Materials and Design 31 (2010) 2422–2434                                                               Contents lists avai...
S. Mahabunphachai, M. Koç / Materials and Design 31 (2010) 2422–2434                                         2423formabili...
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S. Mahabunphachai, M. Koç / Materials and Design 31 (2010) 2422–2434                                                      ...
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S. Mahabunphachai, M. Koç / Materials and Design 31 (2010) 2422–2434                                    2427rate. All thre...
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S. Mahabunphachai, M. Koç / Materials and Design 31 (2010) 2422–2434                               2429                  2...
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S. Mahabunphachai, M. Koç / Materials and Design 31 (2010) 2422–2434                                                      ...
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S. Mahabunphachai, M. Koç / Materials and Design 31 (2010) 2422–2434                                        2433          ...
!!Investigations on forming of aluminum 5052 and 6061 sheet alloys at
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!!Investigations on forming of aluminum 5052 and 6061 sheet alloys at

  1. 1. Materials and Design 31 (2010) 2422–2434 Contents lists available at ScienceDirect Materials and Design journal homepage: on forming of aluminum 5052 and 6061 sheet alloys atwarm temperaturesS. 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, Thailanda r t i c l e i n f o a b s t r a c tArticle history: In an ongoing quest to realize low-mass transportation vehicles with enhanced fuel efficiency, deforma-Received 1 September 2009 tion characteristics of Al5052 and Al6061 were investigated. In the first part of this study, material behav-Accepted 23 November 2009 ior of Al5052 and Al6061 sheet alloys were investigated under different process (temperature and strainAvailable online 26 November 2009 rate) 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 levelsKeywords: in tensile tests. The microstructure analysis showed that the change of grain size due to the effects of ele-Aluminum sheet vated temperatures and strain rates were not significant; therefore, it was concluded that the decrease inFormabilityLightweight material the flow stress at high temperature levels was mainly due to the thermally activated dislocation lines. InHydroforming the second part, the effect of the temperature and the pressure on the formability was further investi-Warm forming gated in a set of closed-die warm hydroforming experiments. The test results showed that a linearlyAA5052 increasing pressure profile up to $20 MPa levels did not have a significant effect on the die filling ratiosAA6061 and thinning of the parts when a uniform temperature distribution of 300 °C was applied. Finally, in the third part of the study, finite element models were developed for the same closed-die hydroforming geometry using the material behavior models obtained from bulge and tensile tests. Flow stress curves obtained from tests were compared in terms of predicting the cavity filling ratios and thinning profiles from the experiments. Based on the comparison, it was revealed that flow stress curves obtained from e the warm hydraulic bulge tests provided accurate predictions at high strain levels (i.e., 0:4, when part filling 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 indicated that 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 whenever available in order to conduct accurate numerical analyses for warm sheet hydroforming where complex geometry and loading conditions exist. Ó 2009 Elsevier Ltd. All rights reserved.1. Introduction solution, the car manufacturers along with various research groups have been investigating the fabrication of structural and body parts With an increasing awareness and effects of global warming, out of lightweight materials such as aluminum and magnesium al-and the scanty fossil fuel resources left when compared to the ever loys [1–3]. On the other hand, despite the obvious advantages ofincreasing demand of oil, car manufacturers have been seeking for the lightweight alloys, they have a notable drawback in that theiralternative and sustaining solutions to the fuel efficiency problem. formability is significantly lower than traditional steel alloys atMany believe that the next generation cars must run on alternative room temperature conditions, which is usually caused by the highand clean fuels (e.g., hydrogen via fuel cells) to prevent further in- alloy percentages that are required for high strength [4,5]. Forcrease of harmful emissions. However, this approach appears to be example, the formability of aluminum alloys is only about two-more of a solution that may not be practically and economically third of a deep drawing steel grade, their Young’s modulus is aboutrealized in short term (i.e., $10 years). On the other hand, another one-third of the steel, which in turn causes higher susceptibility ofprominent approach that is sustaining, effective, and sooner would wrinkling and springback [6], and their elongation is about half ofbe the realization of low-mass vehicles. In the pursuit of latter steel’s [7]. The inferior formability of aluminum alloys makes it more difficult and expensive to use them in mass production of structural and body parts, which requires high levels of elongation * Corresponding author. Tel.: +1 804 827 7029. E-mail address: (M. Koç). and ductility to be formed into complex shapes. Nevertheless, the0261-3069/$ - see front matter Ó 2009 Elsevier Ltd. All rights reserved.doi:10.1016/j.matdes.2009.11.053
  2. 2. S. Mahabunphachai, M. Koç / Materials and Design 31 (2010) 2422–2434 2423formability 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, where CCD Camerasadditional 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, Temp.were also shown to further enhance the formability of the alumi- controller ARAMISnum 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 hydroforming Hydraulicprocess has been used for an increasing number of structural and Pump Die Setbody applications as it enlarges the forming limit windows ofmaterials due to the biaxial and frictionless loading conditions, Die Insertby 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 and CCDhydroforming [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 Laser(2) determination of the optimal process parameters (i.e., temper- Silicone based O-ring Sensorature level and distribution, pressure and blank holding profiles). In this study, our objectives were to: (1) determine proper Fig. 1. Warm hydraulic bulge test setup.experimental methodologies to accurately characterize the mate-rial behavior under warm hydroforming conditions (i.e., comparewarm tensile and warm hydraulic bulge tests), (2) experimentally AZ-2-180HPU-LW), pressure controller (Marsh Bellofram Typeunderstand and quantify the effects of process parameters, such as 3510), and pressure transducer (OMEGA PX605), (2) a set of bulg-temperature and internal pressure, on the part formability into ing die: upper and lower die with a bulge diameter of 100 mm andrepresentative die cavities with reasonably complex geometries, each with a die corner radius of 6.5 mm, clamping and sealingand (3) develop finite element models (FEM) to determine the mechanism (silicone based O-ring and copper O-ring), (3) heatingapplicability of bulge and tensile test findings to accurately predict system: cartridge heaters, temperature controller (OMEGAthe part formability. For this purpose, two commonly used alumi- CN616tc1), and thermocouples (Type K), and (4) in-die non-con-num alloys (5052 and 6061) were selected for experimentation. tact measurement systems: laser sensor (Keyence LK-G402) and In the next section, experimental setup and conditions for the stereoscopic system (two CCD cameras with GOM ARAMIS systemwarm tensile and warm hydraulic bulge tests are presented. In by Trilion). The non-contact measurement systems were used tothe third section, material test results are presented and compared avoid any temperature gradient at the contact location, whichin terms of achievable strain and stress levels (i.e., flow stress can influence the material behavior [4]. In order to avoid damagecurves). In the fourth section, experimental conditions and results on the laser sensor and CCD cameras due to the splashing of theof a design of experiment (DOE) study conducted using a set of hot pressurized oil (Marlotherm SH), a thick glass was placed onclosed-dies are presented and discussed to quantify the effect of the housing roof.process parameters on the cavity filling (i.e., formability) in warm The process parameters of interest in this test were the effect ofhydroforming. Part profile, die filling ratio, and thinning on the temperature and strain rate on the flow stress behavior of thewarm hydroformed parts were reported and compared. A regres- materials. The temperature of each die half was monitored andsion analysis was conducted to reveal the significance of pressure controlled independently using two separate sets of cartridge heat-and temperature parameters. In the fifth section, a finite element ers and thermocouples (t/c) attached to each die half as shown inmodel (FEM) of the warm hydroforming process is developed and Fig. 2. With this type of control loop, the temperature variationvalidated by comparing the predictions with the experimental find- during the test was below 5 °C. The heating cycle was made asings. The flow stress curves obtained from both the tensile and short as possible, where the cycle time depends on the set temper-bulge tests are used in the FEA validation in order to determine ature value. On the other hand, a holding time of 5–10 min waswhich set of material test data is more accurate in predicting the used to allow the uniform temperature distribution on the blankpart formability (i.e., cavity filling and thinning). Finally, a summary and the oil in the die cavity. The strain rate (SR or e) was also con- _of the results and conclusions are presented in the sixth section. trolled using a feedback loop with a PID controller. Based on the difference between the pre-calculated dome height profile (refer- ence value) and the instantaneous dome height value from the la-2. Material characterization experiments ser measurement, the control signals were sent to the pressure controller to regulate the air input pressure and flow rate at the2.1. Hydraulic bulge test setup pump inlet. The pressure and flow rate of the discharge fluid (oil) at the pump outlet were directly proportional to this controlled For warm hydraulic bulge tests, a specially designed and built air flow at the inlet. The schematic of the control loops of the pneu-system composed of four major sub-systems was used as depicted matic/hydraulic system and the non-contact measurement systemin Fig. 1: (1) a pneumatic/hydraulic system: pump (Hydratron (laser sensor) is presented in Fig. 2.
  3. 3. 2424 S. Mahabunphachai, M. Koç / Materials and Design 31 (2010) 2422–2434 Laser sensor strain rate levels are plotted in Fig. 3. These profiles were used as a reference input signal in the feedback control loop. With this test setup, the bulging pressure and dome height CCD Cameras heaters could be continuously measured and recorded using the pressure t/c transducer and the non-contact measurement systems (laser sen- sor and CCD cameras) during the test. Curvature of the bulging Upper die Temp. was also measured using the ARAMIS system with the CCD cam- Controller Hot oil eras. However, the dome height was found to be the same with Lower die the laser measurements as explained in detail in another study t/c P [18]. These pressure and dome height data were later synchronized together by the time stamp, and used for the flow curve determi- nation. For each testing case, three specimens were tested. Overall, LabVIEW P Transducer the variation among these three repeats was small. Hence, all the results reported in the next section are an average of these three repeats. P controller Pump Fig. 2. Schematic diagram of the warm bulge test setup. 2.2. Tensile tests For the warm tensile tests, a 10-kN electromechanical MTS ma- chine equipped with a furnace (max operating temperature of A pre-calculated dome height (hd) profile was used to obtain a 315 °C) was used as illustrated in Fig. 4. A K-type thermocoupleconstant strain rate during the tests. It was derived based on the was placed in contact with the tensile specimen at the middle togeometrical relationships in a circular bulge testing of thin sheet measure the specimen temperature continuously. The cross-headblanks as follows [17]: speed, v, was calculated based on the target strain rate, SR, value t0 (i.e., v = SR Ã l0, where l0 is the initial gauge length, which is arounde ¼ ln ð1Þ 50.8 mm). For each testing condition, three specimens were used. td !2 2 dctd ¼ t0 2 2 ð2Þ 2.3. Material preparation dc þ 4hd Two different aluminum alloys, Al5052-H32 and Al6061-T6,where e is the equivalent strain, t0 is the initial sheet thickness, td is were tested in this study. Both had an initial thickness of 2.03 mm.the instantaneous apex thickness, dc is the bulge diameter, and hd is The compositions of these alloys are presented in Table 1. Bulgethe instantaneous dome height. In addition, since strain rate ðeÞ is _ specimens were prepared into a hexagonal shape by trimming fourthe rate of change in strain, one can write: corners of 150 Â 150 mm square blanks, while the tensile speci-e¼eÁt _ ð3Þ mens were prepared according to the ASTM standard E8-04.where t is time. Combining Eq. (1)–(3), a relationship between the 3. Material testing results and discussioninstantaneous dome height (hd) and the strain rate ðeÞ can be ob- _tained as: 3.1. Bulge test results dc pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffihd ¼ eeÁt=2 À 1 _ ð4Þ Before using the measured and recorded pressure and dome 2 height data for flow curve calculations, first, the accuracy of theThe relationship in Eq. (4) was used for plotting the reference hd measurement system and strain rate control were evaluated.profile as a function of time. Typical dome height profiles at various 60 MTS Machine 50 0.13 s-1 Dome height, hd (mm) 40 0.013 s-1 Specimen 30 20 0.0013 s-1 10 0 0 100 200 300 400 Furnace time (s) Fig. 3. Typical dome height profiles for different strain rate levels. Fig. 4. Warm tensile test setup.
  4. 4. S. Mahabunphachai, M. Koç / Materials and Design 31 (2010) 2422–2434 2425 40 35 SR = 0.013 s-1 Temp. = 200°C 30 Dome Height, hd (mm) 25 20 SR = 0.0013 s-1 Temp. = 100°C 15 10 Al5052, t 0 = 2.03mm Pre-cal. 5 Laser sensor CCD cameras 0 0 200 400 600 Time (sec.) Fig. 5. Dome height value comparisons between the reference input (pre-cal), the laser sensor and the CCD cameras. 2Measured values of dome height (hd) obtained from both the laser ða þ Rc Þ2 þ hd À 2Rc hdsensor and the CCD cameras are compared as shown in Fig. 5. The R¼ ð5Þ 2hdcomparison indicates same measurements by both laser and CCD 2 sin asensors; and hence, reliable to use in further calculations. In addi- td ¼ t0 ð6Þtion, the measured hd and the pre-calculated hd were shown to be a aalmost identical as illustrated in Fig. 5; thus, a constant strain rate a ¼ sinÀ1 ð7Þ(SR or e) during each test could be expected. Some of the bulged _ Rsamples are depicted in Fig. 6. The calculation of the flow stress PR r¼ ð8Þwas carried out based on the measured dome height (hd) and the 2t d bulging pressure (P) according to the following set of equations t e ¼ ln 0 ð9Þthat were validated in another study [18]: tdTable 1Typical compositions of commercial Al5052 and Al6061 alloys ( wt.% Al Cr Cu Fe Mg Mn Si Ti Zn Other, each Other, total AI5052 95.7–97.7 0.15–0.35 Max 0.1 Max 0.4 2.2–2.8 Max 0.1 Max 0.25 0–0.05 Max 0.1 0.05 0.15 AI6061 95.8–98.6 0.04–0.35 0.15–0.40 Max 0.7 0.8–1.2 Max 0.15 0.4–0.8 Max 0.15 0.25 0.05 0.15 Room temp. 100°C 200°C 300°C SR=0.0013 s-1 Al5052 SR=0.013 s-1 SR=0.0013 s-1 Al6061 SR=0.013 s-1 Room temp. 100°C 200°C 240°C Fig. 6. Samples of bulged specimens.
  5. 5. 2426 S. Mahabunphachai, M. Koç / Materials and Design 31 (2010) 2422–2434 is the initial thickness. The equivalent stress and strain were then (a) 450 Tensile-Room, 0.0013 [1/s] Al5052 Tensile-100C, 0.0013 [1/s] combined to construct the material flow curves for different testing 400 Tensile-200C, 0.0013 [1/s] Tensile-300C, 0.0013 [1/s] conditions as shown in Fig. 7. Tensile-Room, 0.013 [1/s] 350 Tensile-100C, 0.013 [1/s] Since the assumption of the perfect spherical bulge shape, True stress [MPa] Tensile-200C, 0.013 [1/s] 300 Tensile-300C, 0.013 [1/s] Bulge-Room, 0.0013 [1/s] which is one of the key assumptions in deriving the Eqs. (5)–(9) Bulge-100C, 0.0013 [1/s] Bulge-200C, 0.0013 [1/s] for calculation of the apex thickness (td) and the dome height 250 Bulge-300C, 0.0013 [1/s] Bulge-Room, 0.013 [1/s] (hd), is far from being true at the beginning of the bulge test [19], 200 Bulge-100C, 0.013 [1/s] only the measured values where hd/a 0.2 (i.e., e ! 0:08) were Bulge-200C, 0.013 [1/s] Bulge-300C, 0.013 [1/s] 150 used for the calculation of the flow curve in this study. Therefore, the initial flow stress value starts at around 0.08 strain as depicted 100 in Fig. 7. Note that the flow curves shown in Fig. 7 represent the 50 average values of the three samples tested at the same testing 0 condition. 0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 The results in Fig. 7 reconfirm a general trend of the tempera- True strain ture and strain rate effects on the flow stress of aluminum alloys; in that the flow stress decreases with increasing temperature and/ (b) 450 Al6061 Tensile-Room, 0.0013 [1/s] Tensile-100C, 0.0013 [1/s] or with decreasing strain rate; therefore, improving the formabil- 400 Tensile-200C, 0.0013 [1/s] Tensile-300C, 0.0013 [1/s] ity. However, there is an inconsistency with this trend; that is in 350 Tensile-Room, 0.013 [1/s] the case of bulging Al5052 at a low temperature level (i.e., below Tensile-100C, 0.013 [1/s] True stress [MPa] 300 Tensile-200C, 0.013 [1/s] Tensile-300C, 0.013 [1/s] 100 °C), the flow stress was observed to decrease with increasing Bulge-Room, 0.0013 [1/s] strain rate, a phenomenon that is usually caused by the solute drag 250 Bulge-100C, 0.0013 [1/s] Bulge-200C, 0.0013 [1/s] and dynamic strain aging [8,15]. To elaborate deeper on the results Bulge-300C, 0.0013 [1/s] 200 Bulge-Room, 0.013 [1/s] in Fig. 7, the strain rate effect was observed to be more pronounced Bulge-100C, 0.013 [1/s] 150 Bulge-200C, 0.013 [1/s] in the case of Al5052, especially at the elevated temperature levels, Bulge-300C, 0.013 [1/s] than on the Al6061. This low strain rate sensitivity in the 6xxx and 100 7xxx alloys has also been mentioned in the literature by Johansson 50 et al. [20]; in their study they showed that the effect of the strain 0 rate on the 6xxx and 7xxx alloys would not be observed before 0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 the strain rate exceeding 1000 sÀ1. In addition, with a higher per- True strain centage of Mg content in Al5052 than in Al6061, the ductility ofFig. 7. Comparison of flow stress curves for: (a) Al5052 and (b) Al6061 from both Al5052 at elevated temperatures (i.e., 200 °C) was shown to betensile and bulge tests. higher than that of Al6061, which was caused by the increasing number of the slip planes in the hexagonal structure of Mg at ele- vated temperatures. A similar observation on the effect of Mg con-where R is the curvature of the bulge radius, a is half bulge diameter tent on the ductility of aluminum alloys at elevated temperatures(dc/2), Rc is the die corner radius, hd is the instantaneous height at was also reported in a tensile test study of four different aluminumthe dome apex, td is the apex thickness, a is the angle that can be alloys in [7].determined using Eq. (7), r is the equivalent flow stress, P is the Another interesting observation from the bulge test resultsinstantaneous bulging pressure, e is the equivalent strain, and t0 comes from the case of bulging Al6061 at 300 °C at 0.013 sÀ1 strain Mid-point Apex 25µm Apex location elliptical shape Base 25µm Mid-point location elongated shape 25µm Base location granular shape [Images from Al5052-SR0.0013-100C#1 sample] Fig. 8. Grain shape and distribution at different locations along the center line.
  6. 6. S. Mahabunphachai, M. Koç / Materials and Design 31 (2010) 2422–2434 2427rate. All three specimens tested at this condition were fractured at up to the respective UTS point for each testing condition since thethe die corner region rather than at the dome apex. Thus, the flow material data after this point is no longer meaningful for the even-curve in this case has a maximum strain of less than 0.2. Nonethe- tual and further use in analyses. Similar effects of the temperatureless, when Al6061 specimens were bulged at the lower strain rate and strain rate on the flow stress curves were also observed; that(0.0013 sÀ1) at 300 °C, the maximum strain value is shown to be as is the flow stress decreases with increasing temperature andhigh as 0.9. Therefore, at a low strain rate, the formability of decreasing strain rate. All three specimens that were pulled underAl6061 may continue to improve with increasing temperature the same testing conditions provided almost identical flow curves,even above 200 °C. showing very reliable and repeatable results. In addition, the max- imum elongation of Al5052 under the uniaxial loading condition3.2. Tensile test results was found to increase considerably at elevated temperature levels between 200 °C and 300 °C, however, such an increase was not Due to the limitations of the electromechanical MTS system, a observed in the case of Al6061 alloy after 200 °C. These observa-constant strain rate control during the tests was found to be tions agreed well with the reported results by Novotny and Geigerrather difficult. Therefore, in this study a constant cross-head [5,6] and Li and Ghosh [8]. In their tensile test study, the formabil-speed was used to provide initial strain rates of 0.0013 and ity of Al5xxx continuously increases with the temperature up to0.013 sÀ1, which were the same strain rates used in the bulge 300 °C, while that of the Al6xxx would increase up to 200 °Ctests. The calculated true-stress–strain curves from the tensile and the maximum elongation starts to decrease with further in-tests are shown in Fig. 7. The flow curves are only presented here crease in temperature. Al5052 – base location Al6061 – base location Room Temp. 25µm 25µm 100°C 25µm 25µm 200°C 25µm 25µm 240°C N/A 25µm Fig. 9. Temperature effect on grain structure at the base location.
  7. 7. 2428 S. Mahabunphachai, M. Koç / Materials and Design 31 (2010) 2422–2434 35 Table 2 Al5052 DOE of closed-die hydroforming. 30 Al6061 Runorder Stdorder Temp. Pressure Grain size [micrometer] 1 17 1 0 25 23.4 23.6 22.5 21.8 2 13 À1 À1 21.4 19.4 3 10 1 À1 20 4 8 À1 0 5 9 À1 1 15 6 4 1 À1 7 16 1 À1 10 8 1 À1 À1 9 14 À1 0 10 5 1 0 5 11 12 1 1 12 6 1 1 0 13 7 À1 À1 Room Temp. 100°C 200°C 14 3 À1 1 15 15 À1 1 Fig. 10. Effect of temperature on the material grain size. 16 18 1 1 17 11 1 0 18 2 À1 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 tensile 3.3. Microstructure analysis on bulged samplestests, it was possible to obtain reliable flow stress values at lowstrain values (below 0.2), which was not possible in the bulge tests In order to better understand the effect of temperature anddue to the limitations dictated by the spherical assumptions strain rate on the response of these Al alloys, a microstructure(i.e., low h/a ratio). analysis was performed on the bulged samples. The grain structure Al5052 – apex location Al6061 – apex location SR=0.0013 s-1 200°C 25µm Room Temp. 25µm SR=0.013 s-1 200°C 25µm Room Temp. 25µm Fig. 11. Strain rate effect on grain structure at the apex location. Die insert Die insert Fig. 12. Geometries of closed-die insert.
  8. 8. S. Mahabunphachai, M. Koç / Materials and Design 31 (2010) 2422–2434 2429 20 The effect of temperature on the grain structure was investi- gated at the base location of the cut strips. The base location was selected because the grains in this region did not undergo any 15 deformation (strain and strain rate independent); thus, represent- Pressure [MPa] ing the microstructure changes caused merely by the temperature effect. The microscopic images of the grain structures are illus- trated in Fig. 9 for both Al5052 and Al6061. Note that the grain 10 Al5052-300C-15MPa structures from the specimens that were bulged at 240 °C and Al5052-300C-20MPa higher could not be clearly seen, which may have been caused by Al6061-200C-10MPa the significant microstructural changes, most likely recrystalliza- 5 Al6061-200C-15MPa tion and growth of grains or precipitates, as the bulging tempera- Reference Profile ture enters into the ‘‘warm forming” regime (i.e., temperature 0.3Tm, where Tm indicates melting point). For the samples that 0 the grain boundary could be clearly indicated, the grain size was 0 20 40 60 80 100 Time [sec.] measured based on the ASTM Standard E112-88 (i.e., Mean Lineal Intercept or Heyn’s method). The grain size measurement results Fig. 13. Hydroforming pressure profiles. are presented in Fig. 10 where slightly larger grains were observed at elevated temperature levels although the difference is statisti- cally insignificant. According to the Hall–Petch relation [21,22],(i.e., size, shape, and distribution) was investigated as the grain material with larger grain size was predicted to have lowerstructure was known to largely influence the overall material re- strength. Despite the fact that such a case was observed in thissponse. The bulged specimens were cut along the centerline and study, it is believed that the drop in the flow stress curve is notsmall sample strips were removed from the center region as shown due to the slightly and statistically indifferent grains, but mainlyin Fig. 8. These sample strips were polished and etched in Keller’s due to the additional slip lines activated due to the elevatedreagent (2.5 ml HNO3, 1.5 ml HCl, 1 ml HF, and 95 ml water) to re- temperature.veal the undeformed grain structure at the ‘‘base location” and the The effect of the strain rate was also investigated through thedeformed grain structure at the ‘‘mid-point” and ‘‘apex” regions as grain structure analysis. Unlike in the temperature effect study,depicted in Fig. 8. Most of the grains were found to have granular the location of interest on the cut specimen was shifted to be atstructure at the base location (i.e., undeformed grains), while elon- the apex of the dome rather than at the base location as mostgated and elliptical grain structures were observed at the mid- deformation dependent characteristic could be seen most in thispoint and the apex locations, respectively. The grains at these region. The microscopic images of the specimens bulged at differ-regions were elongated or stretched as they underwent a large ent strain rates are shown in Fig. 11. Unfortunately, no significantplastic deformation amount during the bugle tests. difference was observed in terms of the grain structure between (a) 200°C Al5052 300°C 10MPa 15MPa 20MPa (b) 200°C Al6061 300°C 10MPa 15MPa 20MPa Fig. 14. Closed-die hydroformed samples: (a) A5051 and (b) Al6061.
  9. 9. 2430 S. Mahabunphachai, M. Koç / Materials and Design 31 (2010) 2422–2434 25 Al5052 20 height [mm] 15 200C-10MPa 200C-15MPa 10 200C-20MPa 300C-10MPa 300C-15MPa 5 300C-20MPa Die Profile 0 0 20 40 60 80 100 120 width [mm] 25 Al6061 20 height [mm] Fig. 16. Response surface plots. 15 200C-10MPa 10 200C-15MPa slope of 0.22 MPa/s was used as a reference input (Fig. 13). The ac- 200C-20MPa tual hydroforming pressure profiles, recorded during the tests 300C-10MPa using a pressure transducer, were shown to closely follow the ref- 5 300C-20MPa erence pressure input profile (Fig. 13). Die Profile After each test, the hydroformed parts were measured using the 0 stereoscopic CCD cameras with ARAMIS software to obtain full sur- 0 20 40 60 80 100 120 face profiles as illustrated in Figs. 14 and 15. For the assurance of width [mm] process repeatability, three experiments were conducted for each case. An average value is reported unless otherwise is stated in thisFig. 15. Profiles of hydroformed parted at different temperature and pressurelevels. section. 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,the two strain rate values used in this study (0.0013 and 0.013 sÀ1). both Al5052 and Al6061 sheet blanks showed poor formability,Nonetheless, the effect of the strain rates on the material response and a fracture occurred in the area of the die radius at the center(i.e., flow curve) and the formability (i.e., maximum elongation) when the pressure was increased from 10 to 15 MPa for Al5052was obvious, especially at elevated temperature levels as discussed and from 15 to 20 MPa for Al6061. As the temperature was in-in the previous section. Hence, it can be concluded that it is not the creased to 300 °C, the formability of Al5052 sheet blanks appearedchange in grain size, but the thermally activated dislocation mo- to improve, and no premature rupture was observed up to 20 MPa.tion that causes the formability increase in warm forming. However, for Al6061, an increasing temperature only reduced the material strength (i.e., higher profile at the same forming pressure when the temperature was increased), while the elongation prop-4. Closed-die hydroforming experiments erties did not change. With the increasing temperature, all Al6061 specimens showed fractures at the central die radius area once a A set of closed-die warm hydroforming experiments were con- certain part height was reached. This observation agrees well withducted on the same alloys using a die insert as shown in Fig. 12. the flow curve plots of Al6061 (Fig. 7), in which the maximumThese experiments were conducted under a design of experiment strain (i.e., elongation) value did not change as temperature in-(DOE) plan, as tabulated in Table 2 (i.e., 18 runs for each alloy), creased, but the flow stress values were observed to significantlyto obtain a quantified understanding of the effect of temperature decrease. Furthermore, based on the comparison of the part pro-and pressure on the die cavity filling and thinning (forming limits). files in Fig. 15, it is clearly shown that Al5052 has a better formabil-During the experiments, a constant blank holder force of 1000 kN ity when compared to Al6061 under these forming conditions (i.e.,was used, as in the bulge tests, to clamp the specimens at the uniform temperature distribution at 200 and 300 °C and linearperiphery. A linearly increasing (ramp-up) pressure profile with a pressure profiles up to 15–20 MPa). Finally, it is important to pointTable 3Percentage of die filling under different process conditions. 10 MPa (%) 15 MPa (%) 20 MPa (%) Al5052 200 °C 77.2 94.2 93.7 300 °C 95.2 95.6 96.2 Al6061 200 °C 60.7 80.8 83.2 300 °C 94.3 n/a 94.5
  10. 10. S. Mahabunphachai, M. Koç / Materials and Design 31 (2010) 2422–2434 2431 60 Table 4 Al5052 200C-10MPa Material parameters for Al5052 and Al6061 used in FEA. 50 300C-15MPa Material parameters AI5052 AI6061 300C-20MPa Die Profile Modulus of elasticity, E (MPa) 70,300 68,900 40 Poisson’s ratio, v 0.33 0.33 % Thinning Mass density, q (Mg/mm3) 2.68E À 09 2.70E À 09 30 Yield strength, r0 (MPa) 89.6 55.2 20 Table 5 ee _ Material constants ðr ¼ K n m Þ from tensile and bulge tests at different temperatures. 10 Test Material Temp. (°C) K (MPa) n m 0 0 10 20 30 40 50 60 Bulge test AI5052 23 455 0.14 0.010 200 412 0.33 0.075 Radial distance [mm] 300 401 0.55 0.135 AI6061 23 483 0.11 0.013 50 100 497 0.17 0.020 Al6061 200C-10MPa 200 503 0.12 0.075 200C-15MPa 40 Die Profile Tensile test AI5052 23 777 0.45 0.000 100 966 0.50 0.027 % Thinning 30 200 437 0.28 0.051 300 253 0.09 0.151 AI6061 23 979 0.38 À0.006 20 100 1058 0.41 0.009 200 880 0.36 0.046 300 474 0.21 0.114 10 0 filling ratio and thinning) were measured and reported in Table 3 0 10 20 30 40 50 60 and in Figs. 16 and 17. Note that for thinning measurements, the Radial distance [mm] hydroformed specimens were cut into two halves with an offset Fig. 17. Thickness profiles in radial direction. of 10 mm from the center line, and the thickness of the bigger half was measured using a micrometer attached with conical shape tips at several locations along the radial directions. In addition, for aout that the die corner rupture was observed in the case of Al6061 meaningful comparison of the thinning, only the specimens with-blanks formed at 300 °C and 15 MPa. As a result, their profiles are out the fracture were used.excluded in Fig. 15 as well as in the rest of the analysis. When cavity filling comparisons in Table 3 and Fig. 16 are In order to quantify the effects of the temperature and pressure considered, at 200 °C, an increasing pressure leads to an increaseon the formability of the hydroformed parts, two variables (the die in the cavity filling ratio for A5052 (77–93%) and Al6061 (60–83%). Total Equivalent Plastic Strain Fig. 18. Two-dimensional axisymmetric numerical modeling of the closed-die warm hydroforming: FE model, predicted and actual deformed part.
  11. 11. 2432 S. Mahabunphachai, M. Koç / Materials and Design 31 (2010) 2422–2434 25 (i.e., for Al5052: p = 0.238 for temperature and p = 0.246 for pres- Al5052 sure; for Al6061: p = 0.151 for temperature and p = 0.350 for pressure). 20 In terms of thinning comparisons as shown in Fig. 17, first of all, the highest thinning, and hence fracture in some cases, occurred Height [mm]. 15 around the die radius at the center region of the part. Second, 200C-10MPa-Exp the effect of temperature and pressure were as expected (i.e., 200C-10MPa-FEA-Bulge increasing pressure and temperature leads to increasing thinning 10 200C-10MPa-FEA-Tensile in general). 300C-15MPa-Exp Based on this result discussion, the use of a uniform tempera- 300C-15MPa-FEA-Bulge 5 ture distribution (i.e., isothermal conditions) and a linearly increas- 300C-15MPa-FEA-Tensile ing pressure profile (i.e., ramp-up pressure input) may not be the Die Profile most efficient approach to increase the cavity filling and reduce 0 thinning (i.e., part formability). Thus, process optimization investi- 0 10 20 30 40 50 60 gation is needed to determine optimal process conditions (e.g., var- Width [mm] iable loading profiles: pressure and blank holder force, and 25 temperature). This optimization study would require the use of fi- Al6061 nite element analysis (FEA) tool. In the following section, FEA mod- els of the hydroforming process will be developed and the material 20 properties obtained from the bulge and tensile tests will be vali- dated for their accuracy and applicability in predicting the Height [mm]. closed-die profiles and thinning values as measured and reported 15 in this section. 200C-10MPa-Exp 10 200C-10MPa-FEA-Bulge 5. Numerical modeling, validation and comparison of material 200C-10MPa-FEA-Tensile behaviors 200C-15MPa-Exp 5 200C-15MPa-FEA-Bulge 200C-15MPa-FEA-Tensile In this section, Finite element models of the closed-die warm hydro- Die Profile forming process (Fig. 18) were developed using MSC.Marc2007r1 soft- 0 0 10 20 30 40 50 60 ware. Since the problem at hand was an axisymmetric type, only a 2-D half-model analysis was developed. In the model, the sheet blank was Width [mm] modeled using deformable, solid, quad-4 elements. Four elementsFig. 19. Comparison of hydroformed part profiles for A5052 and Al6061 alloys at across the blank thickness were used. Both ends of the blank were fixeddifferent process conditions. (no displacement) to reflect on the actual boundary condition of the process where the blank was tightly clamped between the upper and lower dies to prevent any radial flow of the material into the die cavity.However, for 300 °C, the same cannot be said. A similar observation Hydroforming pressure was applied from the bottom side of the blankis made for an increasing temperature at low pressure value with an increasing rate of 0.22 MPa/s until the pressure reached the pre-(10–15 MPa). However, in general, when a regression analysis was set values. According to the actual pressure profiles recorded during themade for the entire ranges of temperature and pressure, their effect experiments, the pressure pump and regulator provided close control ofon the cavity filling ratio was found to be statistically insignificant the pressure profile as shown in Fig. 13. Thus, a ramp pressure inputTable 6Simulation cases. Exp. case FEA run Material Temp. (°C) Pressure (MPa) Mat. flow curve Case 1 1 AI5052 200 10 Bulge 2 AI5052 200 10 Tensile Case 2 3 AI5052 300 15 Bulge 4 AI5052 300 15 Tensile Case 3 5 AI6061 200 10 Bulge 6 AI6061 200 10 Tensile Case 4 7 AI6061 200 15 Bulge 8 AI6061 200 15 TensileTable 7Percentage of die filling comparisons. Material Temp. (°C) Pressure (MPa) Percentage of die filling Experiment FEA-bulge (%error) FEA-tensile (%error) Al5052 200 10 77.2 86.1 (11.5) 79.2 (2.6) 300 15 95.6 98.2 (2.7) 98.2 (2.7) Al6061 200 10 60.7 67.6 (11.4) 59.8 (1.5) 200 15 80.8 83.8 (3.7) 75.1 (7.1)
  12. 12. S. Mahabunphachai, M. Koç / Materials and Design 31 (2010) 2422–2434 2433 which makes the thickness measurement physically possible. The other output chosen for comparison purpose was the part profiles (or percentage of die filling). In addition, with the two flow curves obtain from tensile and bulge tests at each temperature level, a total of eight simulation runs were carried out and their results are 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 was found the tensile flow curves provided better profile predictions e at low pressure (i.e., low strain) levels (10 MPa, where 0:2), while the bulge flow curves did so at the high pressure (i.e., high e strain) levels (15 MPa, where 0:4). The observation could be Fig. 20. Thickness measurement. well explained by the limitations and assumptions associated with each test method and the derivations of the material flow curveswith a slope of 0.22 MPa/s was utilized in all validation cases. Coulomb based on the raw test data. Specifically, for tensile tests conductedfriction model was selected with a friction coefficient of 0.1 for all con- at 200 °C, the maximum strain levels were found to be around 0.2tact surfaces in this study. for Al5052, and 0.15 for Al6061. Therefore, FEA predictions for high Material input parameters: modulus of elasticity (E), poisson’s strain levels would require extrapolation of the material flowratio (v), mass density (q), and yield strength (r0) for both curves. On the other hand, a higher strain levels could be achievedAl5052 and Al6061 are given in Table 4, while the material flow in the bulge tests, e.g., at 200 °C, the achievable strain was aroundcurves obtained from the tensile and bulge tests at different tem- 0.5 for Al5052 and about 0.35 for Al6061; and thus, the materialperature levels were modeled using Field–Backofen equation (i.e., data from the bulge test provided better FEA predictions at high ee _rate power law: r ¼ K n m ) as tabulated in Table 5. strain or pressure levels. Furthermore, with the assumption of a In order to validate the FE models, four experimental cases non-spherical dome shape of the bulge specimen below the h/a va-(Table 6) were selected which included the closed-die hydroform- lue of 0.2 (corresponding to a strain value of 0.08), the materialing results for both Al5052 and Al6061 at elevated temperatures data below this threshold value was excluded for the flow curve(200–300 °C) and two different pressure levels (10 MPa and determination in bulge testing case, which in turns made the FEA15 MPa). The hydroformed parts in these cases were fracture-free, predictions based on the bulge flow curves less accurate at low (a) 60 200C-10MPa-Exp 200C-10MPa-FEA-Bulge 50 200C-10MPa-FEA-Tensile 300C-15MPa-Exp 300C-15MPa-FEA-Bulge 40 300C-15MPa-FEA-Tensile Die Profile % Thinning 30 20 Al5052 200C-10MPa 300C-15MPa 10 Δ%Thinning FEA-Bulge FEA-Tensile FEA-Bulge FEA-Tensile Ave. 2.8 0.8 5.5 5.3 Max. 6.2 3.8 18.2 18.0 0 0 10 20 30 40 50 60 Radial distance [mm] (b) 60 Al6061 200C-10MPa 200C-15MPa Δ%Thinning FEA-Bulge FEA-Tensile FEA-Bulge FEA-Tensile 200C-10MPa-Exp Ave. 0.1 2.5 1.4 6.8 200C-10MPa-FEA-Bulge 50 Max. 6.1 4.6 4.1 12.3 200C-10MPa-FEA-Tensile 200C-15MPa-Exp 40 200C-15MPa-FEA-Bulge % Thinning 200C-15MPa-FEA-Tensile Die Profile 30 20 10 0 0 10 20 30 40 50 60 Radial distance [mm] Fig. 21. Thinning comparisons in radial direction for: (a) Al5052 and (b) Al6061.