- Email: [email protected]
Effect of process parameters on deformation behavior of AA 5052 sheets in stretch flanging process
Available online at www.sciencedirect.com
ScienceDirect Materials Today: Proceedings 4 (2017) 9316–9326
www.materialstoday.com/proceedings
ICAAMM-2016
Effect of process parameters on deformation behavior of AA 5052 sheets in stretch flanging process Yogesh Dewanga*, M.S. Horab, S.K. Panthic a
Assistant Professor, Department of Mechanical Engineering,Lakshmi Narain College of Technology, Bhopal,462021, India b Professor, Department of Civil Engineering,Maulana Azad National Institute of Technology, Bhopal, 462003, India c Scientist,Computer Simulation & Design Centre,Advanced Materials Processes and Research Institute,Bhopal,462024, India
Abstract Aluminum alloys are widely used in manufacturing of aircrafts and in automotive industry. In this paper aluminum alloy 5052 of 0.5 mm thickness was utilized to investigate deformation behavior in stretch flanging process. The effect of process parameters such as initial flange length, punch-die clearance and blank-holding force were investigated using FE simulation and experimentation. The results show that the circumferential strain and radial strain increases with increase in initial flange length and blank-holding force and decreases with increase in punch-die clearance. Experimental investigations were also carried out to validate FE simulation results. A very good agreement was obtained between FE simulation results and experimental in terms of punch load and edge crack location and propagation. © 2017 Elsevier Ltd. All rights reserved. Selection and Peer-review under responsibility of the Committee Members of International Conference on Advancements in Aeromechanical Materials for Manufacturing (ICAAMM-2016). Keywords: stretch flanging; finite element simulation; punch-die clearance; initial flange length; blank-holding force; edge crack.
* Corresponding author. Tel.:+91-0755-2457105; fax: +91-0755-2457042. E-mail address:[email protected]
2214-7853© 2017 Elsevier Ltd. All rights reserved. Selection and Peer-review under responsibility ofthe Committee Members of International Conference on Advancements in Aeromechanical Materials for Manufacturing (ICAAMM-2016).
Yogesh Dewang et al / Materials Today: Proceedings 4 (2017) 9316–9326
9317
1. Introduction Flanging is one of the important sheet metal forming processes which is applied in automobile industry. Flanging is a sheet metal bending operation in which sheet is bent usually to 900 for providing a smooth rounded edge, higher or strength to the edge of sheet-metal parts. It is also used for making hidden joints and assembly of automobile parts. Flanging can be classified on the basis of deformation features into two major types as shrink and stretch flanging. In shrink flanging, the curvature of flange is convex and metal is in compression in circumferential direction. The arc length of final flange is smaller than arc length of original element. The compression is greatest at the top of flange and least at die profile radius. In stretch flanging, the curvature of flange is opposite to shrink flanging, i.e. concave and sheet metal undergoes tension in circumferential direction. The tension is greatest at top of flange resulting in increase in the length of arc. The feasibility of any sheet metal forming process depends upon various processes, geometric as well as material parameters. To obtain defects free products different parameters need to be optimized. Different types of failure like necking, wrinkling and fracture are involved in sheet metal forming processes. With the growth and development of commercially available FEM software it is become easier to predict these types of defects. Finite element simulation is an accurate and faster method for prediction of failure in sheet metal forming processes because it captures realistic phenomena with less cost and time involved. Researchers in past carried out work in the area of the failure analysis in sheet metal forming processes using finite element simulation. Brunet et al. [1] used a gurson-tvergaard ductile damage based model for prediction of necking and fracture for anisotropic sheet metals. Teixeira et al. [2] coupled Hill’s orthotropic plasticity criterion with Lemaitre’s ductile damage evolution law for prediction of ductile fracture in sheet metal forming processes. In deep drawing process, usually two types of failure modes namely fracture and wrinkling were encountered. Li et al. [3] applied modified Mohr-Coulomb fracture criterion in the deep drawing of AHSS using both circular and square punch for prediction of shear induced fracture in terms of crack initiation, propagation and location of cracks. Wu et al. [4] found that that low blank-holding force will give rise to wrinkling while increase in blank-holding force will lead to fracture during deep drawing of sheet metal with nickel. Nguyen et al. [5] predicted wrinkling and fracture in deep drawing of SPCC material by using forming limit diagram ductile fracture criterion. On the other hand, in case of flanging process the major type of defects were necking, fracture and wrinkling. The major types of flanging are straight, stretch, shrink and hole-flanging in which studies were carried out. Kacem et al. [6] predicted necking, damage, fracture and limit of flanging using a macroscopic ductile fracture criterion by taking into account the triaxiality of material in hole-flanging of AA 1050-H-14 and AA-6061.In another type of flanging process i.e., shrink flanging wrinkling was observed as failure. Wang et al. [7] predicted wrinkling by using energy approach based on stress-based criterion. Some studies were also carried out in the domain of flanging process in order to study the effect of various parameters on formability of flange using various theoretical, experimental and FEM based approaches. Hu et al. [8] proposed analytical models for introfelxion/stretch and outcurve/shrink flange which takes into account the effect of planar anisotropy in flanging process for prediction of blank’s trim line. Dudra and Shah [9] developed analytical model for development of trimline for axisymmetric case and utilized FEM method for non-axisymmetric case for prediction of trimline and peak strain. Bao and Huh [10] implemented one-step analysis in trimming die design for effective design procedure. Li et al. [11] used an equivalent 2D FEM program for parametric study in stretch flanging of V-shaped blanks. Wang et al. [12] utilized an elastic-plastic FEM program in order to study the effect of various materials and process parameters on the formability of flange in stretch flanging of V-shaped blanks. Yeh et al. [13] determined optimum blank shape by using forward-inverse prediction scheme which integrates explicit dynamic FEM, true strain method and adaptive-network-based fuzzy inference system in stretch flanging. In stretch flanging, the predominant mode of failure was along edge as fracture and localized necking. Wang et al. [14] proposed localized necking for stretch flanging of AKDQ steels based on modification of Hill’s instability criterion by taking into account the effect of strain-hardening and plastic anisotropy. Asnafi [15] found that fracture limit was obtained by increasing plastic strain ratio, strain hardening exponent and uniform strain in stretch flanging through fluid forming of aluminum alloys.Worswick and Finn [16] found through numerical simulation that strain predictions were more accurate with Barlat-89 and Hill-48 yield criteria than von-Mises yield criteria in stretch flange forming. Feng et al. [17] recommended small flange height and small ratio of straight side length to the curved range radius in stretch curved flanging in order to avoid fracture at edge of flange.
9318
Yogesh Dewang et al / Materials Today: Proceedings 4 (2017) 9316–9326
In finite element simulation of stretch flange forming the concept of continuum mechanics was employed by various researchers for prediction of failure prediction of failure. Chen et al. [18] employed a multi-scale finite element damage percolation model to simulate stretch flange forming of AA 5182 and AA 5754. Butcher et al. [19] used a lower bound damage–based model for prediction of radial and circumferential cracks in stretch z-flanges of AA 5182. Simha et al. [20] utilized continuum mechanics based approach for prediction of radial and circumferential crack in AA 5182 stretch flanges on the basis of extended stress-based forming limit curve. Dewang et al. [21] reviewed the different failure criterions for prediction of failure in stretch flanging process. Dewang et al. [22] found that circumferential strain and edge cracking increases with increase in coefficient of friction, initial flange length, die profile radius while decreases with increase in punch-die clearance and punch profile radius in stretch flange forming. The objective of present work is to investigate the deformation behavior of AA 5052 of 0.5 mm thickness in stretch flanging process considering the effect of process parameters such as punch-die clearance, initial flange length and blank-holding force through FEM simulation and experimentation. The results of FE simulation are presented in terms of circumferential strain, radial strain, edge crack location and propagation and maximum punch load during stretch flanging. 2. Mechanical Properties of material The mechanical properties of aluminum alloy 5052 sheet was evaluated as per ASTM standard (ASTM E8/E8M– 11)[22].The tensile specimens are tested on a computerized universal testing machine INSTRON at a strain rate of 0.16667 s-1 along the rolling direction at room temperature. The following true stress-true strain curve is obtained as shown in Figure 1. Table 1 and table 2 show the tested chemical composition and mechanical properties of AA 5052 of thickness 0.5mm. Element Wt.%
Si 0.1224
Table 1. Chemical composition of AA 5052[22] Fe Mg Mn Cu Cr 0.1964 2.418 0.07688 0.03618 0.1688 Table 2. Mechanical Properties of AA 5052 [22] Mechanical Property Material AA 5052 Mass density
2680 kg/m3
Young’s Modulus
70.3 GPa
Poisson’s ratio
0.33
Fig.1. True stress-true strain curve for AA 5052 [22]
Ni 0.00398
Zn 0.04454
Al 95.70
Yogesh Dewang et al / Materials Today: Proceedings 4 (2017) 9316–9326
9319
3. Finite Element Modeling & Simulation In the current research work, FE simulation of stretch flanging process is carried out using FEM based software ABAQUS/Explicit [23]. Figure 2 shows the FE model of non-axisymmetric stretch flanging process. Rectangular blank of 120 mm × 25 mm × 0.5 mm size is used. Die, punch and blank-holder are considered as rigid body tools and are modelled with 4-noded bilinear quadrilateral discrete rigid 3D (R3D4) elements. The rectangular sheet metal blank is discretized by using 3D solid (C3D8R) elements. Isotropic hardening rule with ductile damage parameter is considered during FE simulation. Prediction of failure is demonstrated by defining ductile damage initiation criteria as a function of equivalent plastic strain. The initiation and propagation of edge cracks in the flange is defined using by damage evolution feature with element deletion technique. Crack initiation in the stretch flange is observed when ductile damage parameter (D) becomes unity. The whole process of simulation of stretch flanging process is carried out in number of small steps.
Fig.2. Finite element model of non-axisymmetric stretch flanging process [23]
4. Experimental work of stretch flanging process In order to validate the results of FE simulations, experiments are performed on a small punch testing machine equipped with a load cell of capacity 1 KN. The experimental tools and workpiece used in this study are shown in Figure 3. The experimental tool consists of punch, die and blank-holder. A rectangular blank of 120 mm × 25 mm × 0.5 mm size are used in this study as work piece material. Non-axisymmetric stretch flanging process is realized experimentally by clamping the work piece between blank-holder and die by using four bolts M10. A constant blank-holding force is maintained during flanging by tightening the bolts with help of mechanical torque wrench. A constant speed of 10 mm/min is given to the punch head during stretch flanging experiments. The results of punch displacements and punch load and of the load cell are continuously stored in a data acquisition system.
9320
Yogesh Dewang et al / Materials Today: Proceedings 4 (2017) 9316–9326
Fig.3. Experimental tool and work piece [23]
5. Results & Discussion The following section describes the effect of constant blank holding forces with low intensity on circumferential strain and radial strain. Besides this, other FE simulation results are also presented for radius of die of 20 mm for different process parameters in terms of the maximum circumferential strain and the maximum radial strain. The intensity of blank holding forces which is utilized lies in the range of 80 N to 160 N (in a step of 20 N). 5.1 Effect of low intensity constant blank holding forces on circumferential strain The circumferential strain is one the important parameters which affects the formability of stretch flange. Figure 4 shows that the circumferential strain distribution along die profile radius from end of the sheet up to sheet center. It is clear from Figure 6 that circumferential strain reaches the maximum value at corner end of flange, while it decreases gradually up to a position near to the corner edge and then becomes more or less constant up to sheet center at the lowest values of strain. A similar trend is found for all cases of constant blank holding forces. It is important to note that circumferential strain increases by nearly 64 % with increase in blank holding force from 80 N to 160 N at edge of stretch flange. The circumferential strain at the edge is found to be maximum for each blank holding force as well as circumferential strain at the corner of the stretch flange is increasing with increase in blank holding force because with increase in blank holding force, the blank will experience greater contact with interacting surfaces and as a result of which circumferential strain increases due to increment in coefficient of friction.
Yogesh Dewang et al / Materials Today: Proceedings 4 (2017) 9316–9326
9321
Fig.4 Effect of blank holding force on circumferential strain distribution
Figure 5 shows the different contour plots of circumferential strain developed in the stretch flange on application of low intensity constant blank holding forces. It is found that circumferential strain increases with increase in blank holding forces. The edge crack is absent in case of three blank holding forces of intensity 80 N, 100N and 120 N. On the other hand, edge cracks start to appear on application of blank holding force of 140 N and it further propagates more in case of blank holding force of 160 N. This indicates that blank holding force has a major role in forming of stretch flange which is governed by another vital parameter i.e., circumferential strain and excess of which will lead to edge cracks. The coefficient of friction also plays an important role as increase in blank holding force the interacting coefficient of friction increases which reduce the possibilities of slipping of blank and it also increases the chances of edge cracking with increment of circumferential strain too.
(a)
(b)
(c)
(d) (e) Fig.5 Contour plots for circumferential strain at blank holding force of (a) 80N (b) 100 N (c) 120 N (d) 140 N (e) 160 N
9322
Yogesh Dewang et al / Materials Today: Proceedings 4 (2017) 9316–9326
5.2 Effect of blank holding force on maximum circumferential strain (initial flange length-basis) Figure 6 shows the effect of blank holding force on the maximum circumferential strain at a punch- die clearance of 1mm for three different initial flange lengths. It is found that the maximum circumferential strain increases by nearly 17 % at the maximum blank holding force of 160 N with increase in initial flange length from 20 mm to 40 mm. The increment in circumferential strain with increase in initial flange length and blank holding force increases the degree of deformation and holding capacity of the blank between the blank-holder. In addition to this, the coefficient of friction plays a major role in increment of circumferential strain with increase in initial flange length and blank holding force.
Fig.6 Effect of blank holding force on the maximum circumferential strain (c = 1 mm)
5.3 Effect of blank holding force on maximum circumferential strain (punch-die clearance base) Circumferential strain increases non- linearly with increase in blank holding force, as shown in Figure 7, for an initial flange length of 30 mm and a punch die clearance from 1 mm to 3 mm. The maximum circumferential strain decreases by nearly 22% with increase in punch die clearance from 1 to 3 mm at a peak blank holding force of 160 N. The increment in punch-die clearance decreases the coefficient of friction, while increment in blank holding increases the coefficient of friction among the interacting surfaces. Therefore the highest circumferential strain is obtained for highest blank holding force and least punch-die clearance, which also signifies the role of coefficient of friction in formation of stretch flange.
Fig.7 Effect of blank holding force on the maximum circumferential strain (L = 30 mm)
5.4 Effect of blank holding force on radial strain (initial flange length base) Figure 8 shows that the maximum radial strain increases with increase in blank holding force from 80 N to 160 N, at a clearance of 1 mm for the three different initial flange lengths. It is found that the maximum radial strain for initial flange length of 40 mm becomes nearly double that of the maximum radial strain for initial flange length of 20 mm, at a peak blank holding force of 160 N. The increment in radial strain is occurring in the stretch flange because the higher will be the initial flange length and blank holding force the greater will be the degree of deformation, which increases the coefficient of friction and as a result of which the radial strain increases.
Yogesh Dewang et al / Materials Today: Proceedings 4 (2017) 9316–9326
9323
Fig.8 Effect of blank holding force on radial strain (c = 1mm)
5.5 Effect of blank holding force on radial strain (punch-die clearance base) Figure 9 shows the variation of maximum radial strain with blank holding force for five different punch-die clearances at an initial flange length of 30 mm and width of sheet of 20 mm at die radius of 20 mm. It is found that the maximum radial strain increases nonlinearly with increase in blank holding force for all the five punch-die clearances. It is also found that the maximum radial strain is obtained for the minimum punch die clearance while the minimum radial strain is obtained for the greatest punch die clearance. Besides this, it is found that there is a decrease in the maximum radial strain by nearly 23% when punch die clearance varies from 1.5 mm to 3 mm. The maximum radial strain at punch die clearance of 3 mm becomes nearly 0.14 times to that obtained for 1 mm punchdie clearance. The increment in radial strain occurring in the stretch flange is occurring due to the increment in blank holding force and in decrement of punch die clearance. The higher will be punch-die clearance the blank will experience lesser coefficient of friction. Therefore the maximum radial strain is obtained for least punch-die clearance and highest blank holding force.
Fig.9 Effect of blank holding force on maximum radial strain (L=30 mm)
Increment in maximum radial strain occurring in the stretch flange with increase in blank holding force and different initial flange lengths is basically due to the increment in greater degree of deformation. Both blank holding force increases the possibilities of greater contact and again coefficient of friction plays important role in occurrence of greater radial strain in the flange. Besides, this the maximum radial strain also increases with increase in initial flange length. The highest radial strain is obtained for the highest blank holding force, highest initial flange and least punch-die clearance.
9324
Yogesh Dewang et al / Materials Today: Proceedings 4 (2017) 9316–9326
5.6 Effect of blank-holding force on punch load and forming kinematics In this section, five different constant blank holding forces are applied on the blank holder. It is clear from Figure 10 that for all the five cases of constant blank holding forces, the punch load increases in order to start bending the sheet from unbent position and reaches to a maximum value. Further the punch load decreases due to crack initiation and crack length increases and finally punch load decreases to attain a constant lowest value. It is important to note that punch load increases by nearly 14 % with increase in blank holding force from 80 N to 160 N. The maximum punch load is evaluated through numerical simulation of stretch flanging process under the influence of blank holding forces in combination with initial flange length and punch die clearance. The reason for increment in punch load with increase in blank holding force is due to greater contact of blank with the interacting surfaces and blank holding force will keep the sheet in contact for greater deformation of blank. Again coefficient of friction plays an important role in forming of flange as with increment in blank holding force, the chance of slipping of blank also decreases. Therefore higher will be the blank holding force, higher punch load will be required for formation of flange through punch over a given range of punch displacement.
Fig.10 Punch load versus punch displacement for constant blank holding forces
6. Validation of results 6.1 Comparison of edge crack location and propagation (punch-die clearance basis) Fig.11 shows the comparison of edge crack location and propagation between FE simulation and experimentation on increment in punch-die clearance. It is found that edge cracking decreases with increase in punch-die clearance. This is occurring due to the fact that the circumferential strain decreases with increase in punch-die clearance which in turn actually decreases the edge cracking in the stretch flange. A very good agreement is found between experimental results and FE simulation results in terms of edge crack location and propagation.
(a) c = 1 mm (b) c = 2 mm (c) c = 3 mm (d) c = 1 mm (e) c = 2 mm (f) c = 3 mm Fig.11. Comparison of edge crack location for different punch-die clearance between experimental (a-c) and FEM simulation (d-f)
Yogesh Dewang et al / Materials Today: Proceedings 4 (2017) 9316–9326
9325
6.2 Comparison of edge crack location and propagation (initial flange length basis) Fig.12 shows the comparison of edge crack location and propagation between FE simulation and experimentation on increment in punch-die clearance. It is found that edge cracking increases with increase in initial flange length. This is occurring due to the fact that the circumferential strain increases with increase in initial flange length as greater flange length is available for deformation in flange, which in turn actually increases the edge cracking in the stretch flange. A very good agreement is found between experimental results and FE simulation results in terms of edge crack location and propagation.
(a) FEM (L= 40 mm) (b) Experimental (L= 40 mm) (c) FEM (L= 40 mm) (d) Experimental (L= 40 mm) Fig.12 (a-d) Comparison of edge crack location for different initial flange length between FE simulation and experiments
6.3 Comparison of edge crack location and propagation (blank holding force basis) Another case used for validation is shown in Figure.13 in which different blank holding forces are applied by introducing sheet metal blanks for providing gaps between the work-piece and blank holder in the range of D = 1 mm to 3 mm (D = 1mm, 2mm ,3 mm). It is found, both from simulation and experimentation that crack length decreases with increase in gap between blank holder and sheet metal blank because of decrease in circumferential strain. It is observed that for D = 1 mm edge crack appears both in simulation and experiments. It is also pertinent to point out that no edge crack appears on increasing the gap (D) further from 1.5 mm to 3 mm, both in simulation and experiments. It is noticed that more and more material accumulates along die profile radius on increasing the gap between sheet metal blank and blank holder which can be seen both in simulation and experimental results as bulging of sheet metal along die profile radius due to decrement in blank holding force. Bulging is found to increase with increase in gap between sheet metal blank and blank holder due to the decrease in blank holding force on sheet metal blank. Very good agreement is again obtained between the results obtained through FE simulation and experiments.
(a) D = 1 mm (b) D = 2 mm (c) D = 3 mm (d) D = 1 mm (e) D = 2 mm (f) D = 3 mm Fig. 13 Comparison of edge crack location and propagation for different blank holding forces between FE simulation (a-c) and experimental (d-f)
9326
Yogesh Dewang et al / Materials Today: Proceedings 4 (2017) 9316–9326
7. Conclusions It can be concluded from this study of stretch flanging process that blank holding force, punch die-clearance and initial flange length are found to be important process parameters which greatly influence strain distribution and edge crack. It is found that edge crack and circumferential strain increases with increase in blank holding force and initial flange length while it decreases with punch-die clearance. Therefore, FEM simulation is quite helpful in better designing of stretch flanging process to obtained defect free products. References [1] Brunet M, Sabourin F, Touchal S M. The prediction of necking and failure in 3D. sheet forming analysis using damage variable. J De Physique 1996; 6:473-482. [2] Teixeira P, Santos A D, Pires FMA, Cesar JMA. Finite element prediction of ductile fracture in sheet metal forming processes. J Mater Proc Technol 2006; 177: 278-281. [3] Li Y,Luo M ,Gerlach J ,Wierzbicki T. Prediction of shear-induced fracture in sheet metal forming. J Mater Proc Technol 2010; 210:18581869. [4] Wu J,Ma Z ,Zhou Yichun ,Lu C.Prediction of failure modes during deep drawing of metal sheets with nickel coating. J Mater Sci Technol 2013; 29:1059-1066. [5] Nguyen D T , Dinh D K, Nguyen H M T, Banh T L ,Kim Y S. Formability improvement and blank Shape definition for deep drawing of cylindrical cup with complex curve profile from spcc sheets using FEM. J Central South University 2014;21:27-34. [6] Kacem A, Krichen A ,Manach P Y, Thuillier S, Yoon J W. Failure prediction in the hole-flanging process of aluminum alloys, Eng Frac Mech 2013;99:251-265. [7] Wang X, Cao J, Li M .Wrinkling analysis in shrink flanging. J Manuf Sci Eng 2001; 123:426-432. [8] Hu P,Li DY,Li Y X.Analytical models of stretch and shrink flanging. Int J of Mach Tools & Manuf (2003); 43:1367–1373. [9] Dudra S, Shah S.Stretch flanges: formability and trimline development. J. Mater Shap Technol 1998; 6(2): 91-101. [10] Bao Y D, Huh H. Optimum design of trimming line by one-step analysis for auto body parts. J Mater Proc Technol 2007; 187-188:108–112. [11] Li D, Luo Y, Peng Y, Hu P. The numerical and analytical study on stretch flanging of V-shaped sheet metal.J Mater Proc Technol 2007; 189: 262–267. [12] Wang N M, Johnson L K, Tang S C .Stretch flanging of “V”-shaped sheet metal blanks. J Appl Metal Working 1984;3: 281-291. [13] Yeh F H, Wu M T., Li CL. Accurate optimization of blank design in stretch flange based on a forward–inverse prediction scheme. Int J Mach Tools & Manuf 2007; 47:1854–1863. [14] Wang C T, Kinzel G, Altan T. Failure and wrinkling criteria and mathematical modeling of shrink and stretch flanging operations in sheet– metal forming. J Mater Proc Technol 1995; 53:759-780. [15] Asnafi N.On Stretch and shrink flanging of sheet aluminium by fluid Forming. J Mater Proc Technol 1999:96:198-214. [16] Worswick M J,Finn M J. The numerical simulation of stretch flange forming. Int J Plast 2000; 16:701-720. [17] Feng X,Zhongqin L ,Shuhui L ,Weili X. Study on the influences of geometrical parameters on the formability of stretch curved flanging by numerical simulation. J Mater Proc Technol 2004; 145:93-98. [18] Chen Z, Worswick M J, Pilkey A K, Lloyd D J. Damage percolation during stretch flange forming of aluminum alloy sheet. J Mech Phy Solids 2005; 53:2692-2717. [19] Butcher C, Chen Z, Worswick W. A lower bound damage-based finite element simulation of stretch flange forming of Al-Mg alloys. Int J Frac 2006; 142:289-298. [20] Simha CHM,Grantab R ,Worswick M J. Application of an extended stress-based forming limit curve to predict necking in stretch flange forming. J Manuf Sci Eng 2008; 130: 1-11. [21] Dewang Y, Hora MS and Panthi SK. Review on finite element analysis of sheet metal stretch flanging process. ARPN J Eng and Appl Sci 2014: 9(9): 1565-1579. [22] Dewang Y, Hora M S, Panthi S K. Prediction of crack location and propagation in stretch flanging process of aluminum alloy AA-5052 sheet using FEM simulation. T. Nonferr. Metal Soc 2015; 25 (7):2308-2320.
ScienceDirect Materials Today: Proceedings 4 (2017) 9316–9326
www.materialstoday.com/proceedings
ICAAMM-2016
Effect of process parameters on deformation behavior of AA 5052 sheets in stretch flanging process Yogesh Dewanga*, M.S. Horab, S.K. Panthic a
Assistant Professor, Department of Mechanical Engineering,Lakshmi Narain College of Technology, Bhopal,462021, India b Professor, Department of Civil Engineering,Maulana Azad National Institute of Technology, Bhopal, 462003, India c Scientist,Computer Simulation & Design Centre,Advanced Materials Processes and Research Institute,Bhopal,462024, India
Abstract Aluminum alloys are widely used in manufacturing of aircrafts and in automotive industry. In this paper aluminum alloy 5052 of 0.5 mm thickness was utilized to investigate deformation behavior in stretch flanging process. The effect of process parameters such as initial flange length, punch-die clearance and blank-holding force were investigated using FE simulation and experimentation. The results show that the circumferential strain and radial strain increases with increase in initial flange length and blank-holding force and decreases with increase in punch-die clearance. Experimental investigations were also carried out to validate FE simulation results. A very good agreement was obtained between FE simulation results and experimental in terms of punch load and edge crack location and propagation. © 2017 Elsevier Ltd. All rights reserved. Selection and Peer-review under responsibility of the Committee Members of International Conference on Advancements in Aeromechanical Materials for Manufacturing (ICAAMM-2016). Keywords: stretch flanging; finite element simulation; punch-die clearance; initial flange length; blank-holding force; edge crack.
* Corresponding author. Tel.:+91-0755-2457105; fax: +91-0755-2457042. E-mail address:[email protected]
2214-7853© 2017 Elsevier Ltd. All rights reserved. Selection and Peer-review under responsibility ofthe Committee Members of International Conference on Advancements in Aeromechanical Materials for Manufacturing (ICAAMM-2016).
Yogesh Dewang et al / Materials Today: Proceedings 4 (2017) 9316–9326
9317
1. Introduction Flanging is one of the important sheet metal forming processes which is applied in automobile industry. Flanging is a sheet metal bending operation in which sheet is bent usually to 900 for providing a smooth rounded edge, higher or strength to the edge of sheet-metal parts. It is also used for making hidden joints and assembly of automobile parts. Flanging can be classified on the basis of deformation features into two major types as shrink and stretch flanging. In shrink flanging, the curvature of flange is convex and metal is in compression in circumferential direction. The arc length of final flange is smaller than arc length of original element. The compression is greatest at the top of flange and least at die profile radius. In stretch flanging, the curvature of flange is opposite to shrink flanging, i.e. concave and sheet metal undergoes tension in circumferential direction. The tension is greatest at top of flange resulting in increase in the length of arc. The feasibility of any sheet metal forming process depends upon various processes, geometric as well as material parameters. To obtain defects free products different parameters need to be optimized. Different types of failure like necking, wrinkling and fracture are involved in sheet metal forming processes. With the growth and development of commercially available FEM software it is become easier to predict these types of defects. Finite element simulation is an accurate and faster method for prediction of failure in sheet metal forming processes because it captures realistic phenomena with less cost and time involved. Researchers in past carried out work in the area of the failure analysis in sheet metal forming processes using finite element simulation. Brunet et al. [1] used a gurson-tvergaard ductile damage based model for prediction of necking and fracture for anisotropic sheet metals. Teixeira et al. [2] coupled Hill’s orthotropic plasticity criterion with Lemaitre’s ductile damage evolution law for prediction of ductile fracture in sheet metal forming processes. In deep drawing process, usually two types of failure modes namely fracture and wrinkling were encountered. Li et al. [3] applied modified Mohr-Coulomb fracture criterion in the deep drawing of AHSS using both circular and square punch for prediction of shear induced fracture in terms of crack initiation, propagation and location of cracks. Wu et al. [4] found that that low blank-holding force will give rise to wrinkling while increase in blank-holding force will lead to fracture during deep drawing of sheet metal with nickel. Nguyen et al. [5] predicted wrinkling and fracture in deep drawing of SPCC material by using forming limit diagram ductile fracture criterion. On the other hand, in case of flanging process the major type of defects were necking, fracture and wrinkling. The major types of flanging are straight, stretch, shrink and hole-flanging in which studies were carried out. Kacem et al. [6] predicted necking, damage, fracture and limit of flanging using a macroscopic ductile fracture criterion by taking into account the triaxiality of material in hole-flanging of AA 1050-H-14 and AA-6061.In another type of flanging process i.e., shrink flanging wrinkling was observed as failure. Wang et al. [7] predicted wrinkling by using energy approach based on stress-based criterion. Some studies were also carried out in the domain of flanging process in order to study the effect of various parameters on formability of flange using various theoretical, experimental and FEM based approaches. Hu et al. [8] proposed analytical models for introfelxion/stretch and outcurve/shrink flange which takes into account the effect of planar anisotropy in flanging process for prediction of blank’s trim line. Dudra and Shah [9] developed analytical model for development of trimline for axisymmetric case and utilized FEM method for non-axisymmetric case for prediction of trimline and peak strain. Bao and Huh [10] implemented one-step analysis in trimming die design for effective design procedure. Li et al. [11] used an equivalent 2D FEM program for parametric study in stretch flanging of V-shaped blanks. Wang et al. [12] utilized an elastic-plastic FEM program in order to study the effect of various materials and process parameters on the formability of flange in stretch flanging of V-shaped blanks. Yeh et al. [13] determined optimum blank shape by using forward-inverse prediction scheme which integrates explicit dynamic FEM, true strain method and adaptive-network-based fuzzy inference system in stretch flanging. In stretch flanging, the predominant mode of failure was along edge as fracture and localized necking. Wang et al. [14] proposed localized necking for stretch flanging of AKDQ steels based on modification of Hill’s instability criterion by taking into account the effect of strain-hardening and plastic anisotropy. Asnafi [15] found that fracture limit was obtained by increasing plastic strain ratio, strain hardening exponent and uniform strain in stretch flanging through fluid forming of aluminum alloys.Worswick and Finn [16] found through numerical simulation that strain predictions were more accurate with Barlat-89 and Hill-48 yield criteria than von-Mises yield criteria in stretch flange forming. Feng et al. [17] recommended small flange height and small ratio of straight side length to the curved range radius in stretch curved flanging in order to avoid fracture at edge of flange.
9318
Yogesh Dewang et al / Materials Today: Proceedings 4 (2017) 9316–9326
In finite element simulation of stretch flange forming the concept of continuum mechanics was employed by various researchers for prediction of failure prediction of failure. Chen et al. [18] employed a multi-scale finite element damage percolation model to simulate stretch flange forming of AA 5182 and AA 5754. Butcher et al. [19] used a lower bound damage–based model for prediction of radial and circumferential cracks in stretch z-flanges of AA 5182. Simha et al. [20] utilized continuum mechanics based approach for prediction of radial and circumferential crack in AA 5182 stretch flanges on the basis of extended stress-based forming limit curve. Dewang et al. [21] reviewed the different failure criterions for prediction of failure in stretch flanging process. Dewang et al. [22] found that circumferential strain and edge cracking increases with increase in coefficient of friction, initial flange length, die profile radius while decreases with increase in punch-die clearance and punch profile radius in stretch flange forming. The objective of present work is to investigate the deformation behavior of AA 5052 of 0.5 mm thickness in stretch flanging process considering the effect of process parameters such as punch-die clearance, initial flange length and blank-holding force through FEM simulation and experimentation. The results of FE simulation are presented in terms of circumferential strain, radial strain, edge crack location and propagation and maximum punch load during stretch flanging. 2. Mechanical Properties of material The mechanical properties of aluminum alloy 5052 sheet was evaluated as per ASTM standard (ASTM E8/E8M– 11)[22].The tensile specimens are tested on a computerized universal testing machine INSTRON at a strain rate of 0.16667 s-1 along the rolling direction at room temperature. The following true stress-true strain curve is obtained as shown in Figure 1. Table 1 and table 2 show the tested chemical composition and mechanical properties of AA 5052 of thickness 0.5mm. Element Wt.%
Si 0.1224
Table 1. Chemical composition of AA 5052[22] Fe Mg Mn Cu Cr 0.1964 2.418 0.07688 0.03618 0.1688 Table 2. Mechanical Properties of AA 5052 [22] Mechanical Property Material AA 5052 Mass density
2680 kg/m3
Young’s Modulus
70.3 GPa
Poisson’s ratio
0.33
Fig.1. True stress-true strain curve for AA 5052 [22]
Ni 0.00398
Zn 0.04454
Al 95.70
Yogesh Dewang et al / Materials Today: Proceedings 4 (2017) 9316–9326
9319
3. Finite Element Modeling & Simulation In the current research work, FE simulation of stretch flanging process is carried out using FEM based software ABAQUS/Explicit [23]. Figure 2 shows the FE model of non-axisymmetric stretch flanging process. Rectangular blank of 120 mm × 25 mm × 0.5 mm size is used. Die, punch and blank-holder are considered as rigid body tools and are modelled with 4-noded bilinear quadrilateral discrete rigid 3D (R3D4) elements. The rectangular sheet metal blank is discretized by using 3D solid (C3D8R) elements. Isotropic hardening rule with ductile damage parameter is considered during FE simulation. Prediction of failure is demonstrated by defining ductile damage initiation criteria as a function of equivalent plastic strain. The initiation and propagation of edge cracks in the flange is defined using by damage evolution feature with element deletion technique. Crack initiation in the stretch flange is observed when ductile damage parameter (D) becomes unity. The whole process of simulation of stretch flanging process is carried out in number of small steps.
Fig.2. Finite element model of non-axisymmetric stretch flanging process [23]
4. Experimental work of stretch flanging process In order to validate the results of FE simulations, experiments are performed on a small punch testing machine equipped with a load cell of capacity 1 KN. The experimental tools and workpiece used in this study are shown in Figure 3. The experimental tool consists of punch, die and blank-holder. A rectangular blank of 120 mm × 25 mm × 0.5 mm size are used in this study as work piece material. Non-axisymmetric stretch flanging process is realized experimentally by clamping the work piece between blank-holder and die by using four bolts M10. A constant blank-holding force is maintained during flanging by tightening the bolts with help of mechanical torque wrench. A constant speed of 10 mm/min is given to the punch head during stretch flanging experiments. The results of punch displacements and punch load and of the load cell are continuously stored in a data acquisition system.
9320
Yogesh Dewang et al / Materials Today: Proceedings 4 (2017) 9316–9326
Fig.3. Experimental tool and work piece [23]
5. Results & Discussion The following section describes the effect of constant blank holding forces with low intensity on circumferential strain and radial strain. Besides this, other FE simulation results are also presented for radius of die of 20 mm for different process parameters in terms of the maximum circumferential strain and the maximum radial strain. The intensity of blank holding forces which is utilized lies in the range of 80 N to 160 N (in a step of 20 N). 5.1 Effect of low intensity constant blank holding forces on circumferential strain The circumferential strain is one the important parameters which affects the formability of stretch flange. Figure 4 shows that the circumferential strain distribution along die profile radius from end of the sheet up to sheet center. It is clear from Figure 6 that circumferential strain reaches the maximum value at corner end of flange, while it decreases gradually up to a position near to the corner edge and then becomes more or less constant up to sheet center at the lowest values of strain. A similar trend is found for all cases of constant blank holding forces. It is important to note that circumferential strain increases by nearly 64 % with increase in blank holding force from 80 N to 160 N at edge of stretch flange. The circumferential strain at the edge is found to be maximum for each blank holding force as well as circumferential strain at the corner of the stretch flange is increasing with increase in blank holding force because with increase in blank holding force, the blank will experience greater contact with interacting surfaces and as a result of which circumferential strain increases due to increment in coefficient of friction.
Yogesh Dewang et al / Materials Today: Proceedings 4 (2017) 9316–9326
9321
Fig.4 Effect of blank holding force on circumferential strain distribution
Figure 5 shows the different contour plots of circumferential strain developed in the stretch flange on application of low intensity constant blank holding forces. It is found that circumferential strain increases with increase in blank holding forces. The edge crack is absent in case of three blank holding forces of intensity 80 N, 100N and 120 N. On the other hand, edge cracks start to appear on application of blank holding force of 140 N and it further propagates more in case of blank holding force of 160 N. This indicates that blank holding force has a major role in forming of stretch flange which is governed by another vital parameter i.e., circumferential strain and excess of which will lead to edge cracks. The coefficient of friction also plays an important role as increase in blank holding force the interacting coefficient of friction increases which reduce the possibilities of slipping of blank and it also increases the chances of edge cracking with increment of circumferential strain too.
(a)
(b)
(c)
(d) (e) Fig.5 Contour plots for circumferential strain at blank holding force of (a) 80N (b) 100 N (c) 120 N (d) 140 N (e) 160 N
9322
Yogesh Dewang et al / Materials Today: Proceedings 4 (2017) 9316–9326
5.2 Effect of blank holding force on maximum circumferential strain (initial flange length-basis) Figure 6 shows the effect of blank holding force on the maximum circumferential strain at a punch- die clearance of 1mm for three different initial flange lengths. It is found that the maximum circumferential strain increases by nearly 17 % at the maximum blank holding force of 160 N with increase in initial flange length from 20 mm to 40 mm. The increment in circumferential strain with increase in initial flange length and blank holding force increases the degree of deformation and holding capacity of the blank between the blank-holder. In addition to this, the coefficient of friction plays a major role in increment of circumferential strain with increase in initial flange length and blank holding force.
Fig.6 Effect of blank holding force on the maximum circumferential strain (c = 1 mm)
5.3 Effect of blank holding force on maximum circumferential strain (punch-die clearance base) Circumferential strain increases non- linearly with increase in blank holding force, as shown in Figure 7, for an initial flange length of 30 mm and a punch die clearance from 1 mm to 3 mm. The maximum circumferential strain decreases by nearly 22% with increase in punch die clearance from 1 to 3 mm at a peak blank holding force of 160 N. The increment in punch-die clearance decreases the coefficient of friction, while increment in blank holding increases the coefficient of friction among the interacting surfaces. Therefore the highest circumferential strain is obtained for highest blank holding force and least punch-die clearance, which also signifies the role of coefficient of friction in formation of stretch flange.
Fig.7 Effect of blank holding force on the maximum circumferential strain (L = 30 mm)
5.4 Effect of blank holding force on radial strain (initial flange length base) Figure 8 shows that the maximum radial strain increases with increase in blank holding force from 80 N to 160 N, at a clearance of 1 mm for the three different initial flange lengths. It is found that the maximum radial strain for initial flange length of 40 mm becomes nearly double that of the maximum radial strain for initial flange length of 20 mm, at a peak blank holding force of 160 N. The increment in radial strain is occurring in the stretch flange because the higher will be the initial flange length and blank holding force the greater will be the degree of deformation, which increases the coefficient of friction and as a result of which the radial strain increases.
Yogesh Dewang et al / Materials Today: Proceedings 4 (2017) 9316–9326
9323
Fig.8 Effect of blank holding force on radial strain (c = 1mm)
5.5 Effect of blank holding force on radial strain (punch-die clearance base) Figure 9 shows the variation of maximum radial strain with blank holding force for five different punch-die clearances at an initial flange length of 30 mm and width of sheet of 20 mm at die radius of 20 mm. It is found that the maximum radial strain increases nonlinearly with increase in blank holding force for all the five punch-die clearances. It is also found that the maximum radial strain is obtained for the minimum punch die clearance while the minimum radial strain is obtained for the greatest punch die clearance. Besides this, it is found that there is a decrease in the maximum radial strain by nearly 23% when punch die clearance varies from 1.5 mm to 3 mm. The maximum radial strain at punch die clearance of 3 mm becomes nearly 0.14 times to that obtained for 1 mm punchdie clearance. The increment in radial strain occurring in the stretch flange is occurring due to the increment in blank holding force and in decrement of punch die clearance. The higher will be punch-die clearance the blank will experience lesser coefficient of friction. Therefore the maximum radial strain is obtained for least punch-die clearance and highest blank holding force.
Fig.9 Effect of blank holding force on maximum radial strain (L=30 mm)
Increment in maximum radial strain occurring in the stretch flange with increase in blank holding force and different initial flange lengths is basically due to the increment in greater degree of deformation. Both blank holding force increases the possibilities of greater contact and again coefficient of friction plays important role in occurrence of greater radial strain in the flange. Besides, this the maximum radial strain also increases with increase in initial flange length. The highest radial strain is obtained for the highest blank holding force, highest initial flange and least punch-die clearance.
9324
Yogesh Dewang et al / Materials Today: Proceedings 4 (2017) 9316–9326
5.6 Effect of blank-holding force on punch load and forming kinematics In this section, five different constant blank holding forces are applied on the blank holder. It is clear from Figure 10 that for all the five cases of constant blank holding forces, the punch load increases in order to start bending the sheet from unbent position and reaches to a maximum value. Further the punch load decreases due to crack initiation and crack length increases and finally punch load decreases to attain a constant lowest value. It is important to note that punch load increases by nearly 14 % with increase in blank holding force from 80 N to 160 N. The maximum punch load is evaluated through numerical simulation of stretch flanging process under the influence of blank holding forces in combination with initial flange length and punch die clearance. The reason for increment in punch load with increase in blank holding force is due to greater contact of blank with the interacting surfaces and blank holding force will keep the sheet in contact for greater deformation of blank. Again coefficient of friction plays an important role in forming of flange as with increment in blank holding force, the chance of slipping of blank also decreases. Therefore higher will be the blank holding force, higher punch load will be required for formation of flange through punch over a given range of punch displacement.
Fig.10 Punch load versus punch displacement for constant blank holding forces
6. Validation of results 6.1 Comparison of edge crack location and propagation (punch-die clearance basis) Fig.11 shows the comparison of edge crack location and propagation between FE simulation and experimentation on increment in punch-die clearance. It is found that edge cracking decreases with increase in punch-die clearance. This is occurring due to the fact that the circumferential strain decreases with increase in punch-die clearance which in turn actually decreases the edge cracking in the stretch flange. A very good agreement is found between experimental results and FE simulation results in terms of edge crack location and propagation.
(a) c = 1 mm (b) c = 2 mm (c) c = 3 mm (d) c = 1 mm (e) c = 2 mm (f) c = 3 mm Fig.11. Comparison of edge crack location for different punch-die clearance between experimental (a-c) and FEM simulation (d-f)
Yogesh Dewang et al / Materials Today: Proceedings 4 (2017) 9316–9326
9325
6.2 Comparison of edge crack location and propagation (initial flange length basis) Fig.12 shows the comparison of edge crack location and propagation between FE simulation and experimentation on increment in punch-die clearance. It is found that edge cracking increases with increase in initial flange length. This is occurring due to the fact that the circumferential strain increases with increase in initial flange length as greater flange length is available for deformation in flange, which in turn actually increases the edge cracking in the stretch flange. A very good agreement is found between experimental results and FE simulation results in terms of edge crack location and propagation.
(a) FEM (L= 40 mm) (b) Experimental (L= 40 mm) (c) FEM (L= 40 mm) (d) Experimental (L= 40 mm) Fig.12 (a-d) Comparison of edge crack location for different initial flange length between FE simulation and experiments
6.3 Comparison of edge crack location and propagation (blank holding force basis) Another case used for validation is shown in Figure.13 in which different blank holding forces are applied by introducing sheet metal blanks for providing gaps between the work-piece and blank holder in the range of D = 1 mm to 3 mm (D = 1mm, 2mm ,3 mm). It is found, both from simulation and experimentation that crack length decreases with increase in gap between blank holder and sheet metal blank because of decrease in circumferential strain. It is observed that for D = 1 mm edge crack appears both in simulation and experiments. It is also pertinent to point out that no edge crack appears on increasing the gap (D) further from 1.5 mm to 3 mm, both in simulation and experiments. It is noticed that more and more material accumulates along die profile radius on increasing the gap between sheet metal blank and blank holder which can be seen both in simulation and experimental results as bulging of sheet metal along die profile radius due to decrement in blank holding force. Bulging is found to increase with increase in gap between sheet metal blank and blank holder due to the decrease in blank holding force on sheet metal blank. Very good agreement is again obtained between the results obtained through FE simulation and experiments.
(a) D = 1 mm (b) D = 2 mm (c) D = 3 mm (d) D = 1 mm (e) D = 2 mm (f) D = 3 mm Fig. 13 Comparison of edge crack location and propagation for different blank holding forces between FE simulation (a-c) and experimental (d-f)
9326
Yogesh Dewang et al / Materials Today: Proceedings 4 (2017) 9316–9326
7. Conclusions It can be concluded from this study of stretch flanging process that blank holding force, punch die-clearance and initial flange length are found to be important process parameters which greatly influence strain distribution and edge crack. It is found that edge crack and circumferential strain increases with increase in blank holding force and initial flange length while it decreases with punch-die clearance. Therefore, FEM simulation is quite helpful in better designing of stretch flanging process to obtained defect free products. References [1] Brunet M, Sabourin F, Touchal S M. The prediction of necking and failure in 3D. sheet forming analysis using damage variable. J De Physique 1996; 6:473-482. [2] Teixeira P, Santos A D, Pires FMA, Cesar JMA. Finite element prediction of ductile fracture in sheet metal forming processes. J Mater Proc Technol 2006; 177: 278-281. [3] Li Y,Luo M ,Gerlach J ,Wierzbicki T. Prediction of shear-induced fracture in sheet metal forming. J Mater Proc Technol 2010; 210:18581869. [4] Wu J,Ma Z ,Zhou Yichun ,Lu C.Prediction of failure modes during deep drawing of metal sheets with nickel coating. J Mater Sci Technol 2013; 29:1059-1066. [5] Nguyen D T , Dinh D K, Nguyen H M T, Banh T L ,Kim Y S. Formability improvement and blank Shape definition for deep drawing of cylindrical cup with complex curve profile from spcc sheets using FEM. J Central South University 2014;21:27-34. [6] Kacem A, Krichen A ,Manach P Y, Thuillier S, Yoon J W. Failure prediction in the hole-flanging process of aluminum alloys, Eng Frac Mech 2013;99:251-265. [7] Wang X, Cao J, Li M .Wrinkling analysis in shrink flanging. J Manuf Sci Eng 2001; 123:426-432. [8] Hu P,Li DY,Li Y X.Analytical models of stretch and shrink flanging. Int J of Mach Tools & Manuf (2003); 43:1367–1373. [9] Dudra S, Shah S.Stretch flanges: formability and trimline development. J. Mater Shap Technol 1998; 6(2): 91-101. [10] Bao Y D, Huh H. Optimum design of trimming line by one-step analysis for auto body parts. J Mater Proc Technol 2007; 187-188:108–112. [11] Li D, Luo Y, Peng Y, Hu P. The numerical and analytical study on stretch flanging of V-shaped sheet metal.J Mater Proc Technol 2007; 189: 262–267. [12] Wang N M, Johnson L K, Tang S C .Stretch flanging of “V”-shaped sheet metal blanks. J Appl Metal Working 1984;3: 281-291. [13] Yeh F H, Wu M T., Li CL. Accurate optimization of blank design in stretch flange based on a forward–inverse prediction scheme. Int J Mach Tools & Manuf 2007; 47:1854–1863. [14] Wang C T, Kinzel G, Altan T. Failure and wrinkling criteria and mathematical modeling of shrink and stretch flanging operations in sheet– metal forming. J Mater Proc Technol 1995; 53:759-780. [15] Asnafi N.On Stretch and shrink flanging of sheet aluminium by fluid Forming. J Mater Proc Technol 1999:96:198-214. [16] Worswick M J,Finn M J. The numerical simulation of stretch flange forming. Int J Plast 2000; 16:701-720. [17] Feng X,Zhongqin L ,Shuhui L ,Weili X. Study on the influences of geometrical parameters on the formability of stretch curved flanging by numerical simulation. J Mater Proc Technol 2004; 145:93-98. [18] Chen Z, Worswick M J, Pilkey A K, Lloyd D J. Damage percolation during stretch flange forming of aluminum alloy sheet. J Mech Phy Solids 2005; 53:2692-2717. [19] Butcher C, Chen Z, Worswick W. A lower bound damage-based finite element simulation of stretch flange forming of Al-Mg alloys. Int J Frac 2006; 142:289-298. [20] Simha CHM,Grantab R ,Worswick M J. Application of an extended stress-based forming limit curve to predict necking in stretch flange forming. J Manuf Sci Eng 2008; 130: 1-11. [21] Dewang Y, Hora MS and Panthi SK. Review on finite element analysis of sheet metal stretch flanging process. ARPN J Eng and Appl Sci 2014: 9(9): 1565-1579. [22] Dewang Y, Hora M S, Panthi S K. Prediction of crack location and propagation in stretch flanging process of aluminum alloy AA-5052 sheet using FEM simulation. T. Nonferr. Metal Soc 2015; 25 (7):2308-2320.