!!Investigations on forming of aluminum 5052 and 6061 sheet alloys at

Materials and Design 31 (2010) 2422–2434

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Materials and Design

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Investigations on forming of aluminum 5052 and 6061 sheet alloys atwarm temperatures

S. Mahabunphachai a,b, M. Koç a,*

a NSF I/UCRC Center for Precision Forming (CPF), Virginia Commonwealth University (VCU), Richmond, VA 23284, USAb National Metal and Materials Technology Center, Pathumthani, Thailand

a r t i c l e i n f o

Article history:Received 1 September 2009Accepted 23 November 2009Available online 26 November 2009

Keywords:Aluminum sheetFormabilityLightweight materialHydroformingWarm formingAA5052AA6061

0261-3069/$ - see front matter � 2009 Elsevier Ltd. Adoi:10.1016/j.matdes.2009.11.053

* Corresponding author. Tel.: +1 804 827 7029.E-mail address: [email protected] (M. Ko

a b s t r a c t

In an ongoing quest to realize low-mass transportation vehicles with enhanced fuel efficiency, deforma-tion characteristics of Al5052 and Al6061 were investigated. In the first part of this study, material behav-ior of Al5052 and Al6061 sheet alloys were investigated under different process (temperature and strainrate) and loading (uniaxial vs. biaxial) conditions experimentally. With the biaxial, hydraulic bulge tests,flow stress curves up to 60–70% strain levels were obtained whereas it was limited to �30% strain levelsin tensile tests. The microstructure analysis showed that the change of grain size due to the effects of ele-vated temperatures and strain rates were not significant; therefore, it was concluded that the decrease inthe flow stress at high temperature levels was mainly due to the thermally activated dislocation lines. Inthe second part, the effect of the temperature and the pressure on the formability was further investi-gated in a set of closed-die warm hydroforming experiments. The test results showed that a linearlyincreasing pressure profile up to �20 MPa levels did not have a significant effect on the die filling ratiosand thinning of the parts when a uniform temperature distribution of 300 �C was applied. Finally, in thethird part of the study, finite element models were developed for the same closed-die hydroforminggeometry using the material behavior models obtained from bulge and tensile tests. Flow stress curvesobtained from tests were compared in terms of predicting the cavity filling ratios and thinning profilesfrom the experiments. Based on the comparison, it was revealed that flow stress curves obtained fromthe warm hydraulic bulge tests provided accurate predictions at high strain levels (i.e., �e > 0:4, when partfilling is above 80%) while the flow stress curves from the tensile tests did so at low strain levels (i.e.,�e < 0:2, when cavity filling is below 80%). On the other hand, comparison of thinning values indicatedthat flow stress curves from bulge tests yielded good agreement with the experimentally measured val-ues in general. Therefore, it can be recommended that the bulge test results should be used wheneveravailable in order to conduct accurate numerical analyses for warm sheet hydroforming where complexgeometry and loading conditions exist.

� 2009 Elsevier Ltd. All rights reserved.

1. Introduction

With an increasing awareness and effects of global warming,and the scanty fossil fuel resources left when compared to the everincreasing demand of oil, car manufacturers have been seeking foralternative and sustaining solutions to the fuel efficiency problem.Many believe that the next generation cars must run on alternativeand clean fuels (e.g., hydrogen via fuel cells) to prevent further in-crease of harmful emissions. However, this approach appears to bemore of a solution that may not be practically and economicallyrealized in short term (i.e., �10 years). On the other hand, anotherprominent approach that is sustaining, effective, and sooner wouldbe the realization of low-mass vehicles. In the pursuit of latter

ll rights reserved.

ç).

solution, the car manufacturers along with various research groupshave been investigating the fabrication of structural and body partsout of lightweight materials such as aluminum and magnesium al-loys [1–3]. On the other hand, despite the obvious advantages ofthe lightweight alloys, they have a notable drawback in that theirformability is significantly lower than traditional steel alloys atroom temperature conditions, which is usually caused by the highalloy percentages that are required for high strength [4,5]. Forexample, the formability of aluminum alloys is only about two-third of a deep drawing steel grade, their Young’s modulus is aboutone-third of the steel, which in turn causes higher susceptibility ofwrinkling and springback [6], and their elongation is about half ofsteel’s [7]. The inferior formability of aluminum alloys makes itmore difficult and expensive to use them in mass production ofstructural and body parts, which requires high levels of elongationand ductility to be formed into complex shapes. Nevertheless, the

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CCD Cameras

Hydraulic Pump Die Set

Temp. controllerARAMIS

CCD

Laser Sensor

Die Insert

Silicone based O-ring

Fig. 1. Warm hydraulic bulge test setup.

S. Mahabunphachai, M. Koç / Materials and Design 31 (2010) 2422–2434 2423

formability of the aluminum alloys has been shown to increasewith an increasing forming temperature up to the recrystallizationtemperature, e.g., 300 �C for Al5xxx, and 200 �C for Al6xxx, whereadditional sliding planes are activated in the material [3–6,8–10].Selective and localized heating strategies on the forming dies,causing an inhomogeneous temperature distribution on the blank,were also shown to further enhance the formability of the alumi-num alloys [9–13]. In addition, high elongation could be obtainedwhen low strain rate is used because these materials have intrinsi-cally high strain rate sensitivity, especially at elevated temperaturelevels [7,14,15].

In addition to the forming at elevated temperature levels, alter-native process technologies have been investigated to be used forcomplex and consolidated part manufacturing. The hydroformingprocess has been used for an increasing number of structural andbody applications as it enlarges the forming limit windows ofmaterials due to the biaxial and frictionless loading conditions,by which necking or thinning are delayed, and thus, elongationlimits are extended [4,5,10,13]. The hybrid warm hydroformingprocess combines the advantages of both warm forming andhydroforming [12,16]. However, it is still considered as a relativelynew and unknown technology waiting to be proven and validated.There are two vital aspects of this hybrid technology that demandfurther investigations: (1) understanding and characterization ofthe material behavior under warm hydroforming conditions, and(2) determination of the optimal process parameters (i.e., temper-ature level and distribution, pressure and blank holding profiles).

In this study, our objectives were to: (1) determine properexperimental methodologies to accurately characterize the mate-rial behavior under warm hydroforming conditions (i.e., comparewarm tensile and warm hydraulic bulge tests), (2) experimentallyunderstand and quantify the effects of process parameters, such astemperature and internal pressure, on the part formability intorepresentative die cavities with reasonably complex geometries,and (3) develop finite element models (FEM) to determine theapplicability of bulge and tensile test findings to accurately predictthe part formability. For this purpose, two commonly used alumi-num alloys (5052 and 6061) were selected for experimentation.

In the next section, experimental setup and conditions for thewarm tensile and warm hydraulic bulge tests are presented. Inthe third section, material test results are presented and comparedin terms of achievable strain and stress levels (i.e., flow stresscurves). In the fourth section, experimental conditions and resultsof a design of experiment (DOE) study conducted using a set ofclosed-dies are presented and discussed to quantify the effect ofprocess parameters on the cavity filling (i.e., formability) in warmhydroforming. Part profile, die filling ratio, and thinning on thewarm hydroformed parts were reported and compared. A regres-sion analysis was conducted to reveal the significance of pressureand temperature parameters. In the fifth section, a finite elementmodel (FEM) of the warm hydroforming process is developed andvalidated by comparing the predictions with the experimental find-ings. The flow stress curves obtained from both the tensile andbulge tests are used in the FEA validation in order to determinewhich set of material test data is more accurate in predicting thepart formability (i.e., cavity filling and thinning). Finally, a summaryof the results and conclusions are presented in the sixth section.

2. Material characterization experiments

2.1. Hydraulic bulge test setup

For warm hydraulic bulge tests, a specially designed and builtsystem composed of four major sub-systems was used as depictedin Fig. 1: (1) a pneumatic/hydraulic system: pump (Hydratron

AZ-2-180HPU-LW), pressure controller (Marsh Bellofram Type3510), and pressure transducer (OMEGA PX605), (2) a set of bulg-ing die: upper and lower die with a bulge diameter of 100 mm andeach with a die corner radius of 6.5 mm, clamping and sealingmechanism (silicone based O-ring and copper O-ring), (3) heatingsystem: cartridge heaters, temperature controller (OMEGACN616tc1), and thermocouples (Type K), and (4) in-die non-con-tact measurement systems: laser sensor (Keyence LK-G402) andstereoscopic system (two CCD cameras with GOM ARAMIS systemby Trilion). The non-contact measurement systems were used toavoid any temperature gradient at the contact location, whichcan influence the material behavior [4]. In order to avoid damageon the laser sensor and CCD cameras due to the splashing of thehot pressurized oil (Marlotherm SH), a thick glass was placed onthe housing roof.

The process parameters of interest in this test were the effect oftemperature and strain rate on the flow stress behavior of thematerials. The temperature of each die half was monitored andcontrolled independently using two separate sets of cartridge heat-ers and thermocouples (t/c) attached to each die half as shown inFig. 2. With this type of control loop, the temperature variationduring the test was below 5 �C. The heating cycle was made asshort as possible, where the cycle time depends on the set temper-ature value. On the other hand, a holding time of 5–10 min wasused to allow the uniform temperature distribution on the blankand the oil in the die cavity. The strain rate (SR or _e) was also con-trolled using a feedback loop with a PID controller. Based on thedifference between the pre-calculated dome height profile (refer-ence value) and the instantaneous dome height value from the la-ser measurement, the control signals were sent to the pressurecontroller to regulate the air input pressure and flow rate at thepump inlet. The pressure and flow rate of the discharge fluid (oil)at the pump outlet were directly proportional to this controlledair flow at the inlet. The schematic of the control loops of the pneu-matic/hydraulic system and the non-contact measurement system(laser sensor) is presented in Fig. 2.

Temp. Controller

heatersCCD Cameras

Hot oil

P

P TransducerLabVIEW

Upper die

Lower die

t/c

Laser sensor

P controller Pump

t/c

Temp. Controller

heatersCCD Cameras

Hot oil

P

P TransducerLabVIEW

Upper die

Lower die

t/c

Laser sensor

P controller Pump

t/c

Fig. 2. Schematic diagram of the warm bulge test setup.

2424 S. Mahabunphachai, M. Koç / Materials and Design 31 (2010) 2422–2434

A pre-calculated dome height (hd) profile was used to obtain aconstant strain rate during the tests. It was derived based on thegeometrical relationships in a circular bulge testing of thin sheetblanks as follows [17]:

e ¼ lnt0

td

� �ð1Þ

td ¼ t0d2

c

d2c þ 4h2

d

!2

ð2Þ

where e is the equivalent strain, t0 is the initial sheet thickness, td isthe instantaneous apex thickness, dc is the bulge diameter, and hd isthe instantaneous dome height. In addition, since strain rate ð _eÞ isthe rate of change in strain, one can write:

e ¼ _e � t ð3Þ

where t is time. Combining Eq. (1)–(3), a relationship between theinstantaneous dome height (hd) and the strain rate ð _eÞ can be ob-tained as:

hd ¼dc

2

ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffie _e�t=2 � 1

pð4Þ

The relationship in Eq. (4) was used for plotting the reference hd

profile as a function of time. Typical dome height profiles at various

0

10

20

30

40

50

60

time (s)

Dom

e he

ight

, hd

(mm

) 0.13 s-1

0.013 s-1

0.0013 s-1

0 100 200 300 400

Fig. 3. Typical dome height profiles for different strain rate levels.

strain rate levels are plotted in Fig. 3. These profiles were used as areference input signal in the feedback control loop.

With this test setup, the bulging pressure and dome heightcould be continuously measured and recorded using the pressuretransducer and the non-contact measurement systems (laser sen-sor and CCD cameras) during the test. Curvature of the bulgingwas also measured using the ARAMIS system with the CCD cam-eras. However, the dome height was found to be the same withthe laser measurements as explained in detail in another study[18]. These pressure and dome height data were later synchronizedtogether by the time stamp, and used for the flow curve determi-nation. For each testing case, three specimens were tested. Overall,the variation among these three repeats was small. Hence, all theresults reported in the next section are an average of these threerepeats.

2.2. Tensile tests

For the warm tensile tests, a 10-kN electromechanical MTS ma-chine equipped with a furnace (max operating temperature of315 �C) was used as illustrated in Fig. 4. A K-type thermocouplewas placed in contact with the tensile specimen at the middle tomeasure the specimen temperature continuously. The cross-headspeed, v, was calculated based on the target strain rate, SR, value(i.e., v = SR � l0, where l0 is the initial gauge length, which is around50.8 mm). For each testing condition, three specimens were used.

2.3. Material preparation

Two different aluminum alloys, Al5052-H32 and Al6061-T6,were tested in this study. Both had an initial thickness of 2.03 mm.The compositions of these alloys are presented in Table 1. Bulgespecimens were prepared into a hexagonal shape by trimming fourcorners of 150 � 150 mm square blanks, while the tensile speci-mens were prepared according to the ASTM standard E8-04.

3. Material testing results and discussion

3.1. Bulge test results

Before using the measured and recorded pressure and domeheight data for flow curve calculations, first, the accuracy of themeasurement system and strain rate control were evaluated.

MTS Machine

Furnace

Specimen

Fig. 4. Warm tensile test setup.

0

5

10

15

20

25

30

35

40

0 200 400 600Time (sec.)

Pre-cal.Laser sensorCCD cameras

Dom

e H

eigh

t, h d

(mm

)Al5052, t 0 = 2.03mm

SR = 0.0013 s-1

Temp. = 100°C

SR = 0.013 s-1

Temp. = 200°C

Fig. 5. Dome height value comparisons between the reference input (pre-cal), the laser sensor and the CCD cameras.


Measured values of dome height (hd) obtained from both the lasersensor and the CCD cameras are compared as shown in Fig. 5. Thecomparison indicates same measurements by both laser and CCDsensors; and hence, reliable to use in further calculations. In addi-tion, the measured hd and the pre-calculated hd were shown to bealmost identical as illustrated in Fig. 5; thus, a constant strain rate(SR or _e) during each test could be expected. Some of the bulgedsamples are depicted in Fig. 6. The calculation of the flow stresswas carried out based on the measured dome height (hd) and thebulging pressure (P) according to the following set of equationsthat were validated in another study [18]:

Table 1Typical compositions of commercial Al5052 and Al6061 alloys (www.matweb.com).

wt.% Al Cr Cu Fe Mg Mn

AI5052 95.7–97.7 0.15–0.35 Max 0.1 Max 0.4 2.2–2.8 MaAI6061 95.8–98.6 0.04–0.35 0.15–0.40 Max 0.7 0.8–1.2 Ma

Al5052

Al6061

Room temp. 100°C

Room temp. 100°C

Al5052

Al6061

Room temp. 100°CRoom temp. 100°C

Room temp. 100°CRoom temp. 100°C

Fig. 6. Samples of bu

R ¼ ðaþ RcÞ2þ h2d � 2Rchd

2hdð5Þ

td ¼ t0sin aa

� �2

ð6Þ

a ¼ sin�1 aR

� �ð7Þ

r ¼ PR2td

ð8Þ

e ¼ lnt0

td

� �ð9Þ

Si Ti Zn Other, each Other, total

x 0.1 Max 0.25 0–0.05 Max 0.1 0.05 0.15x 0.15 0.4–0.8 Max 0.15 0.25 0.05 0.15

SR=0.0013 s-1

SR=0.0013 s-1

SR=0.013 s-1

SR=0.013 s-1

200°C 240°C

200°C 300°C

SR=0.0013 s-1

SR=0.0013 s-1

SR=0.013 s-1

SR=0.013 s-1

200°C 240°C200°C 240°C

200°C 300°C200°C 300°C

lged specimens.

http://www.matweb.com

0

50

100

150

200

250

300

350

400

450

0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8True strain

True

stre

ss [M

Pa]

Tensile-Room, 0.0013 [1/s]Tensile-100C, 0.0013 [1/s]Tensile-200C, 0.0013 [1/s]Tensile-300C, 0.0013 [1/s]Tensile-Room, 0.013 [1/s]Tensile-100C, 0.013 [1/s]Tensile-200C, 0.013 [1/s]Tensile-300C, 0.013 [1/s]Bulge-Room, 0.0013 [1/s]Bulge-100C, 0.0013 [1/s]Bulge-200C, 0.0013 [1/s]Bulge-300C, 0.0013 [1/s]Bulge-Room, 0.013 [1/s]Bulge-100C, 0.013 [1/s]Bulge-200C, 0.013 [1/s]Bulge-300C, 0.013 [1/s]

Al5052

0

50

100

150

200

250

300

350

400

450

0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8True strain

True

stre

ss [M

Pa]

Tensile-Room, 0.0013 [1/s]Tensile-100C, 0.0013 [1/s]Tensile-200C, 0.0013 [1/s]Tensile-300C, 0.0013 [1/s]Tensile-Room, 0.013 [1/s]Tensile-100C, 0.013 [1/s]Tensile-200C, 0.013 [1/s]Tensile-300C, 0.013 [1/s]Bulge-Room, 0.0013 [1/s]Bulge-100C, 0.0013 [1/s]Bulge-200C, 0.0013 [1/s]Bulge-300C, 0.0013 [1/s]Bulge-Room, 0.013 [1/s]Bulge-100C, 0.013 [1/s]Bulge-200C, 0.013 [1/s]Bulge-300C, 0.013 [1/s]

Al6061

(a)

(b)

Fig. 7. Comparison of flow stress curves for: (a) Al5052 and (b) Al6061 from bothtensile and bulge tests.


where R is the curvature of the bulge radius, a is half bulge diameter(dc/2), Rc is the die corner radius, hd is the instantaneous height atthe dome apex, td is the apex thickness, a is the angle that can bedetermined using Eq. (7), r is the equivalent flow stress, P is theinstantaneous bulging pressure, e is the equivalent strain, and t0

Base location granular shape [Image

Base

Mid-pointApex

Mid-poin25µm

Base

Mid-pointApex

25µm

Fig. 8. Grain shape and distribution at dif

is the initial thickness. The equivalent stress and strain were thencombined to construct the material flow curves for different testingconditions as shown in Fig. 7.

Since the assumption of the perfect spherical bulge shape,which is one of the key assumptions in deriving the Eqs. (5)–(9)for calculation of the apex thickness (td) and the dome height(hd), is far from being true at the beginning of the bulge test [19],only the measured values where hd/a > 0.2 (i.e., e � 0:08) wereused for the calculation of the flow curve in this study. Therefore,the initial flow stress value starts at around 0.08 strain as depictedin Fig. 7. Note that the flow curves shown in Fig. 7 represent theaverage values of the three samples tested at the same testingcondition.

The results in Fig. 7 reconfirm a general trend of the tempera-ture and strain rate effects on the flow stress of aluminum alloys;in that the flow stress decreases with increasing temperature and/or with decreasing strain rate; therefore, improving the formabil-ity. However, there is an inconsistency with this trend; that is inthe case of bulging Al5052 at a low temperature level (i.e., below100 �C), the flow stress was observed to decrease with increasingstrain rate, a phenomenon that is usually caused by the solute dragand dynamic strain aging [8,15]. To elaborate deeper on the resultsin Fig. 7, the strain rate effect was observed to be more pronouncedin the case of Al5052, especially at the elevated temperature levels,than on the Al6061. This low strain rate sensitivity in the 6xxx and7xxx alloys has also been mentioned in the literature by Johanssonet al. [20]; in their study they showed that the effect of the strainrate on the 6xxx and 7xxx alloys would not be observed beforethe strain rate exceeding 1000 s�1. In addition, with a higher per-centage of Mg content in Al5052 than in Al6061, the ductility ofAl5052 at elevated temperatures (i.e., 200 �C) was shown to behigher than that of Al6061, which was caused by the increasingnumber of the slip planes in the hexagonal structure of Mg at ele-vated temperatures. A similar observation on the effect of Mg con-tent on the ductility of aluminum alloys at elevated temperatureswas also reported in a tensile test study of four different aluminumalloys in [7].

Another interesting observation from the bulge test resultscomes from the case of bulging Al6061 at 300 �C at 0.013 s�1 strain

Apex location elliptical shape

s from Al5052-SR0.0013-100C#1 sample]

t location elongated shape

25µm

25µm

25µm

25µm

ferent locations along the center line.


rate. All three specimens tested at this condition were fractured atthe die corner region rather than at the dome apex. Thus, the flowcurve in this case has a maximum strain of less than 0.2. Nonethe-less, when Al6061 specimens were bulged at the lower strain rate(0.0013 s�1) at 300 �C, the maximum strain value is shown to be ashigh as 0.9. Therefore, at a low strain rate, the formability ofAl6061 may continue to improve with increasing temperatureeven above 200 �C.

3.2. Tensile test results

Due to the limitations of the electromechanical MTS system, aconstant strain rate control during the tests was found to berather difficult. Therefore, in this study a constant cross-headspeed was used to provide initial strain rates of 0.0013 and0.013 s�1, which were the same strain rates used in the bulgetests. The calculated true-stress–strain curves from the tensiletests are shown in Fig. 7. The flow curves are only presented here

Al5052 base location

Room Temp.

100°C

25µm25µm

25µm25µm

200°C

240°C N/A

25µm25µm

–

Fig. 9. Temperature effect on grain

up to the respective UTS point for each testing condition since thematerial data after this point is no longer meaningful for the even-tual and further use in analyses. Similar effects of the temperatureand strain rate on the flow stress curves were also observed; thatis the flow stress decreases with increasing temperature anddecreasing strain rate. All three specimens that were pulled underthe same testing conditions provided almost identical flow curves,showing very reliable and repeatable results. In addition, the max-imum elongation of Al5052 under the uniaxial loading conditionwas found to increase considerably at elevated temperature levelsbetween 200 �C and 300 �C, however, such an increase was notobserved in the case of Al6061 alloy after 200 �C. These observa-tions agreed well with the reported results by Novotny and Geiger[5,6] and Li and Ghosh [8]. In their tensile test study, the formabil-ity of Al5xxx continuously increases with the temperature up to300 �C, while that of the Al6xxx would increase up to 200 �Cand the maximum elongation starts to decrease with further in-crease in temperature.

Al6061 base location

25µm25µm

25µm25µm

25µm25µm

25µm25µm

–

structure at the base location.

19.421.4 21.822.5 23.4 23.6

0

5

10

15

20

25

30

35

Room Temp. 100°C 200°C

Gra

in s

ize

[mic

rom

eter

]

Al5052Al6061

Fig. 10. Effect of temperature on the material grain size.

Table 2DOE of closed-die hydroforming.

Runorder Stdorder Temp. Pressure

1 17 1 02 13 �1 �13 10 1 �14 8 �1 05 9 �1 16 4 1 �17 16 1 �18 1 �1 �19 14 �1 0

10 5 1 011 12 1 112 6 1 113 7 �1 �114 3 �1 115 15 �1 116 18 1 117 11 1 018 2 �1 0


As for the comparison of the hydraulic bulge and tensile testresults, Fig. 7, flow stress curves from tensile test are limited tolower strain levels, particularly for Al6061 (�20%) when comparedto flow stress curves from bulge test (�60%). However, with tensiletests, it was possible to obtain reliable flow stress values at lowstrain values (below 0.2), which was not possible in the bulge testsdue to the limitations dictated by the spherical assumptions(i.e., low h/a ratio).

SR=0.013 s-1

Al5052 apex location

SR=0.0013 s-1

200°C

200°C

25µm25µm

25µm25µm

–

Fig. 11. Strain rate effect on grain

Die insertDie insert

Fig. 12. Geometries of

3.3. Microstructure analysis on bulged samples

In order to better understand the effect of temperature andstrain rate on the response of these Al alloys, a microstructureanalysis was performed on the bulged samples. The grain structure

Al6061 apex location

Room Temp.

Room Temp.

25µm25µm

25µm25µm

–

structure at the apex location.

closed-die insert.

0

5

10

15

20

0 20 40 60 80 100Time [sec.]

Pres

sure

[MPa

]

Al5052-300C-15MPaAl5052-300C-20MPaAl6061-200C-10MPaAl6061-200C-15MPaReference Profile

Fig. 13. Hydroforming pressure profiles.


(i.e., size, shape, and distribution) was investigated as the grainstructure was known to largely influence the overall material re-sponse. The bulged specimens were cut along the centerline andsmall sample strips were removed from the center region as shownin Fig. 8. These sample strips were polished and etched in Keller’sreagent (2.5 ml HNO3, 1.5 ml HCl, 1 ml HF, and 95 ml water) to re-veal the undeformed grain structure at the ‘‘base location” and thedeformed grain structure at the ‘‘mid-point” and ‘‘apex” regions asdepicted in Fig. 8. Most of the grains were found to have granularstructure at the base location (i.e., undeformed grains), while elon-gated and elliptical grain structures were observed at the mid-point and the apex locations, respectively. The grains at theseregions were elongated or stretched as they underwent a largeplastic deformation amount during the bugle tests.

200°C

300°C

10MPa

200°C

300°C

10MPa

(a)

200°C

300°C

10MPa

200°C

300°C

10MPa

(b)

Fig. 14. Closed-die hydroformed sam

The effect of temperature on the grain structure was investi-gated at the base location of the cut strips. The base location wasselected because the grains in this region did not undergo anydeformation (strain and strain rate independent); thus, represent-ing the microstructure changes caused merely by the temperatureeffect. The microscopic images of the grain structures are illus-trated in Fig. 9 for both Al5052 and Al6061. Note that the grainstructures from the specimens that were bulged at 240 �C andhigher could not be clearly seen, which may have been caused bythe significant microstructural changes, most likely recrystalliza-tion and growth of grains or precipitates, as the bulging tempera-ture enters into the ‘‘warm forming” regime (i.e., temperature>0.3Tm, where Tm indicates melting point). For the samples thatthe grain boundary could be clearly indicated, the grain size wasmeasured based on the ASTM Standard E112-88 (i.e., Mean LinealIntercept or Heyn’s method). The grain size measurement resultsare presented in Fig. 10 where slightly larger grains were observedat elevated temperature levels although the difference is statisti-cally insignificant. According to the Hall–Petch relation [21,22],material with larger grain size was predicted to have lowerstrength. Despite the fact that such a case was observed in thisstudy, it is believed that the drop in the flow stress curve is notdue to the slightly and statistically indifferent grains, but mainlydue to the additional slip lines activated due to the elevatedtemperature.

The effect of the strain rate was also investigated through thegrain structure analysis. Unlike in the temperature effect study,the location of interest on the cut specimen was shifted to be atthe apex of the dome rather than at the base location as mostdeformation dependent characteristic could be seen most in thisregion. The microscopic images of the specimens bulged at differ-ent strain rates are shown in Fig. 11. Unfortunately, no significantdifference was observed in terms of the grain structure between

15MPa 20MPa

Al5052

15MPa 20MPa

Al5052

15MPa 20MPa

Al6061

15MPa 20MPa

Al6061

ples: (a) A5051 and (b) Al6061.

0

5

10

15

20

25

0 20 40 60 80 100 120

width [mm]

heig

ht [m

m]

200C-10MPa200C-15MPa200C-20MPa300C-10MPa300C-15MPa300C-20MPaDie Profile

Al5052

0

5

10

15

20

25

0 20 40 60 80 100 120

width [mm]

heig

ht [m

m]

200C-10MPa200C-15MPa200C-20MPa300C-10MPa300C-20MPaDie Profile

Al6061

Fig. 15. Profiles of hydroformed parted at different temperature and pressurelevels.

Fig. 16. Response surface plots.


the two strain rate values used in this study (0.0013 and 0.013 s�1).Nonetheless, the effect of the strain rates on the material response(i.e., flow curve) and the formability (i.e., maximum elongation)was obvious, especially at elevated temperature levels as discussedin the previous section. Hence, it can be concluded that it is not thechange in grain size, but the thermally activated dislocation mo-tion that causes the formability increase in warm forming.

4. Closed-die hydroforming experiments

A set of closed-die warm hydroforming experiments were con-ducted on the same alloys using a die insert as shown in Fig. 12.These experiments were conducted under a design of experiment(DOE) plan, as tabulated in Table 2 (i.e., 18 runs for each alloy),to obtain a quantified understanding of the effect of temperatureand pressure on the die cavity filling and thinning (forming limits).During the experiments, a constant blank holder force of 1000 kNwas used, as in the bulge tests, to clamp the specimens at theperiphery. A linearly increasing (ramp-up) pressure profile with a

Table 3Percentage of die filling under different process conditions.

10 MPa (%)

Al5052 200 �C 77.2300 �C 95.2

Al6061 200 �C 60.7300 �C 94.3

slope of 0.22 MPa/s was used as a reference input (Fig. 13). The ac-tual hydroforming pressure profiles, recorded during the testsusing a pressure transducer, were shown to closely follow the ref-erence pressure input profile (Fig. 13).

After each test, the hydroformed parts were measured using thestereoscopic CCD cameras with ARAMIS software to obtain full sur-face profiles as illustrated in Figs. 14 and 15. For the assurance ofprocess repeatability, three experiments were conducted for eachcase. An average value is reported unless otherwise is stated in thissection.

The effect of temperature and pressure levels on the sheet form-ability can clearly be seen in Figs. 14 and 15. Specifically, at 200 �C,both Al5052 and Al6061 sheet blanks showed poor formability,and a fracture occurred in the area of the die radius at the centerwhen the pressure was increased from 10 to 15 MPa for Al5052and from 15 to 20 MPa for Al6061. As the temperature was in-creased to 300 �C, the formability of Al5052 sheet blanks appearedto improve, and no premature rupture was observed up to 20 MPa.However, for Al6061, an increasing temperature only reduced thematerial strength (i.e., higher profile at the same forming pressurewhen the temperature was increased), while the elongation prop-erties did not change. With the increasing temperature, all Al6061specimens showed fractures at the central die radius area once acertain part height was reached. This observation agrees well withthe flow curve plots of Al6061 (Fig. 7), in which the maximumstrain (i.e., elongation) value did not change as temperature in-creased, but the flow stress values were observed to significantlydecrease. Furthermore, based on the comparison of the part pro-files in Fig. 15, it is clearly shown that Al5052 has a better formabil-ity when compared to Al6061 under these forming conditions (i.e.,uniform temperature distribution at 200 and 300 �C and linearpressure profiles up to 15–20 MPa). Finally, it is important to point

15 MPa (%) 20 MPa (%)

94.2 93.795.6 96.2

80.8 83.2n/a 94.5

0

10

20

30

40

50

60

0 10 20 30 40 50 60Radial distance [mm]

% T

hinn

ing

200C-10MPa300C-15MPa300C-20MPaDie Profile

Al5052

0

10

20

30

40

50

0 10 20 30 40 50 60

Radial distance [mm]

% T

hinn

ing

200C-10MPa200C-15MPaDie Profile

Al6061

Fig. 17. Thickness profiles in radial direction.

Table 4Material parameters for Al5052 and Al6061 used in FEA.

Material parameters AI5052 AI6061

Modulus of elasticity, E (MPa) 70,300 68,900Poisson’s ratio, v 0.33 0.33Mass density, q (Mg/mm3) 2.68E � 09 2.70E � 09Yield strength, r0 (MPa) 89.6 55.2

Table 5Material constants ð�r ¼ K�en _�emÞ from tensile and bulge tests at different temperatures.

Test Material Temp. (�C) K (MPa) n m

Bulge test AI5052 23 455 0.14 0.010200 412 0.33 0.075300 401 0.55 0.135

AI6061 23 483 0.11 0.013100 497 0.17 0.020200 503 0.12 0.075

Tensile test AI5052 23 777 0.45 0.000100 966 0.50 0.027200 437 0.28 0.051300 253 0.09 0.151

AI6061 23 979 0.38 �0.006100 1058 0.41 0.009200 880 0.36 0.046300 474 0.21 0.114


out that the die corner rupture was observed in the case of Al6061blanks formed at 300 �C and 15 MPa. As a result, their profiles areexcluded in Fig. 15 as well as in the rest of the analysis.

In order to quantify the effects of the temperature and pressureon the formability of the hydroformed parts, two variables (the die

Total Equivalent P

Fig. 18. Two-dimensional axisymmetric numerical modeling of the closed-d

filling ratio and thinning) were measured and reported in Table 3and in Figs. 16 and 17. Note that for thinning measurements, thehydroformed specimens were cut into two halves with an offsetof 10 mm from the center line, and the thickness of the bigger halfwas measured using a micrometer attached with conical shape tipsat several locations along the radial directions. In addition, for ameaningful comparison of the thinning, only the specimens with-out the fracture were used.

When cavity filling comparisons in Table 3 and Fig. 16 areconsidered, at 200 �C, an increasing pressure leads to an increasein the cavity filling ratio for A5052 (77–93%) and Al6061 (60–83%).

lastic Strain

ie warm hydroforming: FE model, predicted and actual deformed part.

0

5

10

15

20

25

0 10 20 30 40 50 60Width [mm]

Hei

ght [

mm

] .

Al5052

0

5

10

15

20

25

0 10 20 30 40 50 60Width [mm]

Hei

ght [

mm

] .

200C-10MPa-Exp200C-10MPa-FEA-Bulge200C-10MPa-FEA-Tensile200C-15MPa-Exp200C-15MPa-FEA-Bulge200C-15MPa-FEA-TensileDie Profile

Al6061


Fig. 19. Comparison of hydroformed part profiles for A5052 and Al6061 alloys atdifferent process conditions.


However, for 300 �C, the same cannot be said. A similar observationis made for an increasing temperature at low pressure value(10–15 MPa). However, in general, when a regression analysis wasmade for the entire ranges of temperature and pressure, their effecton the cavity filling ratio was found to be statistically insignificant

Table 6Simulation cases.

Exp. case FEA run Material T

Case 1 1 AI5052 22 AI5052 2

Case 2 3 AI5052 34 AI5052 3

Case 3 5 AI6061 26 AI6061 2

Case 4 7 AI6061 28 AI6061 2

Table 7Percentage of die filling comparisons.

Material Temp. (�C) Pressure (MPa) Pe

Ex

Al5052 200 10 77300 15 95

Al6061 200 10 60200 15 80

(i.e., for Al5052: p = 0.238 for temperature and p = 0.246 for pres-sure; for Al6061: p = 0.151 for temperature and p = 0.350 forpressure).

In terms of thinning comparisons as shown in Fig. 17, first of all,the highest thinning, and hence fracture in some cases, occurredaround the die radius at the center region of the part. Second,the effect of temperature and pressure were as expected (i.e.,increasing pressure and temperature leads to increasing thinningin general).

Based on this result discussion, the use of a uniform tempera-ture distribution (i.e., isothermal conditions) and a linearly increas-ing pressure profile (i.e., ramp-up pressure input) may not be themost efficient approach to increase the cavity filling and reducethinning (i.e., part formability). Thus, process optimization investi-gation is needed to determine optimal process conditions (e.g., var-iable loading profiles: pressure and blank holder force, andtemperature). This optimization study would require the use of fi-nite element analysis (FEA) tool. In the following section, FEA mod-els of the hydroforming process will be developed and the materialproperties obtained from the bulge and tensile tests will be vali-dated for their accuracy and applicability in predicting theclosed-die profiles and thinning values as measured and reportedin this section.

5. Numerical modeling, validation and comparison of materialbehaviors

In this section, Finite element models of the closed-die warm hydro-forming process (Fig. 18) were developed using MSC.Marc2007r1 soft-ware. Since the problem at hand was an axisymmetric type, only a 2-Dhalf-model analysis was developed. In the model, the sheet blank wasmodeled using deformable, solid, quad-4 elements. Four elementsacross the blank thickness were used. Both ends of the blank were fixed(no displacement) to reflect on the actual boundary condition of theprocess where the blank was tightly clamped between the upper andlower dies to prevent any radial flow of the material into the die cavity.Hydroforming pressure was applied from the bottom side of the blankwith an increasing rate of 0.22 MPa/s until the pressure reached the pre-set values. According to the actual pressure profiles recorded during theexperiments, the pressure pump and regulator provided close control ofthe pressure profile as shown in Fig. 13. Thus, a ramp pressure input

emp. (�C) Pressure (MPa) Mat. flow curve

00 10 Bulge00 10 Tensile




rcentage of die filling

periment FEA-bulge (%error) FEA-tensile (%error)

.2 86.1 (11.5) 79.2 (2.6)

.6 98.2 (2.7) 98.2 (2.7)

.7 67.6 (11.4) 59.8 (1.5)

.8 83.8 (3.7) 75.1 (7.1)

Fig. 20. Thickness measurement.


with a slope of 0.22 MPa/s was utilized in all validation cases. Coulombfriction model was selected with a friction coefficient of 0.1 for all con-tact surfaces in this study.

Material input parameters: modulus of elasticity (E), poisson’sratio (v), mass density (q), and yield strength (r0) for bothAl5052 and Al6061 are given in Table 4, while the material flowcurves obtained from the tensile and bulge tests at different tem-perature levels were modeled using Field–Backofen equation (i.e.,rate power law: �r ¼ K�en _�em) as tabulated in Table 5.

In order to validate the FE models, four experimental cases(Table 6) were selected which included the closed-die hydroform-ing results for both Al5052 and Al6061 at elevated temperatures(200–300 �C) and two different pressure levels (10 MPa and15 MPa). The hydroformed parts in these cases were fracture-free,

0

10

20

30

40

50

60

0 10 20Radial di

% T

hinn

ing

Al5052Δ%Thinning FEA-Bulge FEA-Tensile FEA-Bulge FE

Ave. 2.8 0.8 5.5Max. 6.2 3.8 18.2

200C-10MPa 300C-15M

(a)

(b)

0

10

20

30

40

50

60

0 10 20Radial di

% T

hinn

ing

Al6061Δ%Thinning FEA-Bulge FEA-Tensile FEA-Bulge F

Ave. 0.1 2.5 1.4Max. 6.1 4.6 4.1

200C-15200C-10MPa

Fig. 21. Thinning comparisons in radial dir

which makes the thickness measurement physically possible. Theother output chosen for comparison purpose was the part profiles(or percentage of die filling). In addition, with the two flow curvesobtain from tensile and bulge tests at each temperature level, atotal of eight simulation runs were carried out and their resultsare shown and compared in Fig. 19 and Table 7.

According to the part profile comparisons between the experi-mental measurements and the FEA predictions in Fig. 19, it wasfound the tensile flow curves provided better profile predictionsat low pressure (i.e., low strain) levels (10 MPa, where �e < 0:2),while the bulge flow curves did so at the high pressure (i.e., highstrain) levels (>15 MPa, where �e > 0:4). The observation could bewell explained by the limitations and assumptions associated witheach test method and the derivations of the material flow curvesbased on the raw test data. Specifically, for tensile tests conductedat 200 �C, the maximum strain levels were found to be around 0.2for Al5052, and 0.15 for Al6061. Therefore, FEA predictions for highstrain levels would require extrapolation of the material flowcurves. On the other hand, a higher strain levels could be achievedin the bulge tests, e.g., at 200 �C, the achievable strain was around0.5 for Al5052 and about 0.35 for Al6061; and thus, the materialdata from the bulge test provided better FEA predictions at highstrain or pressure levels. Furthermore, with the assumption of anon-spherical dome shape of the bulge specimen below the h/a va-lue of 0.2 (corresponding to a strain value of 0.08), the materialdata below this threshold value was excluded for the flow curvedetermination in bulge testing case, which in turns made the FEApredictions based on the bulge flow curves less accurate at low

30 40 50 60stance [mm]


A-Tensile5.3

18.0

Pa

30 40 50 60stance [mm]


EA-Tensile6.8

12.3

MPa

ection for: (a) Al5052 and (b) Al6061.


strain or pressure levels. Nonetheless, both bulge and tensile mate-rial flow curves were shown to provide reasonable predictions ofthe hydroformed part profiles as can be confirmed by the compar-isons of the percentage die filling in Table 7 with the maximum er-rors of 1.5–11% based on both types of flow curves.

The second comparison is based on the thinning levels and dis-tributions. The hydroformed specimens were cut in half as shownin Fig. 20 along a line that was about 1 cm off from the centerline ofthe part. The measurement was then performed using a digitalmicrometer at several critical locations with numerous repetitions.Thinning distribution comparisons between the experimentalmeasurements and the FEA predictions are shown in Fig. 21. Thecomparisons were found to be in a good agreement, except forthe case of Al5052 at 300 �C and 15 MPa where the maximum thin-ning percentage was higher than 40% and the instability, whichwas not considered in the FEA, had occurred. In addition, forAl6061 specimens it was clear that the thickness values predictedbased on the bulge flow curves were more accurate than those ob-tained from the tensile flow curves. This is due mainly to the load-ing condition of the bulge test (i.e., bi-axial stress) that is morerelevant to the actual state of stress and strain in the hydroformingprocess. Therefore, both tests were shown to be appropriate forobtaining the material properties for the eventual use in FEA ofwarm hydroforming. However, for the parts with complex geome-tries or when large deformation is expected, numerical modelsbased on bulge tests provided closer predictions.

In summary, the FEA models of the hydroforming process devel-oped in this section along with the material flow curves obtainedfrom both bulge and tensile tests in the previous section have beenshown to provide reasonable and reliable numerical predictions interms of both part profiles and thinning distributions.

6. Conclusions

In this study, mechanical characteristics of Al5052 and Al6061sheet blanks were characterized using both tensile and bulge test-ing methods at temperature levels up to 300 �C, and at the strainrates of 0.0013 and 0.013 s�1. In addition to the expected and gen-eral trend of decreasing material flow curves with increasing tem-perature and/or decreasing strain rate, it was found that the flowstress curves from tensile test were limited to lower strain levels,particularly for Al6061 (<15%) when compared to the flow stresscurves from the bulge tests (�30–60%). However, with tensiletests, it was possible to obtain reliable flow stress values at strainvalues lower than 2%, which was not possible in the bulge tests dueto the limitations dictated by the spherical bulging assumptions(i.e., low h/a ratio). The microstructure analysis showed that thechange of grain size was not significant at different temperatureand strain rates, which leads us to conclude that the decrease inthe flow stress at high temperature levels was mainly due to thethermally activated dislocation lines.

The effects of the temperature and the pressure on the sheetformability were further investigated in a set of closed-die hydro-forming experiments. At 200 �C, an increasing pressure leads to anincrease in the cavity filling ratio for A5052 (77–93%) and Al6061(60–83%). However, for 300 �C, the same cannot be stated. A simi-lar observation is made for an increasing temperature at low pres-sure value (10 MPa), but not at high pressure level (20 MPa). Thetest results along with the regression analysis showed that theuse of a uniform temperature distribution and a ramping pressureinput do not have a significant effect on the percentage of die fillingvalues. Therefore, process optimization investigation is needed to

determine the optimal process conditions (e.g., variable loadingprofiles: pressure and blank holder force, and temperature) formaximize the part formability.

FE modeling findings and comparison with closed-die hydro-forming experiments based on the material flow stress curves fromboth bulge and tensile tests at different temperature, pressure andstrain rate conditions indicated that, in general, flow curves fromboth bulge and tensile tests are in good agreement with experi-mental measurements in terms of predicting the part profile, cavityfilling and thinning comparisons. Overall prediction errors wereless than 10–12%. However, when examined closely, it was re-vealed that flow curves from bulge tests resulted in better predic-tion accuracy, particularly at high strain (pressure) levels.

Acknowledgements

Authors are thankful to National Science Foundation (NSF) forthe partial support on this project through NSF ENG/CMMI Grants0703912 and NSF IIP IUCRC Grant 0638588.

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