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Journal of Materials Processing Technology 84 (1998) 97–106

FEM simulations and experimental analysis of parameters of influence in the blanking process

M. Samuel

Faculty of Engineering , Mansoura Uni ersity, Post No. 35516 , Mansoura, Egypt

Received 8 July 1997

Abstract

This paper summarizes the results of simulating the blanking process based on the multi-purpose FEM code and MARC-2D.

The characteristics of this simulation are four node axisymmetric elements, rigid dies, and an updated lagrangian approach. The

work-hardening behaviour of the material has been derived from a relationship between the equivalent plastic strain, the

equivalent Von Mises stress, and the Vickers hardness in the shear zone. Several blanking simulations are performed and the

results compared with those obtained from an experimental study. The simulation results obtained under the effect of process

variables (e.g. punch–die clearance, tool geometries, and material properties) are in good agreement with the experimental results.

Furthermore, the load–stroke curves obtained from the simulation study show the same changes as those obtained from the

experimental study. However, a difference in the range of about 14% is noted. © 1998 Elsevier Science S.A. All rights reserved.

Keywords: Blanking process; Multi-purpose FEM code; MARC-2D

1. Introduction

In the blanking of metals, it is most important to

suppress the generation of fracture zones in the sheared

surface. The enhancement of the compressive stress in

the shear zones is effective in suppressing the fracture

or initiation of cracking [1].

The operation of blanking has been studied widely

with respect to sheared edge morphology [2,3], shearing

parameters optimization or forming defects [4,5]. More-

over, blanking is a constrained shearing operation thatinvolves elastic deflection, plastic deformation and frac-

ture of the work materials. The criteria for assessing the

quality of the sheared component are the finish of the

sheared surface and the deviation from the nominal

dimensions of the blank or the blank holder. Several

factors such as punch and die clearance, punch velocity,

and the mechanical properties of the materials under

investigation influence the constrained shearing opera-

tion. Some of these factors have been investigated

extensively [6–10]. However, the influence of the radius

on the tools has not yet been defined, in particular, theinfluence of the tool edge geometry on the shearing

force and the profile of the sheared surface on the

component. Recent investigations [11 – 15] have estab-

lished the relationship between the process signature,

the sheared surface profile, and the punch and die radii.

However, several aspects of constrained shearing (with

different shearing edge geometry of the blanking tool)

have yet to be evaluated: these aspects are reported

herein.

In this paper, experimental analysis using punches

and dies with different radii, for work materials under

different conditions are considered. To be competitive

in the global market, manufactures strive to produceparts at lower cost, in less time, and of higher quality.

The use of FEM simulation is increasing for investiga-

tion and optimizing the blanking process. It is possible

to drastically reduce the lead time of new parts and

products by proper implementation of a simulation

technique into development and research. Many time-

consuming experiments can be replaced by computer

simulations [16–19]. Therefore, highly accurate results

of sheet metal forming may be obtained by applying the

FEM simulation. In the work reported here, a sim-

plified simulation model will be created using MARC.This model describes the blanking process of a circular

blank using a 40 mm diameter die. Moreover, the FEM

0924-0136/98/$ - see front matter © 1998 Elsevier Science S.A. All rights reserved.

PII S0924-0136(98)00083-1

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M . Samuel / Journal of Materials Processing Technology 84 (1998) 97–106 98

model of the material under investigation is calculated

for two cases: one with a small die clearance (C /T )=

2%, and the other with a large die clearance (C /T )=

23%. In addition, to verify the model, a series of

blanking tests will be performed for the calculated

geometries.

2. Material model

A cold-rolled strip of aluminium killed (Al. K.) steel

is used in the simulation work. One batch of the

specimens was annealed for 2 h at 850°C and then

cooled in a furnace. The mechanical properties of the

test specimens are shown in Table 1.

During the blanking process, large equivalent plastic

strain occurs. It is assumed that the material is isotropic

and that yielding follows the Von Mises yield criterion.

The following equation is used to model the material

behaviour in MARC:

̄ =K ¯ n (1)

where ̄ is the effective stress, ¯ is the effective strain, K

is the material constant, and n is the strain-hardening

exponent. Eventually, the hardness and the equivalent

Von Mises stress can be determined by postulating a

linear relationship between stress and hardness [17] as:

Table 2

Geometric and kinamatic parameters investigated

CommentsDimensionsValueParameter

(C /T )Clearance % C = (D−d )/2

Punch velocity mm s−11.0

mm 2% ClearancePunch diameter d 1=39.900

d 2=38.850 mm 23% Clearance

Die diameter mmD=40.000Rp1, Rd1=Punch and die mm

radii 0.20

Rp2, Rd2= mm

0.40

0.10Friction coeffi-

cient

Fig. 2. Mesh concentration in the shearing zone.Table 1

Mechanical properties of the test specimens

Cold rolledParameter AnnealedaUnits

Ultimate tensile strength 733MPa 618

433630Yield strength Mpa

Material constant (K ) 483510Mpa

201206Modulus of elasticity (E ) GPa

0.510.46 — Strain hardening (n)

0.25 0.27Poisson’s ratio () —

Vickers hardness (H V 100) 152235 —

mm 2.5 2.5Nominal thickness (T )

a Annealing procedure: 850°C for 2 h and then slow cooling in the

furnace.

Fig. 1. Simulation model.

e=aH V+b (2)

where a and b are constants, H V is the hardness, and eis the Von Mises stress. Eq. (2) will be used in the strain

region from 0 to 1 with the data in Table 1. Based on

the linear relation between e and H V [17], and Eq. (2),

the hardness data can be transformed into equivalent

stresses. In this way a set of data concerning stress and

strain can be collected. These data will be substituted

into a model based on Voce given as:

eH V=

e−b

a (3)

Eq. (3) represents the final model used in the FEM

simulation.

After performing several preliminary simulation tests

to determine the reliability and repeatability of the

modified program, the following blanking parameters

are investigated: punch-die clearance, tool geometry,and material properties.

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Fig. 3. The blanking tool: (a) with sharp die and punch corners; (b) with radiused die and punch corners.

3. FEM model

The multi-purpose software package MARC served

to make the FEM model. The blanking process is

axisymmetric, which means that it requires only a 2-

D model. Thus, it was sufficient to model only one

half of the tooling. Because of the large displacements

that occur, the lagrangian approach combined with arezoning technique is adopted. In this way, a new

mesh is generated each time an incremental step is

taken, this being needed due to the considerable mesh

distortion caused by small deformation. Fig. 1 shows

the simulation model. The actual tooling was used for

blanking round discs of 40 mm diameter. The sheet

material is considered as a plastic object whereas the

punch and die are defined as rigid bodies. In correla-

tion with actual practice, a blank holder with an in-

denter was moulded for guiding and holding down

the stock. Depending on the process that is beingsimulated, a loaded blank-holder that clamps the ma-

terial whilst blanking could be easily implemented.

An indenter was used to eliminate the sudden release

of the elastic energy of the press subsequent to the

initiation of fracture. Table 2 summarizes the geomet-

ric and kinematic values used in the simulation.

The number of elements used in this work is very

critical for each simulation. A large number of ele-

ments increases the accuracy of the results but dra-matically increases the calculation time. Therefore, a

very dense mesh was defined in the area of the shear-

ing zone and relatively large elements for the remain-

der of the part (Fig. 2). A total of 6000 axisymmetric

quadrilateral elements were used.

During the last phase of the blanking process, duc-

tile fracture occurs. A model was used to describe

ductile fracture that is based on the nucleation,

growth, and coalescence of voids in the matrix mate-

rial. Many experimental studies to establish ductile

fracture criteria in order to calculate the formabilitylimits of different materials have been reported in the

literature [18– 20]. Gurson emphasised the stage of

void growth in his original model reported in [21].

Tvergaard [22] improved Gurson’s model by taking

into account the effect of void interaction. Void nu-

cleation (such as in the works of Chu and Needleman

[23]) and coalescence (such as in the work of Tver-

gaard and Needleman [24–26]) have also been incor-

porated into the original model. Gurson’s model in

various forms has recently been applied to metal

forming processes which include sheet forming and

blanking processes [23]. In this work, Gurson’s mixedhardening model is used. This model is expressed as:Fig. 4. Definition of the portions of the sheared surfaces.

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Fig. 5. Experimental load–stroke curves, for different tool geometry, clearance, and material conditions.

=3

2

eF

2+2 fq1 cosh

3B m2F

− (1+q1 f

2) (4)

where F is the equivalent stress for a mixed material, f

is the void fraction, q1 is the constant to improve theresponse of the original Gurson model, which is taken

as q1=1.5 based on Tvergaard [26], and B m is the

spherical portion of shifted stress. Evolution of voids is

assumed to consist of growth and nucleation portions

expressed as:

f = f growth+ f nucleation

where:

f growth=3(1− f )ē p (5)

in which ē

is the mean plastic strain rate. For theimplementation of Eq. (5) in MARC, a strain-con-

trolled nucleation is assumed, then:

f nucleation= f

S N 2exp −

1

2

¯ p− NS N

2n (6)

where f

is the volume fraction of void-forming parti-

cles, S N is the S.D. in the equivalent plastic strain rate,

¯ p is the accumulated plastic strain, and N is the mean

equivalent plastic strain. It is difficult to determine the

material parameters for such a model because they are

outside the scope of this study. Therefore the data

reported in [25,26] will be used in this study. Onlyparameter has been determined from testing the

microstructure, by counting the level of inclusions of

the material. Finally, a higher value of S.D. S N repre-

sents a material with a more homogeneous distribution

for void nucleation, so that voids nucleation also occursin the case of large strains.

4. Experimental procedure

A schematic diagram of the blanking tool is shown in

Fig. 3 whilst the blanking process parameters are given

in Table 2. The blank holder with an indenter shown in

Fig. 3 (b) was used to elimination the sudden release of

the elastic energy of the press subsequent to the initia-

tion of fracture. A load cell and a displacement trans-ducer were used to measure the blanking force and

punch displacement, respectively. A universal Instron

machine was used at a constant ram speed of 1.0 mm

s−1 for all tests.

5. Results and discussion

Fig. 4 shows that the shear edge contains four dis-

tinct zones, these being indentation; burnished surface;

fracture surface; and burr zones.

Examples of blanking force versus punch displace-ment for materials tested for different tool geometry

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Fig. 6. Influence of the tool geometry, the punch–die clearance, and the material condition on the equivalent Von Mises stress: (on left) cold

rolled; (on right) annealed; for: (top) Rp, Rd=0.2 mm and C /T =2%; (middle) Rp, Rd=0.2 mm and C /T =23%; and (bottom) Rp, Rd=0.4 mm

and C /T =2%.

and clearance are shown in Fig. 5. The signatures

show that the maximum blanking force and punchdisplacement at crack initiation and the load required

to separate the blank from the stock are sensitive to

tool geometry, clearance, and condition of the mate-rial.

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Fig. 7. As for Fig. 6, but for total equivalent plastic strain.

Figs. 6 and 7 show the equivalent Von Mises stress

and total equivalent plastic strain for different values of

punch and die radii, and clearance, for two materialconditions described earlier. The maximum value of the

stress and strain are 62.7 MPa and 1.84 mm mm−1 for

cold rolled Al. K. steel for Rp, Rd=0.2 mm and

C /T =2%, whilst the stress and strain values are de-creased by about 1.8 and 35%, respectively, when the

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Fig. 8. Influence of the tool geometry, the clearance, and the work material on the experimental cross-section of the blank: (top) cold rolled Al /

k. steel; (bottom) annealed Al. k. steel; (left) Rp, Rd=0.2 mm and C /T =2%; (centre) Rp, Rd=0.2 mm and C /T =23%; and (right) Rp, Rd=0.4

mm and C /T =2%.

punch and die radius increases for cold rolling at the

same value of clearance (C /T =2%). Also, the value of

Von Mises stress is decreased by about 3% when the

clearance increases (C /T =23%). In the above men-

tioned explanation, it is noted that the values of Von

Mises stress for both of the materials are affected by

the clearance and the tool geometry, Moreover, con-

cerning the influence of the clearance on the totalstrain, it was found that the plastic strain increases as

the clearance increases for each cold rolled and an-

nealed Al. k. steel (Fig. 7). This is a consequence of

plastic deformation of the surface material in the shear

zone. As regards the effect of the material properties on

the Von Mises equivalent stress and plastic strain, it is

clear that the Von Mises stress and plastic strain reduce

by about 8–20% and 2–16%, respectively, compared to

the cold rolled Al. K. steel (Figs. 6 and 7). The form of

crack propagation that dominates the final shape of the

sheared edge and the specific shearing resistance isshown in Fig. 8. From this figure the influence of the

tool geometry and clearance on the accuracy of the

sheared edge is seen clearly for each material. The

increase of the punch and die radius causes increased

plastic deformation before shearing starts. The result is

an increased zone of indentation (a rounded edge) and

burr height, whilst the shear zone (the burnished sur-

face) decreases for both the cold rolled and the an-

nealed Al. K. steel. Furthermore, Fig. 8 shows the

effect of mechanical properties on the accuracy of the

sheared edges from the aspect of the fracture surface or

the angle of the fracture zone and burr height. Theangle of the fracture zone and the burr height are

greater in the blanking of annealed steel than in the

blanking of cold-rolled Al. K. steel. From the above

explanation it can be said that the accuracy of the

sheared edge is dependent on the tool geometry, the

clearance, and the material specifications.

Finally, the effects of strain hardening, tool geome-

try, and clearance on the evolutions of void fractions

and coalescence based on Gurson’s work as improvedby Needleman and Tvergaard are shown in Figs. 9 and

10. Comparing the different cases in Fig. 9, it is noted

that the gradients along the sheared edges are consider-

ably higher in the calculated hardness than in the

experimental hardness data. This can be explained as

follows: hardness can not be measured extremely close

to the edges. In order to do so, extrapolation towards

the edges is necessary. Extrapolation, in this case by

means of the least squares method in combination with

singular value decomposition, entails risks. In order to

overcome this problem, a first-order polynomial fit isused. However, as a consequence, the gradients mea-

sured on the edges are inevitably smaller than the real

hardness gradients. It can also be noted from Fig. 9

that the strain-hardened volume and the average strain

hardening increase with increasing clearance.

Lastly, the effects of clearance on the growth of the

voids is shown in Fig. 10 In the actual simulation, for

the value of clearance (C /T =2%) the void growth is

much greater at an early stage compared with that for

the large clearance (C /T =23%). However, in practice,

void growth will lead to initial cracks. These initial

cracks stop growing very soon and do not initiatefurther cracks.

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Fig. 9. Calculated and measurement of the microhardness of materials tested with different tool geometry and clearance: (top) cold-rolled Al. k.steel (Rp, Rd=0.2 mm and C /T =2%); (middle) cold-rolled Al. k. steel (Rp, Rd=0.2 mm and C /T =23%); and (bottom) cold-rolled Al. k. steel

(Rp, Rd=0.4 mm and C /T =2%).

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Fig. 10. Influence of the clearance size on the growth of the voids at constant punch and die radii: (on left) C /T =2%; and (on right) C /T =23%.

6. Conclusions

From the analysis of the blanking process under

several factors such as punch and die radii, die clear-

ance, and material properties, using FEM simulations

and comparison with experimental tests, the following

conclusions may be drawn: (1) The punch penetration

prior to fracture increases with the increase of the

punch and die radii. Also, the values of equivalent Von

Mises stress and total plastic strain decreases as the

punch and die radius increase. (2) The plastic strain

increases when the clearance increases. (3) Large punch

and die radii suppress the initiation of cracks from the

shearing edge of the tool, increase the tool penetration,

the burr height and the roughness but reduce the extent

of the burnished surface. The angle of the fracture

zone and the burr height are greater in the blanking of

annealed than cold-rolled Al. k. steel. (4) The model

used to calculate crack formation is capable of deter-

mining crack initiation, but it is clearly not capable of

predicting crack propagation. Generally, the stimula-tion responds to parameter changes such as different

clearance, material specifications, and tool geometry

with a relatively good agreement with the experiment

results, the difference being about 14%. Due to the

importance of strain rate effects on the blanking pro-

cess, further study is being conducted, the results of

which will be published later.

Acknowledgements

The author wishes to express his sincere thanks anddeepest gratitude to Professor Dr Eng Jerzy Gronos-

tjski for valuable guidance during the work reported in

this paper. Thanks are also due to Dr Eng Z. Zimniak

for many discussions and to the staff of the workshop

of the Metal Forming Department, Institute of Ma-

chine Building Technology, Technical University of

Wroclaw, Poland.

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.

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