simulation and experiments to evaluate edge …forming technology forum 2019 september 19 – 20,...

Forming Technology Forum 2019 September 19 – 20, 2019, utg, TUM, Germany

* Corresponding author: 1971 Neil Avenue, Room 339, +1 614 292 9267, [email protected]

SIMULATION AND EXPERIMENTS TO EVALUATE EDGE

FRACTURE USING THE DOUBLE BENDING TEST

David Diaz-Infante*, Berk Aykas, Advaith Narayanan and Taylan Altan

Center For Precision Forming, The Ohio State University,

Columbus, Ohio, USA.

ABSTRACT: Edge fracture is a common problem when forming AHSS. Different methods exist to eval-

uate edge fracture: the Hole Expansion, Half Dome, Collar Forming, Edge Fracture Tensile and Double Bend-

ing tests. Each test has advantages and disadvantages regarding practicality, time, cost and emulation of

flanging conditions observed in sheet metal components.

This study aims to suggest a practical methodology to evaluate edge quality for various AHSS, including

GEN3 materials, NEXMET1000/1.26 mm and 980xG3/0.98 mm; the zones of edges trimmed with different

cutting clearances are evaluated (e.g. shear zone). Double Bending Tests (DBT) are investigated using FE

simulations to suggest an adequate tool configuration and the corresponding sample size. Finally, an attempt

is done to correlate edge stretchability to strain hardening along the cut edge, calculated by FE simulation.

Thinning and strains at fracture are reported and can serve as a reference for edge fracture prediction in simu-

lation. The DBT is proposed as an alternative for edge stretchability since this technique is not affected by

blank/tool friction or necking; furthermore, in DBT, trimming along a straight line is used. Consequently, the

cutting clearances are presumably easier to maintain more uniformly, as compared to edge fracture tests where

round piercing is used.

KEYWORDS: Edge fracture, Advanced High Strength Steels, Double Bending Tests,.

1 INTRODUCTION

The increase in use of Advanced High Strength

Steels (AHSS) in the automotive industry comes

with manufacturing challenges; a particular case is

the edge fracture phenomenon. This type of fracture

is generated by tensile stresses along the edge of the

material during some sheet metal forming processes

(e.g. flanging). Significant efforts have been made

in order to predict and prevent edge fracture and

hence, several testing methods are currently availa-

ble to estimate the limits of a given edge under ten-

sion.

The Hole Expansion Test (HET), described by

the ISO standard 16630 [1], is probably the most

popular among these testing methods; despite a sig-

nificant scattering in Hole Expansion Ratio (HER)

results reported by different researchers [2,3]. The

main challenges of the HET lie on the edge prepara-

tion (e.g. tool conditions or uniform cutting clear-

ance along the hole perimeter); moreover, the frac-

ture detection during the expansion process intro-

duces an additional significant error (e.g. using

high-speed cameras coupled to the punch stroke or

simply detected by the operator in real time).

The Collar Forming Test (CFT), described by Braun

et al. [4], serves as a variant of the HET. The CFT

allows the flexibility required regarding the cutting

clearance and initial hole size, limited in the HET by

the ISO16630 to 12% of sheet thickness and 10 mm

respectively. This is especially important since, as

pointed out by Larour et al. [5], the best cutting

clearances for some AHSS can be higher than 12%.

Moreover, the CFT also reduces the fracture detec-

tion issue, since this action is taken only after the

collar has been formed as opposed to the HET where

the fracture must be detected accurately during the

punch stroke. Nevertheless, the challenges on the

edge preparation remain, even when the tool stiff-

ness can be improved by using a larger initial hole

size and thus, a bigger and stiffer die set.

Some additional testing options are the Half Dome

Test (HDT) [6], the Edge Fracture Tensile Test

(EFTT) [7], and the Double Bending Test (DBT)

[8]. These tests have the advantage of an edge pro-

duced by trimming along a straight line; hence, the

cutting clearance can be controlled better. The HDT

is gaining popularity due to its relative simplicity re-

garding sample preparation. Nevertheless, the out-

of-plane type of deformation on the edge produced

during this test may not match the one observed in

operations where edge fracture is common. The use

of EFTTs has also increased; Golovashchenko and


Ilinich [7] proposed a half-dog-bone specimen ge-

ometry, which was later simplified to a rectangular

geometry by Ilinich et al. [9]. Feistle et al. [10] used

more sophisticated, but also more expensive method

to produce, dog-bone specimens. The EFTT elimi-

nates the possible effect of friction in a HDT; how-

ever, in some cases, especially for relatively ductile

materials, necking occurrence can be observed be-

fore the edge fracture and therefore, the use of EFTT

is limited. The DBT eliminates the necking issue,

according to Bouaziz et al. [8], and the sample prep-

aration is relatively simple, such as in the HDT. The

DBT leads to a fracture similar to a Side Bending

Test (SBT), described by Yoshida et al. [11], while

reducing the amount of additional tooling to prevent

buckling on the specimen during the tests. To the

best of the author’s knowledge, the literature related

to the DBT is limited to publications by Bouaziz et

al. [8] and Dietsch et al. [12] and there is no detailed

information regarding the best practices to conduct

the DBT.

One thing that is common to all the aforementioned

methods, and to mechanical cutting processes in

general, is the importance of maintaining a uniform

cutting clearance along the cut perimeter. As it can

be inferred, the nominal clearances may vary under

loading conditions. Thus, to find a method that can

provide an estimation of the actual cutting clearance

during the trimming process is of great interest. A

typical methodology, suggested by Larour et al. [5],

is to measure the distance between the possible last

location of the punch and die from a micrograph as

shown in Fig. 1a; this can be a time-consuming

method and therefore not common in practice out-

side of a laboratory. Larour et al. [5] have also illus-

trated the usefulness of a portable USB microscope

to evaluate the cut edges without major equipment

requirements, Fig. 1b.

Fig. 1 a) Use of micrographs to determine the ef-fective local cutting clearance (cutting clear-ance under loading conditions) and b) use a portable USB-microscope for quick detec-tion of edge defects [5].

The objective of this study is to extend the

knowledge regarding the DBT while evaluating the

edge stretchability of GEN 3 AHSS, 1.26 mm thick

NEXMET1000 and 0.98 mm thick 980xG3, com-

pared to a reference 1.21 mm thick Dual-Phase steel

with 980 MPa Ultimate Tensile Strength (UTS). Ad-

ditionally, the authors seek to illustrate the difficul-

ties of characterizing AHSS by its geometry due to

variations along the trimmed length.

2 APPROACH

The Double Bending Test (DBT) was selected as the

method to evaluate and compare the stretchability of

GEN 3 AHSS against a reference DP980 AHSS;

mechanical properties of the selected materials are

listed in Table 1.

The DBT in Fig. 2, described by Bouaziz et al. [8],

consists of an initially rectangular sample which

should be bent, in a first operation, 90 degrees along

a line parallel to the edge to be evaluated. Later, a

second bending operation perpendicular to the first

bending is necessary. This second operation leads to

tensile strains along the edge to be analyzed and thus

to edge fracture.

Table 1: Mechanical properties for different AHSS, including GEN 3 AHSS. Properties obtained along Rolling and Transverse direction using tensile tests. Values reported are average of three samples tested using the ASTM E8 standard [13].

Fig. 2 Schematic of a Double Bending Test (DBT) by Bouaziz et al. [8]. a) First bending oper-ation and b) second bending operation.


2.1 DETERMINING THE ADEQUATE DBT

SAMPLE SIZE AND TOOLING

CONFIGURATION

In order to expand the information provided by

Bouaziz et al. [8] and Dietsch et al. [12], a group of

Finite Element (FE) simulations was conducted us-

ing shell elements and rigid tools modeled in PAM-

STAMP software. DP980/1.21 mm material was

used as an example in simulations and the trends

found are assumed to be valid for every material.

The coefficient of friction was kept as 0.1 in simu-

lations since it did not have any effect on the thin-

ning and forming force calculations (tested range

from 0.06 to 0.14). The flow stress curves for the

tested materials was obtained from tensile tests until

the uniform elongation and later extrapolated using

the Hollomon’s law; hence, the simulations are used

only as a guideline and some differences between

simulation and experimental results are expected

due to this approximation.

For a proposed initial blank size of 80 mm by 50

mm, variations were simulated for the opening be-

tween lower dies (𝐶𝑑), the punch radius (𝑅), sheet

thickness (𝑡) and the flange height (ℎ), Fig. 3; vari-

ations for the bending radius of the first operation

(𝑅𝑑1) and the sliding radius of the lower die (𝑅𝑑2)

were not included. As indicated in Table 2, when the

die opening increases, the punch force decreases but

the buckling increases. A decrease in flange height

will lead to a lower punch force and to lower possi-

bilities of buckling on the sample flange; however,

depending on the material, thickness and tools avail-

able, the first bending operation can become very

difficult for small flanges due to the small leverage.

Hence, a flange height (ℎ) of 10 mm was selected as

a balanced value. A reduction in punch radius has

no effect upon the punch force while a reduction of

sheet thickness may lead to more buckling.

Fig. 3 Schematic of some of the parameters in-volved in the Double Bending Test (DBT).

The following combination of parameters was se-

lected considering the equipment and tooling avail-

able:

Die Opening (𝐶𝑑) = 20 mm

Flange Height (ℎ) = 10 mm

Punch radius (𝑅) = 4 mm

Table 2: Simulation Matrix.

2.2 SAMPLE PREPARATION

2.2.1 Trimming

A 600 Ton Verson mechanical production press was

used to trim five rectangular samples along 80 mm

(𝐿) perpendicular to the rolling direction of the

tested materials, as shown in Fig. 4. The press ram

was set to a speed of 20 SPM, which led to an ap-

proximate cutting speed of 50 mm/s at punch/blank

contact. The nominal cutting clearance was set to

0.20 mm and it was verified under non-loading con-

ditions using feeler gauges; this value leads to cut-

ting clearances between 16 and 20% of the sheet

thickness of the tested materials, Table 3. The cut-

ting clearance was selected within this range based

on results by Golovashchenko et al. [14] where the

best relative edge stretchability for trimmed

DP980/1.4 mm is between 10 and 20%; it is hypoth-

esized that these values are also adequate for the

other GEN 3 AHSS with similar tensile strength.

The corner radii of the cutting tools (punch and die)

was manufactured as a sharp edge (i.e. corner radii

near zero); however, it should be mentioned that the

actual radii during the experiments was not meas-

ured. The samples were fully clamped using nitro-

gen cylinders and the trimmed length (𝑙) was 10 mm

in all cases, Fig. 4.

Fig. 4 Schematic, not to scale, of the trimming process.


2.2.2 First Bending Operation

A double flange sample design is proposed, instead

of the single flange design used in [8] and [12], to

help to balance the punch while preventing the part

from sliding out of the tools as the punch moves

downwards. Hence, the initial blank geometry was

extended from 80 by 50 mm to 80 by 96.2 mm. The

first bending operation of the five rectangular sam-

ples was performed along the dotted lines in Fig. 5,

using a finger brake machine with a bending radius

of 1.2 mm (measured from the bent sample); no

fracture along the bent radius was observed when

the samples were inspected under the portable USB

microscope. The flange heights (ℎ) and the bending

angles (𝛽), set to 10 mm and 90 degrees respec-

tively, were verified at four locations to ensure a

proper edge alignment. Small variations of ±0.2 mm

were measured in the flange length while variations

of ±1 degree occurred for the bending angle.

Fig. 5 Schematic of blank geometry before and after first bending operation.

2.3 DOUBLE BENDING TESTS

The DBTs were conducted using a 5500 INSTRON

Tensile Machine with 50 kN capacity and the punch

was set to move at 5 mm/min while the force versus

displacement was recorded; five samples per mate-

rial were tested.

The edge stretchability of the tested materials was

determined by the stroke and thinning at fracture.

The fracture of the material was characterized by a

sudden drop in the punch load as shown in Fig. 6.

The tests were stopped immediately after fracture

was observed on the sample. The thinning was

measured at the closest possible location to the frac-

tured region using a micrometer. In case of exces-

sive buckling, the tests were immediately stopped.

Fig. 6 Sudden drop in the punch load indicates edge fracture on the material.

2.4 EDGE ZONES CHARACTERIZATION

It has been reported, extensively, that clearances be-

tween 10 and 20% of sheet thickness are adequate

when trimming AHSS. Nevertheless, it would be in-

teresting to be able to detect variations in cutting

clearances, within such range, due to elastic deflec-

tions during the trimming process. Thus, additional

trimming experiments were performed for the same

materials listed in Table 1. A nominal cutting clear-

ance of 0.12 mm is set as the lower limit of the stud-

ied range; this value led to clearances of 10-12% of

the sheet thicknesses tested, Table 3.

Table 3: Nominal cutting clearance value per sheet thickness.

In spite of having a greater distance between ana-

lyzed locations, additional samples trimmed along

250 mm, also perpendicular to the rolling direction,

were used for this part of the study, Fig. 7. The 80

mm samples mentioned earlier and the additional

250 mm ones were centered at the same position un-

der the blade for consistency.

The trimmed edges of the samples 250 mm long

were analyzed using a portable USB microscope at

three locations, shown as 𝐿1, 𝐿2 and 𝐿3 in Fig. 7. A

micrometer was placed at the selected locations in

order to have a reference thickness value, which was

used to measure the different edge zones as indi-

cated in Fig. 8.


Fig. 7 Schematic of trimmed samples for edge

characterization using locations 𝐿1, 𝐿2 and

𝐿3 as guidelines.

The rollover, shear and fracture zones were meas-

ured and calculated as percentage of sheet thickness,

which was considered as the summation of rollover,

shear and fracture zone lengths (RZL, SZL and FZL

respectively); the burr length was neglected when

measuring the sheet thickness. Later, an attempt was

done to correlate the RZL, SZL and FZL to the cut-

ting clearances.

Fig. 8 “Front” picture of a trimmed edge, taken with a portable USB microscope. A mi-crometer was placed as reference for the measured zone lengths.

2.5 STRAIN HARDENING AND EDGE

FRACTURE OCCURRENCE

In order to correlate strain hardening near a

trimmed, pierced, or blanked edge to edge fracture

occurrence, as suggested by Diaz-Infante [15],

trimming FE simulations were conducted using

DEFORM 2D for NEXMET 1000/1.26 mm. Cutting

clearances of 10% and 16% of sheet thickness were

simulated considering a punch and die corner radii

of 0.02 mm. In a similar way to the described in

section 2.1, the flow stress curve of the material was

obtained from tensile tests until uniform elongation

and later extrapolated using the Hollomon’s law. At

this point, the material model is assumed to be

independent of strain rate and tempearture; further

research is required to validate this assumption. A

coefficient of friction of 0.1 was used as it has it

showed negligible effect on strain calculations for a

0.06-0.14 range simulated.

In order to find the appropriate Critical Damage

Value (CDV), which defines the onset of fracture

during the FE simulations, the measured shear zone

length in experiments and the corresponding value

calculated in simulations should be matched thru

iterations; a Rice & Tracy fracture criterion was

used to define the CDV, Eq. 1.

∫ 𝑒𝛼𝜎𝑚

�̅� 𝑑𝜀̅𝜀𝑓̅̅ ̅

0= 𝐶𝐷𝑉 (1)

Where 𝜎𝑚= Hydrostatic stress, 𝜎 = Effective Stress

𝜀 ̅= Effective strain and 𝛼 = 2.9

The strains were calculated along a center line at the

shear and fracture zones as illustrated in Fig. 9.

Fig. 9 Strains calculated along Lines 1 and 2 as an indication of damage near the edge and hence as an indication of edge stretchabil-ity.

3 RESULTS

3.1 EDGE STRETCHABILITY

As an initial step and in order to validate the pro-

posed DBT procedure, using the samples trimmed

with a 0.2 mm nominal cutting clearance (values as

% of sheet thickness in Table 3), the thinning at frac-

ture measurements for the DP980 steel were com-

pared with corresponding values in the literature.

The maximum thinning measured at edge fracture

was about 8.1% (average of five measured samples).

This measured edge fracture thinning is higher than

the 6.4% reported in literature [16]. This amount of

variation is attributed to the difference in defor-

mation type at the edge; hence, the procedure is con-

sidered adequate.

As expected, for the tested clearances in Table 3, the

GEN 3 AHSS were able to withstand larger punch

strokes before fracture at the edge than the DP980

steel, as it is shown in Fig. 10. It should be men-

tioned that buckling was observed for 980xG3 steel

after about 6 mm of punch stroke and the experiment

was stopped at that point. Due to the non-occurrence

of fracture, it was not possible to determine the thin-

ning limit for this material; nevertheless, the maxi-

mum thinning measured for 980xG3 was about 6%

on the area where the edge fracture was expected.

This value serves as an indication of a safe thinning

at the edge for the material even when it can not be

directly compared with the remaining tested materi-

als.

The punch stroke before fracture, initially, seems to

indicate that the both GEN 3 AHSS have a better


edge stretchability than DP980. However, as afore-

mentioned, this conclusion can not be firmly stated

for 980xG3. Nevertheless, a better edge stretchabil-

ity of NEXMET 1000 was confirmed by a higher

thinning at fracture than the one for DP980, 9.8%

versus 8.1% respectively.

Fig. 10 Stroke and thinning at fracture for tested AHSS (average of 5 measured samples) and the error bars represent the max. and min. measured values.

Additional samples of NEXMET 1000 were pre-

pared following the procedure described in section

2.2 but using a 0.12 mm nominal cutting clearance

(10% of sheet thickness). Later, DBTs were con-

ducted and compared with the aforementioned re-

sults for NEXMET1000, using a 0.2 mm cutting

clearance (16% of sheet thickness). As it can be ob-

served in Fig. 11, the thinning at fracture is very

similar regardless of the cutting clearance. This

comparison is also approached using FE simulations

as reported in section 3.3.

Fig. 11 Thinning at fracture for NEXMET 1000 /1.26 mm trimmed along 80 mm using two different cutting clearances and tested by Double Bending Tests (DBT). Error bars represent max. and min. measured values.

3.2 CORRELATION BETWEEN EDGE

ZONES AND CUTTING CLEARANCE

After taking pictures of the five trimmed samples

per material, a similar phenomenon to the pointed

out by Nasheralahkami et al. [17] was observed for

the analyzed materials, non-uniform shear SZL

along the trimmed edge. In a few cases, observed

variations in the SZL reached values up to 80%, Fig.

12, but it is possible that even larger variations exist

in non-reviewed edge locations; nevertheless, most

of the analyzed samples showed relatively uniform

edge zones (i.e. about 10-20% variation).

Fig. 12 Shear zone length variation at the same lo-cation on the sample for different samples.

Fig. 13 Rollover Zone Length (RZL), Shear Zone Length (SZL) and Fracture Zone Length (FZL) measurements for different AHSS using two different cutting clearances.


Considering that all samples were located approxi-

mately at the same location under the blade, tool

wear is discarded since, for a particular location (e.g.

𝐿2), the variations of shear zone length are different

depending on the sample analyzed. In order to es-

tablish a procedure for edge zones measurements, a

horizontal line was traced so that it can represent

most of the observed shear zone. The values re-

ported in Fig. 13 were measured using this line as a

reference. The large scatter shown in Fig. 13, sug-

gests that the proposed edge characterization

method is not suitable to discriminate cutting clear-

ances for the tested materials; since the potential dif-

ference in zone lengths due to the cutting clearances

is covered by the non-uniformity of the shear zone

length.

3.3 CORRELATION BETWEEN STRAIN

HARDENING AND EDGE FRACTURE

The CDV of NEXMET 1000/1.26 mm was found to

be 1.35, consistently for the cutting clerances

simulated (10% and 16% ), using the Rice & Tracy

fracture criterion. This value was found by matching

the average shear zone length of the material,

reported in Fig. 13.

For the simulated conditions described in section

2.5, at the shear and fracture zones of the trimmed

edge, Fig. 9, the calculated strain hardening on the

material turned out to be very similar as illustrated

in Fig. 14. The calculated strains are used as an

indication of material hardness which has been

related to the edge fracture [18]; i.e. the higher the

hardness near the cut edge, the lower the edge

stretchability.

Fig. 14 Strains near the trimmed edge at the shear and fracture zones calculated by FE simu-lation using DEFORM 2D.

This observation is aligned with the experimental

results which indicated a similar edge stretchability

for NEXMET 1000/1.26 mm. Hence, the strain

hardening calculations using FE simulations is

considered a good indication of the expected edge

stretchability for given cutting conditions.

Nevertheless, it should be noticed that this method

only evaluates the deformation due to the cutting

process and not the possible fractures due to the

microstructure of the material.

Fig. 14 also helps to understand the approximate

length of the edge damage and hence to estimate the

amount of shaving that may be used to increase the

edge stretchability as shown by Feistle et al. [19].

4 CONCLUSIONS

Based on the results presented for the Double Bend-

ing Test (DBT) and the edge characterization, the

following conclusions are drawn:

Punch stroke may not be a good method to eval-

uate edge stretchability when comparing differ-

ent materials, since, in some cases, at similar

strokes different materials may have different

thinning/strains at the edge. However, this

method could be useful when comparing differ-

ent trimmed edges for the same material, i.e. the

larger the punch stroke, the better the stretcha-

bility.

NEXMET 1000/1.26 mm showed a higher thin-

ning at fracture than DP980/1.21 mm for a 17%

cutting clearance (% of sheet thickness); punch

strokes at fracture were about 6 and 5 mm re-

spectively. No fracture was observed for

980xG3/0.98 mm as the tests were stopped due

to buckling (at about 6 mm punch stroke).

The fracture occurrence for NEXMET 1000

was very similar for clearances 10 and 17%;

similar to information in literature for DP980

[14].

Potential buckling is a significant disadvantage

and further tool and sample geometry optimiza-

tion is required to minimize this issue without

the use of additional devices. This problem is

more noticeable as the material thickness de-

creases (e.g. no buckling observed for 1.2 mm

sheets but observed for 1.0 mm sheets, for

tested configuration).

Traditional methods such as micrographs, lim-

ited to a very localized area of trimmed edge,

may not be enough to characterize the edge ge-

ometry of AHSS due to significant variations

along of the shear zone length, for example.

Even when the tools are sharp and the clear-

ances uniform, these variations may occur due

to the non-uniform microstructures of the

AHSS; however, the analysis of microstruc-

tures is not in the scope of this research.

The potential difference in zone lengths due to

the cutting clearances is covered by the non-

uniformity of the shear zone length. Hence, it

was not possible to distinguish between cutting

clearances within the studied range for the

tested materials, 10% to 20% of sheet thickness.


A greater number of pictures using the USB mi-

croscope and a statistical model may help with

this differentiation.

5 ACKNOWLEDGEMENT

The authors would like to thank Dr. Ali Falla-

hiarezoodar, from Shiloh Industries, Inc., who pro-

vided the tools and press where the trimming exper-

iments were conducted; his help and advice are

greatly appreciated. Also, thanks to U.S. Steel and

AK steel for providing enough materials for the ex-

periments.

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