reinforced with carbon fibres (carbon fibre fabric 3k T300 Twill 2×2 by Toray). The composite is then produced by hot curing the resin at 120 °C for 1 h 30 min and then continuing with a further cure phase at 140 °C for 2 h. The composite panels produced have dimensions of 100x100mm and a thickness of 3mm. Further information on the manufacturing of vitrimer-based composite are presented in previous works of the authors [16–18].

In order to determine the most suitable side for deposition, confocal analyses were carried out on the composite surface. This was done because each face of the surface exhibited different surface characteristics due to the contact with the Teflon mould. In fact, the surface characteristics of the substrate, could strongly influence the CS deposition, as assessed in previous studies [19].

Once the panels were formed, a metallization process was applied to them using Cold Spray (CS) technology. The chosen powders for the metallization were micrometric powders of AlSi10Mg (LPW South Group), specifically selected for their compatibility with composite materials based on previous experiments conducted by the authors [19,20]. These powders were found to be highly suitable for deposition on the panels, even at low temperatures.

The metallization process was carried out using a low-pressure cold spray machine, specifically the Dycomet 423 CS equipment. Nitrogen gas was used as the carrier gas in this process.

Based on a preliminary sensitivity analysis, not presented here for the sake of brevity, the most significant process parameters have been identified and are outlined in Table 1.

This analysis underscored the importance of temperature and Standoff Distance (SoD) in the deposition process for composite panels with a vitrimer matrix. Previous studies have confirmed that vitrimers exhibit enhanced flexibility when heated above their Tv, instead of becoming brittle [15,21]. Therefore, it is essential to employ a carrier gas at an appropriate temperature and distance from the nozzle to effectively heat the substrate and achieve the desired temperature.

Two temperature values, T₁=200 °C and T₂=300 °C, and two values of SoD, SO₁=70 mm and SO₂=100mm, have been selected, which represent the change in behaviour of the vitrimeric matrix.

<table><caption>Table 1 – Selected process parameters for CS deposition of AlSi10Mg powders on composite vitrimer-based substrates.</caption><thead><tr><th></th><th>C<sub>T1SO1</sub></th><th>C<sub>T2SO1</sub></th><th>C<sub>T1SO2</sub></th><th>C<sub>T2SO2</sub></th></tr></thead><tbody><tr><td>Inlet gas temperature [°C]</td><td>200</td><td>300</td><td>200</td><td>300</td></tr><tr><td>Gas pressure [bar]</td><td>6</td><td>6</td><td>6</td><td>6</td></tr><tr><td>SoD [mm]</td><td>70</td><td>70</td><td>100</td><td>100</td></tr><tr><td>Gun traverse speed [mm/s]</td><td>1</td><td>1</td><td>1</td><td>1</td></tr></tbody></table>

A single layer of particles was deposited to the composite panel employing the four spraying conditions abovementioned, producing 50 mm long metal tracks.

To assess the characteristics of the metallized panels, various microscopic analysis techniques were employed. A Hirox ST-AS microscope was used to examine the coating width, while a confocal Olympus OLS5100 confocal microscope was utilized to assess variations in surface characteristics.

## Results and Discussion

As highlighted in the previous section, it is crucial to determine the surface on which to deposit. In fact, the presence of the Teflon mould ensures different surface properties for the top face and the back face. The acquisitions of height surfaces through confocal microscopy of the two surfaces before deposition are shown in Fig.1.
reinforced with carbon fibres (carbon fibre fabric 3k T300 Twill 2×2 by Toray). The composite is then produced by hot curing the resin at 120 °C for 1 h 30 min and then continuing with a further cure phase at 140 °C for 2 h. The composite panels produced have dimensions of 100x100mm and a thickness of 3mm. Further information on the manufacturing of vitrimer-based composite are presented in previous works of the authors [16–18].

In order to determine the most suitable side for deposition, confocal analyses were carried out on the composite surface. This was done because each face of the surface exhibited different surface characteristics due to the contact with the Teflon mould. In fact, the surface characteristics of the substrate, could strongly influence the CS deposition, as assessed in previous studies [19].

Once the panels were formed, a metallization process was applied to them using Cold Spray (CS) technology. The chosen powders for the metallization were micrometric powders of AlSi10Mg (LPW South Group), specifically selected for their compatibility with composite materials based on previous experiments conducted by the authors [19,20]. These powders were found to be highly suitable for deposition on the panels, even at low temperatures.

The metallization process was carried out using a low-pressure cold spray machine, specifically the Dycomet 423 CS equipment. Nitrogen gas was used as the carrier gas in this process.

Based on a preliminary sensitivity analysis, not presented here for the sake of brevity, the most significant process parameters have been identified and are outlined in Table 1.

This analysis underscored the importance of temperature and Standoff Distance (SoD) in the deposition process for composite panels with a vitrimer matrix. Previous studies have confirmed that vitrimers exhibit enhanced flexibility when heated above their Tv, instead of becoming brittle [15,21]. Therefore, it is essential to employ a carrier gas at an appropriate temperature and distance from the nozzle to effectively heat the substrate and achieve the desired temperature.

Two temperature values, T₁=200 °C and T₂=300 °C, and two values of SoD, SO₁=70 mm and SO₂=100mm, have been selected, which represent the change in behaviour of the vitrimeric matrix.

<table><caption>Table 1 – Selected process parameters for CS deposition of AlSi10Mg powders on composite vitrimer-based substrates.</caption><thead><tr><th></th><th>C<sub>T1SO1</sub></th><th>C<sub>T2SO1</sub></th><th>C<sub>T1SO2</sub></th><th>C<sub>T2SO2</sub></th></tr></thead><tbody><tr><th>Inlet gas temperature [°C]</th><td>200</td><td>300</td><td>200</td><td>300</td></tr><tr><th>Gas pressure [bar]</th><td>6</td><td>6</td><td>6</td><td>6</td></tr><tr><th>SoD [mm]</th><td>70</td><td>70</td><td>100</td><td>100</td></tr><tr><th>Gun traverse speed [mm/s]</th><td>1</td><td>1</td><td>1</td><td>1</td></tr></tbody></table>

A single layer of particles was deposited to the composite panel employing the four spraying conditions abovementioned, producing 50 mm long metal tracks.

To assess the characteristics of the metallized panels, various microscopic analysis techniques were employed. A Hirox ST-AS microscope was used to examine the coating width, while a confocal Olympus OLS5100 confocal microscope was utilized to assess variations in surface characteristics.

## Results and Discussion

As highlighted in the previous section, it is crucial to determine the surface on which to deposit. In fact, the presence of the Teflon mould ensures different surface properties for the top face and the back face. The acquisitions of height surfaces through confocal microscopy of the two surfaces before deposition are shown in Fig.1.
<figure><img src="image_2368.png" /><figcaption>Fig. 1 – Confocal acquisition of the surface of the bare composite before the CS deposition a) rough back surface b) smooth top surface of the composite</figcaption></figure>

Upon examination of the image and analyzing the values of Arithmetical Mean Height (Sa) of the surfaces, it can be concluded that the lower surface shown in Fig.1a, which is in contact with the mould, appears to be rougher. (Sa=28.80 µm). Furthermore, it is possible to notice from the figure that the presence of fibres is clearly visible. This could hinder the deposition process and potentially cause damage to the fibres, thereby reducing the composite's properties. Conversely, the other surface appears to be smoother (Sa=18.73 µm) and therefore better suited for CS deposition.

Hence, the depositions were performed on the upper surface. As mentioned previously, process parameters were utilized to emphasize the shift of the vitrimer's behaviour from being glass-like to becoming ductile. This transition can be clearly observed by examining the images captured by a confocal microscope, as depicted in Fig. 2.
<figure><img src="image_2359.png" /><figcaption>Fig. 1 – Confocal acquisition of the surface of the bare composite before the CS deposition a) rough back surface b) smooth top surface of the composite</figcaption></figure>

Upon examination of the image and analyzing the values of Arithmetical Mean Height (Sa) of the surfaces, it can be concluded that the lower surface shown in Fig.1a, which is in contact with the mould, appears to be rougher. (Sa=28.80 µm). Furthermore, it is possible to notice from the figure that the presence of fibres is clearly visible. This could hinder the deposition process and potentially cause damage to the fibres, thereby reducing the composite's properties. Conversely, the other surface appears to be smoother (Sa=18.73 µm) and therefore better suited for CS deposition.

Hence, the depositions were performed on the upper surface. As mentioned previously, process parameters were utilized to emphasize the shift of the vitrimer's behaviour from being glass-like to becoming ductile. This transition can be clearly observed by examining the images captured by a confocal microscope, as depicted in Fig. 2.
<figure><img src="image_2369.png" /><figcaption>Fig.2 – Confocal images of the coated panels: a) Specimen $C_{T1SO1}$ b) Specimen $C_{T2SO1}$
c) Specimen $C_{T1SO2}$ d) Specimen $C_{T2SO2}$</figcaption></figure>

It can be observed that when the lowest temperature is selected (Fig.2a and Fig.2c), only a small number of particles can be seen on the vitrimer substrate. In fact, this temperature of 200 °C is actually lower than the vitrimer's topology freezing transition temperature; therefore, the vitrimer vitrimer will not be able to deform enough to accommodate the impinging particles properly. As a result, cracks could form on the surface resulting in poor adhesion.

On the other hand, when higher temperatures (namely 300 °C) are employed and Tv is reached a rather homogeneous coating can be observed on the entire surface (Fig.2b and Fig.2d). Specifically when a lower standoff value is used (Fig.2b), the coating appears even thicker and more uniform. This is because the vitrimer is heated further, allowing it to better accommodate the particles. It is important to note that lower standoff values can also lead to surface degradation, so finding the right balance is crucial. This can be further confirmed by observing the height acquisitions of the surfaces of the coating, presented in Fig.3.
<figure><img src="image_2360.png" /><figcaption>Fig.2 – Confocal images of the coated panels: a) Specimen $C_{T1SO1}$ b) Specimen $C_{T2SO1}$
c) Specimen $C_{T1SO2}$ d) Specimen $C_{T2SO2}$</figcaption></figure>

It can be observed that when the lowest temperature is selected (Fig.2a and Fig.2c), only a small number of particles can be seen on the vitrimer substrate. In fact, this temperature of 200 °C is actually lower than the vitrimer's topology freezing transition temperature; therefore, the vitrimer vitrimer will not be able to deform enough to accommodate the impinging particles properly. As a result, cracks could form on the surface resulting in poor adhesion.

On the other hand, when higher temperatures (namely 300 °C) are employed and Tv is reached a rather homogeneous coating can be observed on the entire surface (Fig.2b and Fig.2d). Specifically when a lower standoff value is used (Fig.2b), the coating appears even thicker and more uniform. This is because the vitrimer is heated further, allowing it to better accommodate the particles. It is important to note that lower standoff values can also lead to surface degradation, so finding the right balance is crucial. This can be further confirmed by observing the height acquisitions of the surfaces of the coating, presented in Fig.3.
<figure><img src="image_2370.png" /><figcaption>Fig.3 – Confocal acquisition of the surface of the coated panels: a) Specimen $C_{T1SO1}$ ($Sa$=86.34 µm) b) Specimen $C_{T2SO1}$ ($Sa$=41.37 µm) c) Specimen $C_{T1SO2}$ ($Sa$=92.87 µm) d) Specimen $C_{T2SO2}$ ($Sa$=39.33 µm)</figcaption></figure>

It is noticeable that all process parameters result in higher Sa values compared to uncoated surfaces. This indicates that even in cases where the coating is not successfully produced (Fig.3a and Fig.3c), the surface still experiences damage and the formation of cracks, leading to an increase in the depth of the surface valleys.

When the coating is successfully applied on the surfaces, lower Sa values can be observed compared to those obtained with the lowest temperature values. For instance, comparing Fig.3a and Fig.3b it is possible to observe lower Sa value for the second set of process parameters. This is due to a homogeneous arrangement of the coating on the entire surface. However, the Sa value of Fig.3b is still higher than that of the bare surface, suggesting that the particles are unable to deform enough to create a completely continuous coating.

In order to compare the quality and characteristics of the produced coatings, obtained with the most effective process parameters (samples $C_{T2SO1}$ and $C_{T2SO2}$) further analyses were carried out using a Hirox microscope. The images of the coatings are presented in Fig.4.
<figure><img src="image_2361.png" /><figcaption>Fig.3 – Confocal acquisition of the surface of the coated panels: a) Specimen $C_{T1SO1}$ ($Sa$=86.34 µm) b) Specimen $C_{T2SO1}$ ($Sa$=41.37 µm) c) Specimen $C_{T1SO2}$ ($Sa$=92.87 µm) d) Specimen $C_{T2SO2}$ ($Sa$=39.33 µm)</figcaption></figure>

It is noticeable that all process parameters result in higher Sa values compared to uncoated surfaces. This indicates that even in cases where the coating is not successfully produced (Fig.3a and Fig.3c), the surface still experiences damage and the formation of cracks, leading to an increase in the depth of the surface valleys.

When the coating is successfully applied on the surfaces, lower Sa values can be observed compared to those obtained with the lowest temperature values. For instance, comparing Fig.3a and Fig.3b it is possible to observe lower Sa value for the second set of process parameters. This is due to a homogeneous arrangement of the coating on the entire surface. However, the Sa value of Fig.3b is still higher than that of the bare surface, suggesting that the particles are unable to deform enough to create a completely continuous coating.

In order to compare the quality and characteristics of the produced coatings, obtained with the most effective process parameters (samples $C_{T2SO1}$ and $C_{T2SO2}$) further analyses were carried out using a Hirox microscope. The images of the coatings are presented in Fig.4.
<figure><img src="image_2371.png" /><figcaption>Fig.4 – Hirox image of the Aluminium coating on vitrimer-based panel a) $C_{T2SO1}$ coating obtained with $T=300$ °C and $S.O=70$ mm b) $C_{T2SO2}$ coating obtained with $T=300$ °C and $S.O=100$ mm</figcaption></figure>

It is possible to observe that the coating obtained at the lowest standoff distance value appears to cover the surface more thoroughly. In fact, when the distance of the nozzle from the substrate is low the combined effect of the ductile matrix accommodating the impacting particles and the stiffening of the panel due to the presence of fibres leads to a more homogeneous coating where the particles are able to deform and produce the coating.

On the other hand, for the highest standoff value, only the areas between the fibres appear to be coated. These areas are indeed richer in vitrimeric matrix, which becomes ductile at high temperatures. The particles, for such high standoff values, do not seem to have enough energy to effectively adhere to the surface. Further analyses of the produced coatings under the Hirox microscope are shown in Fig.5 and Fig.6.

<figure><img src="image_2372.png" /><figcaption>Fig.5 - Hirox 3D image and coating height of the specimen obtained with spraying condition $T=300$ °C and $S.O=70$ mm</figcaption></figure>
<figure><img src="image_2362.png" /><figcaption>Fig.4 – Hirox image of the Aluminium coating on vitrimer-based panel a) $C_{T2SO1}$ coating obtained with $T=300$ °C and $S.O=70$ mm b) $C_{T2SO2}$ coating obtained with $T=300$ °C and $S.O=100$ mm</figcaption></figure>

It is possible to observe that the coating obtained at the lowest standoff distance value appears to cover the surface more thoroughly. In fact, when the distance of the nozzle from the substrate is low the combined effect of the ductile matrix accommodating the impacting particles and the stiffening of the panel due to the presence of fibres leads to a more homogeneous coating where the particles are able to deform and produce the coating.

On the other hand, for the highest standoff value, only the areas between the fibres appear to be coated. These areas are indeed richer in vitrimeric matrix, which becomes ductile at high temperatures. The particles, for such high standoff values, do not seem to have enough energy to effectively adhere to the surface. Further analyses of the produced coatings under the Hirox microscope are shown in Fig.5 and Fig.6.

<figure><img src="image_2363.png" /><figcaption>Fig.5 - Hirox 3D image and coating height of the specimen obtained with spraying condition $T=300$ °C and $S.O=70$ mm</figcaption></figure>
<figure><img src="image_2373.png" /><figcaption>Fig. 6 - Hirox 3D image and coating height of the specimen obtained with spraying condition $T=300$ °C and $S.O=100$ mm</figcaption></figure>

Through analysis of the Hirox 3D images, one can examine the width of the coating profiles. Comparing Fig. 5 and Fig. 6 it can be observed that the coating trace on the $C_{T2SO1}$ sample is wider than that on the $C_{T2SO2}$ sample. This is because as the Standoff Distance (SoD) increases, the particles impact at lower velocities and it is more likely that the particles in the peripheral zones of the flow (which impact with a velocity component perpendicular to the substrate) will have a lower velocity than the critical velocity, which is the minimum velocity required for coating adhesion [22].

## Conclusions

Based on the findings and discussions from the conducted experiments, it can be inferred that:

* The process of cold spray deposition successfully applied aluminium particles onto composite panels containing a vitrimer material as the matrix.

* The use of epoxy vitrimer as the matrix material appeared to be promising for cold spray metallization, as it exhibited similar properties to a thermoplastic material above its transition temperature (Tv), while still maintaining its mechanical and thermal characteristics below this temperature.

* Two key factors for achieving successful deposition on vitrimer-based composites were identified as the gas temperature and standoff distance. Specifically, it was crucial for the gas temperature to be above the topology freezing transition temperature.

* For temperatures lower than the transition temperature, it is not possible to achieve a homogeneous coating on the surface, which is damaged by the impacting particles.

* Using lower standoff values, however, increases the likelihood of achieving a more uniform coating and wider metallized traces, as it allows for a greater number of particles to have velocities higher than the critical velocity.

## Acknowledgements

The authors would like to thank the MUR for funding the research activities through the PRIN2022 PNRR - FRONTIERS grant.
<figure><img src="image_2364.png" /><figcaption>Fig. 6 - Hirox 3D image and coating height of the specimen obtained with spraying condition $T=300$ °C and $S.O=100$ mm</figcaption></figure>

Through analysis of the Hirox 3D images, one can examine the width of the coating profiles. Comparing Fig. 5 and Fig. 6 it can be observed that the coating trace on the $C_{T2SO1}$ sample is wider than that on the $C_{T2SO2}$ sample. This is because as the Standoff Distance (SoD) increases, the particles impact at lower velocities and it is more likely that the particles in the peripheral zones of the flow (which impact with a velocity component perpendicular to the substrate) will have a lower velocity than the critical velocity, which is the minimum velocity required for coating adhesion [22].

## Conclusions

Based on the findings and discussions from the conducted experiments, it can be inferred that:

* The process of cold spray deposition successfully applied aluminium particles onto composite panels containing a vitrimer material as the matrix.

* The use of epoxy vitrimer as the matrix material appeared to be promising for cold spray metallization, as it exhibited similar properties to a thermoplastic material above its transition temperature (Tv), while still maintaining its mechanical and thermal characteristics below this temperature.

* Two key factors for achieving successful deposition on vitrimer-based composites were identified as the gas temperature and standoff distance. Specifically, it was crucial for the gas temperature to be above the topology freezing transition temperature.

* For temperatures lower than the transition temperature, it is not possible to achieve a homogeneous coating on the surface, which is damaged by the impacting particles.

* Using lower standoff values, however, increases the likelihood of achieving a more uniform coating and wider metallized traces, as it allows for a greater number of particles to have velocities higher than the critical velocity.

## Acknowledgements

The authors would like to thank the MUR for funding the research activities through the PRIN2022 PNRR - FRONTIERS grant.
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[16] B. Palmieri, F. Cilento, E. Amendola, T. Valente, S. Dello Iacono, M. Giordano, A. Martone, Influence of Catalyst Content and Epoxy/Carboxylate Ratio on Isothermal Creep of Epoxy Vitrimers, Polymers (Basel) 15 (2023) 3845 https://doi.org/10.3390/polym15183845

[17] B. Palmieri, F. Cilento, E. Amendola, T. Valente, S. Dello Iacono, M. Giordano, A. Martone, An Investigation of the Healing Efficiency of Epoxy Vitrimer Composites Based on Zn2+ Catalyst, Polymers (Basel) 15 (2023) 3611 https://doi.org/10.3390/polym15173611

[18] B. PALMIERI, Viscoelastic characterization of reformable epoxy vitrimers composites, in: 2023: pp. 1871–1878 https://doi.org/10.21741/9781644902479-202

[19] A.S. Perna, A. Viscusi, R. Della Gatta, A. Astarita, Integrating 3D printing of polymer matrix composites and metal additive layer manufacturing: surface metallization of 3D printed composite panels through cold spray deposition of aluminium particles, International Journal of Material Forming 15 (2022) 15 https://doi.org/10.1007/S12289-022-01665-9

[20] R. Della Gatta, A. Viscusi, A.S. Perna, A. Caraviello, A. Astarita, R. Della Gatta, A. Viscusi, A.S. Perna, A. Caraviello, A. Astarita, R. Della Gatta, S. Perna, Cold spray process for the production of AlSi10Mg coatings on glass fibers reinforced polymers, Materials and Manufacturing Processes 36 (2021) 106–121 https://doi.org/10.1080/10426914.2020.1813895

[21] J. Zheng, Z.M. Png, S.H. Ng, G.X. Tham, E. Ye, S.S. Goh, X.J. Loh, Z. Li, Vitrimers: Current research trends and their emerging applications, Materials Today 51 (2021) 586–625 https://doi.org/10.1016/j.mattod.2021.07.003

[22] W.Y. Li, C. Zhang, X.P. Guo, G. Zhang, H.L. Liao, C.J. Li, C. Coddet, Effect of Standoff Distance on Coating Deposition Characteristics in Cold Spraying, Mater. Des. 29 (2008) 297–304 https://doi.org/10.1016/j.matdes.2007.02.005
# Applicability of productivity enhanced sinking-EDM of cemented carbide forming tools for bipolar plates

PETERSEN Timm¹,ᵃ*, VOIGTS Herman¹,ᵇ, HERRIG Tim¹,ᶜ and
BERGS Thomas¹,²,ᵈ

¹Manufacturing Technology Institute (MTI) of RWTH Aachen University, Campus-Boulevard 30, 52074 Aachen, Germany

²Fraunhofer Institute for Production Technology IPT, Steinbachstraße 17, 52074 Aachen, Germany

ᵃt.petersen@mti.rwth-aachen.de, ᵇh.voigts@mti.rwth-aachen.de, ᶜt.herrig@mti.rwth-aachen.de, ᵈt.bergs@mti.rwth-aachen.de

**Keywords:** Cemented Carbide, Sinking-EDM, Coining, Freeform Surface, Bipolar Plate

**Abstract.** The present research uses a previously published method to reduce the duration to machine cemented carbide forming tools by SEDM. Because of the forming setup a relatively thick sheet metal was used, which is why a forming simulation was conducted beforehand to achieve a valuable design, derived from bipolarplates, to observe wear in the forming experiments.

**Introduction**

One of the enablers to reduce the amount of greenhouse gas emissions in the transportation sector are fuel cells [1]. The large number of bipolarplates needed in a fuel cell and their tight tolerances make them cost drivers [2,3]. Therefore, it is of importance to improve the manufacturing process to reduce costs and to enhance their degree of effectiveness [4]. This will allow for a more sustainable and economically acceptable transformation to green transportation.

The effectiveness is enhanced with a rectangular flowfield geometry, which is responsible for the distribution of the working media. This geometric demand and the tight tolerances call for tool materials of high rigidity that are able to transfer their shape to the workpiece accurately and repeatably [5]. A class of materials, which is able to withstand the large compressive stresses needed to comply with these demands are cemented carbides. Another advantageous property of cemented carbides, regarding the challenging tool geometries found in the filigree channel structure of the flowfield, is their large resistance to wear. The high hardness of cemented carbides requires machining technologies like electrical discharge machining (EDM). Its thermophysical working principle enables for machining workpieces independently of their mechanical properties. Albeit, there are two concerns regarding the EDM of cemented carbides, which are subject of several recent publications. On the one hand, the machining of cemented carbides is less productive in comparison to steel and on the other hand the thermal removal mechanism can cause cracks and flaws that would reduce the longevity of the workpieces – in this case the forming tools. Previously conducted research has described the effects sinking EDM has on the surface integrity of cemented carbides [6]. Additionally, the resulting workpiece functionality has been examined [6,7]. Finally, the process has been altered in order to reduce machining duration and the corresponding effect on the surface integrity has been found to be negligible [8]. As a final step in this series of research, the altered sinking EDM process is compared to the standard process in a real application. Due to the necessity of large scale tests, the forming of bipolarplates was chosen as the application test case. It is therefore an additional aim of this research to investigate the wear mechanisms of cemented carbide forming tools that were machined by sinking EDM.

<figure><img src="image_2374.png" /></figure>
# Applicability of productivity enhanced sinking-EDM of cemented carbide forming tools for bipolar plates

PETERSEN Timm¹,ᵃ*, VOIGTS Herman¹,ᵇ, HERRIG Tim¹,ᶜ and
BERGS Thomas¹,²,ᵈ

¹Manufacturing Technology Institute (MTI) of RWTH Aachen University, Campus-Boulevard 30, 52074 Aachen, Germany

²Fraunhofer Institute for Production Technology IPT, Steinbachstraße 17, 52074 Aachen, Germany

ᵃt.petersen@mti.rwth-aachen.de, ᵇh.voigts@mti.rwth-aachen.de, ᶜt.herrig@mti.rwth-aachen.de, ᵈt.bergs@mti.rwth-aachen.de

**Keywords:** Cemented Carbide, Sinking-EDM, Coining, Freeform Surface, Bipolar Plate

**Abstract.** The present research uses a previously published method to reduce the duration to machine cemented carbide forming tools by SEDM. Because of the forming setup a relatively thick sheet metal was used, which is why a forming simulation was conducted beforehand to achieve a valuable design, derived from bipolarplates, to observe wear in the forming experiments.

**Introduction**

One of the enablers to reduce the amount of greenhouse gas emissions in the transportation sector are fuel cells [1]. The large number of bipolarplates needed in a fuel cell and their tight tolerances make them cost drivers [2,3]. Therefore, it is of importance to improve the manufacturing process to reduce costs and to enhance their degree of effectiveness [4]. This will allow for a more sustainable and economically acceptable transformation to green transportation.

The effectiveness is enhanced with a rectangular flowfield geometry, which is responsible for the distribution of the working media. This geometric demand and the tight tolerances call for tool materials of high rigidity that are able to transfer their shape to the workpiece accurately and repeatably [5]. A class of materials, which is able to withstand the large compressive stresses needed to comply with these demands are cemented carbides. Another advantageous property of cemented carbides, regarding the challenging tool geometries found in the filigree channel structure of the flowfield, is their large resistance to wear. The high hardness of cemented carbides requires machining technologies like electrical discharge machining (EDM). Its thermophysical working principle enables for machining workpieces independently of their mechanical properties. Albeit, there are two concerns regarding the EDM of cemented carbides, which are subject of several recent publications. On the one hand, the machining of cemented carbides is less productive in comparison to steel and on the other hand the thermal removal mechanism can cause cracks and flaws that would reduce the longevity of the workpieces – in this case the forming tools. Previously conducted research has described the effects sinking EDM has on the surface integrity of cemented carbides [6]. Additionally, the resulting workpiece functionality has been examined [6,7]. Finally, the process has been altered in order to reduce machining duration and the corresponding effect on the surface integrity has been found to be negligible [8]. As a final step in this series of research, the altered sinking EDM process is compared to the standard process in a real application. Due to the necessity of large scale tests, the forming of bipolarplates was chosen as the application test case. It is therefore an additional aim of this research to investigate the wear mechanisms of cemented carbide forming tools that were machined by sinking EDM.

<figure><img src="image_2365.png" /></figure>
Punch design and preliminary simulations

For wear testing, a die setup from [9] was used. However, the process was designed for coining thus only the embossing pattern is of relevance. The punch contour was used for guidance in an existing tool setup. A bipolarplate geometry, provided by Fraunhofer IPT, Aachen was used as a reference for the present experiments. Additionally, two features were added to allow for an evaluation of different radii and coining heights, cf. Fig. 1.

<figure><img src="image_2375.png" /><figcaption>Fig. 1: Reference bipolarplate (left), punch design with section of the flowfield of the bipolarplate (centre) and isometric view of the embossing pattern (right).</figcaption></figure>

In order to reduce the tool life time, an analogue setup for tribological testing of the channel structures of the embossing pattern was used with the sheet metal strip thyssenkrupp precidur (AISI 52100/1.3505) with a thickness of 3.7 mm and UTS of 660 MPa. This ensures increased wear rates at a similar load collective with significantly higher stress levels (contact pressure: 600 MPa) compared to bipolarplate pressing (contact pressure: 85 MPa). The material was selected due to its similar tribological behavior compared to thin foil used for bipolarplates such as AISI 316L/1.4404. However, the coining setup, led to the question whether the cavity between the channels of the flowfield would fill properly to see friction on the flanks of the channel structure. The software Forge 4.0 was used to design material flow and stresses in the embossing pattern. Fig. 2 displays the equivalent stress for a directly derived flowfield geometry.

As shown in the top view, the stresses are the highest around the corners. In the enlarged cross sectional view the colored workpiece material does not reach the punch contour. The height of the

<figure><img src="image_2376.png" /><figcaption>Fig. 2: Simulation results for both geometries with top view (top left), cross sectional view (bottom left) and enlarged cross sectional view of original (top right) and alternative design (bottom right).</figcaption></figure>
Punch design and preliminary simulations

For wear testing, a die setup from [9] was used. However, the process was designed for coining thus only the embossing pattern is of relevance. The punch contour was used for guidance in an existing tool setup. A bipolarplate geometry, provided by Fraunhofer IPT, Aachen was used as a reference for the present experiments. Additionally, two features were added to allow for an evaluation of different radii and coining heights, cf. Fig. 1.

<figure><img src="image_2366.png" /><figcaption>Fig. 1: Reference bipolarplate (left), punch design with section of the flowfield of the bipolarplate (centre) and isometric view of the embossing pattern (right).</figcaption></figure>

In order to reduce the tool life time, an analogue setup for tribological testing of the channel structures of the embossing pattern was used with the sheet metal strip thyssenkrupp precidur (AISI 52100/1.3505) with a thickness of 3.7 mm and UTS of 660 MPa. This ensures increased wear rates at a similar load collective with significantly higher stress levels (contact pressure: 600 MPa) compared to bipolarplate pressing (contact pressure: 85 MPa). The material was selected due to its similar tribological behavior compared to thin foil used for bipolarplates such as AISI 316L/1.4404. However, the coining setup, led to the question whether the cavity between the channels of the flowfield would fill properly to see friction on the flanks of the channel structure. The software Forge 4.0 was used to design material flow and stresses in the embossing pattern. Fig. 2 displays the equivalent stress for a directly derived flowfield geometry.

As shown in the top view, the stresses are the highest around the corners. In the enlarged cross sectional view the colored workpiece material does not reach the punch contour. The height of the

<figure><img src="image_2367.png" /><figcaption>Fig. 2: Simulation results for both geometries with top view (top left), cross sectional view (bottom left) and enlarged cross sectional view of original (top right) and alternative design (bottom right).</figcaption></figure>
channels in the flowfield of the workpiece only reach 45 % of the designated height. Because of the curvature of the material within the channels, there is even less material in the vicinity of the flanks of the flowfield of the punch. In order to improve the material filling between the channels, the distance between channels was increased. The number of channels was reduced to six. The simulation was repeated with the adapted geometry. It is apparent that the contour was sufficiently filled to create material flow over the radii of the channel pattern. The height of the channels in the workpiece reached 78 % of the designated height. The material appears to create enough friction thus wear around the corners and flanks in analogy to flowfield drawing.

The final geometry that was used for the derivation of the sinking-EDM tool is displayed in Fig. 3. The features 1 and 2 have a height of 0.7 mm, whereas the flowfield only has a height of 0.3 mm. Both features have a bottom radius of 0.25 mm, feature 1 a head radius of 0.05 mm and feature 2 has a head radius of 0.3 mm. All the corners have different radii as well to allow for a differentiation between wear mechanisms. The channels 1 to 6 all have the same corner radius, which was taken from the initial geometry. They have increasing head radii by an increment of 0.02 mm, starting at 0.1 mm for channel 1 until 0.2 mm for channel 6. The bottom radii increase accordingly, however, starting at 0.2 mm for channel 1 up to 0.3 mm for channel 6. The channel height and different radii were chosen according to research conducted regarding the formability of foils used for bipolarplate manufacturing [10].

<figure><img src="image_2377.png" /><figcaption>Fig. 3: Final punch design with a distance between the flowfield channels of 1.7 mm (left), cross sectional view of the features with corresponding radii (top right) and cross sectional view of the flowfield with exemplary channel head and bottom radii (bottom right).</figcaption></figure>

## Machining of cemented carbide punches

The punches were machined from EDM-blocks of the tungsten carbide grade Ceratizit CTF35 (average grain size 0.8 µm, Cobalt-content 17.5 wt-%). First the outer shape of the punches was cut via wire-EDM on a GFMS P550 with 4 trim cuts up to a surface roughness of Ra = 0.4 µm. Additionally, the guide for the forming tool has been cut in a PM steel on the same machine tool. Subsequently, the desired shape was machined via sinking-EDM on a GFMS Form 2000 VHP. The machining sequence and the machining parameters are displayed in Fig. 4. Even though the electrode material was chosen as copper-tungsten, with 30 wt-% copper, the wear was expected to be of a magnitude that prevents a proper shape transfer. Therefore, three electrodes were milled with a machining duration of 16 h each. In order to spread the wear evenly, roughing and super-finishing started with punch 1 and the intermediate finishing operation started reversely with punch 2. The necessity of this altered sequence can also be seen from the machining duration of machining steps 1 and 2. Because of the significant linear wear of approximately 15 %, electrode 1
channels in the flowfield of the workpiece only reach 45 % of the designated height. Because of the curvature of the material within the channels, there is even less material in the vicinity of the flanks of the flowfield of the punch. In order to improve the material filling between the channels, the distance between channels was increased. The number of channels was reduced to six. The simulation was repeated with the adapted geometry. It is apparent that the contour was sufficiently filled to create material flow over the radii of the channel pattern. The height of the channels in the workpiece reached 78 % of the designated height. The material appears to create enough friction thus wear around the corners and flanks in analogy to flowfield drawing.

The final geometry that was used for the derivation of the sinking-EDM tool is displayed in Fig. 3. The features 1 and 2 have a height of 0.7 mm, whereas the flowfield only has a height of 0.3 mm. Both features have a bottom radius of 0.25 mm, feature 1 a head radius of 0.05 mm and feature 2 has a head radius of 0.3 mm. All the corners have different radii as well to allow for a differentiation between wear mechanisms. The channels 1 to 6 all have the same corner radius, which was taken from the initial geometry. They have increasing head radii by an increment of 0.02 mm, starting at 0.1 mm for channel 1 until 0.2 mm for channel 6. The bottom radii increase accordingly, however, starting at 0.2 mm for channel 1 up to 0.3 mm for channel 6. The channel height and different radii were chosen according to research conducted regarding the formability of foils used for bipolarplate manufacturing [10].

<figure><img src="image_2368.png" /><figcaption>Fig. 3: Final punch design with a distance between the flowfield channels of 1.7 mm (left), cross sectional view of the features with corresponding radii (top right) and cross sectional view of the flowfield with exemplary channel head and bottom radii (bottom right).</figcaption></figure>

## Machining of cemented carbide punches

The punches were machined from EDM-blocks of the tungsten carbide grade Ceratizit CTF35 (average grain size 0.8 µm, Cobalt-content 17.5 wt-%). First the outer shape of the punches was cut via wire-EDM on a GFMS P550 with 4 trim cuts up to a surface roughness of Ra = 0.4 µm. Additionally, the guide for the forming tool has been cut in a PM steel on the same machine tool. Subsequently, the desired shape was machined via sinking-EDM on a GFMS Form 2000 VHP. The machining sequence and the machining parameters are displayed in Fig. 4. Even though the electrode material was chosen as copper-tungsten, with 30 wt-% copper, the wear was expected to be of a magnitude that prevents a proper shape transfer. Therefore, three electrodes were milled with a machining duration of 16 h each. In order to spread the wear evenly, roughing and super-finishing started with punch 1 and the intermediate finishing operation started reversely with punch 2. The necessity of this altered sequence can also be seen from the machining duration of machining steps 1 and 2. Because of the significant linear wear of approximately 15 %, electrode 1
was 0.14 mm shorter before the start of machining of punch 2 and was hence significantly faster. Due to the additional wear, punch 2 was missing approximately 0.21 mm from the original machining depth of 1 mm. This difference was compensated with the finishing operation. As indicated in Fig. 4 (right), the roughing steps are conducted again with both subsequent electrodes. This ensures an equal spread of the wear and that both punches reach the desired machining depth. The wear of the electrodes was measured with the sinking EDM machine tool after machining both punches and was 0.21 mm, 0.04 mm and 0.01 mm for the roughing, finishing and super-finishing electrode respectively.

<table><caption>Fig. 4: Description of machining sequence with machining duration (left) and machining parameters with resulting surface roughness (right).</caption><thead><tr><th></th><th></th><th>Punch 1</th><th>Punch 2</th></tr></thead><tbody><tr><td rowspan="3">Roughing</td><td>Machining step</td><td>1</td><td>2</td></tr><tr><td>Electrode</td><td>1</td><td>1</td></tr><tr><td>Time / min</td><td>435</td><td>384</td></tr><tr><td rowspan="3">Finishing</td><td>Machining step</td><td>4</td><td>3</td></tr><tr><td>Electrode</td><td>2</td><td>2</td></tr><tr><td>Time / min</td><td>72</td><td>155</td></tr><tr><td rowspan="3">Super-Finishing</td><td>Machining step</td><td>5</td><td>6</td></tr><tr><td>Electrode</td><td>3</td><td>3</td></tr><tr><td>Time / min</td><td>71</td><td>70</td></tr></tbody></table><table><thead><tr><th></th><th></th><th>I / A</th><th>T / μs</th><th>P / μs</th><th>U / V</th><th>SBOX</th><th>Ra / μm</th></tr></thead><tbody><tr><td rowspan="3">Finishing</td><td></td><td>4.7</td><td>7</td><td>15</td><td>-200</td><td>74</td><td>1.58</td></tr><tr><td></td><td>4.7</td><td>5</td><td>15</td><td>-200</td><td>69</td><td>1.26</td></tr><tr><td></td><td>4.7</td><td>4</td><td>10</td><td>-200</td><td>67</td><td>1</td></tr><tr><td rowspan="3">Finishing</td><td></td><td>3.2</td><td>3.2</td><td>7</td><td>-200</td><td>246</td><td>0.79</td></tr><tr><td></td><td>2.8</td><td>3.2</td><td>4.5</td><td>-200</td><td>219</td><td>0.63</td></tr><tr><td></td><td>2.6</td><td>3.2</td><td>3.6</td><td>-200</td><td>199</td><td>0.5</td></tr><tr><td rowspan="3">Finishing</td><td></td><td>2.6</td><td>2</td><td>3.6</td><td>-200</td><td>196</td><td>0.4</td></tr><tr><td></td><td>1.2</td><td>2</td><td>3.6</td><td>-170</td><td>195</td><td>0.31</td></tr><tr><td></td><td>1.2</td><td>1.6</td><td>3.6</td><td>170</td><td>195</td><td>0.25</td></tr><tr><td rowspan="3">Finishing</td><td></td><td>0.8</td><td>1.6</td><td>3.6</td><td>120</td><td>193</td><td>0.2</td></tr><tr><td></td><td></td><td></td><td></td><td></td><td></td><td></td></tr></tbody></table>

The machining parameters were adjusted according to the approach described in [8]. However, the expected reduction of machining duration was not observed in this setup. Since this was attributed to the wear, which caused different machining durations for both punches, a comparative study according to the setup described in [8] was conducted with the material relevant for this investigation. It was found that the machining time was reduced by 12%, which is in the order of magnitude that was expected. Additionally, the machining duration and the wear in the present research are relatively high, because of the filigree structures that were machined. In order to make sure that the shape with its varying radii is properly transferred from the electrodes to the punches, the undersize was only 0.05 mm per side. However, the wear of the electrodes was stronger than expected, which is why the final surfaces were not all the same. A measurement with the laser scanning microscope Keyence VK-X has revealed that the bottom surfaces of the punches reached a surface roughness of $Ra = 0.2 \ \mu m$. In contrast, the head surfaces of the flow channels of the punches have only reached a surface roughness of $Ra = 0.5 \ \mu m$. This is exactly the surface roughness that was reached after the finishing operation. Because the electrodes have worn differently in these different areas, the super-finishing operation was not anymore conducted on the head surface. This is illustrated by Fig. 5, where the part of the electrode that machines the bottom of the workpiece is the first to interact with the workpiece. With ongoing machining, this part is worn, even when the rest of the electrode has not yet started to machine. Only after the height difference in the electrode has been surpassed, the wear takes place in all regions quasi equally. After machining, the workpiece resembles the negative of the shape of the worn electrode and when a new electrode is used, there is a gap between the areas that were machined with a lesser worn electrode than those that were machined with parts of the electrodes that were subject to more wear. Even after using several electrodes, this effect can still be seen, since it is repeated with subsequent electrodes, however, with another magnitude.
was 0.14 mm shorter before the start of machining of punch 2 and was hence significantly faster. Due to the additional wear, punch 2 was missing approximately 0.21 mm from the original machining depth of 1 mm. This difference was compensated with the finishing operation. As indicated in Fig. 4 (right), the roughing steps are conducted again with both subsequent electrodes. This ensures an equal spread of the wear and that both punches reach the desired machining depth. The wear of the electrodes was measured with the sinking EDM machine tool after machining both punches and was 0.21 mm, 0.04 mm and 0.01 mm for the roughing, finishing and super-finishing electrode respectively.

<table><caption>Fig. 4: Description of machining sequence with machining duration (left) and machining parameters with resulting surface roughness (right).</caption><thead><tr><th></th><th></th><th>Punch 1</th><th>Punch 2</th></tr></thead><tbody><tr><td rowspan="3">Roughing</td><td>Machining step</td><td>1</td><td>2</td></tr><tr><td>Electrode</td><td>1</td><td>1</td></tr><tr><td>Time / min</td><td>435</td><td>384</td></tr><tr><td rowspan="3">Finishing</td><td>Machining step</td><td>4</td><td>3</td></tr><tr><td>Electrode</td><td>2</td><td>2</td></tr><tr><td>Time / min</td><td>72</td><td>155</td></tr><tr><td rowspan="3">Super-Finishing</td><td>Machining step</td><td>5</td><td>6</td></tr><tr><td>Electrode</td><td>3</td><td>3</td></tr><tr><td>Time / min</td><td>71</td><td>70</td></tr></tbody></table><table><thead><tr><th></th><th></th><th>I / A</th><th>T / µs</th><th>P / µs</th><th>U / V</th><th>SBOX</th><th>Ra / µm</th></tr></thead><tbody><tr><td rowspan="3">Finishing</td><td>Machining step</td><td>4.7</td><td>7</td><td>15</td><td>-200</td><td>74</td><td>1.58</td></tr><tr><td>Electrode</td><td>4.7</td><td>5</td><td>15</td><td>-200</td><td>69</td><td>1.26</td></tr><tr><td>Time / min</td><td>4.7</td><td>4</td><td>10</td><td>-200</td><td>67</td><td>1</td></tr><tr><td rowspan="3">Finishing</td><td>Machining step</td><td>3.2</td><td>3.2</td><td>7</td><td>-200</td><td>246</td><td>0.79</td></tr><tr><td>Electrode</td><td>2.8</td><td>3.2</td><td>4.5</td><td>-200</td><td>219</td><td>0.63</td></tr><tr><td>Time / min</td><td>2.6</td><td>3.2</td><td>3.6</td><td>-200</td><td>199</td><td>0.5</td></tr><tr><td rowspan="3">Finishing</td><td>Machining step</td><td>2.6</td><td>2</td><td>3.6</td><td>-200</td><td>196</td><td>0.4</td></tr><tr><td>Electrode</td><td>1.2</td><td>2</td><td>3.6</td><td>-170</td><td>195</td><td>0.31</td></tr><tr><td>Time / min</td><td>1.2</td><td>1.6</td><td>3.6</td><td>170</td><td>195</td><td>0.25</td></tr><tr><td rowspan="3">Finishing</td><td>Machining step</td><td>0.8</td><td>1.6</td><td>3.6</td><td>120</td><td>193</td><td>0.2</td></tr><tr><td>Electrode</td><td></td><td></td><td></td><td></td><td></td><td></td></tr><tr><td>Time / min</td><td></td><td></td><td></td><td></td><td></td><td></td></tr></tbody></table>

The machining parameters were adjusted according to the approach described in [8]. However, the expected reduction of machining duration was not observed in this setup. Since this was attributed to the wear, which caused different machining durations for both punches, a comparative study according to the setup described in [8] was conducted with the material relevant for this investigation. It was found that the machining time was reduced by 12%, which is in the order of magnitude that was expected. Additionally, the machining duration and the wear in the present research are relatively high, because of the filigree structures that were machined. In order to make sure that the shape with its varying radii is properly transferred from the electrodes to the punches, the undersize was only 0.05 mm per side. However, the wear of the electrodes was stronger than expected, which is why the final surfaces were not all the same. A measurement with the laser scanning microscope Keyence VK-X has revealed that the bottom surfaces of the punches reached a surface roughness of $Ra = 0.2 \ \mu m$. In contrast, the head surfaces of the flow channels of the punches have only reached a surface roughness of $Ra = 0.5 \ \mu m$. This is exactly the surface roughness that was reached after the finishing operation. Because the electrodes have worn differently in these different areas, the super-finishing operation was not anymore conducted on the head surface. This is illustrated by Fig. 5, where the part of the electrode that machines the bottom of the workpiece is the first to interact with the workpiece. With ongoing machining, this part is worn, even when the rest of the electrode has not yet started to machine. Only after the height difference in the electrode has been surpassed, the wear takes place in all regions quasi equally. After machining, the workpiece resembles the negative of the shape of the worn electrode and when a new electrode is used, there is a gap between the areas that were machined with a lesser worn electrode than those that were machined with parts of the electrodes that were subject to more wear. Even after using several electrodes, this effect can still be seen, since it is repeated with subsequent electrodes, however, with another magnitude.
<figure><img src="image_2378.png" /><figcaption>Fig. 5: Worn roughing electrode with corresponding linear material reduction (left) and corresponding error propagation (right): Electrode and workpiece before machining, worn electrode after first machining operation with corresponding workpiece and same workpiece with new, not worn electrode.</figcaption></figure>

The previously described effect can also be seen in the images taken via scanning electron microscopy (SEM). Fig. 6 depicts one of the flow channels which has the described features. On the one hand the bottom surfaces as well as the flank surfaces appear to be smooth. On the other hand, the head surface of the channel appears to be a lot rougher. Additionally, there are some single craters, which are significantly larger than those that make up the surrounding surfaces. These craters probably result from a concentration of the electric field on corners and the corresponding ignition of discharges with higher energies in earlier machining steps with a larger gap width. These single craters can be seen frequently on the corners of all machined features and might have been removed, if the frontal undersize of the electrodes had been adjusted according to the occurring wear.

<figure><img src="image_2379.png" /><figcaption>Fig. 6: Image for orientation (top left), SEM image of the second flow channel on the second punch.</figcaption></figure>

## Experimental setup for coining tests

The profile was coined into the sheet metal strip in an automatic stamping process for wear testing. A full coil of the aforementioned material AISI 52100/100Cr6 with thickness 3.7 mm and UTS of 660 MPa was used for a total of 7360 strokes. The experiments were performed on a servomechanical blanking press Feintool XFT 2500speed at stroke rates between 15 min⁻¹ and 45 min⁻¹. In order to coin the full profile into the sheet metal, the die was replaced with a rigid
<figure><img src="image_2369.png" /><figcaption>Fig. 5: Worn roughing electrode with corresponding linear material reduction (left) and corresponding error propagation (right): Electrode and workpiece before machining, worn electrode after first machining operation with corresponding workpiece and same workpiece with new, not worn electrode.</figcaption></figure>

The previously described effect can also be seen in the images taken via scanning electron microscopy (SEM). Fig. 6 depicts one of the flow channels which has the described features. On the one hand the bottom surfaces as well as the flank surfaces appear to be smooth. On the other hand, the head surface of the channel appears to be a lot rougher. Additionally, there are some single craters, which are significantly larger than those that make up the surrounding surfaces. These craters probably result from a concentration of the electric field on corners and the corresponding ignition of discharges with higher energies in earlier machining steps with a larger gap width. These single craters can be seen frequently on the corners of all machined features and might have been removed, if the frontal undersize of the electrodes had been adjusted according to the occurring wear.

<figure><img src="image_2370.png" /><figcaption>Fig. 6: Image for orientation (top left), SEM image of the second flow channel on the second punch.</figcaption></figure>

## Experimental setup for coining tests

The profile was coined into the sheet metal strip in an automatic stamping process for wear testing. A full coil of the aforementioned material AISI 52100/100Cr6 with thickness 3.7 mm and UTS of 660 MPa was used for a total of 7360 strokes. The experiments were performed on a servomechanical blanking press Feintool XFT 2500speed at stroke rates between 15 min⁻¹ and 45 min⁻¹. In order to coin the full profile into the sheet metal, the die was replaced with a rigid
anvil instead of a counterholder. Fig. 7 shows the tool used for the experiments. Coining took place in the top dead center (TDC) of the underneath drive. The stroke length was calibrated in the beginning for a full penetration of the sheet metal according to the process design from the simulations. The die setup was a 2-up die with two cavities. Punch 1 was EDMed with a reference technology whereas punch 2 was manufactured with the adapted technology (see above).

<figure><img src="image_2380.png" /><figcaption>Fig. 7: Punch 1 after SEDM (left) and tool setup with blue arrows indicating the movement of the underneath drive and red indicating the feed direction of the sheet metal (left).</figcaption></figure>

Coined specimens were extracted every 500 strokes for further analysis of forming results and eventual traces resulting from wear on the punch. For this, the coined surface was measured and evaluated using laser scanning microscopy. The tool state was analyzed using scanning electron microscopy (SEM) in initial state, after 1000 strokes and after the total 7360 strokes. The quality of the S-EDMed surface in all radii and surfaces was assessed visually. Traces of adhering sheet metal as well as cracks and surface wear indications were looked for.

## Results and Discussion

During the first 50 strokes, feature 1 broke from both punches. As mentioned above, the small radii were designed as edge cases. The small head radius and the flank experience tensile load during the return stroke because of clamping between the feature and sheet metal according to [9]. The tensile load causes crack formation in the bottom radius of the feature as can be seen on the non-broken features 2 in Fig. 8. Another explanation might be shear load on the features away from the middle of the strip due to lateral material flow in the compression phase.

<figure><img src="image_2381.png" /><figcaption>Fig. 8: Overview for orientation of corner on feature 2 of punch 1 (left) and detailed view of crack development on the bottom radius after 1000 strokes (left and centre) and after 7360 strokes (right).</figcaption></figure>
anvil instead of a counterholder. Fig. 7 shows the tool used for the experiments. Coining took place in the top dead center (TDC) of the underneath drive. The stroke length was calibrated in the beginning for a full penetration of the sheet metal according to the process design from the simulations. The die setup was a 2-up die with two cavities. Punch 1 was EDMed with a reference technology whereas punch 2 was manufactured with the adapted technology (see above).

<figure><img src="image_2371.png" /><figcaption>Fig. 7: Punch 1 after SEDM (left) and tool setup with blue arrows indicating the movement of the underneath drive and red indicating the feed direction of the sheet metal (left).</figcaption></figure>

Coined specimens were extracted every 500 strokes for further analysis of forming results and eventual traces resulting from wear on the punch. For this, the coined surface was measured and evaluated using laser scanning microscopy. The tool state was analyzed using scanning electron microscopy (SEM) in initial state, after 1000 strokes and after the total 7360 strokes. The quality of the S-EDMed surface in all radii and surfaces was assessed visually. Traces of adhering sheet metal as well as cracks and surface wear indications were looked for.

## Results and Discussion

During the first 50 strokes, feature 1 broke from both punches. As mentioned above, the small radii were designed as edge cases. The small head radius and the flank experience tensile load during the return stroke because of clamping between the feature and sheet metal according to [9]. The tensile load causes crack formation in the bottom radius of the feature as can be seen on the non-broken features 2 in Fig. 8. Another explanation might be shear load on the features away from the middle of the strip due to lateral material flow in the compression phase.

<figure><img src="image_2372.png" /><figcaption>Fig. 8: Overview for orientation of corner on feature 2 of punch 1 (left) and detailed view of crack development on the bottom radius after 1000 strokes (left and centre) and after 7360 strokes (right).</figcaption></figure>
Another major event during the forming tests was the partial demolition of channel 1 on punch 1, which was machined with the reference technology. The breakage did indicate itself with some major cracks along channel 1. Fig. 9 highlights the imprint of these cracks on a specimen after 750 strokes. Furthermore, after another 2000 strokes channel 1 broke off partially as visible by the specimen after 2750 strokes. The final stage of the demolition process was reached after 4000 strokes. Another ~3300 strokes were conducted without additional fractures.

<figure><img src="image_2382.png" /><figcaption>Fig. 9: Detailed view of the imprint of the cracks developing on punch 1 after 750 strokes (left), first specimen with the imprint of the first stage of the partially breakage of channel 1 (centre) and specimen with imprint of the intact punch 2 and the imprint of the final breakage state of punch 1 after 7000 strokes (right).</figcaption></figure>

Fracture of features and tribological wear are a result of a high compressive load collective. In previous work regarding cemented carbide wear during stamping, such fractures were explained by tensile and bending stresses during the stripping phase. High stresses in the coining phase create spring back of the formed surfaces causing a clamping situation between the coining features and the sheet metal. Additional lateral stresses due to material flow towards the strip edges as a result of upsetting the sheet metal are an explanation for the observed fractures. The sheet metal is compressed by the punch force and blankholder force, see Fig. 10.

<figure><img src="image_2383.png" /><figcaption>Fig. 10: Process forces during coining representing high compressive load collective on channel pattern for the second stroke (left) and the 7010th stroke (right).</figcaption></figure>

The axial load on both punches adds to almost 800 kN which equals about 600 MPa contact stress. The contact stress is in the range of the UTS of the sheet metal (660 MPa). The axial load is responsible for wear to be observed on the channel flanks. After 1000 strokes, the rate has been increased from 15 to 45 min⁻¹ to increase wear.

A profilometer was used to measure the depth of the imprinted profile of extracted workpieces. Fig. 11 depicts one exemplary sheet metal strip and the measurements of the depths of the imprints of both punches. The red lines on the left show the position and direction of the measurements, approximately 1 mm from the bottom end of the flow fields. As visible in the graphs on the right-
Another major event during the forming tests was the partial demolition of channel 1 on punch 1, which was machined with the reference technology. The breakage did indicate itself with some major cracks along channel 1. Fig. 9 highlights the imprint of these cracks on a specimen after 750 strokes. Furthermore, after another 2000 strokes channel 1 broke off partially as visible by the specimen after 2750 strokes. The final stage of the demolition process was reached after 4000 strokes. Another ~3300 strokes were conducted without additional fractures.

<figure><img src="image_2373.png" /><figcaption>Fig. 9: Detailed view of the imprint of the cracks developing on punch 1 after 750 strokes (left), first specimen with the imprint of the first stage of the partially breakage of channel 1 (centre) and specimen with imprint of the intact punch 2 and the imprint of the final breakage state of punch 1 after 7000 strokes (right).</figcaption></figure>

Fracture of features and tribological wear are a result of a high compressive load collective. In previous work regarding cemented carbide wear during stamping, such fractures were explained by tensile and bending stresses during the stripping phase. High stresses in the coining phase create spring back of the formed surfaces causing a clamping situation between the coining features and the sheet metal. Additional lateral stresses due to material flow towards the strip edges as a result of upsetting the sheet metal are an explanation for the observed fractures. The sheet metal is compressed by the punch force and blankholder force, see Fig. 10.

<figure><img src="image_2374.png" /><figcaption>Fig. 10: Process forces during coining representing high compressive load collective on channel pattern for the second stroke (left) and the 7010th stroke (right).</figcaption></figure>

The axial load on both punches adds to almost 800 kN which equals about 600 MPa contact stress. The contact stress is in the range of the UTS of the sheet metal (660 MPa). The axial load is responsible for wear to be observed on the channel flanks. After 1000 strokes, the rate has been increased from 15 to 45 min⁻¹ to increase wear.

A profilometer was used to measure the depth of the imprinted profile of extracted workpieces. Fig. 11 depicts one exemplary sheet metal strip and the measurements of the depths of the imprints of both punches. The red lines on the left show the position and direction of the measurements, approximately 1 mm from the bottom end of the flow fields. As visible in the graphs on the right-
hand side of Fig. 11, both positions yielded different depths. The flowfield on the right-hand side of the sheet metal strip, whose corresponding data is depicted in the upper graph, averaged to a depth of 0.255±0.045 mm. The other flowfield only reached a depth of 0.184±0.022 mm, averaged over all channels and investigated specimens. The reason for this difference was not attributed to different punch heights, since the position of the punches was swapped after 1000 strokes. However, this explains the slight drift in all measurements from specimen 1000 to specimen 1250. The reason for the difference between the two positions is suspected to be found in the tool setup. As depicted in Fig. 7, the punches are positioned behind one another. This is also indicated by the positions of the red lines in Fig. 11. The blue arrow indicates the feed direction of the sheet metal strip. After one stroke is conducted, the relatively small feed of the sheet metal strip causes the guide plate to press down on the flowfield that has already been imprinted on the left-hand side. This might have caused the filigree structure of the flowfield to be compressed.

<figure><img src="image_2384.png" /><figcaption>Fig. 11: Sheet metal strip with imprinted flowfield and Feature 2 as well as red line indicating position of profile measurements (left) and depths of the indicated channels as a function of the number of strokes (right).</figcaption></figure>

<figure><img src="image_2385.png" /><figcaption>Fig. 12: Laser scanning microscope image of sheet metal strip with imprinted flowfield with red and blue lines indicating positions of profile measurements (top), profile taken on the blue line containing the flowfield (bottom left) and profile taken on the red line indicating the shape of the sheet metal strip (bottom right).</figcaption></figure>
hand side of Fig. 11, both positions yielded different depths. The flowfield on the right-hand side of the sheet metal strip, whose corresponding data is depicted in the upper graph, averaged to a depth of 0.255±0.045 mm. The other flowfield only reached a depth of 0.184±0.022 mm, averaged over all channels and investigated specimens. The reason for this difference was not attributed to different punch heights, since the position of the punches was swapped after 1000 strokes. However, this explains the slight drift in all measurements from specimen 1000 to specimen 1250. The reason for the difference between the two positions is suspected to be found in the tool setup. As depicted in Fig. 7, the punches are positioned behind one another. This is also indicated by the positions of the red lines in Fig. 11. The blue arrow indicates the feed direction of the sheet metal strip. After one stroke is conducted, the relatively small feed of the sheet metal strip causes the guide plate to press down on the flowfield that has already been imprinted on the left-hand side. This might have caused the filigree structure of the flowfield to be compressed.

<figure><img src="image_2375.png" /><figcaption>Fig. 11: Sheet metal strip with imprinted flowfield and Feature 2 as well as red line indicating position of profile measurements (left) and depths of the indicated channels as a function of the number of strokes (right).</figcaption></figure>

<figure><img src="image_2376.png" /><figcaption>Fig. 12: Laser scanning microscope image of sheet metal strip with imprinted flowfield with red and blue lines indicating positions of profile measurements (top), profile taken on the blue line containing the flowfield (bottom left) and profile taken on the red line indicating the shape of the sheet metal strip (bottom right).</figcaption></figure>
The measurements of the right-hand side flowfield reveal that not all channels, but only channels 3, 5 and 6, reached the expected depth of 0.3 mm, cf. Fig. 11. One possibility is that the larger radii towards the middle of the flowfield allowed for a less restricted material flow. However, the measurements taken with a laser scanning microscope reveal, that the metal sheet has a profiled surface, see Fig. 12. The magnitude of the bulge is the same as that of the missing depth. The measurements of channel 1 on the right-handed side, in the upper graph of Fig. 11, also indicate the aforementioned partial fracture of channel 1. The depths of the channels closest to the broken one appear to become larger after it broke off. Overall, the profile measurements show that the imprint is even better than expected by the simulation. This might be caused by the pocket effect of the EDMed surface and the resulting constant supply of lubricants in the contact area as previously discussed by [11].

Besides fractures, wear characteristics were analyzed using SEM-imaging. Fig. 13 presents the tool surface of the channel pattern after 7360 strokes. The microstructure of the EDMed as well as alteration due to tribological load and adhesion of sheet metal on the surface.

<figure><img src="image_2386.png" /><figcaption>Fig. 13: SEM images of tribologically loaded head radius of channel pattern.</figcaption></figure>

Overall, the surface only shows minor characteristics of abrasive wear. A slight texture in sliding direction can be seen on the SEM images of the head radius and flanks. The abrasive wear does not exceed a slight surface polish. Moreover, sheet metal steel was found adhering on the surface and mechanically disposed in valleys of the surface profile. The adhesion is no wear as to the definition as material loss on the technical surface. However, adhesion indicates that the tribological load on the surface was increased.

## Conclusion

The present paper has conducted one step towards a possible new manufacturing route for bipolarplates. First of all, it has shown that the previously published method to enhance the productivity of sinking EDM of cemented carbides is transferable to different cemented carbide grades. Hence, this method is a valuable tool to improve the machining of cemented carbides in general. Furthermore, it was stated that this method does not harm the surface integrity of the workpieces. The comparison of the two punches did not indicate an increased number of flaws for the alternative machining technology. In contrast, the punch which was machined with the reference technology experienced major fractures. However, the fracture of the feature is attributed to an overly ambitious geometry whereas the fracture of the flowfield channel to the occurrence of shearing and lateral movement and forces of the sheet metal, due to its thickness. On the other hand, the examination of the tribological wear has not shown any significant issues. A slight polishing of the punch and some adhesion of the sheet metal was observed. In general, these findings indicate that cemented carbide tools are suited for the filigree structures present in bipolarplates and that even areas with non-optimal surface roughness are able to withstand the
The measurements of the right-hand side flowfield reveal that not all channels, but only channels 3, 5 and 6, reached the expected depth of 0.3 mm, cf. Fig. 11. One possibility is that the larger radii towards the middle of the flowfield allowed for a less restricted material flow. However, the measurements taken with a laser scanning microscope reveal, that the metal sheet has a profiled surface, see Fig. 12. The magnitude of the bulge is the same as that of the missing depth. The measurements of channel 1 on the right-handed side, in the upper graph of Fig. 11, also indicate the aforementioned partial fracture of channel 1. The depths of the channels closest to the broken one appear to become larger after it broke off. Overall, the profile measurements show that the imprint is even better than expected by the simulation. This might be caused by the pocket effect of the EDMed surface and the resulting constant supply of lubricants in the contact area as previously discussed by [11].

Besides fractures, wear characteristics were analyzed using SEM-imaging. Fig. 13 presents the tool surface of the channel pattern after 7360 strokes. The microstructure of the EDMed as well as alteration due to tribological load and adhesion of sheet metal on the surface.

<figure><img src="image_2377.png" /><figcaption>Fig. 13: SEM images of tribologically loaded head radius of channel pattern.</figcaption></figure>

Overall, the surface only shows minor characteristics of abrasive wear. A slight texture in sliding direction can be seen on the SEM images of the head radius and flanks. The abrasive wear does not exceed a slight surface polish. Moreover, sheet metal steel was found adhering on the surface and mechanically disposed in valleys of the surface profile. The adhesion is no wear as to the definition as material loss on the technical surface. However, adhesion indicates that the tribological load on the surface was increased.

## Conclusion

The present paper has conducted one step towards a possible new manufacturing route for bipolarplates. First of all, it has shown that the previously published method to enhance the productivity of sinking EDM of cemented carbides is transferable to different cemented carbide grades. Hence, this method is a valuable tool to improve the machining of cemented carbides in general. Furthermore, it was stated that this method does not harm the surface integrity of the workpieces. The comparison of the two punches did not indicate an increased number of flaws for the alternative machining technology. In contrast, the punch which was machined with the reference technology experienced major fractures. However, the fracture of the feature is attributed to an overly ambitious geometry whereas the fracture of the flowfield channel to the occurrence of shearing and lateral movement and forces of the sheet metal, due to its thickness. On the other hand, the examination of the tribological wear has not shown any significant issues. A slight polishing of the punch and some adhesion of the sheet metal was observed. In general, these findings indicate that cemented carbide tools are suited for the filigree structures present in bipolarplates and that even areas with non-optimal surface roughness are able to withstand the
tremendous loads. Moreover, the analysis of the profile has revealed that the imprint is even better than expected by the simulation, which would be helpful to create a proper sealing of the bipolarplates and increase their degree of effectiveness.

## Acknowledgements

The authors thank the German Research Foundation (DFG) for funding the DFG project BE 2542/42-2: "Analysis of the discharge-dependent surface integrity of cemented carbide forming dies machined by sinking EDM and its influence on the tribological characteristics and the resulting fatigue behaviour" (project number 290130034) and the Bundesministerium für Wirtschaft und Klimaschutz (BMWK, Federal Ministry for Economic Affairs and Climate Action; Projectnumber 20N2204D – KeyTech2GreenPower).

## References

[1] C. Baum, H. Janssen, C. Brecher et al. (2021) The Relevance of Fuel Cells for Mobility Applications. Discussion Paper. Fraunhofer-Gesellschaft, Aachen. https://doi.org/10.24406/IPT-N-633591

[2] M. Miotti, J. Hofer, C. Bauer (2017) Integrated environmental and economic assessment of current and future fuel cell vehicles. Int J Life Cycle Assess 22:94–110. https://doi.org/10.1007/s11367-015-0986-4

[3] B. D. James, J. M. Huya-Kouadio, C. Houchins et al. (2018) Fuel Cell Vehicle Cost Analysis: Annual Progress Report - Hydrogen and Fuel Cell Programm 2017. V.E.5

[4] L. Peng, P. Yi, X. Lai (2014) Design and manufacturing of stainless steel bipolar plates for proton exchange membrane fuel cells. International Journal of Hydrogen Energy 39:21127–21153. https://doi.org/10.1016/j.ijhydene.2014.08.113

[5] U. Engel, J. Groenback, C. Hinsel (2011) Tooling solutions for challenges in cold forging. Umformtechnik

[6] T. Bergs, T. Petersen, U. Tombul et al. (2020) Analysis of the Influence of Surface Integrity of Cemented Carbides Machined by Sinking EDM on Flexural Fatigue. Procedia CIRP 87:456–461. https://doi.org/10.1016/j.procir.2020.02.096

[7] T. Petersen, U. Küpper, T. Herrig et al. (2021) Fracture Toughness and Tribological Properties of Cemented Carbides Machined by Sinking Electrical Discharge Machining. ESAFORM 2021:13. https://doi.org/10.25518/esaform21.1518

[8] T. Petersen, U. Küpper, A. Klink et al. (2022) Discharge energy based optimisation of sinking EDM of cemented carbides. Procedia CIRP 108:734–739. https://doi.org/10.1016/j.procir.2022.03.113

[9] H. Voigts, R. Hild, A. Feuerhack et al. (2021) Investigation of Failure Mechanisms of Cemented Carbide Fine Blanking Punches by Means of Process Forces and Acoustic Emission. In: Forming the Future. Springer, Cham, pp 1173–1187. https://doi.org/10.1007/978-3-030-75381-8_98

[10] D. Qiu, L. Peng, P. Yi et al. (2018) Flow channel design for metallic bipolar plates in proton exchange membrane fuel cells: Experiments. Energy Conversion and Management 174:814–823. https://doi.org/10.1016/j.enconman.2018.08.070

[11] T. Petersen, U. Küpper, T. Herrig et al. (2023) Coefficient of friction of cemented carbides machined by sinking EDM. Materials Research Proceedings 28:1765–1774. https://doi.org/10.21741/9781644902479-191
tremendous loads. Moreover, the analysis of the profile has revealed that the imprint is even better than expected by the simulation, which would be helpful to create a proper sealing of the bipolarplates and increase their degree of effectiveness.

## Acknowledgements

The authors thank the German Research Foundation (DFG) for funding the DFG project BE 2542/42-2: "Analysis of the discharge-dependent surface integrity of cemented carbide forming dies machined by sinking EDM and its influence on the tribological characteristics and the resulting fatigue behaviour" (project number 290130034) and the Bundesministerium für Wirtschaft und Klimaschutz (BMWK, Federal Ministry for Economic Affairs and Climate Action; Projectnumber 20N2204D – KeyTech2GreenPower).

## References

[1] C. Baum, H. Janssen, C. Brecher et al. (2021) The Relevance of Fuel Cells for Mobility Applications. Discussion Paper. Fraunhofer-Gesellschaft, Aachen. https://doi.org/10.24406/IPT-N-633591

[2] M. Miotti, J. Hofer, C. Bauer (2017) Integrated environmental and economic assessment of current and future fuel cell vehicles. Int J Life Cycle Assess 22:94–110. https://doi.org/10.1007/s11367-015-0986-4

[3] B. D. James, J. M. Huya-Kouadio, C. Houchins et al. (2018) Fuel Cell Vehicle Cost Analysis: Annual Progress Report - Hydrogen and Fuel Cell Programm 2017. V.E.5

[4] L. Peng, P. Yi, X. Lai (2014) Design and manufacturing of stainless steel bipolar plates for proton exchange membrane fuel cells. International Journal of Hydrogen Energy 39:21127–21153. https://doi.org/10.1016/j.ijhydene.2014.08.113

[5] U. Engel, J. Groenback, C. Hinsel (2011) Tooling solutions for challenges in cold forging. Umformtechnik

[6] T. Bergs, T. Petersen, U. Tombul et al. (2020) Analysis of the Influence of Surface Integrity of Cemented Carbides Machined by Sinking EDM on Flexural Fatigue. Procedia CIRP 87:456–461. https://doi.org/10.1016/j.procir.2020.02.096

[7] T. Petersen, U. Küpper, T. Herrig et al. (2021) Fracture Toughness and Tribological Properties of Cemented Carbides Machined by Sinking Electrical Discharge Machining. ESAFORM 2021:13. https://doi.org/10.25518/esaform21.1518

[8] T. Petersen, U. Küpper, A. Klink et al. (2022) Discharge energy based optimisation of sinking EDM of cemented carbides. Procedia CIRP 108:734–739. https://doi.org/10.1016/j.procir.2022.03.113

[9] H. Voigts, R. Hild, A. Feuerhack et al. (2021) Investigation of Failure Mechanisms of Cemented Carbide Fine Blanking Punches by Means of Process Forces and Acoustic Emission. In: Forming the Future. Springer, Cham, pp 1173–1187. https://doi.org/10.1007/978-3-030-75381-8_98

[10] D. Qiu, L. Peng, P. Yi et al. (2018) Flow channel design for metallic bipolar plates in proton exchange membrane fuel cells: Experiments. Energy Conversion and Management 174:814–823. https://doi.org/10.1016/j.enconman.2018.08.070

[11] T. Petersen, U. Küpper, T. Herrig et al. (2023) Coefficient of friction of cemented carbides machined by sinking EDM. Materials Research Proceedings 28:1765–1774. https://doi.org/10.21741/9781644902479-191
# Effect of process parameters on final geometry and quality in hot tube spinning process

GUNES Murat¹,ª*, ERDEM Nagihan¹,ᵇ* and YIGIT Mehmet¹,ᶜ

¹Repkon Machine and Tool Industry and Trade Inc., 34980 Şile, Istanbul, Turkey

**Keywords:** Hot Tube Spinning, Hot Metal Forming, Finite Element Analysis

**Abstract.** Spinning is one of the oldest metal forming processes that produce rotationally axisymmetric and seamless final products by using plates or tubes [1]. Preform can be shaped via rollers, hot or cold and it has a big advantage in terms of manufacturing complex geometries [2]. It is used in many different industries around the world. However, the wall thickness of the output (final geometry) might be uneven or the roller path might be different than expected. It is clear that all of those results have a negative impact on the quality of the process, as they make it difficult to obtain the desired final geometry. Within the context of this study, the cylindrical tube will be used in the hot spinning metal forming process. It will be investigated how the reduction ratio, the roller geometry and the initial temperature of the preform affect wall thickness of the final geometry with the help of finite element analysis software.

**Introduction**

It is commonly acknowledged that the finite element method is an effective tool for metal forming analysis since with this technique it is possible to determine the fields of displacements, stresses, and strains, as well as the distributions of loads and pressures, at every stage of the deformation and this allows the analyst to optimize the forming parameters by adjusting certain process parameters [1] such as the geometry of the dies, process temperature, friction coefficient and along with others. In this paper, the authors work with FORGE®, which is a software that can perform hot and cold forming processes.

In recent years, hot spinning has been emphasized since the process force is relatively lower, although the formability of the process is still higher to produce complex part [3]. It can be performed as either hot or cold (room temperature) process, yet hot tube spinning process looks more complicated due to the fact that it has different deformations ways at the same time in the deformation zones, such as bending, shear, thickening and thinning, compression and elongation [3]. Although hot tube spinning, mentioned in this paper, is a different process, in literature, tube spinning and flow forming are used synonymously [4], and this might cause misleading in the context of this study. In tube spinning (flow forming) process, a tubular preform rotates with a mandrel and deforms by rollers [5] while hot tube spinning, in the presented study, does not require any mandrel and is conducted at elevated temperature. A number of researchers have carried out experimental and numerical studies regarding the impact of different parameters on the spinning process. In this paper, temperature, roller geometry and, reduction ratio, known as the percentage of deformation in the thickness direction [6], will be analyzed to understand how to optimize the hot tube spinning process with the numerical outputs drawn from commercial software FORGE®.

<figure><img src="image_2387.png" /></figure>

<footnote>ªmurat.gunes@repkon.com.tr, ᵇnagihan.erdem@repkon.com.tr,</footnote>

<footnote>ᶜmehmet.yigit@repkon.com.tr</footnote>
# Effect of process parameters on final geometry and quality in hot tube spinning process

GUNES Murat¹,ª*, ERDEM Nagihan¹,ᵇ* and YIGIT Mehmet¹,ᶜ

¹Repkon Machine and Tool Industry and Trade Inc., 34980 Şile, Istanbul, Turkey

**Keywords:** Hot Tube Spinning, Hot Metal Forming, Finite Element Analysis

**Abstract.** Spinning is one of the oldest metal forming processes that produce rotationally axisymmetric and seamless final products by using plates or tubes [1]. Preform can be shaped via rollers, hot or cold and it has a big advantage in terms of manufacturing complex geometries [2]. It is used in many different industries around the world. However, the wall thickness of the output (final geometry) might be uneven or the roller path might be different than expected. It is clear that all of those results have a negative impact on the quality of the process, as they make it difficult to obtain the desired final geometry. Within the context of this study, the cylindrical tube will be used in the hot spinning metal forming process. It will be investigated how the reduction ratio, the roller geometry and the initial temperature of the preform affect wall thickness of the final geometry with the help of finite element analysis software.

**Introduction**

It is commonly acknowledged that the finite element method is an effective tool for metal forming analysis since with this technique it is possible to determine the fields of displacements, stresses, and strains, as well as the distributions of loads and pressures, at every stage of the deformation and this allows the analyst to optimize the forming parameters by adjusting certain process parameters [1] such as the geometry of the dies, process temperature, friction coefficient and along with others. In this paper, the authors work with FORGE®, which is a software that can perform hot and cold forming processes.

In recent years, hot spinning has been emphasized since the process force is relatively lower, although the formability of the process is still higher to produce complex part [3]. It can be performed as either hot or cold (room temperature) process, yet hot tube spinning process looks more complicated due to the fact that it has different deformations ways at the same time in the deformation zones, such as bending, shear, thickening and thinning, compression and elongation [3]. Although hot tube spinning, mentioned in this paper, is a different process, in literature, tube spinning and flow forming are used synonymously [4], and this might cause misleading in the context of this study. In tube spinning (flow forming) process, a tubular preform rotates with a mandrel and deforms by rollers [5] while hot tube spinning, in the presented study, does not require any mandrel and is conducted at elevated temperature. A number of researchers have carried out experimental and numerical studies regarding the impact of different parameters on the spinning process. In this paper, temperature, roller geometry and, reduction ratio, known as the percentage of deformation in the thickness direction [6], will be analyzed to understand how to optimize the hot tube spinning process with the numerical outputs drawn from commercial software FORGE®.

<figure><img src="image_2378.png" /></figure>

<footnote>ªmurat.gunes@repkon.com.tr, ᵇnagihan.erdem@repkon.com.tr,</footnote>

<footnote>ᶜmehmet.yigit@repkon.com.tr</footnote>
## Building of the Finite Element Model

Two different preform initial temperatures, three different reduction ratios and two different forming geometries were determined as analysis parameters as in the Table 1 below.

<table><caption>Table 1. Finite element analysis parameters</caption><thead><tr><th>Temperature</th><th>Reduction ratio</th><th>Forming Roller</th></tr></thead><tbody><tr><td>1100 °C (T1100)</td><td>2 mm (R2)</td><td>R15 (F1)</td></tr><tr><td rowspan="2">750 °C (T750)</td><td>4 mm (R4)</td><td>R27.5 (F2)</td></tr><tr><td>6 mm (R6)</td><td></td></tr></tbody></table>

For the finite element analysis 12 sets with different combinations have been created in FORGE® to understand the effect of the temperature, reduction ratio and geometry of the forming roller, as can be seen in the Table 2.

<table><caption>Table 2. All setups used for the finite element analysis</caption><thead><tr><th>Temperature °C</th><th>Reduction ratio (mm)</th><th>Forming Roller</th></tr></thead><tbody><tr><td>T1100</td><td>R2</td><td>F1</td></tr><tr><td>T1100</td><td>R4</td><td>F1</td></tr><tr><td>T1100</td><td>R6</td><td>F1</td></tr><tr><td>T1100</td><td>R2</td><td>F2</td></tr><tr><td>T1100</td><td>R4</td><td>F2</td></tr><tr><td>T1100</td><td>R6</td><td>F2</td></tr><tr><td>T750</td><td>R2</td><td>F1</td></tr><tr><td>T750</td><td>R4</td><td>F1</td></tr><tr><td>T750</td><td>R6</td><td>F1</td></tr><tr><td>T750</td><td>R2</td><td>F2</td></tr><tr><td>T750</td><td>R4</td><td>F2</td></tr><tr><td>T750</td><td>R6</td><td>F2</td></tr></tbody></table>

The CAD model of the simulation contains two forming rollers, one preform (workpiece) and one headstock.

### Rollers

Rollers in many different forms are used in practice. In this study, the effect of the radius on the roller geometry shown in Fig. 1 on the process outputs (thickness and force) will be examined.

The rollers were chosen as rigid bodies in the analysis, and they have fine meshes locally since the radius of the rollers has the main role in forming. The roller geometry has been analyzed in terms of their two different radii, which are 15.0 mm (F1) and 27.5 mm (F2), shown in Fig. 2a and Fig. 2b, respectively.
## Building of the Finite Element Model

Two different preform initial temperatures, three different reduction ratios and two different forming geometries were determined as analysis parameters as in the Table 1 below.

<table><caption>Table 1. Finite element analysis parameters</caption><thead><tr><td>Temperature</td><td>Reduction ratio</td><td>Forming Roller</td></tr></thead><tbody><tr><td>1100 °C (T1100)</td><td>2 mm (R2)</td><td>R15 (F1)</td></tr><tr><td rowspan="2">750 °C (T750)</td><td>4 mm (R4)</td><td>R27.5 (F2)</td></tr><tr><td>6 mm (R6)</td><td></td></tr></tbody></table>

For the finite element analysis 12 sets with different combinations have been created in FORGE® to understand the effect of the temperature, reduction ratio and geometry of the forming roller, as can be seen in the Table 2.

<table><caption>Table 2. All setups used for the finite element analysis</caption><thead><tr><td>Temperature °C</td><td>Reduction ratio (mm)</td><td>Forming Roller</td></tr></thead><tbody><tr><td>T1100</td><td>R2</td><td>F1</td></tr><tr><td>T1100</td><td>R4</td><td>F1</td></tr><tr><td>T1100</td><td>R6</td><td>F1</td></tr><tr><td>T1100</td><td>R2</td><td>F2</td></tr><tr><td>T1100</td><td>R4</td><td>F2</td></tr><tr><td>T1100</td><td>R6</td><td>F2</td></tr><tr><td>T750</td><td>R2</td><td>F1</td></tr><tr><td>T750</td><td>R4</td><td>F1</td></tr><tr><td>T750</td><td>R6</td><td>F1</td></tr><tr><td>T750</td><td>R2</td><td>F2</td></tr><tr><td>T750</td><td>R4</td><td>F2</td></tr><tr><td>T750</td><td>R6</td><td>F2</td></tr></tbody></table>

The CAD model of the simulation contains two forming rollers, one preform (workpiece) and one headstock.

### Rollers

Rollers in many different forms are used in practice. In this study, the effect of the radius on the roller geometry shown in Fig. 1 on the process outputs (thickness and force) will be examined.

The rollers were chosen as rigid bodies in the analysis, and they have fine meshes locally since the radius of the rollers has the main role in forming. The roller geometry has been analyzed in terms of their two different radii, which are 15.0 mm (F1) and 27.5 mm (F2), shown in Fig. 2a and Fig. 2b, respectively.
<figure><img src="image_2388.png" /><figcaption>Fig. 1. Forming roller geometry</figcaption></figure>

The generic press is assigned to the rollers. When using the "generic press" model, the motion characteristics will be a function of the rotational movements around two distinct axes combined with a translational motion [7].

<figure><img src="image_2389.png" /><figcaption>Fig. 2. Different forming roller geometries used a) F1: radius for 15 mm for b) F2: radius for 27.5 mm</figcaption></figure>

## Preform

The preform material is AISI 4140 (42CrMo4, DIN 1.7225) whose chemical composition is shown in the Table 3 below:

<table><caption>Table 3. AISI 4140 material properties</caption><thead><tr><td>Properties</td><td>Metric</td></tr></thead><tbody><tr><td>Tensile Strength</td><td>655 [MPa]</td></tr><tr><td>Yield Strength</td><td>415 [MPa]</td></tr><tr><td>Elastic Modulus</td><td>190-210 [GPa]</td></tr><tr><td>Poisson's Ratio</td><td>0.27-0.30</td></tr><tr><td>Density</td><td>7.85 [g/cm³]</td></tr></tbody></table>

The dimensions of the preform were chosen from the DIN 2448 standard for seamless carbon steel pipes and tubes whose outer diameter is 114.30 mm and wall thickness is 5 mm, as shown in Fig. 3 below. If the face of the preform in contact with the HS is considered as the zero point, the distance defines the axial distance of a point on the preform to zero point. The length of the preform is 300 mm. However, forming will start at distance 100 mm. In other words forming length is 200 mm, as shown in Fig. 3.
<figure><img src="image_2379.png" /><figcaption>Fig. 1. Forming roller geometry</figcaption></figure>

The generic press is assigned to the rollers. When using the "generic press" model, the motion characteristics will be a function of the rotational movements around two distinct axes combined with a translational motion [7].

<figure><img src="image_2380.png" /><figcaption>Fig. 2. Different forming roller geometries used a) F1: radius for 15 mm for b) F2: radius for 27.5 mm</figcaption></figure>

## Preform

The preform material is AISI 4140 (42CrMo4, DIN 1.7225) whose chemical composition is shown in the Table 3 below:

<table><caption>Table 3. AISI 4140 material properties</caption><thead><tr><td>Properties</td><td>Metric</td></tr></thead><tbody><tr><td>Tensile Strength</td><td>655 [MPa]</td></tr><tr><td>Yield Strength</td><td>415 [MPa]</td></tr><tr><td>Elastic Modulus</td><td>190-210 [GPa]</td></tr><tr><td>Poisson's Ratio</td><td>0.27-0.30</td></tr><tr><td>Density</td><td>7.85 [g/cm³]</td></tr></tbody></table>

The dimensions of the preform were chosen from the DIN 2448 standard for seamless carbon steel pipes and tubes whose outer diameter is 114.30 mm and wall thickness is 5 mm, as shown in Fig. 3 below. If the face of the preform in contact with the HS is considered as the zero point, the distance defines the axial distance of a point on the preform to zero point. The length of the preform is 300 mm. However, forming will start at distance 100 mm. In other words forming length is 200 mm, as shown in Fig. 3.
<figure><img src="image_2390.png" /><figcaption>Fig. 3. Preform geometry</figcaption></figure>

In this study, the FORGE® material library was used to define material properties. The Hansel-Spittel law applies to describe the viscoplastic behavior of the material. It is temperature-dependent and takes strain hardening or softening phenomena into account. It also takes the strain rate of the material into account as it is represented in Eq. 1:

$$ \sigma = A e^{m_1 T} T^{m_2} \varepsilon^{m_3} e^{\frac{m_4}{\varepsilon}} (1 + \varepsilon)^{m_5 T} e^{m_7 \varepsilon} \dot{\varepsilon}^{m_3} \dot{\varepsilon}^{m_8 T} \quad (1) $$

where σ is stress, ε is strain, Ḍ is strain rate, T is the temperature given in Celsius m1 and m9 define the material's sensitivity to temperature, m5 term coupling temperature and strain, m8 term coupling temperature and strain rate, m2, m4, and m7 define the material's sensitivity to strain, m3 depends on the material's sensitivity to the strain rate.

For the preform temperature (temperature parameter), a gradient transition is preferred instead of a sharp change from 1100 °C and 750 °C to room temperature, as shown in Fig. 4 below.

<figure><img src="image_2391.png" /><figcaption>Fig. 4. Preform temperature distribution in finite element analysis</figcaption></figure>

As in the actual hot tube spinning process, the preform has one degrees of freedom, which is the rotational movement around the z-axis. The forming rollers can move along the z-axis and y-axis while they are also rotating around the z-axis as shown in Fig. 5. Coulomb limited Tresca friction model which is considered the frictional shear stress at the part-tool interface is a fraction of the material's maximum shear [7] was used between the rollers and the preform, and water-graphite was selected as the lubricant. The angular speed of the preform is 200 rpm.
<figure><img src="image_2381.png" /><figcaption>Fig. 3. Preform geometry</figcaption></figure>

In this study, the FORGE® material library was used to define material properties. The Hansel-Spittel law applies to describe the viscoplastic behavior of the material. It is temperature-dependent and takes strain hardening or softening phenomena into account. It also takes the strain rate of the material into account as it is represented in Eq. 1:

$$ \sigma = A e^{m_1 T} T^{m_2} \varepsilon^{m_3} e^{\frac{m_4}{\varepsilon}} (1 + \varepsilon)^{m_5 T} e^{m_7 \varepsilon} \dot{\varepsilon}^{m_3} \dot{\varepsilon}^{m_8 T} \quad (1) $$

where σ is stress, ε is strain, ˙ε is strain rate, T is the temperature given in Celsius m1 and m9 define the material's sensitivity to temperature, m5 term coupling temperature and strain, m8 term coupling temperature and strain rate, m2, m4, and m7 define the material's sensitivity to strain, m3 depends on the material's sensitivity to the strain rate.

For the preform temperature (temperature parameter), a gradient transition is preferred instead of a sharp change from 1100 °C and 750 °C to room temperature, as shown in Fig. 4 below.

<figure><img src="image_2382.png" /><figcaption>Fig. 4. Preform temperature distribution in finite element analysis</figcaption></figure>

As in the actual hot tube spinning process, the preform has one degrees of freedom, which is the rotational movement around the z-axis. The forming rollers can move along the z-axis and y-axis while they are also rotating around the z-axis as shown in Fig. 5. Coulomb limited Tresca friction model which is considered the frictional shear stress at the part-tool interface is a fraction of the material's maximum shear [7] was used between the rollers and the preform, and water-graphite was selected as the lubricant. The angular speed of the preform is 200 rpm.
<figure><img src="image_2392.png" /><figcaption>Fig. 5. Coordinate system of the model and Roller Path</figcaption></figure>

As a default element type of FORGE® P1, the tetrahedron element type is used. To decrease the computational time mesh refinement was applied on the preform and the forming rollers. The shape of the edge of the rollers, which contacts the preform in real processes, was created, and these areas are meshed finely regenerating meshes shown in Fig. 6 below.

<figure><img src="image_2393.png" /><figcaption>Fig. 6. Preform initial meshing</figcaption></figure>

To determine the roller velocity in the direction of the y and z axis, circular interpolation was used. After the hot spinning process step design is made, the roller movements are transferred to the CNC controller via the G code (RS-274) structure. Since the center of rotation of the rollers follows an arc-shaped direction, G02 or G03 circular interpolation is used [7,8]. This code structure defines the starting point, end point, radius and constant tangential velocity of the arc. In circular interpolation, the simultaneous motion of two axes generates a circular arc at a constant tangential velocity or feed rate $V_0$. The axial velocities satisfy the Eq. 2 and Eq. 3 below:

$$V_x(t) = V_0 \sin \theta(t) \qquad (2)$$

$$V_y(t) = V_0 \cos \theta(t) \qquad (3)$$

where

$\theta(t) = \left(\frac{V_0}{R}\right) t$ and R is the radius of the circular arc.

For the different reduction ratios (2 mm, 4 mm, and 6 mm) which will be used in the simulation, axial velocities were determined with a constant tangential velocity of 1600 mm/min (26.67 mm/sec) shown in Fig. 7. After all velocity values are drawn they are added to the FORGE® press definition.
<figure><img src="image_2383.png" /><figcaption>Fig. 5. Coordinate system of the model and Roller Path</figcaption></figure>

As a default element type of FORGE® P1, the tetrahedron element type is used. To decrease the computational time mesh refinement was applied on the preform and the forming rollers. The shape of the edge of the rollers, which contacts the preform in real processes, was created, and these areas are meshed finely regenerating meshes shown in Fig. 6 below.

<figure><img src="image_2384.png" /><figcaption>Fig. 6. Preform initial meshing</figcaption></figure>

To determine the roller velocity in the direction of the y and z axis, circular interpolation was used. After the hot spinning process step design is made, the roller movements are transferred to the CNC controller via the G code (RS-274) structure. Since the center of rotation of the rollers follows an arc-shaped direction, G02 or G03 circular interpolation is used [7,8]. This code structure defines the starting point, end point, radius and constant tangential velocity of the arc. In circular interpolation, the simultaneous motion of two axes generates a circular arc at a constant tangential velocity or feed rate $V_0$. The axial velocities satisfy the Eq. 2 and Eq. 3 below:

$$V_x(t) = V_0 \sin \theta(t) \qquad (2)$$

$$V_y(t) = V_0 \cos \theta(t) \qquad (3)$$

where

$\theta(t) = \left(\frac{V_0}{R}\right) t$ and R is the radius of the circular arc.

For the different reduction ratios (2 mm, 4 mm, and 6 mm) which will be used in the simulation, axial velocities were determined with a constant tangential velocity of 1600 mm/min (26.67 mm/sec) shown in Fig. 7. After all velocity values are drawn they are added to the FORGE® press definition.
<figure><img src="image_2394.png" /><figcaption>Fig. 7. Roller velocity in the direction of y and z</figcaption></figure>

## Results and Discussion

The simulation analysis is performed to obtain any relationship between wall thickness after the hot spinning process and the parameters, which are temperature, roller geometry and the reduction ratio. With the context of this purpose, the impact of these parameters will be interpreted. Before evaluating the effect of the parameters on the wall thickness, it ought to be noted that temperature has a major influence on the process force in the direction of the y-axis. When temperature increases, the related force decreases by more than half, as can be seen in Fig. 8.

<figure><img src="image_2395.png" /><figcaption>Fig. 8. Effect of the temperature on the roller force in direction of y-axis</figcaption></figure>

While roller geometry has no effect on the process force shown in Fig. 9, the reduction ratio affects the force by 15-30 %, as shown in Fig. 10.

<figure><img src="image_2396.png" /><figcaption>Fig. 9. Effect of the roller geometry on the roller force in direction of y-axis</figcaption></figure>
<figure><img src="image_2385.png" /><figcaption>Fig. 7. Roller velocity in the direction of y and z</figcaption></figure>

## Results and Discussion

The simulation analysis is performed to obtain any relationship between wall thickness after the hot spinning process and the parameters, which are temperature, roller geometry and the reduction ratio. With the context of this purpose, the impact of these parameters will be interpreted. Before evaluating the effect of the parameters on the wall thickness, it ought to be noted that temperature has a major influence on the process force in the direction of the y-axis. When temperature increases, the related force decreases by more than half, as can be seen in Fig. 8.

<figure><img src="image_2386.png" /><figcaption>Fig. 8. Effect of the temperature on the roller force in direction of y-axis</figcaption></figure>

While roller geometry has no effect on the process force shown in Fig. 9, the reduction ratio affects the force by 15-30 %, as shown in Fig. 10.

<figure><img src="image_2387.png" /><figcaption>Fig. 9. Effect of the roller geometry on the roller force in direction of y-axis</figcaption></figure>
<figure><img src="image_2397.png" /><figcaption>Fig. 10. Effect of the reduction ratio on the roller force in the direction of the y-axis
Effect of Parameters on Wall Thickness</figcaption></figure>

1. Temperature

In order to understand the effect of temperature, analyzes with the same reduction ratio and cylinder geometry yet different temperature values were taken as it is shown in Fig. 11.

<figure><img src="image_2398.png" /><figcaption>Fig. 11. Effect of the temperature on the wall thickness</figcaption></figure>

When the wall thickness is examined in Fig. 11, it is observed that it increases up to a certain point and then becomes thinner. This is valid for all analysis combinations. Accordingly, it can be concluded that changing the reduction ratio, temperature and roller geometry changes the thickening rate but does not change the effective thickness orientation.

If the analysis pairs using the F1 roller, where only the preform temperature was changed, were evaluated, it was seen that increasing the initial temperature reduced the wall thickness thickening. No similar results were observed in the analyses performed with the F2 roller.
<figure><img src="image_2388.png" /><figcaption>Fig. 10. Effect of the reduction ratio on the roller force in the direction of the y-axis
Effect of Parameters on Wall Thickness</figcaption></figure>

1. Temperature

In order to understand the effect of temperature, analyzes with the same reduction ratio and cylinder geometry yet different temperature values were taken as it is shown in Fig. 11.

<figure><img src="image_2389.png" /><figcaption>Fig. 11. Effect of the temperature on the wall thickness</figcaption></figure>

When the wall thickness is examined in Fig. 11, it is observed that it increases up to a certain point and then becomes thinner. This is valid for all analysis combinations. Accordingly, it can be concluded that changing the reduction ratio, temperature and roller geometry changes the thickening rate but does not change the effective thickness orientation.

If the analysis pairs using the F1 roller, where only the preform temperature was changed, were evaluated, it was seen that increasing the initial temperature reduced the wall thickness thickening. No similar results were observed in the analyses performed with the F2 roller.
## 2. Reduction Ratio

In order to understand the effect of reduction ratio, analyzes with the same temperature and cylinder geometry yet different reduction ratios were taken as in Fig. 12.

<figure><img src="image_2399.png" /><figcaption>Fig. 12. Influence of reduction ratio on the wall thickness</figcaption></figure>

When the analysis pairs where only the reduction ratio changed were evaluated as shown in Fig. 12, it was seen that the thickening was greatest in cases where the reduction ratio was highest. In addition to thickening, it was observed that the external form of the formed part was different from the path followed by the roller, and deviations occurred, as represented in Fig. 13. As the reduction ratio increases, this amount of deviation also increases. Deviation from ideal geometry affects the dimensional tolerances of the final part. Considering that dimensional tolerances are a quality parameter, it may not be possible to meet quality requirements in processes with high reduction rates. To meet the dimensional quality requirements, it is recommended that a multi-step process be preferred and the reduction ratios of the final steps be smaller.

<figure><img src="image_2400.png" /><figcaption>Fig. 13. Amount of the deviation for T750 R2 F1</figcaption></figure>
## 2. Reduction Ratio

In order to understand the effect of reduction ratio, analyzes with the same temperature and cylinder geometry yet different reduction ratios were taken as in Fig. 12.

<figure><img src="image_2390.png" /><figcaption>Fig. 12. Influence of reduction ratio on the wall thickness</figcaption></figure>

When the analysis pairs where only the reduction ratio changed were evaluated as shown in Fig. 12, it was seen that the thickening was greatest in cases where the reduction ratio was highest. In addition to thickening, it was observed that the external form of the formed part was different from the path followed by the roller, and deviations occurred, as represented in Fig. 13. As the reduction ratio increases, this amount of deviation also increases. Deviation from ideal geometry affects the dimensional tolerances of the final part. Considering that dimensional tolerances are a quality parameter, it may not be possible to meet quality requirements in processes with high reduction rates. To meet the dimensional quality requirements, it is recommended that a multi-step process be preferred and the reduction ratios of the final steps be smaller.

<figure><img src="image_2391.png" /><figcaption>Fig. 13. Amount of the deviation for T750 R2 F1</figcaption></figure>
### 3. Roller Geometry

In order to understand the effect of roller geometry, analyzes with the same temperature and reduction ratio yet different roller geometry were taken, as shown in Fig. 14. When the pairs of analyzes where only the roller geometry changed were evaluated, as shown in Fig. 14. It has been observed that the roller geometry changes the thickening character at the tip. While the thickness increased again at some point when roller F1 was used, it decreased steadily when roller F2 was used.

<figure><img src="image_2401.png" /><figcaption>Fig. 14. Influence of roller geometry on the wall thickness</figcaption></figure>

### Conclusion

In this study, a well-founded finite element model was established in FORGE® to make a prediction regarding the effect of the temperature, reduction ratio and roller geometry on the wall thickness. As an auxiliary finding, the impacts of these parameters were also observed on the force in the y-axis. According to the analysis results, it was observed that as the temperature increased, the increase in wall thickness decreased, and the force in the y-axis direction decreased by almost half. Another observed parameter, the reduction ratio, has a directly proportional effect on the wall thickness. In other words, the thickening was greatest in cases where the reduction ratio was highest. As the reduction ratio increases, the amount of deviation from the ideal geometry also increases, the quality of the desired geometry is not achieved. It was observed that the roller geometry changed the thickening character in the end region of the part. If a roller with a larger forming zone radius is used, the thickness of the end zone gradually decreases.

It would be very worthwhile to test the results of the analysis with the real process. It is planned that this will definitely be done in future studies as an extended version of this research.

### References

[1] N. Alberti, L. Cannizzaro, E. L. Valvo and F. Micari, Analysis Of Metal Spinning Processes By The Adina Code, Comput Struct. 32 (1989) 517-525. https://doi.org/10.1016/0045-7949(89)90343-X
### 3. Roller Geometry

In order to understand the effect of roller geometry, analyzes with the same temperature and reduction ratio yet different roller geometry were taken, as shown in Fig. 14. When the pairs of analyzes where only the roller geometry changed were evaluated, as shown in Fig. 14. It has been observed that the roller geometry changes the thickening character at the tip. While the thickness increased again at some point when roller F1 was used, it decreased steadily when roller F2 was used.

<figure><img src="image_2392.png" /><figcaption>Fig. 14. Influence of roller geometry on the wall thickness</figcaption></figure>

### Conclusion

In this study, a well-founded finite element model was established in FORGE® to make a prediction regarding the effect of the temperature, reduction ratio and roller geometry on the wall thickness. As an auxiliary finding, the impacts of these parameters were also observed on the force in the y-axis. According to the analysis results, it was observed that as the temperature increased, the increase in wall thickness decreased, and the force in the y-axis direction decreased by almost half. Another observed parameter, the reduction ratio, has a directly proportional effect on the wall thickness. In other words, the thickening was greatest in cases where the reduction ratio was highest. As the reduction ratio increases, the amount of deviation from the ideal geometry also increases, the quality of the desired geometry is not achieved. It was observed that the roller geometry changed the thickening character in the end region of the part. If a roller with a larger forming zone radius is used, the thickness of the end zone gradually decreases.

It would be very worthwhile to test the results of the analysis with the real process. It is planned that this will definitely be done in future studies as an extended version of this research.

### References

[1] N. Alberti, L. Cannizzaro, E. L. Valvo and F. Micari, Analysis Of Metal Spinning Processes By The Adina Code, Comput Struct. 32 (1989) 517-525. https://doi.org/10.1016/0045-7949(89)90343-X
[2] M. Zhan, H. Yang, J. Guo, X. Wang, Review on hot spinning for difficult to deform lightweight metals, TNMSC 25 (2005) 1732-1743. https://doi.org/10.1016/S1003-6326(15)63778-5

[3] H. Zoghi, A.F. Arezoodar, M. Sayeaftabi, Enhanced finite element analysis of material deformation and strain distribution in the spinning of 42CrMo steel tubes at elevated temperature, Mater. Des. 47 (2013) 234-242. https://doi.org/10.1016/j.matdes.2012.11.049

[4] M.S.J. Hashmi, Hot Tube-Forming, in: S. Hashmi, C.J. Van Tyne, G.F. Batalha, B. Yilbas (Eds), Comprehensive Materials Processing, Elsevier, Colorado, 2014, pp. 326.

[5] F.A. Hua, Y.N. Zhang, M.H. Guo, D.Y. Guo, W.H. Tong, Z.Q. Hu, Three-dimensional finite element analysis of tube spinning, J. Mater. Process. 168 (2005) 68-74. https://doi.org/10.1016/j.jmatprotec.2004.10.014

[6] M. Mutlu, A. Ozsoy, T.O. Fenercioglu, A. Karakas, M. Baydogan, Effect of reduction ratio in flow forming process on microstructure and mechanical properties of a 6082 Al alloy. In: Proceedings of the 26th International ESAFORM Conference on Material Forming: ESAFORM 2023, 2023, pp.1015. https://doi.org/10.21741/9781644902479-111

[7] FORGE® Documentation- Part 5 Process Data, 2018.

[8] O. Masory and Y. Koren, Reference-Word Circular Interpolators for CNC Systems, J.Eng.Ind. 104 (1982). https://doi.org/10.1115/1.3185849

[9] K. Erkorkmaz, Y. Altintas, High-speed CNC system design. Part I: jerk limited trajectory generation and quintic spline interpolation Int. J. Mach. Tools Manuf. 41 (2001). https://doi.org/10.1016/S0890-6955(01)00002-5
[2] M. Zhan, H. Yang, J. Guo, X. Wang, Review on hot spinning for difficult to deform lightweight metals, TNMSC 25 (2005) 1732-1743. https://doi.org/10.1016/S1003-6326(15)63778-5

[3] H. Zoghi, A.F. Arezoodar, M. Sayeaftabi, Enhanced finite element analysis of material deformation and strain distribution in the spinning of 42CrMo steel tubes at elevated temperature, Mater. Des. 47 (2013) 234-242. https://doi.org/10.1016/j.matdes.2012.11.049

[4] M.S.J. Hashmi, Hot Tube-Forming, in: S. Hashmi, C.J. Van Tyne, G.F. Batalha, B. Yilbas (Eds), Comprehensive Materials Processing, Elsevier, Colorado, 2014, pp. 326.

[5] F.A. Hua, Y.N. Zhang, M.H. Guo, D.Y. Guo, W.H. Tong, Z.Q. Hu, Three-dimensional finite element analysis of tube spinning, J. Mater. Process. 168 (2005) 68-74. https://doi.org/10.1016/j.jmatprotec.2004.10.014

[6] M. Mutlu, A. Ozsoy, T.O. Fenercioglu, A. Karakas, M. Baydogan, Effect of reduction ratio in flow forming process on microstructure and mechanical properties of a 6082 Al alloy. In: Proceedings of the 26th International ESAFORM Conference on Material Forming: ESAFORM 2023, 2023, pp.1015. https://doi.org/10.21741/9781644902479-111

[7] FORGE® Documentation- Part 5 Process Data, 2018.

[8] O. Masory and Y. Koren, Reference-Word Circular Interpolators for CNC Systems, J.Eng.Ind. 104 (1982). https://doi.org/10.1115/1.3185849

[9] K. Erkorkmaz, Y. Altintas, High-speed CNC system design. Part I: jerk limited trajectory generation and quintic spline interpolation Int. J. Mach. Tools Manuf. 41 (2001). https://doi.org/10.1016/S0890-6955(01)00002-5
# Optimizing powder compaction for enhanced relative density: Insights from multi-particle finite element simulations and genetic algorithm

GHORBANI-MENGHARI Hossein<fnref n="1" />,ª, KIM Hwi-Jun<fnref n="2" />,ᵇ, CHOI Hyunjoo<fnref n="3" />,ᶜ, CHA Pil-Ryung<fnref n="3" />,ᵈ and KIM Ji Hoon<fnref n="1" />,ᵉ*

**Keywords:** Powder Compaction, Element Number Per Particle, Particle Size Distribution, Multi-Particle Finite Element Modeling

**Abstract.** In this study, multi-particle finite element simulations in powder compaction were performed to analyze the effects of the size of the representative volume element (RVE), the number of elements per particle, and particle size distribution. Simulation parameters were calibrated to accurately predict the relative density of compacts derived from two types of powders. The influence of RVE size across four mixtures was examined to obtain its relationship with relative density. The impacts of particle size distribution and element number per particle were studied. The results indicate a decline in relative density with increased element size. Moreover, a genetic algorithm is employed to determine the optimum mixture composition yielding the highest relative density at 1400 MPa.

## Introduction

Powder compaction is a manufacturing process used to shape powdered materials into a desired form by applying pressure. This process is widely employed in industries such as powder metallurgy, pharmaceuticals, ceramics, and cosmetics, among others. The primary goal of powder compaction is to transform a loose powder into a solid, dense product with specific geometrical and mechanical properties.

The final characteristics of tablets are primarily influenced by stresses on the powder during compaction. Factors affecting this stress can directly alter tablet outcomes. The study by Mazel et al. [1] examined the role of friction between tooling and powder on die-wall pressure during compaction, using both flat and concave punches. Tests were conducted on microcrystalline cellulose, complemented by finite element method (FEM) simulations. Both approaches showed that increased friction leads to a rise in die-wall pressure, a finding contrasting some existing literature. For flat punches, the stress changes were influenced mainly by die-powder friction. However, with concave punches, friction alterations between punches and powder also impacted die-wall pressure, which is relevant in scenarios where friction changes due to film deposits on punches. The densification of Cu-Al mixed metal powder in double-action die compaction was studied through numerical modeling by Wang et al. [2]. Comparisons were made between single-action and double-action compaction, with the latter enhancing densification, aligning with the Van Der Zwan-Siskens equation. Different initial packing structures showed that denser setups resulted in better compacts. The compact density rose as Al content increased, but stress dropped. Conversely, with consistent Al content, increased compaction pressures raised both density and stress. The impact of friction on compaction was evident in its effect on powder flow and stress

<figure><img src="image_2402.png" /></figure>

<footnote n="1">¹School of Mechanical Engineering, Pusan National University, South Korea</footnote>

<footnote n="2">²Korea Institute of Industrial Technology, Incheon, South Korea</footnote>

<footnote n="3">³School of Advanced Materials Engineering, Kookmin University, Seoul, South Korea</footnote>

<footnote>ªen.ghorbani@gmail.com, ᵇkhj@kitech.re.kr, ᶜhyunjoo@kookmin.ac.kr,</footnote>

<footnote>ᵈcprdream@kookmin.ac.kr, ᶜkimjh@pusan.ac.kr</footnote>
# Optimizing powder compaction for enhanced relative density: Insights from multi-particle finite element simulations and genetic algorithm

GHORBANI-MENGHARI Hossein<fnref n="1" />,ª, KIM Hwi-Jun<fnref n="2" />,ᵇ, CHOI Hyunjoo<fnref n="3" />,ᶜ, CHA Pil-Ryung<fnref n="3" />,ᵈ and KIM Ji Hoon<fnref n="1" />,ᵉ*

**Keywords:** Powder Compaction, Element Number Per Particle, Particle Size Distribution, Multi-Particle Finite Element Modeling

**Abstract.** In this study, multi-particle finite element simulations in powder compaction were performed to analyze the effects of the size of the representative volume element (RVE), the number of elements per particle, and particle size distribution. Simulation parameters were calibrated to accurately predict the relative density of compacts derived from two types of powders. The influence of RVE size across four mixtures was examined to obtain its relationship with relative density. The impacts of particle size distribution and element number per particle were studied. The results indicate a decline in relative density with increased element size. Moreover, a genetic algorithm is employed to determine the optimum mixture composition yielding the highest relative density at 1400 MPa.

## Introduction

Powder compaction is a manufacturing process used to shape powdered materials into a desired form by applying pressure. This process is widely employed in industries such as powder metallurgy, pharmaceuticals, ceramics, and cosmetics, among others. The primary goal of powder compaction is to transform a loose powder into a solid, dense product with specific geometrical and mechanical properties.

The final characteristics of tablets are primarily influenced by stresses on the powder during compaction. Factors affecting this stress can directly alter tablet outcomes. The study by Mazel et al. [1] examined the role of friction between tooling and powder on die-wall pressure during compaction, using both flat and concave punches. Tests were conducted on microcrystalline cellulose, complemented by finite element method (FEM) simulations. Both approaches showed that increased friction leads to a rise in die-wall pressure, a finding contrasting some existing literature. For flat punches, the stress changes were influenced mainly by die-powder friction. However, with concave punches, friction alterations between punches and powder also impacted die-wall pressure, which is relevant in scenarios where friction changes due to film deposits on punches. The densification of Cu-Al mixed metal powder in double-action die compaction was studied through numerical modeling by Wang et al. [2]. Comparisons were made between single-action and double-action compaction, with the latter enhancing densification, aligning with the Van Der Zwan-Siskens equation. Different initial packing structures showed that denser setups resulted in better compacts. The compact density rose as Al content increased, but stress dropped. Conversely, with consistent Al content, increased compaction pressures raised both density and stress. The impact of friction on compaction was evident in its effect on powder flow and stress

<figure><img src="image_2393.png" /></figure>

<footnote n="1">¹School of Mechanical Engineering, Pusan National University, South Korea</footnote>

<footnote n="2">²Korea Institute of Industrial Technology, Incheon, South Korea</footnote>

<footnote n="3">³School of Advanced Materials Engineering, Kookmin University, Seoul, South Korea</footnote>

<footnote>ªen.ghorbani@gmail.com, ᵇkhj@kitech.re.kr, ᶜhyunjoo@kookmin.ac.kr,</footnote>

<footnote>ᵈcprdream@kookmin.ac.kr, ᶜkimjh@pusan.ac.kr</footnote>
distribution. Extended dwell times aided densification, while a larger height-to-diameter ratio impeded it. Compacting powders at high loads is challenging due to particle movement and breakage. While tableting is critical in pharmaceuticals, a deeper connection between large-scale powder behavior and micro-properties is sought. Using a compaction simulator, Cabiscol et al. [3] studied the behavior of limestone powders under different pressures. The compatibility of limestone was analyzed using the traditional Heckel model and the newer Wünsch model, with the latter showing enhanced adaptability. High-pressure results revealed that the role of large particle interlocking, evident at lower pressures, diminishes. Notably, tablet strength remained consistent for particles with a median size below 10 µm, but for larger particles, strength reduced as size increased. Binder jetting additive manufacturing (AM) offers a cost-effective way to create intricate metal parts, but achieving total density without notable sintering shrinkage is challenging. To address this, Bai et al. [4] studied bimodal powder mixtures in binder jetting copper. The results showed that compared to using uniform fine powders, bimodal mixtures increased packing density by 8.2% and sintered density by 4.0% while improving flowability by 10.5% and reducing sintering shrinkage by 6.4%.

In this study, multi-particle finite element simulations in powder compaction were performed to analyze the effects of the size of the representative volume element (RVE), the number of elements per particle, and particle size distribution. Simulation parameters were calibrated to accurately predict the relative density of compacts derived from two types of powders. The influence of RVE size across four mixtures was examined to obtain its relationship with relative density. The impacts of particle size distribution and element number per particle were studied. The results indicate a decline in relative density with increased element size.

## Material Property

In this study, the powder under investigation consists of Fe-Si-Al-P. The modeling approach involves implementing a parametric Python script to generate random particles and simulate the powder compaction process. The simulation is conducted using the dynamic explicit analysis, allowing for a comprehensive exploration of the compaction dynamics. Compression testing was executed using the Instron 8501 Servo Hydraulic Machine, which utilizes a servo-regulated hydraulic mechanism to exert compressive forces on specimens. Table 1 illustrates the material characteristics of this model.

<table><caption>Table 1 Material properties.</caption><thead><tr><td>Parameter</td><td>Value</td></tr></thead><tbody><tr><td>Elastic modulus [GPa]</td><td>170</td></tr><tr><td>Poisson's ratio</td><td>0.3</td></tr><tr><td>ρ (kg/m³)</td><td>7,130</td></tr></tbody></table>

To accurately capture the material behavior during compaction, the stress-strain curve is fitted using a power law Swift hardening model, expressed as $\sigma=K(\varepsilon+\varepsilon_0)m$. The power law Swift parameters are set as follows: K (material constant) is assigned a value of 1380, m (strain hardening exponent) is determined to be 0.245, and $\varepsilon_0$ (initial strain) is specified as 0.0354. Fig. 1 illustrates the Stress-strain curve of the Fe-Si-Al-P alloy powder.
distribution. Extended dwell times aided densification, while a larger height-to-diameter ratio impeded it. Compacting powders at high loads is challenging due to particle movement and breakage. While tableting is critical in pharmaceuticals, a deeper connection between large-scale powder behavior and micro-properties is sought. Using a compaction simulator, Cabiscol et al. [3] studied the behavior of limestone powders under different pressures. The compatibility of limestone was analyzed using the traditional Heckel model and the newer Wünsch model, with the latter showing enhanced adaptability. High-pressure results revealed that the role of large particle interlocking, evident at lower pressures, diminishes. Notably, tablet strength remained consistent for particles with a median size below 10 µm, but for larger particles, strength reduced as size increased. Binder jetting additive manufacturing (AM) offers a cost-effective way to create intricate metal parts, but achieving total density without notable sintering shrinkage is challenging. To address this, Bai et al. [4] studied bimodal powder mixtures in binder jetting copper. The results showed that compared to using uniform fine powders, bimodal mixtures increased packing density by 8.2% and sintered density by 4.0% while improving flowability by 10.5% and reducing sintering shrinkage by 6.4%.

In this study, multi-particle finite element simulations in powder compaction were performed to analyze the effects of the size of the representative volume element (RVE), the number of elements per particle, and particle size distribution. Simulation parameters were calibrated to accurately predict the relative density of compacts derived from two types of powders. The influence of RVE size across four mixtures was examined to obtain its relationship with relative density. The impacts of particle size distribution and element number per particle were studied. The results indicate a decline in relative density with increased element size.

## Material Property

In this study, the powder under investigation consists of Fe-Si-Al-P. The modeling approach involves implementing a parametric Python script to generate random particles and simulate the powder compaction process. The simulation is conducted using the dynamic explicit analysis, allowing for a comprehensive exploration of the compaction dynamics. Compression testing was executed using the Instron 8501 Servo Hydraulic Machine, which utilizes a servo-regulated hydraulic mechanism to exert compressive forces on specimens. Table 1 illustrates the material characteristics of this model.

<table><caption>Table 1 Material properties.</caption><thead><tr><td>Parameter</td><td>Value</td></tr></thead><tbody><tr><td>Elastic modulus [GPa]</td><td>170</td></tr><tr><td>Poisson's ratio</td><td>0.3</td></tr><tr><td>ρ (kg/m³)</td><td>7,130</td></tr></tbody></table>

To accurately capture the material behavior during compaction, the stress-strain curve is fitted using a power law Swift hardening model, expressed as $\sigma=K(\varepsilon+\varepsilon_0)m$. The power law Swift parameters are set as follows: K (material constant) is assigned a value of 1380, m (strain hardening exponent) is determined to be 0.245, and $\varepsilon_0$ (initial strain) is specified as 0.0354. Fig. 1 illustrates the Stress-strain curve of the Fe-Si-Al-P alloy powder.
<figure><img src="image_2403.png" /><figcaption>Fig. 1 Stress-strain curve of the Fe-Si-Al-P alloy powder</figcaption></figure>

## Multi-Particle Finite Element Model

In this approach, each particle is individually depicted as a solid component and is discretized using the four-node tetrahedral elements (C3D4). Providing a fine mesh is important for thoroughly analysing particle behavior during compaction. Meanwhile, the punch and die are other components represented as discrete 3D rigid bodies.

Generating the RVE is a crucial step in ensuring the accuracy of the simulation. A customized MATLAB code [5,6], complemented by a Python script, was developed. The goal was to create a realistic representation of each particle fully encapsulated by the die. The initial relative density of the RVE is set at 40 percent. This parameter is essential for accurately reflecting the initial state of the powder and has implications for the subsequent compaction behavior. Furthermore, special attention is given to the composition of the RVE. It is essential that the RVE composition accurately mirrors the proportions of fine to coarse particles in the actual powder system. Achieving this accurate representation is crucial for capturing the diverse characteristics of the powder mixture and simulating the compaction process with a high degree of fidelity. Fig. 2 shows the generated particles and die.

<figure><img src="image_2404.png" /><figcaption>Fig. 2 3D MPFEM model implemented in Abaqus for [75,25] mixture</figcaption></figure>
<figure><img src="image_2394.png" /><figcaption>Fig. 1 Stress-strain curve of the Fe-Si-Al-P alloy powder</figcaption></figure>

## Multi-Particle Finite Element Model

In this approach, each particle is individually depicted as a solid component and is discretized using the four-node tetrahedral elements (C3D4). Providing a fine mesh is important for thoroughly analysing particle behavior during compaction. Meanwhile, the punch and die are other components represented as discrete 3D rigid bodies.

Generating the RVE is a crucial step in ensuring the accuracy of the simulation. A customized MATLAB code [5,6], complemented by a Python script, was developed. The goal was to create a realistic representation of each particle fully encapsulated by the die. The initial relative density of the RVE is set at 40 percent. This parameter is essential for accurately reflecting the initial state of the powder and has implications for the subsequent compaction behavior. Furthermore, special attention is given to the composition of the RVE. It is essential that the RVE composition accurately mirrors the proportions of fine to coarse particles in the actual powder system. Achieving this accurate representation is crucial for capturing the diverse characteristics of the powder mixture and simulating the compaction process with a high degree of fidelity. Fig. 2 shows the generated particles and die.

<figure><img src="image_2395.png" /><figcaption>Fig. 2 3D MPFEM model implemented in Abaqus for [75,25] mixture</figcaption></figure>
## Results and Discussion

Influence of element size on the relative density. In examining the influence of the number of elements per particle on the relative density within the context of powder compaction, a 3D linear tetrahedral element (C3D4) Finite Element (FE) model was employed. The powder system, characterized by a particle diameter of 125.5 µm, was simulated within an RVE measuring 1200x1200x1200 µm³ and comprising 507 particles. A range of element sizes, determined by varying seed sizes, was investigated, leading to distinct elements per particle (Elem/Particle) ratios and corresponding edge lengths. Notably, as the Elem/Particle increased—from 26 for a seed size of 1000 µm to 2252 for a seed size of 15 µm—a discernible trend emerged: a consistent rise in relative density. This suggests that finer discretization, reflected in higher Elem/Particle ratios, plays a crucial role in enhancing the accuracy of the simulation, enabling a more precise representation of the powder compaction process. These findings underscore the importance of carefully selecting the element size to balance computational efficiency, and ensuring the accuracy of results in simulations of powder compaction is crucial. Fig. 3 depicts the influence of element/particle on the relative density.

<figure><img src="image_2405.png" /><figcaption>Fig. 3 The influence of elements/particles on the relative density.</figcaption></figure>

### Effect of RVE Size on the Relative Density.
The influence of the RVE size on the relative density was analyzed. RVE sizes of 240, 305, and 350 micrometers for the [50,50] mixture were considered. Fig. 4 illustrates the impact of RVE size on relative density for the [50,50] mixtures. Increasing the RVE size results in a corresponding increase in initial density. However, it is essential to note that while this variation becomes more significant at lower pressures, the difference decreases and becomes relatively insignificant in areas where higher pressures are applied.

### Effect of Particle Size Distribution on the Relative Density.
To investigate the influence of particle size distribution on relative density, experimental tests with four distinct mixtures were conducted: [0,100], [25,75], [75,25], and [100,0], where the first number represents the percentage of fine particles and the second denotes the percentage of coarse particles. Fig. 5 demonstrates the relationship between pressure and relative density for various particle size distributions. As evident from Fig. 5, the relative density of pure particles—fine or coarse—is notably lower than the rest of the mixtures. Among the mixtures, the [0,100] composition exhibits the lowest relative density, surpassing even the [100,0] mixture. Notably, when comparing the [75,25] and [25,75] mixtures,
## Results and Discussion

Influence of element size on the relative density. In examining the influence of the number of elements per particle on the relative density within the context of powder compaction, a 3D linear tetrahedral element (C3D4) Finite Element (FE) model was employed. The powder system, characterized by a particle diameter of 125.5 µm, was simulated within an RVE measuring 1200x1200x1200 µm³ and comprising 507 particles. A range of element sizes, determined by varying seed sizes, was investigated, leading to distinct elements per particle (Elem/Particle) ratios and corresponding edge lengths. Notably, as the Elem/Particle increased—from 26 for a seed size of 1000 µm to 2252 for a seed size of 15 µm—a discernible trend emerged: a consistent rise in relative density. This suggests that finer discretization, reflected in higher Elem/Particle ratios, plays a crucial role in enhancing the accuracy of the simulation, enabling a more precise representation of the powder compaction process. These findings underscore the importance of carefully selecting the element size to balance computational efficiency, and ensuring the accuracy of results in simulations of powder compaction is crucial. Fig. 3 depicts the influence of element/particle on the relative density.

<figure><img src="image_2396.png" /><figcaption>Fig. 3 The influence of elements/particles on the relative density.</figcaption></figure>

### Effect of RVE Size on the Relative Density.
The influence of the RVE size on the relative density was analyzed. RVE sizes of 240, 305, and 350 micrometers for the [50,50] mixture were considered. Fig. 4 illustrates the impact of RVE size on relative density for the [50,50] mixtures. Increasing the RVE size results in a corresponding increase in initial density. However, it is essential to note that while this variation becomes more significant at lower pressures, the difference decreases and becomes relatively insignificant in areas where higher pressures are applied.

### Effect of Particle Size Distribution on the Relative Density.
To investigate the influence of particle size distribution on relative density, experimental tests with four distinct mixtures were conducted: [0,100], [25,75], [75,25], and [100,0], where the first number represents the percentage of fine particles and the second denotes the percentage of coarse particles. Fig. 5 demonstrates the relationship between pressure and relative density for various particle size distributions. As evident from Fig. 5, the relative density of pure particles—fine or coarse—is notably lower than the rest of the mixtures. Among the mixtures, the [0,100] composition exhibits the lowest relative density, surpassing even the [100,0] mixture. Notably, when comparing the [75,25] and [25,75] mixtures,
[75,25] stands out with the highest relative density, particularly at lower pressure levels. However, as pressure increases, the distinction in relative density between [75,25] and [25,75] diminishes, becoming relatively insignificant. This behavior is intricately tied to the plastic deformation of particles, with coarse particles experiencing less plastic deformation than fine particles. The result is greater elastic recovery, especially at lower pressures, contributing to the observed variations in relative density across different mixtures.

<figure><img src="image_2406.png" /><figcaption>Fig. 4 Effect of RVE size on the relative density for [50,50]</figcaption></figure>

Fig. 6 illustrates a relative density comparison between EXP and MPFEM at 1600 MPa. The results obtained from MPFEM closely align with the experimental findings, demonstrating a similar trend. This consistency in trends between the simulation and experimental data reinforces the reliability of the MPFEM results in capturing the relative density behavior during loading at 1600 MPa.

Finding the Optimum Mixture with the highest Relative Density Using a Genetic Algorithm. Genetic Algorithms (GAs), inspired by natural selection and genetics, offer a robust approach to solving complex optimization problems [7]. The GA procedure involves several key steps:

1. Initial Population:

* Generate a set of potential solutions represented by chromosomes, $P=\{p_1, p_2, ..., p_s\}$, consisting of random real values.

2. Evaluation (Fitness Calculation):

* Define a fitness function, $g(P)$, to assess the performance of each chromosome in the population based on the optimization objective.

3. Selection:

* Arrange chromosomes by their fitness values.

* Select two parents for crossover and mutation based on their fitness.
[75,25] stands out with the highest relative density, particularly at lower pressure levels. However, as pressure increases, the distinction in relative density between [75,25] and [25,75] diminishes, becoming relatively insignificant. This behavior is intricately tied to the plastic deformation of particles, with coarse particles experiencing less plastic deformation than fine particles. The result is greater elastic recovery, especially at lower pressures, contributing to the observed variations in relative density across different mixtures.

<figure><img src="image_2397.png" /><figcaption>Fig. 4 Effect of RVE size on the relative density for [50,50]</figcaption></figure>

Fig. 6 illustrates a relative density comparison between EXP and MPFEM at 1600 MPa. The results obtained from MPFEM closely align with the experimental findings, demonstrating a similar trend. This consistency in trends between the simulation and experimental data reinforces the reliability of the MPFEM results in capturing the relative density behavior during loading at 1600 MPa.

Finding the Optimum Mixture with the highest Relative Density Using a Genetic Algorithm. Genetic Algorithms (GAs), inspired by natural selection and genetics, offer a robust approach to solving complex optimization problems [7]. The GA procedure involves several key steps:

1. Initial Population:

* Generate a set of potential solutions represented by chromosomes, $P=\{p_1, p_2, ..., p_s\}$, consisting of random real values.

2. Evaluation (Fitness Calculation):

* Define a fitness function, $g(P)$, to assess the performance of each chromosome in the population based on the optimization objective.

3. Selection:

* Arrange chromosomes by their fitness values.

* Select two parents for crossover and mutation based on their fitness.
4. Genetic Operators:

* Create new chromosomes or offspring (C₁ and C₂), from the selected parents using genetic operators.

5. Crossover:

* Exchange information between two parents to generate new offspring.

6. Mutation:

* Introduce changes to the genes of the crossed offspring chromosomes.

This iterative process continues until a convergence criterion is met or a specified number of generations is reached. The GA aims to explore the solution space efficiently, providing a robust methodology for determining the optimum mixture composition that yields the highest relative density at 1400 MPa.

$$ 
\begin{array}{l}
\text{maximize} \quad F(M) = \text{Obj}_s(m_1, m_2) \\
0 \le m_1 \le 100 \\
0 \le m_2 \le 100 \\
m_1 + m_2 = 100
\end{array}
$$

where $m_1$ and $m_2$ represent the volume fractions of the two powders, $F(M)$ is the maximum relative density as the objective function at 1400 MPa. Table 2 illustrates the optimum mixture with the highest relative density at 1400 MPa.

<figure><img src="image_2407.png" /><figcaption>Fig. 5 Relative density for various particle size distributions</figcaption></figure>
4. Genetic Operators:

* Create new chromosomes or offspring (C₁ and C₂), from the selected parents using genetic operators.

5. Crossover:

* Exchange information between two parents to generate new offspring.

6. Mutation:

* Introduce changes to the genes of the crossed offspring chromosomes.

This iterative process continues until a convergence criterion is met or a specified number of generations is reached. The GA aims to explore the solution space efficiently, providing a robust methodology for determining the optimum mixture composition that yields the highest relative density at 1400 MPa.

$$ 
\begin{array}{l}
\text{maximize} \quad F(M) = \text{Obj}_s(m_1, m_2) \\
0 \le m_1 \le 100 \\
0 \le m_2 \le 100 \\
m_1 + m_2 = 100
\end{array}
$$

where $m_1$ and $m_2$ represent the volume fractions of the two powders, $F(M)$ is the maximum relative density as the objective function at 1400 MPa. Table 2 illustrates the optimum mixture with the highest relative density at 1400 MPa.

<figure><img src="image_2398.png" /><figcaption>Fig. 5 Relative density for various particle size distributions</figcaption></figure>
<figure><img src="image_2408.png" /><figcaption>Fig. 6 Relative density comparison within EXP and MPFEM at 1600 MPa</figcaption></figure>

<table><caption>Table 2 The optimum mixture with the highest relative density at 1400 MPa</caption><tr><td>Mixture [Fine, Coarse]</td><td>[66.74, 33.26]</td></tr><tr><td>Relative Density</td><td>85.6107</td></tr></table>

## Conclusion

This study explored the intricate details of powder compaction through advanced simulation techniques and optimization algorithms. The objective was to fine-tune simulation parameters to accurately predict the relative density of compacts derived from two types of powders. Key findings include the significant impact of element size, RVE size, and particle distribution on relative density. The delicate balance between these factors is crucial in accurately predicting the compaction process. Using a genetic algorithm to identify the optimum mixture composition for maximum relative density at 1400 MPa further demonstrates the potential for advanced optimization techniques in powder compaction studies. These insights contribute to the evolving understanding of powder compaction dynamics and offer valuable guidance for optimizing the process in various industrial applications.

## References

[1] V. Mazel, H. Diarra, P. Tchoreloff, Effect of friction between powder and tooling on the die-wall pressure evolution during tableting: Experimental and numerical results for flat and concave punches, Int. J. Pharma, 554 (2019) 116-124. https://doi.org/10.1016/j.ijpharm.2018.11.003

[2] W. Wang, H. Qi, P. Liu, Y. Zhao, H. Chang, Numerical simulation of densification of cu-al mixed metal powder during axial compaction, Int. J. Metals. 8 (2018) 537. https://doi.org/10.3390/met8070537

[3] R. Cabiscol, H. Shi, I. Wünsch, V. Magnanimo, J.H. Finke, S. Luding, A. Kwade, Effect of particle size on powder compaction and tablet strength using limestone, Int. J. Adv. Powder. Tech. 31 (2020) 1280-1289. https://doi.org/10.1016/j.apt.2019.12.033
<figure><img src="image_2399.png" /><figcaption>Fig. 6 Relative density comparison within EXP and MPFEM at 1600 MPa</figcaption></figure>

<table><caption>Table 2 The optimum mixture with the highest relative density at 1400 MPa</caption><tr><td>Mixture [Fine, Coarse]</td><td>[66.74, 33.26]</td></tr><tr><td>Relative Density</td><td>85.6107</td></tr></table>

## Conclusion

This study explored the intricate details of powder compaction through advanced simulation techniques and optimization algorithms. The objective was to fine-tune simulation parameters to accurately predict the relative density of compacts derived from two types of powders. Key findings include the significant impact of element size, RVE size, and particle distribution on relative density. The delicate balance between these factors is crucial in accurately predicting the compaction process. Using a genetic algorithm to identify the optimum mixture composition for maximum relative density at 1400 MPa further demonstrates the potential for advanced optimization techniques in powder compaction studies. These insights contribute to the evolving understanding of powder compaction dynamics and offer valuable guidance for optimizing the process in various industrial applications.

## References

[1] V. Mazel, H. Diarra, P. Tchoreloff, Effect of friction between powder and tooling on the die-wall pressure evolution during tableting: Experimental and numerical results for flat and concave punches, Int. J. Pharma, 554 (2019) 116-124. https://doi.org/10.1016/j.ijpharm.2018.11.003

[2] W. Wang, H. Qi, P. Liu, Y. Zhao, H. Chang, Numerical simulation of densification of cu-al mixed metal powder during axial compaction, Int. J. Metals. 8 (2018) 537. https://doi.org/10.3390/met8070537

[3] R. Cabiscol, H. Shi, I. Wünsch, V. Magnanimo, J.H. Finke, S. Luding, A. Kwade, Effect of particle size on powder compaction and tablet strength using limestone, Int. J. Adv. Powder. Tech. 31 (2020) 1280-1289. https://doi.org/10.1016/j.apt.2019.12.033
[4] Y. Bai, G. Wagner, C.B. Williams, Effect of particle size distribution on powder packing and sintering in binder jetting additive manufacturing of metals, Int. J. Manuf. Sci. Eng. 139 (2017) 081019. https://doi.org/10.1115/1.4036640

[5] P. Kahhal, H. Ghorbani-Menghari, H.J. Kim, H. Choi, P.R. Cha, J.H. Kim, Metaheuristic optimization of powder size distribution in powder forming process using multi-particle finite element method coupled with artificial neural network and genetic algorithm, J. Mater. Trans. 64 (2023) 2648-2655. https://doi.org/10.2320/matertrans.MT-MI2022006

[6] P. Kahhal, J. Jung, Y.C. Hur, Y.H. Moon, J.H. Kim, Analysis of powder compaction process using multi-particle finite element method, J. Mater. Trans. 63 (2022) 1576-1582. https://doi.org/10.2320/matertrans.MT-MB2022012

[7] R.B. Carmona-Paredes, R. Domínguez-Mora, M.L. Arganis-Juárez, M.L., E. Juan-Diego, R. Mendoza-Ramírez, E. Carrizosa-Elizondo, use of evolutionary computation and guide curves to optimize the operating policies of a reservoir system established to supply drinking water, J. Appl. Water. Sci. 13 (2023) 2. https://doi.org/10.1007/s13201-022-01807-z
[4] Y. Bai, G. Wagner, C.B. Williams, Effect of particle size distribution on powder packing and sintering in binder jetting additive manufacturing of metals, Int. J. Manuf. Sci. Eng. 139 (2017) 081019. https://doi.org/10.1115/1.4036640

[5] P. Kahhal, H. Ghorbani-Menghari, H.J. Kim, H. Choi, P.R. Cha, J.H. Kim, Metaheuristic optimization of powder size distribution in powder forming process using multi-particle finite element method coupled with artificial neural network and genetic algorithm, J. Mater. Trans. 64 (2023) 2648-2655. https://doi.org/10.2320/matertrans.MT-MI2022006

[6] P. Kahhal, J. Jung, Y.C. Hur, Y.H. Moon, J.H. Kim, Analysis of powder compaction process using multi-particle finite element method, J. Mater. Trans. 63 (2022) 1576-1582. https://doi.org/10.2320/matertrans.MT-MB2022012

[7] R.B. Carmona-Paredes, R. Domínguez-Mora, M.L. Arganis-Juárez, M.L., E. Juan-Diego, R. Mendoza-Ramírez, E. Carrizosa-Elizondo, use of evolutionary computation and guide curves to optimize the operating policies of a reservoir system established to supply drinking water, J. Appl. Water. Sci. 13 (2023) 2. https://doi.org/10.1007/s13201-022-01807-z
Polymer processing and
thermomechanical properties

<figure><img src="image_2409.png" /></figure>
Polymer processing and
thermomechanical properties

<figure><img src="image_2400.png" /></figure>
# Utilizing thermal imaging for non-destructive thermoformability assessment in vacuum-air pressure thermoforming of plastic-coated paperboards

AFSHARIANTORGHABEH Sanaz¹,ᵃ*, KÄRKI Timo¹,ᵇ and LEMINEN Ville¹,ᶜ

¹Lappeenranta-Lahti University of Technology (LUT), Mechanical Engineering, Yliopistonkatu 34, 53850 Lappeenranta, Finland

ᵃsanaz.afshariantorghabeh@lut.fi, ᵇtimo.karki@lut.fi, ᶜville.leminen@lut.fi

**Keywords:** Thermoforming, Thermoformability, Plastic-Coated Paperboard, Shape Conformability, Thermal Analysis

**Abstract.** This study addressed the limitations of traditional post-production analyses in refining thermoforming operation by employing non-destructive, real-time thermal analysis, specifically employing thermal imaging. The focus was on assessing the thermoformability of plastic-coated paperboards, a recent area of interest in manufacturing. Three paperboards underwent vacuum and air pressure thermoforming, with continuous temperature monitoring. Findings revealed correlations between the temperature distributions, the thermal profiles, and the material shape formability. Direct analysis of the thermal images enabled accurate measurement of contact areas between the mold and material. Furthermore, temperature profiles were closely related to shape profiles, particularly concerning the depth, which might be due to exothermic response of the studied materials during the induced stretching process.

**Introduction**

A thermoforming process stands out as a widely used manufacturing technique that allows the rapid shaping of a variety of products by using air pressure, vacuum, or mechanical force [1]. The thermoforming of fiber-based materials has recently gained increased attention as a part of the global sustainability transition. Nevertheless, thermoforming such materials presents inherent challenges due to their limited stretchability [2]. Therefore, determining the rate at which these materials can assume the desired shape has emerged as a key area of study [3].

The thermoformability of fiber-based materials has previously been assessed using a variety of approaches. These methods include evaluating depth and shape conformability through three-dimensional (3D) profile analysis, utilizing techniques such as microscopy or straightforward caliper measurements [2, 4]. Even though these post-production analyses provide valuable insights for optimizing subsequent trials, relying solely on trial-and-error methodologies may make refining optimal thermoforming parameters both complex and costly. Consequently, exploring non-destructive real-time methods that can be used in-line to provide precise insights into material performance is of great value.

Using thermal analysis, specifically thermal imaging, has long been a staple method for the real-time assessment of deformations and temperatures during the processing across a variety of material structures [5, 6, 7]. In thermoforming, thermal imaging has been used to observe temperature variations during the heating phase in order to fully understand deformations and enhance factors such as optimal thickness distributions [8, 9, 10]. Further, thermal recordings were used to determine the optimal thermoforming window for thermoplastic fiber-reinforced composites based on temperature distributions [11]. For thermoforming paperboard materials, similar thermal recordings were performed to determine the areas of contact between the material and the mold during the thermoforming process to further clarify the forming depth obtained [4].

<figure><img src="image_2410.png" /></figure>

Content from this work may be used under the terms of the Creative Commons Attribution 3.0 license. Any further distribution of this work must maintain attribution to the author(s) and the title of the work, journal citation and DOI. Published under license by Materials Research Forum LLC.
# Utilizing thermal imaging for non-destructive thermoformability assessment in vacuum-air pressure thermoforming of plastic-coated paperboards

AFSHARIANTORGHABEH Sanaz¹,ᵃ*, KÄRKI Timo¹,ᵇ and LEMINEN Ville¹,ᶜ

¹Lappeenranta-Lahti University of Technology (LUT), Mechanical Engineering, Yliopistonkatu 34, 53850 Lappeenranta, Finland

ᵃsanaz.afshariantorghabeh@lut.fi, ᵇtimo.karki@lut.fi, ᶜville.leminen@lut.fi

**Keywords:** Thermoforming, Thermoformability, Plastic-Coated Paperboard, Shape Conformability, Thermal Analysis

**Abstract.** This study addressed the limitations of traditional post-production analyses in refining thermoforming operation by employing non-destructive, real-time thermal analysis, specifically employing thermal imaging. The focus was on assessing the thermoformability of plastic-coated paperboards, a recent area of interest in manufacturing. Three paperboards underwent vacuum and air pressure thermoforming, with continuous temperature monitoring. Findings revealed correlations between the temperature distributions, the thermal profiles, and the material shape formability. Direct analysis of the thermal images enabled accurate measurement of contact areas between the mold and material. Furthermore, temperature profiles were closely related to shape profiles, particularly concerning the depth, which might be due to exothermic response of the studied materials during the induced stretching process.

**Introduction**

A thermoforming process stands out as a widely used manufacturing technique that allows the rapid shaping of a variety of products by using air pressure, vacuum, or mechanical force [1]. The thermoforming of fiber-based materials has recently gained increased attention as a part of the global sustainability transition. Nevertheless, thermoforming such materials presents inherent challenges due to their limited stretchability [2]. Therefore, determining the rate at which these materials can assume the desired shape has emerged as a key area of study [3].

The thermoformability of fiber-based materials has previously been assessed using a variety of approaches. These methods include evaluating depth and shape conformability through three-dimensional (3D) profile analysis, utilizing techniques such as microscopy or straightforward caliper measurements [2, 4]. Even though these post-production analyses provide valuable insights for optimizing subsequent trials, relying solely on trial-and-error methodologies may make refining optimal thermoforming parameters both complex and costly. Consequently, exploring non-destructive real-time methods that can be used in-line to provide precise insights into material performance is of great value.

Using thermal analysis, specifically thermal imaging, has long been a staple method for the real-time assessment of deformations and temperatures during the processing across a variety of material structures [5, 6, 7]. In thermoforming, thermal imaging has been used to observe temperature variations during the heating phase in order to fully understand deformations and enhance factors such as optimal thickness distributions [8, 9, 10]. Further, thermal recordings were used to determine the optimal thermoforming window for thermoplastic fiber-reinforced composites based on temperature distributions [11]. For thermoforming paperboard materials, similar thermal recordings were performed to determine the areas of contact between the material and the mold during the thermoforming process to further clarify the forming depth obtained [4].

<figure><img src="image_2401.png" /></figure>

Content from this work may be used under the terms of the Creative Commons Attribution 3.0 license. Any further distribution of this work must maintain attribution to the author(s) and the title of the work, journal citation and DOI. Published under license by Materials Research Forum LLC.
This study aims to gain a deeper understanding of the potential utility of thermal analysis for obtaining insight into the thermoformability of plastic-coated paperboards, in particular their shape conformability. This analysis is intended to facilitate real-time optimization of the production of such materials. This objective was achieved by thermoforming three plastic-coated paperboards using vacuum and air pressure techniques, with temperature distributions being continuously recorded as the parameters of the forming process were altered. The analysis of shape conformability was approached from two perspectives: contact area and depth acquisition. A subsequent investigation sought to establish a correlation between temperature distributions, thermal profiles, and material shape conformability.

## Materials

For this study, three paperboards were selected, each with a polyethylene (PE) or polyethylene terephthalate (PET) coating. Paperboard 1 includes four alternating layers: two board layers, an internal PE layer sandwiched between the board layers, and another PE coating layer. Paperboard 2 is a PE-coated substrate, whereas Paperboard 3 PET-coated substrate. Paperboards with different structures and coatings were selected to determine whether the thermal analysis is affected by either the type of board material or the coating used. A prior study contains detailed information on the microscopic structures of the chosen materials [2].

Table 1 summarizes the properties of the materials studied. All materials were stored in a constant humidity chamber at 23°C and 80% relative humidity (RH) to ensure consistent experimental conditions. Measurements were conducted at a temperature of 23°C ± 1°C and a relative humidity of 50% ± 2% RH. Using an analytical balance (Kern & Sohn, Balingen, Germany), grammage measurements were conducted in accordance with ISO 536. The total grammage of the materials was determined, thereafter the substrate grammage was determined by subtracting the known grammage of plastic layer(s) from the measured total grammage of a material. The thickness of the material was measured using a digital micrometer (Messmer Büchel model 49-56, Veenendaal, Netherlands) as per ISO 534.

<table><caption>TABLE 1. Measured properties of materials.</caption><thead><tr><th>Properties</th><th>Paperboard 1</th><th>Paperboard 2</th><th>Paperboard 3</th></tr></thead><tbody><tr><td>Composition</td><td>PE/Paperboard/PE/Paperboard</td><td>PE/Paperboard</td><td>PET/Paperboard 474</td></tr><tr><td>Thickness [µm]</td><td>437</td><td>429</td><td>395</td></tr><tr><td>Total grammage [g/m2]</td><td>412</td><td>310.5</td><td>355</td></tr><tr><td>grammage of paperboard [g/m2]</td><td>357</td><td>280.5</td><td>40</td></tr><tr><td>grammage of plastic [g/m2]</td><td>coating layer 60 + internal layer 5</td><td>30</td><td></td></tr></tbody></table>

## Methods

The experiments were conducted on the VARIOVAC Primus thermoforming line (Zarrentin, Germany). The process involved feeding a roll of 422-mm-wide material into the machine, which was then securely clamped in the frame. Subsequently, the material was heated from the bottom until it reached a softened state. Afterward, the material was quickly moved to the forming station, where vacuum and air pressure were used to shape the material over female molds. With vacuum applied from the bottom side and air pressure applied from the top side, the materials were shaped with the stretching principle. The length of the molds was aligned with the machine direction (MD) of the materials, while the width was aligned with the cross direction (CD) of the materials. Three molds were employed on each forming step: the lateral molds shared the same geometry (Mold 1), whereas the central mold had four grooves in addition (Mold 2) (Fig. 1). Next, the formed part was moved to the cutting section of the machine for the cutting procedure.
This study aims to gain a deeper understanding of the potential utility of thermal analysis for obtaining insight into the thermoformability of plastic-coated paperboards, in particular their shape conformability. This analysis is intended to facilitate real-time optimization of the production of such materials. This objective was achieved by thermoforming three plastic-coated paperboards using vacuum and air pressure techniques, with temperature distributions being continuously recorded as the parameters of the forming process were altered. The analysis of shape conformability was approached from two perspectives: contact area and depth acquisition. A subsequent investigation sought to establish a correlation between temperature distributions, thermal profiles, and material shape conformability.

## Materials

For this study, three paperboards were selected, each with a polyethylene (PE) or polyethylene terephthalate (PET) coating. Paperboard 1 includes four alternating layers: two board layers, an internal PE layer sandwiched between the board layers, and another PE coating layer. Paperboard 2 is a PE-coated substrate, whereas Paperboard 3 PET-coated substrate. Paperboards with different structures and coatings were selected to determine whether the thermal analysis is affected by either the type of board material or the coating used. A prior study contains detailed information on the microscopic structures of the chosen materials [2].

Table 1 summarizes the properties of the materials studied. All materials were stored in a constant humidity chamber at 23°C and 80% relative humidity (RH) to ensure consistent experimental conditions. Measurements were conducted at a temperature of 23°C ± 1°C and a relative humidity of 50% ± 2% RH. Using an analytical balance (Kern & Sohn, Balingen, Germany), grammage measurements were conducted in accordance with ISO 536. The total grammage of the materials was determined, thereafter the substrate grammage was determined by subtracting the known grammage of plastic layer(s) from the measured total grammage of a material. The thickness of the material was measured using a digital micrometer (Messmer Büchel model 49-56, Veenendaal, Netherlands) as per ISO 534.

<table><caption>TABLE 1. Measured properties of materials.</caption><thead><tr><th>Properties</th><th>Paperboard 1</th><th>Paperboard 2</th><th>Paperboard 3</th></tr></thead><tbody><tr><td>Composition</td><td>PE/Paperboard/PE/Paperboard</td><td>PE/Paperboard</td><td>PET/Paperboard 474</td></tr><tr><td>Thickness [µm]</td><td>437</td><td>429</td><td>395</td></tr><tr><td>Total grammage [g/m2]</td><td>412</td><td>310.5</td><td>355</td></tr><tr><td>grammage of paperboard [g/m2]</td><td>357</td><td>280.5</td><td>40</td></tr><tr><td>grammage of plastic [g/m2]</td><td>coating layer 60 + internal layer 5</td><td>30</td><td></td></tr></tbody></table>

## Methods

The experiments were conducted on the VARIOVAC Primus thermoforming line (Zarrentin, Germany). The process involved feeding a roll of 422-mm-wide material into the machine, which was then securely clamped in the frame. Subsequently, the material was heated from the bottom until it reached a softened state. Afterward, the material was quickly moved to the forming station, where vacuum and air pressure were used to shape the material over female molds. With vacuum applied from the bottom side and air pressure applied from the top side, the materials were shaped with the stretching principle. The length of the molds was aligned with the machine direction (MD) of the materials, while the width was aligned with the cross direction (CD) of the materials. Three molds were employed on each forming step: the lateral molds shared the same geometry (Mold 1), whereas the central mold had four grooves in addition (Mold 2) (Fig. 1). Next, the formed part was moved to the cutting section of the machine for the cutting procedure.
<figure><img src="image_2411.png" /><figcaption>Fig 1. Molding unit used in thermoforming.</figcaption></figure>

Three coated paperboards were thermoformed using a preheating/forming temperature of 110°C, a preheating/forming time of 3 seconds, and a combination of vacuum and 1 bar air pressure. In discussions, these parameters were referred to as reference values. In addition, forming parameters underwent additional adjustments for Paperboard 1, which included changing the temperature to 90°C, the time to five seconds, and the pressure to two bars. Changes in each parameter were made independently, while other parameters were maintained at their reference values. A modification of the parameters for Paperboard 1 was undertaken to determine the practicability of thermal analysis when the conditions of forming are varied. It is important to note that the selection of the parameter range was informed by a previous study performed on the same materials and using the same process [2]. Three identical samples were produced in each of the trial and the average of the results were reported.

The temperature distribution of the formed samples was captured immediately after they were released from the forming chamber using a thermal camera (FLIR A8201sc, Täby, Sweden) placed above the machine. Considering the enclosed nature of the forming chamber, it was not possible to record thermal data during the forming process. The thermal camera employed had a resolution of 1024 x 1024 and a sensitivity below 20 mK, equipped with a 25 mm lens. The spectral range of the camera is 3-5 µm. In this range of wavelengths, the emissivity values for white Paperboards 1 and 3 were 0.76, while brown Paperboard 2 had an emissivity value of 0.74. These values were determined based on previous research conducted using the same spectral range for imaging moistened machine-made coated paperboards from the coated side [12]. In this study, the thermal camera's temperature detection range was set to -10°C and 150°C. Further examination of the temperature distributions was conducted using FLIR Research's IR analysis software.

In this analysis, temperature profiles are obtained over the selected areas of interest, as depicted in Fig.2. The choice of specified lines was linked to being in the deepest parts of the shape, as achieving the required depth was identified as a challenge in thermoforming coated paperboards [2]. The line 1 corresponds to the midline of the lengthwise cross direction and represents the maximum depth change of the samples. The line 2 corresponds to the midline of the first step in the shape within Mold 2. The selection of line 2 is crucial for evaluating thermoformability in plastic-coated paperboards, as this line represents the deepest part within the most demanding section for these molds which is the groove area. Next, the photos were imported into Microsoft Visio for quantitative analysis. Following the cutting and removal of the samples from the thermoforming line, a shape conformability analysis was conducted using a 3D measurement system (Keyence VR-3200, Osaka, Japan). Using this measurement, the shape profiles were determined in the same areas of interest as the temperature profiles. By using data analysis tools in Microsoft Excel, the results of the two analyses were statistically compared.
<figure><img src="image_2402.png" /><figcaption>Fig 1. Molding unit used in thermoforming.</figcaption></figure>

Three coated paperboards were thermoformed using a preheating/forming temperature of 110°C, a preheating/forming time of 3 seconds, and a combination of vacuum and 1 bar air pressure. In discussions, these parameters were referred to as reference values. In addition, forming parameters underwent additional adjustments for Paperboard 1, which included changing the temperature to 90°C, the time to five seconds, and the pressure to two bars. Changes in each parameter were made independently, while other parameters were maintained at their reference values. A modification of the parameters for Paperboard 1 was undertaken to determine the practicability of thermal analysis when the conditions of forming are varied. It is important to note that the selection of the parameter range was informed by a previous study performed on the same materials and using the same process [2]. Three identical samples were produced in each of the trial and the average of the results were reported.

The temperature distribution of the formed samples was captured immediately after they were released from the forming chamber using a thermal camera (FLIR A8201sc, Täby, Sweden) placed above the machine. Considering the enclosed nature of the forming chamber, it was not possible to record thermal data during the forming process. The thermal camera employed had a resolution of 1024 x 1024 and a sensitivity below 20 mK, equipped with a 25 mm lens. The spectral range of the camera is 3-5 µm. In this range of wavelengths, the emissivity values for white Paperboards 1 and 3 were 0.76, while brown Paperboard 2 had an emissivity value of 0.74. These values were determined based on previous research conducted using the same spectral range for imaging moistened machine-made coated paperboards from the coated side [12]. In this study, the thermal camera's temperature detection range was set to -10°C and 150°C. Further examination of the temperature distributions was conducted using FLIR Research's IR analysis software.

In this analysis, temperature profiles are obtained over the selected areas of interest, as depicted in Fig.2. The choice of specified lines was linked to being in the deepest parts of the shape, as achieving the required depth was identified as a challenge in thermoforming coated paperboards [2]. The line 1 corresponds to the midline of the lengthwise cross direction and represents the maximum depth change of the samples. The line 2 corresponds to the midline of the first step in the shape within Mold 2. The selection of line 2 is crucial for evaluating thermoformability in plastic-coated paperboards, as this line represents the deepest part within the most demanding section for these molds which is the groove area. Next, the photos were imported into Microsoft Visio for quantitative analysis. Following the cutting and removal of the samples from the thermoforming line, a shape conformability analysis was conducted using a 3D measurement system (Keyence VR-3200, Osaka, Japan). Using this measurement, the shape profiles were determined in the same areas of interest as the temperature profiles. By using data analysis tools in Microsoft Excel, the results of the two analyses were statistically compared.
<figure><img src="image_2412.png" /><figcaption>Fig 2. Regions of analysis for temperature profiles and shape profiles (the thermal image displayed pertains to Paperboard 1 formed with reference condition).</figcaption></figure>

**Results and Discussion**

*Contact Area Analysis:* The thermoforming procedure involves cooling the material upon contact with the cold mold [13]. Therefore, by identifying the cold regions within temperature distributions, one can determine the zones of contact. Particularly, in paperboard materials, there is typically a curled profile if the material fails to conform to the mold and solidifies while being cooled. This methodology was previously applied to a qualitative assessment of thermoformed plastic-coated paperboards [4]. The purpose of this section is to make this comparison quantitative, with a view to estimating the accuracy of measuring the contact area by analyzing the temperature distributions that were captured.

An illustrative instance of such analysis is depicted in Fig. 3, demonstrating Paperboard 1 formed under the reference conditions. The left side of Fig. 3 illustrates the temperature distribution across mold 1 immediately after the thermoforming of Paperboard 1. The right side of depiction in Fig. 3 illustrates the corresponding measurement lines derived from shape microscopy analysis of the formed samples. In thermal image, the coldest areas or contact regions are represented by blue or dark green hues. To enhance the visualization of temperature distributions, a broader spectrum of temperatures has been selected for display. However, to detect the area of contact, the spectrum of display can be attenuate, thereby facilitating a clearer identification process. The thermal image was imported into Microsoft Visio in its original size, where the lines of contact were measured accurately.

In microscopic analysis, the x-axis represents the length of the lines, whereas the y-axis constitutes the change in depth of the lines. The variation away from zero in the y-axis is attributed to the inherent nature of thermoforming paperboard material, which tends to exhibit curvature rather than a perfectly flat line. It is important to note that, shape measurements were conducted with the samples turned upside down on a microscope to flatten curled edges, allowing for precise depth measurements. As depicted in Fig. 3, The comparison of measurements on the left and right side of Figure 3 indicates a close alignment between thermal and microscopic measurements, with some variance likely resulting from inherent measurement errors. The significance of this observation lies in its application to fine-tuning production processes, which would allow real-time adjustments to be made based on contact area measurements, even as samples are still in the production line prior to their release.
<figure><img src="image_2403.png" /><figcaption>Fig 2. Regions of analysis for temperature profiles and shape profiles (the thermal image displayed pertains to Paperboard 1 formed with reference condition).</figcaption></figure>

## Results and Discussion

*Contact Area Analysis:* The thermoforming procedure involves cooling the material upon contact with the cold mold [13]. Therefore, by identifying the cold regions within temperature distributions, one can determine the zones of contact. Particularly, in paperboard materials, there is typically a curled profile if the material fails to conform to the mold and solidifies while being cooled. This methodology was previously applied to a qualitative assessment of thermoformed plastic-coated paperboards [4]. The purpose of this section is to make this comparison quantitative, with a view to estimating the accuracy of measuring the contact area by analyzing the temperature distributions that were captured.

An illustrative instance of such analysis is depicted in Fig. 3, demonstrating Paperboard 1 formed under the reference conditions. The left side of Fig. 3 illustrates the temperature distribution across mold 1 immediately after the thermoforming of Paperboard 1. The right side of depiction in Fig. 3 illustrates the corresponding measurement lines derived from shape microscopy analysis of the formed samples. In thermal image, the coldest areas or contact regions are represented by blue or dark green hues. To enhance the visualization of temperature distributions, a broader spectrum of temperatures has been selected for display. However, to detect the area of contact, the spectrum of display can be attenuate, thereby facilitating a clearer identification process. The thermal image was imported into Microsoft Visio in its original size, where the lines of contact were measured accurately.

In microscopic analysis, the x-axis represents the length of the lines, whereas the y-axis constitutes the change in depth of the lines. The variation away from zero in the y-axis is attributed to the inherent nature of thermoforming paperboard material, which tends to exhibit curvature rather than a perfectly flat line. It is important to note that, shape measurements were conducted with the samples turned upside down on a microscope to flatten curled edges, allowing for precise depth measurements. As depicted in Fig. 3, The comparison of measurements on the left and right side of Figure 3 indicates a close alignment between thermal and microscopic measurements, with some variance likely resulting from inherent measurement errors. The significance of this observation lies in its application to fine-tuning production processes, which would allow real-time adjustments to be made based on contact area measurements, even as samples are still in the production line prior to their release.
<figure><img src="image_2413.png" /><figcaption>Fig 3. Conducting contact area analysis using thermal imaging and microscopic measurements.</figcaption></figure>

**Depth Acquisition:** It is an inherent challenge in the thermoforming of plastic-coated paperboards to achieve desired depths during the forming process [4]. Real-time analysis of depth achievement can significantly contribute to the optimization of the process. This study observed a significant correlation between temperature changes in the materials and the rate at which they stretched, which correlated with their depth achievements. To elaborate on this observation, a thermal profile was first produced across the designated regions of interest (Fig. 2) to quantify the temperature change along the selected lines immediately following the release of the samples from the forming chamber. These data were then compared with the depth measurements obtained by shape the microscopy.

Fig. 4 depicts the methodology applied in this instance, illustrating the thermoforming process of Paperboard 1 under the reference forming conditions. In this specific example, the maximum temperature changes along the Line 1 (spanning from the edge to the edge) were found to be 55.6 degrees Celsius, calculated as the difference between maximum and minimum value in the thermal profile. Similarly, through shape profile analysis, the maximum depth change was determined to be 10.918 mm out of 15 mm depth of the mold.
<figure><img src="image_2404.png" /><figcaption>Fig 3. Conducting contact area analysis using thermal imaging and microscopic measurements.</figcaption></figure>

**Depth Acquisition:** It is an inherent challenge in the thermoforming of plastic-coated paperboards to achieve desired depths during the forming process [4]. Real-time analysis of depth achievement can significantly contribute to the optimization of the process. This study observed a significant correlation between temperature changes in the materials and the rate at which they stretched, which correlated with their depth achievements. To elaborate on this observation, a thermal profile was first produced across the designated regions of interest (Fig. 2) to quantify the temperature change along the selected lines immediately following the release of the samples from the forming chamber. These data were then compared with the depth measurements obtained by shape the microscopy.

Fig. 4 depicts the methodology applied in this instance, illustrating the thermoforming process of Paperboard 1 under the reference forming conditions. In this specific example, the maximum temperature changes along the Line 1 (spanning from the edge to the edge) were found to be 55.6 degrees Celsius, calculated as the difference between maximum and minimum value in the thermal profile. Similarly, through shape profile analysis, the maximum depth change was determined to be 10.918 mm out of 15 mm depth of the mold.
<figure><img src="image_2414.png" /><figcaption>Fig 4. An example of temperature changes measurement and depth measurement for Paperboard 1.</figcaption></figure>

The above-mentioned measurements had been performed for all the trials included in this study, and the data obtained was analyzed using regression statistics. As shown in Fig. 5, Observed temperature changes from the edge to edge of samples is significantly correlated with their corresponding depth changes. Indeed, the temperature changes closely reflect the rate at which the material is stretched, suggesting that a deeper shape tends to have a higher temperature difference (albeit without a complete linear tendency). This observation aligns with prior research on the stretching process of different types of materials, such as metal and fiber, where it was noted that stretching results in the release of heat, inducing a temperature change [14, 15]. Their findings elucidated that the stretching process exhibits an endothermic nature at small deformations, while transitioning to an exothermic characteristic at larger deformations.

The threshold between 'large' and 'small' deformation rates in thermoforming is challenging to define due to material behavior complexities. The assessment of deformation magnitude in this study was based on specific properties of the plastic-coated paperboards used. For instance, increasing the mold depth slightly beyond 15 mm—a depth that was used in this study—leads to material rupture. Similarly, a pressure of 2 bars, which was utilized in one trial, is at the upper limit of what the material can withstand without failure. Thus, within material's processing capabilities of this study, the used conditions might represent the boundary of large deformation for materials specific composition and characteristics. Therefore, it could cause the exothermic behavior in induced forming rate.
<figure><img src="image_2405.png" /><figcaption>Fig 4. An example of temperature changes measurement and depth measurement for Paperboard 1.</figcaption></figure>

The above-mentioned measurements had been performed for all the trials included in this study, and the data obtained was analyzed using regression statistics. As shown in Fig. 5, Observed temperature changes from the edge to edge of samples is significantly correlated with their corresponding depth changes. Indeed, the temperature changes closely reflect the rate at which the material is stretched, suggesting that a deeper shape tends to have a higher temperature difference (albeit without a complete linear tendency). This observation aligns with prior research on the stretching process of different types of materials, such as metal and fiber, where it was noted that stretching results in the release of heat, inducing a temperature change [14, 15]. Their findings elucidated that the stretching process exhibits an endothermic nature at small deformations, while transitioning to an exothermic characteristic at larger deformations.

The threshold between 'large' and 'small' deformation rates in thermoforming is challenging to define due to material behavior complexities. The assessment of deformation magnitude in this study was based on specific properties of the plastic-coated paperboards used. For instance, increasing the mold depth slightly beyond 15 mm—a depth that was used in this study—leads to material rupture. Similarly, a pressure of 2 bars, which was utilized in one trial, is at the upper limit of what the material can withstand without failure. Thus, within material's processing capabilities of this study, the used conditions might represent the boundary of large deformation for materials specific composition and characteristics. Therefore, it could cause the exothermic behavior in induced forming rate.
<figure><img src="image_2415.png" /></figure>

<table><caption>Fig 5. Correlation between maximum depth changes and temperature changes of studied paperboards in Line 1.</caption><thead><tr><th>Trial</th><th>Average of Sample depth (mm)</th><th>Average of Temperature change (°C)</th></tr></thead><tbody><tr><td>Paperboard 1- 90°C,3 sec, 1 bar</td><td>10,92 ± 0,55</td><td>42,57 ± 3,29</td></tr><tr><td>Paperboard 1- 110°C,3 sec, 1 bar</td><td>11,16 ± 0,28</td><td>54,83 ± 0,76</td></tr><tr><td>Paperboard 1- 110°C,3 sec, 2 bars</td><td>11,42 ± 0,25</td><td>56,13 ± 3,30</td></tr><tr><td>Paperboard 1- 110°C,5 sec, 1 bar</td><td>11,66 ± 0,06</td><td>59,97 ± 0,35</td></tr><tr><td>Paperboard 2- 110°C,3 sec, 1 bar</td><td>9,84 ± 0,5</td><td>4,47 ± 0,90</td></tr><tr><td>Paperboard 3- 110°C,3 sec, 1 bar</td><td>7,71 ± 0,2</td><td>7,70 ± 0,57</td></tr></tbody></table>

For further validation of the abovementioned correlation, the depth of the grooves was measured along the Line 2 (Fig. 2) from the shape profile, and the corresponding temperature difference was measured from the thermal profiles. A representative example of such measurement for Paperboard 1 under the reference forming conditions is shown in Fig. 6. Next, a statistical analysis of all the data points is presented in Fig. 7. The regression analysis yet revealed another strong correlation, further verifying the established relationship.

Nevertheless, from the regression coefficients displayed in Figures 5 and 7, it is evident that there is no exact linear relationship. This may be attributed to the spring-back and gradual recovery of a viscoelastic component over time, which materials like paper undergo [16]. The current shape measurements were taken on samples removed from the forming process and allowed some rest time. Due to the time gap between thermoforming and measurement, there may be a slight variation in the linear relationship between immediate temperature change after the process and depth change measured later. Notably, Line 2 shows higher regression coefficients than Line 1, possibly due to a higher rate of spring-back occurring as material contacts with mold. Analysis of temperature imaging indicates lack of contact with grooves in blue or dark green color areas; however, for Line 1 cases show material reaching bottom of mold at times. A previous study revealed contact with a cold mold being a contributing factor to spring-back in fiber-based structures [4]. Consequently, there might be a higher rate of spring-back over time in Line 1 where material contacted the mold compared to Line 2 where those not contacting the mold.

Overall, the significance of the correlations presented lies in their potential to optimize production process in real-time. Obtaining reference data that covers temperature changes, shape and depth values allows for the prediction of subsequent trials in real time. By doing so, the need for post-production analysis can be substantially reduced. It is nevertheless important to recognize that, despite the validation of this analysis focusing on depth and contact areas, other aspects of formability might still need to be examined post-production.