Proposal for an Experimental-Numerical Method for Friction Description in Sheet Metal Forming

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Strojniški vestnik - Journal of Mechanical Engineering 61(2015)6, 383-391 Received for review: 2015-01-03
© 2015 Journal of Mechanical Engineering. All rights reserved. Received revised form: 2015-03-30
DOI:10.5545/sv-jme.2015.2404 _Original Scientific Paper_ Accepted for publication: 2015-04-22

# Proposal for an Experimental-Numerical Method

# for Friction Description in Sheet Metal Forming

Trzepieciński, T. – Lemu, H.G.
Tomasz Trzepieciński 1 - Hirpa G. Lemu 2,*

1 Rzeszow University of Technology, Department of Materials Forming and Processing, Poland
2 University of Stavanger, Department of Mechanical and Structural Engineering and Materials Technology, Norway

_During the forming of draw pieces with complex shapes, such as car bodies, it is necessary to use dies with varying outline curvatures and_
_varying radii of edge fillets. In such conditions, using the cylindrical counter-samples for friction modelling on the die fillet surface does not fully_
_reflect the friction conditions on the edge of the die. In this article, a proposal for an experimental numerical method of friction determination_
_on the rounded profiles of punch edges is presented. The method involves a bending under tension (BUT) test that is devised to simulate_
_the frictional conditions over the punch radii. Numerical simulations of BUT tests were conducted with ABAQUS software. An elastic-plastic_
_approach with Hill’s (1948) anisotropic material formulation and isotropic strain hardening were considered. Results of both experimental and_
_numerical approaches confirmed the usefulness of the proposed method._
Keywords: bending-under-tension, BUT test, finite element method, friction, friction coefficient

Highlights

_•_ _An experimental-numerical method of friction determination on the rounded profiles of forming punch edges is proposed_

_•_ _A bending under tension test is devised to simulate the frictional conditions over the punch radii._

_•_ _The numerical simulations of bending under tension tests are conducted with ABAQUS software._

_•_ _The surface topographies of DDQ steel sheet at different strain states and lubrication conditions are studied._

0 INTRODUCTION

Frictional resistance in sheet metal forming depends
on physical and chemical factors, among others,
acting on the contact surface, the roughness and
topography of sheet and tools, the dynamics of the
loads, lubrication conditions, and the temperature

[1] and [2] . Many kinds of friction tests have been
developed to determine the coefficient of friction in
sheet metal forming. Among the available methods,
the most frequently used are the strip-drawing test,
the draw bead simulator friction test and bending the
under tension test.

The bending under tension (BUT) test is a
group of tests in which a work piece is stretched
and bent during the experiment so that the friction
coefficient is measured under conditions similar to
those encountered in the real forming process. The
traditional way of performing BUT tests is with
differential measurements in which two tests are
carried out after each other, one by drawing the work
piece over a fixed circular cylindrical tool-pin, the
other over a freely rotating pin, with the understanding
that no sliding takes place. The difference in front
tension measured in two tests gives an estimate of the
friction. One drawback of this method is the stochastic
variations, which may cause large scatter, and the fact
that steady-state conditions must be present while
measuring.

Many BUT tests have contributed to the
knowledge about sheet-forming tribology [3] and

[4] . A new methodology to determine the friction
coefficient via a BUT test, considering the pressure
non-uniformity, has been suggested by Kim et al. [5] .
Miguel et al. [6] used a BUT test to evaluate pressure
in the die radius and the corresponding lubricated
regime under multi-axial stress conditions.

Deng and Lovell [7] and Lovell and Deng [8]
studied the influence of stamping process parameters
and wear mechanisms on the characterization of
interfacial friction in coated steel sheets. In these
studies, it was determined that the lubricant properties
had little influence on the final surface roughness of
the deformed sheet. Considering the influence of the
pin size, the results showed that the friction coefficient
increased with decreasing pin radius values.

Intending to improve the chance of success
before production trials of more environmentally
benign tribo-systems, many researchers [9] and [10]
described the equipment for realization of BUT tests
and developed a methodology for off-line evaluation
of tribo-systems for sheet metal forming. Furthermore,
the BUT investigations conducted on the friction of
galvannealed interstitial free sheet steels by Garza
and Van Tyne [11] shown that the coefficient of
friction also increases as the amount of iron content
in the coating decreased. Alingner and Van Tyne

[12] examined the change in the die profile of five

*Corr. Author’s Address: University of Stavanger, 4036 Stavanger, Norway; Hirpa.g.lemu@uis.no 383

Strojniški vestnik - Journal of Mechanical Engineering 61(2015)6, 383-391

coated die materials tested in conjunction with three
automotive steel sheets. The zirconia die material
exhibited the least amount, but the most significant
change.

Based on the experimental results of the BUT
test, Wiklund et al. [13] developed a theoretical model
to predict that the effective hardness of a surface is
reduced by the presence of underlying plastic flow.
The results of investigations of Pereira et al. [14]
provide a qualitative description of the evolution
and distribution of contact pressure at the die radius
for a typical channel-forming process. The proposed
contact pressure behaviour for the BUT test was
compared to other results available in the literature.
Both experimental and numerical approaches to the
BUT test carried out by the authors [15] showed that,
although the sheet was drawn over a cylindrical pin,
the plane strain state assumed over the entire width of
the sample is affected by the behaviour of the sample
during bending. The stretch-drawing type tests (i.e.
BUT) are widely used to investigate the influence of
several material and technological parameters [16] and

[17] .

In this article, we propose an experimentalnumerical method of friction determination on a
punch edge with a rounded profile as a modification
of the classic BUT test method.

1 DESCRIPTION OF BUT TEST METHOD

The BUT test is performed in two steps. First, a strip
is drawn over a freely turning roller (Fig. 1), and the
values of both pulling _F_ 1 and back tension _F_ 2 forces
are determined. The difference between _F_ 1 _t_ and _F_ 2 _t_
forces is the force _Fb_ due to bending and unbending of
the sample over the roller. A second strip is then drawn
over a fixed roller, and the corresponding pulling _F_ 1
and back tension forces _F_ 2, are determined.

Assuming that there is a constant friction
coefficient _µ_ in the contact region and that the wrap
angle _γ_ (Fig. 1) is constant during the test according to
the equilibrium of all forces acting on an elemental cut
of the strip _dγ_, it can be shown that:

_F_ + _q_ µ _wRd_ γ     - ( _F_ + _dF_ ) =,0 (1)

_d_ γ _d_ γ
_qwRd_ γ - _F_ sin - ( _F_ + _dF_ )sin = 0, (2)
2 2

where _q_ is the unit normal contact pressure and _w_ is
the width of the strip.

2 MOTIVATIONS

Separation of frictional resistance and bending
resistance in the BUT test is only possible by using a
cylindrical counter-sample. The use of non-cylindrical
counter-samples does not allow carrying out the test in
terms of rotational counter-sample.

Assuming the plane strain state of the sample in
the theoretical analysis of friction tests is, therefore,
a simplification. During the forming of draw pieces
with complex shapes, such as car bodies, there are
dies with varying outline curvatures ( _R_ ) and varying
radii of edge fillets ( _r_ ) (Fig. 2).

Fig. 1. _Forces acting on an elemental cut of the strip_

For a very small _dγ_ one can assume that

γ _d_ γ
sin _d_ ≈ and _dF_ << _F_ . Thus, combining Eq. (1)
2 2

and Eq. (2) gives:

_dF_
µ _d_ γ = . (3)

_F_

Integrating Eq. (3) and taking into account
_γ_ = _π_ /2, the coefficient of friction is determined to be:

µ= 2 ln( _F_ 1 ). (4)
π _F_ 2

Formulas given by several authors take into
consideration roller radius _R_ and sheet thickness _g_ .
For a 90° bend angle, the friction coefficient _µ_ can be
determined from the formulae:





_R_ + _g_
_R_


 

2  _R_ +

µ=
π 



 _F_ 1 −






/ 2 ln  _F_ 1 − _Fb_ , (5)

   

2  _R_ + _g_ / 2 ln  _F_ 1 − _Fb_ 

π  _R_    _F_ 2 

_F_ 1 − _F_
_F_ 2

384 _Trzepieciński, T. – Lemu, H.G._

Strojniški vestnik - Journal of Mechanical Engineering 61(2015)6, 383-391

Fig. 2. _Example of a die shape for sheet forming with varying_
_outline curvature (_ _Ri_ _) and varying radii of edge fillet of die (_ _ri_ _)_

In such conditions, using a cylindrical countersample for friction modelling on the die fillet surface
does not fully reflect the friction conditions on the
edge of the die, which outline changes along the
perimeter of the die hole. During the standard BUT
test, the flat sample is bent in the plane perpendicular
to the axis of the counter-sample (Fig. 1). Analysing
the simplest case of forming of the axisymmetric
draw-piece, it is seen that tri-axial stress occurs state
on the die edge with additional peripheral compressive
stresses (Fig. 3). The tri-axial state of stress in the
contact area of the sheet with the non-cylindrical
counter-sample makes the nature of the changes in the
surface topography of the sheet closer to reality than
when using a cylindrical counter-sample. The stress
state at the rounded edges of the tools, depending on
the geometry of these edges, has a significant effect
on the nature of the plastic deformation of sheet over
the rounded die and the change of its topography.

Fig. 3. _Stress state during forming of axi-symmetrical draw piece;_

_σr_ _,_ _σt_ _,_ _σn_ _- stresses: radial, tangential and normal, respectively_

3 METHOD OF FRICTIONAL RESISTANCE DETERMINATION

A major issue is the development of a realistic test
that allows the separation of the frictional resistance
and the bending resistance of the sample on rounded

profiles of the die and the punch. The proposed
method provides the ability to determine the
coefficient of friction at the rounded edge of the tools.
The formula for determining the friction coefficient is
in accordance with the method for cylindrical countersamples. The shape of the proposed counter-samples
makes it necessary to use a different method for
determining the test forces in the condition of friction
elimination via the following steps:

- experimental determination of pulling and back
tension forces during the friction test using the
counter-sample with the non-cylindrical profile,

- numerical determination of the pulling and back
tension forces under frictionless conditions
( _µ_ = 0) using the counter-sample with the noncylindrical profile.
The proposed method and counter-sample shapes
for friction coefficient evaluation in the BUT test are
not covered in the literature. Assuming that, during
the friction test, the sample is bent at an angle of
90°, the coefficient of the friction is determined by
modification of Eq. (5) as:








_R_ + _g_
_R_


 

_F_ 2

2  _R_ +

µ=
π 



 _F_ 1exp - ( _F_ 1 _num_ - _F_ 2 _num_ )






_F_ 1exp - ( _F_ 1 _num_ - _F_ 2

/ 2   _F_ 1exp - ( _F_ 1 _num_ - _F_ 2 _num_ ) 

  ln,

(6)

2  _R_ + _g_ / 2   _F_ 1exp - ( _F_ 1 _num_ - _F_ 2

exp

where _R_ is the radius of countersample rounding, _g_ the
sheet thickness, _F_ 1 _exp_ the experimentally determined
pulling force, _F_ 1 _num_ the numerically determined
pulling force, _F_ 2 _exp_ the experimentally determined
back tension force, _F_ 2 _num_ the numerically determined
back tension force.

During the experiment, the sheet metal is drawn
on the stationary counter-sample with the convex
(Fig. 4a) or the concave (Fig. 4b) profile.

Fig. 4. _Proposed methods of determination of friction coefficient_

_on die edge with a) convex and b) concave profile_

_Proposal for an Experimental-Numerical Method for Friction Description in Sheet Metal Forming_ 385

Strojniški vestnik - Journal of Mechanical Engineering 61(2015)6, 383-391

The greater the curvature of the tool edge, the
more the strain state occurring in the contact area
deviates from the plane strain state occurring in a
classic BUT test. The value of the bending force
increases nonlinearly with increasing degree of
deformation of the sheet resulting from the strainhardening phenomenon. To separate the frictional
and bending resistances, the pulling force must be
determined in conditions that eliminate the friction
between the sheet and the counter-sample.

4 NUMERICAL SIMULATION OF THE PROPOSED APPROACH

The shape of the die and the punch profile determines
the occurrence of the specific distribution and the
value of strains in the sample, and the distribution
of contact shear stress. To compare the mentioned
parameters, the numerical simulations for the three
profiles of counter-samples (cylindrical, convex
and concave) were carried out. The radius of the
die edge rounding was 7.5 mm, and the radius of
the curvature for counter-samples with concave and
convex contours was 40 mm. The material is assumed
to be deep drawing quality (DDQ) steel sheet with
a thickness of 1.5 mm. The width and the length of
the sample were 85 mm and 11 mm, respectively.
The sample was modelled with eight-node reduced
integration, general-purpose linear brick elements,
called C3D8R in ABAQUS terminology. Kim et
al. [18] and Ramezani et al. [19] successfully used
these types of elements to study the contact pressure
distribution in BUT test.

Table 1. _The mechanical properties of DDQ steel sheet_ [20]

An elasto-plastic material model approach has
been implemented. The plastic behaviour of the metal
is described by the von Mises yield criterion. Further,
an isotropic strain hardening described by Hollomon’s
law with _C_ and _n_ parameter values according to the
values in Table 1 are assumed. The rest of mechanical
parameter values are assumed as follows:

- Young’s modulus _E_ = 210000 MPa,

- Poisson’s ratio ν = 0.3,

- density _ρ_ = 7860 kg·m –3 .
To analyse the influence of different countersample profiles on the strain and pressure distribution,
the contact phenomena in all cases are considered
the same. The friction properties are described by
Coulomb’s law of friction coefficient whose value is
equal to 0.2. One end of the sheet is restrained, while
a displacement equal to 4 mm is applied to the second
end.

The strain distribution for the third analysed
counter-sample profiles, corresponding to this
displacement, is shown in Fig. 5. The counter-sample
profile determines the distribution and values of
equivalent plastic strain of the sheet; for the concave
profile (Fig. 5b) the maximum pressure is greater,
whereas for the convex profile (Fig. 5c) it is less than
the maximum plastic strain obtained for the cylindrical
profile (Fig. 5a).

As depicted in Fig. 6, in the cross section of the
bend sample a neutral layer exists on which the sign
of the longitudinal strains is changed.

In the middle part of the sample width on the
internal side, there are longitudinal compressive

Ultimate tensile strength

_Rm_ [MPa]

Hardening coefficient

_C_ [MPa]

Hardening exponent

_n_

Lankford’s
coefficient

Sample
orientation

Yield stress

_Re_ [MPa]

0° 162 310 554 0.21 1.55
45° 163 322 542 0.20 1.27
90° 168 312 530 0.21 1.67
average value 164.3 314.7 542 0.207 1.497

Fig. 5. _Distribution of equivalent plastic strain after sheet elongation of 4 mm for edges of a) cylindrical, b) concave and c) convex profile_

386 _Trzepieciński, T. – Lemu, H.G._

Strojniški vestnik - Journal of Mechanical Engineering 61(2015)6, 383-391

stresses, and the strain increment is negative; on the
external side, the longitudinal stresses are negative,
and the strain growth is negative.

At the sample edge, the aforementioned stress
and strain state is disturbed (Fig. 6) by increasing
the longitudinal dimensions of the tensional layers
and reducing the corresponding dimensions of
the compressed layers of the sample. The use of
cylindrical counter-samples in the BUT test disturbs
the initial rectangular cross-section of the sample, and
the middle part of the sample width is also curved.

stage, which exhibits steady state contact conditions at
the die radius.

Fig. 6. _Stress and strain state during bending of sheet strip,_

_σ_ _– stress,_ _ε_ _- strain_

After the elongation of the sample by 2 mm, the
distribution of contact pressure at surface nodes (Fig.
7) is not uniform across the width of the sheet, even
when using a cylindrical counter-sample (Fig. 8). The
large ratio of the width to the thickness of the sample
(20/1) determines the occurrence of a specific stress
and strain state on the sample thickness.

There is a significant difference in the value of
the maximum pressure for the analysed profiles,
from about 146 MPa for cylindrical counter-sample
(Fig. 7a) to about 178 MPa for counter-sample with
concave profile (Fig. 7c).

Such a large difference in the pressure between
these profiles confirms the hypothesis of the existence
of different conditions of friction and necessitates
using the proposed method for determining the
coefficient. After the samples were stretched to 4
mm, the difference between the maximum values
of pressure for all counter-samples varied between
8 (Figs. 7a and 8a) and 15 MPa (Figs. 7c and 8c).
It is also noted that there is a clear difference in the
location of maximum pressure for the convex profile
(Figs. 7b and 8b) and concave profile (Figs. 7c and
8c). The increase of the sample strain implies that
contact pressure peaks appear near the entry and exit
regions of the strip (Fig. 8), which is in agreement
with the results of Kim et al. [5] . Contact pressure
evolution at the die radius in sheet metal forming is
studied by Pereira et al. [21], who identified three
distinct phases through the numerical analysis of the
BUT test: bending, an intermediate stage when the
region of the sheet that was deformed at the start of the
die radius has not reached the side-wall, and the final

Fig. 7. _Distribution of contact pressure [MPa] on the sheet_
_surface after sheet elongation of 2 mm for edge of a) cylindrical,_

_b) convex and c) concave profile; d) location of contact pressure_

_measurement_

Fig. 8. _Distribution of contact pressure (MPa) on the sheet_
_surface after sheet elongation of 4 mm for edge of a) cylindrical,_

_b) convex and c) concave profile; d) location of contact pressure_

_measurement_

5 VERIFICATION OF METHOD

5.1 Experimental Procedure

The schematic view of the test device is shown
in Fig. 11. A test strip was held at one end in a grip
supported by a load cell. A specimen of 8 mm width

_Proposal for an Experimental-Numerical Method for Friction Description in Sheet Metal Forming_ 387

Strojniški vestnik - Journal of Mechanical Engineering 61(2015)6, 383-391

and 1 mm thick (no. 3 in Fig. 9) is wrapped around
a cylindrical fixed roll with diameter of 20 mm and
loaded in a tensile testing machine ensuring contact
at an angle of 90°. The application of the fixed pin
allows setting up the rolls in four positions to utilize
the full circumference of the roll. The test was carried
out using the roll made of tool steel with roughness
qualities _Ra_ = 0.32 µm measured parallel with the roll
axis.

The forces _F_ 1 and _F_ 2 were measured
simultaneously during the test. A major advantage of
this test apparatus is that strain does not have to be
measured to determine the coefficient of friction. For
some tests, the effect of strain on the value of the
coefficient of friction may be of interest.

Fig. 9. _Schematic view of testing device; 1 – machine base,_
_2 – device frame, 3 – specimen, 4 and 5 – tension members,_
_6 – working roll, 7 - fixing pin, 8 and 9 – strain gauges_

5.2 Numerical Modelling

We simulated a numerically frictionless BUT test
by assuming that the friction coefficient value
between the sheet and the counter-sample was _μ_ = 0.
The sample was modelled with eight-node reduced
integration C3D8R brick elements. To the best
possible description of the frictionless conditions of a
real BUT test, Hill’s anisotropic (1948) [22] material
formulation with isotropic strain hardening was
implemented in the finite element (FE) model. The
rest of the parameters and the mechanical property
values are given in Section 4.

6 DISCUSSION OF RESULTS

The values of both the pulling _F_ 1 and back tension
_F_ 2 forces increase linearly after sample yielding and
then increase proportionally (Fig. 10). Furthermore,
the relation between forces and the friction condition
does not change for a sheet’s true strain above about

0.15. In the case of small contact pressures (true strain
below 0.15), the friction coefficient increases with
a true strain, which is in agreement with results of
Ramezani et al. [19] . Friction coefficient of materials
with little hardening capacity quickly increases at
relatively small pressure. An additional comparison
of pulling and back tension forces determined for
different roughness qualities _Ra_ (0.32 µm, 0.63 µm
and 2.5 µm) and friction conditions (dry – dry friction,
oil – lubrication conditions) shows that the _F_ 1/ _F_ 2 ratio
depends on pin roughness and friction conditions. The
blue line in Fig. 11 represents when both forces are
equal, i.e. _F_ 1 = _F_ 2. The _F_ 1/ _F_ 2 ratio value increases
faster for higher values of pulling force.

Fig. 10. _Values of forces during friction tests conducted in fixed_

_roller conditions_

For bending force value evaluated as the
difference between pulling and back tension force
values in the case of a freely rotating pin, similar
characteristics are observed (Fig. 12). The freely

Fig. 11. _Comparison of pulling F1 and back tension F2 forces_

_determined for different roughness qualities_
_Ra (0.32, 0.63 and 2.5 µm) and friction conditions_

_(dry – dry friction, oil – lubrication conditions)_

388 _Trzepieciński, T. – Lemu, H.G._

Strojniški vestnik - Journal of Mechanical Engineering 61(2015)6, 383-391

rotating pin did not eliminate the friction between
the counter-sample and its fastening, so the bending
force value varied during the test. In the case of an
FE model, there are small variations of bending force
in the whole range of values of sample true strains,
because it is possible to eliminate the friction in the
test in the numerical model. This is one advantage of
the numerical evaluation of bending force in a BUT
test.

Fig. 12. _Comparison of variation of bending force value_

_determined experimentally and numerically_

The increasing of sample strain allows an
intensification of flattening of surface peaks. The
scratches parallel to the sliding direction allow the
identification of abrasion by grooving as the main
contact mechanism. Simultaneously, it leads to a
reduction of the volume of the valleys in the surface
topography, which are mainly working as oilreservoirs (Fig. 13).

The effectiveness of the supply of lubrication at
the contact zones described by the WC index is studied
by Wihlborg and Crafoord [3]. The increasing of real
contact area is related to peaks flattening and the

change of the surface topography because of sample
stretching. Furthermore, during the deformation of
the sheet metal over a tool, contact occurs only at
the peak asperities of both surfaces, as concluded
by Figueiredo et al. [23]. The frictional response in
a BUT test under mixed lubricated condition was
measured by Wiklund et al. [24], who found the larger
the number of oil pockets, the lower the friction. The
lubricant is squeezed out on the contact spots and
thus the friction is decreased. The contact conditions
of anisotropical surfaces are probably not highly
dependent on the sliding direction, but the increasing
of active oil pockets (Fig. 13) can considerably
decrease the friction. Podgornik et al. [25] and

[26] investigated and compared different surface
modification techniques in terms of galling properties
when applied to forming tools. They concluded that
the galling tendency can be greatly reduced by proper
polishing of the tool surface. Based on the results of
friction tests in different friction conditions, Sedlaček
et al. [27] have also found the most important
surface roughness parameters describing work-piece
topography.

Based on the process forces of the experimental
BUT test realized in fixed roller conditions and
bending forces determined by experiment and
numerical method, the friction coefficient value (Fig.
14) was determined by using Eq. (6). For small values
of the sample true strain, a transient region of variation
of the friction coefficient value is observed, because
the used strain gauges can produce inaccurate forces
values in the case of small measured values of forces.

In the range of the true strain of 0.01 to 0.04,
the average friction coefficient value determined
experimentally equals about 0.15. The test forces in
this range increase continuously but the difference
between them ∆F is quite similar ∆ _F_ 1 ~ ∆ _F_ 2 ~ ∆ _F_ 3 (Fig.
12). This relation is in excellent agreement with the

Fig. 13. _Surface topography of DDQ steel sheet: a) original, b) after sheet true strain 0.15 and c) after sheet true strain 0.3; oil lubrication_

_conditions: 1 – oil pockets, 2, 3 – scratches_

_Proposal for an Experimental-Numerical Method for Friction Description in Sheet Metal Forming_ 389

Strojniški vestnik - Journal of Mechanical Engineering 61(2015)6, 383-391

results of Fratini et al. [28], who studied the influence
of different lubricants on the reduction of friction in
the BUT test. The main impact on the value of the
friction coefficient for the analysed counter-sample
shape has the values of pulling and back tension
forces determined in case of the fixed pin. To take
into consideration in Eq. (6) the numerically evaluated
bending force F _bnum_ = _F_ 1 _num_ - F2 _num_ allows preserving
the trend of changes of the friction coefficient value
(region A in Fig. 14).

Fig. 14. _Variation of friction coefficient value determined_

_experimentally and numerically_

7 SUMMARY

The proposed concept in this article allows the
determination of friction coefficient in the case of dies
with varying outline curvatures and varying radii of
edge fillets. The state of stress at the rounded edges
of the die and punch is determined by the geometry of
these edges and has a significant impact on the plastic
deformation nature of the sheet and consequently
changes its topography.

The counter-sample profile determines the
distribution and values of equivalent plastic strain
of the sheet: for the concave profile the maximum
pressure is greater, whereas for the convex profile
there is less maximum plastic strain obtained for
the cylindrical profile. Different friction conditions
of all analysed profiles are also determined by the
distribution and value of contact pressure. The
difference in the value of maximum contact pressure
depends on the sample elongation and varies between
10 MPa and 30 MPa for all analysed die rounded
profiles.

The values of pulling and back tension forces
determined in the case of the fixed pin have the main
impact on the value of friction coefficient for the
analysed counter-sample shape. The bending force

value has a smaller effect on the accuracy of friction
coefficient value determination in the BUT test. The
experimental investigations and numerical approach
confirmed that the proposed experimental-numerical
method can be useful for the determination of the
friction coefficient in the rounded profile of dies.

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