevated temperatures
(Fig. 16.14e and f). Therefore, to increase the formability, the flow stress
and the strain should be locally controlled by regulating both the temperature and the strain rate (Fig. 16.14a). In the warm deep drawing of 5xxx
aluminum alloys, it is known that the flange temperature should be kept
high to ease the material flow into the die cavity. On the contrary, the punch
temperature should be kept cold to increase the material strength which
will increase the drawability. In an elevated temperature region such as
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Hydroforming for advanced manufacturing
Punch
speed (V)
Blankholder
force (BHF)
187 °C 218 °C
156 °C
125 °C
100 °C 250 °C
93 °C
Blankholder force
BHF
62 °C
Pi
25 °C
31 °C
Hydraulic pressure
t
Punch speed
V
t
t
p Hydraulic
pressure (Pi)
16.15 Temporal control of loading profiles: a controllable variables
(blankholder force, hydraulic pressure and punch speed), b temporal
control of loading profiles.
the flange region, the strain rate should also be considered to increase the
part formability (Fig. 16.14e and f). The strain rate can be controlled by
changing the punch speed or the rate of change in the hydraulic pressure.
After determining the optimal temperature distribution spatially, the
optimal change of punch speed or hydraulic pressure by time should be
conducted.
16.6.2 Loading control
For a successful warm hydroforming process, many process variables should
be controlled appropriately. In a previous study (Choi and Koç, 2007), a
methodology to determine optimal temperature distributions of the tooling
and hydraulic medium was developed. It was demonstrated that there exist
optimal temperature conditions resulting in better formability of lightweight materials. In addition to the methodology to rapidly determine
the optimal temperature distribution, it was concluded that loading parameters should be optimized in conjunction with an optimal temperature
distribution.
Many researchers have developed control strategies for loading parameters in sheet and tube hydroforming processes at cold state. Shulkin and
co-workers (1997) built an 8-point blank holder force control system and
conducted sheet viscous forming experiments. FEA was also conducted to
fine-tune its control system. Ray and McDonald (2004) determined the
optimal loading path for tube hydroforming processes using a fuzzy load
control algorithm and FEA. They developed process windows for hydraulic
pressure and axial feeding in x- or t-branch tube hydroforming. Lundqvist
(2004) investigated adaptive loading algorithms and presented a thorough
literature survey. Abedrabbo and co-workers (2005) investigated the
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Warm hydroforming of lightweight materials
371
wrinkling behavior of aluminum sheet during the cold sheet hydroforming
process numerically and experimentally. Aydemir and co-workers (2005)
presented an adaptive method to obtain a more efficient process control
for tube hydroforming processes. Process parameters such as hydraulic
pressure and axial feeding were adjusted during FE simulation via a fuzzy
knowledge based controller.
On the basis of pre-determined optimal temperature distribution for the
warm hydromechanical deep drawing (HMD) case, loading profiles such as
hydraulic pressure and blankholder force profiles were investigated by Choi
and Koç based on the adaptive FEA technique coupled with a fuzzy control
algorithm (2007). Control strategies of the loading profiles are established
to satisfy geometrical criteria (thinning, wrinkling, floating and contact) and
material characteristics (strain rate sensitivity) of the warm hydroformed
parts. The optimal loading profiles of hydraulic pressure and blank holder
force are determined for the warm HMD process to satisfy geometric
criteria. Similarly, the effect of punch speed conditions is studied with
respect to formability and temperature distribution.
For the development of the warm hydroforming process, loading profiles
of hydraulic pressure, blankholder force and punch speed can be determined through many trial and error experiments and/or preceding experience. However, it would require a long time and tremendous manpower. To
shorten the development time, an adaptive FE analysis with fuzzy control
algorithm was developed in this study. Using this approach, the loading
profiles could be determined by running only few adaptive FEA simulations including the tuning of fuzzy control parameters.
In order to describe the proposed method and its application in warm
hydroforming, first the controllable process variables of this problem have
to be explained. Controllable process variables in the warm hydroforming
process can be grouped into two major categories according to their time
responses (Fig. 16.13). The first group includes variables that have a slow
response on deformation characteristics of the material (i.e., in the range of
1–3 h). An example of such a variable is the temperature of the tooling and
hydraulic medium, which in turn affects the temperature distribution of the
blank material. Generally, the tooling and hydraulic mediums have large
thermal capacity which results in very slow thermal responses, in the order
of hours. Therefore, the temperature of the tooling and hydraulic medium
cannot be controlled temporally during each forming cycle, which is usually
in the range of 10–30 s depending on the part size. Instead, in order to ensure
increased part formability, an appropriate approach would be to control the
temperature spatially with selective heating or cooling devices as this was
shown to be very important and effective. The second group consists of
variables that have rapid responses on part forming characteristics such as
blank holder force, hydraulic pressure and punch speed (Fig. 16.15). These
WPNL2204
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Hydroforming for advanced manufacturing
Blankholder
force
U2, vertical displacement
Punch
speed
U, U2 (m)
+7.570e–03
+6.623e–03
+5.677e–03
+4.731e–03
+3.785e–03
+2.839e–03
+1.892e–03
+9.462e–04
+0.000e+00
–3.000e–02
p
Hydraulic
pressure
1 Thinning
2 Wrinkle
Blankholder
force
Punch
speed
a
156 °C187 °C 218 °C
125 °C
4 Corner
floating
93 °C 100 °C
3 Wall
contact 62 °C
250 °C
25 °C
31 °C
p
Hydraulic
pressure
b
16.16 Criteria used to ensure increased formability and net-shape:
a thinning and wrinkling criteria, b punch wall contact and corner
floating.
variables can be controlled during each forming cycle very quickly (i.e., in
the order of 10−2 s). Therefore, these variables can be controlled temporally
by the use of pre-determined optimal temperature distributions. In this
study, adaptive FE analysis with a fuzzy control algorithm is used to obtain
the loading profiles with a single simulation.
Figure 16.16 depicts the common criteria used to control hydraulic pressure and blank holder force profiles with the fuzzy control algorithm. First,
for an increased formability, the maximum thinning of the entire blank
material should be checked during the forming process while the hydraulic
pressure and blank holder force profiles (Fig. 16.16c) are controlled by the
fuzzy control algorithm. Second, to make sure the blank materials conform
to the designed shape of the dies, their wrinkling should also be checked at
the flange area as shown in Fig. 16.16d. In the warm hydroforming process,
the blank material along the punch wall should sustain a high enough
strength to draw the material into the die cavity from the flange region.
Accordingly, the blank material should be cooled at this region to enhance
WPNL2204
Warm hydroforming of lightweight materials
373
material strength by the cold punch. Hence, in order to secure a continuous
conductive heat transfer between the punch and blank material along the
punch wall, contact should be ensured during the simulation (Fig. 16.16e).
One of the many benefits of the warm hydroforming process is the relief of
the friction force on the blank material around the die corner region. Floating the blank material over the die corner because of the appropriate application of hydraulic pressure will reduce the friction force between the blank
material and die corner as shown in Fig. 16.16f. The floating condition also
helps to increase curvature of the blank material near the die corner region,
which increases formability. This will be explained in later sections.
In Fig. 16.17, the strain rate sensitivity of the magnesium alloy (AZ31B)
is illustrated based on material tests conducted by Droeder (1999). As
shown in Fig. 16.17c, flow stress is not sensitive to strain rate at room
temperature. However, at elevated temperatures, the strain rate sensitivity
of the lightweight materials increases considerably (Fig. 16.17d∼e). Therefore, the strain rate should also be checked during the warm hydroforming
process. In the warm HMD case (Fig. 16.17b), the blank material near the
punch region maintains a low temperature. As a result, flow stress is not
affected by strain rate. However, at elevated temperatures such as in the
flange region, flow stress of the blank material becomes very sensitive to
strain rate. In this case, strain rate can be controlled by changing the punch
speed. Fig. 16.17c shows a simple warm sheet hydroforming case. The rugged
punch is stationary and maintained at room temperature. The temperature
of the flange region is maintained at 220 °C. The initial temperature of the
blank is 220 °C. At the beginning of the forming process, the blank material
maintains elevated temperature. Therefore, at this stage, flow stress is very
sensitive to strain rate before it contacts the stationary cold punch. After
contacting the fixed cold punch, the material temperature drops by contact
heat transfer between the punch and blank material. In this case, strain rate
sensitivity diminishes because of the decreased temperature. For the warm
sheet hydroforming case, strain rate at the elevated temperature region can
be controlled by internal hydraulic pressure or flow rate of hydraulic
medium.
16.7
Characterization of materials for warm
hydroforming conditions
Formability of lightweight materials has been shown to increase with
increasing forming temperature up to their respective recrystallization
temperatures (e.g. 350 °C for Mg AZ31, 300 °C for Al5XXX, and 200 °C for
Al6XXX). A selective or localized heating strategy was also shown to
further enhance the formability of the aluminum sheets. In addition, higher
elongation could be obtained from these materials when lower strain
WPNL2204
Flow stress (MPa)
374
400
350
300
250
200
150
100
50
0
Hydroforming for advanced manufacturing
Material : AZ31B, s0 = 1.0 mm, T = 25 °C
.
j = 0.2 s–1
.
j = 2.0 s–1
1
.
j = 0.02 s–1
.
j = 0.002 s–1
0.05 0.10 0.15 0.20 0.25 0.30 0.35 0.40
Logarithmic strain j
Material: AZ31B (s0 = 1.0 mm), T = 200 °C
300
Flow stress kf (MPa)
0
250
.
j = 2.0 s–1
.
j = 0.2 s–1.
j = 0.02 s–1
.
j = 0.002 s–1
2
200
150
3
100
50
0
0
0.05
0.1
0.15 0.2 0.25 0.3
Logarithmic strain j
0.35
0.4
a
Blank
holder
force
Punch
speed
187 °C 218 °C
156 °C
2, 3
125 °C
93 °C 100 °C 250 °C
62 °C
25 °C
p Hydraulic
pressure
31 °C 1
b
Fixed counter punch
25 °C
171 °C195 °C
149 °C
1 93 °C122 °C
2 , 3
Blank
holder
force
195 °C
171 °C
195 °C
220 °C
p
Hydraulic
pressure
2 , 3
220 °C
c
16.17 Material characteristic (strain rate sensitivity): a flow stress
curves at room and elevated temperature with respect to the strain
rate (Dröder, 1999), b sensitivity to the strain rate at different locations
during the warm HMD, c sensitivity to the strain rate at different
locations during the warm sheet hydroforming,c; not sensitive (at
room temperature), d∼e; sensitive (at the elevated temperature).
WPNL2204
Warm hydroforming of lightweight materials
CCD Cameras
375
Die set
CCD
ARAMIS
Temperature
controller
Silicone based O-ring
Laser
sensor
Hydraulic
pump
Die set
Temperature controller
16.18 Warm hydraulic bulge testing setup.
rate is used because these materials have an intrinsically high strain rate
sensitivity, especially at elevated temperatures. Thus, it is very important to
have complete data on materials available at different temperatures and
strain rates for accurate analysis and design of the forming processes.
In order to construct such a material database, in addition and as an
alternative to tensile tests at elevated temperature levels, warm hydraulic
bulge testing systems were developed to characterize the material behavior
of sheet material blanks at different temperature and strain rate levels.
The hydraulic bulge test is preferred over tensile test method because it
represents a biaxial loading condition which is more relevant to metal
forming processes than the uniaxial loading condition as in a conventional
tensile test. The bulge test results enable construction of a material database
(flow curves) for each material at the selected testing temperatures and
strain rates.
16.7.1 Bulge test at elevated temperature levels
A warm hydraulic bulge testing apparatus, Fig. 16.18, is composed of four
major systems which are (1) a pneumatic/hydraulic system: pump (Hydratron AZ-2–180HPU-LW), pressure controller (Marsh Bellofram Type 3510),
and pressure transducer (OMEGA PX605), (2) a set of bulging die: upper
and lower die with a bulge diameter of 100 mm, clamping and sealing
mechanism (silicone based O-ring), (3) heating system: cartridge heaters,
temperature controller (OMEGA CN616tc1), and thermocouples (Type
K), and (4) in-die non-contact measurement systems: laser sensor (Keyence
LK-G402) and CCD cameras (GOM ARAMIS System by Trilion). The
non-contact measurement systems were used rather than a contact type
to avoid any temperature gradient due to the heat transfer at the contact
location at the dome apex, and to prevent the equipment damages from the
hot pressurized oil (Marlotherm SH) at bursting.
The process parameters of interest in this study are the temperature
and the strain rate. The temperature of each die half was monitored and
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Hydroforming for advanced manufacturing
Laser sensor
CCD cameras
Heaters
Thermocouple
Upper die
Temperature
controller
Hot oil
Lower die
Thermocouple
P
Lab VIEW
P transducer
Pump
P controller
16.19 Schematic diagram of the warm bulge test setup.
controlled independently using two separate sets of cartridge heaters and
thermocouples attached to each die half as shown in Fig. 16.19. With this
type of control loop, the temperature variation during the test was below
5 °C with respect to the set value. On the other hand, the strain rate (SR
or ė¯ ) was carefully controlled in a feedback loop with a PID controller.
Based on the difference between the pre-calculated dome height profile
(reference value) and the instantaneous dome height value from the laser
measurement, the control signals are sent to the pressure controller to
regulate the air input pressure and flow rate to be sent to the pump. The
pressure and flow rate of the discharge fluid (oil) at the pump outlet are
directly proportional to the air flow at the inlet. The connection loop of the
pneumatic/hydraulic system and the non-contact measurement system
(laser sensor) is also shown in Fig. 16.19.
The pre-calculated dome height (hd) profile for a constant strain rate
control was derived from the geometrical relationships in a circular bulge
testing of thin sheet blanks as follows:
t
ε = ln ⎛ 0 ⎞
⎝ td ⎠
2
⎛ d
⎞
td = t0 ⎜ 2 c 2 ⎟
⎝ dc + 4 hd ⎠
[16.1]
2
[16.2]
WPNL2204
Warm hydroforming of lightweight materials
Room temperature
100 °C
200 °C
377
300 °C
SR=0.0013 s–1
AI5052
SR=0.013 s–1
SR=0.0013 s–1
AI6061
SR=0.013 s–1
Room temperature
100 °C
200 °C
240 °C
16.20 Samples of hydraulic bulged specimens at several temperature
and strain rate levels.
where ē is the equivalent strain, t0 the initial sheet thickness, td the instantaneous apex thickness, dc the bulge diameter, and hd the instantaneous
dome height. In addition, since strain rate (ė¯ ) is the rate of change in strain,
we can write:
ē = ė¯ t
[16.3]
where t is time. Combining equation [16.1–16.3] gives a relationship between
the instantaneous dome height (hd) and the strain rate (ė¯ ) as:
hd =
dc
2
e ε t/ 2 − 1
[16.4]
During each test, the bulging pressure and dome height were continuously measured and recorded using the pressure transducer and the noncontact measurement systems (laser sensor and CCD cameras). These
pressure and dome height data were synchronized by the time stamp during
the test and later used for the flow curve calculation.
Two aluminum alloys (Al5052 and Al6061) were tested at different temperature and strain rate levels. Some of the bulged samples are depicted in
Fig. 16.20.
With the ARAMIS system, in addition to the dome height data, the
strain and strain distribution were also measured during the experiment as
depicted in Fig. 16.21. As expected, the maximum strain value was found to
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Hydroforming for advanced manufacturing
0.8
SR = 0.013 s–1
–1
Temperature = 200°C SR = 0.0013 s
Temperature = 100°C
0.7
Equivalent strain
0.6
0.5
0.4
.
e = Δe = slope
Δt
0.3
0.2
AI5052, t0 = 2.03 mm
0.1
Calculation
ARAMIS
0.0
0
200
400
600
Time (s)
16.21 Strain and strain distribution from ARAMIS system.
be at the dome apex and the lowest value at the blank edge. The strain rates
could also be determined from the slope of the plot in Fig. 16.21, which in
this specific case were calculated to be 0.0013 and 0.013 s−1, showing a very
accurate strain rate control in the test.
The calculation of the flow stress was carried out based on the measured
dome height (hd) and the bulging pressure (P) according to:
σ=
PR
2 td
[16.6]
R=
a2 + hd2
2 hd
[16.7]
where s̄ is the equivalent flow stress, R the curvature of the bulge radius, td
the apex thickness which can be calculated using equation (16.2), a half
bulge diameter (dc/2), and hd is the instantaneous height at the dome apex.
The equivalent strain (ē) calculation in equation (16.1) and the equivalent
stress s̄ calculation in equation (16.6) were combined to construct the material flow curves for different testing conditions as shown in Fig. 16.22.
WPNL2204
Warm hydroforming of lightweight materials
Equivalent stress (MPa)
500
Room-SR0.0013
100 °C-SR0.0013
200 °C-SR0.0013
300 °C-SR0.0013
600
Room-SR0.013
100 °C-SR0.013
200 °C-SR0.013
300 °C-SR0.013
500
Equivalent stress (MPa)
600
400
Corner
bursting
300
200
100
0
0.0
0.4
0.6
0.8
Equivalent strain
a
Room-SR0.013
100 °C-SR0.013
200 °C-SR0.013
240 °C-SR0.013
300 °C-SR0.013
400
300
Corner bursting
200
100
Corner
bursting
0.2
Room-SR0.0013
100 °C-SR0.0013
200 °C-SR0.0013
240 °C-SR0.0013
300 °C-SR0.0013
379
1.0
0
0.0
0.2
0.4
0.6
0.8
Equivalent strain
b
1.0
16.22 Equivalent stress–strain curves of a A5052 and b A6061 alloys at
different temperature and strain-rate conditions. Results for bulge
tests and tensile tests for the same alloys under the same conditions
are also given.
Since the assumption of the perfect spherical bulge shape, which was one
of the key assumptions in deriving the formula in equations (16.2) and
(16.7) for calculation of the apex thickness (td) and the dome height (hd), is
far from being true at the beginning of the bulge test, only the flow stress/
curve in the plastic region were calculated and presented in Fig. 16.22. The
results in Fig. 16.22 showed that the flow stress decreases with increasing
temperature and/or with decreasing strain rate. The effect of strain rate was
observed to be more pronounced at the elevated temperature levels. This
trend was rather consistent except in the case of Al5052 at 100 °C, where
the lower strain rate showed higher flow stress. Note that three replications
were used in this study for each testing condition.
Another interesting finding from the test results was in the case of
bulging Al6061 at 300 °C with 0.013 s−1 strain rate. All three specimens
tested at this condition were fractured at the die corner radius region
rather than bursting at the middle of the blank where the maximum
strain was expected. Thus, the flow curve for this case was not included
in the plot in Fig. 16.22 as the maximum strain was well below 0.2. A
similar phenomenon was also observed in the study by Novotny and
Geiger (2004) in their warm tensile and bulge tests of Al5XXX and
Al6XXX aluminum alloys. In their study, the formability of Al5XXX
WPNL2204
380
Hydroforming for advanced manufacturing
continuously increased with the temperature, while that of the Al6XXX
would increase up to a certain temperature (200 °C) after which the
maximum elongation started to decrease with increasing temperature.
This showed that there exists an optimal forming temperature for Al6061
beyond which the formability would be worse. However, this trend
appeared to be contradicted in the case of bulging Al6061 at 300 °C, but
with a lower strain rate of 0.0013 s−1. At this strain rate, the formability
was shown to be continuously improved with increasing temperature
even after 240 °C (maximum strain close to 1 and bursting occurred at
the dome apex).
Nonetheless, based on these test results, it was confirmed that, in general,
the formability of aluminum sheet alloys could be considerably enhanced
by forming the sheet at elevated temperatures or using low strain rates. The
material flow curves in Fig. 16.22 were constructed from the material database for the two Al alloys at temperatures between room temperature and
300 °C, and for strain rates between 0.0013 and 0.013 s−1.
16.8
References
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WPNL2204
Warm hydroforming of lightweight materials
381
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e. doege and k. dröder (1997) ‘Processing of magnesium sheet metals by deep
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k. g. dröder (1999), Untersuchungen zum umformen von veinblenchen aus magnesiumknetlegiereungen, PhD dissertation, University of Hannover.
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temperatures’, Trans. ASM, 36, 254.
h. friedrich and s. schuman (2001), ‘Research for a new age magnesium in the
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analysis of aluminum auto body panel for electric vehicle’, Japan. Soc. Automotive
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128, 622–633.
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deep drawing with uniform pressure onto the blank’, Int. J. Machine Tools Manuf.,
44, 649–657.
s. lee, y. h. chen and j. y. wang (2002), ‘Isothermal sheet formability of Mg alloy
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WPNL2204
382
Hydroforming for advanced manufacturing
d. li and a. ghosh (2003), ‘Tensile deformation behavior of aluminum alloys at warm
forming temperatures’, Mater. Sci. Eng. A, 352(1–2), 279–286.
j. lundqvist (2004), ‘Numerical simulation of tube hydroforming – adaptive loading
paths’, PhD Dissertation, Lulea University of Technology.
k. manabe and h. nishimura (1983) ‘Influence of material properties in forming of
tubes’, Bander Bleche Rohre, 9(5), 235.
k. manabe, s. mori, k. suzuki and h. nishimura (1984), ‘Bulge forming of thin-walled
tubes by micro-computer controlled hydraulic press’, Adv. Technol. Plast., 1,
279–284.
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to increase the deep drawability of aluminum 1050 sheet’, Int. J. Machine Tools
Manuf. 41, 1283–1294.
t. naka and f. yoshida (1999), ‘Deep drawability of type 5083 aluminum–magnesium
alloy sheet under various conditions of temperature and forming speed’, J. Mater.
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k. nakamura (1997), ‘Warm deep drawability with hydraulic counter pressure of
1050 Al sheets’, J. Japan Inst. Light Metals, 47(6), 323–328.
s. novotny and m. geiger (2004), ‘Process design for hydroforming of lightweight
metal sheets at elevated temperatures’, J. Mater. Process. Technol., 138, 594–599.
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p. ray and b. j. mcdonald (2004), ‘Determination of the optimal loading path for
tube hydroforming processes using a fuzzy load control algorithm and finite
element analysis’, Finite Elements Analysis Design, 41, 173–192.
w. rimkus, h. bauer and m. j. a. mihsein (2000), ‘Design of load curves for hydroforming applications’, J. Mater. Process. Technol., 108, 97–105.
d. schmoeckel, c. hessler and b. engel (1992), ‘Pressure control in hydraulic tube
forming’, Annal. CIRP, 41(1), 311–314.
d. schmoeckel, c. hielscher, r. huber and m. prier (1997), ‘Internal high pressure
forming at PtU’, PtU der Technischen Hochschule Darmstadt (in German).
l. b. shulkin, r. a. posteraro, m. a. ahmetoglu, g. l. kinzel and t. altan (2000),
‘Blank holder force (BHF) control in viscous pressure forming (VPF) of sheetmetal, J. Mater Process. Technol., 98, 7–16.
k. siegert, s. jager and m. vulcan (2001), ‘Pneumatic bulging of magnesium AZ31
sheet metals at elevated temperatures’, University of Stuttgart, Germany,
pp. 1–4.
k. siegert, s. jager and m. vulcan (2003), ‘Pneumatic bulging of magnesium AZ31
sheet metals at elevated temperatures’, Annals CIRP, 52(1), pp. 241–244.
k. siegert and s. jäger (2004), Warm forming of magnesium sheet metal, SAE 2004
World Congress and Exhibition, 2004-01-1043, March 2004, Detroit, MI, USA.
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WPNL2204
Warm hydroforming of lightweight materials
383
f. shehata, m. j. painter and r. pearce (1978), ‘Warm forming of aluminum/
magnesium alloy sheet’, J. Mech. Working Technol. 2, 279–290.
m. sugamata, j. kaneko, h. usagawa and m. suzuki (1987), ‘Effect of forming
temperature on deep drawability of aluminum alloy sheets’, Adv. Technol. Plast.,
1275–1281.
k. takata, t. ohwue, m. saga and m. kikuchi (2000), ‘Formability of Al–Mg alloys at
warm temp.’, Mater. Sci. Forum, 331, 631–636.
h. takuda, t. morishita, t. kinoshita and n. shirakawa (2005), ‘Modelling of formula
for flow stress of a magnesium alloy AZ31 sheet at elevated temperatures’,
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h. takuda (2002), ‘Finite element simulation of deep drawing of aluminum alloy
sheet with accounting for heat conduction’, J. Mater. Process. Technol., 120,
412–418.
h. takuda, s. kikuchi, k. kubota and n. hatta (1999), ‘Tensile properties and press
formability of a Mg–Li alloy sheet’, Proc. 6th ICTP, 1429–1434.
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of warm deep drawing of aluminum alloy sheet when accounting for heat conduction’, J. Mater. Process. Technol. 120, 412–418.
f. vollertsen (1999), ‘Warm forming with liquid pressure’, Personal communication
and presentation.
WPNL2204
Index
accessibility, hole creation and 203
accuracy, hole creation and 203
Active Hydromec 231
adaptive loading (AL) 137, 141
advanced high strength steels
(AHSS) 352, 353
aerospace industry see fluid cell
pressing
aluminum
applications 261–2, 361
lightweight 77, 342, 352
process chains 83
seam-welded 243–4
sheet 233, 371
tube bending 248–51, 253
tubes 190–1, 257–8, 259
aluminum alloys 110
high-strength 77, 325, 329
lightweight 16–17, 353, 354, 377
aluminum extruded alloys 81–3,
239–43
chemical composition limits 240
mechanical properties 241, 246–7
aluminum sheet alloys 243–4
chemical composition 245
aluminum-killed-draw-quality
(AKDQ) 100–3, 104, 112
angular springback 187–8
anisotropy 80
annealing 78, 79, 94, 276, 349
application, hole creation and 203
ARAMIS system 377–8
ASEA (company) 316
assemblies 289–90
costs 52
ASTM standards 100, 275
attachment points 349
Auto/Steel Partnership (A/SP) 96,
100–2, 104–5, 109
automobiles see lightweighting
automobiles
automotive industry 277
aluminum/magnesium and 238, 243,
257
applications 260–3
components 52, 77, 83, 333–4, 353
axial compression 107
axial cylinders 15
axial feeding 10, 130
design 135–6
difficulties 128, 129
internal pressure and 22, 137
THF 126–7
axial loads 64–5
axial punches 15
bearing area 239
bend
die 182–3
radius 98, 276, 310
bending see tube bending
bismuth/tin eutectic alloys 328
bladders 46
blankholder 234–6
force (BHF) 10, 144, 152, 223–7
passim, 370–2
384
WPNL2204
Index
blanks
double 25, 219, 221, 232–3, 356
single 217
size/testing 331
body parts 4
boost assist 98–9, 182–3, 192
boundary
conditions 156, 159–60, 162
lubrication 154, 156, 158, 162
surfaces 172–3
brass alloys 77
brehm shearing 211–12
bridge dies 239
Brinell test methods 87
buckling 65, 127, 146, 186
bulge forming of tubes (BFTs) 8
bulge tests 16, 80–1, 84, 85–6, 217–18
elevated temperature 360, 362,
375–80
bulging 130–1
pneumatic 359
burnthroughs, welding 349
burst (free-expansion) tests 105–7, 253
bursting 62–4, 80, 146
calibration pressure 133
CamSys Grid Pattern Analyzer
(GPA) 249
casting 354
centreline shape, tube 128–30
chassis parts 3
clamp die 182–3
clamping
actuators 36–7
devices 10–13
forces 13
closed-loop control system 44–5
closing force 135
CNC see computer numerical control
(CNC)
coating 10, 17–21
-based lubricants 161
cold hydroforming 357, 363–7
sheet 371
cold punch temperature 368
cold working 348
385
cold-forming process 77, 94, 186
computer experiments 138–9
computer numerical control
(CNC) 182, 187, 214
rotary draw bending 182
mandrel 276
tube benders 187
computer simulations 21–4, 114–15,
172–7
conditions
boundary 156, 159–60, 162
dry 154
hydrodynamic 155
process 121–2
cone test 87, 100
confidentiality 13
conical tubes 24
contact stress levels 17
continuous casting (CC) 243
Continuum Theory of Plasticity 54–5,
58, 62
contraction, surface 148–9
controls 43–5
process 66–9, 367–73
valves 42
conventional deep drawing
hydraulic stretch forming and 232–4
vs hydromechanical deep
drawing 222–7, 230
cooling 36
copper alloys 77
corner
cross-section 281–3, 295, 303, 307
fill test 84, 86–7, 168–9
floating 372–3
radius 281–3, 357
cost effectiveness
clamping forces 13
cross-section shaping 343
lubricants 148
production 24
thin wall castings and 353
costs 8
advantage 1
annealing processes 78
assembly 52
WPNL2204
386
Index
comparative 288–9
efficiency and 319
fasteners 349
fluid cell pressing 334
heat treatment 348
hole creation 203
hydroforming 291
laser cutting 212
lightweighting 337–8, 347
lubricant 20
materials 3
power 312
presses 235
process 305–6
production 186
punch units 308
reductions 4, 5, 9, 27, 243
seamless tube 275
stampings vs roll forming 290
THF 267
tooling 261, 316–17, 325
tubing 274
vs performance 287
Coulomb friction law 19, 69, 144, 146,
164
counter cylinder limits 15
counter pressure 224, 227–30, 360
counterpunch 15, 124–6, 131
countersink 205–8
cracking 307, 331, 345
cross-section
corners 281–3, 295, 303, 307
expansion 340
hole-piercing 282–4
ovality 190–1
pre-forming 277–85
shape 127–8, 279–81, 343–5
size 342–3
crushing 128, 135, 193–7
see also pre-forming
cycle time 312–13
cylinders 15, 37
damage, incidental 350
deflection 340–1
deformation 98
deformation in THF 52–74
forming limits 62–6
preceding forming operations
69–74
process control 66–9
stress and strain 54–62
denting 350
design
axial feeding 135–6
dies 204
flexibility 304–5
hole creation and 203
joint 339
loading paths 23–4
parts 197–9
principles 351
process 367–73
Design and Analysis of Computer
Experiments (DACE) 13, 139
Design of Experiments (DOE)
techniques 138
design in THF 121–41
decision sequence 140–1
example 139–41
FEA strategies 136–9
guidelines for 131–6
hydroformability and 127–31
periphery 278–9
process classifications 123–7
die and mandrel process 239
die-workpiece interface 144, 145
dies 326, 330, 332
bridge 182–3
clamp 182–3
design 204
expansion 326–7, 330
hydromechanical deep drawing
234–6
port hole 239
pressure 182–3
wiper 182–3
dimensional stability 261
direct-chill (DC) casting 243
distortion, welding 214–15
docking rod sealing styles 49
documentation 331
WPNL2204
Index
double blanks 25, 219, 221, 232–3, 356
double tubing 24
double-acting hydraulic press 34–5
draw ratio 323–5, 359–60
drawn over mandrel (DOM) tube
275
dry conditions 154
dry lubricants 20
dual-phase (DP) steel 271–2, 341–2
ductile materials 94, 323–5
electrical resistance welding
(ERW) 95, 275
electrohydraulic forming 27
electromagnetic force forming
(EMF) 27
elevated temperature 373
flow stress and 369, 374–5
hydroforming at 361–2
THF and SHF at 361–2
tube bending and 256–7
see also warm hydroforming
end cutting 209–12
inside shearing 211–12
laser cutting 212
outside shearing 210–11
saw cutting 209–10
end feeding 192, 345
energy consumption 312
engine components 4
Engineering Research Centre (Ohio
State University) 166
environment 352
exhaust system parts 3
expansion 80, 135
-driven THF 163
cross-section size and 344
die 326–7, 330
high-pressure 302–3, 303–4, 306–7
low-pressure 300–1
maximum tolerable 133–4
process 161
surface 148–9
test 84
tube section 284–5
zone tribo tests 168–9, 171–2
387
explosive forming 5
extended stress-based forming limit
criterion (XSFLC) 113
extrusion 81–3, 354
processes 95, 248
extrusion alloys 244, 246–8
5xxx 242–3
6xxx 239, 241–2
failures 146–7
feed actuators 48–50
feed-driven THF 163
lubrication 155–9
filtration 36
Finite Element Analysis (FEA) 21–3,
38, 50, 198, 303
lightweighting and 340, 345, 346, 350,
360, 363, 370–2
for THF design 131, 135, 136–9
tribological aspects and 169, 173–4,
176
Finite Element Method (FEM) 54–5,
59–61, 64–5, 67, 70, 72, 74
Process Simulations 219, 350,
371
simulations 122, 126, 132
flange temperature 368
flare test 100
flattening test 84
flexforming 8
floor space 312, 313
flow curves 217
flow stress 373, 374, 378–9
fluid
handling 40, 42–3
pressure design 135–6
fluid cell pressing 315–34
aerospace metals 324
cycle 319
ductile materials 323–5
parts 325–6, 330–1
process of 317–19, 320
recent developments 319–23
stages of 320
technology evolution 315–17
tooling 326–30
WPNL2204
388
Index
forces
blankholder (BHF) 10, 144, 152,
223–7 passim, 370–2
counter pressure related 227–30
friction 164, 222, 225, 227
punch 222–3
formability testing 84–7, 131
principles of 84
see also tube formability
forming
stages 194
strain 251–2
technologies 353–4
zone 184
forming limit curve (FLC) 64, 86, 108–
16 passim, 198–9, 217
stress space (SFLC) 111–13, 114–15,
115–16, 117
forming limit diagram (FLD) 64, 108–
9, 110, 198–9, 253–4
forming limits, tube 62–6, 108–13
process window programs
(PWDs) 110–11
strain-based 108–9, 109–10
stress-based 111–13
four-column hydraulic press 33–6
fracture 127, 128
risk 129
free-expansion (burst) test 105–7, 253
friction 10, 17–21, 63, 129, 248, 360
coefficient 164, 167, 169, 170
Coulomb law 19, 69, 144, 146, 164
effect 365
forces 164, 222, 225, 227
variables 148–53
zones 151–3
full film lubrication 155
fuzzy control 371–2
galvanneal steel 271
Garvin, Martin (first patent) 5
gas forming
aluminum tubes 257–8, 259
magnesium tubes 258–60
gas hydroforming 41
gas metal arc welding (GMAW) 213
German Institute for Standardization
(DIN) 52
guided profile bending 97
guiding zone 156, 158, 159
tribo tests 164–8, 169–71, 172
hand forming 315
hardness tests 84, 87
hardwood, laminated
resin-impregnated 328
heat treatment 348, 354
retrogression 242
HEATforming 25–6, 298, 358
high-frequency (HF) welding 79, 95
high-pressure sheet hydroforming
(SHF) 7–8, 358
high-pressure tube hydroforming
(HPH) 267–85 passim, 292–302
passim, 340, 344, 347
advantages 302–8
comparative process factors 308–13
high-strength aluminum alloys 77, 325,
329
high-strength steels (HSS) 271–2, 302–
3, 341–2, 348, 350
advanced (AHSS) 352, 353
low-alloy (HSLA) 96–7, 271–2, 302,
309, 329
holes
geometry 203
-piercing 282–4, 307–8, 332
self-threading 349
hot forging 354
hot forming (HF) 298
hot spots 340, 346
hydraulic actuator 39–40
hydraulic bulge tests 16
hydraulic cylinders 15
hydraulic medium 30–1
hydraulic presses 10–13, 33–9
fluid cell 317–19
hydraulic power supply (HPS) 36–7
hydromechanical deep drawing
234–6
sizing 37–9
styles of 33–6
WPNL2204
Index
hydraulic pressure 370, 371–2
forming (HPF) 8
hydraulic stretch forming 217–22
conventional deep drawing and
232–4
Hydro Aluminum (company) 244
HydroDieForming (HDF) 268, 269,
271, 272, 279
hydrodynamic conditions 155
hydrodynamic lubrication 156, 158
hydroforming 194
applications 2–4, 260–3
comparisons 310
defined 1
developments in 24–7
disadvantages 261
future trends 260–3
history 5
materials for 15–17
principle of 53
systems see systems, hydroforming
tooling features 14–15
tribology in 17–21
types/classifications 5–10
warm 25–7
Hydroforming Materials and
Lubricants task force 96
hydromechanical deep drawing
(HMD) 6–7, 145, 231, 371
counter pressure in 227–30
dies/presses 234–6
vs conventional deep drawing 222–7,
230
warm 358–60, 373
hydromechanical forming 25
hydropiercing 202–9
inward 205–9
outward 209
preliminaries 202–4
processes 204–5
hydrostatic lubrication 155, 156, 158
hydrostatic pressure forming 8
Industrial Research and Development
Institute (IRDI) 168
inside shearing 211–12
389
Institute for Metal Forming Technology
(IFU) 234–5
intensifiers, pressure 39–41
interface
die-workpiece 144, 145
pressure 148, 149, 153, 157
Interlaken 5000-kN press 257
internal high-pressure forming
(IHPF) 8
internal pressure 10, 22, 99, 106–7, 137,
299, 357
pierced hole 284
inward hydropiercing 204–5
joints
design 339
eliminating 339
Keeler-Brazier prediction 109
kicker cylinders 37
kriging approximation 139
lasers
cutting 212
welding 80–1, 88, 95, 213, 214, 275,
349
Levi stress correlations 57–8
lightweight materials 352–80
advantages/disadvantages 352–3
alloys 27
characterization 373–80
cold hydroforming and 363–7
forming technologies for 353–4
process design and control 367–73
structures 4
warm hydroforming and 354–63,
363–7, 367–73, 373–80
lightweighting automobiles 335–51
difficulties 336–7
limitations 348–50
methods of 337–8
motivation for 352
THF and 338–48
limited dome height (LDH) 217
limiting drawing ratio (LDR) 26–7,
354, 360, 363
WPNL2204
390
Index
limits see forming limits
linear variable displacement
transducers (LVDT) 107
liquid bulge forming (LBF) 8
liquid impact forming (LIF) 8,
295–7
loading control 370–3
loading paths (curves) 13–41 passim,
153
design 23–4
THF and 175
longitudinal strain 103
longitudinal yield stress 102
low carbon steel 108–9, 302
low-pressure tube hydroforming
(LPH) 267–86, 293–5
comparative process factors 304–8
passim, 308–13
cross-section pre-forming 277–85
disadvantages 299–302
HydroDieForming (HDF) 269, 271
lightweighting and 341–2, 344,
347
part characteristics 271–7
sequenced pressure (SPH) 267–8,
268–9, 270
lubricants 10, 17–21, 160–1
attributes 147–8
cost effectiveness 148
development 163–72
evaluation 164–9, 169–72
formulation 164–5
metal-forming 163, 165
pocket volume 176
test matrix 169
tribo tests 169–72
lubrication mechanisms 154–63
expansion-driven THF 159–61
feed-driven THF 155–9, 163
SHF 162–3
THF 155
tool 330
variables 148–53
machining 354
magnesium 83, 352
alloys 353, 354, 373
applications 262–3
tubes 255–7, 258–60
magnesium extruded tubes 244, 246–8
tensile properties 246
mandrel
balls 190–1
die and 239
rotary draw bending 182–3
manganese 247
Mann-Kendall (M-K) analysis 113
Marciniak-Kuczynski (M-K)
analysis 110
mass reduction 260–1
materials
characterization 373–80
ductile 323–5
properties 107, 150, 273, 323
raw 16, 353
reducing joint 339
required characteristics 15–17
rotary draw bending behavior 183–
93
specifications 309
strength 348
type 271–3
see also lightweight materials
materials in THF 77–90
aluminum 81–3
formability testing 84–7, 131
future trends 87–90
magnesium 83
steel 78–81
mechanical lock style press 37
mechanical properties 96, 99–106
passim, 109
mechanical tubeforming 267
melt-spinning 354
Membrane Theory 54–5, 58, 62
metal
aerospace 324
alloys 94
foams 90
forming processes 288
see also specific metals and alloys
metal inert gas (MIG) 213
WPNL2204
Index
microplasto hydrodynamic lubrication
(MPHDL) 156–7, 158–9, 161,
162
microplasto hydrostatic lubrication
(MPHSL) 156–7, 158–9, 161,
162
mild steel 325, 342, 353
mixed layer lubrication 155, 156, 158
multiple tubing 24
multi-stage forming process 94–9
necking 62–4, 99, 108
down 186
neutral axis 183–6, 189, 192
nickel-based alloys 325
numerical simulations 21–4, 105, 114–
15, 172–7
Ohio State University 166, 168
oil-canning effect 349–50
open-loop clamp control system 43–4
optical strain measurement system 360
optimization 138–9
original equipment manufacturers
(OEMs) 316, 319
outside shearing 210–11
outward hydropiercing 204–5
parts
aerospace 325–6, 330–1, 347
body 4
characteristics 271–7
chassis 3
complex 304
consolidation 260
design 197–9
exhaust system 3
hole-piercing 282–4, 307–8
large 299
lighter 308, 347
material properties 273
material type 271–3
rigidity 339, 348
series production 333
simple 299–300
stability 303–4
391
strength 348
structural integrity 204
tubes 273–4, 274–5, 275–7
pastes 20
pear-shaped expansion test 168–9,
171–2
periphery design 278–9
constant 278
HydroDieForming (HDF) and 279
variable 278
pinching 124, 268, 292, 340, 344
Pines rotary draw bending
machine 256
planned optimization 138–9
plasticity theory 111
plastics 328
plumbing 42
pneumatic bulging 359
port hole dies 239
power supply (HPS), hydraulic 36–7
power train components 4
prebulging 231
see also hydraulic stretch forming
preceding forming operations
69–74
pre-forming 10, 73–4, 181–200, 310
cross-section 277–85
crushing 128, 135, 193–7
material behavior 183–93
part design and tube
formability 197–9
phase 123–4
pre-flattening 193–7
principles of 74
rotary draw bending 182–3
presses
costs 235
double-acting hydraulic 34–5
features of 13
Interlaken 5000-kN 257
single-action 34–5, 235–6
straight-sided hydraulic 33–6
warm gas forming 258
see also hydraulic presses
press-forming 95
pressure
WPNL2204
392
Index
counter 224, 227–30
die 182–3
interface 157
internal see internal pressure
levels 317
pressure intensification systems 15,
39–43
control valves/plumbing 42
fluid handling 40, 42–3
intensifiers 39–41
pressure sequence hydroforming
(PSH) 181, 197
process
capabilities, limited 300
classifications, THF 123–7, 136
comparative factors 308–13
conditions/variables 121–2, 301–2
control 66–9, 367–73
costs 305–6
expansion 161
FEA optimization 138–9
flow diagrams 308–9
sequence 134–5
stability 302
window diagrams (PWDs) 110–11,
128
process design 367–73
fluid pressure/axial feed 135–6
guidelines for 131–6
production
costs 186, 353
plant analysis 115–16
programmable logic controller
(PLC) 43
protrusions 130–1
bulging 130–1, 359
height 172
see also bulge tests
punch 15, 235–6
axis 47
speed 370, 371
temperature 359, 368–9
tool 326, 329, 330
units 308
wall 372–3
push-assist 193
radial springback 189
rake angle 182–3
rapid feasibility evaluation 132–4
raw materials
characteristics of 16
costs 353
recrystallization temperatures 373
repeatability, hole creation and 203
response surface method (RSM) 138
retrogression heat treatment 242, 253
reverse drawing 231, 233
ring tensile test 100
ring-hoop tension (RHT) test 103–5
Rockwell test methods 87
roll bending 97
profile 97
roll forming 85, 95–6, 105, 289
tubes 79–81
room temperature
flow stress and 369, 374–5
formability and 353, 354
hydroforming at 259, 361–2
THF and SHF at 361–2
tube bending 255–6, 259
rotary draw bending 71–2, 97–8, 182–3,
276
CNC 182, 276
material behavior 183–93
process 248–9, 256
rubber forming 5
rubber pad
pressing 315–16
tools 321
safety components 4
saw cutting 209–10
sealing 10, 15
seamless tubing 95, 275
seam-welded tubes 243–4
5xxx sheet alloys 244
6xxx sheet alloys 244
segment-elastic blankholder 235–6
self-feeding (SF) approach 137
sequenced-pressure hydroforming
(SPH) 267–8, 268–9, 270
sequential optimization 138–9
WPNL2204
Index
sheet forming technology 59
sheet hydroforming (SHF) 5–8, 144,
145, 338
cold state 370, 370–1
elevated temperature 361
friction zones 152
high-pressure 7–8, 358
hydromechanical deep drawing dies/
presses 234–6
lubrication mechanisms and
162–3
processes 216–34
room temperature 361
tooling 45–7
SHF see sheet hydroforming (SHF)
shimmy shearing 211–12
simulations, numerical 21–4, 105, 114–
15, 172–7
single-action press 34–5, 235–6
sliding velocity 150–5
slugs 203, 282–4
soaps 20
splitting 99, 115, 250
spot welding 213, 214
spray deposition 354
springback 186–9, 303, 322, 353
elastic 74, 82
stability
dimensional 261
parts 303–4
process 302
stainless steel 16–17, 42, 138, 271,
325
alloys 341
stamping 1, 289–90, 307, 315
statistical process control (SPC) 45
steel 77, 78–81, 89, 198–9, 233
alloys 271–2
dual-phase (DP) 271–2, 341–2
low carbon 108–9, 302
mild 325, 342, 353
process chains 80
tubes 190–1, 274
see also high-strength steels (HSS)
Stoughton’s general procedure 116
straight-sided hydraulic press 33–6
393
strain 183–4, 198–9
analyses 84, 86
curve 191–2
distribution 98, 116, 248–51, 365–7
forming 251–2
longitudinal 103
stress and, in THF 54–62
strain rate 373, 374–5, 378–80
aluminum alloys 377
strain-rate, -dependent flow 369
strain-based forming limits 108–9
path-dependency of 109–10
stress 183–4, 226
attachment points 349
distribution 99, 303, 364
reassessment 340–1
-strain curves 96, 100–1, 105
strain and, in THF 54–62
stress-based forming limit curve
(SFLC) 111–13, 114–15, 115–16,
117
extended (XSFLC) 113
stretch forming, hydraulic 217–22
Stribeck curve 154
stripout torque 349
structures 4
integrity 204
performance 260–1
subsurface deformation 176
superplastic forming 5
surface
evolution 157, 161
expansion 148–9, 175
quality 146
strain 175
systems, hydroforming 10–15,
33–51
controls 43–5
further information 51
future trends 50
hydraulic presses 33–9
pressure intensification 39–43
tooling 45–50
tailor-welded tubes 24
tapered tubes 24
WPNL2204
394
Index
temperature
aluminum alloys 377
distribution 368–70
recrystallization 373
see also elevated temperature; room
temperature
temperature distribution
blank material 371
tooling and hydraulic medium
370–1
warm hydroforming 364–6
tensile strains 98
tensile tests 16, 84–5, 96, 100–3
warm 88–9
testing methods 17–21, 84–7
blank 331
bulge 16, 80–1, 84, 85–6, 217–18, 360,
362, 375–80
burst (free-expansion) test 105–7,
253
lubricants 164–9
pear-shaped expansion 168–9, 171–2
ring-hoop tension (RHT) 103–5
tensile 16, 84–5, 88–9, 96, 100–3
Theory of Shells 54, 58–9
thermomechanically coupled
simulation 360
THF see tube hydroforming (THF)
thickness distribution 365–6
thinning 127, 331, 372
distribution 256–7, 365
wall 344–5, 350
throughput, hole creation and 203
TIG welding 275
titanium 325
tooling 13–15, 45–50
costs 261, 316–17, 325
fluid cell 319–23
fluid cell pressing 326–30
lubrication 330
material 328–9
selecting 329–30
SHF styles 45–7
split line 47–8
temperature distributions 370–1
THF layout 124–6
THF styles 47–50
types 326–8
transportation industry 352
Tresca’s flow hypothesis 218
tribological aspects 17–21, 144–77
die-workpiece interface 144, 145
friction laws 144, 146
friction/lubrication/wear
variables 148–53
hydroforming failures 146–7
lubricants 147–8, 154–63, 163–72
numerical simulation 172–7
trimming 331–2
TRIP steel 271–2, 341–2
tube bending 134, 194
aluminum 248–51
comparisons 309–10
effect of 10
influence of 97–9
LPH 275–7
magnesium 255–7
material thickness and 70–3, 340
strains 98
terms and forming zone 184
tooling for 249
tube formability 93–117
future trends 117
hydroforming and 99
measuring 99–107
metal alloys 94
numerical simulation analysis
114–15
part design and 197–9
production plant analysis 115–16
tube bending and 97–9
tube making and 95–7
see also forming limits, tube
tube forming 289, 290–8
comparative costs 289
tube hydroforming (THF) 5, 8–10, 144,
145, 299
advantages and disadvantages 291
classification 123–7, 136
cold state 370
computer simulations 21–4
costs 267
WPNL2204
Index
elevated temperature 361–2
lightweighting and 338–48
lubricants and 155–9, 159–61,
163–72
press, features of 13
processes 174–6, 291–8
room temperature 361–2
stress and strain in 54–62
tooling 47–50, 124–6
see also deformation in THF; design
in THF; high-pressure tube
hydroforming (HPH); lowpressure tube hydroforming
(LPH); materials in THF; tube
formability
tubes
-section expansion 284–5
diameter 273–4
making 95–7, 274–5
mill process 243
specifications 309
tubular blanks 95
University of Darmstadt 164
University of Paderborn 167
University Stuttgart 235
Vari-Form Inc. 181
variables
process 121–2
wear 148–53
vehicles
crash performance 341
operating characteristics 335
weight 9, 181
Vickers test methods 87
von Mises stress correlations 57–8
von Mises Yield Criterion 56, 208
wall thickness
control 345–7
deformation mechanism and 55–8,
59–60, 62–4, 66–8, 70–1, 74
design and 128, 132
formability and 95, 97, 98–9, 100, 102,
105, 107
395
material characteristics and 79–80,
88
minimizing 340–1, 342–3
strain distributions 191–3
tube diameter and 273–4
uneven 293
wall thickness distribution 310–11
lubricants and 171–2
wall thinning 344–5, 350
warm bulging test 360, 362,
375–80
warm deep drawing 360
warm forming 364, 366–7
gas 257–60
warm hydroforming 25–7, 257–60
challenges 355–6
materials characterization
373–80
process design/control 367–73
review 354–63
tooling and products 358
vs cold hydroforming 363–7
warm hydromechanical deep drawing
(warm HMD) 358–60, 373
waxes 20
weak spots 105–7
wear 17–21
variables 148–53
weight 4, 5
car components 77
tubing and 274
vehicle 9, 181
see also lightweight materials;
lightweighting automobiles
weld
annealing 349
seam 52, 95–7, 100, 106
welding 212–15
burnthroughs 349
distortion 214–15
eliminating 339
gas metal arc (GMAW) 213
laser 213, 214
spot 213, 214
tubes 80–1
wet lubricants 20
WPNL2204
396
Index
WPNL2204
wiper die 182–3
work hardening 77, 94–9 passim, 105,
107, 186, 341
wrinkling 146, 250
aluminum sheet 371
comparative techniques and 186, 193,
231, 293
deformation mechanism and 64–5,
74
design and 127, 137
fluid cell pressing and 322, 331
lightweighting and 353, 372
yield strength 88
yield stress 67–8, 99, 103, 105
longitudinal 102
Young’s modulus 188
zinc 247
alloys 329
WPNL2204