Evaluation of Formability and Drawability of Al 5182-O Using a Servo Drive Press DISSERTATION

Transcription

1 Evaluation of Formability and Drawability of Al 5182-O Using a Servo Drive Press DISSERTATION Presented in Partial Fulfillment of the Requirements for the Degree Doctor of Philosophy in the Graduate School of The Ohio State University By Tingting Mao Graduate Program in Industrial and Systems Engineering The Ohio State University 2014 Dissertation Committee: Dr. Taylan Altan, Advisor Dr.Jerald Brevick Dr. Blaine Lilly

2 Copyright by Tingting Mao 2014

3 Abstract The part drawability in stamping process can be improved by many factors, including blank geometry, lubricant, surface finish, the deformation speed and the blank holder force. Lowering the friction between the tool and the workpiece can reduce fracture potential in the part. Friction is affected by contact pressure, temperature, sliding velocity, die material/coating and surface texturing. In this study, only the effects of lubrication and surface were investigated. Various lubricants such as dry film lubricant and wet lubricants, which were developed for aluminum alloys, were evaluated using cup draw test. Tribological test such as cup draw test was utilized to evaluate the lubricity of the selected lubricants. It was found out that dry film lubricant performed better than wet lubricants in the deep drawing of aluminum alloys. Mill finish and Electrical Discharge Texturing (EDT) that are surface topographies for aluminum alloys, were evaluated by conducting cup draw test in the study. It was concluded that EDT can provide better lubrication condition at the interface between workpiece and die, compared to Mill Finish. Blank geometry can be optimized to improve the part drawability. In this study, different blank geometries were studied by using FE simulations to optimize the blank geometry to be used in the servo press tryouts. It was found out blank with an optimal blank shape, such as four corners chamfered blank shape, could give less wrinkling. ii

4 Designing a proper tooling also can improve the part drawability by selecting the optimum punch/ die radii and punch and die clearance. In this study, FE simulations were conducted to determine the optimum punch and die radii and clearance between punch and die. By using a mechanical servo press, which can 1) vary the ram speed at any slide position and 2) vary the blank holder force through the stroke can improve the part by selecting the optimal deformation speed and blank holder force. In this study, the effect of ram speed and variable blank holder profile on part drawability was investigated. It was found out that a higher speed can form a better part. iii

5 Dedication This document is dedicated to my husband (Yang Wang), my daughter (Alice Wang) and my son (Michael Wang) iv

6 Acknowledgments I would like to express the deepest appreciation to my advisor, Dr.Taylan Altan for giving me the opportunity to pursue my graduate studies under his guidance. His encouragement and guidance are the main reasons for this research work to even be possible. I sincerely thank my co-advisor, Dr. Jerald Brevick, for all his inputs and kind support during last four years. I also thank my committee member Dr. Blaine Lily for his inputs and kind support. I would like to thank Dennis O connor and Milan Jurich from Honda for providing all the support, especially for providing the aluminum blanks and tooling for conducting experiments required for my study. I would also like to thank Shrini Patil from AIDA and Cliff Hoschouer from Shiloh Inc. for providing AIDA s press and Shiloh die tooling for conducting experiments. I thank my colleagues of the CPF, Adam Groseclose, Xi Yang, Ali Fallahiarezoodar, Siddharth Kishore, Ganapathy Srinivasan, Suraj Appachu, Chong Li and Xia Jin for the discussion and assistance during the course of my study. I especially want to think Long Ju for his discussion and help for my project. I thank my family and friends for all their support in this endeavor. At last, I would like to thank my husband Yang Wang and two of my kids (Alice Wang and Michael Wang) for their support during my PHD program. I would never have this achievement without their support. v

7 Vita B.S. Mechanical Engineering, Henan University of Technology M.Sc. Mechanical Department, Southern Illinois University Edwardsville 2010 to present...phd student, Integrated Systems Engineering, The Ohio State University Publications Mao, T., and Altan, T., Aluminum sheet forming for automotive applications, Part 1, Stamping Journal, January/February issue, p 12. Mao, T., and Altan, T., Aluminum sheet forming for automotive application, Part II, Stamping Journal, March/April issue, p 12. Mao, T., et al., Applications of FE simulation for Stamping and Blanking of Sheet Materials, Int. Conf. on New Developments in Sheet Metal Forming, May 13-14, 2014, IFU-Stuttgart. Fields of Study Major Field: Industrial and Systems Engineering vi

8 Table of Contents Abstract... ii Dedication... iv Acknowledgments... v Vita... vi Publications... vi Fields of Study... vi Table of Contents... vii List of Tables... xiii List of Figures... xvi Chapter 1. Introduction... 1 Chapter 2. Research Objectives and Approach... 3 Chapter 3. Literature Review on Aluminum forming Material properties and formability Wrought alloys designations Aluminum 5182-O material Determination of materials properties by the uniaxial tensile test vii

9 3.2.1 Limitations of the tensile test Determination of material properties by the viscous pressure bulge (VPB) test Conditions at tool/material interface Lubricants Surface finish of aluminum sheets Tests for evaluation of lubricants Cup drawing test (CDT) Other tests available for evaluation performance of lubricants Design rules for tooling (deep drawing and stretch forming) Simulation/deformation zone Servo press application on aluminum forming Chapter 4. Determination of Material Properties of Al 5182-O using Uniaxial Tensile Test, Viscous Pressure Bulge Test (VPB tests) and Dome Test Determination of sheet material properties by the uniaxial tensile test Determination of sheet material properties by the VPB test Comparison of flow stress data obtained from tensile test and bulge test Chapter 5. Experimental Evaluation of Selected Lubrication Conditions to Improve Stamping Quality Objectives viii

10 5.2 Approach Preparation of Experimental Evaluation for Various Lubrication Conditions Before starting cup drawing tests the following tasks are completed Procurement of Sheet Materials and Lubricant Samples Preliminary Evaluation of Lubrication Conditions Lubrication Tests, Analysis and Results Principles of CDT-Cup Draw Test Preliminary Finite Element Simulation of the Cup Draw Test (CDT) Experimental Set-up Cup Draw Tests Experimental Results FE simulations to predict Coefficient of friction (CoF) at tool-work piece interface Comparison of cup thickness variation obtained from FE simulations and experiments Summary and Conclusions Chapter 6. Design of a Tooling for the Servo Press Studies in Forming Aluminum Alloys FE simulations for tooling design Determination of die radii ix

11 6.1.2 Determination of clearance between punch and die The effect of CoF on the formability of the part FE simulation results The optimum parameters of tool design Summary and conclusions FE simulations for protrusion design FE simulations for simulations # (Angle fixed, radii varied) FE simulations for simulations # (Radii fixed, angle varied) Conclusions Final die design and first tryouts with 820KN blank holder force using a Honda servo press Final Die design First tryouts conducted at Honda with this final die set Summary Summary Chapter 7. Validation of FE model and Prediction of Future Tryouts with Honda Die using FE Simulation FE simulations of Honda part with 820KN blank holder force (without spacers) FE model set up in Pamstamp x

12 7.1.2 Material model in Pamstamp Comparison between Simulation results and experiment (curved shape, no spacer) Conclusions FE simulation of Honda part with 820KN blank holder force (with spacers) FE simulation of drawing process with spacers FE simulation results with curved blank shape (Case III) Conclusions FE simulation results with 250KN blank holder force for AIDA s tryouts (without spacer) FE simulation set up FE simulation results Summary and Conclusions Chapter 8. Conduct Tryouts on the Servo Press using Shiloh Die Shiloh die tooling and AIDA s 300 ton servo press FE simulations to predict the process parameters will be used in the tryouts FE mode set up Investigate the effects of blank shapes on part drawability Investigate the effect of lubrication xi

13 8.2.4 Investigate the effect of BHF Conduct tryouts on the servo press Preparation for experiments Preliminary tryout results (blank applied with dry film lubricant) Second tryouts (using blanks without applying lubricants) Third tryouts (using blanks with dry film lubricant) Observation and discussions Chapter 9. Summary, Conclusions and Future Work Summary Evaluation of lubrication Tool design for servo press s study Investigation of the effect of ram speed and blank holder force on part drawability Conclusions Evaluation of lubrication Tool design for servo press s study Investigation of ram speed Future work References xii

14 List of Tables Table 3.1: Mechanical properties of 5000 series aluminum sheet alloys (SAKURAI, 2008)... 7 Table 3.2: Two lubricants, KTLN 16 and Drylube E1 (Wagner et al., 2002) Table 3.3: Different types of lubricants used for aluminum sheet metal forming (Meiler and Jaschke, 2005) Table 3.4: Ra values of EDT, DF and MF samples (Hartfield-Wunsch et al., 2011) Table 4.1: Tensile test data provided by Honda R&D Table 4.2: Parameters used in the viscous bulge test Table 4.3: K and n values obtained from tensile and bulge tests Table 5.1: Lubricants and their details as provided by lube companies Table 5.2: FE Simulation Parameters for obtaining BHF range for CDT Table 5.3: Measurement results using Microderm CMS Table 5.4: Parameters of Cup Drawing Tests (CDT) Table 5.5: Prediction of Flange Perimeters by FE simulations Table 6.1: List of Simulations to determine the punch/die radii Table 6.2: Input parameters to FE simulations Table 6.3: List of Simulations to determine the clearance between punch and die Table 6.4: List of simulations to investigate the effect of CoF on the formability of the part xiii

15 Table 6.5: The optimum design of the tool Table 6.6: list of simulations to determine R1, R2 and α Table 7.1: Input parameters for Barlat 2000 yield criteria, obtained from (Carsley et al., 2013) Table 7.2: Input parameters for FE simulation Table 7.3: Simulation matrix for rectangular blank shape (the die, punch and blank holder are assumed to be rigid) Table 7.4: Simulation matrix for curved blank shape (the die, punch and blank holder are assumed to be rigid) Table 7.5: Simulation matrix for rectangular blank shape (the die, punch and blank holder are assumed to be rigid) Table 8.1: Simulation inputs (blank shapes A&B are shown in Figure 8.4) Table 8.2: FE simulation case study (blank shapes A&B are shown in Figure 8.4) Table 8.3: Simulation matrix for friction study Table 8.4: Maximum thinning for simulation cases I-IV Table 8.5: Simulation matrix of variable BHF study (blank shape B is shown in Figure 8.4) Table 8.6: Test matrix of Task I Table 8.7: Test matrix of Task II Table 8.8: Test matrix for preliminary test (Blank shapes are shown in Figure 8.4) Table 8.9: Investigation on the effect of speed (crank motion) on Aluminum part formability for second tryouts xiv

16 Table 8.10: Investigation on the effect of speed (constant speed) on Aluminum part formability Table 8.11: Investigation the effect of ram speed on part drawability xv

17 List of Figures Figure 3.1: Wrought aluminum alloy designation system (TheAluminumAutomotiveManual, 2002)... 6 Figure 3.2: The specific Banana diagram (DuckerWorldwide, 2011)... 6 Figure 3.3: Al 5xxx and 6xxx alloys used in car body components (Hirsch, 2004)... 7 Figure 3.4: 5xxx series alloys used for hood, decklid inner and fender reinforcements (Lidguard, 2009)... 8 Figure 3.5: 6xxx series alloys (6111-T4) used for hood and fender skins in Jaguar (Lidguard, 2009)... 9 Figure 3.6: Stretcher strains occurred in Al 5182-O (K.Siegert and S.Wagner, 1994) Figure 3.7: Geometry before and after the test (t0 = original sheet thickness, td = sheet thickness at the apex, hd = bulge height, P = hydraulic pressure, rc = die fillet radius, dc = die cavity diameter, rd=radius of curvature) (Venkatesan, 2010) Figure 3.8: Schematic on the left shows the sheet clamped between the upper and lower die. Schematic on the right shows the sheet bulged by the pressurizing medium into the upper die cavity (Venkatesan, 2010) Figure 3.9: Increasing process window by reducing friction (Meiler and Jaschke, 2005) Figure 3.10: Oil base and dry film lubricant applications (Meiler and Jaschke, 2005) xvi

18 Figure 3.11: The maximum possible drawing depth and maximum applicable blank holder forces for the two lubricants (Wagner et al., 2002) Figure 3.12: Drawing performance of different lubricants. The larger the blank holder force used in deep drawing cups without fracture, the better is the lubricant (Meiler and Jaschke, 2005) Figure 3.13: SEM image of an Electrical Discharge Texture surface (left) and mill finish surface (right) on AA 6016 sheet (TheAluminumAutomotiveManual, 2002) Figure 3.14: Schematic of cup drawing (Kim, 2008) Figure 3.15: Schematic of CDT Tooling at CPF/EWI (Subramonian et al., 2011) Figure 3.16: Schematic of deep drawing (Altan, 2001) Figure 3.17: Strains of at least 2 percent are necessary when stamping aluminum to reduce springback (Thomas and Altan, 1999) Figure 3.18: Panel formed at (a) 41 mm/s and (b) 103 mm/s [Hayashi et al, 2009] Figure 3.19: Different slide motions used to form aluminum panels [Hayashi et al, 2009] Figure 3.20: Examples of different slide motions of servo press (Groseclose, 2008) Figure 3.21: The flexibility of slide motion in servo presses (Miyoshi, 2004) Figure 3.22: Four basic BHF formability windows (Kitayama et al., 2010) Figure 4.1: Dimensions of tensile test specimen, provided by Honda R&D Figure 4.2: True stress-true strain curves in rolling, diagonal and transverse directions. 34 Figure 4.3: Bulge test data provided by EWI/CPF (Mao, 2013) xvii

19 Figure 4.4: Forming limit curve of Al 5182-O, with thickness of 1.1 mm, obtained from (Carsley et al., 2013) Figure 4.5: Comparison on flow stress curve between tensile and bulge tests Figure 5.1: Quarter model used in simulation of Cup Drawing Test using PAM-STAMP Figure 5.2: Microderm CMS tool used to measure the thickness of applied lubricants coating (UPA, 2013) Figure 5.3: Test Procedure for the evaluation of lubricants Figure 5.4: Flange perimeter recorded for 13 lubricant tests at 16 tons BHF (L2, L4 and L8 failed at 16 ton BHF ) Note : The y-axis on the graph does not start from 0. The error bands show the deviation between samples Figure 5.5: Flange perimeter recorded for 5 lubricant tests at 17 tons BHF (L1, L3, L5, L10 L13 failed at 17 ton BHF ) Note : The y-axis on the graph does not start from 0. The error bands show the deviation between samples Figure 5.6: Comparison between MF and EDT at 16 tons BHF Figure 5.7: Comparison between MF and EDT at 17 tons BHF Figure 5.8: Comparison of flange perimeters obtained from simulation and experiment to predict the coefficient of friction at 16 ton BHF for AL5182O/MF with 1.5mm thickness Figure 5.9: Comparison of flange perimeters obtained from simulation and experiment to predict the coefficient of friction at 17 ton BHF for AL5182O/MF with 1.5mm thickness xviii

20 Figure 5.10: Cup samples deep drawn with lubricants F and I Figure 5.11: Thickness comparison between experiment and simulation in rolling direction (cup drawn with lubricant F) Figure 5.12: Thickness comparison between experiment and simulation in transverse direction (cup drawn with lubricant F) Figure 5.13: Thickness comparison between experiment and simulation in rolling direction (cup drawn with lubricant I) Figure 5.14: Thickness comparison between experiment and simulation in transverse direction (cup drawn with lubricant I) Figure 6.1: 2SL Rdpnl (Provided by Honda) Figure 6.2: FE model and the dimensions of the blank Figure 6.3: Thinning distribution of simulations #1.1,1,2,1,3 and Figure 6.4: Thinning distribution of simulations #1.1,2.1 and Figure 6.5: Rupture risk distribution of simulations #1.1,2.1 and Figure 6.6: Thinning distribution of simulations 3.1 (CoF=0.08), 3.2 (CoF=0.12) and 2.1 (CoF=0.1) Figure 6.7: Rupture risk distribution of simulations 3.1 (CoF=0.08), 3.2 (CoF=0.12) and 2.1 (CoF=0.1) Figure 6.8: Material draw-in distribution of simulations 3.1 (CoF=0.08), 3.2 (CoF=0.12) and 2.1 (CoF=0.1) Figure 6.9: Punch with protrusion features Figure 6.10: Honda protrusion design based on inner door panel xix

21 Figure 6.11: A protrusion feature designed by CPF Figure 6.12: A-A section view of protrusion Figure 6.13: Thinning distribution for simulations # Figure 6.14: Fracture probability plot using the FLD for simulations # Figure 6.15: Thinning distribution for simulations # Figure 6.16: Fracture probability plot using the FLD for simulations # Figure 6.17: Overall view of tool Figure 6.18: Punch dimensions Figure 6.19: Material draw-in when the die set does not have draw beads (draw depth is 155 mm) Figure 6.20: Blank holder plate dimensions Figure 6.21: 2D section view of draw beads design Figure 6.22: Protrusion design Figure 6.23: (a) Rectangular shape blank and (b) curved shape blank Figure 6.24: Material draw-in with curved blank shape, provided by Honda EGA Figure 6.25: This part was drawn up to 100 mm depth and crack occurred afterwards, provided by Honda EGA Figure 7.1: Thinning distribution of the part (no spacers, curved shape), at stroke of 80 mm (max thinning is 44%) Figure 7.2: Schematic of a spacer used during forming process [Faass et al, 2008] Figure 7.3: Material draw-in comparison between simulation and experiment xx

22 Figure 7.4: Thinning distribution of the part (Case I), at stroke of 100 mm (max thinning is 18%) Figure 7.5: Thinning distribution of the part (Case I), at stroke of 150 mm (max thinning is 22%) Figure 7.6: Thinning distribution of the part (Case II), at stroke of 100 mm (max thinning is 19%) Figure 7.7: Thinning distribution of the part (Case II), at stroke of 150 mm (max thinning is 24%) Figure 7.8: Material draw-in comparison between simulation and experiment Figure 7.9: Thinning distribution at stroke of 155 mm, with curved blank shape Figure 7.10: Rectangular shape with four corners chamfered Figure 7.11: Thinning distribution of simulations A,B,C and D Figure 7.12: Thinning comparison along A to B between simulations A,B,C and D Figure 7.13: Probability of Fracture (using FLD) from Simulations C (left one) and D (right one) Figure 8.1: Shiloh die installed in the AIDA s press Figure 8.2: Shiloh die front and cross-section view (Adam Groseclose, 2014) Figure 8.3: FE modeling and inputs in Pam-stamp Figure 8.4: Blank shapes designed for future tryouts Figure 8.5: Thinning distributions at stroke of 75mm, BHF=150KN Figure 8.6: Max. thinning values at selected strokes, under Cof 0.1 and Cof Figure 8.7: Thinning values at locations A and B xxi

23 Figure 8.8: Thinning distribution under BHF 250KN Figure 8.9: Profiles of designed blank holder force Figure 8.10: Simulation results (thinning) when different blank holder profiles (A, B and C, see Figure 8.9) were used Figure 8.11: Guides used to position the blanks (Blank shape B (see Figure 8.4) was used in this figure) Figure 8.12: Different ram speed profiles will be used in Task I Figure 8.13: Varied blank holder force profiles prepared for Task II Figure 8.14: Samples with different test conditions for the preliminary tests (see Table 8.8 for the test conditions) Figure 8.15: Comparison on samples 7_a, 8_a and 6_a (see Table 8.9 for the test conditions) Figure 8.16: Locations A and B in the aluminum sample Figure 8.17: Comparison between samples 9_a and 10_a by using different constant speeds during deformation (310 mm/s and 50.5 mm/s), see Table 8.10 for the test conditions Figure 8.18: Ram speed vs punch stroke curves used in the tryouts (output data from AIDA s press) Figure 8.19: Press total load vs slide position when 18 spm and 10 spm were used Figure 8.20: Formed parts with constant speeds of 50.6 mm/s and 310 mm/s during the deformation (see Figure 8.11 for the test conditions) xxii

24 Figure 8.21: Formed parts with ram speed of 10 spm and 1 spm (see Figure 8.11 for the test conditions) Figure 8.22: the flow stress curves under different strain rates (1 to 10-4 ) (Lademo et al., 2012) Figure 8.23: flow stress curves under different temperatures (25 o C to 280 o C) (Abedrabbo et al., 2007) Figure 8.24: Cof vs sliding speed (Fenske et al, 2008) xxiii

25 Chapter 1. Introduction Due to regulations to improve fuel efficiency and reduce greenhouse gas emissions, the amount of aluminum used in vehicles is increasing. However, stamping of Al alloys presents new challenges in obtaining good part definition (corner and fillet radii) and formability (fracture and resulting scrap), due to the higher cost (raw material cost: Steel $700/t, Al $2900/t) and the lower formability of the material. The tensile elongation of 5xxx at room temperature is usually lower than 30% (TheAluminumAutomotiveManual, 2002); while the elongation of aluminum-killed steels is around 50% (Wagoner and Laukonis, 1983). Therefore, stamping of Al alloys presents new challenges in obtaining good part definition (corner and fillet radii) and formability (fracture and resulting scrap). FE simulations are being used increasingly over the past decade in the field of metal forming, to reduce design errors and reduce the cost of tooling and experimentation. However, FE simulations require the use of accurate information about the manufacturing process and the ability to infer the simulation results. The following are some of the important aspects of a metal forming simulation: 1) Accurate material characterization and 2) Accurate Coefficient of Friction in tool interface. The uniaxial tensile properties of Al alloys reported in the literature are not very accurate predictors of the response of Al alloys as they are exposed to predominantly biaxial tensile loading conditions in sheet metal forming operations. Therefore, biaxial tests, such as the bulge test and dome test, should be used to improve the accuracy of FE simulations. 1

26 The right lubricant can reduce or even avoid failures associated with wrinkling and premature, thus increasing the process window (Meiler and Jaschke, 2005).Thus, to improve the stamping operations, it is necessary to determine and use the best combination of lubricant and sheet surface texture to improve the lubrication condition between blank and dies. In order to estimate the coefficient of friction (CoF) used in Aluminum stamping simulations, various tribo-tests (such as strip draw test, draw-bead test and cup draw test) were studied in order to emulate the near-production condition. By using the FE inverse analysis, the CoF between the workpiece and tools will be determined. The freedom to use any velocity profiles used to be restricted only to FE simulations, till recently- until the advent of servo presses. It is known that the formability of aluminum alloys can be improved with the increase in forming rate. The acceleration rate dependency of material properties and the change of frictional characteristic may also effect the aluminum forming performance. However, the detailed investigation has not done yet. We will look into the effect of forming rate on the formality of aluminum part by changing the forming speed, by using a servo press. A servo-drive press is relatively new type of sheet metal stamping machine using an electric servo-motor to drive and actuate the press. Servodrive presses enable much greater control of press velocity vs. stroke compared to traditional mechanical and hydraulic presses (Hayashi and Nishimura, 2009).As a result, servo-presses can be tuned to optimize a given stamping process for the sheet material being formed, and the specific forming operation being performed (shape forming or drawing). 2

27 Chapter 2. Research Objectives and Approach The overall objective of this study is to improve the drawability of aluminum sheet (Al 5182-O) by using the optimum slide velocity and forming conditions (lubrication, blank holder pressure, die design) in a servo press. The specific objectives are: Material property: determine material properties (flow stress, yield stress, ultimate tensile stress, elongation, r values) for selected Al alloys. (also input parameters for FE simulations) Lubrication condition: identify the best lubricants and surface conditions (texture) for reducing friction. Determine the values of coefficient of friction (COF) for inputs to FE simulations. Design guideline: identify guidelines for designing and forming stampings for Al alloys, i.e. minimum corner radii, maximum thinning and max draw depth. Ram velocity (servo-press): optimize the ram velocity (velocity-stroke curves) and blank holder force in stamping Al alloy sheets in a servo press. In order to achieve the objectives the following approach is followed: Phase I: Characterization of Sheet Material Properties. During this phase, the material properties (flow stress, yield stress, ultimate tensile stress, elongation, r values) for selected Al alloys will be determined. Meanwhile, the flow stress data will also be obtained from bulge test and will be used for FE simulations, to improve the accuracy of FE prediction. 3

28 Phase II: Experimental evaluation of selected lubrication conditions to improve stamping quality. During this phase, the tribotest will be determined to identify the best lubricants and surface conditions (texture) to reduce friction between work piece and tools. The method of estimating the values of coefficient of friction (COF) will be studied for inputs to FE simulations. Phase III: Design a die set for use in the servo press. In this phase, FE simulations will be conducted to determine the best design for this die set, such as the optimum die/punch radii, the clearance between punch and die, and the maximum draw depth. The effect of CoF on the FE simulation results will also be investigated Phase IV: Conduct simulations and tryouts on the servo press using the die geometry of Phase III. In this phase, FE simulations of the tooling (designed in Phase III), will be conducted to optimize the slide motion for the best forming in a servo-press. From this Task, the output should be a know-how of how to optimize the slide motion for best aluminum forming by using servo press. It should be noted that since the die (Honda Die) we helped to design is not available to run the experiments in this July. Hence, we used a Shiloh die, which was built by Shiloh Inc. for Advance d High Strength Steel (AHSS) s tryout using a servo press, to run some preliminary aluminum tryouts. Phase V: Develop design guidelines using results from Phase IV. 4

29 Chapter 3. Literature Review on Aluminum forming 3.1 Material properties and formability Wrought alloys designations Wrought aluminum is identified with a four-digit number (see Figure 3.1). The first digit defines the principle alloying element. Wrought alloys have seven different series. 5xxx, 6xxx and 7xxx aluminum alloys are used most commonly for automotive applications. The specific banana diagram (elongation vs specific tensile strength) in Figure 3.2 shows that aluminum alloys have higher strength-to-weight ratio than steel. This characteristic indicates that the aluminum alloys could be used in some applications where weight saving are worth the higher cost. The mechanical properties of some 5xxx and 6xxx alloys are listed in Table 3.1, and their specific properties and main differences are shown in Figure

30 Figure 3.1: Wrought aluminum alloy designation system (TheAluminumAutomotiveManual, 2002) Figure 3.2: The specific Banana diagram (DuckerWorldwide, 2011) 6

31 Figure 3.3: Al 5xxx and 6xxx alloys used in car body components (Hirsch, 2004) Table 3.1: Mechanical properties of 5000 series aluminum sheet alloys (SAKURAI, 2008) Mechanical properties alloy UTS YS EL n- R- (MPa/ksi) (MPa/ksi) (%) value value AA /40 135/ series AA /41 135/ AA /38 125/ AA /27 90/ AA /31 90/

32 5xxx alloys are non-heat treatable alloys with ultimate tensile strength of 18 to 51 ksi (125 to 350 MPa). The amount of magnesium in 5xxx alloys is in the range of % and these alloys have the highest strength among the non-heat treatable alloys. 5xxx alloys are exceptionally formable but are susceptible to the formation of Luder s bands during forming. Therefore, these alloys are used for inner panels. Al 5182 and 5754 are the principle 5xxxx series alloys for auto body panels. Al 5754 is usually recommended for relatively high temperature applications. Jaguar uses Al 5182 and 5754 for hood, decklid inner and fender reinforcements in Jaguar XJ, as shown in Figure 3.4. Figure 3.4: 5xxx series alloys used for hood, decklid inner and fender reinforcements (Lidguard, 2009) 8

33 6xxx series alloys are heat treatable with ultimate tensile strength of ksi (125 to 400 MPa). The amount of magnesium and silicon in the 6xxx alloys is around 1.0% is the basis for developments in Europe; 6111 is the basis for development in USA. 6xxx series alloys are free of Luder s band and these alloys are mostly used for outer panels, where the formability is less important. Al 6022 and 6111 are the principle 6xxx series alloys for body panel applications. Figure 3.5 shows one example using Al 6111-T4 for hood and fender skins in Jaguar XJ. Figure 3.5: 6xxx series alloys (6111-T4) used for hood and fender skins in Jaguar (Lidguard, 2009) Aluminum 5182-O material The use of annealed Al 5182 alloy is of special interest to automotive users to make automotive panels because the material represents a cost-effective solution for reducing 9

34 weight while maintaining functional requirements for structural strength and crash resistance. It has the good balance between ductility and formability. This annealed aluminum sheet goes through the fully annealing process (fully recrystallized, fully soft O-temper) to reach the final gauge, by removing the cold work stored in the material. It does not have further treatment after processing ( 805). The composition of Al 5182-O mainly includes Aluminum (~95%) and Magnesium (~4.5%) ( The advantages of alloying with Mg are 1) provides solid solution strengthen and 2) reduces density. There are two types of stretcher strains occurred during deformation when al 5182-O alloy is used, see Figure 3.6. Type A is observed when the deformation is non-homogeneous at small plastic region; while type B is observed at higher degrees of deformation in the stressstrain curve. Due to these two types of stretch strains, the uses of Al 5182 alloy are limited to the inner panels, brackets and supports which are not visible in the final structure. 10

35 Figure 3.6: Stretcher strains occurred in Al 5182-O (K.Siegert and S.Wagner, 1994) 3.2 Determination of materials properties by the uniaxial tensile test The standard uniaxial tensile test is commonly used to determine the mechanical properties of the sheet materials. The procedure of determining material properties by uniaxial tensile test is described in (Altan and Tekkaya, 2012) Limitations of the tensile test Uniaxial state of stress The tensile test only provides ductility and work hardening under uniaxial tension conditions. This deformation condition does not represent the material behavior in unequal biaxial stretching that usually occurs in practical sheet metal forming operations. Therefore, it is necessary to test the material under biaxial deformation conditions using the biaxial bulge test. Limited strain range 11

36 The flow stress curve obtained from the tensile test is limited to the point at which local necking is observed. This maximum strain at local necking is small compared to the strains encountered in actual stamping operations. Sheet material properties (flow stress data) for the parts undergoing large deformations (large strains) are critical to the accuracy of predictions. Extrapolation of flow stress data from the tensile test to large strains is an approximation and not necessarily accurate. In the biaxial bulge test, the maximum effective strain achievable without local necking is much larger (usually twice) than that in the tensile test. Therefore flow stress curves obtained from the biaxial bulge test are over a larger strain range, compared to tensile test. 3.3 Determination of material properties by the viscous pressure bulge (VPB) test In the bulge test, the sheet metal clamped at its edges is stretched against circular die using viscous medium, as shown in Error! Reference source not found.. The sheet metal bulges into a hemispherical dome and eventually it bursts. The tooling at the CPF/EWI is designed for a single action hydraulic press that uses a punch to pressurize the viscous medium (see Figure 3.8). The upper die is connected to the slide of the press and the lower die is connected to the cushion of the press. The punch in the lower die is fixed to the press table. At the beginning, the tooling is open and the blank sheet is placed between the upper and the lower dies. Then the dies closed to clamp the blank material and the slides moves down together with the entire die set. 12

37 Consequently, the viscous medium is pressurized by the stationary punch and the sheet is bulged by the viscous medium flowing into the upper die as shown in Figure 3.8. With the tooling at CPF/EWI, it is possible to measure the dome height and the pressure of the viscous medium during the experiment. The procedure of estimation of the flow stress curve using bulge test is described in (Venkatesan, 2010). Viscous medium Figure 3.7: Geometry before and after the test (t0 = original sheet thickness, td = sheet thickness at the apex, hd = bulge height, P = hydraulic pressure, rc = die fillet radius, dc = die cavity diameter, rd=radius of curvature) (Venkatesan, 2010). Figure 3.8: Schematic on the left shows the sheet clamped between the upper and lower die. Schematic on the right shows the sheet bulged by the pressurizing medium into the upper die cavity (Venkatesan, 2010). 13

38 3.4 Conditions at tool/material interface In aluminum forming, lubrication is critical because the surface of aluminum sheet is smoother than that of steel. Better lubrication can increase the process window, see Figure 3.9. Besides lubricants, surface finish is also important in forming of aluminum sheets. Figure 3.9: Increasing process window by reducing friction (Meiler and Jaschke, 2005) Lubricants There are mainly two types of lubricants: liquid lubricants and dry film lubricants. Liquid lubricants are usually applied in the press shop, while dry film lubricants are applied in the rolling mill (Meiler et al., 2003). The different production processes used with liquid based and dry film lubricants are shown in Figure It can be seen that the dry film lubricant process does not need the panel washing process. Thus, it reduces the cost of product. Recent studies have shown that dry-film lubricants could provide better lubrication conditions compared to oil based lubricants. In October 2003, BMW group switched to 14

39 water free dry film lubricants in forming aluminum front fender and hood assembly parts in the BMW 7 series (Meiler et al., 2004). Dry Film Lubricant is a wax component suspended in an aqueous solvent (Erickson and Folkins, 1993). It reduces galling, seizing and fretting and improve corrosion resistance. The dry film lubricants are divided into two categories: water-soluble dry film lubricants and water-free dry film lubricants (Meiler and Jaschke, 2005). 1. Water-soluble dry film lubricants are applied in amounts from 0.5 to about 1.5 grams per square meter at the rolling mill (see Figure 3.10). They stick to the panel's surface and offer sufficient corrosion protection, but they are not compatible with most adhesives used in automotive body construction. The typical resulting coefficient of friction is about 0.04 to 0.15 (TheAluminumAssociation, 2006). 2. Water-free dry film lubricants (or hotmelts) also are applied to the sheet material in small amounts at the rolling mill. They provide good drawing performance and are compatible with almost all commonly used adhesives. The coefficient of friction can be 0.02 to 0.05 (TheAluminumAssociation, 2006). Table 3.2: Two lubricants, KTLN 16 and Drylube E1 (Wagner et al., 2002) KTLN 16 Mineral oil-based lubricant, viscosity 160 mm2/s@40 o C, application in 1.0 g/m2 at the press shop, good corrosion protection Drylube E1 Water-free dry film lubricant (hotmelts), viscosity 100 mm2/s@40 o C, application in 1.0 g/m2 at the rolling mill, excellent corrosion protection 15

40 Figure 3.10: Oil base and dry film lubricant applications (Meiler and Jaschke, 2005) (Wagner et al., 2002) evaluated two lubricants: KTLN 16 (wet lubricant) and Drylube E1 (dry film lubricant), by using the deep drawing test with Al A 500 x 400 mm sheet blank was drawn up to the maximum drawing depth of 100 mm, at a drawing velocity of 64 mm per second. Figure 3.11 shows the maximum possible drawing depth and the maximum applicable blank holder forces for these two lubricants listed in Table 3.2. It can be seem that the dry film lubricant, E1, gives a larger process window, compared to KTLN 16. The suppliers of various lubricants, cited in this report, are given in Appendix. 16

41 Figure 3.11: The maximum possible drawing depth and maximum applicable blank holder forces for the two lubricants (Wagner et al., 2002) (Meiler and Jaschke, 2005) studied the effect of four different lubricants (Drylube E1, Houghton DF 25 B, Drylube C2 and KTLN 16) on the drawability (in terms of blank holder force - BHF). The types of lubricants used for this study are shown in Table 3.3. The alloy is AlMg3.5Mn with an Electrical Discharged Texture (EDT) of a roughness (Ra) of 0.8 µm. The diameters of the blank sheet and punch are 100 mm and 50 mm respectively. Figure 3.12 shows the blank holder forces with different types of lubricants and different amounts of lubricants applied on the sheet. It can be seen that for 1.5 g/m 2 application the best performance is water-based product Drylube C2, followed by Drylube E1 and Houghton DF 25 B. Also, it can be seen from Figure 3.12 that the dry film lubricants (Drylube E1 and Drylube C2) can be applied on the sheet samples with less amount of 17

42 lubricant compared to KTLN 16. The supplies of various lubricants, cited in this report, are given in Appendix. Table 3.3: Different types of lubricants used for aluminum sheet metal forming (Meiler and Jaschke, 2005) 18

43 Figure 3.12: Drawing performance of different lubricants. The larger the blank holder force used in deep drawing cups without fracture, the better is the lubricant (Meiler and Jaschke, 2005) Surface finish of aluminum sheets In addition to lubricants, surface finish of the sheet also affects the friction behavior during sheet metal forming. Texture transfer occurs during the last stage of cold rolling, which is used to flatten the sheet and impart the final surface texture. Aluminum sheet is commercially available in three surface finishes listed as below: Mill Finish (MF) on sheet finish rolled with ground rolls used for interior parts Electric Discharge Texture (EDT) - on sheet finish rolled with EDT rolls used for skin panels 19

44 Dull Finish (DF) on sheet finish rolled with shot blasted rolls EDT surface texturing is commonly used in Europe for aluminum sheets. It can improve paint appearance and formability due to the presence of lubricant pockets and isotropic surface (see Figure 3.13). The EDT and DF samples have higher roughness values (Ra) than MF samples (see Table 3.4). MF is the standard surface used in North America for aluminum sheets. Besides these three surface methods, there are some other surface texturing for aluminum sheets, such as Precision texturing (PRETEX) and laser Texturing (LT) (Groche and Callies, 2005). Table 3.4: Ra values of EDT, DF and MF samples (Hartfield-Wunsch et al., 2011) Surface finish Ra value in transverse rolling direction Ra value in longitudinal rolling direction EDT ~0.75 ~0.75 DF ~0.52 ~0.51 MF ~0.25 ~

45 Figure 3.13: SEM image of an Electrical Discharge Texture surface (left) and mill finish surface (right) on AA 6016 sheet (TheAluminumAutomotiveManual, 2002) 3.5 Tests for evaluation of lubricants Various tribological tests are used to evaluate the performance of lubricants for aluminum under laboratory conditions. Such as cup drawing test (CDT) and strip draw test: Cup drawing test can emulate the production condition such as contact pressure and speed. Other tribotests (such as strip draw test) are also available to emulate the performance of lubricants Cup drawing test (CDT) In the cup drawing test, a hollow is formed by forcing a flat, circular blank into a die using a punch (see Figure 3.14) (Kim, 2008). The initial blank is constrained by the blank holder, while central portion of the sheet is pushed into a die opening with a punch, to draw the metal into the desired shape without causing wrinkles or splits in the drawn part. Figure 3.15 shows the schematic of CDT tooling at CPF/EWI. 21

46 Figure 3.14: Schematic of cup drawing (Kim, 2008) Figure 3.15: Schematic of CDT Tooling at CPF/EWI (Subramonian et al., 2011) 22

47 Performance evaluation criteria for cup drawing test include: i. Maximum applicable blank holder force (BHF) without failure of the cup (The higher BHF that is applied without fracture in the drawn cup, the better the lubrication condition) ii. The measurement of draw-in length, Ld, and perimeter of flange in a drawn cup (The larger the draw-in length or the smaller the perimeter, the better the lubrication condition) Using FE and inverse analysis, it is also possible to determine the Coefficient of Friction (COF) for input into FE simulation Other tests available for evaluation performance of lubricants Various tribotests can be used to evaluate the lubricity of lubricants in the laboratory. These include the strip draw test (SDT) and draw bead test (Andreasen et al., 1997) & (Dalton and Schey, 1992). Limit dome height test (LDH) can also be used to evaluate lubricants for stamping (Schey, 1984). However, the above mentioned tests have limitations in emulating the process conditions in terms of temperature, contact pressure and speed (Altan and Tekkaya, 2012). Studies have shown that the cup draw tests (CDT) can emulate the near-production conditions well (Kim et al., 2007) & (Subramonian et al., 2011). Therefore, CDT was selected to evaluate the stamping lubricants for this study. 3.6 Design rules for tooling (deep drawing and stretch forming) Many die design guidelines can be applied to achieve good part quality in aluminum forming (Kazanowski, 2006) & (Altan, 2001). i. Deep drawing: Figure 3.16 shows the schematic of a simple cup deep drawing operation. a) Die design 23

48 Tooling for steel deep drawing can be used for aluminum deep drawing. However, there should be some modifications on the tooling. Figure 3.16: Schematic of deep drawing (Altan, 2001) b) Die clearance Die clearance 'ud' is the distance between the punch surface and the die surface. In deep drawing, if the die clearance is too large, the component will be drawn as a cone and not a cylinder. If it is too small, then ironing will occur. Recommended clearance for deep drawing of aluminum round cups is: Where: s0 = sheet thickness c) Die corner radius ud = s (10s0) (1/2) The die corner radius rd depends on the thickness of the workpiece. 24

49 Recommended values of die radii for aluminum stampings are about 4 to 8 times of the sheet thickness. If the die radius is too large, it may lead to wrinkling. If it is too less, the probability of fracture may increase. d) Punch corner radius The recommended punch radii values for aluminum stampings are in the range of 8 to 10 times of the sheet thickness. Punch radii greater than 10 times of the sheet thickness may cause wrinkling, because of high compressive hoop stresses that occur while the sheet wraps around the punch radius. e) Surface finish on tools for deep drawing Draw dies and punches should have a surface finish of 0.4 μm (16 μin.) or less for most applications. 3.7 Simulation/deformation zone Some critical parameters that are different for aluminum and that affect formability are: Elastic modulus: E-modulus of aluminum is about one third that of steel. Thus, more springback is observed in the part. By increasing the blank holder force towards the end of the deformation stroke, the amount of stretching is slightly increased, while the sheet thickness and the springback can be reduced. To control springback, the forming operation must be optimized to ensure at least 2 percent stretch throughout the part, as seen in Figure

50 Figure 3.17: Strains of at least 2 percent are necessary when stamping aluminum to reduce springback (Thomas and Altan, 1999) Friction: Friction between tool and aluminum sheet is expected to be higher compared to steel sheet because the surface roughness (Ra) of aluminum ranges from 0.25 to 0.38 micron. In comparison, the Ra of steel sheet is around 0.63 to Thus, the smoother texture of aluminum tends not to Trap the lubricants as well as steel. Therefore, dry wax-like lubricants are often used. Yield criteria: Aluminum has low anisotropy values (the value is almost half of that of steel). Consequently, yield criteria that consider the effect of anisotropy, such as Barlat p, Barlat d or Hill 90, should be used. Among these three criteria, Barlat p can predict the most accurate simulation results for aluminum forming. However it requires 18 parameters and the computation time is very long. Barlat d is 26

51 recommended to be used for aluminum alloy sheets, since it only needs 8 input parameters and the accuracy of Barlat d is close to Barlat p. Three uniaxial tension tests at 0, 45 and 90 o from rolling direction, one balanced biaxial bulge test and one disk compression test are needed, in order to calculate the Barlat d function coefficients (Barlat et al., 2003). 3.8 Servo press application on aluminum forming Aluminum has lower formability compare to that of steel. Therefore, the application of aluminum sheets is limited to relative simple shapes, like hood panel and trunk lid panels. It is known that the formability of aluminum alloys can be improved with the increase in forming rate. Figure 3.18 shows the modified inner door panels formed by using a servopress at different slide motion patterns. It shows that the part can be formed with a large fracture with an average velocity of 41 mm/s, and can be formed without any fracture and necking with an average velocity of 103 mm/s (see Figure 3.19). (a) (b) Figure 3.18: Panel formed at (a) 41 mm/s and (b) 103 mm/s [Hayashi et al, 2009] 27

52 Figure 3.19: Different slide motions used to form aluminum panels [Hayashi et al, 2009] The change of frictional characteristic when using high forming speed may also effect the aluminum forming performance. However, the detailed investigation has not done yet. We will look into the effect of forming rate on the formality of aluminum part by changing the forming speed, by using a servo press. A servo-drive press is relatively new type of sheet metal stamping machine using an electric servo-motor to drive and actuate the press. Servo-drive presses enable much greater control of press velocity vs. stroke compared to traditional hydraulic presses (Hayashi and Nishimura, 2009), as shown in Figure As a result, servo-presses can be tuned to optimize a given stamping process for the sheet material being formed, and the specific forming operation being performed (shape forming or drawing, for example), as shown in Figure

53 Figure 3.20: Examples of different slide motions of servo press (Groseclose, 2008) 29

54 Figure 3.21: The flexibility of slide motion in servo presses (Miyoshi, 2004) In deep drawing process, a low blank holder force will cause wrinkling while a large blank holder force will cause cracking. Thus, it is important to find an appropriate blank holder force to form the designed part. Usually, there are four types of BHF formability windows as shown in Figure For the type a, we may be able to form the part to the designed draw depth using a constant blank holder force (see red dotted line in Figure 3.22 (a)); while for types c and d, we may need use the variable blank holder force (VBHF) approach (see red dotted line in Figure 3.22 (b) and (d)) to form the part to a certain draw depth, which means that we may need change the blank holder force through stroke for form the part. In this study, we will investigate the effect of variable blank holder force (VBHF) on the aluminum part s formability by using AIDA s 300 tons Servo press. 30

55 Figure 3.22: Four basic BHF formability windows (Kitayama et al., 2010) 31

56 Chapter 4. Determination of Material Properties of Al 5182-O using Uniaxial Tensile Test, Viscous Pressure Bulge Test (VPB tests) and Dome Test In this chapter, the elastic and plastic sheet material properties were obtained by viscous pressure bulge (VPB) test and tensile test, which were conducted by CPF/OSU and Honda R&D, respectively. These material properties are essential inputs to finite element analysis that will be used to simulate aluminum forming processes. 4.1 Determination of sheet material properties by the uniaxial tensile test The material properties (flow stress curve, Lankford values, Yield stress) were obtained from the uniaxial tensile test at Honda R&D. The dimensions of tensile test specimen are given in Figure 4.1. The results of tensile test provided by Honda R&D are shown in Figure 4.2. Two specimens were tested in each direction (rolling, diagonal and transverse). The true stress-true strain curves in these three directions are shown in Figure 4.2. It can be observed that there is not much difference on the flow stress curves in rolling, diagonal and transverse directions (see Figure 4.2). 32

57 Figure 4.1: Dimensions of tensile test specimen, provided by Honda R&D Table 4.1: Tensile test data provided by Honda R&D Sample Maximum Tensile stress Tensile stress at Yield (Offset 0.2 %) Young's Plastic Strain Ratio at r- Value (Strain 15 % ) Thickness (MPa) (MPa) (%) mm AL 5182 RD AL 5182 RD AL 5182 DD AL 5182 DD AL 5182 TD AL 5182 TD

58 Figure 4.2: True stress-true strain curves in rolling, diagonal and transverse directions 4.2 Determination of sheet material properties by the VPB test Biaxial bulge tests at EWI/CPF were conducted to determine the flow stress. The test parameters used are shown in Table 4.2. It should be noted that we tested the Al 5182-O with 1 mm thickness instead of 1.5 mm thickness, since 1.5 mm sheets were not available when we were running the bulge test. The thicknesses and the anisotropy values, R0 and R90, considered in calculation of flow stress curves are given in Table 4.1. The detailed description of the procedure for flow stress determination with hydraulic bulge test, can be found in (Gutscher et al., 2004). The bulge test result is shown in Figure 4.3. It should be noted that all the simulations we have conducted in this study, the material properties provided by Honda R&D and CPF were used, except that the forming limit 34

59 curve (FLC) was obtained from (Carsley et al., 2013), see Figure 4.4, since we don t have our own FLC yet. The simulation results will be updated once we have our own FLC. Table 4.2: Parameters used in the viscous bulge test Ram speed Clamping force Radius of fillet of cavity Size of the test sample Sheet thickness 5 mm/sec 500 kn 6.4 mm 254mm x 254mm 1.0mm Figure 4.3: Bulge test data provided by EWI/CPF (Mao, 2013) 35

60 Major strain Minor strain Figure 4.4: Forming limit curve of Al 5182-O, with thickness of 1.1 mm, obtained from (Carsley et al., 2013) 4.3 Comparison of flow stress data obtained from tensile test and bulge test The flow stress data obtained from tensile and bulge tests were compared in Figure 4.5. It can be observed that the maximum strain obtained in uniaxial tensile test is around 0.2. However, the maximum strain obtained in biaxial bulge test is around The K and n values obtained from tensile and bulge tests by curve fit are listed in Table 4.3. The difference on simulation results by using different K and n values will be investigated and explained in the future study. Table 4.3: K and n values obtained from tensile and bulge tests K n Tensile test Bulge test Since biaxial test can provide larger maximum strain, the flow stress data obtained from bulge test was used in our FE simulations. 36

61 True stress (MPa) It should be noted that the thickness of blank used for tensile test is 1.5 mm; while the thickness of the blank used for bulge test is 1 mm y = x Tensile test, 1.5 mm Bulge test, 1.0 mm y = x Al 5182-O True strain Figure 4.5: Comparison on flow stress curve between tensile and bulge tests 37

62 Chapter Objectives Experimental Evaluation of Selected Lubrication Conditions to Improve Stamping Quality The overall objective of this lubrication study is evaluate the lubricity of the selected lubricants for forming Al 5182-O sheets, provided with two different surface texturing (Mill Finish and Electro Discharge Texturing). The specific objectives are to: Select the lubricants that perform best in stamping of aluminum sheet (Al 5182-O) Compare the friction behavior between Mill Finish (MF) and Electro Discharge Texturing (EDT) sheets. Determine the coefficient of friction (CoF) at the tool-work piece interface under various lubrication conditions by comparing experimental results and FE simulation results. Thus, the value of the CoF can be used as input into FE simulation of stamping with complex geometries used in automotive production, such as hoods, decklids and doors. 5.2 Approach In order to achieve the objectives listed above, a step by step procedure is followed. First, all the sheet materials and lubricants are procured. Cleanability tests are conducted by a lubricant company as preliminary evaluations of the lubricants. Tensile and Viscous Pressure Bulge (VPB) tests are conducted on the sheet material in order to obtain the material properties, used as input to FE simulations. Preliminary FE simulations are conducted to determine the experimental parameters for Cup Draw Test (CDT). The different lubrication conditions are evaluated using CDT to determine the best lubricants of the ones evaluated in this study. The preliminary FE simulations of the CDT are used to 38

63 determine the experimental parameters for the tests and also to determine the coefficient of friction at the tool-workpiece interface. 5.3 Preparation of Experimental Evaluation for Various Lubrication Conditions Before starting cup drawing tests the following tasks are completed Procurement of Sheet Materials and Lubricant Samples 14 lubricants (see Table 5.1) are procured for this study. The 5182-O aluminum sheets were available with two different surface texturing (MF and EDT). In this study, all the lubricants are used as received, except lubricant I, which had to be diluted with water in the ratio of 1:3 (lubricant/water). The coating weight applied on the blank was 1 gram/m 2 for all lubricants. Draw down bar #0 is used to apply the lubricant on the blank Preliminary Evaluation of Lubrication Conditions The sheet samples, as supplied were coated with mill oil. Therefore, the lubrication conditions tested were mill oil plus lubricants, listed in Table 5.1. The cleanability of all the lubricants was evaluated by Henkel. The evaluation shows that all of lubricants, evaluated in this study, passed the cleanability test. 39

64 Table 5.1: Lubricants and their details as provided by lube companies Lubricant code A B C D E F G H I J K L M N Lubricant type Petroleum oil Petroleum oil Petroleum oil, Additive blend Petroleum oil, Additive blend Emulsified Oil (semi synthetic) Dry film Mineral oil based Mineral oil based Mineral oil based Water based Water based Water based Dry film Dry film Application method draw down bar #0 draw down bar #0 draw down bar #0 draw down bar #0 draw down bar #0 draw down bar #0 Applied by Lub company draw down bar #0 draw down bar #0 draw down bar #0 draw down bar #0 draw down bar #0 Applied by lub company Applied by lub company 40

65 5.4 Lubrication Tests, Analysis and Results All the lubricants in this study were evaluated using the cup draw test (CDT) Principles of CDT-Cup Draw Test The principle of cup draw test was described in section Preliminary Finite Element Simulation of the Cup Draw Test (CDT) The objective of performing preliminary finite element simulations of CDT is to obtain a workable range for the Blank Holder Force (BHF) for the test. The commercial software PAM-STAMP was used for the preliminary simulations. Forming Limit Curve (FLC) obtained from (Carsley et al., 2013), shown in Figure 4.4, was used as the criterion to determine the working range of BHF, i.e. to estimate the values of BHF where fracture, as indicated by the FLD, would not occur. The FE model and the parameters used in the preliminary FE simulations are shown in Figure 5.1 and Table 5.2, respectively. A quarter model is used because of symmetry (see Figure 5.1). The outer surfaces of the punch, die and blank holder are meshed using shell elements. Based on the FE simulation results, it is concluded that tons BHF should be a good working range for conducting the CDT experiments. 41

66 Table 5.2: FE Simulation Parameters for obtaining BHF range for CDT Parameter Sheet material Description AL5182-O, 1.5 mm (Provided by Honda Engineering) Anisotropic values R0=0.71 R45=1.09 R90=0.84, provided by Honda R&D ( Table 4.1) Flow stress curve Obtained from the spherical bulge test at EWI/CPF (see Figure 4.3) Blank Size Blank Holder Force Coefficient of friction (between blank mm diameter 16 tons, 18 tons, 20 tons, 22 tons, 24 tons 0.1,0.11,0.12 holder and sheet, die and sheet) (assumed) Punch (or die ) stroke 80 mm 42

67 Figure 5.1: Quarter model used in simulation of Cup Drawing Test using PAM-STAMP Experimental Set-up Description of the tooling and test procedure The cup drawing tooling used in this study, operates with a 160-ton hydraulic press that has a maximum ram speed of 300 mm/sec (12 inch/sec). The schematic of the tooling is shown in Figure 3.15, in section A draw die attached to the upper ram moves down and forms a cup sample with a stationary punch. The preset constant blank holder force is provided by the CNC-controlled hydraulic cushion pins. During the test, the punch force is measured by a load cell located at the bottom of punch and the displacement of die is recorded by an infrared laser sensor. The draw ratio (blank diameter / punch diameter) was selected to be 2.0. The drawing depth was selected to be 80 mm (3.15 inch) to leave some flange area for measuring the draw-in length and the perimeter of the flange. Round blanks of mm (=12 inch) diameter and 43

68 1.5 mm (=0.059 inch) thickness were cut from Al 5182-O sheet. The cup draw test procedure is described in detail in elsewhere (Subramonian et al., 2011) Lubricant Selection and Application Method Application of lubricant on the blank is done by using a pipette and draw down bar #0. The lubricant was spread all over the blank surface on both sides uniformly. Three drops of the lubricant were applied on each side of blank to reach the required coating weight 1.0g/m 2 (+/- 0.3 g/m 2 ). The coating weight/amount was verified by measuring a few samples using the lab balance before and after lube application. Additionally, the applied lubricant thickness was verified by using a portable tool (Microderm CMS, see Figure 5.2) provided by Honda. The measurement results obtained by using Microderm CMS tool are shown in Table 5.3. Figure 5.2: Microderm CMS tool used to measure the thickness of applied lubricants coating (UPA, 2013) 44

69 5.4.4 Cup Draw Tests The blank holder forces for the CDT are selected based on the FE simulation results conducted in section The 16 tons was set as the minimum blank holder force (BHF). 17 tons was found to be the maximum blank holder force (BHF) from experiments, since all the cups drawn with wet lubricants cracked when BHF was 18 tons. The test parameters for the CDT are given in Table 5.4. Table 5.3: Measurement results using Microderm CMS Lubricant Coating weight Mill oil +Lube A 2.4 g/m 2 Mill oil 1.16 g/m 2 The coating weight measured by using Microderm CMS tool is around 1.24 g/m 2, which is close to the required coating weight (1 g/m 2 ). 45

70 Table 5.4: Parameters of Cup Drawing Tests (CDT) Parameter Sheet Material Blank Size Blank Holder Force Ram Speed Punch Stroke Number of Samples (for each test Description AL5182-O (1.5mm thickness) mm diameter 16 tons & 17 tons 40 mm/sec 80 mm 5 condition) Tested Lubricants All lubricants listed in Table Experimental Results The performance of the lubricants is evaluated by measuring the perimeter of the flange of the drawn cup. A lubrication condition is said to have failed at a certain BHF if three consecutive cups cracked. The test procedure for the evaluation of all lubricants is shown in Figure 5.3. At 16 tons BHF, all lubricants listed in Table 5.1 were tested. The blank holder force was increased from 16 tons to 17 tons, in order to select the better lubricants. Lubricants that cracked at 16 tons were not considered for 17 tons test. The results of 17 tons are used in the final evaluation, since all lubricants cracked at 18 tons. 46

71 Cups were deep drawn with all lubricants except four (B, D, H and N). The variations in perimeter are shown in Figure 5.4. Lubricants F in the graph could be categorized as best performing lubricant when blank holder force (BHF) is16 tons. Figure 5.3: Test Procedure for the evaluation of lubricants At 17 tons BHF, cups with all lubricants cracked, except with lubricants F, G, K, L and M. The results at 17 tons BHF are shown in Figure 5.5. It can be concluded that G, K and L lubricants could be categorized as better performing lubricants, while the lubricant F performed best at 16 and 17 tons BHFs. Thus, lubricant F can be considered to be the best performing lubricant. 47

72 There is a good correlation between the results obtained at 16 tons BHF (see Figure 5.4) and 17 tons (see Figure 5.5). F, K, L and M performed well at 16 tons BHF also performed well at 17 tons BHF. Figure 5.4: Flange perimeter recorded for 13 lubricant tests at 16 tons BHF (L2, L4 and L8 failed at 16 ton BHF ) Note : The y-axis on the graph does not start from 0. The error bands show the deviation between samples. 48

73 Figure 5.5: Flange perimeter recorded for 5 lubricant tests at 17 tons BHF (L1, L3, L5, L10 L13 failed at 17 ton BHF ) Note : The y-axis on the graph does not start from 0. The error bands show the deviation between samples. The friction behavior of Al 5182-O with EDT surface texturing is also compared to that of Al 5182-O with MF surface texturing. Lubricants K, J and L were tested at 16 tons BHF and Lubricant J cracked when BHF was increased from 16 tons to 17 tons. The comparison between MF surface texturing and EDT texturing is shown in Figure 5.6 and Figure 5.7. It can be concluded that EDT performs better than MF, when wet lubricants are applied. 49

74 Figure 5.6: Comparison between MF and EDT at 16 tons BHF Figure 5.7: Comparison between MF and EDT at 17 tons BHF 50

75 5.4.6 FE simulations to predict Coefficient of friction (CoF) at tool-work piece interface The purpose of this simulation is to estimate the values of friction coefficient for the tested lubricants. The perimeter of the flange obtained at the end of stroke in the simulation is compared to the perimeter of the flange obtained in the experiment and used to determine the coefficient of friction (CoF) from the simulation. PAMSTAMP was used to predict the CoF at the tool-blank interface, the parameters input to FE simulation are the same as parameters listed in Table 5.2, except that 1) BHFs used in the simulation are 16 tons and 17 tons and 2) the coefficient of friction (CoF) are assumed to be 0.06, 0.08, 0.1 and FEA results are shown in Table 5.5. From Figure 5.8 and Figure 5.9, it can be concluded that i) the CoFs of lubricants G and K is in the range of , ii) the CoF of lubricant F is around 0.08 and iii) the CoF of the lubricant I is around

76 Table 5.5: Prediction of Flange Perimeters by FE simulations. Blank Holder Force (Tons) Coefficient of Friction (CoF) Flange perimeter (mm)

77 Figure 5.8: Comparison of flange perimeters obtained from simulation and experiment to predict the coefficient of friction at 16 ton BHF for AL5182O/MF with 1.5mm thickness 53

78 Figure 5.9: Comparison of flange perimeters obtained from simulation and experiment to predict the coefficient of friction at 17 ton BHF for AL5182O/MF with 1.5mm thickness Comparison of cup thickness variation obtained from FE simulations and experiments In order to verify the simulation results (coefficient of friction-cof), the thickness distribution in the drawn cup obtained from experiments and simulations is compared. Lubricants F and I are selected to verify simulations. As it mentioned in section 5.4.6, the CoF of F is around 0.08 and the CoF of I is around The cup samples drawn with lubricants F and I were cut in rolling and transverse directions, see Figure The thickness measurement along with rolling and transverse directions was completed by using Starrett Micrometer. The thickness comparison between experiment and simulation 54

79 is shown in Figure 5.11 to Figure It can be observed that the max prediction error is around 5% (see Figures 19 to 22), which shows that thickness distribution trends are predicted with reasonable accuracy. Figure 5.10: Cup samples deep drawn with lubricants F and I 55

80 Figure 5.11: Thickness comparison between experiment and simulation in rolling direction (cup drawn with lubricant F) 56

81 Figure 5.12: Thickness comparison between experiment and simulation in transverse direction (cup drawn with lubricant F) 57

82 Figure 5.13: Thickness comparison between experiment and simulation in rolling direction (cup drawn with lubricant I) 58

83 Figure 5.14: Thickness comparison between experiment and simulation in transverse direction (cup drawn with lubricant I) 5.5 Summary and Conclusions In this study, the lubricities of wet lubricants are compared with those of dry film lubricants, under the same operation conditions. Based on the cup draw tests, the following conclusions can be drawn: Cup Draw Tests (CDT) were used to evaluate the good lubricants and find out that G, K and L lubricants could be categorized as better performing wet lubricants and lubricant F can be considered to be the best performing lubricant among all the lubricants evaluated in the tests. 59

84 The friction behavior of dry film lubricants is compared with that of wet lubricants. It is found out that in this application, dry film lubricants perform better than wet lubricants, when the same amount of coating weight (1 gram/m 2 ) is applied. It is possible that dry film lubricants can be applied more uniformly by using electrostatic applicators. They reduce the friction and consequently, increases the process windows. However, dry film lubricants are in general difficult to clean after the forming process. This study shows that the friction behavior of Al 5182-O with EDT surface texturing is slightly better than that of Al 5182-O with MF surface texturing. The reason may be that the EDT finish has small pockets on the sheet surface and helps to maintain an even distribution of lubricant. CoF was predicted using FE simulations conducted using PAMSTAMP for the different experimental conditions. It can be concluded that i) the CoF of lubricant F is around 0.08; ii) the CoF of lubricants L and K is in the range of and iii) the CoF of lubricant C is around

85 Chapter 6. Design of a Tooling for the Servo Press Studies in Forming Aluminum Alloys 6.1 FE simulations for tooling design The objective is to make a flexible tooling that can be tried at various press conditions. Part shape was based on 2SL Rdpnl (see Figure 6.1) due to its height. Straight vertical walls were used to make the die flexible to obtain the max draw depth. Design was scaled down (see Figure 6.2 ) in order to fit AIDA s 300 ton servo press presses. Figure 6.1: 2SL Rdpnl (Provided by Honda) 61

86 Successful deep drawing process depends on the selection of process parameters as listed below: Die/punch radii Friction (CoF) Clearance between punch and die Blank holder force (BHF) In this study, the effects of die/punch radii, CoF and the clearance between punch and die on the formability of deep drawing process are investigated. The BHF was assumed to be the maximum value (250KN) that is available in the AIDA s 300 ton servo press, which will be used in the Al forming trials Determination of die radii The overall view of the initial die set, selected for the servo-press forming trials, is shown in Figure 6.2. It should be noted that this die design will be updated, since it does not include the protrusion features. The punch and upper die radii (Rp,Ru) influence the deformation of material during the deep drawing process. If the radii of the punch and die cavity edges are too large, wrinkling in the cup wall can occur. If the radii are too small, the blank is prone to tearing because of the high stresses. Hence, FE simulations listed in Table 6.1 were performed with different radii to identify the best punch and upper die radii FE model set up The FE simulations are conducted using PAMSTAMP The blank material is Al O, with 1.2 mm thickness. The shape and dimensions of blank, and the geometry of the tool are shown in Figure 6.2. The blank dimension of 1000 mm x 700 mm is used in 62

87 all the simulations. This blank size may need to be modified later. The blank shape is rectangular, to avoid additional blanking/shearing. It should be noted that there are no protrusion and draw beads in this tool. First, the blank holder force is applied on the blank. Then the upper die moves down and draws the blank against the punch and die cavity to form the part. Parameters used in the FE model are given in Table 6.2. Figure 6.2: FE model and the dimensions of the blank Material model Hill 48 yield function was used. The flow stress data (experimental data points) used in this simulation was obtained from bulge test, see Figure 4.3. The anisotropic values (r 0, r 45, r 90 ), yield stress (YS) and young s modulus were provided by Honda R&D, as shown in Table 4.1, in Chapter 4. Since we don t have our own forming limit curve for this material, the forming limit curve data provided by (Carsley et al., 2013) was used, see Figure 4.4. However, we will update our simulations once we have our own FLC of Al 5182-O. Other input simulation parameters are shown in Table

88 Table 6.1: List of Simulations to determine the punch/die radii ID CoF Punch radius (Rp, mm) Die radius (Ru, mm) Clearance between punch and die %t %t %t %t FE simulation results The thinning distribution of sheet metal with diverse values of the upper die and punch, are shown in Figure 6.3. It can be observed that the value of maximum thinning is reduced from 29% to 27% when punch radius is increased from 10 mm to 15 mm, as seen in Figure 6.3. It also should be noted that when upper die radius is increased from 20 mm to 30mm, as seen in Figure 6.3, the material on the both long sides of the part flows much more into the die cavity, which brought wrinkles in the side wall. Based on the simulations we performed, it is shown that, upper die radius of 20 mm and punch radius of 15 mm are recommended to reduce the maximum thinning. 64

89 Table 6.2: Input parameters to FE simulations Simulation parameters Descriptions Material Al 5182-O Material model Punch, blank holder and the upper die are assumed as rigid bodies Thickness (t), mm 1.2 Coefficient of friction (COF), µ 0.1 Blank Holder Force (BHF), KN 250 Stroke of upper die, mm 155 Upper die velocity (Virtual), mm/s 10 Anisotropic values, r0, r45 and r90 Flow stress data (actual experimental data points) Provided by Honda R&D, see Table 4.1 Obtained from CPF bulge test for 1.0 mm thick blank, see Figure

90 Figure 6.3: Thinning distribution of simulations #1.1,1,2,1,3 and Determination of clearance between punch and die The clearance between punch and die influences the deformation of material during the deep drawing process. If the clearance is too large, wrinkles may occur on the sidewalls of the drawn part. If the die clearance is too small, ironing can take place, which will increase the drawing load and the danger of cracking. Hence, FE simulations listed in Table 6.3 were conducted to identify the optimum clearance between punch and die. 66

91 FE model set up and material model The FE model and parameters used in the FE simulations of #2.1 and 2.2 are the same as for simulation #1.1, in section 6.1.1, except that 1) the punch and die radii are 10 mm and 20 mm, respectively and 2) the clearance will be varied from 110%t to 130%t. Table 6.3: List of Simulations to determine the clearance between punch and die ID COF Punch radius(mm) Die radius Clearance between (mm) punch and die (% of sheet thickness) % % % FE simulation results The thinning distribution of sheet metal with diverse clearance values are shown in Figure 6.4. It can be observed that the value of maximum thinning is reduced from 29% to 26% when the clearance is reduced from 130%t to 110%t. The probability of fracture predicted by using forming limit diagram (FLD) also shows that 110%t of clearance between punch and upper die is recommended, since the red region shifts down (the risk of rupture is reduced, see Figure 6.5). 67

92 Figure 6.4: Thinning distribution of simulations #1.1,2.1 and

93 Figure 6.5: Rupture risk distribution of simulations #1.1,2.1 and The effect of CoF on the formability of the part Friction plays an important role, which influences the stresses and strains in the work piece material and hence the quality of the part. FE simulations listed in Table 6.4 were performed with different CoFs to investigate the effect of CoF on the formability of this deep drawing process. 69

94 Table 6.4: List of simulations to investigate the effect of CoF on the formability of the part ID COF Punch radius(mm) Die radius (mm) Clearance between punch and die (% of sheet thickness) % % % FE model set up and material model The FE model and parameters used in the FE simulations of #3.1 and 3.2 are the same as for simulation #2.1, in section 6.1.1, except the Cofs (see Table 6.4) FE simulation results The thinning and rupture risk distributions of sheet metal with different CoFs are shown in Figure 6.6 and Figure 6.7. It can be observed that when the value of CoF increases, the value of maximum thinning occurred in the part decreases. The maximum thinning is around 25%, when CoF is The rupture distribution in Figure 6.7 shows that when the value of CoF decreases, the risk of rupture is reduced (red region shifts down). Therefore, it can be concluded that when the part is drawn with a better lubricant (the value of CoF is smaller), the formability of the part can be improved. 70

95 The material draw-in distributions of simulations #3.1 (CoF=0.08), 3.2 (CoF=0.12) and 2.1 (CoF=0.1) were plotted in Figure 6.8, which indicated that material flows much more in the die cavity when the smaller value of CoF is used, because the friction force is reduced. It should be noted, however, that the results are not entirely reliable because in the simulations, the die and blank hold plate are assumed to be rigid. This is not the case under the actual test conditions. Figure 6.6: Thinning distribution of simulations 3.1 (CoF=0.08), 3.2 (CoF=0.12) and 2.1 (CoF=0.1) 71

96 Figure 6.7: Rupture risk distribution of simulations 3.1 (CoF=0.08), 3.2 (CoF=0.12) and 2.1 (CoF=0.1) 72

97 Figure 6.8: Material draw-in distribution of simulations 3.1 (CoF=0.08), 3.2 (CoF=0.12) and 2.1 (CoF=0.1) The optimum parameters of tool design Based on all the simulations we have done so far, the optimum parameters for this tool could be set as listed in Table 6.5. Table 6.5: The optimum design of the tool ID COF Punch radius(mm) Die radius (mm) Clearance between punch and die %t 73

98 6.1.6 Summary and conclusions We studied the effects of die/punch radii, CoF and the clearance between punch and die on the formability of deep drawing process, assuming a constant BHF=250KN (max available in the servo press). The critical parameters that influence the deformation of the sheet material, are die/punch radii (Ru and Rp) and the clearance between punch and die. FE simulations were conducted to estimate the die/punch radii and the clearance. Summary and conclusions drawn from this study are as follows: 1) FE simulations were conducted with COFs of 0.08, 0.1 and FE results indicated that the rupture risk can be reduced, when CoF is changed from 0.12 to 0.08, which indicated that when the part is drawn with a better lubricant (the value of CoF is smaller), the formability of the part can be improved. These results may not be entirely reliable because in the simulations we assumed rigid die and blank holder. This assumption will not hold in actual tests because the tooling will show some elastic deflection. 2) FE simulations were conducted with clearances of 110%t, 120%t and 130%t. FE results indicated that 110%t clearance could have better forming results (less thinning) compared to those of 120%t and 130%t. 3) FE simulations were conducted with punch radii of 10 mm and 15mm and upper die radii of 20 mm and 30 mm with all possible combinations. FE results show that with the combination of punch radius of 15mm and upper die radius of 20 mm, the formability of the part can be improved (less thinning occurs). Therefore, clearance of 110%t, punch radius of 15mm and upper radius of 20 mm are recommended for this tool design. 74

99 6.2 FE simulations for protrusion design In order to investigate the forming of a complex part like inner door panel, two protrusions are added in the die sets as shown in Figure 6.9. The protrusion design is based on the Honda inner door panel design for steel. The protrusion radii vary from 6 mm to 9.6 mm. The angle α (see Figure 6.10) is The height of the protrusion is 10 mm (see Figure 6.10). The values of R1, R2 and α (see Figure 6.10) will need to be modified to make it suitable for Aluminum. In order to determine the optimum values for R1, R2 and angle α, a simple protrusion feature is assumed as shown in Figure 6.11 and Figure We assumed R1 is equal to R2 in order to minimize the variables. The values of R3 and R4 are calculated by changing the off-set between punch and die. The equations are listed as below: R3 = R1 1.1 t R4 = R t where 1) t is the blank thickness and 2) R1=R2. A list of simulations were conducted as shown in, to investigate which factor is more important between R1 and α. For simulations # 1.1 to 1.4, the radii (R1, R2) of the protrusion on the punch are varied and angle α is fixed; while for simulations # 2.1 to 2.4, the radii of the protrusion on the punch are fixed and angle α is varied. It should be noted that we assume R1 is equal to R2. 75

100 Figure 6.9: Punch with protrusion features Figure 6.10: Honda protrusion design based on inner door panel 76

101 Figure 6.11: A protrusion feature designed by CPF. Figure 6.12: A-A section view of protrusion 77

102 Table 6.6: list of simulations to determine R1, R2 and α FE simulations for simulations # (Angle fixed, radii varied) FE simulation set up For simulations #1.1 to 1.4, the punch radius varied from 3mm to 10 mm (R1=R2=3mm, 6mm, 8mm and 10mm). The thickness of blank is 1.2mm and the clearance between punch and die is 1.32 mm. The depth of the protrusion is chosen to be 10mm (h=10mm). The angle (α) between bottom and side wall of protrusion is 90 o, see Table 6.6. The simulation input parameters are the same as for simulations in section The material property is obtained from tensile test and bulge test, which was described in Chapter 4. Forming limit curve (FLC) is obtained from (Carsley et al., 2013), see Figure FE simulations results The thinning distribution of simulations # on the formed blank is given in Figure The same legend bar is used. It can be observed that the maximum thinning value on protrusion area decreases when R1 increases. Simulation results show that there is not too 78

103 much difference on thinning distribution when the value of R1 is larger than 8 mm. The probability of fracture predicted by using forming limit curve (FLC) is also plotted in Figure 6.14, which shows that 1) crack may occur on the formed part, when R1 is equal to 3 mm or 6 mm and 2) we may be able to form this part if R1 is equal to 8 mm or 10 mm. The thinning distribution and rupture distribution indicate that the optimum R1 value could be around 8 mm. Figure 6.13: Thinning distribution for simulations #

104 Figure 6.14: Fracture probability plot using the FLD for simulations # FE simulations for simulations # (Radii fixed, angle varied) FE simulations set up For simulations #2.1 to 2.4, the radii of protrusion on punch and die are fixed and angle α is varied. The radius on the punch will be redesigned to be 6mm (R1=R2=6mm). The depth of the protrusion h is chosen to be 10mm. The input parameters for FE simulation are the same as for simulations # FE simulations results The thinning distribution of simulations # on the formed blank is given in Figure The same legend bar is used. It can be observed that the maximum thinning value on protrusion area increases when angle α increases. When angle α is varied between 30 0 and 60 0, the simulation results give better results (less thinning) compared to that when angle 80

105 α is The probability of fracture predicted by using forming limit diagram (FLC) is also plotted in Figure 6.16, which shows that we may be able to form this part if angle α is equal to 30 0, 45 0 or The thinning distribution and rupture distribution indicate that the optimum α value could be around Conclusions The radius R1 is the main factor that leads to cracking. The minimum value of R1 we could use is around 8 mm for protrusion design. The max. thinning at the protrusion increases when α increases. The optimum α value could be 45 0 based on our FE simulation results. Figure 6.15: Thinning distribution for simulations #

106 Figure 6.16: Fracture probability plot using the FLD for simulations # Final die design and first tryouts with 820KN blank holder force using a Honda servo press Final Die design The overall dimensions of the experimental die (to be used in servo press forming) are illustrated. This final die design (provided by Honda) has 3 draw bead lines and two protrusion attributes, as shown in Figure Punch dimensions and blank holder plate dimensions are shown in Figure 6.18 and Figure 6.20, respectively. Punch radius of 10 mm and die radius of 20 mm were used in the final die design. The reason to add draw beads in this die set is to restrict the material flow into the die cavity. Otherwise, the material will flow too much into the cavity, as shown in Figure 6.19, 82

107 for a draw depth of 155 mm. By restricting the material flow, more stretching strain can be induced in the center of panel and a tighter appearance can be achieved. The radii (R1=12mm, R2=7mm), height (D=2mm) of draw beads (provided by Honda) and clearance (130%t) between punch and die are shown in Figure 6.20 and Figure Figure 6.22 shows the protrusion design in the final die. It should be noted that the protrusion design was based on our FE simulations in section 6.2. However, Honda selected punch radii and clearance based on 1) their experience and 2) our FE simulations in section 6.1. Honda selected GM338 steel as tool material, in order to avoid the galling issues that aluminum causes. Figure 6.17: Overall view of tool 83

108 Figure 6.18: Punch dimensions Figure 6.19: Material draw-in when the die set does not have draw beads (draw depth is 155 mm) 84

109 Figure 6.20: Blank holder plate dimensions 85

110 Figure 6.21: 2D section view of draw beads design 86

111 Figure 6.22: Protrusion design First tryouts conducted at Honda with this final die set Honda EGA conducted some experimental tryouts using a Honda s servo press, since AIDA s 300 ton servo press was not available at that time. The experimental data was available for two blank shapes (rectangular shape and curved shape, see Figure 6.23). Initially, the blank shape was rectangular, to avoid additional blanking/shearing. However, according to the preliminary experimental results conducted by Honda, this part can only be drawn up to 100 mm with this blank shape. Therefore, Honda EGA modified the blank shape as shown in Figure 6.23 (b). With this blank shape, this part was drawn up to 155mm without crack. The material draw-in information using 87

112 the curved blank shape is shown in Figure 6.24, provided by Honda. 820KN blank holder force was used during experiment, to reduce wrinkles. (a) (b) Figure 6.23: (a) Rectangular shape blank and (b) curved shape blank Figure 6.24: Material draw-in with curved blank shape, provided by Honda EGA 88

113 Figure 6.25: This part was drawn up to 100 mm depth and crack occurred afterwards, provided by Honda EGA Summary In this study, the overall dimensions of this final die design were illustrated. Using this final die set, Honda conducted some tryouts with two blank shapes (curved shape and rectangular shape, Figure 6.23). By using the curved shape blank, this part would be drawn up to 155 mm depth. We have material draw-in information from experiment with this blank shape, provided by Honda. For the rectangular shape, the part would only be drawn up to 100 mm and crack occurred beyond 100 mm. Since we have experimental data for the curved shape, we will validate our FE model in Chapter 7, by comparing our FE simulation results with experiment. Two kinds of FE 89

114 models in Chapter 7 will be used. One includes spacers; while another one does not include spacers. 6.4 Summary FE simulations were conducted to estimate the die/punch radii and the clearance in this Chapter. FE simulations were conducted with clearances of 110%t, 120%t and 130%t. FE results indicated that 110%t clearance could have better forming results (less thinning) compared to those of 120%t and 130%t. FE simulations were also conducted with punch radii of 10 mm and 15mm and upper die radii of 20 mm and 30 mm with all possible combinations. FE results show that with the combination of punch radius of 15mm and upper die radius of 20 mm, the formability of the part can be improved (less thinning occurs). In section 6.2, FE simulations were conducted, to obtain the optimum protrusion shape. It was found that protrusion radius should be around 8 mm and protrusion angle should be around 45 degree. Section 6.3 describes the final die design and the first tryouts conducted by Honda with a Honda servo press. Two blank shapes were tried in their tryouts. Material draw-in data is available for curved blank. For the rectangular blank, the material draw-in data is not available. In the next chapter, we will validate our FE model by comparing the FE simulation results with experiments. 90

115 Chapter 7. Validation of FE model and Prediction of Future Tryouts with Honda Die using FE Simulation 7.1 FE simulations of Honda part with 820KN blank holder force (without spacers) Simulations were conducted by using the final tool design (provided by Honda), with two blank shapes (see Figure 6.23), to validate our FE model by comparing the simulation results with experiments. It should be noted that we didn t include spacers in the FE model for this chapter FE model set up in Pamstamp FE simulations were conducted using PAMSTAMP The blank material is Al O, with 1.2 mm thickness. The geometry of the tool is shown in Figure Initially, the blank holder force is applied on the blank. Then the upper die moves down and draws the blank against the punch and die cavity to form the part. Curved blank shape (see Figure 6.23) was used, since experimental data is available with this shape, provided by Honda Material model in Pamstamp The anisotropy values, bulge test data and forming limit curve (FLC) are obtained from Chapter one. 820 KN blank holder force (suggested by Honda) was used, in order to reduce the wrinkles. Barlat 2000 yield criteria was used, to consider the anisotropic effect. The input parameters for Barlat 2000 is obtained from (Carsley et al., 2013), as shown in Table

116 Table 7.1: Input parameters for Barlat 2000 yield criteria, obtained from (Carsley et al., 2013) Barlat 2000 m=8 a1 a2 a3 a4 a5 a6 a7 a Comparison between Simulation results and experiment (curved shape, no spacer) FE simulation with curved shape (Figure 6.23) was also conducted, to compare FE simulation with experiment. Figure 7.1 shows the thinning distribution of the part, when stroke is 80 mm. The maximum thinning value is around 44%, which indicates that this part already cracked at 80 mm stroke. However, the experiment conducted by Honda shows that this part was able to be drawn up to 155 mm, when this blank shape was used. Therefore, we decided to tune our FE model by adding spacers, since spacers are usually used during the tryouts. FE simulations with spacers will be conducted and the simulation results will be compared with experiments in Chapter 6, to validate the FE model. 92

117 Table 7.2: Input parameters for FE simulation Simulation parameters Descriptions Material Al 5182-O Material model Punch, blank holder and the upper die are assumed as rigid bodies Thickness (t), mm 1.2 Coefficient of friction (COF), µ Blank Holder Force (BHF), KN Stroke of upper die, mm 155 Upper die velocity (Virtual), mm/s 10 Anisotropic values, r0, r45 and Provided by Honda R&D, see r90 Table 4.1 Flow stress data (actual experimental data Obtained from CPF bulge test, see Figure 4.3 points) 93

118 Figure 7.1: Thinning distribution of the part (no spacers, curved shape), at stroke of 80 mm (max thinning is 44%) Conclusions FE simulations with the final die design were conducted, with the curved blank shape. It should be noted that 1) spacers were not added in this study, 2) 820KN blank holder force was applied and 3) experimental data was provided by Honda. By comparing the simulation results with experiments, it was concluded that the FE model without spacers cannot predict results well for Honda s tryouts using the Honda servo press. 7.2 FE simulation of Honda part with 820KN blank holder force (with spacers) Spacers are metal pieces that are placed on the blank holder and generally do not come in contact with the sheet metal during forming process, see Figure 7.2. Spacers are usually used for locally controlling the material flow. Depending the height, the spacers create a 94

119 gap between the blank holder and the die, thereby affecting the pressure distribution on the sheet. Thus, spacers have a very local effect on the material flow, compensate elastic deflections in the dies and press, and they are less costly, compared to other available options. Figure 7.2: Schematic of a spacer used during forming process [Faass et al, 2008] The current practice of using spacers is purely based on experience. In this study, the effect of spacers on the material flow in forming this aluminum part is investigated, by using FE simulations. It should be noted that all the FE simulations conducted in this chapter have the following assumptions as listed below: The FE model has draw beads, see Figure All tools components are assumes to be rigid. However, elastic deflection will occur on the dies in the reality. Hence, we should consider about this elastic deflection when we do tryouts. 95

120 7.2.1 FE simulation of drawing process with spacers FE model set up In this study, the upper die, blank holder plate are modeled as rigid to reduce the computation time. However, it should be noted that there will be some elastic deflection on the die in reality, where spacers are in contact with the dies. FE model set up is the same as for simulations in section 7.1, except that a spacer was added between upper die and blank holder plate. Two blank shapes (see Figure 6.23) were used, since we have experimental data for these two blank shapes FE simulation results with rectangular blank shape (Cases I and II) Two FE simulations with two different spacer heights (1.5 mm and 1.4 mm, see Table 7.3) were conducted, to investigate the effect of different spacer height on formability. Rectangular blank shape was used in these two simulations. Table 7.3: Simulation matrix for rectangular blank shape (the die, punch and blank holder are assumed to be rigid) Case Height of spaces Blank holder CoF Blank shape (mm) force (KN) I 1.5 mm rectangular II 1.4 mm rectangular Thinning distribution, at different strokes (100 mm and 150 mm), are shown in Figure 7.4 and Figure 7.5. The spacer height is 1.5 mm (120% of sheet thickness). It can be seen that 1) the maximum thinning value increases from 18% to 22%, when the die stroke increases from 100 mm to 150 mm and 2) there are some wrinkles in the flange area, due to the spacer between the upper die and blank holder plate. However, it should be noted that the maximum thinning value (22%) is less compared to 30% when the spacer was not added 96

121 in the FE model. The contact pressure is reduced and material flows more easily due to the spacer. Since there are a lot of wrinkles in the flange area, a lower spacer height (1.4 mm) was used, to increase the stretching force on the blank. Figure 7.6 and Figure 7.7 show the thinning distribution of the part, when strokes are 100 mm and 150 mm, respectively. It can be observed that the values of the max thinning increases from 22% to 24%, when the height of spacer is reduced from 1.5 mm to 1.4 mm. It can be explained that with a lower spacer height, the contact pressure on the part is higher and the material will be more restricted to flow into the cavity. The material draw-in comparison between Pamstamp simulation results and Autoform simulation results (by Honda) is shown in Figure 7.3, when the rectangular blank is used. It should be noted that the spacer height is 1.5 mm for this simulation. It can be observed that FE simulation results match well with experiment, see Figure 7.3. However, it should be noted that the material draw-in can be adjusted, when the height of spacer is changed. 97

122 Figure 7.3: Material draw-in comparison between simulation and experiment Figure 7.4: Thinning distribution of the part (Case I), at stroke of 100 mm (max thinning is 18%) 98

123 Figure 7.5: Thinning distribution of the part (Case I), at stroke of 150 mm (max thinning is 22%) Figure 7.6: Thinning distribution of the part (Case II), at stroke of 100 mm (max thinning is 19%) 99

124 Figure 7.7: Thinning distribution of the part (Case II), at stroke of 150 mm (max thinning is 24%) FE simulation results with curved blank shape (Case III) We also conducted one more simulation with curved blank shape, since Honda provided us the experimental data on material draw-in, with this blank shape. Table 7.4: Simulation matrix for curved blank shape (the die, punch and blank holder are assumed to be rigid) Case Height of spaces Blank holder force CoF Blank shape (mm) (KN) III 1.5 mm Curved blank The material draw-in comparison between simulation result and experiment is shown in Figure 7.8. It can be observed that FE simulation results match well with experiment. The thinning distribution is shown in Figure 7.9 and the maximum thinning value is around 20%, when die stroke is 155 mm. This part a) could be drawn up to 155 mm when this 100

125 blank shape was used, according to the experimental data provided by Honda and b) material draw-in comparison between FE simulation and experiment is quite close. Therefore, we can conclude that the maximum allowable thinning before crack occurs may be around 20%. It should be noted that the material draw-in and thinning distribution can be adjusted, when the height of spacer is changed. However, since we don t know the final spacer height and location Honda EGA used, there are some difference between FE simulation results and experiments. We can update our FE simulation when these information is available and the FE simulation can predict more accurate results. Figure 7.8: Material draw-in comparison between simulation and experiment 101

126 Location A Figure 7.9: Thinning distribution at stroke of 155 mm, with curved blank shape Conclusions Spacers were added in the FE model in this study, since it is quite common to add spacers during tryouts. The FE model and input parameters are the same as for FE simulations in section 7.1. FE simulation results with experiments were compared. It was found that FE simulation with the added spacers can predict more accurate results than that when spacers were not added, for Honda s tryouts using the Honda servo press. Based on the material draw-in length comparison between FE simulation results with experiments, it can be concluded that 1) the maximum thinning always occur at the location A as shown in Figure 7 10 and 2) when maximum thinning value exceeds 20% at that location, cracking may occur in the aluminum part. 102

127 7.3 FE simulation results with 250KN blank holder force for AIDA s tryouts (without spacer) FE simulation with 250 KN blank holder force was conducted, to check the feasibility of AIDA s 300 ton servo press on forming this aluminum part, with current die design. 250KN is the maximum blank holder force that AIDA s 300 ton press has. Four different blank shapes were tried to check the effect of blank shape on the part formability. Rectangular shape and curved shape were provided by Honda EGA. Other two blank shapes were modified based on the curved shape, since it would be much easier for us to shear the blank. It should be noted that we didn t include spacers in the FE model, since we want to reduce wrinkles in the part FE simulation set up FE model and the input parameters are the same as for FE simulations in Chapter 5, except that 1) blank holder force is 250 KN and 2) spacer is not included. Four FE simulations were conducted with three different blank shapes, as listed in Table 7.5. Table 7.5: Simulation matrix for rectangular blank shape (the die, punch and blank holder are assumed to be rigid) Case Draw depth (mm) Blank holder force (KN) CoF Blank shape A 155 mm Rectangular, see Figure 7.10 B 155 mm Curved blank, see Figure 7.10 C 155 mm Rectangular with four corners chamfered, see Figure 7.10 D 155 mm Rectangular with four corners chamfered (center of the blank is shifted), see Figure

128 Figure 7.10: Rectangular shape with four corners chamfered FE simulation results The thinning distribution of simulations A, B, C and D are shown in Figure The maximum thinning values are around 19%, 15%, 15% and 15% (see Figure 7.11), for these four simulations, respectively. However, severe wrinkles occurred along I to II region, as shown in Figure In order to compare the wrinkling distribution in the formed part, the thinning distributions along I to II were plotted in Figure 7.12 for these four simulations. It can be observed that 1) severe wrinkling occurred in the flange area when rectangular shape was used, 2) curved blank and rectangular blank with four corners chamfered (Case B and Case C) present similar thinning distribution (see Figure 7.12) and 3) Case D gives the best simulation results (fewer wrinkles and less thinning) among these four simulations. The probability of fracture (using FLD) as obtained from simulations B and C (see Figure 7.13) also show that there is not too much difference between Case B and Case C. Therefore, we can conclude that 1) the rectangular shape with four corners 104

129 chamfered will be a better shape option to run the tryouts using Honda die, since it will be much easier to be sheared compared to the curved blank and 2) we can shift the blank center by 5 mm (see Figure 7.10, Case D) when we run tryouts to reduce wrinkles. Figure 7.11: Thinning distribution of simulations A,B,C and D. 105

130 Figure 7.12: Thinning comparison along A to B between simulations A,B,C and D Figure 7.13: Probability of Fracture (using FLD) from Simulations C (left one) and D (right one) 106

131 7.4 Summary and Conclusions Results of FE simulations with the final die design are discussed in this Chapter. It should be noted that 1) spacers were not added in the simulations in section 7.1, 2) 820KN blank holder force was applied and 3) experimental data was provided by Honda. FE simulations were conducted with curved blank shape. By comparing the simulation results with experiments, it was concluded that the FE model without spacers cannot predict results well for Honda s tryouts using the Honda servo press. Spacers were added in the FE model in the study of section 7.2, since it is quite common to add spacers during tryouts. The FE model and input parameters are the same as for simulations in Section 7.1. FE simulation results with experiments were compared. It was found that the FE simulation when spacers are included, can predict results more accurately than when spacers were not added, for Honda s tryouts using the Honda servo press. After we validated our FE model, we used 250KN blank holder force instead of 820KN, to optimize the blank shape for future tryouts, using AIDA s 300 ton servo press. Four FE simulations were conducted with four different blank shapes as described in section 7.3. It was found that 1) the rectangular shape with four corners chamfered can be formed with similar results as the curved shape and 2) by shifting the blank 5mm away from die center, we can reduce wrinkles. Therefore, the rectangular shape with four corners chamfered will be a better option to run the tryouts using Honda die, since the samples will be much easier to shear compared to curved blank. 107

132 Chapter 8. Conduct Tryouts on the Servo Press using Shiloh Die In this chapter, we summarized the results of the tryouts using AIDA s 300 tons servo press with Shiloh die, since the tooling we designed (Honda die) was not available until August, Shiloh die was designed by CPF/OSU and manufactured by Shiloh Industries Inc., for Advanced High Strength Steels (AHSS) tryouts using AIDA s 300 ton servo press. 8.1 Shiloh die tooling and AIDA s 300 ton servo press Figure 8.1 shows the Shiloh die in the AIDA s 300 ton servo press. The front view and cross-section view of the Shiloh die are shown in Figure 8.2. The maximum depth of this die is around 80 mm. The detailed information about this Shiloh die can be found in CPF report 1.4/14/01 (Adam Groseclose, 2014). It should be noted that the clearance between punch and die of the Shiloh die is 1.6 mm; while the thickness of aluminum samples is only around 1.2 mm. The clearance between punch and die is almost 133% t, which may be too large for this material. It may cause a lot of wrinkles in the part when we use this die for aluminum forming. AIDA s 300 ton servo press is a servo motor driven press. The maximum cushion force is 250 KN and the cushion force can be varied through the stroke. The maximum die cushion stroke is 150 mm. It also has the pre-acceleration function, which can be controlled to several values (without, strong, medium and weak). This press has data measurement and storing function, which can record slide and DC positions, slide and DC speeds, press load 108

133 and DC load s data. For this study, we used this data storing function to record all the data for each test. Before we conducted the Aluminum tryouts, we conducted FE simulations to estimate the blank size, blank holder force and draw depth for future tryouts using Shiloh die. Figure 8.1: Shiloh die installed in the AIDA s press 109

134 Figure 8.2: Shiloh die front and cross-section view (Adam Groseclose, 2014) 110

135 8.2 FE simulations to predict the process parameters will be used in the tryouts In this Task, FE simulations of the tooling, provided by Shiloh Industries, are conducted to estimate the work range of blank holder force, blank size and draw depth for the aluminum forming by using AIDA s 300 ton servo-press. It is necessary to make sure that whether forming is possible or not at a very early stage before we do any try-outs FE mode set up The FE model is illustrated as Table 8.4. Blank thickness is 1.2 mm. The input material properties were obtained from bulge test, described in section 4.2. The dry film lubricant selected from the lubrication study (CPF report 2.2/13/04), was pre-coated on both sides of the blank by Honda EGA. Based on the lubricant study in section 5.4.6, the coefficient of friction is assumed as 0.1 for FE simulations. The maximum limitation of blank holder force (BHF) of the AIDA S press is 250KN. Simulation inputs are listed in Table 8.1. In order to investigate the work range of BHF, several simulations were conducted listed in Table 8.2. Both blank shapes (A and B, see Figure 8.4) were simulated under BHFs of 50 KN, 100 KN, 150 KN and 250 KN. The simulation matrix is given in Table 8.2. Figure 8.3: FE modeling and inputs in Pam-stamp 111

136 Table 8.1: Simulation inputs (blank shapes A&B are shown in Figure 8.4) Material Al 5182-O Thickness 1.2 mm Flow stress data CFP bulge test Max. BHF 250 KN CoF 0.1 Draw depth 75 mm Blank shape A B Rectangular Chamfered Table 8.2: FE simulation case study (blank shapes A&B are shown in Figure 8.4) CASE Blank shape Blank holder force (KN) Draw depth (mm) I A/B 50 II A/B 100 III A/B IV A/B

137 8.2.2 Investigate the effects of blank shapes on part drawability In this study, two blank shapes were designed, and the corresponding draw depth was 75 mm (the limit stroke of the designed tooling is 90 mm). The dimension for rectangular blank is 720*500 mm. In order to release the compression stress at four corners, the chamfered blank shape was also designed, as shown in Figure 8.4. (A) Rectangular (B) Chamfered Figure 8.4: Blank shapes designed for future tryouts The thinning distributions with the two shapes were shown in Figure 8.5. It is found that the maximum thinning locations were shown at the same corner for rectangular and chamfered blanks. It can be observed that the difference on thinning distribution by using these two different blank shapes is very small. Thus, either blank shape could be applied for further tryouts. 113

138 (A) Rectangular (B) Chamfered Figure 8.5: Thinning distributions at stroke of 75mm, BHF=150KN Investigate the effect of lubrication The aluminum blanks were provided by Honda EGA. All the blanks were pre-coated with the best dry film lubricant according to the former lubrication study conducted in Chapter 5. Also, the coefficient of friction (Cof) was numerically determined as 0.1. The Cof for non-lubrication conditions was assumed to be 0.2. In this section, two FE simulations were conducted with different Cofs (0.1 and 0.2). The parameters are listed in Table 8.3. The 75 KN blank holder force was used. 114

139 Table 8.3: Simulation matrix for friction study Blank CASE Blank shape Cof holder force (KN) V B 0.1 VI B As we know, friction has a significant effect on forming process. Figure 8.6 illustrates the maximum thinning values at different strokes. It shows a notable increase under Cof 0.2 at stroke 60mm. The maximum thinning is around 20%, which could indicate a fracture according to Honda s experience. At Honda EGA, in the deep drawing of Al5182-O, 20% is a critical maximum thinning value to predict fracture. Usually, the possible fracture locations are the cavity corners around the punch nose. In this case, cracking may occurs at corners A and B (see Figure 8.7), because these two locations locate at the concave side of the die. Compared to location A, location B may have a higher probabilities to have cracking, (see Figure 8.7), since the radius punch of location B is smaller than that of location A (see Figure 8.2). 115

140 Figure 8.6: Max. thinning values at selected strokes, under Cof 0.1 and Cof 0.2 Figure 8.7: Thinning values at locations A and B 116

141 8.2.4 Investigate the effect of BHF Constant BHF through the stroke The blank holder force is a very important process parameter to improve the formability. In deep drawing process, a higher BHF could suppress wrinkles at flanges and walls, however, it also would lead to cracking. In order to investigate the effects of BHF on forming Al 5182-O with current Shiloh die, constant BHFs were applied. FE simulations under constant BHFs from 50KN to 250KN were conducted and the results are given in Table 8.4. Table 8.4: Maximum thinning for simulation cases I-IV CASE Blank shape Cof Blank holder force (BHF, KN) Max. thinning I % B, see II % Figure 0.1 III % 8.4 IV % Figure 8.8 shows the thinning distribution under constant BHF of 250KN. The maximum thinning value is 18.4%, which is lower than 20%. Based on the Honda s experience on aluminum forming, it can be concluded that based on the FE simulation results, the parts coated with dry film lubricant could be formed in the range of 50KN to 250KN BHF. 117

142 Figure 8.8: Thinning distribution under BHF 250KN Variable BHF through the stroke In the practical deep drawing process, a constant BHF is applied over the working stroke. However, recent studies show a better formability could be obtained from a variable BHF. Since AIDA 300 ton servo press can vary the blank holder force through the stroke, we have investigated the effect of variable BHFs on part drawability in our study. BHF profiles G, F and H were designed, as shown in Figure 8.9. The input parameters of FE simulation matrix under BHF profiles G, F and H were given in Table 8.5. Chamfered blank shape was used. The maximum thinning occurred at the corner (location B in the part, see Figure 8.10), which is marked with a dashed line. Element a in this area was selected to show the thinning-stroke curves, under BHF profiles G, F and H when different Cofs (0.2 and 0.1) were used. It is found that, the results (thinning vs stroke curves) are close when blank 118

143 holder profiles G, F and H were used. It should be noted that only blank shape B was used in simulations listed in Figure 8.5. G F H Figure 8.9: Profiles of designed blank holder force Table 8.5: Simulation matrix of variable BHF study (blank shape B is shown in Figure 8.4) CASE Blank holder force profile Cof Blank shape VII F VIII G 0.2 B IX X H F XI G 0.1 XII H 119

144 Figure 8.10: Simulation results (thinning) when different blank holder profiles (A, B and C, see Figure 8.9) were used 8.3 Conduct tryouts on the servo press Preparation for experiments Experiments with the designed blank size and blank holder forces estimated in section 8.2 were tried for the selected aluminum alloy (Al 5182-O, 1.2 mm) on the AIDA s 300 ton servo press. Two blank shapes (see Figure 8.4) were used during tryouts. The blanks were cut to the desired dimensions predicted by FE simulations conducted at section 8.2 and were coated with the best dry film lubricant by Honda EGA. During the tryouts, we used the guides (see Figure 8.11) to position the blanks, to keep the consistence on blank s position. Concerning drawability, the biggest advantages of servo press are to 1) vary the blank holder force through the stroke and 2) change the ram profile through the stroke. Therefore, two tasks will be conducted as listed below: Task I: investigation of the effect of ram speed through stroke on part drawability 120

145 Task II: investigation of the effect of variable BHF profile through stroke on part drawability The test matrix of Task I is shown in Table 8.6. In this task, we will use different ram speed profile as shown in Figure For the speed profiles (A-G-H-J, A-C-D-J and A-E- F-J), when the upper die contacts the sheet, the ram speed will keep constant during deformation. For the speed profile (A-F), when the upper die contacts the sheet, the ram speed will follow the sine curve as shown in Figure The test matrix of Task II is shown in Table 8.7 and Figure Blank shape B (see Figure 8.4) was used for all the tryouts, except that we used one rectangular blank shape in first tryout, to check the part drawability difference by using different blank shapes (A and B, see Figure 8.4 ). It should be noted that the draw depth and cushion force listed in these tables may need to be adjusted according to the try-outs on the servo press. 121

146 Figure 8.11: Guides used to position the blanks (Blank shape B (see Figure 8.4) was used in this figure) 122

147 Table 8.6: Test matrix of Task I Aluminum Ram speed Blank holder Draw depth Number of sample # profile force (Sd, mm) Samples (BHF) 1(a) (b) (c) A-J, see Figure KN (a) (b) (c) A-G-H-J, a higher speed, see Figure KN (a) (b) (c) A-C-D-J, a middle speed, see Figure KN (a) (b) (c) A-E-F-J, a lower speed, see Figure KN

148 Blank holder force (KN) Figure 8.12: Different ram speed profiles will be used in Task I. VBH1 VBH2 contant BHF Stroke (mm) Figure 8.13: Varied blank holder force profiles prepared for Task II 124

149 Table 8.7: Test matrix of Task II Blank holder Best ram speed Draw Number of Aluminum sample # force (BHF, KN) obtained from tests listed in Table 8.6 depth (mm) Samples 5 (a) (b) (c) 150 KN, see Figure mm/s (a) (b) (c) 75 KN 250 KN, see Figure mm/s (a) (b) (c) 250 KN 75 KN, see Figure mm/s Preliminary tryout results (blank applied with dry film lubricant) We did some preliminary tryouts using aluminum blanks coated with the dry film lubricant. Blank shapes A and B (see Figure 8.4) were used. The test matrix is shown in Table 8.8. The formed samples are shown in Figure It can be concluded that based on the test results all the parts coated with dry film lubricant can be formed without causing any cracking. However, when the blank holder force of 50KN was used, there are severe 125

150 wrinkles occurred in the part (see Figure 8.14). Thus, we increased the blank holder force up to 150 KN to reduce wrinkles. The maximum blank holder force of 250 KN was also tried to check the maximum draw depth of this part. It was found out that with this lubricant, this part is able to drawn up to the maximum draw depth of the tool (80.8 mm) without causing any cracking when the blank holder force is in the range of 150 KN to 250KN. It should be noted that we only tested one sample for each test condition. Table 8.8: Test matrix for preliminary test (Blank shapes are shown in Figure 8.4) Aluminum sample # (blank shape) Cushion Force (KN) Blank shape (A/B) Ram speed (spm) depth 1_a 50 B _a 150 B _a 250 B _a 150 A _a 250 B

151 Figure 8.14: Samples with different test conditions for the preliminary tests (see Table 8.8 for the test conditions) Second tryouts (using blanks without applying lubricants) For the second tryouts, we removed the lubricant from blank using Acetone and only tested one sample per set-up. During the tryout, we wiped the dies using paper towels. Blank shape B (see Figure 8.4) was used for all the tests conducted in the second tryouts. We conducted few tests with different draw depth (60.8mm, 65.8mm and 75.8 mm) and found out that this part was able to form at the draw depth of 60.8 mm, when 18 spm and 100 KN blank holder force were used (see Table 8.9). Hence, we selected the sample 6_a as a base and lowered the ram speed while keeping all other parameters (blank holder force and draw depth) the same, to investigate the effect of ram speed on the part formability. Table 8.9 shows the test matrix we used during the tryouts. Figure 8.16 shows the locations A and B 127

152 in the part, where cracking may occur as we predicted in FE simulation in section 8.2. Figure 8.17 shows the comparison between samples 6_a, 7_a and 8_a. It should be noted that for samples 6_a, 7_a and 8_a, we used the crank motion to form aluminum parts. 18 spm, 10 spm and 1 spm were used for samples 6_a, 8_a and 7_a, respectively. The ram speed vs stroke curves and contact speeds are shown in Figure It was found out that 1) using a higher speed (18 spm) there is no cracking on the aluminum part and 2) cracking occurred on the part when we used lower speeds (10 spm and 1 spm). The reason may be that when we increased the speed, the interface condition between workpiece and dies were changed and the blank can be drawn into die cavity much easily. We also used the constant speeds (310 mm/s and 50.6 mm/s, see Figure 8.18) to form the parts during the deformation stroke. In this case, when the upper die contact the blank, ram speed keeps the constant speed during deformation. Table 8.10 shows the test matrix we used during our tests. Figure 8.17 shows that 1) cracking didn t occur on the part when we used a high speed of 310 mm/s and 2) part got cracking when we used a lower speed of 50.6 mm/s. These results also indicate that we may get better formed part when we increase the forming speed. However, it should be noted that we only used one sample for each test condition for this tryout. We will conduct another tryouts using 3 samples for each set-up. The press total load vs slide position curves were plotted in Figure 8.19, when 18 spm and 10 spm were used. It can be observed that the curves are overlapping. It can be concluded that this material is not stain-rate sensitive in the speed range of 18 spm (contact speed is 310 mm/s) to 10 spm (contact speed is 172 mm/s). It should be noted that the press total load is equal to punch force plus blank holder force. 128

153 Table 8.9: Investigation on the effect of speed (crank motion) on Aluminum part formability for second tryouts. Aluminum (sample #) Blank holder Force (KN) Ram speed Peak Force (KN) Contact speed (mm/s) Draw Depth (mm) crack 6_a (max) no 7_a (min) yes 8_a yes 129

154 Figure 8.15: Comparison on samples 7_a, 8_a and 6_a (see Table 8.9 for the test conditions) Figure 8.16: Locations A and B in the aluminum sample 130

155 Table 8.10: Investigation on the effect of speed (constant speed) on Aluminum part formability Aluminum(sample #) Blank holder Force (KN) Ram speed Total Force (KN) Draw Depth (mm) crack 9_a mm/s no 10_a mm/s yes Figure 8.17: Comparison between samples 9_a and 10_a by using different constant speeds during deformation (310 mm/s and 50.5 mm/s), see Table 8.10 for the test conditions 131

156 Figure 8.18: Ram speed vs punch stroke curves used in the tryouts (output data from AIDA s press) 132

157 Sample 8a, spm=10 Sample 6a, spm=18 Figure 8.19: Press total load vs slide position when 18 spm and 10 spm were used Third tryouts (using blanks with dry film lubricant) For the third tryouts, we ran few tests with different draw depth (60.8mm, 65.8mm and 75.8 mm) and found out this part was able to form at the draw depth of 60.8 mm, when 10 spm and 125 KN blank holder force were used. 125 KN was selected since this is the minimum blank holder force that can avoid severe wrinkles mm of draw depth was selected since the part got cracking when the draw depth is larger than 60.8 mm. It can be observed that the results we obtained from this test are different with the results obtained in section The possible reason is listed below: For the third tryouts, we cleaned dies using Acetone; while for the first and second tryouts, we just wiped dies with paper towels. 133

158 Even the results of the third tryouts are different with the results obtained from the preliminary tryouts, we still should be able to compare the difference on the part by using different ram speed as listed in Table Three samples (a, b, c) were used for each test condition. Figure 8.18 shows the different ram speed profiles used in these tryouts. Table 8.11: Investigation the effect of ram speed on part drawability Aluminum (sample #) Blank holder Force (KN) Ram speed Peak Force (KN) Contact speed (mm/s) Draw Depth (mm) Crack (yes/no) 12_a 12_b 12_c spm no 13_a 13_b 13_c spm no 14_a 14_b 14_c mm/s no 15_a 15_b 15_c mm/s no Figure 8.20 and Figure 8.21 show the formed parts when the different ram speed profiles were used. It should be noted that only one sample for each set-up condition was shown in the Figure 8.20 and Figure 8.21, since 3 parts (a, b and c) look similar for each set-up 134

159 condition. Based on the test results, it can be observed that 1) cracking occurred at location B when the constant speed of 50.5 mm/s was used and 2) there is no cracking or necking occurred when the speed of 310 mm/s was used. It should be noted that the formed parts look similar when 10 spm and 1 spm were used. However, it should be noted that 1 spm is very slow, most of mass production products will not use this speed. These results indicate that we can get better formed part when we use a large forming speed (around 310 mm/s) during aluminum deep draw process. The reason may be that when we increased the speed, the interface condition between aluminum workpiece and dies were changed and the blank can be drawn into die cavity much easily. 135

160 Sample 15_a Location B Sample 14_a Location B Figure 8.20: Formed parts with constant speeds of 50.6 mm/s and 310 mm/s during the deformation (see Figure 8.11 for the test conditions) Location B Sample 13_a 10 spm Location B Sample 13_a 1 spm Figure 8.21: Formed parts with ram speed of 10 spm and 1 spm (see Figure 8.11 for the test conditions) 136

161 8.3.5 Observation and discussions Based on the experimental results, it can be concluded that we can get better formed part when we use a large forming speed (around 310 mm/s) during aluminum deep draw process, after we cleaned the dies using acetone. The possible reasons are listed below: Strain rate effect Temperature effect Lubrication effect Strain rate effect Based on the information listed in paper (Lademo et al., 2012), it is shown that the strain rate sensitivity is almost negligible in the strain rate range of 1 to 10-4, see Figure The material is Al 5182-O, with thickness of 1 mm. Therefore, the material properties of Al 5182-O will keep the same even we use a higher speed (310 mm/s) to form the part. Figure 8.22: the flow stress curves under different strain rates (1 to 10-4 ) (Lademo et al., 2012) 137

162 Temperature effect (Abedrabbo et al., 2007) investigated the effect of temperature on the flow stress curves. Figure 8.23 shows that the stress-strain curves are almost overlapping when temperature is lower than 93 o C. The material is Al 5182-O, with thickness of 1.15 mm. Therefore, the effect of temperature on the material properties is also negligible. Figure 8.23: flow stress curves under different temperatures (25 o C to 280 o C) (Abedrabbo et al., 2007) Lubrication effect Friction/wear study conducted by (Fenske et al, 2008) shows that the Cof decreases as sliding speeds increases. It should be noted that the disc-compression test does not present real stamping process, it still can show the trend that when sliding speed increases, Cof decreases. Therefore, it is possible that the lubrication condition between workpiece and 138

163 dies becomes better when a higher forming speed is used. However, in order to verify this idea, the lubrication test at different sliding speeds can be conducted in the future to future analyze the relationship between Cof and sliding speed. Figure 8.24: Cof vs sliding speed (Fenske et al, 2008) 139