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Prediction of process conditions in drawing and ironing of cans
Journal of
ELSEVIER
Journal of Materials Processing Technology 59 (1996) 1-9
Materials Processing Technology
Prediction of process conditions in drawing and ironing of cans M a r c o S c h f . i n e m a n n , M u s t a f a A. A h m e t o g l u * , a n d T a y l a n A l t a n
ERC for Net Shape Manufacturing, Ohio State University, Columbus, Ohio 43210, USA
Industrial S u m m a r y
Deep drawing and ironing are the major processes used today in manufacturing of beverage cans from aluminum. The same technology is utilized in manufacturing of steel cans for the food industry, photosensitive drums and pressure vessels for fire extinguishers. Most of these processes require multi-stage ironing following the deep drawing and redrawing or backward extrusion processes. The practical aspects of this technology are well k n o w n and gained through extensive experimentation and production know-how. H o w e v e r , the fundamental aspects of these processes are relatively less known, especially regarding the temperature developed during deformation and the effect of deformation speed upon temperatures and lubrication. Thus, it is expected that process simulations using FEM techniques would provide additional detailed information that could be utilized to improve the process conditions. This paper illustrates the application of process modeling to deep drawing and ironing operations. A commercially available finite element code DEFORM was used to simulate deep drawing, redrawing and multi-stage ironing operations with varying parameters to illustrate the interactions between process variables.
1. I n t r o d u c t i o n
Ironing, together with drawing and redrawing, is the major process used in can manufacturing. The principle of ironing with a single die and a cylindrical punch is illustrated in Fig. 1. The moving punch, or ram, draws the cup through one or more ironing rings. The thickness of the cup bottom remains almost unchanged, while the thickness of the wall is reduced. Die
Fig. 1: Single ironing with a cylindrical punch [1]
* Corresponding author 0924-0136/96/$15.00 © 1996 Elsevier Science S.A~All rights reserved PI10924-0136 (96) 02280-7
In practice, metal forming operations are classified as either massive or sheet metal forming. In sheet metal forming, sheet blanks are formed into hollow bodies without intended changes of the initial sheet thickness. In ironing, however, the wall thickness is reduced significantly (25-70%). Thus, ironing is not a typical sheet metal forming operation. Frictional stresses at the p u n c h / c u p interface within the deformation zone pull the material of the cup through the die [1;2]. Due to these frictional stresses the deformation force is transmitted by the punch instead of the cross-sectional area which supports the force in wire or tube drawing. Reducing the wall thickness of the drawn cup determines the final height of the can. Ironing also gives the cup the desired mirror surface finish. A major advantage of ironing is the distribution of a very uniform wall thickness. Usually, manufacturing of beverage cans requires a three station process (Fig. 2). The plain sheet is deep drawn into a cup. The pre-drawn cup is fed into a horizontally w o r k i n g press and its diameter is reduced in redrawing. Finally, the cup is ironed in three subsequent steps. In the last operation, the
2
M Schiinemann et al./Journal of Materials Processing Technology 59 (19 96) 1-9
Cup -58
Dimensions in mm
Beveroge can
' o.ssI o.ogs
Ironing punch Redrowing die
1.
'
2.
Ir oning die
3.
1
Doming toot
Fig. 2: Drawing and wall ironing [3]
doming tool forms the can bottom. Due to the nature of the process, intermediate annealing can not be done. Presses with relatively longer strokes (600-750 mm) and higher speeds (400 spm) are used for the operation [4]. A l u m i n u m and its alloys are used as the p r i m a r y can material. The manganese and m a g n e s i u m alloyed a l u m i n u m alloys of the 3xxx series are preferably used. 2. Finite E l e m e n t M e t h o d (FEM) S i m u l a t i o n s
Few attempts were m ade to investigate the influence of process parameters in the ironing process using FEM. The earliest w o r k done by Odell [2]. An elastic-plastic FEM code was used to study the effects of the die angle and the punch and die friction on the m a x i m u m reduction ratio (MRR, ratio of the initial wall thickness to the final wall thickness). Odell found that MRR increases with decreasing die angle, increasing punch friction, and decreasing die friction. A 2-D analysis e m p l o y i n g a thermo elasto-plastic m o d e l w i t h an i n t e r n a l c o n s t i t u t i v e va r i abl e describing the microstructural state of the material has been c o n d u c t e d by Tjotta and H e i m l u n d [5]. A l u m i n u m alloy 1100 has been subjected to a three stage process with cold extrusion and two ironing (thickness was r e d u c e d from 4.25 to 2.25 m m and from 2.25 to 1.25 m m in first and second ironing o p e r a t i o n s , respectively). An objective of the simulation was to examine the feasibility of a single
ironing operat i on to reach the same final wall thickness. In the single pass ironing operation, necking was predicted. A nal yt i cal and e x p e r i m e n t a l i n v e s t i g a t i o n s studies upon the residual stress distribution were carried out by Danckert [6]. Austenitic stainless steel (AISI 304L) was d r a w n , r e d r a w n and ironed. Simulations were carried out using the explicit FEM code Dyna-2D. It was predicted that ironing causes compressive stresses near the outside and the inside surfaces of the cup. The residual axial stresses in the middle are of a tensile nature. Takeuchi [7] has c o n d u c t e d FEM s i m u l a t i o n s for the n e c k e d - i n process, the scoring of the lid and the optimization of the dome. He has found that the use of FEM in combination with new aluminum alloys is a powerful tool to reduce the can weight. The geometrical accuracy in ironing has been analyzed by Kampus et al [8]. A non-isothermal FEM analysis of one deep drawing and one subsequent i r o n i n g o p e r a t i o n s was c o n d u c t e d using the c o m p u t e r code ABAQUS. In these simulations, elastic deflections of both the punch and the die were taken into consideration. According to the author, the geometry of the part is strongly influenced by the heat generated during the process. The distributed heat affects the diameters of the punch and the die, as well as the geometry of the warm part. Since the part is expected to have a near net shape geometry, it is desired to control the temperature.
M. Schiinemann et al. /Journal of Materials Processing Technology 59 (1996) 1-9
3. Multi-pass ironing Sachs and L u b a h n [9] has investigated the influence of the spacing between two ironing dies (Fig. 3). They categorize the tandem ironing process by different spacing between the ironing dies. They f o u n d that m a x i m u m r e d u c t i o n s i n c r e a s e d considerably w h e n the wall thickness is reduced simultaneously in two dies. They have carried out several tests to determine the dependencies of the maximum possible second reduction on the spacing and the first reduction. They concluded that a bigger spacing leads to a larger second reduction. The level of damage through ironing and deep d r a w i n g has been investigated by Baudelet and Grange [10]. For this purpose, the Cockroft/Latham criterion was utilized. Experiments were performed to form aluminum alloy 3104 cups in two operations followed by three ironing stages which reduced the wall thickness from 0.28 to 0.125 mm. Variation of density was m e a s u r e d on the formed parts. The authors found that ironing tends to increase the density and reduce the damage. An optimization method to determine the process parameters, namely the partition of the ironing ratio in two pass ironing, has been developed by Xiaozhen et al [11]. They have found that the drawing load was significantly affected by the partition of the ironing ratio. However, in this optimization m e t h o d , the temperature effects were neglected. Evolution of the mechanical properties during the can making operations, using aluminum alloy 3104 sheets, has been investigated by Merchant et al [12]. In this experimental study, two drawing and three ironing steps were performed. Property changes during forming were f o u n d relatively small since the drawing steps soften the material whereas ironing increases the strength of the material. The strength of the ironed wall was found to be 10-15% higher than the strength of the initial sheet. The detected temperature rise during ironing was about 100°C[12]. A test method to measure the friction coefficient during ironing was d e v e l o p e d by Rajagopal and Misra [13].They tried to determine the punch and die
3
friction coefficient separately. However, the method they d e v e l o p e d is applicable only to small reductions. A c c o r d i n g to Panknin [14], the characteristic parameters for the ironing process are (a) punch force, (b) die angle, (c) friction, (d) wall thickness reduction. He developed design rules for the process considering the interactions between these parameters and predicted no considerable a d v a n t a g e s of s i m u l t a n e o u s ironing. On the contrary, bulging might occur b e t w e e n dies. An extensive literature review on manufacturing of beverage cans is summarized by Tufekci et al. [15]. A parametric study was also conducted in a single stage ironing to illustrate the interactions among process variables. 4. Simulations of small cup and comparison with experimental data from literature Drawing, redrawing, and ironing of aluminum sheet were simulated with DEFORM 2D which is an implicit rigid-plastic finite e l e m e n t p r o g r a m originally written for billet forming applications. However, the applicability of DEFORM 2D to sheet metal forming simulations w a s tested at the ERC/NSM, and the code was found to be quite reliable despite large CPU times. To verify the predictions of DEFORM for ironing, simulations were conducted for the experimental study performed by Saito et al [5] for AA 1100-O. Dimensions of the ironing test cup from the experiments are shown in Fig. 4. In this study, circular blanks of 0.8 mm thick aluminum were first deep drawn, then ironed to 0.275 m m wall thickness, and finally, annealed at a temperature of 350°C for an hour. Then, tension test specimens were cut from the cup wall to determine material properties. Note that cups were pre-ironed to a height of 5 mm to satisfy the design specifications before the actual ironing experiments were conducted.
20
3.275 O
! S=oo
S=O
(A) (B) (C) Fig. 3: Schematic of double ironing, (A: infinite spacing, B: regular spacing, C: zero spacing, S: spacing) [9].
Fig. 4: Dimensions of the ironing test cups [16].
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3/i. Schitnemann et al./Journal o f Materials Processing Technology 59 (1996) 1-9
D i m e n s i o n s of the c u p a n d the die u s e d in the s i m u l a t i o n s are s h o w n in Fig. 5 a n d the p r o c e s s p a r a m e t e r s are s u m m a r i z e d in T a b l e 1. Fig. 6 illustrate the c o m p a r i s o n b e t w e e n the p u n c h l o a d m e a s u r e m e n t s a n d p r e d i c t i o n s b y D E F O R M as a f u n c t i o n of p u n c h s t r o k e for 34.5% t h i c k n e s s reduction. The p r e d i c t i o n s are in g o o d a g r e e m e n t with measurements. The differences between p r e d i c t e d a n d m e a s u r e d p u n c h force v a l u e s at the b e g i n n i n g of the stroke, a n d the difference b e t w e e n the p r e d i c t e d a n d m e a s u r e d l e n g t h of the p u n c h strokes are d u e to n e g l e c t i n g the p r e - i r o n i n g of the test c u p s to a height of 5 m m d u r i n g simulations. The g e o m e t r y of the c u p a n d the dies in the d o u b l e ironing is s h o w n in Fig. 7. In the simulations, dies are d e f i n e d as m o v i n g objects. The distance b e t w e e n the t w o dies (S) is an i m p o r t a n t p a r a m e t e r that n e e d s to be defined d u r i n g process design. The true strain in one die is defined as follows: e,,i¢ = lnt,,
In the literature, several criteria are u s e d to assess the q u a l i t y of a n i r o n i n g p r o c e s s w i t h g i v e n p a r a m e t e r s . Earlier e x p e r i m e n t a l p a p e r s utilize the m a x i m u m p u n c h force or the m a x i m u m r e d u c t i o n ratio, while recent analytical w o r k f a v o r s the wall stresses. The effective stress is practical to investigate the processes in the d e f o r m a t i o n zone, while the axial stress can be u s e d to e v a l u a t e the p r o b a b i l i t y of cracks in the ironed wall. In this p a p e r , b o t h of the stress v a l u e s h a v e b e e n used. The effective or v o n Mises stress is defined as: (3)
w h e r e a~ are p r i n c i p a l stresses. T h e m a x i m u m effective wall stress p r e d i c t e d in the s i m u l a t i o n s is s h o w n in Fig. 8. The n a t u r a l l o g a r i t h m of the true strain ratio has been u s e d as scale for the X-axis.
(1)
2000
t~
w h e r e t~ a n d t o are the final a n d initial wall thickness', r e s p e c t i v e l y a n d ~ is the l o g a r i t h m i c true strain. C o n s e q u e n t l y , ¢ is a n e g a t i v e n u m b e r for the ironing process. The total true strain is the s u m of the strains f r o m s e p a r a t e reductions.
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Fig. 6: Comparison of the ironing predictions and m e a s u r e m e n t s for 34.5% thickness r e d u c t i o n . Measurements are done by Saito et al [16].
P%~%~ K 0.05
r
A
B
Fig. 5: Initial geometry of the cup (A) and the die (B) [Dimensions are in m m ] Table 1: Ironing parameters Punch friction coefficient (m 1) Die friction coefficient (m 7) Punch speed Initial wall thickness Final wall thickness Die land Material: K, n (a=Ke n) Material: K, n (o=KEn)
0.06 - 0.08 0.07 - 0.09 1.2 m / s 0.275 m m 0.123 mm/0.1787 m m 0.175 m m AA 1100-O 154.8 MPa, 0.2 AA 3104-H19 360 MPa, 0.0065
-
{} VDIE r ,
r
Fig. 7: Arrangement of cup and dies in the simulation. (s: effective die distance)
M Schiinemann et al./Journal o f M a t e r . s Processing Technology 59 (1996) 1-9 165
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Fig. 9: Typical load stroke curves for double ironing
lnE1/E2 400
Fig. 8: Maximum effective wall stress for double and single ironing of aluminum alloy 1100-O.
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where ~1 and ~2 are the true strains calculated for the first and the second die, respectively. For equal true strains at first and second die this scale shows zero. For bigger reductions in the first die, the equation yields positive numbers, while it shows negative values for bigger reductions at the second die. The calculated effective wall stresses for the ironing of AA 1100-0 are comparable to stresses predicted for the single reduction (Fig. 7), and the differences between the calculated stresses is small, only within _+10%. Ironing of a l u m i n u m alloy 3104-H19 is also simulated using the same geometry (Fig. 5) and the process parameters (Table 1). Predicted forces are shown in Fig. 9. Three phases of the deformation can be distinguished in this figure. The first phase with only the first die in contact, the second phase with two dies in contact, and finally, the third phase with only the second die in contact with the cup. It can be seen that the total punch force is the sum of the forces acting on the dies. In phases 1 and 3, the total punch force is equal to the force carried by the die in contact. Hence, a die distance (Fig. 7) that exceeds the cup height makes this sum equal to the forces of the contact die. Fig. 10 shows the calculated maximum values for the effective stress, the axial stress and the punch force. While effective stresses do not seem to be affected by varying die distance, axial stresses and punch forces change significantly. The course of the axial stress curve as well as the punch force curve show an upper and a lower level. The transition from one level to the other occurs when the die distance is larger than or equal to the cup height. Thus, an infinite spacing (i.e., cup in contact with one die at a time) is most beneficial for the process.
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-- 200
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1
50 0
I ................ 0
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i .... 10
15
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.... 23
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Die Distance
Fig. 10: Maximum axial stress, maximum effective stress and maximum punch force vs. die distance, S, for the ironing of AA 3104-H19 with 55% reduction ratio. 5. P r o c e s s s i m u l a t i o n s for can m a n u f a c t u r i n g
The f o r m i n g p r o c e s s for b e v e r a g e can manufacturing requires several production steps to form the sheet metal. These processes are (in operation sequence): 1) blanking, 2) drawing, 3) redrawing, 4) first ironing, 5) second ironing, 6) third ironing, 7) doming, 8) f l a n g i n g / n e c k i n g . Nonisothermal simulations of steps 2-6 were carried out using DEFORM 2D. The workpiece was considered rigid-plastic, while tooling was considered rigid. Simulations were conducted using aluminum alloys 3104-H19 and 3003. Flow stress data for AA 3104H19 are given in Table 1. Since heat dependent flow stress curves were not available for AA 3104-H19, AA 3003 was also used in the simulations although it is not a typical material used in can manufacturing. Fig. 11 illustrates the flow stress of AA 3003 for various temperatures. Thermal properties used for the workpiece (AA 3104-H19/AA 3003) and tooling ( H l l tool steel) are
6
A~ Schiinemann et al. /Journal of Matertals Processing Technology 59 (1996) 1-9 250
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Fig. 11: Flow-stress data for AA 3003 at various temperatures. given in Table 2. The thermal properties of AA 3104H19 and AA 3003 could not be obtained, and hence, the data for AA 1100-O were used in the simulations. Process p a r a m e t e r s u s e d in the simulations of deep drawing, r e d r a w i n g and ironing are s u m m a r i z e d in Table 3. Table 2: Thermal properties for the non-isothermal simulations. Material
Thermal Conductivity (kin) N/sec/C o Heat capacity (CJ N/mm2/C ° Emissivity (E n~
Workpiece tAA 1100-0)
Tooling /H 11)
0.22429E+03
0.24575E+02
0.20559E+01 0.05
0.34746E+01 0.6
Table 3 : Parameters for the simulation of deep drawing and redrawing Blank holder friction coefficient (~t~) Punch friction coefficient (~tr/ Die friction coefficient (~tn) Punch speed (v) Initial sheet thickness (tll) Blankholder force: Drawing Redrawing Material
0.09 0.09 0.06 1.0 m/s 0.292 mm 10000 N 2000 N AA 3003 AA 3104
The g e o m e t r y of the initial b l a n k and the cups f o r m e d in d e e p d r a w i n g a n d r e d r a w i n g operations are s h o w n in Fig. 12. Note the location of the points P1, P2 and P3 on the flat sheet. T e m p e r a t u r e and axial stress predictions will be d i s p l a y e d at these points. Tool g e o m e t r i e s for d e e p d r a w i n g a n d r e d r a w i n g o p e r a t i o n s are s u m m a r i z e d in Table 4. R o u n d cup f o r m e d in the r e d r a w i n g operation (whose g e o m e t r y
0.299
Fig. 12: Sheet metal blank and part geometries from the deep drawing and redrawing operations. is s h o w n in Fig. 12) was ironed in three s u b s e q u e n t operations. Part geometries obtained in the ironing o p e r a t i o n s are s u m m a r i z e d in Table 5. The tool geometry for the ironing operations is the same as the one s h o w n in Fig. 5. Table 4: Geometry of the dies and punches used in the drawing and redrawing simulations Drawin~ Redrawin~ Punch diameter (dr) 90.55 mm 66.00 mm Die diameter (d p) 91.21 mm 66.74 mm Punch profile radius (re) 3.175mm 5.080 mm Die profile radius (r~) 6.35 mm 6.35 mm Clearance/c) 0.33 mm 0.37 mm
Table 5: Part geometries obtained in the ironing operations. Ironing Stage
Inside Diameter
Wall Thickness
Hei;~t
/mmt
Immt
1st 2nd 3rd
66.0 66.0 66.0
.229 .183 .152
66.11 82.61 99.53
Cu
5.1. Deep drawing and redrawing simulations
After the d r a w i n g s i m u l a t i o n s , the cup was transferred to the r e d r a w i n g tool. In this p r o c e d u r e , the d e f o r m a t i o n h i s t o r y of the c u p w a s k e p t u n c h a n g e d , w h i l e the n o d a l t e m p e r a t u r e s w e r e initialized a n d set to 20°C. In the m a n u f a c t u r i n g process, the d r a w n cup is p r o d u c e d in a separate tooling and cools d o w n to r o o m t e m p e r a t u r e d u r i n g the transport. The generated heat for selected points,
M. Schiinemann et al. /Journal o f Materdals Processing Technology 59 (1996) 1-9 45
7
14
0
!0
20
30
0
40
10
20
30
40
Punch Stroke [mm]
P u n c h Stroke [mini
Fig. 13: Temperatures of selected points at various locations during the deep drawing simulations of AA 3003.
Fig. 15: Load-stroke curves for the drawing simulations. 12
60
Pointl
Pomt2 mapoilR3
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m
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2 30
0
20
50
60
70
Punch Stroke [mm]
20 0
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20
30
40
50
60
70
Punch Stroke [ram]
Fig. 14: Temperatures of selected points at various locations during the redrawing simulations of AA 3003. and P3 (see Fig. 12) on the cup surface can be seen in Fig. 13 for the drawing and in Fig. 14 for the redrawing of AA 3003. The points were located at 30 m m ( P 1 ) , 50 m m (P2) and 69 m m (P3) radius of the initial sheet (Fig. 12). The variation of the temperature with punch stroke is very similar for drawing and redrawing. At Point 1, located at the cup bottom, after both drawing a n d r e d r a w i n g , the t e m p e r a t u r e increase is insignificant due to the fact that there is not much deformation in this region. The temperature increase at point 2 is caused by the metal deformation, whereas the major part of the generated heat at point 3 is created by frictional interactions. The thickening of the wall during drawing as well as redrawing is significant. P1, P2
10
Fig. 16: Load-stroke curves for the redrawing simulations. The effect of sizing can also be illustrated in the load stroke curves of drawing and redrawing. In Fig. 15, the sizing starts at 28 m m punch stroke for AA 3104-H19 and at 32 m m for AA 3003 for the drawing simulations. The sizing in the redrawing is less significant for both materials (Fig. 16). It starts at approximately 52 m m for both materials. The significantly elevated punch forces for the 3rawing of AA 3104-H19 are caused by the higher initial flow stress due to high prior cold forming. Additionally, the neglected heat dependency of the flow stress curve causes disproportional forces.
5.2. Multi stage ironing simulations According to manufacturers, the ironing is carried out in three subsequent steps with infinite spacing [3]. Thus, throughout the process, only one die is in contact with the cup. From Section 4 we know the
8
M Schi~nemann et al./Journal o f Materials Processing Technology .59 (1996) 1-9 140
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Fig. 17: Load-stroke curves for the 1st ironing step with different ironing ratios and different materials. axial stresses for this a r r a n g e m e n t are l o w e r t h e n those for the simultaneous ironing. T w o sets of three ironing stages w e r e simulated. In the first set, the true strains w e r e -0.498, -0.223, and -0.182, while the true strains in the s e c o n d set w e r e constant-0.301. The reduction in the first step refers to the m a x i m u m wall thickness of the r e d r a w n cup. The thickness of the cup after the r e d r a w i n g is not u n i f o r m due to thickening. The m a x i m u m thickness of the cup wall is f o u n d at the c u p rim. Thus, the ironing ratio is not constant d u r i n g the first ironing. In the design of the ironing process, it is i m p o r t a n t to account for this fact. Also, the tolerances of the u s e d coil material h a v e to be considered. In addition, wall thickening a n d e v e n t u a l l y d i l a t a t i o n i n c r e a s e the a b s o l u t e i r o n i n g ratio a n d cause h i g h e r wall stresses and p u n c h loads. Load-stroke curves for the first ironing step with different i r o n i n g ratios and different materials are illustrated in Fig. 17. An increasing p u n c h load d u r i n g the first ironing step is a result of two effects; 1) the real r e d u c t i o n ratio increases w i t h the cup height, 2) the strain h a r d e n i n g of the cup material increases with the cup height. A f t e r the first i r o n i n g , the w a l l thickness is u n i f o r m o v e r the cup height. C o n s e q u e n t l y , the increase of the p u n c h force d u r i n g the second and third ironing is d u e to strain hardening. The ironing ratio is t h e n constant t h r o u g h o u t the process. Higher flow stresses in the u p p e r part of the cup wall result in higher temperatures, as seen in Fig. 18. Fig. 19 illustrates the t e m p e r a t u r e distribution in the tooling a n d the cup d u r i n g the first ironing step at a p p r o x i m a t e l y 4 / 5 of the ironing stroke. Fig. 19 refers to the i r o n i n g of A A 3003 w i t h an ironing reduction (~) of -0.498. The t e m p e r a t u r e distribution
,
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Fig. 18: Temperatures of the selected points during the ironing simulations of AA 3003 with ironing ratios of e,/2/3=-0.498/-0.223/-0.182. (3bi~tl
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for the ironing simulation of AA 3003. was f o u n d to be representative for all simulations, while the absolute t e m p e r a t u r e s changed. Typically, the cup and the die show the highest t e m p e r a t u r e s in the d e f o r m a t i o n area, w h i l e the h i g h e s t p u n c h t e m p e r a t u r e is f o u n d b e l o w the d e f o r m a t i o n area closer to the cup bottom. H i g h p u n c h s p e e d causes h i g h sliding b e t w e e n cup a n d die. The sliding b e t w e e n cup and p u n c h is significantly lower, hence, the heat transfer b e t w e e n p u n c h and cup h a p p e n s m o s t l y in the area b e h i n d the die o u t l e t b y conduction. The tool area influenced b y the increased t e m p e r a t u r e s is v e r y small, s o m e m i l l i m e t e r s at m a x i m u m , but the tooling is expected to w a r m up if more then one can is produced.
M Schfgnemann et al/Journal of Materials Processing Technology 59 (1996) 1-9
60
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Fig. 20: Axial stresses of selected points for the ironing simulation of AA 3003 with the reduction ratios of E1/2/3=0.498/-0.223/-0.182. The axial stresses calculated for the ironing of AA 3003 are shown in Fig. 20. As expected, the axial stress in the point located at the cup bottom stays very low, while the stresses at the other points raise after the ironing die passes their location. Once the ironing die passed the location of point 2, the axial stress increases to its m a x i m u m until the die loses contact with the cup. The scatter in the curves is c a u s e d by elements leaving and entering the deformation zone. References
[1] Altan, Oh, S., and Gegel, H., Metal FormingF u n d a m e n t a l s and Applications, A m e r i c a n Society of Metals, Metals Park, Ohio (1983). [2] Odell, E. I., "A Study of the Wall Ironing by the Finite Element Technique", J. of Eng. for Ind., vol. 100, pp. 31-36 (1978) [3] Sodeik, M., Taeffner, K., and Weber, F., " F u n d a m e n t a l s of M o d e r n Can Making and Materials D e v e l o p m e n t for Two-Piece Can Manufacturing", Transactions of lSU, vol. 28, pp. 672-677 (1988) [4] Sheet Metal Industries, "Deep Drawing and Ironing of Non-Ferrous Material", Sheet Metal Industries, vol. 65, pp. 284-286, June (1988)
[5] Tjotta, S., and Heimlund, O., "Finite Element Simulation in Cold Forging Process Design", J. of Mat. Proc. Tech., vol. 36, pp. 76-96 (1992) [6] Danckert, J., "The Residul Stresses in the Wll of a Deep D r a w n and Ironed Cup D e t e r m i n e d Experimentally and by FEM", CIRP Annals, vol. 43/1, pp. 249-252 (1994) [7] Takeuchi, H., "Numerical Simulation Technology for Lightweight Aluminum Cans", J. of Mat. Proc. Tech., vol. 38, pp. 675-678 (1993) [8] Kampus, Z., and Kuzman, K., "Analysis of the Geometrical Accuracy in Ironing", A d v a n c e d Technology of Plasticity, Proceedings of the 4th Int. Conf. on Tech. of Plasticity, vol. 2, pp. 10051010 (1993) [9] Sachs, G., and Espey, G., "Effect of Spacing between Dies in the Tandem Drawing of Tubular Parts", Trans. of the ASME, February (1947) [10] Baudelet, B., and Grange, B., "Damage in Deep Drawn and Ironed Can Bodies in an Aluminum Alloy", Scripta Metallurgia et Materialia, vol 26, no. 3, pp. 375-379 (1992) [11] Xiaozhen, M., Yuying, Y., and Yingchun, L., "Experimental Study on Ironing of Stainless Steel and Optimization of Process Parameters", Advanced Technology of Plasticity, Proceedings of the 4th Int. Conf. on Tech. of Plasticity, vol. 3, pp. 1653-1656 (1993) [12] Merchant, H. D., Hodgson, D. S., O'Reilly, I., and Embury, J. D., "Structure and Property Evolution During Drawing and Wall Ironing of AA 3004", Materials Characterization, vol. 25, pp. 251-261 (1990) [13] Rajagopal, S., and Misra, S., "Measurements of Differential Friction Coefficients in Ironing", Proceedings of the 18th NAMRC, pp. 89-95 (1990) [14] Panknin, W., "Grundlagen des Tiefziehens zur Herstellung Zweiteiliger Dosen", Werkstatt und Betrieb, vol 110, no. 5, pp. 313-319 (1977) [15] Tufekci, S. S., Ahmetoglu, M. A., Kinzel, G. L., and Altan, T., "Process Simulation for Can Manufacturing by Deep Drawing and Ironing", SAE Paper No. 950696, International Congress and Exposition, Detroit, Michigan, February 27March 2 (1995) [16]Saito, M., 1989, Saiki, H., and Kawai, N., "Experimental Analysis of Ironing of Thin Metal Cups", Trans. of the ASME, vol. 111, pp. 56-63, February (1989)
ELSEVIER
Journal of Materials Processing Technology 59 (1996) 1-9
Materials Processing Technology
Prediction of process conditions in drawing and ironing of cans M a r c o S c h f . i n e m a n n , M u s t a f a A. A h m e t o g l u * , a n d T a y l a n A l t a n
ERC for Net Shape Manufacturing, Ohio State University, Columbus, Ohio 43210, USA
Industrial S u m m a r y
Deep drawing and ironing are the major processes used today in manufacturing of beverage cans from aluminum. The same technology is utilized in manufacturing of steel cans for the food industry, photosensitive drums and pressure vessels for fire extinguishers. Most of these processes require multi-stage ironing following the deep drawing and redrawing or backward extrusion processes. The practical aspects of this technology are well k n o w n and gained through extensive experimentation and production know-how. H o w e v e r , the fundamental aspects of these processes are relatively less known, especially regarding the temperature developed during deformation and the effect of deformation speed upon temperatures and lubrication. Thus, it is expected that process simulations using FEM techniques would provide additional detailed information that could be utilized to improve the process conditions. This paper illustrates the application of process modeling to deep drawing and ironing operations. A commercially available finite element code DEFORM was used to simulate deep drawing, redrawing and multi-stage ironing operations with varying parameters to illustrate the interactions between process variables.
1. I n t r o d u c t i o n
Ironing, together with drawing and redrawing, is the major process used in can manufacturing. The principle of ironing with a single die and a cylindrical punch is illustrated in Fig. 1. The moving punch, or ram, draws the cup through one or more ironing rings. The thickness of the cup bottom remains almost unchanged, while the thickness of the wall is reduced. Die
Fig. 1: Single ironing with a cylindrical punch [1]
* Corresponding author 0924-0136/96/$15.00 © 1996 Elsevier Science S.A~All rights reserved PI10924-0136 (96) 02280-7
In practice, metal forming operations are classified as either massive or sheet metal forming. In sheet metal forming, sheet blanks are formed into hollow bodies without intended changes of the initial sheet thickness. In ironing, however, the wall thickness is reduced significantly (25-70%). Thus, ironing is not a typical sheet metal forming operation. Frictional stresses at the p u n c h / c u p interface within the deformation zone pull the material of the cup through the die [1;2]. Due to these frictional stresses the deformation force is transmitted by the punch instead of the cross-sectional area which supports the force in wire or tube drawing. Reducing the wall thickness of the drawn cup determines the final height of the can. Ironing also gives the cup the desired mirror surface finish. A major advantage of ironing is the distribution of a very uniform wall thickness. Usually, manufacturing of beverage cans requires a three station process (Fig. 2). The plain sheet is deep drawn into a cup. The pre-drawn cup is fed into a horizontally w o r k i n g press and its diameter is reduced in redrawing. Finally, the cup is ironed in three subsequent steps. In the last operation, the
2
M Schiinemann et al./Journal of Materials Processing Technology 59 (19 96) 1-9
Cup -58
Dimensions in mm
Beveroge can
' o.ssI o.ogs
Ironing punch Redrowing die
1.
'
2.
Ir oning die
3.
1
Doming toot
Fig. 2: Drawing and wall ironing [3]
doming tool forms the can bottom. Due to the nature of the process, intermediate annealing can not be done. Presses with relatively longer strokes (600-750 mm) and higher speeds (400 spm) are used for the operation [4]. A l u m i n u m and its alloys are used as the p r i m a r y can material. The manganese and m a g n e s i u m alloyed a l u m i n u m alloys of the 3xxx series are preferably used. 2. Finite E l e m e n t M e t h o d (FEM) S i m u l a t i o n s
Few attempts were m ade to investigate the influence of process parameters in the ironing process using FEM. The earliest w o r k done by Odell [2]. An elastic-plastic FEM code was used to study the effects of the die angle and the punch and die friction on the m a x i m u m reduction ratio (MRR, ratio of the initial wall thickness to the final wall thickness). Odell found that MRR increases with decreasing die angle, increasing punch friction, and decreasing die friction. A 2-D analysis e m p l o y i n g a thermo elasto-plastic m o d e l w i t h an i n t e r n a l c o n s t i t u t i v e va r i abl e describing the microstructural state of the material has been c o n d u c t e d by Tjotta and H e i m l u n d [5]. A l u m i n u m alloy 1100 has been subjected to a three stage process with cold extrusion and two ironing (thickness was r e d u c e d from 4.25 to 2.25 m m and from 2.25 to 1.25 m m in first and second ironing o p e r a t i o n s , respectively). An objective of the simulation was to examine the feasibility of a single
ironing operat i on to reach the same final wall thickness. In the single pass ironing operation, necking was predicted. A nal yt i cal and e x p e r i m e n t a l i n v e s t i g a t i o n s studies upon the residual stress distribution were carried out by Danckert [6]. Austenitic stainless steel (AISI 304L) was d r a w n , r e d r a w n and ironed. Simulations were carried out using the explicit FEM code Dyna-2D. It was predicted that ironing causes compressive stresses near the outside and the inside surfaces of the cup. The residual axial stresses in the middle are of a tensile nature. Takeuchi [7] has c o n d u c t e d FEM s i m u l a t i o n s for the n e c k e d - i n process, the scoring of the lid and the optimization of the dome. He has found that the use of FEM in combination with new aluminum alloys is a powerful tool to reduce the can weight. The geometrical accuracy in ironing has been analyzed by Kampus et al [8]. A non-isothermal FEM analysis of one deep drawing and one subsequent i r o n i n g o p e r a t i o n s was c o n d u c t e d using the c o m p u t e r code ABAQUS. In these simulations, elastic deflections of both the punch and the die were taken into consideration. According to the author, the geometry of the part is strongly influenced by the heat generated during the process. The distributed heat affects the diameters of the punch and the die, as well as the geometry of the warm part. Since the part is expected to have a near net shape geometry, it is desired to control the temperature.
M. Schiinemann et al. /Journal of Materials Processing Technology 59 (1996) 1-9
3. Multi-pass ironing Sachs and L u b a h n [9] has investigated the influence of the spacing between two ironing dies (Fig. 3). They categorize the tandem ironing process by different spacing between the ironing dies. They f o u n d that m a x i m u m r e d u c t i o n s i n c r e a s e d considerably w h e n the wall thickness is reduced simultaneously in two dies. They have carried out several tests to determine the dependencies of the maximum possible second reduction on the spacing and the first reduction. They concluded that a bigger spacing leads to a larger second reduction. The level of damage through ironing and deep d r a w i n g has been investigated by Baudelet and Grange [10]. For this purpose, the Cockroft/Latham criterion was utilized. Experiments were performed to form aluminum alloy 3104 cups in two operations followed by three ironing stages which reduced the wall thickness from 0.28 to 0.125 mm. Variation of density was m e a s u r e d on the formed parts. The authors found that ironing tends to increase the density and reduce the damage. An optimization method to determine the process parameters, namely the partition of the ironing ratio in two pass ironing, has been developed by Xiaozhen et al [11]. They have found that the drawing load was significantly affected by the partition of the ironing ratio. However, in this optimization m e t h o d , the temperature effects were neglected. Evolution of the mechanical properties during the can making operations, using aluminum alloy 3104 sheets, has been investigated by Merchant et al [12]. In this experimental study, two drawing and three ironing steps were performed. Property changes during forming were f o u n d relatively small since the drawing steps soften the material whereas ironing increases the strength of the material. The strength of the ironed wall was found to be 10-15% higher than the strength of the initial sheet. The detected temperature rise during ironing was about 100°C[12]. A test method to measure the friction coefficient during ironing was d e v e l o p e d by Rajagopal and Misra [13].They tried to determine the punch and die
3
friction coefficient separately. However, the method they d e v e l o p e d is applicable only to small reductions. A c c o r d i n g to Panknin [14], the characteristic parameters for the ironing process are (a) punch force, (b) die angle, (c) friction, (d) wall thickness reduction. He developed design rules for the process considering the interactions between these parameters and predicted no considerable a d v a n t a g e s of s i m u l t a n e o u s ironing. On the contrary, bulging might occur b e t w e e n dies. An extensive literature review on manufacturing of beverage cans is summarized by Tufekci et al. [15]. A parametric study was also conducted in a single stage ironing to illustrate the interactions among process variables. 4. Simulations of small cup and comparison with experimental data from literature Drawing, redrawing, and ironing of aluminum sheet were simulated with DEFORM 2D which is an implicit rigid-plastic finite e l e m e n t p r o g r a m originally written for billet forming applications. However, the applicability of DEFORM 2D to sheet metal forming simulations w a s tested at the ERC/NSM, and the code was found to be quite reliable despite large CPU times. To verify the predictions of DEFORM for ironing, simulations were conducted for the experimental study performed by Saito et al [5] for AA 1100-O. Dimensions of the ironing test cup from the experiments are shown in Fig. 4. In this study, circular blanks of 0.8 mm thick aluminum were first deep drawn, then ironed to 0.275 m m wall thickness, and finally, annealed at a temperature of 350°C for an hour. Then, tension test specimens were cut from the cup wall to determine material properties. Note that cups were pre-ironed to a height of 5 mm to satisfy the design specifications before the actual ironing experiments were conducted.
20
3.275 O
! S=oo
S=O
(A) (B) (C) Fig. 3: Schematic of double ironing, (A: infinite spacing, B: regular spacing, C: zero spacing, S: spacing) [9].
Fig. 4: Dimensions of the ironing test cups [16].
4
3/i. Schitnemann et al./Journal o f Materials Processing Technology 59 (1996) 1-9
D i m e n s i o n s of the c u p a n d the die u s e d in the s i m u l a t i o n s are s h o w n in Fig. 5 a n d the p r o c e s s p a r a m e t e r s are s u m m a r i z e d in T a b l e 1. Fig. 6 illustrate the c o m p a r i s o n b e t w e e n the p u n c h l o a d m e a s u r e m e n t s a n d p r e d i c t i o n s b y D E F O R M as a f u n c t i o n of p u n c h s t r o k e for 34.5% t h i c k n e s s reduction. The p r e d i c t i o n s are in g o o d a g r e e m e n t with measurements. The differences between p r e d i c t e d a n d m e a s u r e d p u n c h force v a l u e s at the b e g i n n i n g of the stroke, a n d the difference b e t w e e n the p r e d i c t e d a n d m e a s u r e d l e n g t h of the p u n c h strokes are d u e to n e g l e c t i n g the p r e - i r o n i n g of the test c u p s to a height of 5 m m d u r i n g simulations. The g e o m e t r y of the c u p a n d the dies in the d o u b l e ironing is s h o w n in Fig. 7. In the simulations, dies are d e f i n e d as m o v i n g objects. The distance b e t w e e n the t w o dies (S) is an i m p o r t a n t p a r a m e t e r that n e e d s to be defined d u r i n g process design. The true strain in one die is defined as follows: e,,i¢ = lnt,,
In the literature, several criteria are u s e d to assess the q u a l i t y of a n i r o n i n g p r o c e s s w i t h g i v e n p a r a m e t e r s . Earlier e x p e r i m e n t a l p a p e r s utilize the m a x i m u m p u n c h force or the m a x i m u m r e d u c t i o n ratio, while recent analytical w o r k f a v o r s the wall stresses. The effective stress is practical to investigate the processes in the d e f o r m a t i o n zone, while the axial stress can be u s e d to e v a l u a t e the p r o b a b i l i t y of cracks in the ironed wall. In this p a p e r , b o t h of the stress v a l u e s h a v e b e e n used. The effective or v o n Mises stress is defined as: (3)
w h e r e a~ are p r i n c i p a l stresses. T h e m a x i m u m effective wall stress p r e d i c t e d in the s i m u l a t i o n s is s h o w n in Fig. 8. The n a t u r a l l o g a r i t h m of the true strain ratio has been u s e d as scale for the X-axis.
(1)
2000
t~
w h e r e t~ a n d t o are the final a n d initial wall thickness', r e s p e c t i v e l y a n d ~ is the l o g a r i t h m i c true strain. C o n s e q u e n t l y , ¢ is a n e g a t i v e n u m b e r for the ironing process. The total true strain is the s u m of the strains f r o m s e p a r a t e reductions.
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Fig. 6: Comparison of the ironing predictions and m e a s u r e m e n t s for 34.5% thickness r e d u c t i o n . Measurements are done by Saito et al [16].
P%~%~ K 0.05
r
A
B
Fig. 5: Initial geometry of the cup (A) and the die (B) [Dimensions are in m m ] Table 1: Ironing parameters Punch friction coefficient (m 1) Die friction coefficient (m 7) Punch speed Initial wall thickness Final wall thickness Die land Material: K, n (a=Ke n) Material: K, n (o=KEn)
0.06 - 0.08 0.07 - 0.09 1.2 m / s 0.275 m m 0.123 mm/0.1787 m m 0.175 m m AA 1100-O 154.8 MPa, 0.2 AA 3104-H19 360 MPa, 0.0065
-
{} VDIE r ,
r
Fig. 7: Arrangement of cup and dies in the simulation. (s: effective die distance)
M Schiinemann et al./Journal o f M a t e r . s Processing Technology 59 (1996) 1-9 165
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t
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Fig. 9: Typical load stroke curves for double ironing
lnE1/E2 400
Fig. 8: Maximum effective wall stress for double and single ironing of aluminum alloy 1100-O.
~
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~ 300
x = In Cl
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........ "--i ................... 0
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where ~1 and ~2 are the true strains calculated for the first and the second die, respectively. For equal true strains at first and second die this scale shows zero. For bigger reductions in the first die, the equation yields positive numbers, while it shows negative values for bigger reductions at the second die. The calculated effective wall stresses for the ironing of AA 1100-0 are comparable to stresses predicted for the single reduction (Fig. 7), and the differences between the calculated stresses is small, only within _+10%. Ironing of a l u m i n u m alloy 3104-H19 is also simulated using the same geometry (Fig. 5) and the process parameters (Table 1). Predicted forces are shown in Fig. 9. Three phases of the deformation can be distinguished in this figure. The first phase with only the first die in contact, the second phase with two dies in contact, and finally, the third phase with only the second die in contact with the cup. It can be seen that the total punch force is the sum of the forces acting on the dies. In phases 1 and 3, the total punch force is equal to the force carried by the die in contact. Hence, a die distance (Fig. 7) that exceeds the cup height makes this sum equal to the forces of the contact die. Fig. 10 shows the calculated maximum values for the effective stress, the axial stress and the punch force. While effective stresses do not seem to be affected by varying die distance, axial stresses and punch forces change significantly. The course of the axial stress curve as well as the punch force curve show an upper and a lower level. The transition from one level to the other occurs when the die distance is larger than or equal to the cup height. Thus, an infinite spacing (i.e., cup in contact with one die at a time) is most beneficial for the process.
iiiiiiiiiiii
ii
5
4o~ .3
-- 200
2 ~. ~ too m
1
50 0
I ................ 0
5
i .... 10
15
20
.... 23
0 3O
Die Distance
Fig. 10: Maximum axial stress, maximum effective stress and maximum punch force vs. die distance, S, for the ironing of AA 3104-H19 with 55% reduction ratio. 5. P r o c e s s s i m u l a t i o n s for can m a n u f a c t u r i n g
The f o r m i n g p r o c e s s for b e v e r a g e can manufacturing requires several production steps to form the sheet metal. These processes are (in operation sequence): 1) blanking, 2) drawing, 3) redrawing, 4) first ironing, 5) second ironing, 6) third ironing, 7) doming, 8) f l a n g i n g / n e c k i n g . Nonisothermal simulations of steps 2-6 were carried out using DEFORM 2D. The workpiece was considered rigid-plastic, while tooling was considered rigid. Simulations were conducted using aluminum alloys 3104-H19 and 3003. Flow stress data for AA 3104H19 are given in Table 1. Since heat dependent flow stress curves were not available for AA 3104-H19, AA 3003 was also used in the simulations although it is not a typical material used in can manufacturing. Fig. 11 illustrates the flow stress of AA 3003 for various temperatures. Thermal properties used for the workpiece (AA 3104-H19/AA 3003) and tooling ( H l l tool steel) are
6
A~ Schiinemann et al. /Journal of Matertals Processing Technology 59 (1996) 1-9 250
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Fig. 11: Flow-stress data for AA 3003 at various temperatures. given in Table 2. The thermal properties of AA 3104H19 and AA 3003 could not be obtained, and hence, the data for AA 1100-O were used in the simulations. Process p a r a m e t e r s u s e d in the simulations of deep drawing, r e d r a w i n g and ironing are s u m m a r i z e d in Table 3. Table 2: Thermal properties for the non-isothermal simulations. Material
Thermal Conductivity (kin) N/sec/C o Heat capacity (CJ N/mm2/C ° Emissivity (E n~
Workpiece tAA 1100-0)
Tooling /H 11)
0.22429E+03
0.24575E+02
0.20559E+01 0.05
0.34746E+01 0.6
Table 3 : Parameters for the simulation of deep drawing and redrawing Blank holder friction coefficient (~t~) Punch friction coefficient (~tr/ Die friction coefficient (~tn) Punch speed (v) Initial sheet thickness (tll) Blankholder force: Drawing Redrawing Material
0.09 0.09 0.06 1.0 m/s 0.292 mm 10000 N 2000 N AA 3003 AA 3104
The g e o m e t r y of the initial b l a n k and the cups f o r m e d in d e e p d r a w i n g a n d r e d r a w i n g operations are s h o w n in Fig. 12. Note the location of the points P1, P2 and P3 on the flat sheet. T e m p e r a t u r e and axial stress predictions will be d i s p l a y e d at these points. Tool g e o m e t r i e s for d e e p d r a w i n g a n d r e d r a w i n g o p e r a t i o n s are s u m m a r i z e d in Table 4. R o u n d cup f o r m e d in the r e d r a w i n g operation (whose g e o m e t r y
0.299
Fig. 12: Sheet metal blank and part geometries from the deep drawing and redrawing operations. is s h o w n in Fig. 12) was ironed in three s u b s e q u e n t operations. Part geometries obtained in the ironing o p e r a t i o n s are s u m m a r i z e d in Table 5. The tool geometry for the ironing operations is the same as the one s h o w n in Fig. 5. Table 4: Geometry of the dies and punches used in the drawing and redrawing simulations Drawin~ Redrawin~ Punch diameter (dr) 90.55 mm 66.00 mm Die diameter (d p) 91.21 mm 66.74 mm Punch profile radius (re) 3.175mm 5.080 mm Die profile radius (r~) 6.35 mm 6.35 mm Clearance/c) 0.33 mm 0.37 mm
Table 5: Part geometries obtained in the ironing operations. Ironing Stage
Inside Diameter
Wall Thickness
Hei;~t
/mmt
Immt
1st 2nd 3rd
66.0 66.0 66.0
.229 .183 .152
66.11 82.61 99.53
Cu
5.1. Deep drawing and redrawing simulations
After the d r a w i n g s i m u l a t i o n s , the cup was transferred to the r e d r a w i n g tool. In this p r o c e d u r e , the d e f o r m a t i o n h i s t o r y of the c u p w a s k e p t u n c h a n g e d , w h i l e the n o d a l t e m p e r a t u r e s w e r e initialized a n d set to 20°C. In the m a n u f a c t u r i n g process, the d r a w n cup is p r o d u c e d in a separate tooling and cools d o w n to r o o m t e m p e r a t u r e d u r i n g the transport. The generated heat for selected points,
M. Schiinemann et al. /Journal o f Materdals Processing Technology 59 (1996) 1-9 45
7
14
0
!0
20
30
0
40
10
20
30
40
Punch Stroke [mm]
P u n c h Stroke [mini
Fig. 13: Temperatures of selected points at various locations during the deep drawing simulations of AA 3003.
Fig. 15: Load-stroke curves for the drawing simulations. 12
60
Pointl
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m
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2 30
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20
50
60
70
Punch Stroke [mm]
20 0
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30
40
50
60
70
Punch Stroke [ram]
Fig. 14: Temperatures of selected points at various locations during the redrawing simulations of AA 3003. and P3 (see Fig. 12) on the cup surface can be seen in Fig. 13 for the drawing and in Fig. 14 for the redrawing of AA 3003. The points were located at 30 m m ( P 1 ) , 50 m m (P2) and 69 m m (P3) radius of the initial sheet (Fig. 12). The variation of the temperature with punch stroke is very similar for drawing and redrawing. At Point 1, located at the cup bottom, after both drawing a n d r e d r a w i n g , the t e m p e r a t u r e increase is insignificant due to the fact that there is not much deformation in this region. The temperature increase at point 2 is caused by the metal deformation, whereas the major part of the generated heat at point 3 is created by frictional interactions. The thickening of the wall during drawing as well as redrawing is significant. P1, P2
10
Fig. 16: Load-stroke curves for the redrawing simulations. The effect of sizing can also be illustrated in the load stroke curves of drawing and redrawing. In Fig. 15, the sizing starts at 28 m m punch stroke for AA 3104-H19 and at 32 m m for AA 3003 for the drawing simulations. The sizing in the redrawing is less significant for both materials (Fig. 16). It starts at approximately 52 m m for both materials. The significantly elevated punch forces for the 3rawing of AA 3104-H19 are caused by the higher initial flow stress due to high prior cold forming. Additionally, the neglected heat dependency of the flow stress curve causes disproportional forces.
5.2. Multi stage ironing simulations According to manufacturers, the ironing is carried out in three subsequent steps with infinite spacing [3]. Thus, throughout the process, only one die is in contact with the cup. From Section 4 we know the
8
M Schi~nemann et al./Journal o f Materials Processing Technology .59 (1996) 1-9 140
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,
20 . . . . . . . . . . . . . i I
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Fig. 17: Load-stroke curves for the 1st ironing step with different ironing ratios and different materials. axial stresses for this a r r a n g e m e n t are l o w e r t h e n those for the simultaneous ironing. T w o sets of three ironing stages w e r e simulated. In the first set, the true strains w e r e -0.498, -0.223, and -0.182, while the true strains in the s e c o n d set w e r e constant-0.301. The reduction in the first step refers to the m a x i m u m wall thickness of the r e d r a w n cup. The thickness of the cup after the r e d r a w i n g is not u n i f o r m due to thickening. The m a x i m u m thickness of the cup wall is f o u n d at the c u p rim. Thus, the ironing ratio is not constant d u r i n g the first ironing. In the design of the ironing process, it is i m p o r t a n t to account for this fact. Also, the tolerances of the u s e d coil material h a v e to be considered. In addition, wall thickening a n d e v e n t u a l l y d i l a t a t i o n i n c r e a s e the a b s o l u t e i r o n i n g ratio a n d cause h i g h e r wall stresses and p u n c h loads. Load-stroke curves for the first ironing step with different i r o n i n g ratios and different materials are illustrated in Fig. 17. An increasing p u n c h load d u r i n g the first ironing step is a result of two effects; 1) the real r e d u c t i o n ratio increases w i t h the cup height, 2) the strain h a r d e n i n g of the cup material increases with the cup height. A f t e r the first i r o n i n g , the w a l l thickness is u n i f o r m o v e r the cup height. C o n s e q u e n t l y , the increase of the p u n c h force d u r i n g the second and third ironing is d u e to strain hardening. The ironing ratio is t h e n constant t h r o u g h o u t the process. Higher flow stresses in the u p p e r part of the cup wall result in higher temperatures, as seen in Fig. 18. Fig. 19 illustrates the t e m p e r a t u r e distribution in the tooling a n d the cup d u r i n g the first ironing step at a p p r o x i m a t e l y 4 / 5 of the ironing stroke. Fig. 19 refers to the i r o n i n g of A A 3003 w i t h an ironing reduction (~) of -0.498. The t e m p e r a t u r e distribution
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200
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Fig. 18: Temperatures of the selected points during the ironing simulations of AA 3003 with ironing ratios of e,/2/3=-0.498/-0.223/-0.182. (3bi~tl
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for the ironing simulation of AA 3003. was f o u n d to be representative for all simulations, while the absolute t e m p e r a t u r e s changed. Typically, the cup and the die show the highest t e m p e r a t u r e s in the d e f o r m a t i o n area, w h i l e the h i g h e s t p u n c h t e m p e r a t u r e is f o u n d b e l o w the d e f o r m a t i o n area closer to the cup bottom. H i g h p u n c h s p e e d causes h i g h sliding b e t w e e n cup a n d die. The sliding b e t w e e n cup and p u n c h is significantly lower, hence, the heat transfer b e t w e e n p u n c h and cup h a p p e n s m o s t l y in the area b e h i n d the die o u t l e t b y conduction. The tool area influenced b y the increased t e m p e r a t u r e s is v e r y small, s o m e m i l l i m e t e r s at m a x i m u m , but the tooling is expected to w a r m up if more then one can is produced.
M Schfgnemann et al/Journal of Materials Processing Technology 59 (1996) 1-9
60
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Fig. 20: Axial stresses of selected points for the ironing simulation of AA 3003 with the reduction ratios of E1/2/3=0.498/-0.223/-0.182. The axial stresses calculated for the ironing of AA 3003 are shown in Fig. 20. As expected, the axial stress in the point located at the cup bottom stays very low, while the stresses at the other points raise after the ironing die passes their location. Once the ironing die passed the location of point 2, the axial stress increases to its m a x i m u m until the die loses contact with the cup. The scatter in the curves is c a u s e d by elements leaving and entering the deformation zone. References
[1] Altan, Oh, S., and Gegel, H., Metal FormingF u n d a m e n t a l s and Applications, A m e r i c a n Society of Metals, Metals Park, Ohio (1983). [2] Odell, E. I., "A Study of the Wall Ironing by the Finite Element Technique", J. of Eng. for Ind., vol. 100, pp. 31-36 (1978) [3] Sodeik, M., Taeffner, K., and Weber, F., " F u n d a m e n t a l s of M o d e r n Can Making and Materials D e v e l o p m e n t for Two-Piece Can Manufacturing", Transactions of lSU, vol. 28, pp. 672-677 (1988) [4] Sheet Metal Industries, "Deep Drawing and Ironing of Non-Ferrous Material", Sheet Metal Industries, vol. 65, pp. 284-286, June (1988)
[5] Tjotta, S., and Heimlund, O., "Finite Element Simulation in Cold Forging Process Design", J. of Mat. Proc. Tech., vol. 36, pp. 76-96 (1992) [6] Danckert, J., "The Residul Stresses in the Wll of a Deep D r a w n and Ironed Cup D e t e r m i n e d Experimentally and by FEM", CIRP Annals, vol. 43/1, pp. 249-252 (1994) [7] Takeuchi, H., "Numerical Simulation Technology for Lightweight Aluminum Cans", J. of Mat. Proc. Tech., vol. 38, pp. 675-678 (1993) [8] Kampus, Z., and Kuzman, K., "Analysis of the Geometrical Accuracy in Ironing", A d v a n c e d Technology of Plasticity, Proceedings of the 4th Int. Conf. on Tech. of Plasticity, vol. 2, pp. 10051010 (1993) [9] Sachs, G., and Espey, G., "Effect of Spacing between Dies in the Tandem Drawing of Tubular Parts", Trans. of the ASME, February (1947) [10] Baudelet, B., and Grange, B., "Damage in Deep Drawn and Ironed Can Bodies in an Aluminum Alloy", Scripta Metallurgia et Materialia, vol 26, no. 3, pp. 375-379 (1992) [11] Xiaozhen, M., Yuying, Y., and Yingchun, L., "Experimental Study on Ironing of Stainless Steel and Optimization of Process Parameters", Advanced Technology of Plasticity, Proceedings of the 4th Int. Conf. on Tech. of Plasticity, vol. 3, pp. 1653-1656 (1993) [12] Merchant, H. D., Hodgson, D. S., O'Reilly, I., and Embury, J. D., "Structure and Property Evolution During Drawing and Wall Ironing of AA 3004", Materials Characterization, vol. 25, pp. 251-261 (1990) [13] Rajagopal, S., and Misra, S., "Measurements of Differential Friction Coefficients in Ironing", Proceedings of the 18th NAMRC, pp. 89-95 (1990) [14] Panknin, W., "Grundlagen des Tiefziehens zur Herstellung Zweiteiliger Dosen", Werkstatt und Betrieb, vol 110, no. 5, pp. 313-319 (1977) [15] Tufekci, S. S., Ahmetoglu, M. A., Kinzel, G. L., and Altan, T., "Process Simulation for Can Manufacturing by Deep Drawing and Ironing", SAE Paper No. 950696, International Congress and Exposition, Detroit, Michigan, February 27March 2 (1995) [16]Saito, M., 1989, Saiki, H., and Kawai, N., "Experimental Analysis of Ironing of Thin Metal Cups", Trans. of the ASME, vol. 111, pp. 56-63, February (1989)