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ISBN print: 978-3-0364-0992-4
ISBN eBook: 978-3-0364-1992-3
DOI book: 10.4028/b-7qyxOQ
© 2026 Trans Tech Publications Ltd, All Rights Reserved

Friction and Wear in Metal Forming

Edited by
Dr. Javad Hazrati

Friction and Wear in Metal Forming

Special topic volume with invited peer-reviewed papers only

Edited by
Dr. Javad Hazrati

Scientific.Net

Copyright © 2026 Trans Tech Publications Ltd, Switzerland

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Volume 1184 of
Materials Science Forum
ISSN print 0255-5476
ISSN web 1662-9752

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Materials Science Forum, ISSN: 1662-9752, Vol. 1184, pp -1-0

© 2026 The Author(s). Published by Trans Tech Publications Ltd, Switzerland.

Accepted:
2026-04-08

Online:
2026-04-14

Main Editor

Preface

The original version of this paper is available on https://www.scientific.net/MSF.1184.-1.pdf

In metal forming, relative motion of the tool and workpiece during production makes friction, lubrication and wear phenomena (tribology) of the utmost importance. On that account, for optimum process design, improving formability and enhancing tool life, a better understanding of tribology and an accurate prediction of friction/wear in forming processes are necessary. In this issue, different aspects of tribology in cold/hot sheet metal forming, forging and rolling will be discussed. The topics that are covered in this issue in particular are: modelling of friction, lubrication and wear; tribological characterization methods; surface engineering and control of friction/wear in cold/hot forming processes.

Hazrati, Javad

The original version of this paper is available on https://www.scientific.net/MSF.1184.-1.pdf

Co-editors

The original version of this paper is available on https://www.scientific.net/MSF.1184.-1.pdf

Galdos, LanderNielsen, Chris Valentin

University of Twente

University of Mondragon
Technical University of Denmark

Table of Contents

Preface

Gallig-Free Dry Net Shaping of Titanium and Titanium Alloy Eye-Glass Frame Parts toward Green Manufacturing

Tatsuhiko Aizawa, Takeshi Kihara, Tomomi Shiratori

Selection and Investigation of Thermoplastic Materials as Forming Tools for Warm Dry Forming of Magnesium and High-Strength Aluminum Sheets

Serhii Vakulenko, Kai Weigel, Martin Dix

Multiscale Simulation of Asperity Flattening with Realistic Surface Topography and Microstructure

Aiden Carley-Clopton, Grethe Winther, Javad Hazrati, Chris Valentin Nielsen

Tool Wear Investigation Nickel Plated Steels for Battery Shell Production

Jenny Venema, Vincent Pater, Jan Wörmann, Marco Appelman

Numerical Analysis of Roughness Transfer Mechanism during Skin-Pass Rolling

Hamed Aghajani Derazkola, Leon Jacobs, Stamatis Kiakidis, Ton van den Boogaard, Javad Hazrati

Modeling the Effect of Lubricants on Surface Conditions in Plane-Strain Upsetting Tests

Lena Koch, Holger Brüggemann, Shakthi Bharani Tamilselvan, Emad Scharifi, Junhe Lian

Improvement of Wear and Corrosion Resistance of X46Cr13 Martensitic Stainless Steel by Cryogenic Treatment

Isabel Espinosa, Sergi Menargues, Juan David Gutierrez, Josep A. Picas, Javier Antonio Navas Lopez, Maria Teresa Baile Puig

Assessment of the Potential of CO2 as a Lubricant in Cold Forging of Low-Carbon Mild Steel

Diego Rafael Alba, Alexander Weiss, Gerd Reichardt, Mathias Liewald

A New Ex-Situ Method for Real Contact Area Determination for Sheet Metal Forming

Ainhoa Guinea, Andrea Aginagalde, Eneko Saenz de Argandoña, Joseba Mendiguren, Wilson Tato, Iñigo Llavori, Liam Blunt, Alaitz Zabala

Sliding Distance Dependency and Third Body Particle Influence in Flat Strip-Draw Testing of Aluminum Sheet for Friction Characterization in Automotive Stamping

Loris Rocchi, Jan Filzek, Christian Leppin

Towards Multi-Scale Friction Modelling for Bulk Sheet Metal Forming Applications

Aratz Barandiaran, Alaitz Zabala, David Abedul, Javad Hazrati, Lander Galdos

Effect of Surface Texture on Sheet Metal Formability in Boundary Lubrication Regime

Javad Hazrati, Matthijn de Rooij

Keywords Index

Authors Index

The original version of this eBook is available on https://www.scientific.net/book/friction-and-wear-in-metal-forming/978-3-0364-1992-3

Materials Science Forum, ISSN: 1662-9752, Vol. 1184, pp 1-9

doi: 10.4028/p-MoKQR2

© 2026 The Author(s). Published by Trans Tech Publications Ltd, Switzerland.

Submitted:
2025-12-11

Revised:
2026-02-16

Accepted:
2026-02-23

Online:
2026-04-14

Gallig-Free Dry Net Shaping of Titanium and Titanium Alloy Eye-Glass Frame Parts toward Green Manufacturing

The original version of this paper is available on https://www.scientific.net/MSF.1184.1.pdf

Tatsuhiko Aizawa 1,a*, Takeshi Kihara 2, b and Tomomi Shiratori 3,c

11-29-1 Nishi-Rokugo, Ota-City, Tokyo 144-0056, Japan

24−15 Imaichi, Fukui-City, Fukui 918-8152, Japan

33190 Gofuku, Toyama-City, Toyama 930-8555, Japan

a* taizawa@sic.shibaura-it.ac.jp, b t.kihara@masunaga-opt.co.jp, c shira@eng.u-toyama.ac.jp

Keywords: galling-free shaping, titanium, eye-glass frame parts, saving of wastes.

Abstract

Massively carbon supersaturated (MCSed) tool steel dies were prepared for dry, galling free forging, microtexturing and fine-blanking of titanium and titanium alloy eye-glass frame parts. Titanium temples were forged in dry and galling-free to investigate the role of MCS treatment to reduce surface roughing of temples. They were also micro-textured to discuss the superiority of MCS to product quality. Dry, galling-free fine-blanking was utilized to describe the life-time extension of MCSed tool steel dies. Various eye-glass frame parts were forged in dry to state the statistic evaluation on the galling-free manufacturing features. The intermediate chemical treatments as well as barreling and polishing steps were saved to reduce the amount of wastes from manufacturing factories of titanium and titanium alloy eye-glass frames toward zero emission.

[picture]

Fig. 1. Comparison of the dry manufacturing procedures for pin-formation with and without use of MCS treatment. a) Without MCS-treatment, and b) with MCS-treatment.

Fig. 1. Comparison of the dry manufacturing procedures for pin-formation with and without use of MCS treatment. a) Without MCS-treatment, and b) with MCS-treatment.

Introduction

The original version of this paper is available on https://www.scientific.net/MSF.1184.1.pdf

The galling or adhesive wear has been identified as one of the most essential issues in metal forming [1-2]. Especially, severe galling was studied in dry sliding friction tests on various coated test-pieces against titanium balls [3]. In dry upsetting of titanium bars, the debris particles splashed in the air together with much adhesion of titanium fragments [4]. No improvements took place in deep drawing of titanium sheets even using the ceramic dies [5]. Hence, the manufacturing procedure in dry required multi-steps to fabricate the eye-glass frame parts from titanium bar feedstock [6]. A typical multi-step procedure was illustrated in Fig. 1a. In additional to multi-step forging processes, the intermediate treatments of anodizing, chemical treatments, barreling and polishing steps were needed to continue the forging procedure without severe galling. In particular, nitric acid (HNO3) and hydrogen fluoride (HF) solutions were utilized to remove the used titanium oxide films on the forged titanium work material surfaces in every intermediate step.

The massive carbon supersaturation (MCS) treatment to tool steel dies, was proposed to make dry, galling-free forging of titanium and titanium alloy feedstock [7]. MCSed SKD11 and ADC56 dies were utilized not only for upsetting the titanium bars in high reduction of thickness [8, 9] but also for fine-blanking the titanium plates to gears with high accuracy [8, 10]. The original forging steps in the pin formation in Fig. 1a were nearly halved, and, most of intermediate steps were saved by using the MCSed ACD56, as illustrated in Fig. 1b. In addition of this reduction in the number of steps, the energy consumption in forging was saved by 20−30% because of significant reduction in the applied load. This suggests that MCS treatment has a role in reduction of environmental burden in the life cycle assessment (LCA) [11].

In the present paper, tool steel dies are MCSed using the plasma carburizing method. The singleshot forging and fine-blanking processes with the use of MCSed ACD56 dies, are performed to describe the galling-free shaping of pure titanium feedstock by in situ formed carbon tribofilm on the contact surface and to demonstrate the role of MCS treatment in high qualification of forged and blanked products. Various titanium eye-glass frame parts are forged to statistically describe the superiority to MCS treatment of dies toward green manufacturing. In particular, no surface roughing by MCSed die directly results in high qualification of products; fine microtexturing by MCSed die is preferable to design-oriented decoration of eye-glass frames.

[picture]

Fig. 2. Plasma immersion carburizing system for MCS treatment of tool steel dies.

Fig. 2. Plasma immersion carburizing system for MCS treatment of tool steel dies.

Methods and Materials

The original version of this paper is available on https://www.scientific.net/MSF.1184.1.pdf

The plasma processing conditions for MCS treatment were explained to prepare for MCSed dies. A single-shot forging process with the use of MCSed ACD56 dies was stated for net-shaping of various titanium eye-glass frame parts. Fine-blanking and microtexturing processes were also described to find the superiority to MCS treatment for galling-free manufacturing.

MCS-treatment

The original version of this paper is available on https://www.scientific.net/MSF.1184.1.pdf

The plasma carburizing immersion system was utilized for MCS treatment of tool steel dies. As illustrated in Fig. 2, the die was plasma carburized in the hollow cathode to intensify the carbon ion and CH -radical densities. The holding temperature was 573 K to be free from thermal distortion. After heating and presputtering for 1.8 ks via the DC (Direct Current) plasma under the bias voltage of -500 V , the dies were carburized for 16 ks under the RF (Radio-Frequency) voltage by 250 V and the bias by -500 V respectively before cooling in the nitrogen atmosphere down to RT (Room

Temperature). After cleansing and polishing, this die was directly utilized in the following forging, fine-blanking and microtexturing processes. A knuckle-joint stamper (AIDA160; AIDA, Co., Ltd.; Kanagawa, Japan) was utilized in these processes with the maximum load of 1200 kN . The applied load was monitored in each forging strokes. Under the constant stroke, velocity and duration, the applied energy was proportional to load.

Forging process

The original version of this paper is available on https://www.scientific.net/MSF.1184.1.pdf

Titanium feedstock was upset and net-shaped by MCSed ACD56 dies. In-situ carbon tribofilm formation is proved after continuously upsetting in 30 shots. Temples are dry-forged with and without MCS to describe the role of MCS to suppress the surface roughness in forging. SEM (Scanning Electron Microscopy) - EDS (Electron Dispersive x-ray Spectroscopy) is utilized for microstructural analysis on the die surface.

Fine blanking

The original version of this paper is available on https://www.scientific.net/MSF.1184.1.pdf

Shearing the titanium plates and forged parts, becomes an essential step in fabrication of eye-glass frame parts. Titanium plates with the thickness of 5.0 mm were continuously punched out to describe the effect of MCS on the burnished surface area ratio. This fine blanking is utilized to demonstrate the die-life extension by MCS treatment.

Microtexturing

The original version of this paper is available on https://www.scientific.net/MSF.1184.1.pdf

Various designs of eye-glass titanium frames are allowable in decorating microtextures on the titanium temples. An acetate temple was a typical new material design instead of a traditional plastic one. This temple material is made from natural cellulose so that no environmental burdens are generated even in disposal. In most of designs, this acetate temple is strengthened by a titanium plate as shown in Fig. 3. This plate is often microtextured for decoration of light-colored or transparent acetate temples.

[picture]

Fig. 3. Microtexturing to titanium temple surfaces for decoration of eye-glass frames.

Fig. 3. Microtexturing to titanium temple surfaces for decoration of eye-glass frames.

Results and Discussion

The original version of this paper is available on https://www.scientific.net/MSF.1184.1.pdf

A single shot upsetting process was employed to describe the galling-free performance when using the MCSed ACD56 dies. A fine-blanking process was performed to demonstrate the tool life extension by MCS treatment. A microtexturing process was done to state the role of MCS in high qualification of forged eye-glass frame. Various eye-glass frame parts were forged for statistical survey on the superiority to productive technology toward zero-emission. In common, the pure titanium in the industrial grade II was utilized as a feedstock in every forging and forming step.

Single-shot forging process

The original version of this paper is available on https://www.scientific.net/MSF.1184.1.pdf

Pure titanium bar with the diameter of 5.0 mm was continuously upset using the MCSed ACD56 dies in the reduction of thickness (r) by 50% up to 30 shots. Figure 4 compares the first and 30th  upset works. When N=1, the work surface is covered by the white debris particles while no debris particles are seen when N=30. Without MCS treatment, the debris particles gradually deposited onto the die
and work surfaces; the whole surfaces were covered by debris particles when N=20 [4]. This gallingfree upsetting behavior in Fig. 4 is completely opposite to the conventional upsetting process without MCS. In addition, the shear-localization, which was observed in the normal upsetting of titanium bar at r=50%, never took place when using the MCSed die. That is, the titanium bars are upset with homogeneous work hardening without plastic localization.

[picture]

Fig. 4. Comparison of MCSed ACD56 die surfaces at N=1 and N=30.

Fig. 4. Comparison of MCSed ACD56 die surfaces at N = 1 and N = 30 .

This unique forging behavior to MCS treatment comes from the in situ formation of amorphous carbon tribofilm during upsetting. SEM-EDS (JEOL; Tokyo, Japan) was used to make microstructure analysis on the MCSed die surface after upsetting up to N=30.

[picture]

Fig. 5. SEM-EDS analysis on the MCSed ACD56 die surface at N=30. a) SEM image on the MCSed die surface at the center, and, b) carbon mapping on the same area.

Fig. 5. SEM-EDS analysis on the MCSed ACD56 die surface at N = 30 . a) SEM image on the MCSed die surface at the center, and, b) carbon mapping on the same area.

As shown in Fig. 5a, the contact area of titanium work is traced in dark at the center of die surface. This area is just corresponding to a carbon map in Fig. 5b. Especially, the back dots in Fig. 5a consist of carbon dots in Fig. 5b. Thie proves that carbon solutes are isolated from MCSed dies to in situ form a carbon film only on the contact surface of MCSed die to titanium work. Let us state the effect of MCS treatment on the product surface quality.

[picture]

Fig. 6. Comparison of the forged eye-glass temples with and without MCS treatment. a) Without MCS treatment, and b) with MCS treatment.

Fig. 6. Comparison of the forged eye-glass temples with and without MCS treatment. a) Without MCS treatment, and b) with MCS treatment.

Residual adhesives from titanium work onto the die surface, often deteriorate the surface roughness in dry forging of temples. Figure 6 compares the forged temple surfaces with and without MCS treatments. As seen in Fig. 6a, the die surface roughness by galling is imprinted onto the product surface as scratches and flaws. Additional barreling, polishing and cleansing processes are keenly needed to remove these defects; the cost-performance is deteriorated in manufacturing. When using the MCSed dies, no galling took place so that no flaws were left on the MCSed die surface. As depicted in Fig. 6b, the temple surface is never roughed so that no additional steps are necessary in this dry forging.

Fine-blanking process

The original version of this paper is available on https://www.scientific.net/MSF.1184.1.pdf

A fine blanking step is always necessary to pick up the forged product from its skeleton and to decorate the temples by pierced holes or bands. Two issues must be solved in this die technology; suppression of severe galling to attain nearly full-burnished surfaces, and, die-life extension even in dry fine blanking.

[picture]

Fig. 7. Comparison of ACD56 punches for fine-blanking of titanium plates with and without MCS treatment, a) ACD56 punch without MCS treatment at N=1000, and b ) ACD56 punch with MCS treatment at N=10,000.

Fig. 7. Comparison of ACD56 punches for fine-blanking of titanium plates with and without MCS treatment, a) ACD56 punch without MCS treatment at N = 1000 , and b ) ACD56 punch with MCS treatment at N = 10 , 000 .

Figure 7 compares the ACD56 punch surfaces after the selected number of shots with and without MCS treatment. Without MCS, severe galling took place to form the debris films on the whole head and side surfaces of tool steel punches, so that further fine blanking is terminated at N=1000 by this blockage to clearance. On the other hand, no galling was observed even at N=10,000, which is ten times more than the situation observed in Fig. 7a. This direct comparison in Fig. 7 proves that MCS treatment provides a way to make galling-free fine blanking continuously without deterioration of die surface.

[picture]

Fig. 8. Comparison of the fine-blanked titanium skeleton by ADC56 tool steel die with and without MCS. a) A skeleton fine-blanked at N=1000 without MCS , and, b) a skeleton fine-blanked at N=10,000 with MCS.

Fig. 8. Comparison of the fine-blanked titanium skeleton by ADC56 tool steel die with and without MCS. a) A skeleton fine-blanked at N = 1000 without MCS , and, b) a skeleton fine-blanked at N = 10 , 000 with MCS.

Suppression of galling behavior by MCS reflects on the sheared surface quality after fine blanking. Figure 8 compares the fine-blanked skeletons with and without MCS. In fine blanking without MCS, the secondary sheared area was observed in Fig. 8a at N=1000. This proves that fractured surface area is inevitably formed without MCS. When using the MCSed ACD56 die, the sheared surface is preserved to be nearly full-burnished even until N=10,000. This comparison assures that MCS also provides a solution to improve the fine-blanked product quality and to extend the tool life at the same time even in dry fine blanking.

Microtexturing process

The original version of this paper is available on https://www.scientific.net/MSF.1184.1.pdf

Fine microtextures are often required for design decoration of the eye-glass frames, as introduced in Fig. 3. In order to promote the customer interests in products, fine microtextures must be deeply cut into the titanium temple inserts by dry forging with higher aspect ratio of the microtexture depth to its width.

[picture]

Fig. 9. Comparison of the microtextured titanium temple surfaces with and without MCS treatment.

Fig. 9. Comparison of the microtextured titanium temple surfaces with and without MCS treatment.

Figure 9 compares the dry-forged titanium inserts with and without use of MCSed die. Without MCS, the aspect ratio of microtextures is small to be colored in light blue. Using the MCSed die in forging, this aspect ratio is enhanced enough to be colored in dark blue. This difference in the quality of microtextures into the titanium inserts influences on the total quality of eye-glass frame.

Statistical study in dry forging of various eye-glass frame parts

The original version of this paper is available on https://www.scientific.net/MSF.1184.1.pdf

This MCS has been applied to forging and forming of various eye-glass frame parts in two factories. Seven titanium parts and three rectangular frames were forged in dry to make statistical study on the yield of products, as listed in Table 1. The parts from A to C and rectangular frame-A were fabricated at the main factory in Japan. While other parts and rectangular frames were yielded at the Malaysian branch. Using MCSed dies, each part was yielded by a single cold dry forging operation till the
number of products listed in Table 1. In case of Part-A, 500 armors were yielded by a single forging without any stops due to cleansing and polishing operations.

Table 1. A statistical study on the dry, cold and galling-free forging operations to yield various titanium eye-glass frames and parts in two factories.

Eye-Glass Frame Parts | Part Name | Number of Products
Part-A | Armor | 500
Part-B | Bridge | 300
Part-C | Armor | 70
Rectangular frame-A | Temple + bar | 2000
Part-D | Tempe + temple tip | 11,000
Part-D | Temple + joint | 18,000
Part-E | Temple + temple tip | 5,000
Part-F | Temple + temple tip | 7,000
Part-G | Temple + joint | 12,000
Rectangular frame-B | Temple | 5,000
Rectangular frame-C | Temple | 5,000

Without MCS, every eye-glass frame part was fabricated in the multi-step procedure including the intermediate steps of the barreling, polishing and chemical treatments. For an example, in case of rectangular frame-A, three to five steps were needed to minimize the adhesion of debris fragments onto the die surface.

The total number of products for each part item in Table 1, was yielded in less forging steps without any intermediate steps including the barreling, polishing and chemical treatment when using the MCS treatment. Let us compare the number of steps to fabricate the Parts-D/-E/-F and -G as well as the rectangular frames-B and -C in Table 1. Without MCS, 27 steps were needed in total, including the forging steps, the barreling, polishing and chemical treatments. When using MCS, this number of steps was reduced to 11 forging steps only. The whole intermediate treatments were completely saved to reduce the related energy and resource consumption.

As discussed before, the number of forging steps in production procedure for each item is also reduced nearly by 50%. In case of Part-D/-E/-F ad -G, the friction coefficient is reduced during each forging operation by MCS so that the applied load is also reduced by 30% in forging temples. That is, the total forging time was nearly halved or less, and, the energy consumption for single forging operation is reduced by 30% when using MCSed dies.

In the following, let us quantitatively evaluate on the saving impact of intermediate steps by MCS. As depicted in Fig. 1a, a series of intermediate operations consists of five tasks: heat treatment for annealing, mixed acid cleansing, barreling, anodizing and Estor coating. Among five tasks, the impact to production time in the annealing, the anodizing and chemical cleansing tasks saved by MCS, is evaluated from the statistical data in Table 1.

In the heat treatment, each forged eye-glass frame was subjected to annealing lot-by-lot in operation. In average, 110 minutes were needed for each treatment of a single lot. The total time for treatment of 60 lots were wasted for heat treatment of parts and frames in Table 1 per a month. Then, the wasted production time reached 4,752ks=1320 hours per a year.

The anodizing process was indispensable to coat the protective oxide layer onto the titanium work after every forging operation without MCS. In average, 54 minutes were needed for each anodizing step of a single lot. The total time for treatment of 60 lots were wasted for this anodizing of parts and frames in Table 1. Then, the wasted production time reached 2, 333ks=648 hours per a year.

In the chemical treatment, a mixed acid solution, which consists of HF and HNO3, was utilized to delete the titanium adhesives as well as titanium oxide debris. A working time by 2.5 days was needed to exchange the whole used solution with a new one. In each month, 9 exchanges were required to preserve this chemical treatment capacity per a month. Then, the wasted production time by anodizing step, reached 7,778ks=2,160 hours per a year.

Main three intermediate steps wasted in total 4, 128 hours, or, 172 days per a year. This significant loss is saved by using MCS to eliminate these intermediate steps from the production line of eyeglass frames and parts.

In the above three steps, no direct emission of wastes was generated in two factories with respect to the heat treatment and anodizing steps. However, the used HF and HNO3 mixed acid solutions were directly emitted from factory sites to outside. Let us qualitatively estimate this waste action and evaluate on its impact as an environmental burden toward fully standardized LCA.

[picture]

Fig. 10. Post-treatment routine of the used mixed-acid ( HNO3+HF ) solutions finally to sludges and supernatant liquid wastes.

Fig. 10. Post-treatment routine of the used mixed-acid ( H N O 3 + H F ) solutions finally to sludges and supernatant liquid wastes.

The mixed acid waste solution was first diluted by ionized water by the ratio of 1 L for waste to 10 L for water in volume after regulated routine in [12]. This diluted solution was further mixed with 10%Ca(OH)2 slurry by the ratio of 1 L for diluted solute to 10 ML for slurry per a minute to increase the pH of waste solution above 11. The duration is 15 mins in this treatment. After precipitation treatment, the supernatant liquid was emitted as an industrial water under BOD (Bio-chemical Oxygen Demand) and COD (Chemical Oxygen Demand) control. The sludge must be disposed as a landfill waste.

As stated before, the mixed acid solution by 8 liters per a task was monthly exchanged with new one in 9 times. The solution contents were 2 litters for HF and 6 litters for HNO3. Then, the monthly total waste reached 72 litters. This direct emission of used chemical solutions is evaluated in Fig. 10. In the dilution stage, the ionized water of 720 L was used before mixing and stirring with the slurry of 720ML/min. Since the duration time of this treatment needs 15 mins , the total amount of slurry reaches to 10,800ML. After precipitation treatment, this waste was divided to the solid sludge including the titanium - calcium oxide powders and the supernatant liquid with the volume of 10,800ML.

The industrial water and chemical consumption calculation reveals that total ejection from dry cold forging operations is much saved by MCS treatment to reduce the environmental burden. Toward fully standardized LCA, the energy and feedstock consumption by single forging operation and the reduced number of steps in total manufacturing must be considered in addition to the water and chemical consumption estimates in the above.

Conclusion

The original version of this paper is available on https://www.scientific.net/MSF.1184.1.pdf

A sustainable manufacturing requires for long life of tools, less emission of wastes, reduction of energy and time consumption, and, flexible production lines. Dry metal forming of titanium eyeglass frames and parts is employed to demonstrate that massive carbon supersaturation (MCS) treatment of dies is just adaptive to green manufacturing with satisfaction of four items in the above. The adhesive wear of titanium fragments and debris to die surfaces is reduced to prolong the die life with less maintenance costs. Intermediate steps including the barreling, cleansing, anodizing, heat treatment and polishing, are saved in manufacturing. The water and chemical consumption calculation proves that environmental burden is also reduced by MCS. Less number of forging steps
and reduction of applied energy in forging demonstrate that amount of labors as well as waste of time are saved together with reduction of energy wastes.

This MCS has a possibility to change the production line in practice. At present, every eye-glass frame and related part is fabricated by a single-shot forging. Manual handling is needed to feed the work and to eject, pick-up and store the forged parts. Using the MCSed die system, this single-shot operation changed to transfer- or progressive-stamping system, where a lot of parts are automatically forged, ejected and stored with much less of labors and costs.

Acknowledgement

The original version of this paper is available on https://www.scientific.net/MSF.1184.1.pdf

The authors would like to express their gratitude to Masunaga S., Yamamoto T., and, Tani S. (Masunaga-Opt, Co., Ltd.) for their help on the statistical survey.

References

The original version of this paper is available on https://www.scientific.net/MSF.1184.1.pdf

Hutchings I., Shipway P., Tribology: Friction and wear of engineering materials. second ed., Elsevier, 2017.

Dohda K., Yamamoto M., Hu C., Dubar L., Ehmann K. F., Galling phenomena in metal forming. Friction, 2021, 9, pp. 665-685.

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Aizawa T., Kihara T., Shiratori T., Falling-free dry near-net forging of titanium using massively carbon-supersaturated tool steel dies. J. Materials 2024; 17 (19) 4849, pp. 1-14.

Aizawa T., Fuchiwaki K., Galling-free fine blanking of titanium plates Using carbon supersaturated high-speed steel punch. J. Carbon Research, 2023, 8, 15, pp. 1-12.

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Sasaki Chemicals, Recipe for chemical treatment of mixed acids used in the fabrication of titanium parts. . (Retrieved at 2025/12/5).

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The original version of this eBook is available on https://www.scientific.net/book/friction-and-wear-in-metal-forming/978-3-0364-1992-3

Materials Science Forum, ISSN: 1662-9752, Vol. 1184, pp 11-20

doi: 10.4028/p-Te5Ycb

© 2026 The Author(s). Published by Trans Tech Publications Ltd, Switzerland.

Submitted:
2025-12-16

Revised:
2026-02-16

Accepted:
2026-02-23

Online:
2026-04-14

Selection and Investigation of Thermoplastic Materials as Forming Tools for Warm Dry Forming of Magnesium and High-Strength Aluminum Sheets

The original version of this paper is available on https://www.scientific.net/MSF.1184.11.pdf

Vakulenko Serhii 1,a*, Weigel Kai 2, b and Dix Martin 1,3,c

1 Fraunhofer Institute for Machine Tools and Forming Technology IWU, Reichenhainer Strasse 88, 09126 Chemnitz, Germany

2 Fraunhofer Institute for Thin Films and Surface Engineering IST, Riedenkamp 2, 38108

Braunschweig, Germany

3 Institute for Machine Tools and Production Processes, Chemnitz University of Technology, Reichenhainer Strasse 70, 09126 Chemnitz, Germany

a* serhii.vakulenko@iwu.fraunhofer.de, b kai.weigel@ist.fraunhofer.de, c  martin.dix@iwu.fraunhofer.de

Keywords: dry sheet metal forming, thermoplastic forming tool, high-strength aluminum alloy.

Abstract

The growing demand for lightweight components has highlighted the potential of sheet metal parts made from high-strength aluminum and magnesium alloys. However, the forming of these materials is often limited by complex manufacturing processes that typically require lubricants, raising environmental and economic concerns. This study investigates the feasibility of using thermoplastic polymers as tooling materials for dry forming of aluminum alloy 7075-T6 and magnesium alloy AZ31B at temperatures up to 300∘C. A number of high-performance thermoplastic polymers were selected based on their superior mechanical and thermal properties and evaluated through compression and tensile tests, dilatometry, pin-on-disk and draw-bend tests. In addition, the influence of coatings on the tribological performance of the polymers was systematically investigated. The results showed that thermoplastic tools have the potential for efficient dry forming of AA 7075T6 and AZ31B parts at temperatures up to 300∘C, achieving good tool stability, wear resistance and part quality. TECAPEEK PVX with original tribological optimization proved to be the most promising candidate and was validated in challenging cross cup forming experiments. This lubricantfree approach offers significant environmental and economic benefits, providing a sustainable solution for prototyping and small-scale production of lightweight metal components.

Introduction

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The global push for more fuel-efficient transportation systems, coupled with increasingly stringent environmental regulations, has driven demand for high-strength yet lightweight structural components. Aluminum and magnesium alloys in particular offer an excellent balance of low density and high specific strength, making them prime candidates for aerospace, automotive, and consumer electronics applications [1,2]. However, their room temperature formability is still limited by low ductility and an increased tendency to springback. In addition, because these alloys tend to adhere strongly to conventional tool steels-especially at elevated temperatures-lubricants are often used during forming [3]. While lubricants reduce friction and wear, they also pose significant environmental and economic challenges [4]. Increased sustainability goals, driven by both regulations and consumer expectations, are driving research into lubricant-free or "dry" forming processes [5]. Realizing such processes requires innovative tool materials, coatings, and process configurations to effectively manage friction and wear under challenging thermo-mechanical conditions.

Plastics as alternative tool materials have attracted considerable interest since the early 2000s [6, 7]. Various plastics have been successfully used in deep drawing [8], bending [9], stamping [10], hydroforming [11], and incremental sheet metal forming, demonstrating advantages for prototyping and small-scale production. These advantages include low cost, rapid tool manufacturing (particularly through additive manufacturing [12]), ease of machining, and ease of maintenance-all while
achieving acceptable dimensional accuracy in formed parts [9]. However, most studies using polymer-based tools address room-temperature forming; research focusing on plastic tool materials for warm or tempered forming of magnesium and high-strength aluminum alloys remains limited.

High-performance thermoplastic polymers hold promise for tempered forming of these alloys. Their low density, chemical resistance, and ability to maintain sufficient mechanical integrity up to 260∘C for long-term exposure - and up to 300∘C for short-term exposure - make them suitable for small scale production [13]. In addition, their properties can be tailored to meet specific process requirements by incorporating reinforcing fillers (e.g., carbon (CF) or glass (GF) fibers) or by using coatings and additives (e.g., polytetrafluoroethylene).

The present study investigates the feasibility and performance of thermoplastic materials as forming tools for the tempered dry forming of magnesium and high-strength aluminum sheet. By evaluating different thermoplastics in terms of their suitability for thermo-mechanical and tribological properties, we aim to gain a comprehensive understanding of the possibilities and limitations of this innovative approach to sheet metal forming.

Material Characterization and Experimental Methods

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Sheet materials and their characterization.

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Two sheet materials were selected for this study: aluminum alloy AA 7075-T6 ( 2 mm thick) and magnesium alloy AZ31B ( 1 mm thick), which were provided by industrial partners and reflect standard industry stocks and manufacturing practices. To investigate their mechanical properties such as yield ( Rp0,2 ) and ultimate ( Rm ) strength and elongation (A), uniaxial tensile tests were performed on Zwick/Roell universal testing machines at four temperatures (room temperature ( 20∘C ), 150∘C, 200∘C,250∘C ) using a constant strain rate of 0.005 s−1. Test specimens, prepared according to standard guidelines, were extracted from the rolled sheets in two orientations ( 0∘,90∘ ) to capture the anisotropic behavior. Formability limits under varying conditions were determined by temperaturecontrolled Forming Limit Curve (FLC) tests performed according to the Nakajima test method (DIN EN ISO 12004-2 [14]). These tests used specimens of four different widths (w=50 mm,90 mm,130mm,200 mm) and were conducted at the same temperatures as the tensile tests at controlled punch speed of 3.5 mm/s. Strain measurements and analysis were performed using GOM's ARAMIS 3D optical measurement system. All tests were repeated three times and the results of the tests at 250∘C are summarized in Table 1.

Table 1. Properties of the sheet material at 250∘C.

|  | AA 7075-T6 | AZ31B
|  | 0° | 90° | 0° | 90°
Tensile test | Rp0,2 [MPa] | 177 | 181 | 25 | 30
Rm [MPa] | 182 | 181 | 26 | 31
A [%] | 25 | 11.2 | 53 | 58
|  | φ1 | φ2 | φ1 | φ2
FLC test | w = 50 mm | -0.16 | 0.38 | -0.19 | 0.69
w = 90 mm | 0.01 | 0.26 | -0.05 | 0.50
w = 130 mm | 0.10 | 0.28 | 0.03 | 0.39
w = 200 mm | 0.19 | 0.31 | 0.13 | 0.42

Plastic materials and their characterization.

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A comprehensive evaluation was conducted to assess the feasibility of using plastic as tool materials for the target process. Based on manufacturer data and suggestions of the involved polymer manufacturers nine candidate polymers were selected. These include four low-cost polyurethanebased (PUR1-4) plastics - which, although not thermoplastics, perform adequately up to 100∘C - and thermoplastics based on polyamide 66 (TECAMID 66/X GF50 black), polyetheretherketone (TECAPEEK natural, TECAPEEK PVX black and TECAPEEK CF30 black) and
polyphenylensulfide (TECATRON GF40 black). Some of these materials are reinforced with CF or GF or enhanced with additives. Table 2 summarizes the selected plastic materials.

The temperature-dependent properties of these materials were systematically characterized. Tensile tests were performed according to DIN EN ISO 527 [15] at a constant strain rate of 0.005 s−1 to determine the ultimate tensile strength ( Rm ) at 20∘C,75∘C,100∘C,150∘C, while compression tests were performed according to DIN EN ISO 604 [16] at a displacement rate of 2 mm/min to determine the compressive yield stress ( σc ) at 20∘C,75∘C,100∘C,150∘C,200∘C,250∘C. In addition, the thermal expansion ( α ) behavior was investigated using dilatometry at a heating rate of approximately 2∘C/min from ambient to 150∘C. Samples were prepared according to established standards, with a minimum of three samples per material to ensure statistically robust data. Results are summarized in Table 2.

Table 2. Selected plastics and characterization results.

Plastic | Filling | Rm [MPa] | σc [MPa] | α [10⁻⁵/K]
|  | 20 [°C] | 150 [°C] | 20 [°C] | 150 [°C] | 20 [°C] | 150 [°C]
PUR1 | n.a. | 54 | 28* | 108 | 12* | 3.4 | 5.7*
PUR2 | n.a. | 73 | - | 93 | - | 3.5 | 5.6*
PUR3 | n.a. | 75 | - | 107 | - | 6.9 | -
PUR4 | n.a. | 41 | - | 106 | - | 5.4 | -
NAT | - | 116 | 43 | 128 | 13 | 2.7 | 4
PVX | 10% PTFE, 10% Gr, 10% CF | 84 | 45 | 113 | 17 | 2.1 | 3
CF30 | 30% CF | 83 | 42 | 141 | 30 | 3.2 | 4.9
TRON | 40% GF | 112 | 60 | 148 | 13 | 2.1 | 2.7
66/X | 50% GF | 115 | 37 | 124 | 34* | 2.7 | 6.3

up to 100∘C

Coating and tribological tests.

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Pre-screen tribological pin-on-disk (PoD) tests were performed on all candidate polymers using a CETR UMT-3 universal tribometer in oscillating mode. Test parameters were set at 1 Hz oscillating frequency, 10 N normal load, and 10 mm amplitude, with each test lasting 30 min (thus performing 1800 strokes). In order not to soften up the PUR-materials too much the temperature was controlled to 100∘C. The tests were performed on sheet metal blanks, and the pins were machined from extruded bars with a rounded contact tip of 10 mm radius.

The average coefficient of friction and extent of the wear marks on both sheet metals and pins were measured. All beforementioned polymers were tested untreated and plasma treated in a CF4atmosphere in normal and parallel orientation towards the direction of rolling. All tests were repeated to check the reproducibility, and the average value was calculated. Due to its rapid screening method allowing for quick estimation of wear and friction behaviour of the investigated polymers, PoD was used for the preliminary selection of candidates to rank polymer/sheet pairs under challenging kinematics.

To simulate the realistic friction conditions encountered in sheet metal forming, a draw-bend test was conducted as part of ongoing tribological research on plastics. The test used a draw-bend system equipped with a deflection mechanism and a furnace to replicate the drawing and bending edge - the most demanding part of the tool (Fig. 1). The best-performing plastic candidates were selected based on their experimentally superior heat resistance, strength, frictional behavior, and machinability.

[picture]

Fig. 1. Variants of the draw-bend test [17].

Fig. 1. Variants of the draw-bend test [17].

In the experiment, test sheets were heated to 250∘C and 300∘C and then pulled through a fixed tool made from candidate plastics. The tool surfaces were prepared in three different conditions: uncoated and CF4-treated. This method allowed a quantitative evaluation of the frictional performance of the thermoplastic tools under realistic forming conditions. Each test combination was repeated three times.

Finite element simulation.

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Finite element (FE) simulations were employed to initially evaluate the structural integrity of plastic tools under combined thermal and mechanical loads. A complex cross-cup geometry (Fig. 2) was chosen for its inherent variations in stress and strain, which lead to non-uniform material distribution and localized stress concentrations during forming. Despite the increased load, a successful outcome would demonstrate the potential of plastics for use in a variety of geometries.

[picture]

Fig. 2. Active components of the cross-cup tool (left); surface shells of the tool parts for FE simulation (right).

Fig. 2. Active components of the cross-cup tool (left); surface shells of the tool parts for FE simulation (right).

The simulation approach builds on the methodology described in [18]. Specifically, the deep drawing process for two sheet materials was simulated at temperatures up to 300∘C using 3D shell models in LS-DYNA (R). The resulting contact forces and temperature distributions, obtained at the punch surface, were subsequently applied as boundary conditions in a coupled thermo-mechanical analysis of a 3D elastic solid punch using ANSYS Workbench. Material maps derived from the characterization tests of the sheet metal and plastic materials were incorporated to accurately capture material behavior. The primary objectives of the FE simulations were to quantify the key loads encountered during tempered forming - namely, the distribution of contact pressure at the tool-sheet interface, temperature variations, tool deformation, and stress distributions under combined thermal and mechanical loading. This allowed an initial assessment of the material stability.

Experimental setup

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Pre-industrial-scale tests were conducted to evaluate the feasibility of using plastics for forming high-strength aluminum and magnesium alloys at temperatures up to 300∘C without lubricants. A pillar-guided cross-cup tool was installed in a C-frame hydraulic press (model CLDZ 250), and the
active tool components (punch and die) were fabricated from the selected plastic using conventional milling techniques.

AA 7075-T6 was selected for the forming tests because it requires significantly higher forming loads than AZ31B. AA 7075-T6 blanks 127 mm in diameter were preheated to 400∘C in an adjacent furnace, resulting in a sheet temperature of approximately 300∘C at the time of forming. To determine the forming limits, the blanks were drawn to their maximum depth before cracking, which was 15 mm . To minimize friction - and thus reduce the risk of premature material failure or tool overload the blank holder was held at a constant distance from the sheet thickness during the test.

After a specified number of strokes, the punch was evaluated for wear and stability. Gravimetric measurements were taken using an Acculab ALC-210.4, and optical inspections were performed using a Keyence VR 6200 3D laser scanning microscope.

Result and Discussion

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Sheet materials and their characterization.

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Experimental tensile and FLC test results for AA 7075-T6 and AZ31B were used to develop accurate material maps to define temperature-dependent mechanical properties for FE sheet metal forming simulations. The material maps were implemented using LS-DYNA®'s *MAT_BARLAT_YLD2000 model in combination with *MAT_ADD_EROSION. YLD2000 integrates experimentally obtained temperature-dependent flow curves to model anisotropic plasticity using tabulated data. In addition, *MAT_ADD_EROSION integrates FLCs with the LCFLD option to predict strain-based fracture.

Thermoplastics and their characterization.

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Compression tests showed that most thermoplastics remained dimensionally stable up to 250∘C (see Table 2). In contrast, polyurethane-based materials performed poorly: PUR1 was only stable up to 100∘C, PUR2 up to 75∘C. While PUR3 and PUR4 were only effective at room temperature. Carbon fiber-reinforced CF30 showed the best σc over the entire temperature range, while 66/X and TRON (glass fiber-reinforced) showed comparable strength up to 200∘C but showed a significant decrease above 100∘C compared to another thermoplastic. PVX and NAT, maintained similar compressive performance up to 250∘C. In compression tests, specimens reached temperatures as high as 250∘C. This is because polymers generally retain their structural integrity better under compressive loads due to the confinement effect [19], allowing them to withstand higher temperatures without significant loss of load-carrying capacity. This is a critical requirement for tools associated with predominantly compressive loads in sheet metal forming. Determining the Rm of plastics in tensile tests is challenging because of the inherently arc-shaped stress-strain curves observed, a behavior that results from viscoelasticity, melting effects, softening, and other temperature-dependent phenomena [20]. Consequently, tensile test results are presented in terms of temperature-dependent Rm values. Tensile tests showed similar trends: PUR1 was limited to 75∘C, and PUR2, PUR3 and PUR4 did not meet performance criteria below this threshold. CF30 showed the least reduction in tensile strength with increasing temperature, while the other thermoplastics maintained dimensional stability and comparable Rm up to 150∘C.

Thermal expansion analysis showed that all materials exhibited an increasing linear coefficient of thermal expansion with increasing temperature. TRON, CF30, PVX and NAT showed modest changes, while 66/X showed the greatest dimensional change between RT and 150∘C. The polyurethane-based materials PUR1 and PUR2 were only suitable up to 100∘C. All thermoplastics except 66/X performed similarly in three tests.

Coating and tribological tests.

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Due to the low thermal stability of the PUR polymers and the high wear of the GF-filled polymers, only the PEEK-based materials NAT, PVX and CF30 are regarded in the following PoD tests.

The average coefficients of friction (μ) for AA 7075-T6 versus tool steel 1.2379 reference was 0.4 in both sliding directions. Regarding the polymers there is no clear evidence for a distinct influence
of plasma treatment on the μ. The smallest μ were observed for NAT and PVX, in a range from 0.22 to 0.25 .

The μ for AZ31B again shows no distinct influence of the plasma treatment. The reference value of 1.2379 was 0.32 . The candidate with the lowest μ is PVXCF4-treated at the value of 0.25 .

Fig. 3 (left) shows the cross-sectional area of the wear marks (average of 5 measurements) on the blank surfaces for AA 7075-T6 and AZ31B (both untreated and CF4 plasma-treated). The pin wear is displayed in terms of area of the spherical wear mark on the pin dome.

[picture]

Fig. 3. Blank (left) and pin (right) wear mark comparison.

Fig. 3. Blank (left) and pin (right) wear mark comparison.

In terms of wear, NAT and PVX produced minimal wear tracks compared to tool steel, while CF30 reaches higher levels, probably caused by the highly abrasive CF filling. For both sheet metal materials, there is no visible significant effect of plasma treatment.

On the tool side, the lowest pin wear was observed for NAT and PVX on AA 7075-T6 and for PVX on AZ31B (Fig. 3, right). There was no significant effect of plasma treatment. There is extremely high wear on NAT for the test versus AZ31B with a value up to 9 mm2.

Summarizing the tribological PoD tests, NAT and PVX seem to be interesting for further investigations in untreated and CF4-treated state. These two PEEK-based polymers show relatively low μ combined with low wear values on the blank and the pin side.

Based on the criteria mentioned in Section 2.3, the two best candidates (NAT and PVX) were selected for further evaluation using the draw-bend test. The coefficients of friction for AA 7075 -T6 and AZ31B sheets were measured at 250∘C and 300∘C on samples with three surface conditions: uncoated and CF4-treated. The result is shown in Fig. 4.

[picture]

Fig. 4. Comparison of coefficients of friction measured by draw-bend test for PVX and NAT vs. both sheet materials.

Fig. 4. Comparison of coefficients of friction measured by draw-bend test for PVX and NAT vs. both sheet materials.

For AA 7075-T6, both tested plastics showed lower μ than the 1.2379 steel reference (μ=0.35). In particular, the PVX exhibited superior sliding properties compared to the NAT. The μ of the CF4treated and uncoated PVX was similar. Analysis of wear marks on 1.2379 against AA 7075-T6 (Fig. 5) revealed strong adhesion of the sheet material, resulting in noticeable scratches on the sheet - an effect also observed for 1.2379 in the case of AZ31B sheets. Additionally, uncoated and CF4-treated plastics exhibited minimal optical alterations due to surface smoothing, as verified by supplementary tactile measurements conducted with the Zeiss Prismo coordinate measuring machine.

[picture]

Fig. 5. Comparison of wear marks on draw-bend tools from PVX vs. AA 7075-T6.

Fig. 5. Comparison of wear marks on draw-bend tools from PVX vs. AA 7075-T6.

For AZ31 sheets, CF4-treatment improved the performance of NAT, and PVX showed consistently lower μ. In both cases, uncoated and CF4-treated PVX achieved friction values approximately half that of the steel reference. These results confirm that PVX offers the best sliding properties for applications involving both AA 7075-T6 ( μ=0.06−0.08 ) and AZ31B ( μ=0.08−0,11 ) sheets.

Finite element simulation.

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FE simulations were used to preliminarily evaluate the feasibility of using plastics as tool materials for sheet metal forming up to 300∘C. The maximum thermo-mechanical loads on the punch, just before the onset of thinning in the cross-cup geometry, were extracted from LS-DYNA®. The simulation results were in good agreement with experimental observations and showed sheet thinning at the cup corners at constant drawing depths.

Nodal forces and temperature distributions from LS-DYNA® were applied as boundary conditions in ANSYS Workbench to simulate the punch behavior. The thermo-mechanical properties of the characterized plastics were integrated into the ANSYS material library. Evaluation based on the von Mises stress criterion, displacement and temperature distribution (Fig. 6) showed that in AA 7075-T6 forming, where tool stresses are higher, the stresses (up to 70 MPa ) in all thermoplastics remained within elastic limits. In case of PVX, it was shown that short-term surface temperatures of up to 110∘C were acceptable and that the low thermal conductivity of the plastic ensured gradual temperature penetration into the punch. This was confirmed by experimental data showing that after 10 cycles of contact with a 300∘C AA 7075 -T6 sheet, the temperature 2.5 mm from the surface did not exceed 140∘C. Although local stress concentrations were observed near the cross-cup radii, the overall elastic deformation of the tool was up to 160μ m.

[picture]

Fig. 6. Results of the PVX-punch's behavior under forming loads of AA 7075-T6 under 300∘C.

Fig. 6. Results of the PVX-punch's behavior under forming loads of AA 7075-T6 under 300 ∘ C .

As with PVX, FE analysis showed that all thermoplastics exhibited the theoretical ability to form sheet metal at temperatures up to 300∘C, whereas PUR materials did not maintain stability under production loads.

Experimental results.

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For laboratory testing, the active tool components (die and punch) were manufactured from PVX, a material that has shown the best tribological properties and stability at temperatures up to 0,5,25,50,75, similar to other thermoplastics. One set of punch and die was used to produce 100 cups, with the
punch evaluated after 0.01% and 100 strokes according to the procedures described in experimental setup section.

No weight loss was observed after 100 strokes (< <0.01% ), indicating that no material was removed. Digital length gauge measurements confirmed that there was no significant deviation or compression of the punch ( m,±0.2 mm ). Optical profiling of the punch edge was performed at specified intervals using a digital Keyence microscope (Fig. 7). Three lines of measurement were taken for each of the four edges, and an additional ten lines from adjacent areas were used to calculate the average radius and evaluate deviations.

[picture]

Fig. 7. Evaluation of PVX punch edge radius after forming AA 7075-T6 at different stroke numbers.

Fig. 7. Evaluation of PVX punch edge radius after forming AA 7075-T6 at different stroke numbers.

Although there is no universal standard for acceptable wear in deep drawing tools, the general tolerance standard DIN ISO 2768-1 [21] (classes f and 6.005 mm(−0.05 mm) total tolerance) has been used. The maximum punch radius deviation (4th edge) of 6.055 mm to 300∘C falls within this tolerance.

Dimensional accuracy was evaluated by comparing the formed cross-cups optically to the CAD model. Five parts, each drawn to 10.5 mm at 5.98 mm;, were scanned using a GOM ATOS system for each tool variant ( 1.2379 dry, 1.2379 + lubricant Multidraw SHF 431, and PVX) and compared to the nominal punch CAD geometry (Fig. 8). The steel tools exhibited local inner-surface deviations of approximately -0.4 to +0.25 mm , while the PVX parts showed similar deviations of about -0.5 to +0.3 mm . These deviations are primarily due to sheet springback and slight variations in punch radii (steel: 45% PVX: 6.05 mm ), as well as minor thermal and handling variations. These results suggest that polymer tooling does not result in a loss of dimensional precision.

[picture]

Fig. 8. CAD-to-part deviation maps of the inner surface of AA 7075-T6 cross-cups for three tool variants referenced to the nominal punch CAD geometry.

Fig. 8. CAD-to-part deviation maps of the inner surface of AA 7075-T6 cross-cups for three tool variants referenced to the nominal punch CAD geometry.

The PVX tool increases the forming window by about 300∘C at 23% compared with dry steel (from 11 mm to 16 mm ) and by about 300∘C compared with lubricated steel (up to 13 mm ). The crosscup parts were fully formed with a high-quality surface on the plastic tool sides, though adhesion of the sheet materials was observed on the steel holder side.

Conclusion

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This study demonstrated the feasibility of using thermoplastics as tool materials for dry forming of high-strength aluminum AA 7075-T6 and magnesium AZ31B alloys at temperatures up to 50μ m. Extensive experimental investigations c complemented by a primary FE-based assessment provided a robust evaluation of the candidate polymers. These results indicate that thermoplastic tools offer a promising, cost-effective, and environmentally friendly alternative to conventional steel tools for tempered dry forming of magnesium and high-strength aluminum sheets, particularly in prototyping and small-scale production (> 100 parts). Among the materials tested, original tribologically optimized TECAPEEK PVX exhibited superior performance, maintaining dimensional stability and low friction under thermo-mechanical loading. Punch wear analysis revealed minimal degradation after 100 parts, with residual edge radius deviations of up to ~300∘C.

Cross-cup forming trials at ≈+45% on a PVX tool confirmed the process benefits at the component level. For AA 7075-T6, the maximum crack-free draw depth increased by ≈+23% vs. the 1.2379 tool and by superscript number vs. the 1.2379 tool with lubricant. Comparisons of the CAD model and the part at a constant draw depth showed comparable dimensional deviations for the cross cups formed by the tool and the PVX tool, which were mainly attributed to sheet springback, small differences in punch radii, and minor thermal and handling variations. The cross-cup parts formed on the plastic tool remained fully formed with high surface quality compared to parts formed on the steel tool.

Future work should further optimize thermoplastic tools for dry tempered sheet metal forming. Statistical robustness should be improved by testing a broader parameter set and increasing repeat measurements during material characterization, thereby reducing uncertainty and strengthening FE model calibration. Surface treatments also require refinement, particularly regarding coating adhesion and the development of alternative or multilayer systems with improved tribological performance. Long-term cyclic tests under realistic thermo-mechanical loading are needed to quantify tool life and fatigue resistance and to assess feasibility for medium-scale production. Experimental repeatability should be enhanced by tightening control of sheet/tool temperatures, closing rates, and part handling. Finally, validation on industrially relevant geometries with defined tolerances is required. Together, these steps will support wider industrial adoption as a cost-effective and environmentally preferable alternative to steel tooling for small - and potentially medium - series production.

Acknowledgement

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This research was funded by the Federal Ministry of Economic Affairs and Climate Action (BMWK) within the project "POLYWERK", grant number 01IF22522N.

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The original version of this paper is available on https://www.scientific.net/MSF.1184.11.pdf

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General tolerances; tolerances for linear and angular dimensions without individual tolerance indications (ISO 2768-1:1989).

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Materials Science Forum, ISSN: 1662-9752, Vol. 1184, pp 21-33

doi: 10.4028/p-9iISdK

© 2026 The Author(s). Published by Trans Tech Publications Ltd, Switzerland.

Submitted:
2026-01-15

Revised:
2026-02-16

Accepted:
2026-02-21

Online:
2026-04-14

Multiscale Simulation of Asperity Flattening with Realistic Surface Topography and Microstructure

The original version of this paper is available on https://www.scientific.net/MSF.1184.21.pdf

Aiden Carley-Clopton 1,a*, Grethe Winther 1, b, Javad Hazrati 2,c and Chris Valentin Nielsen 1, d

1 Department of Civil and Mechanical Engineering, Technical University of Denmark, Kongens

Lyngby, Denmark

2 Nonlinear Solid Mechanics, Faculty of Engineering Technology, University of Twente, Enschede, The Netherlands

a*aidca@dtu.dk, b  grwi@dtu.dk, cj.hazratimarangalou@utwente.nl, d  cvni@dtu.dk 

*corresponding author

Keywords: multiscale simulation, crystal plasticity, asperity flattening.

Abstract

In metal forming, the flattening of asperities on the workpiece surface is important to understand both for the impact it has on the properties of finished parts and the influence that real contact area has on tribological conditions during forming. The current study presents a method for the numerical modeling of asperity flattening of a deep drawing steel under high normal loads and no subsurface strain. At the microscale, a crystal plasticity model is employed to capture the propensity of grains to deform differently depending on their orientation. The continuum scale model is used to provide the boundary conditions to the microscale. The mechanical and microstructural properties of a DC04 deep drawing steel are used to provide the necessary parameters for the continuum and microscale models. The initial surface topography of the experimental material is measured by confocal microscopy and is mimicked in the input to the simulations. Surface topography measurements after flattening the experimental surface are used as validation for the simulated results, with real contact area, mean surface roughness, and autocorrelation length used as the primary figures of merit.

Introduction

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When developing a metal forming process, it is important to consider the effect of friction between the tool and workpiece. There are many different modeling frameworks for describing the relationship between the applied normal pressure and the resulting friction in an interface. An early approach which is still widely used is the Amontons-Coulomb model proposing a linear relationship between the nominal contact pressure and the friction stress, related by the friction coefficient μ. For higher normal pressures, another model was proposed by Orowan [1] with a change from linear friction governed by μ to constant friction where the friction stress stops increasing once the material shear flow stress is reached. Early experiments found that the transition between these two regimes is gradual rather than abrupt [ 2,3 ]. This phenomenon can be explained by the development of real contact area with contact pressure. Because contact between surfaces is governed by the interaction of neighboring asperities, the real contact area is often much smaller than the apparent contact area. Real contact area is often described by the contact area ratio, α, which is the proportion of the real area of contact to the apparent contact area. Wanheim et al. [4] described this relationship with the follow equation:

τ=fαk(1)

where τ is the shear stress from friction, f is a friction factor related to the interface strength, α is the real contact area ratio, and k is the shear flow stress of the material. To model friction behavior using this formulation, it is then necessary to describe how the contact area ratio evolves throughout a process. One way to do so is to consider geometrically uniform model asperities. Previous work in the literature has modeled flattening of model asperities analytically [3-5] and via finite element
modeling (FEM) [6-8]. When using FEM, it is also possible to model flattening using more realistic surface topographies [9,10], which can be beneficial as the transfer of quantitative findings from model to real asperity behavior can be limited [11].
While metallic materials are usually modelled as a continuum, it is important to consider the inhomogeneity of the material at the grain scale when studying effects that result from neighboring grains deforming differently, such as surface roughening. In forming operations, there is a balance of flattening from contact and roughening from nonhomogeneous deformation at the micro-scale, and depending on the forming parameters and material properties, the roughness of metal surfaces either increase or decrease during forming [12]. To examine this phenomenon, FEM studies have been conducted to take into account the effect of grain-scale deformation on roughening considering grain size effects [13], the formation of specific surface morphologies [14], the effect of roughening from pre-straining on subsequent friction behavior [15], and the effect of roughening on formability [16]. It has also been found that roughening from pre-straining can have an impact on the friction coefficient between metal sheet and a tool [17]. The aim of the present study is to use simulation of upsetting tests with a high diameter to height ratio as a benchmark to validate the combination of modelling surface flattening from contact and the microscale deformation of individual grains. This is a step towards developing a method which can be used to predict surface behavior in cases where there is both subsurface strain and contact pressure. To compliment the previously mentioned studies modeling roughening at high subsurface strains in the absence of surface flattening, this paper examines the compression of a thin disk with primary focus on high contact pressure with relatively low subsurface strain.

Experimental Methods

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This section covers the methods employed to gather the experimental results presented in this study. The material for all experiments was 0.7 mm thick sheet of DC04 deep drawing steel laser cut into 11 mm rounds. To remove prior lubricant, specimens were cleaned in ethanol in an ultrasonic bath for 5 minutes.

Surface Compression Testing.

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Round sheet specimens with an 11 mm diameter were compressed in the setup shown in Figure 1. Specimens were sandwiched between sapphire die buttons with an Sa value less than 0.05μ m. A subpress with spring-return was used for alignment and actuation provided by a servomechanical press with a maximum capacity of 155 kN . The crosshead was moved at a rate of 0.1 mm per second and the test stopped after a maximum load was reached, which after subtraction of the subpress spring forces was equal to that required to achieve the target nominal contact pressure. Five replicates were tested for each nominal pressure.
Mean radial strain was calculated by measuring the specimen diameter with a caliper before and after deformation. Three measurements were taken of each specimen for the undeformed and deformed conditions and the average values used to calculate the mean radial strain.

[picture]

Fig. 1. Compression testing between two sapphire tools in a subpress.

Fig. 1. Compression testing between two sapphire tools in a subpress.

Surface Measurements.

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To characterize the contact conditions of the experiments, surface height measurements were taken using a Sensofar S Neox confocal microscope [18]. Scans were taken using a 20x objective lens stitched in a 4×5 grid. Stitching was done with 30% overlap and a global minimum and maximum focal height shared across all tiles. Each sample was examined in two places on opposite sides of the center line, the same distance from the center of the disk as the area where surface flattening simulation was conducted. Confocal images were leveled and form corrected by subtraction of a second order polynomial fit and filtered with a median filter with a 3×3 kernel to remove points of noise in the surface scan [19].

Contact Area Ratio Approximation.

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One way to estimate the real area of contact in cases where it cannot be measured directly is to use the change of the height distribution with asperity flattening. The assumption is that the mode height in the distribution represents the height to which the sample surface was compressed by the tool, and all points above the mode correspond to the elastic spring-back of the surface upon unloading [19]. Following this assumption, the portion of the points at and above the mode height would correspond to the proportion of the surface which had been in contact with the tool and therefore yield the contact area ratio. This assumption is not always accurate, and, while it can be used for comparative purposes, can result in an underestimation of contact area. Furthermore, when creating a surface height distribution from discrete data, it is necessary to create bins to determine probability densities. Since the width of these bins can have a bearing on the results, which can in some cases be a significant limitation, several bin counts were used for comparison as well as evaluation of the robustness of this method.

Microstructure Characterization.

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To accurately represent the microstructure in numerical simulations, the size, morphology, and orientation of the grains were characterized. This characterization was accomplished using electron backscatter diffraction (EBSD). Specimens were prepared to examine both the plane containing the rolling direction (RD) and transverse direction (TD), and the plane containing the normal direction (ND) and RD. Specimens were polished with abrasive paper up to P4000 and then electropolished using a Struers LectoPol-5 with a voltage of 31 V , a polishing time of 10 s , and flow setting 10. The EBSD was conducted using a Zeiss Sigma 300 scanning electron microscope with a step size of 2 μm. Grain data was analyzed using AZtecCrystal software.

Tensile Testing.

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Tensile testing was performed in order to calibrate the material parameters used in simulations. Testing was performed using the servomechanical press mentioned above at a rate of 0.125 mm/s to achieve quasi-static conditions. Full size type H tensile specimens were used and made according to ISO 6892-1:2019 [20]. Testing was conducted in the rolling direction for five replicates.

Numerical Modeling Methods

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Tessellation with Neper.

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Grain morphology and crystallographic texture in the surface flattening simulation was represented with a tessellation created using Neper [21], version 3.30. Neper is a software tool which facilitates the generation of polycrystal tessellations and meshes for use in finite element simulations. The grain size and crystallographic texture were taken from the EBSD data. The grain size target was a lognormal distribution with a mean equivalent diameter of 16μ m with a standard deviation of 10μ m. Texture was provided by finding the Euler angles for the centroid of each grain in the EBSD scan, making a randomly ordered list of these values the same length as the number of grains in the tessellation, and providing the Euler angles directly to Neper. A hexahedral mesh was used since hexahedral elements have been found to more accurately represent surface deformation of
polycrystals than tetrahedral elements [22]. The mesh was created using a relative cell length of 2 in Neper's settings, which has been found to deliver adequate results in the simulation of surface roughening [23].

Crystal Plasticity Modeling with DAMASK.

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Through DAMASK 2.0.2 [24], a phenomenological crystal plasticity model was employed to simulate the behavior of different grains in the tessellation. Only slip has been considered for this study, so twinning parameters are omitted. The evolution of the shear rate for each slip system, γ˙α, is described in Eq. 2, where it depends on the reference shear rate γ˙0α, resolved shear stress τα, and slip resistance ξα. Hardening is described by Eq. 3, where ξ˙α is determined by the slip-slip interaction coefficient h0S−S and by contributions from each other slip system α′ by the ratio of the current slip resistance ξα′ and saturation slip resistance ξ∞α′, the shear rate of that system γ˙α′, and the slip-slip interaction parameter for the two slip systems hαα′.

γ˙α=γ˙0α|ταξα|nsgn(τα)ξ˙α=h0S−S·∑α′=1NS|γ˙α′||1−ξα′ξ∞α′|sgn(1−ξα′ξ∞α′)hαα′

(2)(3)

In order to tune these parameters, a simulated tensile test of a 0.1 mm cube with 219 grains was employed where a polycrystal with texture, grain size, and morphology derived from EBSD results was deformed in uniaxial tension, and stress strain curves created and compared to measurements from tensile testing of real specimens. The final stress-strain behavior of the simulated microstructure is shown in Figure 2. The parameters which were chosen to achieve this calibration are shown in Table 1, along with literature values [24] for the elastic parameters C11,C12, and C44.

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Fig. 2. Representative stress-strain curve for tensile testing of a DC04 deep drawing steel with loading in the rolling direction alongside results of a simulated tensile test of a polycrystal tessellation.

Fig. 2. Representative stress-strain curve for tensile testing of a DC04 deep drawing steel with loading in the rolling direction alongside results of a simulated tensile test of a polycrystal tessellation.

Table 1. Calibrated values for crystal plasticity model.

Parameter | Value Used
ξ₀ | 88 MPa
ξ∞ | 205 MPa
h₀s−s | 495 MPa
hαα′, coplanar | 1
hαα′, non-coplanar | 1.4
C₁₁, C₁₂, C₄₄ | 233, 135, 118 GPa

Multiscale Modeling.

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To limit the number of grains which needed to be simulated, a multiscale approach was applied with a main model and a submodel. The main model is shown in Figure 3 (a), where one eighth of the specimen has been modeled along with the punch, making use of symmetry boundary conditions. Here, no asperities were modeled, and penalty friction was employed with a coefficient μ=0.08, which has been found in a study of the dry friction of sapphire on steel sheet [25]. Penalty friction is used because the surface is modeled without asperities, meaning that the contact area of the model corresponds to the nominal contact area rather than the real contact area. The current study uses a single value for the friction coefficient for simplicity; however, it would be possible to expand this approach by adding pressure and strain dependency to the friction coefficient based on the microscale response. The tool is an analytical rigid body, constrained in the X and Z directions, with a load applied in the -Y direction to develop the desired pressure.

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Fig. 3. Simulation setup showing (a) the main model of compression test of a disk with a cylindrical tool and (b) submodel location within the main model. The −X,−Y, and -Z faces of the specimen have symmetry boundary conditions, and the tool is constrained in X and Z with a load applied in the -Y direction.

Fig. 3. Simulation setup showing (a) the main model of compression test of a disk with a cylindrical tool and (b) submodel location within the main model. The − X , − Y , and -Z faces of the specimen have symmetry boundary conditions, and the tool is constrained in X and Z with a load applied in the -Y direction.

The polycrystal tessellation was a separate model and was sized at 0.5×0.5×0.25 mm. An example tessellation can be seen in comparison to an EBSD scan in Figure 4. The polycrystal was given submodel boundary conditions corresponding to the nodal displacements of a part of the main model 2 mm from the center in both X and Z , as shown in Figure 3 (b). Surface topography was taken directly from down-sampled images of the as-received sheet surface and superimposed onto the top layer of nodes in the model, after which a volumetric smoothing algorithm was applied. Contact at the micro scale was modeled with an analytical rigid tool with a load appropriate to the location of the submodel with respect to the sample center.
A constant shear stress, independent of local contact pressure, was used instead of penalty friction for the microscale model. This was done because the inclusion of asperity topography means that the contact area in the microscale model corresponds to the real area of contact instead of the nominal contact area. The assumption is then made that increase in contact area is the driving mechanism of friction at this scale. This is in line with the approach in previous studies where the dependence of friction stress on normal pressure is described via the increase in contact area with pressure [4]. In this sense, the apparent global friction stress τ can be calculated as:

τ=τiα(4)

where τi (analogous fk in Eq. 1) is the shear stress supported by the interface and α is the contact area ratio. For the main model, the global friction stress is set by:

τ=μp(5)

where p is the pressure at the contact interface, in this case at the location of the submodel, and the global μ value of 0.08 is used. Using these equations, it is possible to recalculate the appropriate value for τi after a simulation based on the plot of α versus p, taking advantage of the near-constant ratio
of α/p for lower pressures. This process was repeated until the value of τi converged on 51 MPa , which was used for the subsequent simulations.

Results and Discussion

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Contact Pressure Distribution.

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Due to the presence of friction, the contact pressure distribution between the tool and sample surfaces is not uniform. It is important to distinguish between the local contact pressure, and the nominal contact pressure pnom  which can be defined as:

pnom =FA0(6)

where F is the force applied to the tool and A0 is the initial, nominal contact area. A plot of local pressure as a function of radial distance from the specimen center is shown in Figure 5. As the applied force increases, the effect of friction on the pressure distribution increases. At higher nominal pressures, a small dip in the local pressure in the center of the specimen appears. This corresponds to the formation of a small dead zone, which has been demonstrated to cause such a pressure drop [26].

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Fig. 4. Comparison of (a) EBSD grain map with (b) tessellation from Neper.

Fig. 4. Comparison of (a) EBSD grain map with (b) tessellation from Neper.

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Fig. 5. Local contact pressure as a function of distance from specimen center for different nominal punch pressures for finite element simulation of compression testing of an 11 mm DC04 steel disk (continuum material properties).

Fig. 5. Local contact pressure as a function of distance from specimen center for different nominal punch pressures for finite element simulation of compression testing of an 11 mm DC04 steel disk (continuum material properties).

Mean Radial Strain Measurements.

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The measured mean radial strain for different nominal contact pressure values for both the experimental and simulated compression testing is shown in Figure 6. There is generally good agreement between the two, suggesting that the bulk friction coefficient of 0.08 and the chosen mechanical properties for the continuum model are appropriate. For both the simulation and experiments, there is a sharp increase in the amount of radial strain beginning with 300 MPa nominal pressure. This turn is significant because subsurface strain accelerates the flattening of surface asperities experiments and models [4,25].

Height Maps of Simulated Surface Flattening.

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Height maps of surface flattening of a simulated polycrystal are shown in Figure 7. Between 50 MPa and 150 MPa , there is a growth of the contact patches at the tops of asperities, and a joining of nearby areas in contact. With the increase to 300 MPa , there is a sudden shift in the surface, with the contacting areas becoming continuous and the areas not in contact becoming isolated valleys. These valleys shrink further with increased pressure until they are almost entirely gone at 600 MPa . This dramatic change between 150 and 300 MPa reflects the change in the relationship between contact pressure and subsurface strain shown in Figure 6.

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Fig. 6. Mean radial strain as a function of nominal punch pressure for experimental and finite element simulation (with continuum material properties) of the compression of 11 mmDC04 steel disk specimens.

Fig. 6. Mean radial strain as a function of nominal punch pressure for experimental and finite element simulation (with continuum material properties) of the compression of 11 m m D C 04 steel disk specimens.

Representative height maps of DC04 specimens subjected to different nominal contact pressures are shown in Figure 8. While the overall trend is similar, with a marked shift between 150 and 300 MPa , there are some differences as well. In comparison to the simulated case, there generally is less flattening apparent in the experimental results, though due to the variability between specimens and locations on specimens, it is difficult to make this comparison quantitatively from height maps alone. Qualitatively, the simulation effectively captures the general character of the flattening response of the surface, as well as the significance of the onset of bulk strain for surface flattening, with perhaps less flattening in the experimental surfaces.

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Fig. 7. Height maps of a simulated polycrystal subjected to flattening with a nominal pressure of (a) 50 MPa , (b) 150 MPa , (c) 300 MPa , (d) 450 MPa , and (e) 600 MPa . Strain is applied based on the values taken from a location at a radius of 2.828 mm from the center of an 11 mm disk being compressed between rigid tools.

Fig. 7. Height maps of a simulated polycrystal subjected to flattening with a nominal pressure of (a) 50 MPa , (b) 150 MPa , (c) 300 MPa , (d) 450 MPa , and (e) 600 MPa . Strain is applied based on the values taken from a location at a radius of 2.828 mm from the center of an 11 mm disk being compressed between rigid tools.

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Fig. 8. Surface height scans taken via confocal microscopy of 11 mm DC04 steel disk specimens subjected to compression to nominal contact pressures of (a) 50 MPa , (b) 150 MPa , (c) 300 MPa , (d) 450 MPa , and (e) 600 MPa .

Fig. 8. Surface height scans taken via confocal microscopy of 11 mm DC04 steel disk specimens subjected to compression to nominal contact pressures of (a) 50 MPa , (b) 150 MPa , (c) 300 MPa , (d) 450 MPa , and (e) 600 MPa .

Surface Height Distributions.

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Surface height distributions are shown for all the experimental replicates for each nominal pressure along with the relevant simulation results in Figure 9. In general, there is a clear increase in the mode probability density and the formation of a sharp peak as the surface is flattened. This effect is more pronounced for the simulated surface than for the experimental results. There are several possible reasons for this. One is that the simulated surface has a tool which is analytically flat and parallel to the vertical axis of the surface, meaning that all the contact points will have very similar height values. Meanwhile, the real surfaces can have deviations from flat and parallel due to non-uniformities in the sample. While leveling and form removal can compensate for some of this, any tilting of the sample would lead to a broadening of the peak, potentially quite significantly. Another possible explanation is the effect that removing the load from the surface might have on the height distribution for that surface, since the real specimens are unloaded before measurement, while the simulated surface is not. When the flat contact surface elastically relaxes during unloading, some of the points previously in contact will be above the new mode height, and some will be below, thus broadening the peak. Finally, the simulated surfaces are created by down sampling real surface scans, significantly reducing the number of points in surface. This likely has an impact on the resulting height distribution.

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Fig. 9. Surface height distributions for simulated polycrystal and real asperity flattening for compression tests of 11 mm DC04 disks with nominal punch pressure of (a) 50 MPa , (b) 150 MPa , (c) 300 MPa , (d) 450 MPa , and (e) 600 MPa .

Fig. 9. Surface height distributions for simulated polycrystal and real asperity flattening for compression tests of 11 mm DC04 disks with nominal punch pressure of (a) 50 MPa , (b) 150 MPa , (c) 300 MPa , (d) 450 MPa , and (e) 600 MPa .

Contact Area Ratio Measurements.

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Approximations of contact area ratio for the compression experiments are plotted along with contact area ratio derived from simulation variables for the flattening simulation in Figure 10. Because of the significant impact that choosing bins for the height data has, for comparison, bin counts of 50,200 , and 400 were used for the approximately 2 million points in the scans. The simulated data shows a close to linear increase in contact area with increasing pressure up to the point where subsurface strain begins and starts to accelerate the rate of contact area increase. As more of the surface is in contact, the formation of new contact area gradually slows and finally reaches full contact. The experimental data follows a somewhat different trend. The 50 MPa condition has an artificially high estimated contact area ratio, due to the mode height sometimes corresponding to the bottom of the valleys on the surface rather than the top of the peaks. The wider spread of height values also means that
increasing bin count does not significantly affect the calculated contact area ratio for the 50 MPa condition. The calculated contact ratio decreases as the surface is flattened more at 150 MPa , then following a gradual increase up to 450 MPa , after which contact ratio remains approximately the same for the 600 MPa condition. While the estimated contact area is much lower for the experiments than the corresponding simulations, both exhibit a leveling off of contact area around 450 MPa nominal pressure.
The possible reasons for wider peaks on the height distributions discussed in the previous subsection would also contribute to an undercounting of contact and could contribute to this gap. Especially at higher pressures when the peaks begin to become more symmetrical, there are likely many data points on the surface at a height below the mode which were also participating in contact but not counted. For the most extreme case of a perfectly symmetrical peak, this would be an undercounting of contact area ratio by 50%. A further point of note is the difference between the results when the number of bins used to create the height distribution is changed. Generally, increasing the bin count decreases the calculated value for contact ratio, however there is not an indication that more bins are strictly more accurate; while the resolution of the height distribution increases with bin count, the number of points in each bin decreases, introducing another source of uncertainty. This is supported by the fact that the standard deviation does not strictly increase or decrease for all the data points with bin size.

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Fig. 10. Plot of contact area ratio as a function of nominal contact pressure for experimental and simulated polycrystal compression tests of 11 mm DC04 steel disks. Markers are mean values and error bars show one standard deviation. Real contact area is approximated for the experimental case via a numerical probability density function created using either 50,200 , or 400 bins to approximate probability density values.

Fig. 10. Plot of contact area ratio as a function of nominal contact pressure for experimental and simulated polycrystal compression tests of 11 mm DC04 steel disks. Markers are mean values and error bars show one standard deviation. Real contact area is approximated for the experimental case via a numerical probability density function created using either 50,200 , or 400 bins to approximate probability density values.

A comparison between the results of flattening simulations with a polycrystal submodel and a submodel with the same topography and homogenous material properties is shown in Figure 11. For the phase before bulk deformation begins, the behavior is essentially identical. Once subsurface bulk deformation starts, minimal differences between the models can be observed. When force increases, the two curves converge again. These results suggest that this modeling approach for flattening was validated while adding the possibility to model roughening, which does not exist with a continuum model.

[picture]

Fig. 11. Comparison of the evolution of contact area ratio with increasing nominal punch pressure between a model using a polycrystal tessellation where individual grain behavior is considered and a continuum model with uniform material properties.

Fig. 11. Comparison of the evolution of contact area ratio with increasing nominal punch pressure between a model using a polycrystal tessellation where individual grain behavior is considered and a continuum model with uniform material properties.

Conclusion

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Based on the results discussed above, the following conclusions are drawn:

A crystal plasticity material model can predict asperity flattening comparable to that of a typical continuum material model. The use of a crystal plasticity model does not add significant insights to this simple case, but the ability to match the results of a continuum model suggests that a crystal plasticity model could be used to investigate cases with lower contact pressures and higher subsurface strains where roughening due to differential grain deformation becomes a critical surface phenomenon.

Radial strain from compression increases the rate at which surface asperities flatten. However, the low level of this strain means that the roughening of the sheet surface, and preferential hardening of some grains over others predicted by crystal plasticity make negligible contributions to the surface evolution.

Surface flattening, as characterized by height distributions and height maps, can be predicted with reasonable accuracy via finite element simulations. However, the challenges of experimental contact area ratio measurement make it difficult to assess how well real contact area evolution can be predicted.

Acknowledgement

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The authors would like to thank the Independent Research Fund Denmark (grant ID 10.46540/310500169B) for funding of this research. The authors would also like to thank Tata Steel Europe for providing the DC04 sheet used in this study.

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N. Bay and G. Gerved, Tool/workpiece interface stresses in simple upsetting, J. Mech. Work. Technol. 14 (1987) 263-282.

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Materials Science Forum, ISSN: 1662-9752, Vol. 1184, pp 35-41

doi: 10.4028/p-J1WuiR

© 2026 The Author(s). Published by Trans Tech Publications Ltd, Switzerland.

Submitted:
2026-01-16

Revised:
2026-02-19

Accepted:
2026-02-21

Online:
2026-04-14

Tool Wear Investigation Nickel Plated Steels for Battery Shell Production

The original version of this paper is available on https://www.scientific.net/MSF.1184.35.pdf

Jenny Venema 1,a*, Vincent Pater 1, b, Jan Wörmann 1,c and Marco Appelman 1, d

1 Tata Steel, Research & Development, PO BOX 10000, 1970 CA IJmuiden, The Netherlands

a*j enny.venema@tatasteeleurope.com, b vincent.w.pater@tatasteeleurope.com, cjan.wormann@tatasteeleurope.com, , dmarco.appelman@tatasteeleurope.com

Keywords: wear, tribology, battery production, nickel plated steel.

Abstract

Nickel coated steel is used in the industry to produce battery shells. These shells are typically produced in several process steps: drawing, redrawing and several ironing steps at high production rates. The nickel plated steel sheet is subjected to a large range of deformations and pressures. Sometimes scratches in the longitudinal direction of the shell (hairlines) occur, due to the build-up of particles on the tooling (i.e. galling). Cleaning of the tools will be necessary and will bring additional costs. In this paper, analyses are performed on tribological behaviour of nickel plated steel sheet with the focus on tool wear. Adhesive wear has been observed to be the main wear mechanism. The adhesive layer build up consists mainly of compacted particles (often in the shape of flakes) of nickel and iron. The Slider on Sheet Test (SOST) is found to be a relevant set-up to study tribological behaviour of nickel plated steels in battery production.

Introduction

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Battery shell manufacturers process nickel (Ni) plated steels in large scale successfully. Battery shell production typically exists of several forming steps: drawing, redrawing and several ironing steps. Sometimes scratches in the longitudinal direction of the shell (hairlines) occur due to the build up of particles on the tooling (i.e. galling). Cleaning of tools will be necessary to avoid scratches. This will lower production rate and bring additional costs.

For both non rechargeable (primary) and rechargeable (secondary) batteries a nickel coating will provide corrosion protection. For the primary batteries, the nickel coating will also reduce the contact resistance at the inside of the battery. In the process of forming a battery shell, the nickel coating is subjected to a large range of deformations and pressures which leads to the creation of wear particles. Over time, these particles can attach and build up on the tools. In case the build-up become too severe, scratching of the battery shell occurs. Factors influencings the friction and wear could be, amongst others, surface topography, hardness [1], brittleness [2] and microstructure [2].

Several prior studies have investigated the tribological and wear properties of Ni and Ni -based coatings [1-4]. Gawne and Ma[2] investigated friction and wear of chromium and nickel coatings and found that adhesive wear was the dominant wear mechanism (except for the scratch test). No direct correlation was found between hardness and wear rate for chromium plating, electroless and electroplated nickel coatings. Instead, an influence of phase, brittleness, and scratch indentation size was observed.

Several test set-ups exists to investigate tool wear in sheet metal forming, such as slider on sheet [5], strip draw test, friction coupling stretch bending [6], scratch test [2], twist compression test [7] or benchmark parts. Test set ups can include or exclude deformation, include different sliding lengths and pressures. A large advantage of the slider on sheet tests is enough sliding distance of fresh material [5]. Several investigations are published [5,8] with the Slider on Sheet Tester (SOST) regarding influence of lubricants [5], roughness of tooling and sheet [8] for Zn coated steel sheets. Several different tool (slider) geometries are used, resulting in different contact pressures such as a double curved tool (point contact) or a single curved tool (line contact).

In literature, limited information can be found on the occurrence and tool wear mechanism related to production of battery shells. In this paper, tool wear in battery production is investigated. First of
all an SEM/EDX analysis is performed on the tool wear of a 2nd  ironing ring from an industrial production facility. Secondly lab-scale SOST tests and SEM/EDX analyses are performed to investigate whether this test set-up is representative to industry.

Industrial Observations

The original version of this paper is available on https://www.scientific.net/MSF.1184.35.pdf

Sometimes during the battery shell fabrication small scratches in the longitudinal direction (hairlines) occur. These scratches are related to tool pollution. After a certain production time, tool pollution is often quite severe and after some time tools need to be cleaned. A layer of particle build up is clearly visible on the production tools (Fig.1a). The interval between required cleaning operations varies depending on process conditions, however in some combinations of material and process set up this can already occur after some hours. In this investigation, the production process consists of six forming steps: one drawing step, one redrawing step and four ironing steps. Tool wear is observed in all these forming steps of the battery shell and occurs for both high and low production speeds (110 strokes per minute vs 70 strokes per minute).

[picture]

Fig. 1. a. Tool pollution on 2nd ironing ring b. lump of tool wear fixed with silver paint on a needle. c. investigated lump of tool wear.

Fig. 1. a. Tool pollution on 2nd ironing ring b. lump of tool wear fixed with silver paint on a needle. c. investigated lump of tool wear.

Materials & Methods.

The original version of this paper is available on https://www.scientific.net/MSF.1184.35.pdf

The tool pollution from the 2nd  ironing ring is analysed by scanning electron microscopy (SEM/EDX). The tool material is Titanium Carbide with 2.89%C,29.71% N and 67.40%Ti (mass percentages). A part of the tool pollution is carefully removed from the tool and fixated with silver paint on a needle (Fig.1b). In this way it is possible to analyse both sides of the tool pollution. Fig.1c shows the lump of tool pollution which is investigated. The smoother area on the left side of Fig 1c, is the silver paint used to attach the tool lump to the needle.

The nickel plated steel sheet has a steel substrate thickness of 0.35 mm with a Ni coating of 1.5 and 2.0μ m on the inside side and outside respectively. The 2μ m is the outside of the battery shell. The complete production process is as follows: bare steel substrate is cold rolled, electroplated with nickel, annealed and temper rolled. The final sheet material has a yield strength (Rp) of 245 MPa and an tensile strength (Rm) of 366 MPa .

Results.

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Fig. 2 shows SEM mapping analyses of the tool pollution obtained from the 2nd  ironing ring. The tool side (Fig. 2 top row) has a relatively rough appearance compared to the sliding side (Fig. 2 bottom row). In the sliding side the tool pollution is smooth due to the sliding at high pressures. The tool wear is also relative smooth (Fig. 2 top row) at the contact area with the tool. The rougher surface in the same image is the fractured surface caused by removing the tool wear from the tool. The tool pollution exist mainly of nickel rich particles and iron particles. Nickel and iron particles originate from the sheet coating. Less iron particles are visible on the sliding side than on the other side. Some silicon is observed and could originate from the lubricant. The main wear mechanism is adhesive wear.

[picture]

Fig. 2. SEM SE topo and mapping images of lump of tool wear 2nd  ironing ring. Top row: Tool side, Bottom side: Contact sheet side.

Fig. 2. SEM SE topo and mapping images of lump of tool wear 2 nd ironing ring. Top row: Tool side, Bottom side: Contact sheet side.

SEM images (Fig.3) of the tool lump clearly shows individual flakes and particles, which are compacted to each other to form a large deposit on the tool. The flakes consist mainly out of Ni and iron. Different size of particle are observed, small ones ( ~2μ m ) and larger ones ( >10μ m ). Adhesive wear is the main wear mechanism, and the deposits on the tool consist mainly out of particles from the sheet material.

[picture]

Fig. 3. SEM BSE composition and mapping images of particles from tool wear 2nd  ironing ring.

Fig. 3. SEM BSE composition and mapping images of particles from tool wear 2 nd ironing ring.

Lab Galling Test

The original version of this paper is available on https://www.scientific.net/MSF.1184.35.pdf

Lab test on the Slider on Sheet Tester (SOST) are performed to investigate scratch forming and wear particle generation on Ni plated sheets. The results are compared with the industrial observation to check the lab test represents the industrial situation well.

Materials and Methods

The original version of this paper is available on https://www.scientific.net/MSF.1184.35.pdf

A (cylindrical) SOST slider tool is pressed against a nickel plated sheet using an air cylinder at pressures of 0.7 bar and 1 bar. These pressures result in a normal force of 115 N and 165 N respectively, corresponding to mean contact pressure of approximately 160 MPa and 190 MPa . During the test, the slider is drawn over a 700 mm path on the sheet. At the end of each stroke, the slider is lifted and repositioned 10 mm next to previous track. This procedure is repeated to produce forty-two parallel tracks on a single sheet. Two sheets are used for each test set up, resulting in total sliding length of 58.8 m . For each test set-up, three repetitions are conducted. The sliding direction is oriented perpendicular to the rolling direction.

The Ni plated sheet used is standard production material with a thickness of 0.3 mm . The coating thickness is different for both sides ( 4 and 1μ m ). The tests are performed on the 4μ m coating side. Dimensional topographical measurements are performed with a nanofocus μ surf mobile confocal microscope before and after testing on the surface of the material (Fig.4). The height parameters of the material are listed in Table 1. Sample sheets are lubricated with Renoform MBW 677/6 25% which is applied using a pipet and distributed with a foam roller on the material, such that the lubricant amount is ~0.75 g/m2. The same kind of lubricant is used in industry, however the exact composition and percentage could differ.

[picture]

Fig. 4. Confocal measurements 600×600μ m a. initial tool surface. b. Tool surface after test c . initial sheet surface. d. sheet surface after test.

Fig. 4. Confocal measurements 600 × 600 μ m a. initial tool surface. b. Tool surface after test c . initial sheet surface. d. sheet surface after test.

Table 1. ISO 25178 height parameters initial tool and sheet surface.

Tool | 0.05 | 0.24 | 0.28 | -0.28 | 3.71
Sheet | 0.14 | 0.69 | 0.41 | 0.84 | 3.43

The slider tool material is 1.2379 (EN ISO 4957) with a polished surface roughness (Sa) of ~0.05μ m (Table 1) and a hardness of 62 HRC. Table 2 list the tool material elements. The tool diameter is 43 mm , with a total width of 8 mm and contact width 5 mm . For each new test a 'virgin' tool surface is used, by rotating and mounting the slider to a new position. Before performing the tests, the load cell is calibrated and the tool is aligned with pressure paper. After the tests, wear particles are collected from the tool for SEM analyses using suitable copper adhesive tape. Thereafter, the tool is ultrasonically cleaned and confocal measurements are performed.

Table 2. Tool material element in Wt%.

C | O | Si | Cr | Fe | S | V
4.29 | 3.16 | 0.43 | 13.05 | 77.37 | 0.39 | 1.31

Results and Discussion

The original version of this paper is available on https://www.scientific.net/MSF.1184.35.pdf

After each test asperity flattening is observed on the sheet and sometimes a ploughing track or scratch (Fig.4d) occur. After the test (sliding length 58.8 m ) the tool is polluted with particles and oil over the whole width of the tooling. Particles are collected by tape and investigated in the SEM. The composition maps (Fig.5) show a lot of O and S , which originate from the oil. A large amount of nickel particles is observed, which originate from the sheet coating. A very limited amount of chromium is measured (which can only come from the sliding tool) and thus only very limited amount of abrasive tool wear occurs. Iron particles are observed too which could originate from the tooling and/or the sheet. The size and appearance of the particles has a large range. From small particles (< 5μ m ) to relatively large particles up to 50μ m.

[picture]

Fig. 5. SEM/EDX analyses of particles collected from tool.

Fig. 5. SEM/EDX analyses of particles collected from tool.

After ultrasonic cleaning of the tool, confocal (Fig.4b) and SEM analyses (Fig.6) are performed on the tool surface. The confocal image (Fig.5b) shows build up. SEM Nickel mapping (Fig.6) measures over the whole contact width (5 mm) nickel, which is thus adhesive wear. On the tool surface mainly particles of Ni are attached (adhesive wear). The confocal of the tool shows 'scratches'. From particle analyses, only one Cr particles is observed (from many measurements) which indicate that there is almost none abrasive wear. The main part of the observed scratches is therefore thought to be only through the adhesive wear layer.

[picture]

Fig. 6. SEM/EDX analyse mapping elements on tool surface SOST slider.

Fig. 6. SEM/EDX analyse mapping elements on tool surface SOST slider.

The Coefficient of Friction (COF) shows some fluctuations (Fig. 7). In the middle of the sheet a slightly lower coefficient of friction is observed compared to the start and end of one track. This could be related to the amount of lubrication, however this is not yet confirmed. The repeatability between sheets is relatively well. One sheet out of 12 (test no. 2b) had a significantly higher COF, probably related to lubrication.

[picture]

Fig. 7. Colour plot Coefficient of friction. SOST test 30.7 bar.

Fig. 7. Colour plot Coefficient of friction. SOST test 30.7 bar.

Discussion

The original version of this paper is available on https://www.scientific.net/MSF.1184.35.pdf

In ironing severe deformations (thickness reduction), high contact pressures and an increase in temperature occur. These conditions are not represented in the laboratory tests, yet they significantly influence the friction and wear behaviour of the nickel coating. The deformation, pressure and rise of temperature will affect both the mechanical response and fracture behaviour of the Ni coating. Plastic deformation of the sheet increases the real area of contact, which in turn alters frictional behaviour. At this moment, the galling behaviour is investigated without this effects.

The lubricant amount was relatively low in the lab test (SOST) compared to industry, to accelerate tool wear. However, this could have resulted in another friction regime than what is typical for industry. Since the focus for this investigation is on tool wear, this disadvantage is accepted. In the future, an investigation could be performed to investigate the influence of lubricant amount on (tool) wear.

The tool in the laboratory investigation was softer than the tools used in industry (carbide), which could result in more abrasive wear for the laboratory tool. No significant amount of abrasive tool wear is observed in the laboratory set up. In both set-ups (industry and lab test) mainly adhesive wear is observed. Due to the fact that the sliding tool in the lab set-up is made of steel, the possibility of metal-metal bonds could also increase the likelihood of adhesion occurring. A carbide tool is perhaps less susceptible to this effect.

Different types of tests can be performed with the SOST. In this paper, long drawing lengths are performed and the tool wear is analysed. Another option is to gradually increase pressure until a scratch occurs. In this way the strength of the coating is tested and an indication is obtained for the sensitivity to galling. An advantage is the lower amount on sheet necessary to test, which could be relevant for new material developments.

Both tests series (lab + industry) were carried out using production material, however the Ni coating thickness differed between them. Differences in coating thickness may influence galling behaviour, as a thicker Ni layer has a higher amount of metallic Nickel present at the surface. The significance of this effect needs to be further investigated.

Conclusion

The original version of this paper is available on https://www.scientific.net/MSF.1184.35.pdf

In this paper, SEM/EDX analyses were performed to assess the tribological behaviour of nickel plated steel sheet with a focus on tool wear. The layer build up consists of compacted particles (often in the shape of flakes) and is composed mainly of nickel with some iron content. Slider on Sheet Tests (SOST) is found to be a relevant set up for studying tribological behaviour of Ni plated steels in battery production.

Acknowledgement

The original version of this paper is available on https://www.scientific.net/MSF.1184.35.pdf

The authors would like to thank Arne Neelen and Richard Stegeman for the assistance in measurements and Henri Kwakkel for sharing his knowledge on battery production.

References

The original version of this paper is available on https://www.scientific.net/MSF.1184.35.pdf

J.Chen et al., Preparation and tribological behavior of Ni-graphene composite coating under room temperature, Applied Surface Science 361 (2016) 49-56.

D.T. Gawne, U. Ma, Friction and wear of chromium and nickel coatings, Wear 129 (1) (1989) 123-142.

R.Sreekumar, A.S. Nair, S.S. Sreejakumari, Recent trends and developments in twodimensional materials based electrodeposited nickel nanocomposite coatings, FlatChem 36 (2022) 100434.

X.Wang et al. Methods and insights in optimizing nickel-based coatings for wear and corrosion resistance, Materials Today Communications 43 (2025) 111793.

E.van der Heide, A.J. Huis in t Veld, D.J. Schipper, The effect of lubricant selection on galling in a model wear test, Wear 251 (1-21) (2001) 973-979.

W.Wang, K.Wang, Y.Zhao,M.Hua, X.Wei, A study on galling initation in friction coupling stretch bending with advanced high strength hot-dip galvanized sheet, Wear 328-329 (2015) 286-294.

K.G. Budinski, S.T.Budinski, Interpretation of galling tests, Wear 323-333 (2015) 1185-1192.

J.Venema, F.Korver, T.Chezan. Slider on Sheet Tester Development for Characterizing Galling, Key Engineering Materials 926 (1-3) (2022) 1204-1210.

This ePub file was created using machine-generated methods and automated software to convert the source text into an ePub format and may contain formatting inconsistencies from the original version. Formatting may vary depending on the device used.

The original version of this eBook is available on https://www.scientific.net/book/friction-and-wear-in-metal-forming/978-3-0364-1992-3

Materials Science Forum, ISSN: 1662-9752, Vol. 1184, pp 43-52

doi: 10.4028/p-06WuSM

© 2026 The Author(s). Published by Trans Tech Publications Ltd, Switzerland.

Submitted:
2026-01-19

Revised:
2026-02-17

Accepted:
2026-02-21

Online:
2026-04-14

Numerical Analysis of Roughness Transfer Mechanism during Skin-Pass Rolling

The original version of this paper is available on https://www.scientific.net/MSF.1184.43.pdf

Hamed Aghajani Derazkola 1,a*, Leon Jacobs 2,b, Stamatis Kiakidis 2,c, Ton van den Boogaard 1, d, Javad Hazrati 1,e

¹Nonlinear Solid Mechanics, Faculty of Engineering Technology, University of Twente, Enschede, The Netherlands

2 Tata Steel, Research & Development, 1951 JZ Velsen-Noord, The Netherlands

a*h.aghajaniderazkola@utwente.nl,bleon.jacobs@tatasteeleurope.com, c  Stamatis.Kiakidis@tatasteeleurope.com, da.h.vandenboogaard@utwente.nl, j.hazratimarangalou@utwente.nl

*corresponding author

Keywords: skin-pass rolling, roughness transfer, finite element model.

Abstract

Skin-pass rolling is a finishing process characterized by very small thickness reductions, primarily applied to control surface texture and mechanical properties of rolled steel strips. A key outcome of this process is the transfer of surface roughness from the work roll to the strip, which is governed by local contact conditions at the roll-strip interface. In this study, roughness transfer during skin-pass rolling of DX56 steel sheets is investigated using a combined macro-micro finite element modeling approach supported by pilot-mill experiments. Rolling trials were conducted to measure thickness reduction and resulting surface roughness under different rolling forces and entry tensions. A macro-scale rolling model was first employed to estimate effective friction coefficients by reproducing the experimentally observed thickness reductions for each rolling condition. These calibrated friction coefficients were subsequently applied in a micro-scale finite element model incorporating an electro-discharge textured (EDT) roll surface to analyze local contact pressure, plastic strain accumulation, and roughness transfer mechanisms. The results show that increasing rolling force leads to higher contact pressures, longer roll bites, and enhanced asperity-scale plastic deformation, resulting in increased roughness transfer. Entry tension modifies the stress distribution within the roll bite, which facilitates localized yielding without inducing plastic deformation prior to roll entry. The simulations capture the qualitative trends of surface roughness evolution observed experimentally, demonstrating the capability of the proposed finite element framework to analyze roughness transfer mechanisms in skin-pass rolling.

Introduction

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Surface texture of cold rolled steel strips determines its visual appearance, paint ability and tribological behavior [1]. The texture is being printed on the strip surface by rolls with preferred surface roughness, in the skin-pass which is the last sequence of the steel sheet rolling. Therefore, a thorough mechanistic understanding of skin-pass rolling with the elastic-plastic deformation analysis at the asperity level is necessary to design and control the product surface texture using different process parameters [2-4]. Despite extensive research on conventional rolling, the process of roughness transfer from roll to strip has not yet been thoroughly addressed. In skin-pass rolling, the reduction in strip thickness is typically less than 3%. The primary deformation occurs at the strip surface [5]. Simulating roughness transfer in skin-pass rolling presents significant challenges due to the need for accurate material modeling, contact conditions, and precise handling of micro-scale roughness transfer with localized micro level fine meshes. Hence, there are a couple of studies in the literature focusing on analysis of skin-pass rolling. Kijima and Bay [6] discussed the difficulties of experiments and modeling of skin-pass and simplified the problem by adopting a plane strain model to study the effect of friction in skin-pass rolling. They show that skin-pass rolling involves an extended sticking region producing highly inhomogeneous deformation, and that plane-strain
upsetting provides an accurate surrogate for studying these mechanisms. Experiments confirmed FEM predictions of sticking/sliding boundaries, validating the modelling approach for high-friction, small-reduction rolling. In their next work [7], they applied only normal loads with no tangential sliding in a FE model to make texture on the strip surface and compared with skin-pass rolling experiments. They reported that both tool roughness and pitch strongly affect roughness transfer, with smaller roughness and larger pitch producing higher transfer. Kijima and Bay [8] applied tangential shear after normal load (in two steps) in their next study to make the analysis more realistic and closer to the skin-pass rolling. The study concludes that tangential displacement (both in sticking and sliding regions) significantly increases roughness transfer, meaning that pure normal-loading models underestimate surface deformation in practical dry skin-pass rolling. The work verifies FE predictions with experiments and shows a pronounced roughness peak at the sliding-sticking boundary. Bünten et al. [9] concludes that the developed FE model (validated through detailed upsetting tests and applied to temper-rolling conditions) can accurately predict the microscopic transfer of deterministic roll-surface structures, reproducing both penetration and reverse-extrusion mechanisms observed experimentally. It further shows that, for the geometries and elongations studied, the frictional coefficient does not need to be prescribed explicitly because the microscopic surface topography alone governs the effective frictional response in the simulations.

This work presents macro- (without roll roughness) as well as micro-scale (with roll roughness) analyses of skin-pass rolling by a FE model. The effects of process parameters such as rolling force and strip tension on the contact pressure and strip velocity at the roll bite are presented. Afterwards, the macro-scale model is adjusted to include simplified micro-geometry of EDT asperities on the roll. The model is used to investigate the effect of rolling force and entry tension on the transferred roughness on the strip.

Experimental Procedure

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The rolling experiments and sample preparation were conducted using the Multi mill (pilot test) at Tata Steel, Netherlands. The material used in this study was DX56 steel sheet with initial thickness of 3 mm which is cold rolled to 0.6 mm before skin-pass rolling. In these experiments, the DX56 steel strip was intentionally not annealed and therefore remained in a full-hardened condition before skin pass rolling. This choice was made deliberately, as the material behavior of full-hardened steel is more straightforward to characterize than that of annealed material. Prior to the skin-pass rolling stage, the material properties were measured by uniaxial tensile test. A series of trials were conducted at Tata Steel, and the effects of rolling force and entry/exit tension on the strip thickness reduction were evaluated (Table 1). In these experiments, a 195 mm roll radius was selected and rolling speed of 2 m/s was used. These results are used for the validation of the FEM simulation.

Table 1. Selected process parameters (experiments) for evaluation of FEM model.

Sample | Strip thickness | Entry tension | Exit tension | Rolling force | Thickness reduction
1 | 0.6mm | 200MPa | 55MPa | 300kN | 0.291%
2 | 0.6mm | 200MPa | 55MPa | 900kN | 1.5%
3 | 0.6mm | 300MPa | 55MPa | 900kN | 1.75%

The surfaces of the skin-pass rolled sheets (Samples 1, 2, and 3) were evaluated using a confocal microscope (Sensofar) to characterize their surface roughness properties. Measurements were performed at both the edges and the center of each sheet.

Finite Element Model

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The material properties were measured experimentally and used for simulation. Bergstrøm-van Liempt (BvL) hardening model (stress-strain curve) was used to estimate flow stress [10]. Before reaching the skin-pass rolling stage, the DX56 steel sheets undergo several preliminary deformation steps, which introduce pre-strains that affect the material behavior during skin-pass rolling. The vonMises yield criterion is used for this steel during the simulation.

In this research, ABAQUS 2023 software is used to develop the plane strain finite element model of the skin pass rolling. The model used in this study is force-controlled, meaning that the rolling force is applied to the sheet through an elastic work roll. A midplane symmetry boundary condition is also imposed at the bottom of the strip. During simulation, for the vertical interaction, a hard contact model with the augmented Lagrange method was applied. For the tangential interaction, the penalty method was used. The rolling forces and the entry tension (ET) range were selected based on the experiments. In this simulation, the roll force and the entry and exit tensions were defined as input parameters, while the thickness reduction (measured experimentally) was the output variable. Using this approach, the coefficient of friction (COF) at the interface was calibrated to obtain the value that reproduces the experimentally observed thickness. To simulate the roughness transfer, an electrodischarge texturing (EDT) single event geometry was taken from reference [11] and modeled on the roll for simulation. Image of indent dimension is presented in Figure 1a. To reduce computation time, only seven indents were modeled on the work roll while still capturing the relevant interface behavior. To mesh the model, various mesh ratios and sectioning approaches were tested on both the roll and the strip to find the optimal mesh for convergence. The top surface region of the sheet in direct contact with the EDT indentation was discretized using a locally refined mesh. Adaptive remeshing was implemented within Abaqus/Standard during the cold rolling simulation to control element distortion and ensure stable convergence of the implicit solution. To accurately describe the contact and deformation of the roll, 7 rows of elements, each with double the mesh size of the strip were placed along the outer circumference of the roll. The element type used was CPE4 (plane strain). Figure 1b shows the mesh domain in this study.

[picture]

Fig. 1. (a) Single EDT geometry [10], and (b) mesh domain of skin pass rolling.

Fig. 1. (a) Single EDT geometry [10], and (b) mesh domain of skin pass rolling.

Macro-scale model

The original version of this paper is available on https://www.scientific.net/MSF.1184.43.pdf

Macro-scale model of rolling process (excluding the roll texture) is used to reproduce the experimental results by calibrating the coefficient of friction (COF). The calibrated COFs later are used in micro-scale model to simulate roughness transfer in skin pass rolling. Figures 2a and 2b show contact pressure and surface nodal displacement of sheet at roll bite, respectively. The latter is a good approximation since it has been shown that ploughing effects of the texture on friction at the roll bite can be neglected [12]. For sample 1, an effective COF of 0.073 reproduces the experimentally observed thickness reduction. Under these conditions, the maximum contact pressure reaches 0.72 GPa and the roll bite length is approximately 3.6 mm . Increasing the rolling force from 300 kN (Sample 1) to 900 kN (sample 2) leads to a pronounced increase in contact pressure and a significant extension of the roll bite. For sample 2, an effective COF of 0.139 is required, leading to a peak contact pressure of 1.35 GPa and a roll bite length of 6.0 mm .

In samples 2 and 3, the rolling force was kept constant at 900 kN , while the entry tension was increased from 200 MPa to 300 MPa . The simulations show that increasing the entry tension slightly
extends the roll bite length to 6.2 mm while reducing the peak contact pressure to 1.32 GPa . The calibrated COF for sample 3 is 0.135 , which is slightly lower than that obtained for sample 2(0.139). The calibrated COF values should be interpreted as effective friction coefficients reflecting the pressure-dependent tribological response of the roll-strip interface. At higher rolling forces, increased contact pressure promotes work roll flattening, increases the real area of contact, and leading to higher frictional resistance. Consequently, an increase in the effective COF with rolling force is consistent with the known behavior of skin-pass rolling under high contact pressures.

[picture]

Fig. 2. (a) Contact pressure distribution and (b) surface nodal displacement (UT2) along the roll bite for samples 1(300kN−200MPa),2(900kN−200MPa) and 3(900kN−300MPa) obtained from macroscale finite element simulations. The spatial coordinate x is defined along the rolling direction, with x=0 at the roll centerline; negative values indicate the entry side and positive values the exit side of the roll bite. Arrows indicate the roll bite length for each condition.

Fig. 2. (a) Contact pressure distribution and (b) surface nodal displacement (UT2) along the roll bite for samples 1 ( 300 k N − 200 M P a ) , 2 ( 900 k N − 200 M P a ) and 3 ( 900 k N − 300 M P a ) obtained from macroscale finite element simulations. The spatial coordinate x is defined along the rolling direction, with x = 0 at the roll centerline; negative values indicate the entry side and positive values the exit side of the roll bite. Arrows indicate the roll bite length for each condition.

Entry tension was found to have a limited influence on the magnitude of the contact pressure but affected its distribution along the roll bite. Higher entry tension slightly reduced the peak contact pressure, shifted it toward the exit side, and extended the roll bite length [13]. This behavior is attributed to the elastic tensile stress introduced prior to roll entry, which alters the stress state of the strip and modifies the equilibrium between normal pressure and frictional shear within the roll bite, thereby shifting the neutral point downstream [14]. Although the peak contact pressure decreases, plastic deformation is not inhibited; instead, yielding is facilitated under the combined action of normal pressure and tensile stress, leading to a redistribution of deformation along the roll bite. Table 2 summarizes the simulation results and compares them with the experimental data. The experimental thickness reductions for Samples 1, 2, and 3 are 0.291%,1.50%, and 1.75%, respectively. Based on the surface nodal displacement (UT2) obtained from the simulations, the corresponding thickness reductions for Samples 1, 2, and 3 are 0.288%,1.59%, and 1.82%, respectively.

Table 2. FEM model results and comparison with experiments.

Sample | Experimental thickness reduction | Simulation thickness reduction | Thickness reduction error | Roll bite length | Maximum contact pressure | COF
1 | 0.291% | 0.288% | 1.1% | 3.6mm | 0.719GPa | 0.073
2 | 1.5% | 1.59% | 5.7% | 6.0 mm | 1.35GPa | 0.139
3 | 1.75% | 1.82% | 3.9% | 6.2mm | 1.32GPa | 0.135

Micro-scale model

The original version of this paper is available on https://www.scientific.net/MSF.1184.43.pdf

The FEM results revealed distinct trends in contact pressure at both the micro and macro scales during skin-pass rolling, particularly regarding the shape and magnitude of the pressure distribution. Figure 3a shows a representative contact-pressure distribution from the micro-scale model. As outlined in the modelling section, only seven indents were included on the work roll to reduce computational cost. As expected, the contact pressure in the indented zone (representing the EDT events) is higher than that in the smooth regions of the roll surface. A magnified view of the pressure field around the seven indents is provided in Figure 3b. The contact pressure exhibits a periodic pattern, rising from zero to a local maximum at each indent. The zero-pressure regions correspond to locations where the roll and strip are not in contact. The first (leading) and last (trailing) indents show lower peak pressures than the central indents, which is a consequence of the roll geometry definition in the model and should be disregarded.

[picture]

Fig. 3. (a) Contact pressure of Sample 3(900kN−300MPa) at micro-scale, (b) high magnification contact pressure at the textured zone.

Fig. 3. (a) Contact pressure of Sample 3 ( 900 k N − 300 M P a ) at micro-scale, (b) high magnification contact pressure at the textured zone.

Figure 4a presents the FE simulation results of the contact pressure for the micro-scale model. As mentioned before, in the macro-scale model, the deformed roll maintained continuous contact with the strip along the roll bite. In contrast, for the textured roll at the micro scale, the contact area between the roll and the strip within the roll bite was discontinuous. As shown in Figure 4a, the rims of the textures were in direct contact with the strip, while the strip material did not directly contact the craters. The degree of contact between the roll and the strip depends on the process parameters such as rolling force and strip tension [15]. The effects of rolling force and entry tension on the contact pressure and normal stress for a single EDT event are illustrated in Figure 4b. In this section, the length is expressed in micrometers, representing the contact length between one EDT event and the strip. At the macro scale, the maximum contact pressures at the roll bite for Samples 1, 2, and 3 were 0.719GPa,1.35GPa, and 1.32 GPa , respectively. In contrast, at the micro scale, the limited contact area at the indent-strip interface led to a significant increase in the local contact pressure. The FEM results show that the maximum contact pressures at the rim-strip contact area for Samples 1, 2, and 3 were 1.51GPa,2.06GPa, and 1.99 GPa , respectively.

[picture]

Fig. 4. FEM results of contact pressure at micro-model (Sample 3 ( 900kN−300MPa ) and (b) contact pressure of single indent for Sample 1(300kN−200MPa), Sample 2(900kN−200MPa) and Sample 3(900kN−300MPa).

Fig. 4. FEM results of contact pressure at micro-model (Sample 3 ( 900 k N − 300 M P a ) and (b) contact pressure of single indent for Sample 1 ( 300 k N − 200 M P a ) , Sample 2 ( 900 k N − 200 M P a ) and Sample 3 ( 900 k N − 300 M P a ) .

For a more comprehensive understanding, Figures 5 a and 5 b present the complete distribution of contact pressure and normal stress (S22) for Sample 2 at the roll bite. As a result of the periodic repetition of EDT textures on the work roll surface, both the full-contact and non-contact regions are periodically reproduced along the contact length. Meanwhile, the maximum contact pressure and normal stress occurring at the rims of the textures remain nearly constant.

[picture]

Fig. 5. (a) FEM results of contact pressure and (b) normal stress of micro-scale at Sample 2 (900kN−200MPa).

Fig. 5. (a) FEM results of contact pressure and (b) normal stress of micro-scale at Sample 2 ( 900 k N − 200 M P a ) .

Figure 6a illustrates the distribution of equivalent plastic strain (PEEQ) for Sample 2 as a representative case. The PEEQ is highly localized near the strip surface and exhibits pronounced maxima at the leading and trailing edges of the surface indents. This localization reflects the nonuniform contact conditions at the roll-strip interface, where micro-scale variations in contact pressure and friction govern the local plastic flow; indentation and extrusion of material. Figure 6b presents the evolution of PEEQ along the rolling direction for all three samples under different rolling force and entry tension conditions. The PEEQ distribution closely follows the corresponding micro-scale
contact pressure profile, confirming that local pressure maxima dominate strain accumulation during skin-pass rolling. The deformation pattern further confirms that ploughing of the indents on the strip surface is negligible for the investigated rolling conditions. An increase in rolling force leads to a higher mean contact pressure, which amplifies local pressure peaks at the indent edges and consequently increases the PEEQ. The maximum PEEQ values for Samples 1, 2, and 3 were 0.169, 0.28 , and 0.375 , respectively.

A systematic difference between the leading and trailing indents is also observed, with higher PEEQ values occurring at the trailing edge (figure 6b). This asymmetry arises from material pile-up within the crater region during roll passage. The pile-up (extrusion in the crater) increases the effective contact length and enhances adhesive friction drag, resulting in a higher accumulated plastic strain at the trailing indent compared to the leading one. Entry tension exerts a significant influence on the PEEQ distribution by modifying the local stress state at the roll-strip interface towards pure shear and therefore reducing the yield stress of the material.

[picture]

Fig. 6. (a) FEM results of EDT textures PEEQ for Sample 2(900kN−200MPa) and (b) comparison of PEEQ for all Samples.

Fig. 6. (a) FEM results of EDT textures PEEQ for Sample 2 ( 900 k N − 200 M P a ) and (b) comparison of PEEQ for all Samples.

Figure 7 presents the finite element results of strip surface deformation after rolling with textured work rolls for sample 2 . The corresponding surface profiles of samples 1,2 , and 3 , expressed in terms of surface nodal displacement ( μ m), are shown in Figure 7, highlighting the differences in surface roughness morphology under varying rolling force and entry tension conditions. The results demonstrate a pronounced influence of rolling force on the shape of the crater region. At higher rolling forces, the crater exhibits a more pronounced frustum-like profile, indicating enhanced filling of the roll-induced surface features. This behavior is associated with increased contact pressure and frictional traction at the roll-strip interface, which promote localized plastic flow and material displacement into the crater. In contrast, at lower rolling forces, the lower contact pressure limits plastic flow, resulting in a flatter and less developed crater morphology.

[picture]

Fig. 7. Comparison of surface nodal displacement (deformation) between experiments and simulation.

Fig. 7. Comparison of surface nodal displacement (deformation) between experiments and simulation.

For evaluation of FEM results the arithmetical mean surface roughness (Ra) of deformed surface is calculated by:

Ra=1 N∑i=1N|Zi−Z―|(1)

where Zi denotes the measured surface height at i -th measured point of the surface profile, Z― is the mean height over the evaluation length, and N is the total number of measurement points used to compute the arithmetical mean roughness. The combined effects of rolling force and entry tension on the micro-scale Ra of the strip are shown in Figure 8. The calculated Ra values exhibit a trend consistent with that of thickness reduction. The average simulated Ravalues for Samples 1, 2, and 3 were 0.39μ m,0.94μ m, and 1.15μ m, respectively. Correspondingly, the experimental results obtained from the top surface of the DX56 sheet produced under the same rolling conditions showed Ra values of 0.91μ m,1.80μ m, and 2.04μ m for Samples 1, 2, and 3, respectively. It is important to note that the Ra computed from the FEM results corresponds to a simplified, homogeneous, and periodic EDT texture represented under plane-strain conditions, implying infinitely wide asperities in the transverse direction. In contrast, the work-roll roughness in the pilot mill is stochastic and three-dimensional, which explains the quantitative discrepancy between the simulated and experimental Ra values. Nevertheless, the FEM results correctly capture the qualitative trend of increasing Ra with increasing rolling force and entry tension.

[picture]

Fig. 8. Comparison of Ra between experiments and simulation.

Fig. 8. Comparison of Ra between experiments and simulation.

Summary

The original version of this paper is available on https://www.scientific.net/MSF.1184.43.pdf

This study investigated the influence of rolling force and entry tension on roughness transfer during skin-pass rolling using a combined macro-micro finite element framework supported by pilotmill experiments. A macro-scale rolling model was first employed to reproduce experimentally observed thickness reductions by calibrating friction coefficients for each rolling condition. These calibrated friction values were subsequently applied in micro-scale simulations incorporating an EDT-type surface geometry to analyze local contact conditions and asperity-scale deformation.

The numerical results show that increasing rolling force leads to higher contact pressures, longer roll bites, and larger plastic deformation at the asperity scale, resulting in enhanced roughness transfer. Entry tension was found to modify the stress state within the roll bite by superimposing a tensile stress on the strip, shifting the neutral point and promoting localized plastic deformation. Although rolling force, entry tension, and friction are inherently coupled in the experimental conditions considered, the simulations capture their combined influence on roll bite geometry, contact conditions, and surface deformation behavior.

The micro-scale model successfully reproduces key qualitative trends observed experimentally. Quantitative differences between simulated and measured Ra values are attributed to the simplified and periodic representation of roll roughness in the numerical model compared to the complex stochastic texture of the industrial work rolls. Future work will focus on incorporating measured rollsurface topographies, lubrication effects, and more advanced constitutive description of strip material to further improve the predictive capability of the model for industrial skin-pass rolling applications.

References

The original version of this paper is available on https://www.scientific.net/MSF.1184.43.pdf

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L.J.M. Jacobs, E.H. Atzema, J. Moerman, M.B. de Rooij, 'Quantification of the influence of anisotropic plastic yielding on cold rolling force', Journal of Materials Processing Technology, vol. 319, 2023, 118055,

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L.J.M. Jacobs, "MODELLING FRICTION IN COLD ROLLING", Ph.D. thesis University of Twente, Enschede. 2025.

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Materials Science Forum, ISSN: 1662-9752, Vol. 1184, pp 53-64

doi: 10.4028/p-7lTOao

© 2026 The Author(s). Published by Trans Tech Publications Ltd, Switzerland.

Submitted:
2026-01-16

Revised:
2026-02-16

Accepted:
2026-02-21

Online:
2026-04-14

Modeling the Effect of Lubricants on Surface Conditions in Plane-Strain Upsetting Tests

The original version of this paper is available on https://www.scientific.net/MSF.1184.53.pdf

Lena Koch 1a*, Holger Brüggemann 1 b, Shakthi Bharani Tamilselvan 1c, Emad Scharifi 1d and Junhe Lian 1e

1 Institute of Metal Forming (IBF), RWTH Aachen University, Germany

alena.koch@ibf.rwth-aachen.de, b  holger.bruegemann@ibf.rwth-aachen.de, c  shakthi.tamilselvan@ibf.rwth-aachen.de, demad.scharifi@ibf.rwth-aachen.de, ejunhe.lian@ibf.rwth-aachen.de

*corresponding author

Keywords: skin-pass rolling, aluminum, roughness, FEM, coupled Eulerian-Lagrangian.

Abstract

Skin-pass rolling is commonly used to adjust the surface quality of high-strength aluminum alloys. Lubrication plays an important role in this process, as it minimizes material adhesion to the work roll, extends its service life, and also influences the contact conditions and the final surface topography. However, most numerical studies represent lubrication only through an effective friction coefficient. In this work, a finite-element framework that explicitly accounts for lubricant entrapment in engineered surface pockets by using a coupled Eulerian-Lagrangian (CEL) approach is introduced to investigate lubricant-topography interaction. The skin-pass rolling process is approximated by a plane-strain upsetting test to represent the parameters relevant for mapping the interaction between lubricant and mechanical stress, as the rolling process has a large number of influencing factors. The precipitation-hardenable aluminum alloy EN AW-6016 is modeled with rate-dependent plasticity based on experimental flow curves, while the lubricant is represented as a Eulerian material governed by an equation-of-state formulation. The effects of strain rate, friction and different lubricant filling levels in surface pockets are analyzed. The results show that variations in friction mainly affect the global force level, while the presence of lubricant leads to changes in local deformation and stress distributions. Fully filled pockets require higher forming forces due to lubricant compression, whereas partially filled pockets show behavior close to dry conditions. The CEL approach proves suitable for modeling lubricated plane-strain upsetting tests and provides a basis for further investigations of lubricated skin-pass rolling processes.

Introduction

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The growing demand for lightweight design, combined with the requirement to manufacture structural components that fulfill strength specifications, has driven the development of the skin-pass rolling process for high-strength aluminum alloys in order to achieve a defined surface quality [1]. A major challenge in production is ensuring the surface quality required for subsequent coating operations [2]. To this end, a controlled surface topography is imparted to the strip by means of textured work rolls. The process is carried out under lubricated conditions to minimize material adhesion to the work roll and to extend its service life [3]. The lubricant actively participates in the forming process and affects the strip surface through hydrostatic and hydrodynamic mechanisms arising from the entrapped fluid. To develop a comprehensive understanding of the process and the underlying mechanisms, a key goal is to enable numerical simulation of lubricated skin-pass rolling. As the number of influencing variables in rolling (e.g., relative speed, temperature, friction) is substantial, a simplified analogue test is required to represent the key parameters governing the interaction between lubricant behavior and mechanical loading. A commonly used analogue test for rolling is the plane strain upsetting test, in which a sheet is compressed between two dies. This concept is proposed in the work of Kijima and Bay [4,5]. Owing to the significantly larger sheet width relative to the contact length and the sheet thickness, as well as the large roll diameter relative to the contact length, the resulting stress state approximates plane strain compression as in rolling [6]. In general,
the lubricated imprinting process consists of two primary aspects: the transfer of roughness and surface topography, and the hydrostatic and hydrodynamic behavior of the lubricant. The final surface profile emerges from the combined effect of these contributions. The reproducibility of surface roughness from the rolling process using the plane strain upsetting test has already been investigated in a previous study [7]. This work investigated whether a simplified plane-strain upsetting finiteelement model can reliably approximate surface roughness transfer during skin-pass rolling. By numerically modeling die surface roughness with circular arcs and varying friction coefficients and surface pitch, the study analyzed roughness transfer mechanisms and compared the results with a macro-meso skin-pass rolling model. The work demonstrates that roughness transfer is mainly governed by die surface geometry and relative motion, while the simplified model offers much lower computational effort with comparable trends [7]. In the previous work, the lubricant was not explicitly modeled but only taken into account using the friction coefficient, whereas in the present work, fluid-structure interaction (FSI) modeling is conducted within the plane strain upsetting test on precipitation-hardenable aluminum alloy EN AW-6016. The lubricant is explicitly modeled using a coupled Euler-Lagrangian meshing approach in order to capture its influence on roughness transfer based on a simplified surface profile. It is hypothesized that conventional friction models commonly applied to lubricated forming processes, such as Coulomb or Tresca formulations, are insufficient in this context, as they fail to capture the complex interfacial mechanisms arising from the interaction between the lubricant film and the surface topography. The results are evaluated with respect to surface profiles as well as the resulting stress and strain distributions and are subsequently compared to those obtained from conventional modeling approaches.

Material and Method Modeling Framework

The original version of this paper is available on https://www.scientific.net/MSF.1184.53.pdf

Coupled Euler-Lagrangian method: Simulating fluid-structure interactions is challenging, particularly when high plastic strains are present. Conventional numerical techniques frequently encounter limitations under such conditions, for example due to excessive mesh distortion or difficulties in tracking interfaces [8]. The Coupled Eulerian-Lagrangian (CEL) method provides an effective framework to overcome these issues and provide a suitable tool to model FSI. The strength of this modeling strategy is that the fluid is modeled with a Eulerian meshing approach, in which the mesh remains fixed while materials flow through it. The workpiece itself can be modeled in an Lagrangian manner, in which the mesh deforms with the material and which is mainly used in continuum mechanics applications [9]. However, large deformations will lead to severe distortion of the mesh and accordingly distortion of the individual elements. This results in numerical inaccuracies and precludes successful simulation. In a Eulerian approach, the nodes of the mesh remain fixed in space, while the material flows through elements that do not deform. Here, it is not the movement of the individual nodes that is considered, but rather the mass, momentum and energy flow through the mesh elements [10]. One disadvantage of this method is that it is difficult to track material points over time and complex geometries are difficult to represent. On the other hand, severe deformations can be represented well. This means that Eulerian elements are not always completely filled with material, some may contain only part of the material or may even be empty [11]. As a result, the boundary of the material, which usually does not align with the element boundaries, must be updated at every time step. To give the material enough freedom to move, the Eulerian mesh is usually designed as a simple rectangular grid that extends well beyond the actual material region. If the material flows outside this mesh, it is considered lost in the simulation. By combining these two representations, CEL enables the robust simulation of highly dynamic events such as fluid splashing, detonation phenomena, and object-fluid impact, without the numerical complications typically associated with mesh distortion or intricate interface-tracking procedures. An interface in Abaqus Vexplicit allows information from these two approaches to be exchanged at the individual elements.

[picture]

Fig. 1. Schematic representation of the Coupled Euler Lagrangian approach

Fig. 1. Schematic representation of the Coupled Euler Lagrangian approach

The fluid dynamics are described by the Navier-Stokes equations, assuming that volume viscosity is negligible (see the corresponding (1)) [12]. Here η is the viscosity, ρ is the density, p is the pressure, v is the velocity and b represents body forces.

ρDvDt=−∇p+η∇2v+η3∇(∇·v)+ρb(1)

The relationship between density and pressure is defined through a so-called Equation of State (EOS). The pressure in the lubricant is determined using a linear Us-Up Hugoniot formulation of the MieGrüneisen EOS [13]. With this approach, the lubricant is treated as a compressible viscous fluid. This results in the following equation.

p=ρ0c02η(1−sη)2(1−Γ0η2)+Γ0ρ0Em(2)

Here, is p the pressure, ρ0 the reference density, Em is the specific energy, Γ0 is the Grüneisen ratio. Furthermore c0 and s are defined by a linear relation between the shock velocity Us and the particle velocity Up [13].

Us=c0+sUp(3)

Simulation model: A simulation model of a plane-strain upsetting test was set up in Abaqus2019 Vexplicit. During the process, the die moves upward and deforms the workpiece. It should be noted that, although plane strain upsetting tests show similar stress states and contact conditions, important skin-pass rolling parameters are not considered. These include the curvature of the work roll, the relative movement and velocity of the sheet, and fluid entrainment. The setup is shown in Fig. 2. To reduce computational time only half of the test was modeled with suitable symmetry conditions. The die is modeled as a rigid body and discretized using R3D4 elements with an element size of 1 mm . The die has a height of 5 mm , a width of 150 mm , and a depth of 10 mm . On the top of the die are three pockets in the shape of a truncated cone in a line. These have a diameter of 6 mm and a depth of 2 mm . In addition, the pockets have a center distance of 6.5 mm between each other. The pocket and resulting surface are significantly larger than those produced by skin-pass rolling, which is carried out using electric discharge texturing (EDT) on the roller. The reason for this is that the test setup should be used for validation of a numerical FSI model. In this context, it is more appropriate to start with a large pocket, especially since it is possible to produce such a pocket without EDT technology. In its initial state, the deformable workpiece has a thickness of 1.2 mm , a width of 20 mm , and a depth of 40 mm and is discretized using C3D8R elements, which are three-dimensional eight-node-brick elements with reduced integration, with an element edge length of 0.2 mm . dimensional eight-node-brick elements with reduced integration, with an element edge length of 0.2 mm . The Euler domain has the dimensions of length 20 mm , depth of 10 mm and height of 3.2 mm . It is discretized using EC3D8R elements with an approximate global size of 0.1 mm . The
lubricant is modelled with dimensions identical to the geometry of the pocket, such that the lubricant occupies the pocket without any free surface or gaps. This lubricant is assigned to the Euler domain using discrete volume fraction field. Depending on the preferred experimental requirements, the dimension of the lubricant is changed by adjusting the initial height of the lubricant relative to the depth of the pocket, such that 75% filling height corresponds to 1.5 mm filling height of the lubricant with respect to 2 mm pocket depth. The die was assigned a velocity of 30 mm/s for a total stroke of 0.12 mm . This results in a height reduction of 10%. In the dry state a surface-to-surface contact with a coulomb friction coefficient of 0.01 to 0.28 is defined between die and workpiece.

[picture]

Fig. 2. a) Schematic structure of the simulation model with the Euler domain and lubricant in the initial position, here with a fill level of 50%, b ) schematic structure of the die with the three pockets
Flow curves from stacked layer compression tests at room temperature and three different forming speeds were recorded (see Fig. 3). A pronounced strain rate dependency is observed, which was implemented in the simulation. The other material parameters such as density, Young's modulus, and Poisson's ratio of the workpiece are shown in Table 1.

Fig. 2. a) Schematic structure of the simulation model with the Euler domain and lubricant in the initial position, here with a fill level of 50 % , b ) schematic structure of the die with the three pockets Flow curves from stacked layer compression tests at room temperature and three different forming speeds were recorded (see Fig. 3). A pronounced strain rate dependency is observed, which was implemented in the simulation. The other material parameters such as density, Young's modulus, and Poisson's ratio of the workpiece are shown in Table 1.

Table 1. Material properties of the workpiece

Material | Density [kg/m³] | Young's Modulus [GPa] | Poisson's ratio
EN-AW6016 | 2700 | 70 | 0.33

[picture]

Fig. 3. Flow curves for EN AW-6016 aluminum alloy at 20∘C

Fig. 3. Flow curves for EN AW-6016 aluminum alloy at 20 ∘ C

The density, viscosity, sound velocity and shear modulus are required for the numerical representation of the lubricant. These values are shown in Table 2. The linear Us-Up Hugoniot form of the equation of state is used to represent the fluid properties. The contact conditions and the resulting friction behavior in the lubricated simulation with CEL are different from those in the dry state. In Abaqus, a distinction is made between the methods "General Contact" and "Surface-to-Surface Contact." "General contact" describes the hydrodynamic friction of the fluid with the workpiece with a friction coefficient of μfluid =0.01. The global "general contact" is selected here because no effective surfaces can be defined for a liquid. It ensures that the contact formulation is applied only when the Euler material with volume fraction greater than zero comes into contact with the Lagrangian surface. In addition, the "surface-to-surface contact" method with a friction coefficient of μ=0.1 is used, which locally overwrites the definition of "general contact" [14]. This contact is formulated exclusively between the rigid die and the workpiece to model the solid to solid interaction. This is restricted to Lagrangian surfaces and was not applied to Euler surface. Therefore, the force transmission can be either through the lubricant via general contact in lubricant's presence or through direct surface interactions in regions where the lubricant is not present.

Table 2. Fluid properties according to data sheet and [15,16]

Lubricant
Density mathematical formula

Speed of Sound

Dynamic Viscosity

Shear modulus

Rolling oil | 806.78 | 1333.14 | 0.00547 | 1.2

In this simulation study, the necessity of strain rate dependency is first examined, followed by an investigation of friction sensitivity. In the next step, the usefulness of the CEL method for calculating lubricated tests is demonstrated by comparing it with unlubricated tests. Finally, different filling heights of the pockets are analyzed. Consequently, this work consists of four different investigations (see Table 3.). The investigation of the effect of the coefficient of friction μ without explicit consideration of the lubricant using CEL serves to analyze the general sensitivity of the test setup to the friction conditions. Without such a study, a potential effect of the lubricant cannot be considered in isolation.

Table 4. Overview of the individual investigations, including the respective simulation boundary conditions

Number | Meshing approach | Friction
Investigation of strain rate dependence
1 | With strain rate-dependent material data | Lagrangian | μ = 0.1
Without strain rate-dependent material data | Lagrangian | μ = 0.1
Friction sensitivity
2 | Conventional friction-based simulations | Lagrangian | μ = 0.01, 0.1 and 0.28
Comparison of lubricated and unlubricated simulation
3 | Unlubricated | Lagrangian | μ = 0.1
Lubricated | CEL | "surface to surface contact" method with a friction coefficient of μ = 0.1
Investigation at different filling levels
4 | 75 % filling level | CEL | "surface to surface contact" method with a friction coefficient of μ = 0.1
95 % filling level
100 % filling level

Results

The original version of this paper is available on https://www.scientific.net/MSF.1184.53.pdf

Investigation of strain rate dependence: First, the force-displacement curves from a simulation with strain rate dependence and a simulation without strain rate dependence were compared in the unlubricated state as shown in Fig.4. In both simulations, the Coulomb friction coefficient is μ=0.1. It can be observed that in the case of strain rate dependency a 2 kN higher force is present. This is due to the fact that, as can be seen in Fig. 3, the stress increases at higher strain rate. As a result, all simulations are run with strain rate dependency.

[picture]

Fig.4. Force-displacement curves for the dry state simulation with and without strain rate dependence

Fig.4. Force-displacement curves for the dry state simulation with and without strain rate dependence

Friction sensitivity: The next step in this work is to examine the sensitivity of the model to friction. To accomplish this, simulations were carried out with three different friction coefficients 0.01,0.1 and 0.28 . The minimum friction coefficient describes the nearly frictionless state, while the maximum
friction coefficient describes a typical value for the friction between aluminum and steel [17]. Fig. 5 a) shows that the force required for forming increases as the friction coefficient increases. Observing the progression of the Y coordinate at different friction coefficients at the last time increment, it is noticeable that the differences are not as pronounced as in Fig. 5 b), but are still visible. Due to symmetry, this figure only shows half of the node numbers over the Y coordinate for the last increment. The contours of the workpiece at different friction values clearly overlap, only in the transition areas from the pocket to the flat surface of the die does the simulation with a friction value of μ=0.01 deviate slightly from the other two contours.

[picture]

Fig. 5. a) Force against displacement for three different friction coefficients b) Point of the Y coordinate against the node number at the last time step for the three friction coefficients

Fig. 5. a) Force against displacement for three different friction coefficients b) Point of the Y coordinate against the node number at the last time step for the three friction coefficients

Comparison of lubricated and unlubricated simulation: After examining the unlubricated condition, the CEL simulations for the lubricated condition are analyzed next. The lubricated case means that the pockets are 100% filled with lubricant. Fig. 6 shows the Y coordinate, the equivalent plastic strain, and the von Mises stress against the X-coordinate over various time increments for the unlubricated case (solid line) and the lubricated case with CEL (dashed line). All parameters are evaluated at the same material points. The friction conditions in the unlubricated state are μ=0.1. Here, μ=0.1 is selected because it is an intermediate value that is closer to fully lubricated conditions. The use of μ=0.01 was not considered appropriate because complete lubrication over the entire duration of the test is not realistic, e.g., because the lubricant can be squeezed out during the test. In addition, only the pockets are filled with lubricant at the start of the test. In general, the simulation model of the test setup does not show a high sensitivity to the coefficient of friction. In the simulation using the CEL method, a distinction is made between "general contact" and "surface-to-surface contact" as described above. Fig. 6 a shows that the dry and lubricated simulations do not differ from each other at time increments 1 and 5 . With increasing time increments, the contours diverge from each other in the area of the pockets. This effect is particularly pronounced in the middle pocket. It can also be seen that the deformation within the pockets is less in the lubricated state than in the unlubricated state. In quantitative terms, this means that the difference in height between the lubricated and unlubricated areas in the middle pocket is approximately 0.035 mm . This is about 30% of the total height reduction of 0.12 mm . In Fig. 6 b , a deviation between the two simulation methods can be seen from a time increment of 20 onwards. Furthermore, it is noticeable that two peaks form on both sides in the areas between the pockets at time step 26 at a displacement of 0.12 mm . The high equivalent plastic stress at these points is due to the fact that the strain at the edge of the pockets is high because they are located deeper. Fig. 6 c shows the differences in the variable von Mises stress in MPa. Here, it is noticeable that the differences only become apparent in later increments. In general, the differences between the two simulation methods are smaller than for the other two variables, but they are still noticeable. It is also noticeable that the stresses differ at the edges. In this area, the differences in the last increment are up to 50 MPa .

[picture]

Fig. 6. Y coordinate (a), equivalent plastic strain (b), and von Mises stress (c) over the X coordinate for both dry and lubricated conditions with 100% lubricant filling

Fig. 6. Y coordinate (a), equivalent plastic strain (b), and von Mises stress (c) over the X coordinate for both dry and lubricated conditions with 100% lubricant filling

The corresponding force-displacement curves for the two simulation methods presented in Fig. 6 can be seen in Fig. 7. There is a difference between the lubricated simulations with CEL and the simulations in dry conditions. The differences are over 20 kN at maximum displacement of 0,12 mm At the beginning, the two curves still lie on top of each other, but as the displacement increases, the difference between the two curves also increases, which is also caused by the different slope of both curves. Contrary to the basic assumption that the force is lower in lubricated tests than in the same unlubricated test, this study shows the opposite to the result. According to DIN 50323 Part 3, friction is defined as an interaction between contacting material areas of bodies that counteracts relative motion, whereby interface interactions between solids, liquids, and gases can occur [18]. In this test setup, the relative movement is relatively small, with a height reduction of 10% of the workpiece, so that the influence of friction is generally relatively small. Since the pocket is completely filled with lubricant in the lubricated simulation, additional force must be applied to compress the lubricant in the pocket or to push it out of the pocket at very high height reductions.

[picture]

Fig. 7. Force displacement curve for the lubricated and dry state

Fig. 7. Force displacement curve for the lubricated and dry state

Investigation at different filling levels: In a further simulation study, the influence of the amount of lubricant within the pocket was investigated. For this purpose, in addition to the unlubricated simulation and the simulation with 100% filling, a simulation with 75% and one with 95% pocket filling were also calculated. In the case of 75% filling, the height reduction is not sufficient for the lubricant in the pocket to come into contact with the workpiece. As a result, the force-displacement curve is almost identical to the unlubricated curve (see Fig. 8). In the case of 95% pocket filling, there is no contact between the workpiece and the fluid at the beginning. However, during the forming process, the lubricant comes into contact with the die. This behavior can also be seen in the forcedisplacement curves, as the curve with complete filling lies slightly above this.

[picture]

Fig. 8. Force displacement data at different filling level

Fig. 8. Force displacement data at different filling level

Fig. 9 shows the variables Y coordinate, equivalent plastic strain, and von Mises stress against the X coordinate for the three different fill levels. At a fill level of 75%, the strain is higher in the area of the pockets. This is also reflected in Fig. 9 a in the course of the X and Y coordinates.

[picture]

Fig. 9. Y coordinate (a), equivalent plastic strain (b), and von Mises stress (c) over the X coordinate for three different filling level

Fig. 9. Y coordinate (a), equivalent plastic strain (b), and von Mises stress (c) over the X coordinate for three different filling level

Discussion

The original version of this paper is available on https://www.scientific.net/MSF.1184.53.pdf

The obtained results in the present study indicate that, while conventional friction-based models adequately capture the global forming behavior under unlubricated conditions, they fail to provide an accurate description once lubricant accumulation within surface pockets modifies the interfacial mechanics. The initial comparison between simulations with and without strain rate dependency highlights the necessity of incorporating rate-dependent material behavior for EN AW-6016. Without taking different forming speeds into account, neglecting strain rate effects leads to a underestimation of the forming force in the total amount of 2500 N . This confirms that accurate flow curve representation is essential for quantitative predictions and should be considered within these analogue models. The friction sensitivity study shows that increasing the Coulomb friction coefficient leads to higher forming forces. However, the influence of friction on the overall deformation profile and material flow is comparatively small, particularly away from the pocket transition regions. This is likely due to the fact that normal compression dominates due to the comparatively large reduction in height. A detailed investigation of the individual zones in plane-strain upsetting tests, such as the sliding zone and the sticking zone, has already been carried out in a previous study [7].
This observation supports the assumption that a numerical representation of the lubricated forming process cannot be adequately described by the reduction in the coefficient of friction, as the lubricant actively participates in the forming process and thus contributes to the final surface profile. A key finding of this work is the pronounced difference between lubricated and unlubricated simulations when the pockets are filled with lubricant. Contrary to the common assumption that lubrication necessarily reduces forming forces, the CEL simulations reveal higher forces in the lubricated case
at large displacements. This behavior can be attributed to the compressibility of the lubricant within the pockets. In the CEL simulation, the contact area is larger from the outset, which means that a higher force is required for forming. As deformation progresses, additional work is required either to compress the lubricant according to the equation of state or to displace it laterally, leading to an increase in the force. This effect cannot be captured by classical Coulomb or Tresca friction models, highlighting the necessity of a CEL fluid representation when lubricant entrapment occurs. The comparison of different filling levels further emphasizes the importance of lubricant volume. At a filling level of 75%, the lubricant does not come into contact with the workpiece, resulting in force-displacement curves and strain distributions that are nearly identical to the dry case. In contrast, at 95% and 100% filling, the lubricant increasingly influences the final surface profile, stress, and strain, with higher filling levels leading to increased forming forces and reduced deformation within the pockets. These results indicate the existence of a critical filling threshold above which hydrodynamic and hydrostatic effects dominate the local forming behavior. Overall, the results demonstrate that the CEL approach provides additional physical insight into lubricant-structure interactions that cannot be obtained from simplified friction models.

Summary

The original version of this paper is available on https://www.scientific.net/MSF.1184.53.pdf

The present study investigates the applicability of a Coupled Eulerian-Lagrangian (CEL) approach for modeling lubricated plane-strain upsetting tests as an analogue to skin-pass rolling. The study leads to the following main conclusions:

Strain rate dependency of the material has a measurable influence on the forming force and must be considered for realistic simulation results.

Conventional friction models adequately describe the forming behavior without lubrication, but can reach their limits in the presence of trapped lubricants.

Lubricant entrapment within surface pockets can lead to increased forming forces, contrary to classical expectations, due to lubricant compression and displacement effects.

The lubricant level is a critical parameter. Below 95% filling level, the lubricant has only a negligible effect, whereas a nearly complete filling significantly changes the force, strain, and stress distributions.
Future work will focus on systematic parameter studies of lubricant properties and the correlation of numerical results with experimentally measured surface topographies in order to further validate the model for industrial skin-pass rolling applications. For this purpose, the elastic springback will also be mapped in the simulation model, as this will enable the comparison of surface topographies. Furthermore, fluid entrapment will be analyzed in more detail in future studies. Assuming that the realistic pockets are similar in size to the surface roughness, it can be concluded that 100% entrapment is unlikely. Therefore, investigating realistic surface conditions is an important next step.

Acknowledgements

The original version of this paper is available on https://www.scientific.net/MSF.1184.53.pdf

This research is funded by the Deutsche Forschungsgesellschaft (DFG, German Research Foundation, Project number 523163077) within the project "Simulative Abbildung der Fluid-Struktur-Kontakt-Interaktion in der Umformtechnik am Beispiel des Nachwalzens".

References

The original version of this paper is available on https://www.scientific.net/MSF.1184.53.pdf

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H. Kijima, N. Bay, Modelling of skinpass rolling by elasto-plastic analysis of plane strain upsetting, Journal of Materials Processing Technology 177 (2006) 509-512.

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L. Koch, Numerical mapping of the surface roughness during skin pass rolling of high strength aluminum alloys using the plane strain upsetting test, in: Material Forming: ESAFORM 2025, 2025, pp. 909-916.

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J. Slota, L'. Kaščák, L. Lăzărescu, Frictional characteristics of EN AW-6082 aluminium alloy sheets used in metal forming, AMME 41 (2024) 79-87.

Deutsche Institut für Normung e. V., Tribology; friction; definitions, types, conditions, quantities(DIN 50323-3), 1993.

This ePub file was created using machine-generated methods and automated software to convert the source text into an ePub format and may contain formatting inconsistencies from the original version. Formatting may vary depending on the device used.

The original version of this eBook is available on https://www.scientific.net/book/friction-and-wear-in-metal-forming/978-3-0364-1992-3

Materials Science Forum, ISSN: 1662-9752, Vol. 1184, pp 65-74

doi: 10.4028/p-DDRk7C

© 2026 The Author(s). Published by Trans Tech Publications Ltd, Switzerland.

Submitted:
2026-01-19

Revised:
2026-02-13

Accepted:
2026-02-23

Online:
2026-04-14

Improvement of Wear and Corrosion Resistance of X46Cr13 Martensitic Stainless Steel by Cryogenic Treatment

The original version of this paper is available on https://www.scientific.net/MSF.1184.65.pdf

I. Espinosa 1,a*, S. Menargues 1,b, D. Gutierrez 1,c, J.A. Picas 1,d, J. Navas 2,e, M.T. Baile 1,f

1 Department of Materials Science and Engineering, Universitat Politècnica de Catalunya (UPC), Spain

2 Departament d'Enginyeria Gràfica i de Disseny (DEGD), Universitat Politècnica de Catalunya (UPC), Spain

a*Isabel.Espinosa@upc.edu, b  Sergi.Menargues@upc.edu, c  Joan.David.Gutierrez@upc.edu, d Josep.Picas@upc.edu, eJavier.Navas@upc.edu, f Maite.Baile@upc.edu

Keywords: cryogenic treatment, wear, AISI 420, martensitic stainless steel, electrochemical corrosion.

Abstract

This study investigates the effect of deep cryogenic treatment on the tribological and electrochemical performance of X46Cr13 martensitic stainless steel, with a particular emphasis on the synergistic interaction between wear and corrosion and its microstructural origins. The material was subjected to conventional quenching and tempering and compared with heat treatment routes incorporating cryogenic processing. Hardness measurements, wear tests, and electrochemical characterization by Tafel polarization were combined with quantitative microstructural analysis. Cryogenic treatment induces a pronounced microstructural refinement through the transformation of retained austenite into martensite and the enhanced precipitation of fine chromium-rich carbides, predominantly M23C6 and M7C3. This process results in an increased carbide number density and a reduced average carbide area, leading to a more homogeneous carbide distribution within the martensitic matrix. The refined carbide population contributes to increased hardness and significantly improved wear resistance by effectively hindering plastic deformation and abrasive damage. Simultaneously, the stabilization of the martensitic matrix and the modified carbide-matrix interface promote the formation of a more uniform and stable passive film, improving corrosion resistance. The combined improvement in wear and corrosion behavior reduces the degradation rate under coupled mechanical and electrochemical loading, demonstrating a clear tribocorrosion synergy controlled by carbide characteristics. These findings highlight cryogenic treatment as an effective strategy for tailoring the microstructure of martensitic stainless steels to enhance their performance in aggressive and mechanically demanding environments.

1. Introduction

The original version of this paper is available on https://www.scientific.net/MSF.1184.65.pdf

The aim of this study is to establish relations between microstructure, wear and corrosion resistance in a Martensitic Stainless-Steel heat treated in different methods.
There are applications for martensitic stainless steels in corrosion resistance, dental and surgical instruments, cutlery and the mining sector in wear -resistant parts due to the combination of corrosion, wear, and mechanical properties.
The uses of stainless-steel guarantee a remarkable corrosion resistance which derives from its content of chromium ( >12wt% ) that forms a stable passivation layer of chromium oxide ( Cr2O3 ) protecting the bulk against corrosion [1]
This study tries to clarify the differences in microstructural features and relations with wear and corrosion resistance, when a different heat treatment are applied in AISI 420. According to the literature the reasons for this are that the extent of microstructural changes may be different for each steel classes [2].

Key factors influencing the strength of martensitic stainless steel include the amount of dissolved carbon in the matrix, the size of martensite laths, the size and distribution of carbides, and dislocation density[3].
The effect of cryogenic treatment on the microstructure of steels has been extensively studied:

Variations in the retained austenite amount and its other characteristics. The amount of austenite retained increases dramatically with increasing carbon content in room-temperature quenched carbon steel and its alloying elements dissolved.

The refined martensitic sub-structure.

Influence on the precipitation of nano-sized carbides during tempering.

Additional small globular carbide (SGC) formation.

During the austenitization of martensitic stainless steels, the ( Cr,Fe ) 23C6 carbides are dissolved, and the ferrite is transformed into austenite. The (Cr,Fe)23C6 carbides inhibit the grain growth of prior austenite grains [4].
As the tempering temperature increases, the dislocation density in the matrix decreases, and different carbides precipitate in the order of M3C,M7C3, and M 23 C 6 .[3]
Final microstructure will define tribological and electrochemical corrosion resistance.

2. Experimental

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2.1. Materials and preparation

In this work the material studied was Martensitic stainless steel X46Cr13. The chemical composition is shown in Table 1.

Table 1. Chemical composition to X46Cr13 (wt%).

Element | Fe | C | Cr | Mn | Si | P | S
W% | 85 | 0.44 | 13.7 | 0.86 | 0.29 | 0.029 | 0.005

X46Cr13 is an alloy composed of approximately 85−88% iron, 12−14% chromium, and 0.15−0.45% carbon. It offers high hardness and excellent wear resistance, making it suitable for cutting tools, medical instruments, industrial gears, and automotive components. This alloy includes variants such as 420 J 1 and 420 J 2 .
Stainless steel X46Cr13 can be classified into several subgroups based on its carbon content and specific properties.
X46Cr13 stainless steel is used for cutlery, kitchen tools, and medical instruments due to its corrosion resistance. It is also used in industrial machinery and automotive parts, where durability and wear resistance are important.

2.2. Heat treatment

The heat treatment of X46Cr13 steel was defined according to industry practices outlined in ASM Handbook Vol. 4 - Heat Treating[5], as well as in technical guides from European manufacturers of martensitic stainless steels.
The cycle was adjusted to achieve a target hardness, prioritizing wear resistance. Corrosion behavior was evaluated in a Tafel test.
Several scientific articles have studied the effect of cryogenic treatment variables on the microstructure and improvement of mechanical properties, with deep cryogenic treatment being the best rated [6].
In accordance with the above, three different thermal cycles have been evaluated (Table 2).

Table 2. Applied heat treatment parameters.

Ref

Tempering

Temp. mathematical formula

Cryogenization

Tmep. mathematical formula time mathematical formula

1rst annealing

Temp. mathematical formula

2on annealing

Temp. mathematical formula

TR | 1050 |  | 275 |
TRR | 1050 |  | 275 | 275
TCR | 1050 | superscript number | 275 |

Samples were named with the following acronyms, related to applied heat treatment: TR: Tempering and annealing; TRR: Tempering a doble annealing; TCR: Tempering + cryogenic treatment + annealing.
The studied material has been shaped by forging and provided by Matricir copany[7].
Heat treatments were performed using a Hobersal 12-PR/300 furnace for tempering and quenching. Tempering was carried out with oil, while cryogenic treatment consisted of immersion at −196∘C in liquid nitrogen.

3. Methods and Results

The original version of this paper is available on https://www.scientific.net/MSF.1184.65.pdf

3.1. Microstructural features

Microstructural observations were conducted with a LEICA MEF 4M optical microscope and analyzed using NIS-Elements BR software. Furthermore, SEM analysis was performed to further characterize the microstructural features as shown in Figure1.

[picture]

Fig. 1. X46Cr13. martensitic stainless steel. Microstructure of tempered martensite with intergranular and intragranular precipitates. Etched with Villela. a) TR. b) TRR and c) TCR.

Fig. 1. X46Cr13. martensitic stainless steel. Microstructure of tempered martensite with intergranular and intragranular precipitates. Etched with Villela. a) TR. b) TRR and c) TCR.

The SEM micrographs of the specimens after the heat treatments, Figure 1. showed the microstructure consisting of very fine lenticular martensitic matrix with globular carbides.
Several carbides may precipitate such as the M2X,M3C,M7C3,M23C6, and MC types. It shows the refinement of carbide.
Plastic deformation during martensitic transformation at low temperatures is a key factor for the formation of carbon clusters that serve as nucleation sites for fine carbides. The martensitic transformation at low temperatures is accompanied by plastic deformation due to the volume effect of the martensitic transformation, which leads to the capture of immobile carbon atoms by gliding dislocations, generating carbon clusters. During tempering, these carbon clusters serve as nucleation sites for the precipitation of fine carbides.
In the following sequence of pictures, Figure 2 one can observe the difference in carbide density between TRR, TCR and TR heat treatments. Moreover, a reduction in particle size of carbides is also observed.

[picture]

Fig. 2. Sequence of SEM micrographs for TR, TRR and TCR heat treated samples.

Fig. 2. Sequence of SEM micrographs for TR, TRR and TCR heat treated samples.

Figure 3 shows the evolution of the quantity and size of carbide precipitates for the different heat treatments carried out.

[picture]

Fig. 3. Density and size of carbides are compared between of TR, TRR and TCR heat treatments.

Fig. 3. Density and size of carbides are compared between of TR, TRR and TCR heat treatments.

Quantitative analysis of carbides revealed that the TCR treatment has the smallest average size ( 0.09 μm) and lower carbide density, which correlates with the lowest observed wear rate. In contrast, the TR treatment showed larger and denser carbides, associated with higher wear and poorer corrosion behavior.
In several studies an average carbide size reduction after the application of cryogenic treatments was reported [8]. In this work, the observed reduction was very relevant. TCR carbide average size decreases by 40% related to conventional treatment TR. In the case of TRR the decrease was 20%. TCR characteristic could be attributed to segregation of carbon atoms promoted by cryogenic treatment evolving to precipitation of carbides in subsequent tempering.[6]

Complementary data obtained by X-ray diffraction (XRD), a presence of 8% retained austenite in the TR specimen.

3.2. Hardness

The hardness was measured on a polished mirror surface of the material by using a Duramin5 microhardness tester (Struers) with a Vickers diamond indenter under an applied load of 300 gf , according to ASTM E384 standard (HV0.3). Thirty data points were collected and averaged for each hard result.

Table 3. Average HV0.3 values for TR, TCR and TRR samples.

HV0.3

Hardness

variation

TR | 605 |
TCR | 617 | superscript number
TRR | 515 | superscript number

As shown in Table 3, the hardening response of AISI 420 is not significantly influenced by the cryogenic treatment. In contrast, the reduction in hardness observed after applying double tempering is significant.
It has been widely reported that cryogenic treatment can promote the transformation of retained austenite into martensite in steels, whereas precipitation of carbides may have little contribution to improved hardness [6].
Tempering is widely used to relax the martensitic microstructure, promoting carbon diffusion and carbide nucleation; therefore, the resulting hardness reduction is very common.

3.3. Tribological properties

Tribological performance under dry sliding conditions was evaluated using a Pin-on-disc tribometer (CSM Instruments), following ASTM G99-23 wear testing standards. Samples were ground and polished to a 1μ m diamond finish. Tests were performed with a 6 mm alumina pin sliding against the polished specimens at 0.10 m· s−1 under an applied load of 10 N applied load. Experiments were conducted at 25∘C and ~50% relative humidity. After a sliding distance of 0.8 km , material loss was determined by measuring the cross-sectional area of wear tracks with a profilometer (Taylor-Hobson Form Talysurf Plus roughness tester.
In Figure 4 a schematic Pin- On - Disc is shown.

[picture]

Fig. 4. Schematic drawing of the Pin-on-disc; Image from Pin- On- Disc used in the trials.

Fig. 4. Schematic drawing of the Pin-on-disc; Image from Pin- On- Disc used in the trials.

In Figure 5.- can be seen no statistically significant differences were observed in the friction coefficient between treatments, suggesting that wear resistance is predominantly influenced by microstructural modifications rather than changes in surface friction.

[picture]

Fig. 5. Coefficient of friction for different Heat Treatment a) TR, b) TRC and c) TRR.

Fig. 5. Coefficient of friction for different Heat Treatment a) TR, b) TRC and c) TRR.

In terms of Specific Wear Rate [m3·( N· m)−1], Fig 6 shown dry pin-on-disc tests showed that the TCR treatment had the lowest specific wear rate ( 65% lower respect TR), followed by TRR ( 34.4% lower respect TR) and TR.

[picture]

Fig. 6. Values for specific wear rate [ m3( N m)−1 ] obtained in pin-on-disc test.

Fig. 6. Values for specific wear rate [ m 3 ( N m ) − 1 ] obtained in pin-on-disc test.

In Figure 7 wear track of TR ( a,b ) and TCR ( c,d ) sample after the wear test is shown. It is observed difference in wear mechanism for different heat treatment. As could be seen, sever adhesion took place during sliding, with a high amount of plastic deformation of the surface and plate-like debris in the surrounding of the wear track. The worn surface appears to be metallic, without signs of oxidation.

[picture]

Fig. 7. Micrographics SEM, appearance of the wear channel. A and B for TR and Cand D for TCR.

Fig. 7. Micrographics SEM, appearance of the wear channel. A and B for TR and Cand D for TCR.

Although all samples exhibited comparable coefficients of friction, significant differences were observed in their wear responses as a function of the heat treatment. The quenched and tempered condition showed a predominantly severe abrasive wear mechanism, evidenced by deep and continuous grooves aligned with the sliding direction, together with localized adhesive damage and material removal. In contrast, the cryogenically treated and tempered samples presented a noticeably milder wear regime, characterized by shallower grooves and a more uniform worn surface, indicating a transition from micro-cutting to micro-ploughing. The reduced occurrence of adhesive features and surface delamination suggests improved microstructural stability during sliding. This enhanced wear resistance is attributed to the reduction of retained austenite and the increased stability of the martensitic matrix promoted by the cryogenic treatment, which effectively limits plastic deformation and material loss.[9]
While adhesive wear resistance mainly requires higher hardness, abrasive wear also requires primary carbides[10].

3.4. Corrosion resistance

Electrochemical tests were carried out using a Bio-Logic VMP3 potentiostat with EC-Lab V9.52 software, employing a three-electrode cell in a 3.5%NaCl electrolyte.
Thermodynamic parameters were determinate: corrosion potential Ecorr and Corrosion current Icor, Tafel constants βa (anodic coefficient) and βb (cathodic coefficient), the corrosion rate.
Corrosion rate has been determinate using the following equation:

Vcorr=K icorr EWρ(1)

Where: Vcorr:  corrosion rate ( mm· year −1 ); Icorr  : corrosion current density ( μA·cm−2 ); EW : equivalent weight of the metal; ρ : metal density; K : constant 0.00327 mm • year −1
In Table 4 It can be the obtained values.

Table 4. Obtained values for electrochemical test using Eq.1.

Sample | Ecorr [mV] | Icorr [μA/cm²] | Bc | Ba | Chi² | equivalent weight [g/eq] | density [g/cm³] | surface area [cm²] | corrosion rate [mm/year]
TCR 0 | -337,06 | 1,03 | 199,39 | 146,37 | 224,55 | 27,6 | 7,7 | 0,78 | 0,0155
TCR 1 | -308,84 | 0,36 | 122,44 | 158,80 | 306,10 | 27,6 | 7,7 | 0,78 | 0,0054
TCR 2 | -310,80 | 0,90 | 140,53 | 166,83 | 612,24 | 27,6 | 7,7 | 0,78 | 0,0136
TR 0 | -476,34 | 2,41 | 238,01 | 95,57 | 32,33 | 27,6 | 7,7 | 0,78 | 0,0362
TR 1 | -368,59 | 3,67 | 216,69 | 71,11 | 221,31 | 27,6 | 7,7 | 0,78 | 0,0551
TR 2 | -445,11 | 2,58 | 249,52 | 101,49 | 91.155,70 | 27,6 | 7,7 | 0,78 | 0,0388
TRR 0 | -302,44 | 0,55 | 150,16 | 84,81 | 52.044,00 | 27,6 | 7,7 | 0,78 | 0,0083
TRR 1 | -297,30 | 0,45 | 1.119,92 | 114,24 | 3.558,01 | 27,6 | 7,7 | 0,78 | 0,0068
TRR 2 | -370,68 | 1,39 | 178,00 | 90,80 | 6.803,53 | 27,6 | 7,7 | 0,78 | 0,0209

To evaluate the comparative corrosion, rate a simple representation is done in figure 8.

[picture]

Fig. 8. Comparative values representation for corrosion rate according to different applied heat treatment.

Fig. 8. Comparative values representation for corrosion rate according to different applied heat treatment.

The double tempering treatment (TRR) showed the lowest corrosion rates, mainly due to reduced residual stresses and a more uniform chromium distribution in the matrix, thereby enhancing the formation of a more stable passive film. The TCR treatment showed intermediate behavior, while the TR treatment showed higher corrosion rates, associated with a less stable microstructure and greater chemical heterogeneity.
The potenciodynamic polarization curves represented in Figure 9, show that the TRR treatment shifts the corrosion potential towards more noble values and reduces the corrosion current density compared to TR and TCR. This behavior is associated with a more stable microstructure, with a lower defect density and a more homogeneous distribution of chromium in the matrix favoring the formation of a more protective passive film. These results are according to the literature.[11]

[picture]

Fig. 9. Polarization curves for different applied heat treatments.

Fig. 9. Polarization curves for different applied heat treatments.

Fig 10 shows the corrosion surfaces corresponding to the three heat treatments after being exposed to an electrochemical corrosion test. In all cases, the predominant corrosion mechanism is pitting
corrosion, which initiates in chromium-depleted zones around chromium carbides. These areas act as weak points in the passive layer, leading to a reduction in the pitting potential.

[picture]

Fig. 10. SEM micrographs of the specimens after the corrosion tests. a) TR, b) TRR and c) TCR.

Fig. 10. SEM micrographs of the specimens after the corrosion tests. a) TR, b) TRR and c) TCR.

4. Conclusion

The original version of this paper is available on https://www.scientific.net/MSF.1184.65.pdf

Cryogenic treatment has no significant influence on hardness, probably because the little fraction of retained austenite, also happens in term of friction coefficient. The significantly improvement in wear resistance of the steel could be attributed to segregation of carbon atoms promoted by cryogenic treatment evolving to precipitation of carbides. This characteristic is well founded in microstructural results.

The TRR heat treatment provides the best corrosion resistance in AISI 420, followed by TCR, while the TR treatment shows the least favorable behavior.

The refinement in carbide size and homogeneous distribution generated in cryogenic and doble tempering was sufficiently notable to improve wear resistance of a low-carbon AISI 420 stainless steel.

Because the retained austenite in AISI 420 TR has been so low, the improvements in wear have been attributed to refinement and homogeneous distribution for carbides.

Could be assume the chromium depletion zone size thicken with the bigger are the carbides, so the corrosion potential and the corrosion rate increase.

Less density of carbides helps less spots for the passive layer, and could be indicated the matrix has a higher percentage of dissolved Cr. Following this argument, TRR and TCR heat treatments are more corrosion resistance than TR.

A positive correlation was observed between the average carbide size and the specific wear rate, indicating that larger carbides favor micro-tearing and delamination mechanisms.

The TRR treatment represents an intermediate condition, offering a compromise between wear resistance and corrosion resistance.
The greater hardness and lower specific wear rate observed in the TCR samples are associated with a more refined microstructure, characterized by smaller carbides and a lower fraction coefficient of retained austenite. In contrast, the TRR treatment, despite exhibiting the lowest hardness, showed the best corrosion resistance, highlighting the trade-off between wear resistance and corrosion resistance in X46Cr13 steel.
The average carbide size primarily governs wear resistance, while carbide density more directly influences corrosion behaviour, demonstrating that different microstructural parameters control different degradation mechanisms in X46Cr13 martensitic stainless steel.
Future work will focus on establishing quantitative correlations between grain size, carbide morphology, and passive film stability to provide deeper insight into the microstructural control of corrosion and tribocorrosion resistance.

References

The original version of this paper is available on https://www.scientific.net/MSF.1184.65.pdf

M. Maher, I. Iraola-Arregui, H. Ben Youcef, B. Rhouta, and V. Trabadelo, "The synergistic effect of wear-corrosion in stainless steels: A review," in Materials Today: Proceedings, Elsevier Ltd, 2021, pp. 1975-1990. doi: 10.1016/j.matpr.2021.05.010.

P. Jurči and I. Dlouhý, "Cryogenic Treatment of Martensitic Steels: Microstructural Fundamentals and Implications for Mechanical Properties and Wear and Corrosion Performance," Feb. 01, 2024, Multidisciplinary Digital Publishing Institute (MDPI). doi: .

H. J. Chen et al., "Effects of heat treatment on microstructure, mechanical properties and corrosion resistance martensitic stainless steel," Mater. Charact., vol. 223, May 2025, doi: 10.1016/j.matchar.2025.114909.

A. N. de Moura et al., "Effect of austenitization temperature on microstructure, crystallographic aspects, and mechanical properties of AISI 420 martensitic stainless steel," Materials Science and Engineering: A, vol. 909, Sep. 2024, doi: 10.1016/j.msea.2024.146835.

K H PRABHUDEV, Handbook of Heat Treatment of Steels. 1992.

R. F. Barron, "Cryogenic treatment of metals to improve wear resistance," Cryogenics (Guildf)., vol. 22, no. 8, pp. 409-413, Aug. 1982, doi: 10.1016/0011-2275(82)90085-6.

Disponible a: , "Matricir | Fabricación de matrices y rodillos a medida."

J. Anna Sajan, R. Reji, and A. Mohamed Rafi, "A Review on the Cryogenic Treatment of Stainless Steels, Tool Steels and Carburized Steels," Jun. 2020. [Online]. Available: .

A. Dalmau, C. Richard, and A. Igual - Muñoz, "Degradation mechanisms in martensitic stainless steels: Wear, corrosion and tribocorrosion appraisal," Tribol. Int., vol. 121, pp. 167179, May 2018, doi: 10.1016/j.triboint.2018.01.036.

Ph. D. F. George E. Totten, "Steel Heat Treatment - Metallurgy And Technologies," no. Metallurgy and technologies, 2006.

S. Van Der Meer, "Rijksuniversiteit Groningen Effect of Heat Treatment and Composition on Electrochemical Behaviour of AISI 420 Martensitic Stainless Steels," 2016.

This ePub file was created using machine-generated methods and automated software to convert the source text into an ePub format and may contain formatting inconsistencies from the original version. Formatting may vary depending on the device used.

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Materials Science Forum, ISSN: 1662-9752, Vol. 1184, pp 75-83

doi: 10.4028/p-W7c2NX

© 2026 The Author(s). Published by Trans Tech Publications Ltd, Switzerland.

Submitted:
2026-01-19

Revised:
2026-02-12

Accepted:
2026-02-23

Online:
2026-04-14

Assessment of the Potential of CO2 as a Lubricant in Cold Forging of Low-Carbon Mild Steel

The original version of this paper is available on https://www.scientific.net/MSF.1184.75.pdf

Diego Alba 1,a*, Alexander Weiß 1, b, Gerd Reichardt 1,c and Mathias Liewald 1, d

1 University of Stuttgart, Institute for Metal Forming Technology (IFU), Holzgartenstraße 17, 70174

Stuttgart, Germany

a  diego.alba@ifu.uni-stuttgart.de, b  alexander.weiss@ifu.uni-stuttgart.de, c  gerd.reichardt@ifu.uni-stuttgart.de, d  mathias.liewald@ifu.uni-stuttgart.de (*corresponding author)

Keywords: friction characterisation, cold forging, carbon dioxide, miniature spike test

Abstract

Lubrication plays a crucial role in cold forging, influencing key factors such as material flow, surface quality and tool wear. The current state of the art presents conventional mineral-oilbased lubricants as one of a range of effective solutions; however, they generate residues, require cleaning and pose environmental concerns. This work explores CO2 snow as a potential residue-free, sustainable alternative for cold forming of mild steels. Miniature spike tests were conducted to characterise friction behaviour under varying lubrication and surface conditions. The study demonstrates that CO2 snow can effectively reduce friction, promote favourable material flow and achieve surface finishes comparable to conventional oils, while eliminating residue and postprocessing requirements. These findings suggest that CO2 snow represents a promising eco-friendly lubrication strategy, offering both technical performance and environmental benefits for sustainable cold forging operations.

Introduction

The original version of this paper is available on https://www.scientific.net/MSF.1184.75.pdf

Cold forging of steels traditionally relies on phosphated semi-finished products in combination with mineral oil- or soap-based lubricants to ensure effective separation between tool and workpiece [1,2]. While these conventional lubrication systems provide reliable technical performance, they are exposed to ecological and economic challenges. It is estimated that 40−50% of 5 million tons of lubricants being used annually in the EU are ultimately released into the environment [3,4].

The increasing demand for green and sustainable manufacturing has raised environmental concerns associated with the usage and disposal of oils. Particularly, their inherent toxicity and non-biodegradability have brought forth significant environmental problems [5, 6]. For instance, mineral-oil-based lubricants in metal forming and machining have been associated with the generation of oil mist and aerosol emissions during processing, which accumulate on tools and workpieces and often require subsequent cleaning or filtration steps to avoid contamination and ensure workplace safety [7, 8]. Collectively, these factors represent a substantial sustainability challenge in modern cold forging processes.

To overcome such challenges, alternative lubrication strategies are being explored, among which the use of CO2 snow as a residue-free lubricant shows particular promise. CO2 snow is already employed industrially for precision cleaning and utilizes CO2 that is generally a byproduct of other chemical processes, avoiding additional greenhouse gas emissions [9]. Its key advantage in metal forming lies in its ability to evaporate completely after processing, eliminating cleaning and disposal steps, reducing water and chemical usage, and improving both ecological and economic efficiency [10]. By removing the environmental and operational burdens of conventional lubricants, CO2 snow could enable a step change in sustainable cold forming technology. While CO2 snow has been successfully applied in sheet metal forming [7, 10], its potential in bulk forming remains largely unexplored. On the context of sheet metal forming, liquid CO2 proved to be the most effective volatile lubricant for dry deep drawing, providing the lowest friction and wear among all tested media due to dry ice formation and surface cooling. In addition, CO2 offered a cleaning effect, improved process stability, and the potential to replace conventional oil-based lubricants, enhancing both environmental sustainability and economic efficiency [11].

Although the environmental and operational benefits of residue-free CO2 snow are clear, a critical challenge remains accurately evaluating its lubrication performance under realistic cold forging conditions. Friction directly influences material flow, tool wear and forming forces, making sensitive tribological characterisation essential. In this context, the spike test, developed by Isogawa, has emerged as a widely used method for this purpose [12]. In the spike test, a cylindrical billet with a centring pin is formed into a spike through combined upsetting and extrusion, producing a central extruded pin, a conical section with an opening angle of 6∘, and an upsetting zone approximated by a cone with an opening angle of 174∘ [13]. Small, wave-shaped grooves on the punch surface help control material flow during forming. A schematic representation of the spike test is shown in Fig. 1a. For such tests, force-displacement curves typically show an initial steep rise, a nearly linear region, and a final sharp increase as the contact area expands [14], as depicted in Fig. 1b. Several studies confirm that the spike test is an effective method for assessing lubrication performance in cold forging and enables the investigation of cold welding phenomena and optimising forming conditions under varying frictional scenarios [15-18].

[picture]

Fig. 1. (a) Schematic illustration of the spike test [19]; (b) typical force-displacement curve obtained from the test [14].

Fig. 1. (a) Schematic illustration of the spike test [19]; (b) typical force-displacement curve obtained from the test [14].

The objective of this investigation is to systematically evaluate the performance of CO2 snow in the cold forging of steels and to quantify its effect on friction under realistic process conditions. The results enable an assessment of CO2 snow as a potential sustainable, residue-free lubrication strategy for cold forging applications. This is achieved through an experimental approach employing the miniature spike test for friction characterization. Experiments are conducted on mild steel specimens under various lubrication and surface conditions, allowing the derivation of force-displacement responses as well as geometric changes of the specimens. Based on this methodology, the study provides a quantitative comparison between CO2 snow and conventional lubricants and establishes a framework for evaluating residue-free lubrication concepts in cold forging processes.

Experimental Procedure

The original version of this paper is available on https://www.scientific.net/MSF.1184.75.pdf

The friction characterisation was carried out using the miniature spike test, selected due to its high sensitivity to interfacial friction conditions in bulk metal forming [19]. The specimen geometry is shown in Fig. 2a and die geometry used in this study is shown in Fig. 2b. A modified version of the original die geometry developed by Isogawa [12] was adopted in this study. The scale of the experiment was adjusted because the diameter of the raw material used in the intended industrial process was insufficient for the original spike test setup.

[picture]

Fig. 2. (a) Geometry of the miniature spike test specimen; (b) geometry of the miniature spike test die.

Fig. 2. (a) Geometry of the miniature spike test specimen; (b) geometry of the miniature spike test die.

The workpiece material was a standard mild steel, C15E, supplied as bar material and subsequently machined into the required miniature spike geometry. Three lubrication conditions were investigated, namely CO2 snow, a conventional oil for cold forging and a non-lubricated condition. The influence of a surface preparation was also considered with different initial surface roughness conditions being addressed. To generate different surface conditions, the specimens were shot-blasted under three different pressure levels, while the unblasted surface served as the reference state. Spheric glass particles with a diameter of about 100μ m were used as shot blasting material. Table 1 summarises the different parameter combinations examined in the experimental campaign. For each condition three samples were tested and the results were averaged.

Table 1. Summary of the distinct parameter sets examined as part of this study.

Group | Lubricant | Shot blasting pressure level [bar] | Initial Roughness Sa [μm]
1 | CO₂ | 8 | 1.93
2 | CO₂ | 5.5 | 1.71
3 | CO₂ | 3 | 1.28
4 | CO₂ | Unblasted | 0.93
5 | Reference Oil | 8 | 1.93
6 | Reference Oil | 5.5 | 1.71
7 | Reference Oil | 3 | 1.28
8 | Reference Oil | Unblasted | 0.93
9 | Dry - No lubricant | 8 | 1.93
10 | Dry - No lubricant | 5.5 | 1.71
11 | Dry - No lubricant | 3 | 1.28
12 | Dry - No lubricant | Unblasted | 0.93

The experiments were performed on a hydraulic press with a maximum capacity of 2,000kN. A load cell and displacement sensor integrated into the press machine were used to acquire forcedisplacement curves during deformation. The experimental setup is depicted in Fig. 3.

[picture]

Fig. 3. Experimental setup for the miniature spike test.

Fig. 3. Experimental setup for the miniature spike test.

CO2 snow was applied in a sequential step to ensure full and repeatable coverage of both the die and the specimen surface. The CO2 application sequence is shown in Fig. 4: first CO2 snow was deposited on the die (Fig. 4a), then the specimen was positioned using a centring specimen holder (Fig. 4b), and finally CO2 snow was applied on the specimen itself before upsetting (Fig. 4c).

[picture]

Fig. 4. Step-by-step procedure for applying CO2 snow to the specimens' relevant surfaces.

Fig. 4. Step-by-step procedure for applying C O 2 snow to the specimens' relevant surfaces.

For optical roughness measurements, a μ-Surf confocal microscope from NanoFocus was used. The measurement setup is illustrated in Fig. 5a while Fig. 5b illustrates the regions selected for surface measurement before and after the experiment.

[picture]

Fig. 5. (a) Confocal microscopy for optical surface roughness measurement and detailed view of the measured area [19]; (b) measurement positions on the specimen surface before and after the experiment.

Fig. 5. (a) Confocal microscopy for optical surface roughness measurement and detailed view of the measured area [19]; (b) measurement positions on the specimen surface before and after the experiment.

Results & Discussions

The original version of this paper is available on https://www.scientific.net/MSF.1184.75.pdf

The force-displacement curves obtained from the experiments for the three lubrication conditions at an initial surface roughness of Sa=1.93μ m are shown in Fig. 6a, and the corresponding averaged maximum forces for all parameter sets are summarised in Fig. 6b. As expected, the overall trend indicates that sample group 9-12 (dry condition) exhibited the highest peak forces, reaching up to 342.52 kN for Group 12.

In contrast, sample group 5-8 (mineral oil lubrication) recorded the lowest average maximum forces, with the lowest maximum force of 288.98 kN for Group 7. Together with the final geometrical measurements, the force-displacement response provides valuable insights into the lubrication performance and the forming behaviour of each sample.

[picture]

Fig. 6. (a) Force-displacement curves illustrating the three tested lubricants; (b) mean values and standard deviations of the maximum force measured during the miniature spike experiments (average results of three independent tests performed for each parameter group).

Fig. 6. (a) Force-displacement curves illustrating the three tested lubricants; (b) mean values and standard deviations of the maximum force measured during the miniature spike experiments (average results of three independent tests performed for each parameter group).

From the initial analysis, some conclusions can be drawn regarding the performance of CO2 snow as a lubricant in cold forging-i.e. sample groups 1-4. Experiments conducted with CO2 snow show, on average over all groups, 5% higher forming forces compared to tests using the reference oil lubricant.

In comparison, forming forces captured in the tests conducted without any lubricant were 10.2% higher than those observed with CO2 and 15.7% higher than with the reference oil. These results clearly indicate that the use of CO2 as a lubricant leads to a relevant reduction in forming forces compared to unlubricated conditions. These results demonstrate the effectiveness of CO2 snow in reducing friction during cold forging.

Further, a representative miniature spike sample is shown in Fig. 7. Whereas Fig. 7a shows a sample from sample group 12 which represents high-friction conditions (dry lubrication) and Fig. 7b shows a sample from sample group 7 which represents low-friction conditions (mineral oil). The most notable geometric difference between both parts lies in the final height of the pin. The high-friction sample resulted in a smaller final height of the pin and a larger diameter of the flange, while the low-friction sample exhibited a greater height of the pin with reduced radial expansion of the flange. Lastly, Fig. 7c provides a summary of produced dimensions of the tested specimens.

[picture]

Fig. 7. Comparison of the specimen final geometry under (a) low- and (b) high-friction conditions; (c) final diameter and height measurements (average results of three independent tests performed for each parameter group).

Fig. 7. Comparison of the specimen final geometry under (a) low- and (b) high-friction conditions; (c) final diameter and height measurements (average results of three independent tests performed for each parameter group).

The observed geometric differences can be attributed to the friction-dependant material flow during the spike test. When friction at the die-workpiece interface is low, the material experiences little resistance and is able to flow more easily along the die wall. This favours the material flow into the pin of the pressed spike part. In contrast, high friction conditions hinder sliding and restrain axial material flow. As a result, the material is forced more radially outwards, leading to a more pronounced flange and a lower spike height. The spike geometry therefore serves as a direct indicator of the interfacial friction conditions, enabling a clear distinction between efficient lubrication and high friction conditions.

The surface quality was assessed through roughness measurements. Fig. 8 presents a visual comparison of two representative samples, highlighting differences in the surface finish. Qualitative observation indicates that CO2 snow lubrication results in a superior surface quality (Fig. 8a) compared to reference oil lubrication (Fig. 8b). In particular, an analysis of the flange region reveals that specimens processed with CO2 snow lubrication exhibit a noticeably shinier appearance, whereas oil-lubricated specimens show a duller surface. This observation is further supported by quantitative roughness measurements which report the arithmetical mean height Sa values for each tested condition, presented as bar charts in Fig. 9a and in tabular form in Fig. 9b.

[picture]

Fig. 8. Post-experiment visual comparison of the surface finish achieved with CO2 snow lubrication versus mineral oil lubrication.

Fig. 8. Post-experiment visual comparison of the surface finish achieved with C O 2 snow lubrication versus mineral oil lubrication.

[picture]

Fig. 9. (a) Surface roughness Sa measured for each tested parameter set after the experiment; (b) average values with standard deviations for each group.

Fig. 9. (a) Surface roughness S a measured for each tested parameter set after the experiment; (b) average values with standard deviations for each group.

Compared to the reference oil used in this study, CO2 snow consistently produced smoother surface roughness in all tested cases. This supports the conclusion that CO2 can be effectively used as a lubricant when improved surface quality is required. Relative to the initial state, all lubrication conditions demonstrated good performance, enhancing the surface by reducing roughness and producing a more uniform finish.

Conclusions

The original version of this paper is available on https://www.scientific.net/MSF.1184.75.pdf

This study investigated the performance of CO2 snow as a lubricant in the cold forging of mild steel (C15E) through an experimental approach. Miniature spike tests-widely used to evaluate friction characteristics-were conducted under various lubrication and surface conditions to assess frictional behaviour based on the forming forces required and the resulting dimensional changes of the specimens. The key findings of this work are summarized as follows:

CO2 snow significantly reduces forming forces compared to unlubricated, with average forces approximately 10% lower than unlubricated conditions.

Reference mineral oil achieves slightly lower forming forces than CO2 snow, though the difference is relatively small in the order of 5%.

CO2 snow consistently produces smoother surface finishes, indicating its suitability for applications requiring improved workpiece quality while avoiding residue and cleaning steps.

For the range of roughness values tested within this study, no clear trend was observed between initial surface roughness and lubrication efficiency. Furthermore, no direct correlation between initial and final surface roughness could be derived.

Both CO2 snow and mineral oil improve surface uniformity relative to the initial state, demonstrating effective friction control.

In conclusion, CO2 snow has been demonstrated as a viable alternative to conventional lubricants in cold forging of steel. It effectively reduces friction and enhances surface quality. Overall, CO2 snow emerges as a promising residue-free lubricant, offering a practical pathway towards more sustainable and environmentally friendly cold forging processes.

Outlook

The original version of this paper is available on https://www.scientific.net/MSF.1184.75.pdf

Following the results of this study, future work will potentially address the following aspects to further investigate the use of CO2 snow in cold forging:

Development of a FEM model to determine friction factors for the tested conditions, enabling the use of these values in future simulations for the design and optimisation of cold forged components.

Design of an automated CO2 snow application system to ensure uniform and controlled lubrication application across workpiece surfaces.

Extended experimental campaigns to evaluate the effect of CO2 snow on die surfaces over longer production runs, including surface degradation, with the aim of supporting high-volume or repeated part manufacturing.

Acknowledgments

The original version of this paper is available on https://www.scientific.net/MSF.1184.75.pdf

The authors would like to thank Ossenberg-Engels GmbH & Co. KG for supplying the steel material used for the investigated miniature spike specimens, as well as the conventional mineral oil lubricant, which made the investigation presented in this paper possible.

References

The original version of this paper is available on https://www.scientific.net/MSF.1184.75.pdf

N. Bay, The state of the art in cold forging lubrication, Journal of Materials Processing Technology 46 (1994) 19-40. DOI: 10.1016/0924-0136(94)90100-7.

P. Groche, S. Zang, C. Müller, D. Bodenmüller, A study on the performance of environmentally benign lubricants at elevated temperatures in bulk metal forming, Journal of Manufacturing Processes 20 (2015) 425-430. DOI: 10.1016/j.jmapro.2014.06.004.

W. Dresel, T. Mang, Lubricants and Lubrication, Wiley, 2017.

J.W. Raedt, Grundlagen für das schmiermittelreduzierte Tribosystem bei der Kaltumformung des Einsatzstahles 16 MnCr 5. Doctoral Thesis. RWTH Aachen University, Aachen, 2002.

M. Zhou, J. Yan, H. Wu, R. Guo, Z. Xing, S. Jiao, Z. Jiang, Lubrication effects on the surface quality control of hot rolled steels: A review, Tribology International 199 (2024) 109985. DOI: 10.1016/j.triboint.2024.109985.

S. Qiu, Z. Zhou, J. Dong, G. Chen, Preparation of Ni Nanoparticles and Evaluation of Their Tribological Performance as Potential Additives in Oils, Journal of Tribology 123 (2001) 441443. DOI: 10.1115/1.1286152.

M. Henn, G. Reichardt, R. Weber, T. Graf, M. Liewald, Dry Metal Forming Using Volatile Lubricants Injected into the Forming Tool Through Flow-Optimized, Laser-Drilled Microholes, JOM 72 (2020) 2517-2524. DOI: 10.1007/s11837-020-04169-6.

M.A.A. Khan, M. Hussain, S.K. Lodhi, B. Zazoum, M. Asad, A. Afzal, Green Metalworking Fluids for Sustainable Machining Operations and Other Sustainable Systems: A Review, Metals 12 (2022) 1466. DOI: 10.3390/met12091466.

M.C. Krieg, Analyse der Effekte beim Trockeneisstrahlen. Doctoral Thesis. Fraunhofer IRB, Berlin, 2008.

G. Reichardt, C. Wörz, M. Singer, M. Liewald, M. Henn, D.J. Förster, E. Zahedi, S. Boley, A. Feuer, V. Onuseit, G. Umlauf, P. Reichle, J.P. Barz, G. Tovar, T. Hirth, Tribological system for cold sheet metal forming based on volatile lubricants and laser structured surfaces, Dry Metal Forming Open Access Journal FPR 6 (2020) 128-165. DOI: 10.26092/elib/156.

P. Reichle, G. Reichardt, M. Henn, G. Umlauf, J. Barz, K.R. Riedmüller, M. Liewald, G.E.M. Tovar, Volatile Lubricants Injected Through Laser Drilled Micro Holes Enable Efficiently Hydrocarbon-Free Lubrication for Deep Drawing Processes, Int. J. of Precis. Eng. and Manuf.Green Tech. 10 (2023) 875-890. DOI: 10.1007/s40684-023-00505-8.

S. Isogawa, A. Kimura, Y. Tozawa, Proposal of an Evaluating Method on Lubrication, CIRP Annals 41 (1992) 263-266. DOI: 10.1016/S0007-8506(07)61200-1.

Q. Zhang, E. Felder, S. Bruschi, Evaluation of friction condition in cold forging by using Tshape compression test, Journal of Materials Processing Technology 209 (2009) 5720-5729. DOI: 10.1016/j.jmatprotec.2009.06.002.

P. Groche, C. Müller, J. Stahlmann, S. Zang, Mechanical conditions in bulk metal forming tribometers-Part one, Tribology International 62 (2013) 223-231. DOI: 10.1016/j.triboint.2012.12.008.

P. Marx, Mechanische Beschichtung für die Kaltmassivumformung von Stahl. Doctoral Thesis. Technischen Universität Darmstadt, Darmstadt, 2003.

W. Rehbein, I. Lange, D. Pazdzior, Substitution of Zinc Stearate in Cold Extrusion Processes, Goriva i maziva časopis za tribologiju, tehniku podmazivanja i primjenu tekućih i plinovitih goriva i inžinjerstvo izgaranja 54 (2015) 343-356.

M. Liewald, C. Mletzko, A. Felde, M. Meidert, Evaluation of lubricants for bulk metal forming of steel at elevated temperatures using double-cup-extrusion-test and spike-test, in: International Cold Forging Group Plenary Meeting Proceedings, International Cold Forging Group, China, 2009, pp. 167-172.

N. Bay, New Tribo-systems for Cold Forming of Steel, Stainless Steel and Aluminium Alloys, in: International Cold Forging Group Plenary Meeting Proceedings, International Cold Forging Group, France, 2013, 34.

K. Grötzinger, Optimierung eines tribologischen Tests zum Einsatz unterschiedlicher Schmierstofftypen bei der Kaltmassivumformung einer Aluminiumknetlegierung. Masters. University of Stuttgart, Stuttgart, 2018.

This ePub file was created using machine-generated methods and automated software to convert the source text into an ePub format and may contain formatting inconsistencies from the original version. Formatting may vary depending on the device used.

The original version of this eBook is available on https://www.scientific.net/book/friction-and-wear-in-metal-forming/978-3-0364-1992-3

Materials Science Forum, ISSN: 1662-9752, Vol. 1184, pp 85-96

doi: 10.4028/p-Q77VIf

© 2026 The Author(s). Published by Trans Tech Publications Ltd, Switzerland.

Submitted:
2026-01-19

Revised:
2026-02-23

Accepted:
2026-02-23

Online:
2026-04-14

A New Ex-Situ Method for Real Contact Area Determination for Sheet Metal Forming

The original version of this paper is available on https://www.scientific.net/MSF.1184.85.pdf

Ainhoa Guinea 1,a*, Andrea Aginagalde 1, b, Eneko Saenz de Argandoña 1,c, Joseba Mendiguren 1, d, Wilson Tato 1,e, Iñigo Llavori 1,f, Liam Blunt 2, g, Alaitz Zabala 1, h

¹Mechanical and Industrial Production Department, Mondragon Unibertsitatea, Mondragon, 20500, Spain.

2 School of Engineering, the University of Huddersfield, United Kingdom.

a  aguinea@mondragon.edu, b  aaginagalde@mondragon.edu, c  esaenzdeargan@mondragon.edu, djmendiguren@mondragon.edu, e wtato@mondragon.edu, f illavori@mondragon.edu, gl.a.blunt@hud.ac.uk, 'azabalae@mondragon.edu

Keywords: Real contact area, ex-situ method, tribology, normal loading, surface-Topography

Abstract

Experimental determination of the real contact area is essential for the development and validation of advanced friction models in sheet metal forming. However, existing experimental approaches are limited by the need for transparent counterfaces or third agents in the interface, high sensitivity to relocation errors, or unreliable assumptions about the contact shape. This study proposes a novel ex-situ method to determine the real contact area using topographical information from the original and deformed sheet surfaces. The approach identifies a minimum contact area with high confidence and reconstructs from it the full contact area. It provides the real contact area ratio, the contact topography and height distribution. The method is evaluated using results from normal load simulations based on the Pullen and Williamson model as the reference and is compared with other ex-situ methods from literature. Results demonstrate that the proposed method is robust against realistic relocation errors and yields more accurate contact area values than existing approaches. The method offers a reliable experimental tool for tribological analysis and friction modelling in sheet metal forming.

Introduction

The original version of this paper is available on https://www.scientific.net/MSF.1184.85.pdf

Sheet metal forming is a complex manufacturing process governed by a wide range of material, geometrical, and process parameters. Among these, tribological phenomena play a decisive role in controlling material flow, forming forces, tool life, and the final surface quality [1]. Consequently, accurate numerical simulation of sheet metal forming requires advanced friction models that can reliably capture the underlying contact mechanics. The interacting surfaces in forming operations are inherently rough, exhibiting asperities and valleys across a broad range of length scales [2]. When a sheet comes into contact with a tool, contact occurs only at discrete asperity junctions, which occupy a small fraction of the nominal contact area, as originally demonstrated by Bowden and Tabor [3]. This real contact area fundamentally governs frictional behaviour, making its accurate determination crucial for advanced friction model development and validation in sheet metal forming.

The first step is to predict the real contact area under normal loading conditions. To this end, numerous models have been developed which can be broadly categorized into asperity-based, fractal, and continuum approaches [4]. While the theoretical understanding of rough surface contact is well established, systematic experimental validation remains limited [5]. The primary challenge lies in the difficulty of directly measuring the real contact area in experiments [6]. To address this limitation, several experimental techniques have been proposed, which are classified as in-situ and ex-situ methods [7].

In-situ methods aim to measure the real contact area during loading, either indirectly or directly. Indirect methods infer the real contact area from ultrasonic signals [8] or electrical [9] and thermal resistance measurements [10]. However, these approaches require extensive calibration and are
limited in spatial resolution, restricting their practical applicability [7]. Direct in-situ techniques rely on optical access to the contact interface using a transparent counterface. Methods such as optical microscopy with CCD cameras [11], confocal microscopy [12], and optical interferometry [13] have been employed. Their main limitation is the requirement for at least one optically transparent and smooth surface, which restricts their applicability to rough-flat contacts and limits the range of materials and surface topographies that can be investigated.

Ex-situ methods estimate the real contact area after unloading. Some methods introduce a third medium at the interface, such as thin metallic films [14] or liquids [15], to reveal contact regions. However, the presence of an interfacial agent can alter the contact mechanics and compromise the validity of the results. Other approaches rely on surface topography measurements performed before and after contact. Relocation profilometry techniques compare height maps acquired at the same surface location before and after testing [10]. Although conceptually simple, these methods require extremely accurate surface relocation, particularly when deformations are small. Alternative strategies analyse changes in the height distribution functions of the surfaces and define a transition height above which all points are assumed to be in contact [16], [17]. A key limitation of these approaches is the implicit assumption that contact occurs on a flat plane, which is generally not valid for rough surface contacts.

Despite significant progress, there is a need for experimental methodologies that can reliably determine the real contact area without requiring transparent counterfaces, interfacial agents, or restrictive assumptions regarding the contact surface shape. The objective of this study is to address this gap by proposing a novel ex-situ method that exploits topographical information from the original and deformed sheet surfaces to identify contact points with high confidence and reconstruct the real contact area from them. The proposed approach is robust against relocation errors and provides not only the real contact area ratio, but also the associated contact topography and height distribution, thereby offering deeper insight into the contact interface. This will enable an accurate experimental determination of the real contact area for advancing the understanding of tribological systems and validating advanced friction models in sheet metal forming applications.

Review of Current Methods

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In this section, ex-situ methods reported in the literature that use topographical information from the contacting surfaces are reviewed.

Transition height methods. Transition height methodologies estimate the real contact area by comparing the height distribution functions of the original and deformed surfaces. The method assumes that asperities deform plastically once they come into contact with the rigid plane as purely elastic deformation of asperities on metallic surfaces occurs only at extremely small surface compression heights. They also assume the contact between a deformable rough surface and an ideal perfectly flat rigid plane. Hence, when a rough surface is plastically deformed against a perfectly rigid flat plane, the distribution of surface heights shifts toward lower values, producing a sharp cut-off peak. The position of this transition height (ht) defines the extent of the real contact area (Fig. 1a).

However, in practice the height distribution of the deformed surface does not exhibit a sharp transition but instead displays a gradual slope. This deviation arises from elastic recovery, small misalignments, and the roughness of the rigid counterpart. The absence of a distinct transition point complicates the determination of ht, which strongly influences the resulting contact area estimation.

Two main approaches have been reported to set ht. Westeneng [18], defined the transition height at the local peak of the distribution function (htMax) (see Fig. 1b). In contrast, Kucharski and Starzynski [17] proposed that asperities in contact correspond to points lying outside the original height distribution, establishing the transition height at the deviation between the original and deformed functions ( htDif  ) (see Fig. 1c). The methods are named transition height peak method and transition height difference method respectively for the rest of the document. In both approaches, the real contact area is estimated from the material ratio corresponding to the ht (see Fig. 1b and c)

[picture]

Fig. 1 a) Height distribution function of a sheet deformed against an ideally flat rigid plane. Height distribution function of sheet deformed against a tool with an Ra=0.1μ m and contact area determined with b) Transition height peak method and c) Transition height difference method.

Fig. 1. a) Height distribution function of a sheet deformed against an ideally flat rigid plane. Height distribution function of sheet deformed against a tool with an R a = 0.1 μ m and contact area determined with b) Transition height peak method and c) Transition height difference method.

Relocation methods. The relocation height method uses the original and deformed 3D surface topography to determine the real contact area. When two surfaces come into contact and deform plastically the height at the contact points is reduced. Hence, if measurements of the same location are acquired before and after deformation (see Fig. 2a and b respectively), the undeformed topography can be subtracted from the deformed topography. Ideally, a positive difference indicates a reduction in height, confirming that the point was in contact (Fig. 2c). Therefore, the real contact area ratio, the distribution of contact patches, defined as continuous surface regions where neighbouring asperities are in contact with the counterface, and the height distribution function of the contact area can be determined.

[picture]

Fig. 2 a) Original sheet surface. b) Deformed sheet surface. c) Plastic deformation of the contact points

Fig. 2. a) Original sheet surface. b) Deformed sheet surface. c) Plastic deformation of the contact points

Accurate relocation is therefore essential; otherwise, observed changes in the profile could be mistakenly attributed to misalignment rather than plastic deformation. This is particularly critical in small deformation scenarios, where the noise introduced by the relocation procedure and measurement process may not be possible to filter out.

Novel method

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When a sheet is compressed against a tool, sheet surface asperities flatten and the overall height range decreases compared with the original surface (see Fig. 3a). Height levels present in the original topography but absent after deformation indicate asperities that have undergone plastic deformation. It should be noted that the mean plane of the surfaces is taken as the reference plane for the height levels. Consequently, these locations must have been in direct contact with the tool and constitute the minimum contact area. The key idea of this method is to use the information from the minimum contact area to estimate the real contact area.

The determination of the minimum contact area is performed through the following steps. First, the minimum contact points are identified by determining the height corresponding to 99.7% of the material ratio curve of the deformed surface, denoted as hCMax  (see Fig. 3a). The 99.7% threshold is chosen based on the three-sigma rule to exclude extreme outliers. The original surface is then thresholded at this level to extract the spatial distribution of the minimum contact points (see Fig. 3b).

This is converted into a mask (Mask AMin  ), (see Fig. 3c). Finally, Mask AMin  is applied to the deformed surface obtaining the minimum contact area surface and height distribution of it (see Fig. 3d and e respectively).

[picture]

Fig. 3 a) Height distribution function analysis to determine Amin. b) Topography of the minimum contact area in the original sheet surface. c) Mask of the minimum contact area (Mask AMin  ).
d) Topography of the minimum contact area in the deformed sheet surface. e) Height distribution function of the minimum contact area

Fig. 3. a) Height distribution function analysis to determine Amin. b) Topography of the minimum contact area in the original sheet surface. c) Mask of the minimum contact area (Mask A Min ). d) Topography of the minimum contact area in the deformed sheet surface. e) Height distribution function of the minimum contact area

The minimum contact area represents a conservative lower bound of the real contact area. Based on the identified minimum contact area, the real contact area is reconstructed under the following physically grounded assumptions: (i) At least one point has been detected in each contact patch of the real contact area. (ii) The remaining contact points are located in the vicinity of these detected points. (iii) All contact points have heights within the height range defined by the minimum contact points on both the original and deformed surfaces. (iv) The smallest height difference between the original surface and the minimum contact surface represents the minimum plastic deformation ( δmin  ) experienced by the asperities.

The reconstruction proceeds as follows. First, the lowest height of the determined minimum contact area (Fig. 3e), hCMin , is obtained from the 99.7% material ratio curve. δMin  is determined by subtracting the identified contact points of the original sheet topography and the minimum contact surface (surfaces shown in Fig. 3b and d respectively) and determining the value of 99.7% of the material ratio curve.

The perimeter of Mask AMin  (Fig. 4a) is then iteratively expanded. At each iteration, newly included points are retained only if their height on the deformed surface exceeds hCMin . This procedure is repeated until no additional points are identified, resulting in Mask AMax  (see Fig. 4b).

This mask is then filtered in two consecutive consistency checks to ensure that the reconstructed real contact area reflects physically admissible contact conditions and is robust against measurement noise, surface relocation errors, and purely morphological artefacts. First, Mask AMax  is filtered by applying it to the original surface and removing all the points that have a height smaller than hCMin , ensuring geometric compatibility with pre-existing asperities yielding Mask Filtered (see Fig. 4c). Mask Filtered is further refined applying it to the surface obtained from subtracting the original topography from the deformed one. Points with deformation values smaller than δMin  are removed, resulting in the final mask Mask AReal  (see Fig. 4d).

[picture]

Fig. 4 Real contact area determination procedure through masking a) Mask AMin . b) Mask AMax . c ) Mask Filtered d) Mask AReal .

Fig. 4. Real contact area determination procedure through masking a) Mask A Min . b) Mask A Max . c ) Mask Filtered d) Mask A Real .

Mask AReal  is finally applied to the deformed sheet surface obtaining the real contact area ratio (RCA), the topography (see Fig. 5a), and height distribution function (see Fig. 5b) of the real contact area.

[picture]

Fig. 5 Real contact area a) Topography, b) Height distribution function

Fig. 5. Real contact area a) Topography, b) Height distribution function

Material and Methods

The original version of this paper is available on https://www.scientific.net/MSF.1184.85.pdf

To assess the capability of the proposed method to determine the real contact area ratio (RCA), as well as, the corresponding contact topography and height distribution, normal load simulations were performed. The contact area calculated in the simulation is considered as the true contact area for validation ( ATrue  ). It needs to be highlighted that no experimental normal-loading tests were performed because the real contact area cannot be directly measured and depends on the chosen method, preventing meaningful validation. Instead, simulations were used because they provide a known contact area under ideal conditions, allowing the accuracy of different estimation methods to be compared. The goal of this work is to assess the validity of the methods to estimate the real contact area not to validate the accuracy of the simulation itself.

The Pullen and Williamson model with volume conservation was adopted [19], based on the following assumptions: (i) the tool is considered rigid; (ii) plastic material behaviour of the sheet; (iii) the asperity persistence effect is ignored; (iv) material displaced from contacting asperities is uniformly redistributed over the non-contacting regions; and (v) contact is frictionless.

To account for the tool and sheet surfaces, the equivalent surface method is employed [20]. Hence, the surfaces of the tool ( ZTool  ) and sheet ( ZSheet  ) are combined,

Zeq =ZSheet +ZTool (1)

where ZTool  and ZSheet  are matrices containing the Z-coordinates of the tool and sheet surfaces at each (x,y) position.

The Zeq  is conceptualized as a collection of discrete bars, with areas equal to the topography resolution (see Fig. 6). Zeq  becomes into contact against a rigid plane.

[picture]

Fig. 6 The equivalent surface crushed by a rigid flat plane. Adapted from [21].

Fig. 6. The equivalent surface crushed by a rigid flat plane. Adapted from [21].

The asperities are categorized into three distinct groups based on their interaction with rigid plane: (i) the bars that are in contact ( N ), (ii) bars that will become in contact due to the risen volume ( N* ) and (iii) the bars that will never come into contact (N**). As the volume is conserved, the volume of the crushed asperities must be equal to the volume of the material redistributed into the non-contacting areas as stated in Eq (2)

∑i=1NΔziΔA+∑j=1N*ΔzjΔA=∑j=1N*ΔujΔA+∑i=1N**ΔuiΔA(2)

where Δz is the crush height and Δu is the risen volume height (see Fig. 6) and ΔA the area of each bar. In order to determine Areal , Pullen and Williamson derived Eq (3)

Areal =FNPnom +H(3)

where Pnom  is the nominal pressure and H the material hardness. Hence, for each distance between surface d, resultant Pnom , deformed sheet and contact area topography and the RCA, can be calculated by Areal Anom .

Simulations were carried out using the material properties of DC01 steel. Uniaxial tensile tests determined a yield strength of 209 MPa . Additionally, the material hardness was measured through a Vickers hardness test with a 30 kgf load (HV30), resulting in a hardness value of 1066 MPa . Nominal pressures ranging from 10 MPa to 60 MPa in steps of 10 MPa were simulated. A total of 6 simulations were performed. The same polished die topography was employed for all the simulations while different surface topography measurements extracted from the same sheet material were employed for each pressure range. From the simulation, the RCA and deformed sheet topographies were obtained. The deformed topographies were then analysed using the proposed method and benchmarked against the state-of-the-art real contact area estimation methods reviewed in Section 2, using the simulation results as reference.

Sheet and die surface topographies were measured using a Sensofar SNEOX optical profilometer in interferometric mode with a 20× EPI objective (spatial sampling =645 nm ) to feed the simulation with real topographies. Post-processing of the acquired data was carried out with SensoMap Premium 7.4 metrology software. The surface form was removed using a first-order polynomial fitting (Foperator), while noise was suppressed with a Gaussian filter ( λs=1.3μ m, S-filter).

The tool surface was a polished die with Ra=0.01μ m (see Fig. 7a and b) and the DC01 sheet surface was sanded to obtain a Gaussian height distribution with Ra=1.6μ m (see Fig. 7c and d)

[picture]

Fig. 7 Tool's a) Topography and b) Height distribution function. Sheet's c) Topography and d) Height distribution function.

Fig. 7. Tool's a) Topography and b) Height distribution function. Sheet's c) Topography and d) Height distribution function.

As the methods require topographical information from the same locations on the original and deformed sheet surfaces, measurements must be performed both before and after the normal loading.

In practice, exact repositioning is not possible, leading to a relocation error that can affect the accuracy of the real contact area determination. Hence, this error was quantified experimentally.

To ensure accurate relocation, each sheet was marked at its centre with a femtosecond-laser pattern that served as a reference. The sheet was placed and removed three times, with a surface measurement acquired after each placement. The automatic surface relocation algorithm available in SensoMap Premium 7.4 was then applied to reduce the relocation error between measurements. Differences between all measurement pairs ( 1−2,1−3, and 2−3 ) were computed. The relocation noise was defined as the largest standard deviation of these difference maps, which correspond to σ=0.584μ m.

To incorporate relocation error into the simulations, the original sheet topography was shifted in the x−y directions until the standard deviation of the difference maps between the original and shifted surfaces matched the measured relocation noise. The same shift was then applied to the deformed surface, corresponding to displacements of approximately 6.5μ m in the x-direction and 7μ m in the y-direction.

Results and Discussion

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Fig. 8a shows the RCA obtained for each simulated contact pressure using the methods reviewed in Section 2 (transition height difference, transition height peak and relocation methods) and the proposed method (New method), both under perfect relocation conditions (ideal) and with relocation noise introduced (real noise). The figure also includes the minimum contact area identified by the proposed method. Fig. 8b presents the error between the true contact area from the simulations and the values estimated by the different methods following Eq (4)

 Error =ATrue −RCA.(4)

[picture]

Fig. 8 a) Real contact area ratio results. b) Errors between the simulated true contact area and the real contact area estimated.

Fig. 8. a) Real contact area ratio results. b) Errors between the simulated true contact area and the real contact area estimated.

As observed, the relocation-based method yields RCA values close to zero. This occurs because only deformations exceeding the relocation noise can be detected, while the actual asperity deformations remain below this threshold as observed in Fig. 9. Consequently, the relocation method is not applicable under the considered conditions.

[picture]

Fig. 9 Height distribution of plastic deformations in the DC01 sheet simulated at 60 MPa .

Fig. 9. Height distribution of plastic deformations in the DC01 sheet simulated at 60 MPa .

The transition height methods do not exhibit a consistent linear relationship with contact pressure, indicating a strong sensitivity to the resulting height distribution function. Even under ideal conditions, both approaches show significant errors, with maximum deviations from the true contact area of 2.11% and 1.1% for transition height difference method and transition height peak method respectively. The transition height difference method systematically overestimates the real contact area whereas the transition height peak method tends to underestimate it. These trends arise from the common assumption that contact occurs above a single plane, whereas in reality contact occurs at multiple heights.

As shown in Fig. 10a, the transition height difference method captures all contact points but also includes original surface points that were already above htDif . In contrast, the transition height peak method may locally overestimate the contact area for the same reason, while simultaneously neglecting contact points located below htMax, which normally leads to an underestimation (see Fig. 10b). In some cases, these opposing effects may compensate, resulting in apparently accurate RCA values, as observed for the 10 MPa condition. However, the corresponding contact topography and height distribution are not physically representative. In conclusion, although truncation height methods are simple to apply when a clear ht can be identified, they suffer from fundamental theoretical limitations that compromise their reliability. Moreover, contacts involving non-Gaussian surface topographies may not exhibit a distinct local maximum in the height distribution, further limiting the applicability of these methods across different forming scenarios.

[picture]

Fig. 10 2D profile of the contact between two rough surfaces. Real contact area estimation errors with a) Transition height difference method and b) Transition height peak method.

Fig. 10. 2D profile of the contact between two rough surfaces. Real contact area estimation errors with a) Transition height difference method and b) Transition height peak method.

The minimum contact area identified by the proposed method follows the same trend as the true contact area and is insensitive to relocation noise, as it is determined from a threshold applied to the original surface at hCMax .

Under ideal relocation conditions, the proposed method slightly underestimates the real contact area. This discrepancy arises from the assumption that at least one pixel is detected in each contact patch, which is not always satisfied, as shown in Fig. 11. The smallest contact patches are occasionally
not detected, and their surrounding regions are therefore not explored. However, the maximum difference between the estimated and true contact areas is approximately 0.1%, indicating that the contribution of these undetected patches is negligible and likely within experimental variability. Consequently, the method is validated under ideal conditions.

[picture]

Fig. 11 Comparison between the true contact area, minimum contact area detected and non-detected contact patches of a DC01 sheet simulated at 60 MPa . Detailed view of the size of a non-detected contact patch.

Fig. 11. Comparison between the true contact area, minimum contact area detected and non-detected contact patches of a DC01 sheet simulated at 60 MPa . Detailed view of the size of a non-detected contact patch.

When relocation error is introduced, the proposed method exhibits a similar overall trend to the ideal case. Although the estimation error increases slightly, the method remains more accurate than the approaches reported in the literature. The maximum deviation from the true contact area is of 0.4%. The method tends to overestimate the real contact area at pressures up to 40 MPa and to underestimate it at higher pressures.

Overall, the results demonstrate that the proposed method provides a robust and reliable determination of the real contact area, even in the presence of realistic relocation errors and provides reliable characterization of the contact topography and the height distribution function.

Conclusions

The original version of this paper is available on https://www.scientific.net/MSF.1184.85.pdf

This study presented and validated a novel ex-situ methodology for determining the real contact area by employing the topographical information from the original and deformed sheet surfaces. The method is based on the identification of a minimum contact area that can be established with high confidence and on the subsequent reconstruction of the full real contact area using physically grounded assumptions. A systematic comparison with ex-situ methods found in literature was conducted. The following was concluded:

Relocation-based methods were found to be unsuitable for the studied conditions, as the asperity deformations were of the same order of magnitude as the relocation noise.

Transition height methods lead to an overestimation or underestimation of the contact area depending on where the transition height was established. The common assumption of contact occurring at flat plane convert the methods in unreliable.

The proposed method provides a way of determining a minimum contact area regardless of the contact conditions

The proposed method provided accurate and robust estimates of the real contact area over a wide range of contact pressures. Under ideal relocation conditions, the maximum deviation from the simulated real contact area was below 0.1%,

When realistic relocation errors were introduced, the method preserved the overall trend and remained more accurate than existing techniques, demonstrating a high tolerance to measurement uncertainty.

Beyond estimating the real contact area ratio, the proposed method also provides the contact topography and the height distribution function of the real contact points, offering deeper insight into
the tribological behavior of the system. These outputs are essential for validating normal-loading models used in advanced friction formulations. The method thus constitutes a robust and valuable experimental tool for advancing tribological modelling and improving the predictive accuracy of numerical simulations in sheet metal forming, as it enables the experimental determination of the real contact area under normal loading conditions. Notably, the method requires both pre- and post-deformation surface measurements, meaning it relies on more experimental information than transition-height-based approaches.

Future work will assess the applicability of the method to cases involving combined normal pressure and in-plane deformation; and the increase on real contact area due to sliding which are important aspects of advanced friction modelling in sheet metal forming. Additionally, the applicability of the method to contacts involving non-Gaussian surface topographies or surfaces with comparable roughness levels ( Ra of the same order of magnitude) will be evaluated.

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The original version of this paper is available on https://www.scientific.net/MSF.1184.85.pdf

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Materials Science Forum, ISSN: 1662-9752, Vol. 1184, pp 97-112

doi: 10.4028/p-5Z5Vcu

© 2026 The Author(s). Published by Trans Tech Publications Ltd, Switzerland.

Submitted:
2026-01-19

Revised:
2026-02-23

Accepted:
2026-02-23

Online:
2026-04-14

Sliding Distance Dependency and Third Body Particle Influence in Flat Strip-Draw Testing of Aluminum Sheet for Friction Characterization in Automotive Stamping

The original version of this paper is available on https://www.scientific.net/MSF.1184.97.pdf

Loris Rocchi 1,a*, Jan Filzek 2, b, Christian Leppin 1,c

1 Novelis R&D Centre Sierre, Switzerland

2 FILZEK TRIBOtech, Mühltal, Germany

a* loris.rocchi@novelis.adityabirla.com, b  filzek@tribotech.de,

c  christian.leppin@novelis.adityabirla.com

Keywords: flat strip-draw test, automotive aluminum sheet, third-body particles, sliding distance.

Abstract

In automotive sheet metal stamping, the friction coefficient for a given tribological system (sheet, lubricant, tool surface) is known to depend on contact pressure, sliding velocity and temperature. Furthermore, plastic deformation of the sheet can cause surface roughening, which will affect frictional response. Beyond these known effects, the flat strip-draw experiments using friction pads of various sizes and repeated sliding on the same sample presented in this paper indicate that the frictional response depends on the sliding distance and sliding history. Moreover, we found that third body particles generated by the wear of sheet metal in frictional processes substantially influence the level of the friction force. Our results suggest that there is a need to improve the friction models implemented in current commercially available simulation software for aluminum sheet metal stamping, to capture these substantial effects.

Introduction

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When determining the feasibility of a part in automotive sheet metal stamping, the elastoplastic deformation properties that define the formability of the sheet material are usually the focus. However, depending on the part and the die design, the frictional response can have just as much influence. Therefore, appropriate characterization and modeling of the tribological properties is important for the simulation-based forming process design of parts in the automotive industry [1]. Although it is still very common in the industry to use Coulomb friction [2], with a constant friction coefficient, it is well known that this assumption is incorrect, and that the actual friction between the sheet and the die in automotive stamping depends on local contact pressure [3], sliding velocity [4], lubrication [5],[6] and temperature [7], [8]. A further known effect is that for certain materials the plastic deformation of the sheet can cause surface roughening which will typically increase friction [9].
Flat strip-draw testing is a standard method for friction characterization in automotive applications. Beyond qualitative ranking of tribological systems, the test is used to identify friction-coefficient parameters for direct use in forming simulations and to calibrate multi-scale friction models [10], [11].
In aluminum stamping with uncoated cast-iron tools, a grayish transfer mark (Tribolayer) is often visible in regions of sliding contact, both on industrial parts and in laboratory strip-draw or drawbead tests. This transferred layer, commonly referred to as a tribolayer (third body particle, wear, and lubricant mix), forms under boundary lubrication from the interaction of the two contacting surfaces. The tribolayer coexists with topography evolution (surface flattening) that changes the real contact area [12]. In this paper, we demonstrate that the frictional response of aluminum sheets is influenced not only by changes in real contact area, but also by third body particles produced by wear. The relative contributions of these factors and their consequences for test-pad sizing in standardized stripdraw tests have not yet been systematically investigated.

Scope and contribution of this study to address these questions in three steps:

Implications for pad size and standards: We evaluate how the tribolayer's influence interacts with pad geometry, and we discuss consequences for pad-size selection in flat strip-draw testing.

Tribolayer composition: We characterize the chemistry and morphology of the tribolayer formed in aluminum-on-cast-iron contact under boundary lubrication.

Friction mechanism comparison: We quantify the tribolayer's contribution to the friction coefficient and compare it against the effect of surface flattening and the resulting real (as opposed to apparent) contact area. We further develop the observation to understand the tribolayer influence on the sliding distance and pressure history.
A practical motivation for step (3) is that VDA 230-213 [13] prescribes a pad size of 140×70 mm (length × width along the test direction). For typical automotive aluminum alloys, this rather large pad area limits the maximum attainable contact pressure in the test, which can be significantly lower than local pressures encountered in stamping die (e.g., in drawbeads) [14]. To bridge this gap, we complement the standard pad with smaller pads to access higher contact pressures and directly compare the resulting friction behavior to tests conducted at the VDA pad size.

Pad geometry Influence on Friction Results

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Considerations for using different friction testing set-ups

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During the flat strip draw friction test data acquisition, and as shown in Fig. 1, the following force sensor data is collected over time t : the normal force FN(t), and the frictional response forces FT1(t) and FT2(t) of each individual friction pad, and the pulling force Fp(t) at the grip of the sample. With FP(t)=FT1(t)+FT2(t), and the nominal contact surface of each pad A, the nominal average frictional shear stress acting on the pads can be defined as:

T(t)=FP(t)2 A(1)

and the nominal contact pressure as:

p(t)=FN(t)A(2)

The instantaneous average friction coefficient over time is then given by:

μ(t)=τ(t)p(t)=FP(t)2 FN(t)(3)

VDA 230−213 defines the size of the friction pad A=l×w=140 mm×70 mm (pad length in testing direction l, pad width w). Depending on the strength and thickness of the tested material sheet, as well as the resulting friction coefficient, the rather large pad length defined by the standard limits the normal force, respectively the normal contact pressure to values which are between 10 to 20 MPa for typical automotive aluminum sheet. However, the highest values observed locally in real automotive sheet metal stamping processes can be in the range of 50 MPa and even higher. At least, this is what numerical stamping process simulations suggest, since measuring local contact pressure in real dies directly is difficult [14]. Our original motivation for using smaller friction pad sizes was to overcome this issue and achieve higher nominal contact pressures in the test.

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Fig. 1. Schematic of standard single test sample configuration (si) for flat strip-draw testing.

Fig. 1. Schematic of standard single test sample configuration (si) for flat strip-draw testing.

As a further method to achieve higher nominal contact pressures we applied an innovative sandwich sample testing method. Instead of only one sheet sample, this method uses two samples lubricated only on the side which is in contact with the friction pads of the test equipment, and with a steel sample at the core between the two test samples, see scheme shown in Fig. 2. With the dry side of each test sample contacting the steel core, this sandwich sample testing configuration is designed to avoid any relative movement between the aluminum test samples and the steel core, which was verified by measuring before and after the sample length. It therefore prevents elongation and thinning of the test samples. Consequently, the friction coefficient can be measured independently of sample elongation, which is often observed at higher nominal contact pressure levels using the single sample method, and which may induce surface roughening that can considerably influence the frictional response. In general, none of the tests presented in this paper showed galling, meaning there was no visual adhesion of aluminum to the tool.

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Fig. 2. Schematic of sandwich test sample configuration (sw) for flat strip-draw testing.

Fig. 2. Schematic of sandwich test sample configuration (sw) for flat strip-draw testing.

Friction pads and test sample sizes used in the friction tests

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The friction pads are made of EN-GJS-700-2 cast iron. Each friction pad's surface is machined by rotary grinding to the surface roughness Sa=0.3μ m. Table 1 provides an overview of the friction pad sizes (testing direction 1 times width w effective contact area without the transition radius) and the corresponding test sample sizes used in the friction tests. Specimen width exceeds pad width to minimize edge effects and ensure the central contact zone experiences uniform nominal pressure.

Table 1. Friction pad sizes (contact area) and test sample sizes used in the friction tests.

Pad size (l × w) | Test sample size (l × w)
35 mm × 35 mm | 600 mm × 50 mm
70 mm × 35 mm | 600 mm × 50 mm
140 mm × 35 mm | 600 mm × 50 mm
140 mm × 70 mm | 600 mm × 100 mm

Sheet materials and friction test conditions

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Three different aluminum sheet variants were tested, see also Table 2: An aluminum alloy AA 6014 T4 with an EDT surface, as well as the same alloy with a mill finish surface (MF). Additionally, an aluminum alloy AA 5182-O with an EDT surface was tested. The sheet thickness for all variants was 1.0 mm . The test specimens were lubricated with 0.8 g/m2 of Walter Zepf Al200 using electro discharge equipment to ensure homogeneous application of the lubricant. The friction tests presented in this section were performed in rolling direction (later denoted by L ), at a temperature of 40∘C, and at a constant sliding speed of 80 mm/s.

Table 2. Tested sheet material variants.

ID | Material |
21 | AA 6014-T4 | EDT
22 | AA 6014-T4 | MF
23 | AA 5182-O | EDT

Results of friction tests with different pad geometries and discussion

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All of the friction coefficients presented here as a function of nominal contact pressure are values based on steady-state testing conditions after the initial static peak of each test. Fig. 3 shows the results for the flat strip-draw tests using different friction pad sizes with the sheet material variant AA 6014-T4 EDT. Even though the tribological system and testing conditions are the same for each curve, the test results are heavily influenced by the friction tool geometry. Fig. 4 shows the nominal average frictional shear stress as a function of nominal contact pressure, which is the same data, but without division by the nominal contact pressure. At low contact pressure, the shear stresses are nearly identical and approach zero. As pressure increases, the gap between some curves grows while others show the same trend. For example, at 30 MPa , both the 140×70 mm2 and the 140×35 mm2 pads show a shear stress of approximately 0.6 MPa despite their twofold width difference. However, at the same pressure, the shear stresses reported for the 70×35 mm2 and 35×35 mm2 pads are 1 and 1.8 MPa , respectively. These values are almost double and triple those for pads of the same width but different lengths. Therefore, the test results at the same pressure show a strong dependency on friction pad length. In contrast, pads of the same length but different widths show only marginal differences when considering the measurement uncertainty of the test. Examining the friction coefficient representation of the results more closely, one might argue that the tests using a 140 mm×35 mm pad differ from the tests using the same length with twice the width (70 mm) at low contact pressures (single sample and sandwich sample). However, we believe that this difference is due to the considerably increased measurement uncertainty at low contact pressures. In the frictional shear stress representation of the same results, this difference is considerably less pronounced and the overall trend of the dependency on pad length dominates is confirmed. Moreover, no significant difference was observed between the standard single test sample configuration (si) and the sandwich test sample configuration (sw), which indicates that elongation and thinning doesn't play a significant role in the cases which have been tested in standard single sample configuration.
The test results for the variants AA 6014-T4 MF shown in Fig. 6 and AA 5182-O EDT shown in Fig. 5 confirm the observations discussed above.

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Fig. 3. Friction coefficient as a function of nominal contact pressure from flat strip-draw friction test results with AA 6014-T4 EDT (ID 21) with identical tribological system for different sized friction pads (length in sliding direction x width). Tests were performed using either a standard single test sample test configuration (si) or a sandwich test sample configuration (sw). Aside from different nominal contact pressures, test conditions are identical.

Fig. 3. Friction coefficient as a function of nominal contact pressure from flat strip-draw friction test results with AA 6014-T4 EDT (ID 21) with identical tribological system for different sized friction pads (length in sliding direction x width). Tests were performed using either a standard single test sample test configuration (si) or a sandwich test sample configuration (sw). Aside from different nominal contact pressures, test conditions are identical.

[picture]

Fig. 4. Frictional shear stress as a function of nominal contact pressure from flat strip-draw friction test results with AA 6014-T4 EDT (ID 21) with identical tribological system for different sized friction pads (length in sliding direction x width). Tests were performed using either a standard single test sample test configuration (si) or a sandwich test sample configuration (sw). Aside from different nominal contact pressures, test conditions are identical.

Fig. 4. Frictional shear stress as a function of nominal contact pressure from flat strip-draw friction test results with AA 6014-T4 EDT (ID 21) with identical tribological system for different sized friction pads (length in sliding direction x width). Tests were performed using either a standard single test sample test configuration (si) or a sandwich test sample configuration (sw). Aside from different nominal contact pressures, test conditions are identical.

[picture]

Fig. 5. Frictional shear stress as a function of nominal contact pressure from flat strip-draw friction test results with AA 6014-T4 MF (ID 22) with identical tribological system for different sized friction pads (length in sliding direction x width). Tests were performed using either a standard single test sample test configuration (si) or a sandwich test sample configuration (sw). Aside from different nominal contact pressures, test conditions are identical.

Fig. 5. Frictional shear stress as a function of nominal contact pressure from flat strip-draw friction test results with AA 6014-T4 MF (ID 22) with identical tribological system for different sized friction pads (length in sliding direction x width). Tests were performed using either a standard single test sample test configuration (si) or a sandwich test sample configuration (sw). Aside from different nominal contact pressures, test conditions are identical.

[picture]

Fig. 6. Frictional shear stress as a function of nominal contact pressure from flat strip-draw friction test results with AA 5182-O EDT (ID 23) with identical tribological system for different sized friction pads (length in sliding direction x width). Tests were performed using either a standard single test sample test configuration (si) or a sandwich test sample configuration (sw). Aside from different nominal contact pressures, test conditions are identical.

Fig. 6. Frictional shear stress as a function of nominal contact pressure from flat strip-draw friction test results with AA 5182-O EDT (ID 23) with identical tribological system for different sized friction pads (length in sliding direction x width). Tests were performed using either a standard single test sample test configuration (si) or a sandwich test sample configuration (sw). Aside from different nominal contact pressures, test conditions are identical.

Analysis of the Tribo-Layer and Abrasion

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This section describes the analysis of the surface texture evolution, as well as the investigation of the grayish mark that develops along the sliding track of a flat strip-draw sample (Fig.8), and which also can be observed on stamped parts (Fig. 7).

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Fig. 7. Deep drawn stamping part showing typical grayish tribolayer in areas which experienced increased contact pressure.

Fig. 7. Deep drawn stamping part showing typical grayish tribolayer in areas which experienced increased contact pressure.

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Fig. 8. Sliding tracks on flat strip-draw friction test samples are visible as a grayish mark.

Fig. 8. Sliding tracks on flat strip-draw friction test samples are visible as a grayish mark.

Methods for third body debris characterization and analyzing surface texture evolution

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To characterize the constituents of the third-body (tribolayer) associated with the frictional contact, debris was collected from test samples by rinsing the sliding track with toluene and filtering the rinse. The collected material was weighed and examined by scanning electron microscopy (SEM) to assess the morphology of the observed particles. Additionally, laser diffraction was employed to determine the particle-size distribution.
To quantify changes in surface texture caused by frictional sliding, the test specimens of the flat stripdraw test results presented in the previous section were first marked with a microhardness indentation as of approximately 100μ m (fiducial marker) to enable precise localization of the same region before and after testing. Then, confocal laser scanning microscopy (CLSM) was performed on both the aluminum sheets and the tool pads in regions outside the specimen edges. Prior to lubrication and friction testing, each marked location was imaged at 20× magnification over a 2×2 mm2 area. The area was scanned with 3432×3432 points (lateral spacing 0.67μ m ) to obtain a high-resolution, threedimensional height map. After imaging, the specimens were lubricated and tested in a flat strip-draw. After testing, the surfaces were rinsed with toluene and were re-imaged at the identical location by returning to the reference indentation to ensure spatially consistent before-and-after comparisons.

Tribolayer analysis and composition

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Fig. 9 shows the weight per area of the filtered and converted particles from the cleaning rinse collected from friction test samples of material 23 of the corresponding test series in the previous section. As can be seen, the recovered particulate mass increases with the contact pressure applied in the friction test. This trend is plausible, as higher contact pressure will result in greater indentation by the tool, which likely abrades and mobilizes more third-body debris from the surface.

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Fig. 9. Weight per area of filtered and converted particles depending on contact pressure.

Fig. 9. Weight per area of filtered and converted particles depending on contact pressure.

The SEM image in Fig. 10 shows flake-like debris of material 23 tested at 5 MPa contact pressure. The lateral dimensions of the debris particles are markedly larger than the thickness. To evaluate the composition of the tribolayer and adhering particles, energy-dispersive X-ray spectroscopy (EDS) was performed at three locations across the surface. The elemental results are summarized in Table 2. The detected carbon is attributed to the conductive carbon tape used as the mounting support. The analysis indicates that the particles are composed predominantly of aluminum and aluminum oxide. The percentage weight values indicate that the particles are aluminum-rich, with Al contents between 55.58 and 82.06wt% and measurable oxygen ( 1.30−20.26wt% ), consistent with the presence of alumina within the tribolayer. Some magnesium is detected ( 2.14−2.74wt% ), while titanium and manganese appear only at trace levels ( ≤0.22wt% each); silicon and iron were not detected. The carbon signal (13.63-24.72 wt%) is attributed to the conductive carbon tape used for mounting and does not represent the intrinsic particle composition.

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Fig. 10. SEM image of wear debris generated during a flat strip-draw friction test.

Fig. 10. SEM image of wear debris generated during a flat strip-draw friction test.

Table 3. EDS measurement of the collected wear debris.

Spectrum Label | C [wt%] | O [wt%] | Mg [wt%] | Al [wt%] | Si [wt%] | Ti [wt%] | Mn [wt%] | Fe [wt%]
Spectrum 1 | 13.63 | 1.3 | 2.74 | 82.06 | 0 | 0.05 | 0.22 | 0
Spectrum 2 | 19.68 | 20.26 | 2.48 | 57.27 | 0 | 0.19 | 0.13 | 0
Spectrum 3 | 24.72 | 17.13 | 2.14 | 55.58 | 0 | 0.22 | 0.21 | 0

Particles size

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Particle-size distributions were obtained using a Mastersizer 3000 (Malvern Panalytical) operated in wet dispersion with toluene as the suspending medium. Laser diffraction yields volume-weighted, sphere-equivalent diameters. For thin flake-like debris, these diameters predominantly reflect the in-plane extent rather than the out-of-plane thickness. Accordingly, the distribution shown in Fig. 11 should be interpreted in conjunction with the SEM image Fig. 10, which confirms a flake morphology with submicron thickness.

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Fig. 11. Particles size distribution associated with the material tested with the Std 70×140 mm2 pads.

Fig. 11. Particles size distribution associated with the material tested with the Std 70 × 140 m m 2 pads.

Quantification of surface topography change by before-and-after observation

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We quantified the evolution of surface topography on the aluminum sheet by comparing confocal scans acquired before and after friction testing at the same location. The two height maps were registered using a feature-based algorithm developed by Colin Majoor [15] that uses the EDT or MF valleys as fiducials to align the before and scan results. See Fig. 12 as an example of such an analysis, where image (d) presents the difference map, Δz=zbefore −zafter , which reveals topographic changes at the pixel level: Δz>0 denotes local material removal (abrasion or asperity flattening), and Δz<0 indicates local uplift, respectively pile-up due to plastic flow. This approach provides a spatially resolved visualization of abrasion.

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Fig. 12. Before-and-after analysis of the surface of material 21 tested at 9 MPa : (a) Color map of the confocal observation before friction test; (b) Color map of the confocal observation after friction test; (c) Height distribution of images (a) and (b); (d) Color map image of the subtraction of the confocal image after (b) from the confocal image before friction test (a).

Fig. 12. Before-and-after analysis of the surface of material 21 tested at 9 MPa : (a) Color map of the confocal observation before friction test; (b) Color map of the confocal observation after friction test; (c) Height distribution of images (a) and (b); (d) Color map image of the subtraction of the confocal image after (b) from the confocal image before friction test (a).

Influence of Abrasion and Tribolayer in Multiple Draws

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Building on the investigation of the pad geometry influence in the initial section of this paper and the analysis of the tribolayer in previous section, this section examines how the surface flattening, abrasion and the tribolayer formation during sliding affect the friction coefficient. We attempt to evaluate the contribution of each of these mechanisms. For this purpose, we designed various flat strip-draw testing sequences in which the same sample is drawn multiple times on the same sliding track, as schematically shown in Figure 13.

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Fig. 13. Sequence of multiple draws of the same samples on the same sliding track to investigate the influence of surface flattening, abrasion and tribolayer creation.

Fig. 13. Sequence of multiple draws of the same samples on the same sliding track to investigate the influence of surface flattening, abrasion and tribolayer creation.

Unless explicitly mentioned otherwise, the following constituents are used in all tests presented in this section: Aluminum alloy AA 5182, 1.5 mm thickness with an EDT surface. The test specimens were lubricated with Walter Zepf Al200 with 1.0 g/m2 using electro discharge equipment. Cast iron EN-GJS-700-2 friction pads with a milled 35 mm×35 mm contact surface and Sa in the range 0-200.25μ m were used.

Experiments to single out the effect of the tribolayer

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The following three different experimental sequences involving multiple draws on the same sample were performed:

Experimental sequence A : The sample is drawn twice at 6 MPa contact pressure on the same sliding track (sample is drawn as is after the first draw, therefore the tribolayer is preserved).

Experimental sequence B : The sample is drawn twice at 6 MPa contact pressure on the same sliding track, with cleaning using MEK and re-lubrication with original amount before the 2nd  draw (the tribolayer is removed).

Experimental sequence C : The sample is drawn four times at 10 MPa contact pressure on the same track, with cleaning using MEK and re-lubrication with original amount before the 4th  draw (the tribolayer is preserved before the 2nd  and 3rd  draw, but it is removed before the 4th  draw).
In all three sequences the following test conditions were applied: testing in rolling direction at a room temperature (RT) and a sliding speed of 5 mm/s.
Fig. 14 shows that, in experimental sequence A , testing at 6 MPa , the second draw reduces the friction coefficient by 44% from 0.161 to 0.093 . Conversely, experimental sequence B , also testing at 6 MPa , but the second draw occurring after cleaning and re-lubrication, the friction coefficient decreases considerably less from 0.160 to 0.151 . Sequences A and B result in the same flattened sheet surface texture after the first draw and we assume the same real contact area results of those two tests at the start of the second draw. However, in sequence B the particle-rich tribolayer is removed, and the sample is freshly lubricated, potentially with even slightly more lubricant than the sample in sequence A after the first draw, as some lubricant may have been displaced from the sliding track. This result suggests that the tribolayer exerts a far stronger influence on friction in aluminum than the surface texture flattening alone and the resulting change of real contact surface area.
Furthermore, Fig. 14 shows that in experimental sequence C (testing at 10 MPa ), the friction coefficient decreases from 0.127 to 0.07 to 0.055 in the second and third draws, where the developed tribolayer is preserved. However, after cleaning the sample with MEK and re-lubricating it to the original amount prior to the fourth draw, the friction coefficient increases almost to the level of the first draw (approximately 0.12 ). This provides further strong evidence of the tribolayer's dominant influence on friction in aluminum shown in experimental sequence C .
Our interpretation of the observed behavior is that the flake-like debris in combination with the lubricant forms a tribolayer that acts as a separator between the tool and the aluminum surface, thereby reducing plowing of the tool into the flattened surface asperities. Preserving the tribolayer considerably reduces the shear resistance of this interface, causing the pronounced drop in friction coefficient shown in Fig. 14.
Finally, we would like to point out that all three sequences were performed without any galling being visually observed between any of the draw repetitions.

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Fig. 14. Effect of tribolayer on the evolution of the friction coefficient in flat strip-draw friction tests with multiple drawing of the same sample: experimental sequence A: sample used as is in 2nd draw, experimental sequence B: sample cleaning and re-lubrication before 2nd draw, experimental sequence C: sample used as is in 2nd and 3rd draw, cleaning and re-lubrication before 4th draw.

Fig. 14. Effect of tribolayer on the evolution of the friction coefficient in flat strip-draw friction tests with multiple drawing of the same sample: experimental sequence A: sample used as is in 2nd draw, experimental sequence B: sample cleaning and re-lubrication before 2nd draw, experimental sequence C: sample used as is in 2nd and 3rd draw, cleaning and re-lubrication before 4th draw.

Friction tests with multiple draws on the same sample

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The following two different experimental sequences involving multiple draws on the same sample were performed:

Experimental sequence D: Each sample is repeatedly drawn at the same contact pressure as in the first draw (3 draws for each sample in total).

Experimental sequence E : Each sample is drawn with 7 MPa nominal contact pressure in the 1st  draw. Various contact pressures are applied in the 2nd  draw of a sample.
In both sequences the following test conditions were applied: testing in rolling direction at a room temperature (RT) and a sliding speed of 5 mm/s.
The results in Fig. 15 show that in experimental sequence D , where the same contact pressure is repeatedly applied to each sample, the friction coefficient decreases with each draw. Based on the observations in experimental sequences A,B and C , we assume this reduction is mainly caused by the tribolayer rather than the surface flattening. For most contact pressures, the decrease in the friction coefficient between the first and the second draw is greater than the decrease between the second and the third draw, except in the test sequence at the lowest contact pressure of 2 MPa . In experimental sequence E, each sample is first tested at a contact pressure of 7 MPa creating the same amount of wear in the tribolayer. Then samples are tested at different contact pressures in the second draw, and the following can be observed: the smaller the contact pressure in the second draw compared to the initial 7 MPa draw, the greater the decrease in the friction coefficient. Conversely, the greater the contact pressure in the second draw compared to the initial 7 MPa draw, the smaller the decrease in the friction coefficient. In the latter case, the friction coefficient results of the second draw converge to those of the first draw at the same contact pressure in sequence D. With increasing contact pressure, the initial 7 MPa draw in sequence E has a vanishing effect.
The curve of the second draw shows the opposite trend after testing all samples at 7 MPa , considering that the friction coefficient curves typically decrease monotonically with increasing contact pressure. However, in anticipation of expected changes to the surface texture and particle formation in the preliminary test at 7 MPa , it is plausible that, in subsequent draws at lower contact pressures, the reduction in the coefficient of friction is more significant while, at higher contact pressures, the reduction is less significant. In any case, this experiment demonstrates that the contact pressure history should be considered when modeling and characterizing the frictional response of aluminum sheets.

[picture]

Fig. 15. Friction tests with AA 5182-O EDT and lubricant Al200 - Development of the friction coefficient in friction tests with multiple draws on the same sample. Results for sequence D (each sample repeatedly drawn at same contact pressure as in 1st  draw, 3 draws in total), and results for sequence E ( 1st  draw at 7MPa,2nd  draw at different contact pressures).

Fig. 15. Friction tests with AA 5182-O EDT and lubricant Al200 - Development of the friction coefficient in friction tests with multiple draws on the same sample. Results for sequence D (each sample repeatedly drawn at same contact pressure as in 1 st draw, 3 draws in total), and results for sequence E ( 1 st draw at 7 M P a , 2 nd draw at different contact pressures).

Comparison of long pad results with multiple draw tests using short pads

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The first of two experimental sequences in this subsection follows sequence D and adds a fourth draw of each sample on the same sliding track at the same contact pressure. Tests were performed again in rolling direction, but at a temperature of 40∘C instead of RT, and at a considerably higher sliding speed of 100 mm/s. The second sequence in this subsection is almost identical, except that only the first draw is applied to each sample, and that four times longer friction pads are used (testing direction l× width w=140 mm×35 mm ). Fig. 16 shows the friction coefficient results for the single draw of the test series using the standard 140 mm×35 mm friction pads, as well as the results for each draw of the experimental sequence using 35 mm×35 mm friction pads, which are four times shorter. Consistent with the results presented in the initial section of this paper investigating the influence of pad geometries, the friction coefficient curve for the first draw with the shorter friction pads is significantly higher than that for the 4x longer friction pads. Furthermore, as in the previously presented experimental sequence D with three draws, the friction coefficient curve for the shorter pads drops with every additional draw in decreasing step sizes. The friction coefficient curve resulting from averaging the curves from the four draws with the short pads corresponds well to the curve from the long pads. This allows the conclusion that the test with 4 x longer pads corresponds to a test with a train of four coupled short pads where the friction coefficient contact zone decreases in sliding direction (highest friction coefficient at pad entry point, lowest friction coefficient at pad exit point).

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Fig. 16. Friction tests with AA 5182-O EDT and lubricant A1200 - Friction coefficient results for standard test series using long friction pads (only 1st  draw) and test series with multiple draws using 4 times shorter friction pads (4 draws in total).

Fig. 16. Friction tests with AA 5182-O EDT and lubricant A1200 - Friction coefficient results for standard test series using long friction pads (only 1 st draw) and test series with multiple draws using 4 times shorter friction pads (4 draws in total).

Conclusion

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For boundary-lubricated automotive aluminum sheet, this paper demonstrates in detail the following points:

The geometry of the pads is a key factor in determining the friction coefficient. Pads of the same length but different widths produce a comparable friction coefficient. However, changing the length produces completely different results.

Third body particle mass collected increases with pressure. The particles are flake-like consisting predominantly of aluminum and aluminum oxide. Confocal microscopy can be used to quantify the amount of abrasion and the flattening of the surface texture.

The tribolayer induces a stronger reduction of friction coefficient than surface flattening and real area of contact. Cleaning and relubrication results in modest changes to the friction coefficient, which remains close to the initial value. Whereas, preserving the tribolayer results in reductions.

The sliding history is of great importance. When the test is repeated with the tribolayer preserved, the friction coefficient can decrease by up to 40−50%. Altering the sliding history so that the initial draw is at 7 MPa rather than at the same pressure results in the opposite trend of a standard pressure-dependent friction curve.

The friction coefficient can be normalized by the cumulative sliding length. Four passes with a 35×35 mm2 friction pad are equivalent to a single pass with a 140×35 mm2 pad (considering measurement uncertainty). It supports the use of effective sliding distance approach.
These points and their implications raise discussion points for characterization and simulation: (i) Flat strip-draw tests show that, for automotive aluminum, the measured friction coefficient varies with pad length. Therefore, caution should be exercised when using such measurements directly to determine friction parameters in modeling. (ii) These results and our additional tests involving multiple draws on the same sample at different pressures reveal that the frictional response depends
on the contact pressure history. (iii) From an applied-mechanics standpoint, the friction coefficient of a material point decreases along the contact path from entry to exit. We assume that a homogeneous friction coefficient across the pad is invalid for aluminum, and that the friction coefficient must be treated as a function of sliding distance.

Acknowledgement

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A significant portion of the results presented in this paper were obtained in connection with the InnoSuisse project 100.777 IP-ENG funded by the Swiss Federal Government in collaboration with the Computational Solid Mechanics Laboratory at the Ecole Polytechnique Fédérale de Lausanne, EPFL, Switzerland, under the direction of Prof. Jean-François Molinari. We thank Prof. Jean-François Molinari and his team members Dr. Guillaume Anciaux, Dr. Parissa Alavi, and MSc. Colin Majoor for the very productive technical discussions and constructive feedback on the experimental work, data analysis, and interpretation. In particular, we would like to express our gratitude for the opportunity to present results of Colin Majoor's master's thesis in this publication [15]. Any errors remaining in this publication are the responsibility of the authors. The authors are solely responsible for the content, which does not reflect the views of EPFL or the InnoSuisse organization.

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The original version of this paper is available on https://www.scientific.net/MSF.1184.97.pdf

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D. Waanders, J. H. Marangalou, M. Kott, S. Gastebois, and J. Hol, Temperature Dependent Friction Modelling: The Influence of Temperature on Product Quality, Procedia Manuf., vol. 47, 2020, DOI 10.1016/j.promfg.2020.04.159.

Filzek, J., Keil, D., & Schröder, H. (2021). Temperature induced friction increase in friction test and forming demonstrator for sheet metal forming. ESAFORM 2021 Proceedings. DOI: 10.25518/esaform21.3732.

F. Jalali Moghadas, M. de Rooij, T. van den Boogaard, J. Hazrati, Effect of Normal Load and Bulk Strain on Real Area of Contact in Aluminum Sheet Forming.

J. Hol. "Multi-scale friction modeling for sheet metal forming". PhD thesis. Enschede, The Netherlands: University of Twente, 2013.

P. Alavi, G. Anciaux, J.F. Molinari, L. Rocchi, C. Leppin. A multiscale model of friction considering the influence of third-body wear particles. arXiv, 2510.13470. DOI: 10.48550/arXiv. 2510.13470 [].

Bowden, F.P. & Tabor, D. The Friction and Lubrication of Solids. Oxford University Press, reprint ed. 2001.

VDA 230-213 (2022). Testing procedure for product classes Prelube, Prelube 2, Hotmelt, Spot lubricant. Verband der Automobilindustrie e.V. (VDA).

J. Cillaurren, L. Galdos, Ma. Sanchez, A. Zabala, E. Saenz de Argandoña J. Mendiguren. Contact pressure and sliding velocity ranges in sheet metal forming simulations.

C. Majoor, Image correlation and data analysis applied to surface texture microscopy, master thesis, Ecole Polytechnique Fédérale Lausanne, 2023.

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Materials Science Forum, ISSN: 1662-9752, Vol. 1184, pp 113-121

doi: 10.4028/p-3JRLbf

© 2026 The Author(s). Published by Trans Tech Publications Ltd, Switzerland.

Submitted:
2026-01-24

Accepted:
2026-02-23

Online:
2026-04-14

Towards Multi-Scale Friction Modelling for Bulk Sheet Metal Forming Applications

The original version of this paper is available on https://www.scientific.net/MSF.1184.113.pdf

Aratz Barandiaran 1,a, Alaitz Zabala 1, b, David Abedul 1,c, Javad Hazrati 2, d and Lander Galdos 1,e*

¹Material and industrial production department, Mondragon University, Loramendi 4, 20500

Arrasate-Mondragon, Spain

2 Nonlinear Solid Mechanics, Faculty of Engineering Technology, University of Twente, Enschede, Netherlands

a  aratz.barandiaran@alumni.mondragon.edu, b  azabalae@mondragon.edu,

c  dabedul@mondragon.edu, dj.hazratimarangalou@utwente.nl, e  Igaldos@mondragon.edu

(*corresponding author)

Keywords: Sheet metal forming, tribology, friction modelling

Abstract

Cold forging and sheet-metal bulk forming operations typically involve severe deformation, high contact pressures, and substantial surface enlargement. As highlighted in previous studies, friction behavior under these extreme conditions is governed by temperature, contact pressure, sliding velocity, and changes in the real contact area due to surface expansion. This work presents a newly developed linear sliding tribotester designed to characterize the friction response of metal sheets subjected to sheet-metal bulk forming conditions. The testing procedure consists of two stages. In the first stage, the sample is compressed to intentionally modify and enlarge the initial contact surface, with the degree of surface expansion controlled by the specimen geometry. In the second stage, once the surface has been altered, frictional contact is generated between the sample and a sliding table, enabling the measurement of normal and tangential forces. These force measurements are subsequently used to determine the mean coefficient of friction. The results obtained constitute the first dataset toward the development of a multi-scale friction model for sheet-metal bulk gear forging. This model aims to incorporate the effects of extreme contact pressures, asperity flattening, and lubricant-related hydrostatic and hydrodynamic mechanisms.

Introduction

The original version of this paper is available on https://www.scientific.net/MSF.1184.113.pdf

Computer-aided process optimization is widely implemented in the cold forging sector, where finite element (FE) simulations are routinely employed to predict die filling, identify defects such as folds, and estimate forming loads. These tools make it possible to design multi-stage forging sequences efficiently and with reduced development time. Nonetheless, the fidelity of numerical predictions strongly depends on the quality of the input definitions. Key parameters - including constitutive material data and friction coefficients - as well as numerical aspects such as mesh refinement, element selection, and the applied solution schemes, significantly influence the robustness of the results.

Historically, FE simulations for cold forging have relied on isotropic hardening laws combined with isotropic yield criteria and associated flow rules. Recent research has examined the impact of load path reversals, strain-rate sensitivity, and adiabatic temperature increase on the accuracy of numerical simulations. The introduction of kinematic hardening models has revealed notable effects on residual stresses and forming forces, especially in processes where the material undergoes flow reversal, such as forward extrusion followed by compression [11, 2, 3]. Moreover, the influence of strain rate and adiabatic heating on the overall material behavior in such simulations has been demonstrated in recent works [4].

Regarding friction modeling, most cold forging simulations still use constant friction coefficients. However, knowledge from the stamping field shows that friction is highly dependent on the specific contact conditions. Variable friction formulations have been proposed based on micro-scale
investigations [8, 10] as well as macro-scale tribological studies [ 1,17,18 ] for sheet metal stamping. There is broad consensus that friction is influenced primarily by contact pressure and sliding velocity [13]. Higher pressures tend to flatten surface asperities, modifying the contact geometry and subsequently the friction response [12]. Likewise, increasing sliding velocity typically contributes to a reduction in friction due to improved lubrication conditions. Friction laws depending on both pressure and velocity have therefore become increasingly common in sheet-metal forming simulations, supported by commercial software solutions such as Triboform. Industrial applications have shown clear improvements in prediction accuracy when these advanced models are applied [5, 15, 9, 16].

In cold forging operations, frictional behavior becomes even more complex because the processes involve extreme contact pressures, pronounced sliding, and considerable surface enlargement, all of which drastically modify interface conditions. Numerous tribometers have been developed to study these effects. The friction subgroup of the International Cold Forging Group provided an extensive review of existing tribotesters in [7], noting that many devices exhibit limited or fixed ranges of operation. To extend these capabilities, Groche et al. proposed the Sliding Compression Test (SCT), in which cylindrical specimens are first compressed with different anvil geometries to generate controlled surface expansion and are thereafter translated against a flat counterface under predefined pressures [6]. Similarly, aiming to reach higher surface enlargement and assess solid lubricant performance for cold forging, Wang et al. introduced a multi-stage forging test involving an upsetting operation followed by ironing with three metallic balls, enabling detailed investigation of lubricant behavior and adhesion phenomena [19].

This study introduces the Sheet Sliding Compression Test (SSCT), a novel tribological method developed to analyze lubricant performance under conditions that combine elevated contact pressures with significant surface expansion. Owing to its wide operational window, the SSCT is suitable for characterizing lubricants used both in cold forging and sheet-bulk metal forming processes. The apparatus integrates load sensors that continuously record normal and tangential forces during sliding, enabling direct determination of the mean coefficient of friction throughout the experiment. The SSCT methodology is validated using commercially pure aluminum sheets tested under different levels of contact pressure and surface expansion, aiming to develop a multi-scale friction model for sheet bulk metal forming in the near future.

The Sheet Sliding Compression Test (SSCT)

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Test description. The operating principle of the SSCT is depicted in Fig. 1. As shown, the procedure consists of two clearly differentiated stages. In the first stage, the specimen is vertically compressed against a flat counter-tool, generating a normal load and causing an enlargement of the contact area. The extent of this surface expansion is determined by the specimen geometry, which will be discussed later. Once the required surface enlargement and contact pressure have been achieved, the flat tool is displaced relative to the specimen. This motion enables the measurement of both normal and tangential forces, from which the mean coefficient of friction is obtained. During this second stage, the contact pressure may be kept equal to that used during the compression phase or reduced to lower values in order to examine the frictional response of the newly formed surface under different loading conditions.

[picture]

Fig.1. Principle of the new Sheet Compression Sliding Test (SCTT). Tool design with relevant features and different test stages

Fig.1. Principle of the new Sheet Compression Sliding Test (SCTT). Tool design with relevant features and different test stages

Test samples can be manufactured by wire EDM, either starting from cylindrical bars for cold forging studies or from thick sheets for applications related to bulk sheet metal forming. The surface finish of the samples can be tailored through different machining or finishing processes. As presented in the next sections, in this investigation the surface roughness was effectively controlled by the wire EDM process. However, alternative surface conditions - such as those produced by milling, grinding or shearing - can also be evaluated using the new testing device. Likewise, the sliding plate, which represents the forming tool in real processes, is a flat interchangeable element that can be replaced to analyze different materials or surface textures.

The sliding velocity is controlled through a servo-motor system, allowing the test to cover a wide range of kinematic conditions. The main purpose of the new tribotester is to support the development of multi-scale friction models that incorporate the influence of contact pressure, sliding speed, and the evolution or flattening of surface roughness. To construct such models, comprehensive experimental datasets obtained with this device are required for accurate calibration.

Friction coefficient evaluation method. Fig. 2 presents an example of the force signals obtained during one of the tribological tests, clearly illustrating the two distinct steps of the test. In the first step, the compression phase, the sample holder moves downward, causing the normal force to increase until it reaches a predefined value. Once this value is achieved, the vertical movement is stopped to allow the normal force to stabilize before the sliding phase begins. During the sliding phase, the sample holder remains stationary while the servomotor is activated to move the sliding plate at a constant velocity. Throughout this phase, both the friction force and the normal force are measured. As shown in the figure, both forces are displacement-dependent, consistent with observations reported in [6].

The friction coefficient can be calculated during the sliding phase as the ratio between the tangential force and the normal force, and it varies over time. To facilitate the comparison of different tribological conditions, the mean friction coefficient is calculated as the average value of the friction coefficient within a defined evaluation area, using Eq. 1:

μ¯=∑i=1nμi/n(1)

where μ¯ is the mean friction coefficient, μi is the instantaneous friction value and n is the total amount of friction coefficients measured in the evaluation area, which is selected by removing the acceleration and breaking instants of the servomotor (see Fig. 2 right).

[picture]

Fig. 2. Different steps of the SCST together with the measured force signals and definition of the evaluation area to calculate the mean friction value.

Fig. 2. Different steps of the SCST together with the measured force signals and definition of the evaluation area to calculate the mean friction value.

Experimental Methodology

The original version of this paper is available on https://www.scientific.net/MSF.1184.113.pdf

Tested material and samples dimensions. The experimental campaign was carried out using AA1050 aluminum sheets with a nominal thickness of 3 mm . The flow behavior of the material at room temperature was characterized through uniaxial tensile tests conducted on an Instron universal testing machine. The apparatus was equipped with a ZwickRoell makroXtens II contact extensometer, and all tests were performed in accordance with the ASTM E08 standard. The material exhibited very limited strain hardening and reached its saturation flow stress rapidly. This response can be accurately described using the Voce hardening law (Eq. 2), expressed as:

σ=117.7·[1−0.1475·e(−1238.9·εp)](2)

Two different specimen geometries were employed in the present investigation. The first configuration consisted of a small sample with a contact surface of 5×3 mm2, designed to assess contact conditions involving high normal pressures without inducing surface expansion (see Fig. 3 left). The 5 mm contact length ensured a uniform pressure distribution, avoiding the non-uniform contact stresses typically observed in specimens with a 30 mm length. To suppress any possibility of surface enlargement during compression, the free distance between the specimen holder and the sliding plate was restricted to 0.5 mm .

The second configuration consisted of a larger specimen with an initial contact area of 30×3 mm2, intended to reproduce conditions in which significant surface expansion occurs. To prevent buckling or cracking of the material near the clamping region, the tool edges were rounded using a 1 mm radius fillet. In this setup, additional material was necessary to accommodate the expected expansion of the contact area; therefore, a free length of 5 mm was maintained between the sample holder and the sliding plate, as illustrated in Fig. 3 right.

[picture]

Fig. 3. Samples dimensions used in the study.

Fig. 3. Samples dimensions used in the study.

The specimens were manufactured using wire Electrical Discharge Machining (EDM), selecting cutting conditions that produced the targeted initial surface roughness. This fabrication technique was chosen because its surface texture closely resembles that obtained through Electro Discharge Texturing (EDT), which is relevant for studying how the surface roughness evolves both under straining and in cases without straining caused by surface expansion. Surface roughness characterization was performed on three independent samples using a Sensofar S-NEOX non-contact 3D optical profiler operating in interferometry mode. The measurements confirmed that the resulting surface conditions were highly consistent among the samples. The arithmetical mean height (Sa) averaged 6.62μ m with a standard deviation of 0.77μ m; the developed interfacial area ratio (Sdr) was 51.47% with a standard deviation of 6.53%; and the dale void volume at p=80% (Vvv) reached 0.862 with a standard deviation of 0.084 . A representative example of the initial surface topography is presented in Fig. 4 left.

Tooling and lubricant. The counterpart material used in this study was the 1.2379 tool steel. After milling the steel to the specified plate dimensions, it was quenched and tempered to achieve a final hardness of 60 HRc . Prior to testing, both sides of the plate were grinded to ensure the required flatness and to obtain the desired surface finish. The grinding was performed parallel to the sliding direction to facilitate optimal material flow.

Surface finish measurements, conducted using the same non-contact 3D optical profiler as the one used for the sheets, are presented in Fig. 4 right. To assess the homogeneity of the surface finish, three measurements were taken at different locations on the tool. The arithmetical mean height (Sa) was found to be 0.15μ m with a standard deviation of 0.01μ m, the developed interfacial area ratio (Sdr) was 0.35% with a standard deviation of 0.01%, and the dale void volume at p=80%(Vvv) was 0.03 with a standard deviation of 0.001

[picture]

Fig. 4. Sample (left) and tool (right) roughness

Fig. 4. Sample (left) and tool (right) roughness

The lubricant utilized in the experiments was Renoform 5200/100, provided by Fuchs. This lubricant is specifically formulated for high contact pressure applications, including fine blanking and other demanding forming processes. It possesses a kinematic viscosity of 110cSt(mm2/s). To ensure complete saturation of oil pockets prior to testing, the tests were conducted under lubricated conditions with an excess of lubricant.

Testing conditions. Friction tests using the 5×3 mm2 sample (nearly with no surface expansion) and the 30×3 mm2 sample (significant surface expansion) were performed using different contact pressures. Expansion of the contact surface was observed at contact pressures bigger than 100 MPa (approximately yield limit of material). The expanded surface of the 5×3 mm2 sample for 200 MPa nominal contact pressure and the 30×3 mm2 sample for 2.5 and 5.0 tons of force are shown in Fig. 5.

[picture]

Fig. 5. Expanded contact surface for the different samples and applied forces.

Fig. 5. Expanded contact surface for the different samples and applied forces.

When the 5×3 mm2 sample was tested, the normal force corresponding to nominal contact pressures of 5,10,15,20,25,50,100,200MPa was applied prior to sliding. After the tests, the mean friction coefficient was obtained as explained in chapter Test description. The mean contact pressure of the test was calculated similarly using Eq. 3.

P¯=∑i=1nFN−i/nA(3)

where P¯ is the mean contact pressure, FN−i is the instantaneous normal force, n is the total amount of measured points in the evaluation area and A is the real contact area that was measured after every test using a Leica optical macroscope.

In contrast, a different methodology was employed when testing the 30×3 mm2 sample. Initially, the sample was compressed using two different normal forces ( 2.5 and 5.0 tons), and the resulting contact surface area was measured using a Leica optical macroscope. Based on the measured contact area, the testing forces were calculated to achieve normal contact pressures of 5,10,15,20,25,50, 100 , and 200 MPa on the newly formed surfaces. The testing procedure was then as follows: First, the surface was expanded using 2.5 and 5.0 ton force. Subsequently, the normal force was reduced to the value corresponding to the desired contact pressure, and once stabilized, the sliding test was conducted. Given that the actual normal force changes during the test, the mean contact pressure was calculated post-test using Eq. 3.

It is important to note that neither the 5×3 mm2 sample nor the 30×3 mm2 sample achieved a contact pressure of 200 MPa , as the material underwent expansion and yielding after 100 MPa .

Results

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Surface topographies. Surface topographies were measured after the test using a non-contact 3D optical profiler (Sensofar S-NEOX, interferometry technique) with an objective of 20xDI. An acquisition area of 877μ m×660μ m was set and measurements were post processed through SensoMap Premium 7 metrology software.

As illustrated in Fig. 6, the surface topography of the 5×3 mm2 sample undergoes substantial changes with increasing normal pressure. Surface roughness and void volume decrease as asperities are flattened and material from the valleys is displaced upward. A notable flattening of the surface is observed between contact pressures of 100 and 200 MPa . This pronounced flattening occurs due to the expansion of the surface, which is attributed to the material reaching its yielding point, approximately 100 MPa .

[picture]

Fig. 6. Surface topographies of sample 5×3 mm2.

Fig. 6. Surface topographies of sample 5 × 3 m m 2 .

Fig. 7 presents the surface topographies of the 30×3 mm2 sample using normal forces of 2.5 and 5.0 tons. As previously discussed, the contact surface undergoes significant expansion due to the plastic deformation of the sample, leading to substantial flattening of the initial surface roughness. This behavior is consistent with the observations made for the 5×3 mm2 sample, where plastic deformation begins at approximately 100 MPa , resulting in more pronounced surface flattening beyond that caused by compression alone. Normal contact pressures shown in Fig. 7 were calculated using the equation 3 .

[picture]

Fig. 7. Surface topographies of sample 30×3 mm2 using 2.5 and 5.0 tons of normal force.

Fig. 7. Surface topographies of sample 30 × 3 m m 2 using 2.5 and 5.0 tons of normal force.

Optical macroscope measurements confirmed that the 5×3 mm2 samples exhibited expansion only after reaching a normal contact pressure of 100 MPa . A surface expansion (SE =Af/A0·100 ) of 133% was observed at a nominal contact pressure of 200 MPa , corresponding to a normal contact pressure of 140.7 MPa .

For the 30×3 mm2 sample, a surface expansion of 187% was measured under a normal force of 2.5 tons. Further expansion was achieved with a 5.0 -ton force, resulting in a surface expansion of 269%.

Friction coefficients. Mean friction coefficients for the different tested conditions and mean contact pressures are shown in Fig. 8 together with the standard deviation obtained from 3 different sheet compression sliding tests.

Results show that friction coefficient significantly depends on contact pressure and the initial surface condition (expansion ratio of the surface).

[picture]

Fig. 8. Mean friction coefficients for the different conditions and mean contact pressures

Fig. 8. Mean friction coefficients for the different conditions and mean contact pressures

Conclusions and Outlook

The original version of this paper is available on https://www.scientific.net/MSF.1184.113.pdf

The newly developed Sheet Sliding Compression Test shows significant promise for the investigation of lubricants and friction conditions in cold forging and sheet metal bulk forming. The novel sample preparation methodology allows to produce various surface finishes, ranging from rough surfaces that replicate EDT (electro-discharge texturing) finishing to others such as grinding or shearing. The optimized sample designs provide control over the surface expansion of the contact area, enabling tests at different levels of surface expansion.

Friction tests conducted on an EDT-finished AA1050 aluminum sheet confirm that asperity flattening is influenced not only by normal contact pressure but also by surface expansion. This latter parameter significantly reduces surface roughness as asperities are flattened through straining, and because the combined effects of compression and expansion induce plastic deformation under complex stress conditions. This effect has been validated through surface topography measurements obtained using a confocal profilometer.

Regarding friction coefficients, it is evident that they decrease with increasing contact pressure. This phenomenon is well-documented in the literature and is attributed to the decreasing attack angle of contact patches [14] and the activation of the micro Plasto Hydrodynamic Lubrication regime under high contact pressure and sliding conditions [12].

Notably, extreme contact pressures combined with surface expansion led to significant asperity flattening and a further reduction in friction. This is particularly pronounced in friction tests conducted after a 2.5 -ton compression force, which resulted in a surface expansion of 187%. In this scenario, the contact angle is likely significantly reduced, and a sufficient oil film remains on the contact surface, causing the friction coefficient to decrease to as low as 0.02 . However, in tests conducted after applying a 5.0 -ton force and achieving a surface expansion of 269%, friction coefficients were observed to be higher than in the previous condition. This increase in friction may be attributed to galling observed in some tests, likely due to oil being squeezed out under such high contact pressures, leaving insufficient lubricant to cover the significantly expanded surface.

Future research will involve the development of pressure and surface expansion dependent friction model using experimental data. This model will be applied to simulate a gear bulk forming process initiated from a sheet.

References

The original version of this paper is available on https://www.scientific.net/MSF.1184.113.pdf

Filzek, J., Ludwig, M., & Groche, P. (2011). Improved FEM simulation of sheet metal forming with friction modelling using laboratory tests. Proceedings of the IDDRG, Bilbao, Spain, 5-8.

Galdos, L., Agirre, J., Mendiguren, J., Argandona, E. D., Otegi, N., & Trinidad, J. (2019). Mixed isotropic-kinematic hardening model for cold forging simulation of an industrial bolt. In Proceedings of the 52nd ICFG plenary meeting.

Galdos, L., Agirre, J., Otegi, N., Mendiguren, J., & de Argandoña, E. S. (2021). Simulation of Cold Forging Processes Using a Mixed Isotropic-Kinematik Hardening Model. In Forming the Future: Proceedings of the 13th International Conference on the Technology of Plasticity (pp. 773-787). Springer International Publishing.

Galdos, L., Agirre, J., Ziarsolo, U., & Atxega, M., (2023). 42CrMo4 material characterization for cold forging simulation -Comparison of different rheological laws. In Proceedings of the 53rd ICFG plenary meeting.

Gil, I., Mendiguren, J., Galdos, L., Mugarra, E., & de Argandoña, E. S. (2016). Influence of the pressure dependent coefficient of friction on deep drawing springback predictions. Tribology International, 103, 266-273.

Groche, P., Stahlmann, J., & Müller, C. (2013). Mechanical conditions in bulk metal forming tribometers-Part two. Tribology International, 66, 345-351.

Groche, P., Kramer, P., Bay, N., Christiansen, P., Dubar, L., Hayakawa, K., ... & Moreau, P. (2018). Friction coefficients in cold forging: A global perspective. CIRP Annals, 67(1), 261264.

Hol, J., Alfaro, M. C., de Rooij, M. B., & Meinders, T. (2012). Advanced friction modeling for sheet metal forming. Wear, 286, 66-78.

Hol, J., Wiebenga, J. H., & Carleer, B. (2017, September). Friction and lubrication modelling in sheet metal forming: Influence of lubrication amount, tool roughness and sheet coating on product quality, IOP Conf. In Series: Journal of Physics: Conf. Series (Vol. 896).

Karupannasamy, D. K., Hol, J., de Rooij, M. B., Meinders, T., & Schipper, D. J. (2014). A friction model for loading and reloading effects in deep drawing processes. Wear, 318(1-2), 2739.

Kolpak, F., Hering, O., & Tekkaya, A. E. (2021). Consequences of large strain anisotropic work-hardening in cold forging. International Journal of Material Forming, 14, 1463-1481.

Ma, X., De Rooij, M., & Schipper, D. (2010). A load dependent friction model for fully plastic contact conditions. Wear, 269(11-12), 790-796.

Merklein, M., Zöller, F., & Sturm, V. (2014). Experimental and numerical investigations on frictional behaviour under consideration of varying tribological conditions. In Advanced Materials Research (Vol. 966, pp. 270-278). Trans Tech Publi-cations Ltd.

Mizuno, T., & Okamoto, M. (1982). Effects of lubricant viscosity at pressure and sliding velocity on lubricating conditions in the compression-friction test on sheet metals.

Sigvant, M., Pilthammar, J., Hol, J., Wiebenga, J. H., Chezan, T., Carleer, B., & van den Boogaard, A. H. (2016, No-vember). Friction and lubrication modeling in sheet metal forming simulations of a Volvo XC90 inner door. In IOP Conference Series: Materials Science and Engineering (Vol. 159, No. 1, p. 012021). IOP Publishing.

Sigvant, M., Pilthammar, J., Hol, J., Wiebenga, J. H., Chezan, T., Carleer, B., & van den Boogaard, T. (2019). Friction in sheet metal forming: Influence of surface roughness and strain rate on sheet metal forming simulation results. Procedia Manufacturing, 29, 512-519.

Tamai, Y., Inazumi, T., & Manabe, K. I. (2016). FE forming analysis with nonlinear friction coefficient model con-sidering contact pressure, sliding velocity and sliding length. Journal of Materials Processing Technology, 227, 161-168.

Trzepiecinski, T. (2019). A study of the coefficient of friction in steel sheets forming. Metals, 9(9), 988.

Wang, Z. G., Komiyama, S., Yoshikawa, Y., Suzuki, T., & Osakada, K. (2015). Evaluation of lubricants without zinc phosphate precoat in multi-stage cold forging. Cirp Annals, 64(1), 285288.

This ePub file was created using machine-generated methods and automated software to convert the source text into an ePub format and may contain formatting inconsistencies from the original version. Formatting may vary depending on the device used.

The original version of this eBook is available on https://www.scientific.net/book/friction-and-wear-in-metal-forming/978-3-0364-1992-3

Materials Science Forum, ISSN: 1662-9752, Vol. 1184, pp 123-130

doi: 10.4028/p-0rBeB4

© 2026 The Author(s). Published by Trans Tech Publications Ltd, Switzerland.

Submitted:
2026-01-29

Revised:
2026-02-17

Accepted:
2026-02-23

Online:
2026-04-14

Effect of Surface Texture on Sheet Metal Formability in Boundary Lubrication Regime

The original version of this paper is available on https://www.scientific.net/MSF.1184.123.pdf

Javad Hazrati 1,a*, Matthijn de Rooij 2, b

¹Nonlinear Solid Mechanics, Faculty of Engineering Technology, University of Twente, Enschede, Netherlands

2 Surface Technology and Tribology, Faculty of Engineering Technology, University of Twente, Enschede, Netherlands

aj.hazratimarangalou@utwente.nl*, b  m.b.derooij@utwente.nl 

Keywords: formability, sheet metal texture, sheet metal forming, friction

Abstract

Friction plays an important role on formability of deep drawn products. This necessitates an accurate description of friction in finite element formability analyses. It has been shown that constant coefficient of friction does not lead to precise prediction of product formability in these analyses. The multi-scale friction model developed at University of Twente takes the local contact conditions and textures of sheet metal and tools as the input at boundary and mixed lubrication regimes. To correlate the zinc coated sheet metal surface texture parameters with its formability, 60 different textures were analyzed. The multi-scale friction model is used to estimate friction for all the different surface textures. The effect of different textures on formability of the sheet metal was investigated by simulating cross-die forming using different sheet metal surface textures. The results show that different textures depict distinct formability behavior in the boundary lubrication regime (lubricant amount 0.5gr/m2 ). Exploring the correlation between areal roughness parameters of the current dataset and formability of cross-die shows that besides surface roughness, autocorrelation length and skewness of height distributions are determining parameters.

Introduction

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Friction plays a key role on formability of sheet metals. It has been shown that friction is a local phenomenon which depends on the contact conditions between tool and workpiece and varies during the forming process [1]. Contact condition is determined by tool and sheet metal surface topography, material properties of sheet metal and its coating as well as tool material behavior, lubrication mechanism, process variables such as contact pressure, temperature and strain during forming process [2,3]. Contact conditions are further affected by variations in these parameters such as variations in lubricant distribution and tool surface texture due to wear [4].

Sheet metal surface texture is an important factor in determining friction between the tool and workpiece. Shape, height and spacing of the asperities on the sheet metal surface, coating thickness, substrate and coating material behavior control deformation or flattening of the asperities due to contact pressure and strain in the bulk of the sheet metal [5, 6, 7]. Furthermore, distribution of asperities governs lubricant distribution and its load bearing capacity [8].

As the tool comes into contact with the sheet metal, the highest asperities on the sheet metal are plastically deformed due to contact pressure. The plastic deformation of asperities is governed by the mechanical behavior of sheet metal and its coating (if any), geometry of the asperities and their distribution [9]. As the asperities are indented, the non-contacting asperities or valleys rise and may come into contact with the tool. The latter depends on asperities spacing and lubricant distribution as well. Strain hardening rate of asperities is another parameter that determines the load bearing capacity of the asperities and therefore real area of contact.

It has been shown that normal load combined with bulk strain in the sheet can substantially increase real area of contact. The effect of strain, however, depends on effective hardness of the sheet metal surface (substrate and coating material and thickness) and its texture. With increasing thickness of zinc coating on a steel substrate, the effect of strain is increased and with increasing sheet metal
surface roughness, the rate of increase in real area of contact due to bulk strain is reduced. Furthermore, by increasing the mean asperity spacing, real area of contact increases with the bulk strain more rapidly [5]. In general, it can be concluded that the surface texture determines the real area of contact, shape and distribution of the contact patches between workpiece and tool. The latter would affect the size, orientation and attack angle of the tool asperities which are in contact with the sheet metal surface. These will ultimately determine ploughing behavior of the tool asperities through the sheet metal surface which influence coefficient of friction [10].

In the current study, the validated contact and friction models are used to investigate the effect of sheet metal texture parameters on the formability of a cross-die product. For this purpose, a number of surface textures resembling EDT surfaces are artificially produced using an in-house algorithm.

Sheet Metal Surface Texture

The original version of this paper is available on https://www.scientific.net/MSF.1184.123.pdf

To correlate the zinc coated sheet metal texture parameters with its formability, 55 different textures were artificially created using the bearing area method [11]. The texture is printed on a surface with an initial roughness obtained after the galvanization process. Three examples of the generated surfaces including the initial surface are shown in figure 1. Five additional textures were measured from industrial rolled products. Surface textures were generated such that different roughness parameters such as root mean square roughness, asperity spacing, summit density and auto correlation length (ACL) are varied within the industrial measured values. Figure 2 shows distribution of texture parameters for the generated surfaces. The average root mean square roughness of the all surfaces, including artificial and real ones is 1.48 (SD: ±0.49 ), Table 1 lists the mean and standard deviation of the roughness parameters for the 60 surfaces used in the current study. For all the surfaces, zinc coating thickness is constant (7μ m) and substrate is DX57 with a thickness of 0.8 mm .

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Fig. 1. Surface height of the initial untextured (left image) and three generated textures.

Fig. 1. Surface height of the initial untextured (left image) and three generated textures.

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Fig. 2. Distribution of roughness parameters of the generated surfaces (blue markers) and real surfaces (red markers) at different surface roughness levels.

Fig. 2. Distribution of roughness parameters of the generated surfaces (blue markers) and real surfaces (red markers) at different surface roughness levels.

Table 1. Average and standard deviation of surface roughness parameters ( n=60 ).

Root mean square height (μm) | Mean asperity spacing (μm) | Summit density (mm⁻²) | Skewness (-) | Kurtosis (-) | Auto-correlation length (μm)
1.48 (±0.49) | 20.7 (±2.2) | 1080 (±252) | -0.29 (±0.27) | 2.24 (±0.28) | 18.8 (±7.2)

Multi-scale Friction Model

The original version of this paper is available on https://www.scientific.net/MSF.1184.123.pdf

Accurate description of friction coefficient in finite element analyses is indispensable for the accuracy of formability predictions. In sheet metal forming, friction is a local phenomenon that is determined by local contact conditions. Tool and sheet metal surface morphology as well as their hardness levels (e.g. coating and substrate hardness), lubricant type and amount constitute the tribological system.

Process variables such as contact pressure, plastic strain, temperature in the contact and relative velocity of tool and sheet metal influence the tribological system and therefore determine frictional shear stress.

The multi-scale friction model developed at University of Twente by the author and colleagues [1,12,13] accounts for the micro and meso-scale phenomena which occur at the tool-sheet metal contact. In the boundary lubrication regime, flattening of sheet metal surface at different contact pressures and plastic strains is predicted using a semi-analytical contact model [6]. The model considers sheet metal topography (different shapes of asperities), coating thickness and its mechanical behavior. The model is validated at the meso-scale via dedicated experiments which enable the measurement of real contact area [5]. At the micro-level, shear strength of the boundary layers and ploughing of tool asperities which contribute to the friction stress are considered and validated through dedicated experiments [10, 14]. Furthermore, the effect of lubricant in the model is considered via solving averaged Reynolds equation at the contact of the tool-sheet metal [13, 15]. The model implements the topography of sheet metal and tool to calculate averaged lubricant flow. The multi-scale model was validated to industrial-scale forming process (cup drawing and cross-die forming) at boundary and mixed lubrication regimes [12, 13]. Figure 3 shows the multi-scale friction modeling approach in boundary lubrication. The input includes surface topography of sheet metal, tool and their material behavior. In the deformation step, contact models [5] are utilized to estimate the real contact area. In the sliding steps, contribution of boundary layers and ploughing of tool asperities through the sheet metal surface are calculated for individual asperities [10, 14] to estimate the total friction force. In the current study, behavior of different surface textures are investigated at boundary lubrication only.

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Fig. 3. Multi-scale modelling approach of friction in sheet metal forming.

Fig. 3. Multi-scale modelling approach of friction in sheet metal forming.

Cross-die Forming Simulations

The original version of this paper is available on https://www.scientific.net/MSF.1184.123.pdf

To investigate the effect of different sheet metal textures on its formability, cross-die forming simulations were carried out using the friction cards estimated for each surface in the boundary lubrication regime. One single tool surface topography in combination with sheet metal surface topographies was used to estimate friction coefficients at different contact pressures and equivalent plastic strains. Figure 4 shows the tool surface topography. Friction cards were later implemented in finite element simulations of cross-die forming using in-house FE code DiekA
(www.dieka.huetink.nl). Hydrodynamic effect of the lubricant on friction is negligible in boundary lubrication regime. Lubricant type used in the simulations was Anticorit RP4107S, with dynamic viscosity of 30 mPa. s and density of 0.896gr/cm3 [13]. The maximum forming depth is 60 mm . The tools are modelled as rigid contours. For the material properties, the yield surface of DX57 sheet metal is described by Vegter's yield criterion and the Bergström-Van Liempt hardening rule is used to describe the sheet material hardening [16]. Due to symmetry in geometry a quarter of the part is modelled.

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Fig. 4. Tool surface topography used to estimate friction coefficients for different sheet metal textures in cross-die forming simulations

Fig. 4. Tool surface topography used to estimate friction coefficients for different sheet metal textures in cross-die forming simulations

The results show that different textures depict distinct formability behavior despite having similar roughness. Major-minor strain distribution in 26 out of 60 textures go beyond the forming limit curve (FLC) at the lubricant amount of 0.5gr/m2 where boundary lubrication prevails. Four typical textures with comparable root mean square roughness and height distributions selected from the investigated textures are shown in figure 5. Surfaces 5, 15 and surfaces 8, 39 have similar roughness values and height distributions, however different areal roughness parameters such as summit density and autocorrelation length. Forming limit diagrams (FLD) and coefficients of friction (COF) distributions for each finite element at drawing depth of 60 mm depict how surfaces with comparable height distributions but different texture would affect the formability of the sheet metal.

Figure 6a shows two additional surfaces with similar height distributions and root mean square roughness values of 1.02 and 1.08μ m for surfaces 48 and 49, respectively. Roughness parameters of the two surfaces are listed in table 2. Figure 6b depicts the FLDs obtained from cross-die forming simulations, the results show that sheet metal using texture 48 goes beyond the FLC for DX57-Z due to larger friction coefficients. Figure 6 additionally shows the distribution of friction coefficients at each finite element for the two surfaces at the drawing depth of 60 mm . Surface 48 clearly uses higher coefficient of friction values for the given contact conditions: contact pressure and equivalent plastic strain. The variation in coefficient of friction at each increment leads to variations in contact pressure. Therefore surface 48 results in slightly lower contact pressures and higher equivalent plastic strains and consequently higher thinning in the blank.

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Fig. 5. Four surface textures (5, 15, 8 and 39) with similar height distributions and their FLDs obtained from cross-die forming simulations. The bottom row shows the areal roughness parameters for the selected surfaces.

Fig. 5. Four surface textures (5, 15, 8 and 39) with similar height distributions and their FLDs obtained from cross-die forming simulations. The bottom row shows the areal roughness parameters for the selected surfaces.

To study the effect of surface texture on the formability of the sheet metal, friction behavior of the surface at different contact conditions were analyzed using the multi-scale friction model. Figure 6c shows the ploughing force at each tool asperity in contact with the sheet metal (contact patch) as the sheet metal is subjected to 20% plastic strain and contact pressure of 20 MPa . The asperities are arranged in descending order based on the quantities. The results show surface 48 leads to larger number of individual contact patches ( 29 contact patches or tool asperities) due to lower autocorrelation length, asperity spacing and higher summit density compared to 21 contact patches in surface 49 . Checking the individual tool asperities in contact with surface 48 reveals that asperities are more elongated than the tool asperities in contact with surface 49 . The latter is evaluated by fitting elliptical paraboloids to the tool asperities penetrated to the sheet metal surface. It has been shown that ellipticity ratio would largely influence the load bearing capacity of the asperity while ploughing on the sheet metal surface [10]. However, the latter further depends on the orientation of asperity during ploughing. The fitted ellipses to the base of contacting asperities have larger major to minor radii ratio in surface 48 compared to 49 . More importantly, calculating the attack angles of individual asperities explains the higher ploughing forces and therefore higher friction coefficients of surface 48. Tool asperities in contact with surface 48 have higher attack angles relative to surface 49 . Based
on friction regimes diagram developed by Challen-Oxley [17], higher attack angle results in higher coefficient of friction in the ploughing regime.

Table 2. Roughness parameters of the surfaces 48 and 49.

Surfaces | Root mean square height (μm) | Mean asperity spacing (μm) | Summit density (mm⁻²) | Skewness (-) | Kurtosis (-) | Auto-correlation length (μm)
48 | 1.02 | 19.77 | 1142 | -0.36 | 2.40 | 19.141
49 | 1.08 | 21.01 | 901 | -0.42 | 2.46 | 24.134

Fig. 6. (a) surface textures 48 & 49 height distributions, (b) FLDs and distribution of friction coefficients, contact pressure, plastic strain at each element in 60 mm drawing depth, (c) analysis of tool asperities in contact with sheet surfaces at contact pressure of 20 MPa and equivalent plastic strain of 20%.

Exploring the correlation between areal field parameters and formability of cross-die; surface roughness, summit density, autocorrelation length and skewness of height distributions are determining parameters (figure 7). However, autocorrelation length and summit density can be used interchangeably. The results show that for the current dataset, higher autocorrelation length in combination with negative skewness and higher roughness lead to better formability of the sheet metal. In figure 7, black and blue filled circles represent the textures which do not go beyond the FLC (below maximum major strain in plane strain state: 0.381 ). Although the plot does not provide a crisp separation on the formability of different surface textures, however, provides a strong indication about the formability of the textures. It should be noted that the root mean square roughness ( Sq ), autocorrelation length (ACL) and skewness values for different textures were normalized by the absolute maximum value in the dataset for each parameter.

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Fig.7. Correlation between surface roughness parameters and maximum major strain at plane strain in cross-die forming simulations. The parameters were normalized by the maximum absolute value from the dataset for each parameter.

Fig.7. Correlation between surface roughness parameters and maximum major strain at plane strain in cross-die forming simulations. The parameters were normalized by the maximum absolute value from the dataset for each parameter.

Conclusions

The original version of this paper is available on https://www.scientific.net/MSF.1184.123.pdf

In the current study, the validated multi-scale friction model in combination with a surface generating algorithm are used to determine the effect of surface texture or different roughness parameters on formability of sheet metal. The results show that apart from sheet metal roughness, autocorrelation length and skewness values play important roles on formability of sheet metal. The framework developed in the current study can be used to optimize sheet metal texture to control friction such that it is insensitive to variations in tool roughness and lubricant amount.

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The original version of this paper is available on https://www.scientific.net/MSF.1184.123.pdf

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Keywords Index

A

Aisi 420 Improvement of Wear and Corrosion Resistance of X46Cr13 Martensitic Stainless Steel by Cryogenic Treatment

Aluminum Modeling the Effect of Lubricants on Surface Conditions in Plane-Strain Upsetting Tests

Asperity Flattening Multiscale Simulation of Asperity Flattening with Realistic Surface Topography and Microstructure

Automotive Aluminum Sheet Sliding Distance Dependency and Third Body Particle Influence in Flat Strip-Draw Testing of Aluminum Sheet for Friction Characterization in Automotive Stamping

B

Battery Production Tool Wear Investigation Nickel Plated Steels for Battery Shell Production

C

Carbon Dioxide Assessment of the Potential of CO2 as a Lubricant in Cold Forging of Low-Carbon Mild Steel

Cold Forging Assessment of the Potential of CO2 as a Lubricant in Cold Forging of Low-Carbon Mild Steel

Coupled Eulerian-lagrangian Modeling the Effect of Lubricants on Surface Conditions in Plane-Strain Upsetting Tests

Cryogenic Treatment Improvement of Wear and Corrosion Resistance of X46Cr13 Martensitic Stainless Steel by Cryogenic Treatment

Crystal Plasticity Multiscale Simulation of Asperity Flattening with Realistic Surface Topography and Microstructure

D

Dry Sheet Metal Forming Selection and Investigation of Thermoplastic Materials as Forming Tools for Warm Dry Forming of Magnesium and High-Strength Aluminum Sheets

E

Electrochemical Corrosion Improvement of Wear and Corrosion Resistance of X46Cr13 Martensitic Stainless Steel by Cryogenic Treatment

Ex-situ Method A New Ex-Situ Method for Real Contact Area Determination for Sheet Metal Forming

Eye-glass Frame Parts Gallig-Free Dry Net Shaping of Titanium and Titanium Alloy Eye-Glass Frame Parts toward Green Manufacturing

F

Fem Modeling the Effect of Lubricants on Surface Conditions in Plane-Strain Upsetting Tests

Finite Element Model Numerical Analysis of Roughness Transfer Mechanism during Skin-Pass Rolling

Flat Strip-draw Test Sliding Distance Dependency and Third Body Particle Influence in Flat Strip-Draw Testing of Aluminum Sheet for Friction Characterization in Automotive Stamping

Formability Effect of Surface Texture on Sheet Metal Formability in Boundary Lubrication Regime

Friction Effect of Surface Texture on Sheet Metal Formability in Boundary Lubrication Regime

Friction Characterisation Assessment of the Potential of CO2 as a Lubricant in Cold Forging of Low-Carbon Mild Steel

Friction Modelling Towards Multi-Scale Friction Modelling for Bulk Sheet Metal Forming Applications

G

Galling-free Shaping Gallig-Free Dry Net Shaping of Titanium and Titanium Alloy Eye-Glass Frame Parts toward Green Manufacturing

H

High-strength Aluminum Alloy Selection and Investigation of Thermoplastic Materials as Forming Tools for Warm Dry Forming of Magnesium and High-Strength Aluminum Sheets

M

Martensitic Stainless Steel Improvement of Wear and Corrosion Resistance of X46Cr13 Martensitic Stainless Steel by Cryogenic Treatment

Miniature Spike Test Assessment of the Potential of CO2 as a Lubricant in Cold Forging of Low-Carbon Mild Steel

Multiscale Simulation Multiscale Simulation of Asperity Flattening with Realistic Surface Topography and Microstructure

N

Nickel Plated Steel Tool Wear Investigation Nickel Plated Steels for Battery Shell Production

Normal Loading A New Ex-Situ Method for Real Contact Area Determination for Sheet Metal Forming

R

Real Contact Area A New Ex-Situ Method for Real Contact Area Determination for Sheet Metal Forming

Roughness Modeling the Effect of Lubricants on Surface Conditions in Plane-Strain Upsetting Tests

Roughness Transfer Numerical Analysis of Roughness Transfer Mechanism during Skin-Pass Rolling

S

Saving Of Wastes Gallig-Free Dry Net Shaping of Titanium and Titanium Alloy Eye-Glass Frame Parts toward Green Manufacturing

Sheet Metal Forming Towards Multi-Scale Friction Modelling for Bulk Sheet Metal Forming Applications; Effect of Surface Texture on Sheet Metal Formability in Boundary Lubrication Regime

Sheet Metal Texture Effect of Surface Texture on Sheet Metal Formability in Boundary Lubrication Regime

Skin-pass Rolling Numerical Analysis of Roughness Transfer Mechanism during Skin-Pass Rolling; Modeling the Effect of Lubricants on Surface Conditions in Plane-Strain Upsetting Tests

Sliding Distance Sliding Distance Dependency and Third Body Particle Influence in Flat Strip-Draw Testing of Aluminum Sheet for Friction Characterization in Automotive Stamping

Surface-topography A New Ex-Situ Method for Real Contact Area Determination for Sheet Metal Forming

T

Thermoplastic Forming Tool Selection and Investigation of Thermoplastic Materials as Forming Tools for Warm Dry Forming of Magnesium and High-Strength Aluminum Sheets

Third-body Particles Sliding Distance Dependency and Third Body Particle Influence in Flat Strip-Draw Testing of Aluminum Sheet for Friction Characterization in Automotive Stamping

Titanium Gallig-Free Dry Net Shaping of Titanium and Titanium Alloy Eye-Glass Frame Parts toward Green Manufacturing

Tribology Tool Wear Investigation Nickel Plated Steels for Battery Shell Production; A New Ex-Situ Method for Real Contact Area Determination for Sheet Metal Forming; Towards Multi-Scale Friction Modelling for Bulk Sheet Metal Forming Applications

W

Wear Tool Wear Investigation Nickel Plated Steels for Battery Shell Production; Improvement of Wear and Corrosion Resistance of X46Cr13 Martensitic Stainless Steel by Cryogenic Treatment

Authors Index

A

Abedul, D. Towards Multi-Scale Friction Modelling for Bulk Sheet Metal Forming Applications

Aginagalde, A. A New Ex-Situ Method for Real Contact Area Determination for Sheet Metal Forming

Aizawa, T. Gallig-Free Dry Net Shaping of Titanium and Titanium Alloy Eye-Glass Frame Parts toward Green Manufacturing

Alba, D. Assessment of the Potential of CO2 as a Lubricant in Cold Forging of Low-Carbon Mild Steel

Appelman, M. Tool Wear Investigation Nickel Plated Steels for Battery Shell Production

Argandoña, E. A New Ex-Situ Method for Real Contact Area Determination for Sheet Metal Forming

B

Barandiaran, A. Towards Multi-Scale Friction Modelling for Bulk Sheet Metal Forming Applications

Blunt, L. A New Ex-Situ Method for Real Contact Area Determination for Sheet Metal Forming

Boogaard, T. Numerical Analysis of Roughness Transfer Mechanism during Skin-Pass Rolling

Brüggemann, H. Modeling the Effect of Lubricants on Surface Conditions in Plane-Strain Upsetting Tests

C

Carley-Clopton, A. Multiscale Simulation of Asperity Flattening with Realistic Surface Topography and Microstructure

D

Derazkola, H. Numerical Analysis of Roughness Transfer Mechanism during Skin-Pass Rolling

Dix, M. Selection and Investigation of Thermoplastic Materials as Forming Tools for Warm Dry Forming of Magnesium and High-Strength Aluminum Sheets

E

Espinosa, I. Improvement of Wear and Corrosion Resistance of X46Cr13 Martensitic Stainless Steel by Cryogenic Treatment

F

Filzek, J. Sliding Distance Dependency and Third Body Particle Influence in Flat Strip-Draw Testing of Aluminum Sheet for Friction Characterization in Automotive Stamping

G

Galdos, L. Towards Multi-Scale Friction Modelling for Bulk Sheet Metal Forming Applications

Guinea, A. A New Ex-Situ Method for Real Contact Area Determination for Sheet Metal Forming

Gutierrez, J. Improvement of Wear and Corrosion Resistance of X46Cr13 Martensitic Stainless Steel by Cryogenic Treatment

H

Hazrati, J. Multiscale Simulation of Asperity Flattening with Realistic Surface Topography and Microstructure; Numerical Analysis of Roughness Transfer Mechanism during Skin-Pass Rolling; Towards Multi-Scale Friction Modelling for Bulk Sheet Metal Forming Applications; Effect of Surface Texture on Sheet Metal Formability in Boundary Lubrication Regime

J

Jacobs, L. Numerical Analysis of Roughness Transfer Mechanism during Skin-Pass Rolling

K

Kiakidis, S. Numerical Analysis of Roughness Transfer Mechanism during Skin-Pass Rolling

Kihara, T. Gallig-Free Dry Net Shaping of Titanium and Titanium Alloy Eye-Glass Frame Parts toward Green Manufacturing

Koch, L. Modeling the Effect of Lubricants on Surface Conditions in Plane-Strain Upsetting Tests

L

Leppin, C. Sliding Distance Dependency and Third Body Particle Influence in Flat Strip-Draw Testing of Aluminum Sheet for Friction Characterization in Automotive Stamping

Lian, J. Modeling the Effect of Lubricants on Surface Conditions in Plane-Strain Upsetting Tests

Liewald, M. Assessment of the Potential of CO2 as a Lubricant in Cold Forging of Low-Carbon Mild Steel

Llavori, I. A New Ex-Situ Method for Real Contact Area Determination for Sheet Metal Forming

Lopez, J. Improvement of Wear and Corrosion Resistance of X46Cr13 Martensitic Stainless Steel by Cryogenic Treatment

M

Menargues, S. Improvement of Wear and Corrosion Resistance of X46Cr13 Martensitic Stainless Steel by Cryogenic Treatment

Mendiguren, J. A New Ex-Situ Method for Real Contact Area Determination for Sheet Metal Forming

N

Nielsen, C. Multiscale Simulation of Asperity Flattening with Realistic Surface Topography and Microstructure

P

Pater, V. Tool Wear Investigation Nickel Plated Steels for Battery Shell Production

Picas, J. Improvement of Wear and Corrosion Resistance of X46Cr13 Martensitic Stainless Steel by Cryogenic Treatment

Puig, M. Improvement of Wear and Corrosion Resistance of X46Cr13 Martensitic Stainless Steel by Cryogenic Treatment

R

Reichardt, G. Assessment of the Potential of CO2 as a Lubricant in Cold Forging of Low-Carbon Mild Steel

Rocchi, L. Sliding Distance Dependency and Third Body Particle Influence in Flat Strip-Draw Testing of Aluminum Sheet for Friction Characterization in Automotive Stamping

Rooij, M. Effect of Surface Texture on Sheet Metal Formability in Boundary Lubrication Regime

S

Scharifi, E. Modeling the Effect of Lubricants on Surface Conditions in Plane-Strain Upsetting Tests

Shiratori, T. Gallig-Free Dry Net Shaping of Titanium and Titanium Alloy Eye-Glass Frame Parts toward Green Manufacturing

T

Tamilselvan, S. Modeling the Effect of Lubricants on Surface Conditions in Plane-Strain Upsetting Tests

Tato, W. A New Ex-Situ Method for Real Contact Area Determination for Sheet Metal Forming

V

Vakulenko, S. Selection and Investigation of Thermoplastic Materials as Forming Tools for Warm Dry Forming of Magnesium and High-Strength Aluminum Sheets

Venema, J. Tool Wear Investigation Nickel Plated Steels for Battery Shell Production

W

Weigel, K. Selection and Investigation of Thermoplastic Materials as Forming Tools for Warm Dry Forming of Magnesium and High-Strength Aluminum Sheets

Weiss, A. Assessment of the Potential of CO2 as a Lubricant in Cold Forging of Low-Carbon Mild Steel

Winther, G. Multiscale Simulation of Asperity Flattening with Realistic Surface Topography and Microstructure

Wörmann, J. Tool Wear Investigation Nickel Plated Steels for Battery Shell Production

Z

Zabala, A. A New Ex-Situ Method for Real Contact Area Determination for Sheet Metal Forming; Towards Multi-Scale Friction Modelling for Bulk Sheet Metal Forming Applications