# Experimental characterisation for compression moulding of hybrid architecture composites using reclaimed prepreg manufacturing waste

QIAN Connie<fnref n="1" />, a*, YUAN Hao<fnref n="1" />, b, BEGUM Hasina<fnref n="1" />, c, HILL Tom<fnref n="1" />, d, GROVES Richard<fnref n="1" />, e, JOESBURY Adam<fnref n="2" />, f, HARPER Lee<fnref n="2" />, g

**Keywords:** Compression Moulding, Hybrid Architecture, Woven Prepreg, Sustainability

**Abstract.** Composites manufacturing using prepreg general a large proportion of waste from ply cutting, which usually ends up in landfill. A novel reuse route for prepreg manufacturing waste is proposed by combining chip-SMC reprocessed from the waste material and the virgin continuous fibre prepreg to create hybrid architecture composites. Experimental studies are performed to investigate the flow behaviour of prepreg chip-SMC under typical compression moulding conditions and benchmark it against a conventional SMC. Process characterisation is also performed for hybrid architecture composites to understand the critical deformation mechanisms of chip-SMC and prepreg, and the interaction between the two materials. Compression moulding trials are performed to further study the material behaviour and process characterises under realistic manufacturing conditions.

## Introduction

Composites manufacturing using prepreg typically generates 30%-50% of waste from ply cutting [1], which usually ends up in landfill, having negative impact on the environment and resulting in substantial costs. Research has shown that the CO₂ emissions associated with manufacturing virgin material is significantly higher (up to 80% [2]) than for retrieving fibres from manufacturing scrap. Developing commercially viable solutions to reclaim and reuse carbon fibre prepreg is therefore of great industrial interest, increasing material efficiency, reducing manufacturing cost, and most importantly avoiding landfill. The key to prepreg reuse is to handle the manufacturing waste “here-and-now” to eliminate further cost and emissions during transportation, developing intermediate products that fit readily within existing manufacturing processes and infrastructures.

Prepreg off-cuts typically contain fibres of a finite length, therefore any potential recycling strategy should focus on improving the processibility of this material to offset the disadvantages of the resulting discontinuous fibres, using intelligent topological design to maximise the mechanical performance. Several approaches have been explored for processing discontinuous prepreg, including chip-based sheet moulding compounds (SMC), which flow under heat and pressure to fill the mould cavity when compression moulded. Commercial examples using virgin prepreg include Hexcel's HexMC and DowAksa's Engineered Moulding Compound (EMC). This approach potentially offers the most robust way of reprocessing prepreg scraps, tolerating any size or shape of material off-cuts, utilising 100% of the waste. Existing prepreg-based SMCs usually consist of fibre lengths similar to conventional SMCs (e.g. 12.5mm, 25mm and 50mm), but the level of in-cavity flow is limited by the higher fibre content (>60% Vf compared to <40% Vf for

<figure><img src="image_598.png" /></figure>

<footnote n="1">¹Warwick Manufacturing Group, University of Warwick, Coventry, United Kingdom, CV4 7AL</footnote>

<footnote n="2">²Composites Research Group, University of Nottingham, Nottingham, United Kingdom, NG7 2GX</footnote>

<footnote>ªConnie.Qian@warwick.ac.uk, ᵇHao.Yuan@warwick.ac.uk, ᶜHasina.Begum.4@warwick.ac.uk, ᵈTom.Hill@warwick.ac.uk, ᶜRichard.E.Groves@warwick.ac.uk, ᵉAdam.Joesbury@nottingham.ac.uk, ᵈLee.Harper@nottingham.ac.uk</footnote>
# Experimental characterisation for compression moulding of hybrid architecture composites using reclaimed prepreg manufacturing waste

QIAN Connie<fnref n="1" />, a*, YUAN Hao<fnref n="1" />, b, BEGUM Hasina<fnref n="1" />, c, HILL Tom<fnref n="1" />, d, GROVES Richard<fnref n="1" />, e, JOESBURY Adam<fnref n="2" />, f, HARPER Lee<fnref n="2" />, g

**Keywords:** Compression Moulding, Hybrid Architecture, Woven Prepreg, Sustainability

**Abstract.** Composites manufacturing using prepreg general a large proportion of waste from ply cutting, which usually ends up in landfill. A novel reuse route for prepreg manufacturing waste is proposed by combining chip-SMC reprocessed from the waste material and the virgin continuous fibre prepreg to create hybrid architecture composites. Experimental studies are performed to investigate the flow behaviour of prepreg chip-SMC under typical compression moulding conditions and benchmark it against a conventional SMC. Process characterisation is also performed for hybrid architecture composites to understand the critical deformation mechanisms of chip-SMC and prepreg, and the interaction between the two materials. Compression moulding trials are performed to further study the material behaviour and process characterises under realistic manufacturing conditions.

**Introduction**

Composites manufacturing using prepreg typically generates 30%-50% of waste from ply cutting [1], which usually ends up in landfill, having negative impact on the environment and resulting in substantial costs. Research has shown that the CO₂ emissions associated with manufacturing virgin material is significantly higher (up to 80% [2]) than for retrieving fibres from manufacturing scrap. Developing commercially viable solutions to reclaim and reuse carbon fibre prepreg is therefore of great industrial interest, increasing material efficiency, reducing manufacturing cost, and most importantly avoiding landfill. The key to prepreg reuse is to handle the manufacturing waste “here-and-now” to eliminate further cost and emissions during transportation, developing intermediate products that fit readily within existing manufacturing processes and infrastructures.

Prepreg off-cuts typically contain fibres of a finite length, therefore any potential recycling strategy should focus on improving the processibility of this material to offset the disadvantages of the resulting discontinuous fibres, using intelligent topological design to maximise the mechanical performance. Several approaches have been explored for processing discontinuous prepreg, including chip-based sheet moulding compounds (SMC), which flow under heat and pressure to fill the mould cavity when compression moulded. Commercial examples using virgin prepreg include Hexcel's HexMC and DowAksa's Engineered Moulding Compound (EMC). This approach potentially offers the most robust way of reprocessing prepreg scraps, tolerating any size or shape of material off-cuts, utilising 100% of the waste. Existing prepreg-based SMCs usually consist of fibre lengths similar to conventional SMCs (e.g. 12.5mm, 25mm and 50mm), but the level of in-cavity flow is limited by the higher fibre content (>60% Vf compared to <40% Vf for

<figure><img src="image_609.png" /></figure>

<footnote n="1">¹Warwick Manufacturing Group, University of Warwick, Coventry, United Kingdom, CV4 7AL</footnote>

<footnote n="2">²Composites Research Group, University of Nottingham, Nottingham, United Kingdom, NG7 2GX</footnote>

<footnote>ªConnie.Qian@warwick.ac.uk, ᵇHao.Yuan@warwick.ac.uk, ᶜHasina.Begum.4@warwick.ac.uk, ᵈTom.Hill@warwick.ac.uk, ᶜRichard.E.Groves@warwick.ac.uk, ᵉAdam.Joesbury@nottingham.ac.uk, ᵈLee.Harper@nottingham.ac.uk</footnote>
conventional SMCs). Little effort has been made to investigate ways to engineer the flow characteristics of these materials, particularly to maximise the out of plane flow behaviour to fill complex geometries.

This paper proposes a practical solution for utilising prepreg waste, by co-moulding continuous fibre prepreg with discontinuous prepreg chips (SMC) to form composites with a hybrid architecture. A series of experimental studies have been performed to characterise the main deformation mechanisms for prepreg chip-SMC and hybrid architecture materials under typical compression moulding conditions. A bespoke squeeze flow rig has been used in this paper for all of the experimental characterisation studies. Compression mouldings of hybrid architecture composites combining virgin prepreg and chip-SMC have been conducted using a flat plaque geometry.

## Material Characterisation Methodologies

Sample Preparation. The majority of the studies in this paper have been conducted using Mitsubishi TRK501 - an epoxy based woven carbon prepreg with 12k tow size, 2x2 twill weave and 55% fibre volume fraction. Due to the interruption to material supply, an alternative carbon/epoxy supplied by DowAksa with the same tow size, weave pattern and fibre volume fraction was used for characterising the pure chip-SMC specimens. The DowAksa material is not commercially available, therefore further details of the material will not be disclosed due to confidentiality reasons. All chips were cut at 0/90° in relation to the direction of the fibres, and three different chip sizes were studied, including 12.5mm, 17.2mm, and 25mm. Lamination of loose prepreg chips was carried out by shuffling the chips within a sealed plastic envelope until the distribution of the chips was visually uniform. Envelopes containing the chips were then positioned on the conveyor belt of a Schmidt & Heinzmann compounding line, and compacted at 3 bar and 50°C inside the compaction chamber of the compounding line. Fig. 1 shows one sheet of chip-SMC consisting of 25mm square chips.

<figure><img src="image_599.png" /><figcaption>Fig. 1: A sheet of manufactured chip-SMC consisting of 25mm square chips</figcaption></figure>

Chip-SMC Characterisation. Flow characterisation for chip-SMCs was performed on a squeeze flow rig shown in Fig. 2, which was originally developed in [3] for characterising the flow behaviour for conventional SMCs. The squeeze flow rig consisted of a pair of 350mm × 350mm testing plates heated by cartridge heaters. The testing plates were attached to a die set with four guide columns to ensure parallelism of the testing plates. An extensometer was attached to the side of the testing plates for accurately measuring the displacement of the moving plate. Circular-shaped squeeze flow testing specimens with a diameter of 100 mm (Fig. 3 (a)) were stamped out from the laminated chip-SMCs using a clicker-press. Chip-SMCs of three different nominal areal masses were tested in this study: 3000gsm, 6000gsm and 8000gsm. All tests were performed at
conventional SMCs). Little effort has been made to investigate ways to engineer the flow characteristics of these materials, particularly to maximise the out of plane flow behaviour to fill complex geometries.

This paper proposes a practical solution for utilising prepreg waste, by co-moulding continuous fibre prepreg with discontinuous prepreg chips (SMC) to form composites with a hybrid architecture. A series of experimental studies have been performed to characterise the main deformation mechanisms for prepreg chip-SMC and hybrid architecture materials under typical compression moulding conditions. A bespoke squeeze flow rig has been used in this paper for all of the experimental characterisation studies. Compression mouldings of hybrid architecture composites combining virgin prepreg and chip-SMC have been conducted using a flat plaque geometry.

## Material Characterisation Methodologies

Sample Preparation. The majority of the studies in this paper have been conducted using Mitsubishi TRK501 - an epoxy based woven carbon prepreg with 12k tow size, 2x2 twill weave and 55% fibre volume fraction. Due to the interruption to material supply, an alternative carbon/epoxy supplied by DowAksa with the same tow size, weave pattern and fibre volume fraction was used for characterising the pure chip-SMC specimens. The DowAksa material is not commercially available, therefore further details of the material will not be disclosed due to confidentiality reasons. All chips were cut at 0/90° in relation to the direction of the fibres, and three different chip sizes were studied, including 12.5mm, 17.2mm, and 25mm. Lamination of loose prepreg chips was carried out by shuffling the chips within a sealed plastic envelope until the distribution of the chips was visually uniform. Envelopes containing the chips were then positioned on the conveyor belt of a Schmidt & Heinzmann compounding line, and compacted at 3 bar and 50°C inside the compaction chamber of the compounding line. Fig. 1 shows one sheet of chip-SMC consisting of 25mm square chips.

<figure><img src="image_610.png" /><figcaption>Fig. 1: A sheet of manufactured chip-SMC consisting of 25mm square chips</figcaption></figure>

Chip-SMC Characterisation. Flow characterisation for chip-SMCs was performed on a squeeze flow rig shown in Fig. 2, which was originally developed in [3] for characterising the flow behaviour for conventional SMCs. The squeeze flow rig consisted of a pair of 350mm × 350mm testing plates heated by cartridge heaters. The testing plates were attached to a die set with four guide columns to ensure parallelism of the testing plates. An extensometer was attached to the side of the testing plates for accurately measuring the displacement of the moving plate. Circular-shaped squeeze flow testing specimens with a diameter of 100 mm (Fig. 3 (a)) were stamped out from the laminated chip-SMCs using a clicker-press. Chip-SMCs of three different nominal areal masses were tested in this study: 3000gsm, 6000gsm and 8000gsm. All tests were performed at
100°C mould temperature, which was lower than the recommended moulding temperature of 150°C in order to slow down the curing of the material and provide the operator sufficient time to load the specimen and start the test. This temperature (100°C) was selected to ensure the resin viscosity remained constant. All tests were performed at a constant speed of 1 mm/s with a 200 kN force limit. At the end of the test, specimens were held between the plates at a constant cavity height for 20 minutes to develop a sufficient degree of cure before removing Fig. 3 (b)).

The flow behaviour of each specimen was assessed in terms of the compressive stress-strain response and the area increase of the specimen after the squeeze flow testing. Detailed methods for determining the stress-strain response and the area increase of the specimen can be found in [3]. VORAFUSE® EMC supplied by DowAksa was selected as a baseline SMC in this study to benchmark the flow behaviour of chip-SMCs against commercial SMCs. The baseline SMC was a UD prepreg-based SMC consisting of the same epoxy resin system as the prepreg, with 27mm – 54mm fibre length, 1048gsm areal mass and 55% fibre volume fraction.

<figure><img src="image_600.png" /><figcaption>Fig. 2: The squeeze flow rig employed for flow characterisation [3]</figcaption></figure>

<figure><img src="image_601.png" /><figcaption>Fig. 3: An 12.5 mm, 3000 gsm chip-SMC specimen (a) before and (b) after the squeeze flow test</figcaption></figure>

Hybrid Architecture Composites Characterisation. Hybrid architecture specimens were also characterised using the squeeze flow rig, where 17.5mm square chip-SMC with 3000gsm areal mass were chosen for all specimens. For these hybrid architecture tests, chip-SMC specimens of 100mm diameter were centrally located on top of two plies of square prepreg with 200mm edge length (see Fig. 4). Two different configurations were studied, as shown in Fig. 4: one with continuous prepreg in both plies, and the other had four 100mm x 100mm prepreg patches with a series of butt joints in the top ply. The smaller prepreg patches were introduced as a third type of architecture which could also be produced using prepreg offcuts, and the characterisation aimed
100°C mould temperature, which was lower than the recommended moulding temperature of 150°C in order to slow down the curing of the material and provide the operator sufficient time to load the specimen and start the test. This temperature (100°C) was selected to ensure the resin viscosity remained constant. All tests were performed at a constant speed of 1 mm/s with a 200 kN force limit. At the end of the test, specimens were held between the plates at a constant cavity height for 20 minutes to develop a sufficient degree of cure before removing Fig. 3 (b)).

The flow behaviour of each specimen was assessed in terms of the compressive stress-strain response and the area increase of the specimen after the squeeze flow testing. Detailed methods for determining the stress-strain response and the area increase of the specimen can be found in [3]. VORAFUSE® EMC supplied by DowAksa was selected as a baseline SMC in this study to benchmark the flow behaviour of chip-SMCs against commercial SMCs. The baseline SMC was a UD prepreg-based SMC consisting of the same epoxy resin system as the prepreg, with 27mm – 54mm fibre length, 1048gsm areal mass and 55% fibre volume fraction.

<figure><img src="image_611.png" /><figcaption>Fig. 2: The squeeze flow rig employed for flow characterisation [3]</figcaption></figure>

<figure><img src="image_612.png" /><figcaption>Fig. 3: An 12.5 mm, 3000 gsm chip-SMC specimen (a) before and (b) after the squeeze flow test</figcaption></figure>

Hybrid Architecture Composites Characterisation. Hybrid architecture specimens were also characterised using the squeeze flow rig, where 17.5mm square chip-SMC with 3000gsm areal mass were chosen for all specimens. For these hybrid architecture tests, chip-SMC specimens of 100mm diameter were centrally located on top of two plies of square prepreg with 200mm edge length (see Fig. 4). Two different configurations were studied, as shown in Fig. 4: one with continuous prepreg in both plies, and the other had four 100mm x 100mm prepreg patches with a series of butt joints in the top ply. The smaller prepreg patches were introduced as a third type of architecture which could also be produced using prepreg offcuts, and the characterisation aimed
to understand how they affect the potential distortion in the continuous fibre prepreg ply beneath
and the flow behaviour of the chip-SMC above.

<figure><img src="image_602.png" /><figcaption>Fig. 4: Specimen configurations for hybrid architecture composites characterisation</figcaption></figure>

Material Characterisation Results

Chip-SMC Characterisation. Fig. 5 presents the area increase after testing for each chip size and nominal areal mass, where the dotted red line indicates the area increase of the baseline SMC specimen. Area increase is a good indicator for the material's ability to flow, which is typically dominated by two deformation mechanisms: compaction and spreading of tows and dispersion of tows. The area increase measures the level of dispersion of tows. However, it should be noted that area increase cannot be directly compared between specimens of different masses, because specimen with larger mass contains more tows, which naturally leads to higher levels of dispersion. According to Figure 5, at the same nominal areal mass, specimens with smaller chips generally experienced larger area increase after testing. The 3000gsm specimens with chip sizes of 12.5mm and 17.5mm both had similar area increase to the baseline SMC (around 200%). However, as the area increase generally decreases with reduced specimen mass, it would be expected that these chip-SMCs would experience lower area increase at an areal mass similar to that of the baseline SMC (1048gsm). This could be explained by the mesoscale fibre architecture: the interlacing tows in chip-SMCs can prohibit the movement of tows, leading to lower level of fibre dispersion. It is also interesting to note that the 17.5mm chip size had 50% more area increase compared to the 25mm chip size, while, the 12.5mm chip size had only 20% more area increase than the 17.5mm chip size. This suggests the possible existence of a threshold for chip sizes below which the material's ability to flow could not be improved anymore. It should be noted that this conclusion is exclusive to square chips manufactured using the specific woven prepreg, other parameters such as chip aspect ratio and the fibre architecture in the prepreg materials can also affect the flow behaviour of chip-SMC, which will be investigated in future work.
to understand how they affect the potential distortion in the continuous fibre prepreg ply beneath
and the flow behaviour of the chip-SMC above.

<figure><img src="image_613.png" /><figcaption>Fig. 4: Specimen configurations for hybrid architecture composites characterisation</figcaption></figure>

Material Characterisation Results

Chip-SMC Characterisation. Fig. 5 presents the area increase after testing for each chip size and nominal areal mass, where the dotted red line indicates the area increase of the baseline SMC specimen. Area increase is a good indicator for the material's ability to flow, which is typically dominated by two deformation mechanisms: compaction and spreading of tows and dispersion of tows. The area increase measures the level of dispersion of tows. However, it should be noted that area increase cannot be directly compared between specimens of different masses, because specimen with larger mass contains more tows, which naturally leads to higher levels of dispersion. According to Figure 5, at the same nominal areal mass, specimens with smaller chips generally experienced larger area increase after testing. The 3000gsm specimens with chip sizes of 12.5mm and 17.5mm both had similar area increase to the baseline SMC (around 200%). However, as the area increase generally decreases with reduced specimen mass, it would be expected that these chip-SMCs would experience lower area increase at an areal mass similar to that of the baseline SMC (1048gsm). This could be explained by the mesoscale fibre architecture: the interlacing tows in chip-SMCs can prohibit the movement of tows, leading to lower level of fibre dispersion. It is also interesting to note that the 17.5mm chip size had 50% more area increase compared to the 25mm chip size, while, the 12.5mm chip size had only 20% more area increase than the 17.5mm chip size. This suggests the possible existence of a threshold for chip sizes below which the material's ability to flow could not be improved anymore. It should be noted that this conclusion is exclusive to square chips manufactured using the specific woven prepreg, other parameters such as chip aspect ratio and the fibre architecture in the prepreg materials can also affect the flow behaviour of chip-SMC, which will be investigated in future work.
<figure><img src="image_603.png" /><figcaption>Fig. 5: Comparison of specimen area increase after squeeze flow testing for chip-SMCs of different chip sizes and nominal areal masses. The red dashed line indicates the baseline SMC of 27mm-54mm fibre length and 1048gsm fibre areal mass.</figcaption></figure>

Fig. 6 shows the compressive stress-strain relationship obtained from the squeeze flow test for all chip-SMC specimens. The dotted red curve is the stress-strain result for the baseline SMC. Specimens with higher areal mass exhibited lower resistant to compressive pressures during testing, possibly because thicker specimens promote higher level of out-of-plane orientation in chips, resulting in higher level of lofting between chips and leading to more voids. Interestingly, the chip size had less effect on the material's stress-strain response, as curves for specimens with the different chip sizes but same areal mass are grouped together, although specimens with larger chip sizes tend to exhibit up to 35% higher compressive stresses at higher compressive strains. In addition, the curve for the baseline SMC sits closely between the curves for 3000 gsm chip-SMC with 12.5 mm and 17.5 mm chips where the three materials also had similar area increase as discussed above.

<figure><img src="image_604.png" /><figcaption>Fig. 6: Comparison of compressive stress-strain response obtained from squeeze flow testing for chip-SMCs of different chip sizes and nominal areal masses. The red dashed line indicates the baseline SMC of 27mm-54mm fibre length and 1048gsm fibre areal mass.</figcaption></figure>

Characterisation of Hybrid Architecture Composites. Fig. 7 (a) presents images of the hybrid architecture specimens after squeeze flow testing, where the top row shows the chip-SMC face and the bottom row shows the continuous prepreg face. Comparison to the chip-SMC face suggests that the prepreg patches restrict the flow of chip-SMC, such that the chip-SMC flow front was further ahead in the specimen without patches compared to the specimen with patches. Areas with
<figure><img src="image_614.png" /><figcaption>Fig. 5: Comparison of specimen area increase after squeeze flow testing for chip-SMCs of different chip sizes and nominal areal masses. The red dashed line indicates the baseline SMC of 27mm-54mm fibre length and 1048gsm fibre areal mass.</figcaption></figure>

Fig. 6 shows the compressive stress-strain relationship obtained from the squeeze flow test for all chip-SMC specimens. The dotted red curve is the stress-strain result for the baseline SMC. Specimens with higher areal mass exhibited lower resistant to compressive pressures during testing, possibly because thicker specimens promote higher level of out-of-plane orientation in chips, resulting in higher level of lofting between chips and leading to more voids. Interestingly, the chip size had less effect on the material's stress-strain response, as curves for specimens with the different chip sizes but same areal mass are grouped together, although specimens with larger chip sizes tend to exhibit up to 35% higher compressive stresses at higher compressive strains. In addition, the curve for the baseline SMC sits closely between the curves for 3000 gsm chip-SMC with 12.5 mm and 17.5 mm chips where the three materials also had similar area increase as discussed above.

<figure><img src="image_615.png" /><figcaption>Fig. 6: Comparison of compressive stress-strain response obtained from squeeze flow testing for chip-SMCs of different chip sizes and nominal areal masses. The red dashed line indicates the baseline SMC of 27mm-54mm fibre length and 1048gsm fibre areal mass.</figcaption></figure>

Characterisation of Hybrid Architecture Composites. Fig. 7 (a) presents images of the hybrid architecture specimens after squeeze flow testing, where the top row shows the chip-SMC face and the bottom row shows the continuous prepreg face. Comparison to the chip-SMC face suggests that the prepreg patches restrict the flow of chip-SMC, such that the chip-SMC flow front was further ahead in the specimen without patches compared to the specimen with patches. Areas with
compacted tows can be observed ahead of the flow front of chip-SMC in both specimens, although it is less obvious in the specimen without patches as the flow front of the chip-SMC reached the edges of the prepreg plies in most areas. Further research will be performed using larger prepreg plies so that this effect can be better investigated. According to a similar study where conventional SMC was used [3], tow compaction was expected to be accompanied by tow spreading in the area underneath the chip-SMC, which can be seen in the bottom prepreg ply in the prepreg-side images in Fig. 7 (a). It was also reported in [3] that tow spreading in the continuous fibre phase was primarily caused by the compressive stresses arisen from SMC flows, and the effects of friction at the interface were negligible, which was assumed to be the same case in this paper. While the prepreg patches affected the flow of chip-SMC, it had little influence on the level of distortion in the bottom prepreg ply.

<figure><img src="image_605.png" /></figure>

<figure><img src="image_606.png" /></figure>

<figure><img src="image_607.png" /></figure>

<figure><img src="image_608.png" /></figure>

<table><tr><td>Chip-SMC</td></tr><tr><td>Prepreg</td></tr><tr><td>Prepreg</td></tr></table>

<table><caption>(a)</caption><tr><td>Chip-SMC</td></tr><tr><td>Prepreg</td><td>Prepreg</td></tr><tr><td>Prepreg</td></tr></table>
compacted tows can be observed ahead of the flow front of chip-SMC in both specimens, although it is less obvious in the specimen without patches as the flow front of the chip-SMC reached the edges of the prepreg plies in most areas. Further research will be performed using larger prepreg plies so that this effect can be better investigated. According to a similar study where conventional SMC was used [3], tow compaction was expected to be accompanied by tow spreading in the area underneath the chip-SMC, which can be seen in the bottom prepreg ply in the prepreg-side images in Fig. 7 (a). It was also reported in [3] that tow spreading in the continuous fibre phase was primarily caused by the compressive stresses arisen from SMC flows, and the effects of friction at the interface were negligible, which was assumed to be the same case in this paper. While the prepreg patches affected the flow of chip-SMC, it had little influence on the level of distortion in the bottom prepreg ply.

<figure><img src="image_616.png" /></figure>

<figure><img src="image_617.png" /></figure>

<figure><img src="image_618.png" /></figure>

<figure><img src="image_619.png" /></figure>

<table><tr><td>Chip-SMC</td></tr><tr><td>Prepreg</td></tr><tr><td>Prepreg</td></tr></table>

<table><caption>(a)</caption><tr><td>Chip-SMC</td></tr><tr><td>Prepreg</td><td>Prepreg</td></tr><tr><td>Prepreg</td></tr></table>
<figure><img src="image_609.png" /><figcaption>Fig. 7: (a) Hybrid architecture specimens of two different configurations after squeeze flow test and (b) the cross-section of a highly distorted location (highlighted in (a))</figcaption></figure>

Fig. 7 (b) shows the micrograph of the cross section of an area where extremely large distortion in the bottom prepreg ply was observed, as indicated by the green boxes in Fig. 7 (a). It is interesting to note that tows remain interlaced in the chip-SMC at the same location, alongside an internal crack seen in Fig. 7 (b). A possible explanation is that the fibres were trapped at this location, prohibiting the break-up of the interlaced tows, resulting in higher local fibre volume fraction. The higher local volume fraction led to higher local compressive stresses in the chip-SMC and therefore higher levels of tow spreading in the prepreg underneath. At the same time, there was no longer sufficient resin to bind the fibres together, hence an internal crack was formed once the mould pressure was released.

## Compression Moulding of Hybrid Architecture Composites

Compression moulding of hybrid architecture composites were performed to further characterise the manufacturing process in more representative scenarios, as well as evaluate the chip-SMCs' abilities to flow out-of-plane, which could not be assessed using in-plane squeeze flow testing. All chip-SMC used in the compression moulding studies consisted of 25mm square chips. A flat plaque geometry of 200mm x 300mm x 2.5mm was chosen for the compression moulding studies.

<figure><img src="image_610.png" /><figcaption>Fig. 8: Prepreg plies and chip-SMC charges in the flat plaque mould cavity on the Langzauner press</figcaption></figure>

Flat plaques were moulded using a Langzauner 150 Tonne multifunctional fast-closing press. Centrally located rectangular initial charges were employed to introduce a 1D flow regime (Fig. 8), using an initial charge coverage of 90% or 50%. For both initial coverage sizes, 2 plies of net-shaped prepreg were used and the prepreg plies were positioned directly on the mould surface as shown in Fig. 8. It should be noted that unlike the material characterisation studies where the chip-
<figure><img src="image_620.png" /><figcaption>Fig. 7: (a) Hybrid architecture specimens of two different configurations after squeeze flow test and (b) the cross-section of a highly distorted location (highlighted in (a))</figcaption></figure>

Fig. 7 (b) shows the micrograph of the cross section of an area where extremely large distortion in the bottom prepreg ply was observed, as indicated by the green boxes in Fig. 7 (a). It is interesting to note that tows remain interlaced in the chip-SMC at the same location, alongside an internal crack seen in Fig. 7 (b). A possible explanation is that the fibres were trapped at this location, prohibiting the break-up of the interlaced tows, resulting in higher local fibre volume fraction. The higher local volume fraction led to higher local compressive stresses in the chip-SMC and therefore higher levels of tow spreading in the prepreg underneath. At the same time, there was no longer sufficient resin to bind the fibres together, hence an internal crack was formed once the mould pressure was released.

## Compression Moulding of Hybrid Architecture Composites

Compression moulding of hybrid architecture composites were performed to further characterise the manufacturing process in more representative scenarios, as well as evaluate the chip-SMCs' abilities to flow out-of-plane, which could not be assessed using in-plane squeeze flow testing. All chip-SMC used in the compression moulding studies consisted of 25mm square chips. A flat plaque geometry of 200mm x 300mm x 2.5mm was chosen for the compression moulding studies.

<figure><img src="image_621.png" /><figcaption>Fig. 8: Prepreg plies and chip-SMC charges in the flat plaque mould cavity on the Langzauner press</figcaption></figure>

Flat plaques were moulded using a Langzauner 150 Tonne multifunctional fast-closing press. Centrally located rectangular initial charges were employed to introduce a 1D flow regime (Fig. 8), using an initial charge coverage of 90% or 50%. For both initial coverage sizes, 2 plies of net-shaped prepreg were used and the prepreg plies were positioned directly on the mould surface as shown in Fig. 8. It should be noted that unlike the material characterisation studies where the chip-
SMC samples were cut from a larger sheet, the initial charges for compression mouldings were manufactured net-shaped. In addition, no compaction was performed to the initial charges. Consequently, the initial charge had high variation in thickness distributions, and strong edge effects such that the areal mass in the centre was significantly higher than around the edges (see Fig. 8).

Fig. 9 presents the images of the moulded flat plaques, where the top row shows the chip-SMC side and the bottom row shows the prepreg side. For both 90% and 50% initial coverages, the chip-SMC achieved complete filling, since no prepreg was visible from the SMC side of the images. Unbroken chips and interlaced tows can be observed in both plaques, particularly around the centre of each plaque where the initial charge was originally located. This indicates a difference in flow mechanism for the chip-SMC compared to conventional SMC. In conventional SMC, the fibre architecture characteristics do not have noticeable variations along the flow direction [4], meaning the material experiences uniform local deformations. The chip-SMC on the other hand experiences higher flow towards the edges of the initial charge, and lower flow in the centre of the initial charge.

It is also interesting to observe from the prepreg-side in Fig. 9 that the high levels of tow spreading in continuous fibres were mainly observed in the areas directly underneath the initial charges, as indicated by the red boxes. This is because the higher thicknesses and areal masses towards the centre of the initial charge led to higher resistance to compression, resulting in higher compressive stresses applied to the prepreg.

<figure><img src="image_611.png" /><figcaption>Fig. 9: Hybrid architecture flat plaques moulded from two different initial charge coverages (highlighted)</figcaption></figure>

Fig. 10 (a) presents the in-plane strain distribution in the prepreg ply for the 50% initial coverage plaque, obtained using the grid strain analysis (GSA) method proposed in [5]. The colour contour shows the equivalent strain, calculated as:

$$ \varepsilon_{eq} = (1/\sqrt{2})[(\varepsilon_x - \varepsilon_y)^2 + (3/2)\gamma_{xy}^2]^{1/2}. \qquad (1) $$
SMC samples were cut from a larger sheet, the initial charges for compression mouldings were manufactured net-shaped. In addition, no compaction was performed to the initial charges. Consequently, the initial charge had high variation in thickness distributions, and strong edge effects such that the areal mass in the centre was significantly higher than around the edges (see Fig. 8).

Fig. 9 presents the images of the moulded flat plaques, where the top row shows the chip-SMC side and the bottom row shows the prepreg side. For both 90% and 50% initial coverages, the chip-SMC achieved complete filling, since no prepreg was visible from the SMC side of the images. Unbroken chips and interlaced tows can be observed in both plaques, particularly around the centre of each plaque where the initial charge was originally located. This indicates a difference in flow mechanism for the chip-SMC compared to conventional SMC. In conventional SMC, the fibre architecture characteristics do not have noticeable variations along the flow direction [4], meaning the material experiences uniform local deformations. The chip-SMC on the other hand experiences higher flow towards the edges of the initial charge, and lower flow in the centre of the initial charge.

It is also interesting to observe from the prepreg-side in Fig. 9 that the high levels of tow spreading in continuous fibres were mainly observed in the areas directly underneath the initial charges, as indicated by the red boxes. This is because the higher thicknesses and areal masses towards the centre of the initial charge led to higher resistance to compression, resulting in higher compressive stresses applied to the prepreg.

<figure><img src="image_622.png" /><figcaption>Fig. 9: Hybrid architecture flat plaques moulded from two different initial charge coverages (highlighted)</figcaption></figure>

Fig. 10 (a) presents the in-plane strain distribution in the prepreg ply for the 50% initial coverage plaque, obtained using the grid strain analysis (GSA) method proposed in [5]. The colour contour shows the equivalent strain, calculated as:

$$ \epsilon_{eq} = (1/\sqrt{2})[(\epsilon_x - \epsilon_y)^2 + (3/2)\gamma_{xy}^2]^{1/2}. \qquad (1) $$
where $ε_x$, $ε_y$ and $γ_{xy}$ are in-plane strain components. Higher strains were mainly observed around the centre of the initial charge, which was in agreement with the higher levels of the tow spreading observed in Fig. 9. Fig. 10 (b) presents the thickness distributions in the chip-SMC layer and prepreg along the cross section indicated by the blue line in Fig. 10 (a), where the thickness values were manually measured using micrographs of the cross-section. The 0 x-coordinate in Fig. 10 (b) indicates the geometric centre of the plaque. In general, higher SMC/prepreg thickness ratios were observed in areas with higher levels of tow spreading, therefore the SMC/prepreg thickness ratios show a decreasing trend from the centre (x=0) to the edge (x=150mm) of the plaque.

<figure><img src="image_612.png" /><figcaption>Fig. 10: (a) Area change of grid elements calculated with grid strain analysis method and (b) thickness distribution of two fibre architectures (SMC and prepreg) along the dashed line indicated in (a)</figcaption></figure>

## Conclusions

Characterisation of chip-SMC materials have shown that smaller chip sizes experience greater levels of area increase during squeeze flow. However, chip-SMCs general experience less area increase compared to the baseline SMC as the interlaced tows prohibit fibre dispersion. Analysis of the compressive stress-strain relationship has suggested that the material's resistance to compression stresses decreases with increasing areal mass, but chip size has little influence on the compressive stress-strain response.
where $ε_x$, $ε_y$ and $γ_{xy}$ are in-plane strain components. Higher strains were mainly observed around the centre of the initial charge, which was in agreement with the higher levels of the tow spreading observed in Fig. 9. Fig. 10 (b) presents the thickness distributions in the chip-SMC layer and prepreg along the cross section indicated by the blue line in Fig. 10 (a), where the thickness values were manually measured using micrographs of the cross-section. The 0 x-coordinate in Fig. 10 (b) indicates the geometric centre of the plaque. In general, higher SMC/prepreg thickness ratios were observed in areas with higher levels of tow spreading, therefore the SMC/prepreg thickness ratios show a decreasing trend from the centre (x=0) to the edge (x=150mm) of the plaque.

<figure><img src="image_623.png" /><figcaption>Fig. 10: (a) Area change of grid elements calculated with grid strain analysis method and (b) thickness distribution of two fibre architectures (SMC and prepreg) along the dashed line indicated in (a)</figcaption></figure>

## Conclusions

Characterisation of chip-SMC materials have shown that smaller chip sizes experience greater levels of area increase during squeeze flow. However, chip-SMCs general experience less area increase compared to the baseline SMC as the interlaced tows prohibit fibre dispersion. Analysis of the compressive stress-strain relationship has suggested that the material's resistance to compression stresses decreases with increasing areal mass, but chip size has little influence on the compressive stress-strain response.
Characterisation of chip-SMC/prepreg hybrid architecture have shown that the flow of chip-SMC introduces tow spreading in the area underneath the SMC, and tow compaction in the area ahead of the flow front of the SMC. An additional study has been conducted to investigate the effects of introducing a third fibre architecture consisting of prepreg patches between the chip-SMC and the continuous prepreg ply. While no significant effects are observed on the level of distortions in the continuous prepreg ply, the prepreg patches experience large distortion and the flow of SMC is restricted, suggesting a reduced overall formability and mechanical properties compared to the hybridisation of chip-SMC and continuous prepreg only. Higher levels of tow spreading and internal cracks tend to appear in areas where interlaced tows remain in the chip-SMC.

Finally, compression mouldings of hybrid architecture composites have been performed using a flat plaque geometry. Variations in the fibre architecture in the chip-SMC and the level of tow spreading in the continuous prepreg have been observed using different initial coverage sizes. These are due to the higher thickness and areal mass variations in the larger initial charges, compared to the smaller specimens used in the material characterisation studies.

## Acknowledgement

The authors would like to acknowledge the financial support from the EPSRC Future Composites Manufacturing Hub (grant number EP/P006701/1) and the material donation from DowAksa.

## References

[1] I. Bianchi, A. Forcellese, M. Marconi, M. Simoncini, A. Vita, V. Castorani, Environmental impact assessment of zero waste approach for carbon fiber prepreg scraps, Sustainable Materials and Technologies. 29 (2021). https://doi.org/10.1016/j.susmat.2021.e00308

[2] G. Nilakantan, S. Nutt, Reuse and upcycling of aerospace prepreg scrap and waste, Reinforced Plastics. 59 (2015) 44-51. https://doi.org/10.1016/j.repl.2014.12.070

[3] H. Yuan, M.A. Khan, K.N. Kendall, C.C. Qian, Process Characterisation for Compression Moulding of Hybrid-architecture Composites, Proceedings of the 23rd International Conference on Composite Materials, Belfast, July 30 – August 4, 2023.

[4] C. Qian, D. Norman, M.A. Williams, K. Kendall, Experimental and numerical characterisation of fibre orientation distributions in compression moulded carbon fibre SMC, Plastics, Rubber and Composites (2022). https://doi.org/10.1080/14658011.2022.2108984

[5] H. Yuan, M. Khan, T. Peijs, C. Qian, Experimental Process Characterisation for High-rate Compression Moulding of Hybrid-architecture Composites: submitted to Composites Part A: Applied Science and Manufacturing (2023). (under review)
Characterisation of chip-SMC/prepreg hybrid architecture have shown that the flow of chip-SMC introduces tow spreading in the area underneath the SMC, and tow compaction in the area ahead of the flow front of the SMC. An additional study has been conducted to investigate the effects of introducing a third fibre architecture consisting of prepreg patches between the chip-SMC and the continuous prepreg ply. While no significant effects are observed on the level of distortions in the continuous prepreg ply, the prepreg patches experience large distortion and the flow of SMC is restricted, suggesting a reduced overall formability and mechanical properties compared to the hybridisation of chip-SMC and continuous prepreg only. Higher levels of tow spreading and internal cracks tend to appear in areas where interlaced tows remain in the chip-SMC.

Finally, compression mouldings of hybrid architecture composites have been performed using a flat plaque geometry. Variations in the fibre architecture in the chip-SMC and the level of tow spreading in the continuous prepreg have been observed using different initial coverage sizes. These are due to the higher thickness and areal mass variations in the larger initial charges, compared to the smaller specimens used in the material characterisation studies.

## Acknowledgement

The authors would like to acknowledge the financial support from the EPSRC Future Composites Manufacturing Hub (grant number EP/P006701/1) and the material donation from DowAksa.

## References

[1] I. Bianchi, A. Forcellese, M. Marconi, M. Simoncini, A. Vita, V. Castorani, Environmental impact assessment of zero waste approach for carbon fiber prepreg scraps, Sustainable Materials and Technologies. 29 (2021). https://doi.org/10.1016/j.susmat.2021.e00308

[2] G. Nilakantan, S. Nutt, Reuse and upcycling of aerospace prepreg scrap and waste, Reinforced Plastics. 59 (2015) 44-51. https://doi.org/10.1016/j.repl.2014.12.070

[3] H. Yuan, M.A. Khan, K.N. Kendall, C.C. Qian, Process Characterisation for Compression Moulding of Hybrid-architecture Composites, Proceedings of the 23rd International Conference on Composite Materials, Belfast, July 30 – August 4, 2023.

[4] C. Qian, D. Norman, M.A. Williams, K. Kendall, Experimental and numerical characterisation of fibre orientation distributions in compression moulded carbon fibre SMC, Plastics, Rubber and Composites (2022). https://doi.org/10.1080/14658011.2022.2108984

[5] H. Yuan, M. Khan, T. Peijs, C. Qian, Experimental Process Characterisation for High-rate Compression Moulding of Hybrid-architecture Composites: submitted to Composites Part A: Applied Science and Manufacturing (2023). (under review)
# Tool path strategies for single point incremental forming of fiber-reinforced thermoplastic sheets

RATH Jan-Erik¹,ª* and SCHÜPPSTUHL Thorsten¹,ᵇ

¹Hamburg University of Technology, Institute of Aircraft Production Technology, Denickestr. 17,
21073 Hamburg, Germany

ªjan-erik.rath@tuhh.de, ᵇschueppstuhl@tuhh.de

**Keywords:** Composite, Fiber Reinforced Plastic, Incremental Sheet Forming, Tool Path

**Abstract.** Continuous fiber-reinforced thermoplastics (FRTP) are gaining increasing interest as a lightweight material. However, production processes such as thermoforming rely on costly molds, making them unsuitable for individual parts or small series production. Therefore, realizing an incremental sheet forming (ISF) process for FRTP is very desirable. As direct application of ISF to FRTP is impossible, researchers placed the FRTP between two metal sheets, allowing it to slide between them during single point incremental forming (SPIF). So far, only traditional tool path strategies for metal SPIF, such as the z-level or the spiral approach, have been used, which do not consider the draping requirements of woven reinforcement fibers. To optimize the forming process and prevent the development of wrinkles or other defects in the organo sheet, we propose and test novel tool path strategies for the SPIF of an FRTP sandwiched between two metal sheets. Results show that the proposed strategies have a positive impact on part quality, as fiber orientation-based tool path approaches show less wrinkling and higher part accuracy in SPIF without support.

**Introduction**

Fiber-reinforced plastic (FRP) composites are increasingly popular due to their high strength-to-weight ratios and other properties. Among the reinforcement materials, woven fibers are frequently used because of their comparably high reinforcing effect, good handling and draping properties, and wide availability [1,2]. In recent years, the somewhat contrary demands for efficiency and flexibility of FRP production processes have increasingly been addressed by automation and robotics, especially in additive processes such as Automated Fiber Placement (AFP), robotic layup, or 3D printing [3]. Still, the hand-layup process, manually draping fiber fabrics on a mold and subsequently impregnating them with a thermoset resin, is commonly used for prototype and small series production [4]. However, fiber-reinforced thermoplastics (FRTP) are gaining more and more relevance in the market due to their superior ductility, damping properties, corrosion resistance, weldability, and reshapeability, among others [1]. As a result of the high viscosity of thermoplastic matrices, pre-impregnated and consolidated flat sheets, called organo sheets, are commonly used as a semi-finished product and formed in thermoforming processes involving at least one part-specific mold. Due to the high cost and effort to produce, maintain, and store these molds, an economical production process for prototypes and small series using organo sheets is missing [5].

One solution to this problem could be incremental sheet forming (ISF), where a metal sheet is clamped on its edges and locally deformed with a standard tool, such as a hemispherical tooltip, guided by a CNC machine or an industrial robot. In addition to single point incremental forming (SPIF), process variations, including a die or a supporting tool on the other side of the sheet, exist [6]. While the ISF of thermoplastics has been demonstrated [7], its direct application to FRTP, especially with continuous fibers, is not readily possible. While metal and thermoplastic sheets deform by stretching and thinning, the continuous fibers would break when clamped on the edges [8]. Therefore, researchers incrementally formed the FRPs together with at least one metal dummy sheet, allowing the fibers to slide relative to it. Fiorotto et al. [9] used a vacuum bag to apply a

<figure><img src="image_613.png" /></figure>
# Tool path strategies for single point incremental forming of fiber-reinforced thermoplastic sheets

RATH Jan-Erik¹,ª* and SCHÜPPSTUHL Thorsten¹,ᵇ

¹Hamburg University of Technology, Institute of Aircraft Production Technology, Denickestr. 17, 21073 Hamburg, Germany

ªjan-erik.rath@tuhh.de, ᵇschueppstuhl@tuhh.de

**Keywords:** Composite, Fiber Reinforced Plastic, Incremental Sheet Forming, Tool Path

**Abstract.** Continuous fiber-reinforced thermoplastics (FRTP) are gaining increasing interest as a lightweight material. However, production processes such as thermoforming rely on costly molds, making them unsuitable for individual parts or small series production. Therefore, realizing an incremental sheet forming (ISF) process for FRTP is very desirable. As direct application of ISF to FRTP is impossible, researchers placed the FRTP between two metal sheets, allowing it to slide between them during single point incremental forming (SPIF). So far, only traditional tool path strategies for metal SPIF, such as the z-level or the spiral approach, have been used, which do not consider the draping requirements of woven reinforcement fibers. To optimize the forming process and prevent the development of wrinkles or other defects in the organo sheet, we propose and test novel tool path strategies for the SPIF of an FRTP sandwiched between two metal sheets. Results show that the proposed strategies have a positive impact on part quality, as fiber orientation-based tool path approaches show less wrinkling and higher part accuracy in SPIF without support.

**Introduction**

Fiber-reinforced plastic (FRP) composites are increasingly popular due to their high strength-to-weight ratios and other properties. Among the reinforcement materials, woven fibers are frequently used because of their comparably high reinforcing effect, good handling and draping properties, and wide availability [1,2]. In recent years, the somewhat contrary demands for efficiency and flexibility of FRP production processes have increasingly been addressed by automation and robotics, especially in additive processes such as Automated Fiber Placement (AFP), robotic layup, or 3D printing [3]. Still, the hand-layup process, manually draping fiber fabrics on a mold and subsequently impregnating them with a thermoset resin, is commonly used for prototype and small series production [4]. However, fiber-reinforced thermoplastics (FRTP) are gaining more and more relevance in the market due to their superior ductility, damping properties, corrosion resistance, weldability, and reshapeability, among others [1]. As a result of the high viscosity of thermoplastic matrices, pre-impregnated and consolidated flat sheets, called organo sheets, are commonly used as a semi-finished product and formed in thermoforming processes involving at least one part-specific mold. Due to the high cost and effort to produce, maintain, and store these molds, an economical production process for prototypes and small series using organo sheets is missing [5].

One solution to this problem could be incremental sheet forming (ISF), where a metal sheet is clamped on its edges and locally deformed with a standard tool, such as a hemispherical tooltip, guided by a CNC machine or an industrial robot. In addition to single point incremental forming (SPIF), process variations, including a die or a supporting tool on the other side of the sheet, exist [6]. While the ISF of thermoplastics has been demonstrated [7], its direct application to FRTP, especially with continuous fibers, is not readily possible. While metal and thermoplastic sheets deform by stretching and thinning, the continuous fibers would break when clamped on the edges [8]. Therefore, researchers incrementally formed the FRPs together with at least one metal dummy sheet, allowing the fibers to slide relative to it. Fiorotto et al. [9] used a vacuum bag to apply a

<figure><img src="image_624.png" /></figure>
woven thermoset prepreg to a metal sheet, protecting the FRP surface and transferring the deformation to the flexible fabric. Xiao et al. [10] and Xu et al. [11] placed a die below the sheet layup in a two-point incremental forming process. Heated continuous fiber-reinforced organo sheets were formed between two metal dummy sheets with PTFE layers to reduce adhesion, by Al-Obaidi et al. [5,12], Emami et al. [13], and Hou et al. [8]. Conte et al. [14] proposed to use the pressure of a heated fluid to apply an FRTP onto a metal sheet during SPIF. Defects such as cracks and wrinkles were often observed, especially at higher wall angles [5,8,13]. One reason could be the tool path strategies used, which have been basic z-level or spiral contouring tool paths of metal SPIF. Hou et al. [8] were the only researchers to additionally try an inside-out tool path, starting forming in the center of the sheet in each layer, and reported on improved part quality.

While the standard tool path strategies have been proven for metal sheets, they do not account for the deformation mechanisms of woven fiber fabrics, which must be sheared to represent a 3D shape with double curvature smoothly. This shear is usually introduced by tensioning the fabric in-plane in fiber or bias direction [4], which is hard to realize with undetermined fiber movement between two metal sheets and in a dieless process. Therefore, Rath et al. [15] developed a draping strategy for the dieless forming of woven FRP and tested it in various alternative processes for forming FRP without a mold [16,17]. The strategy involves a 3D tool movement in fiber direction along the target geometry, starting at a fixed point and moving towards the edge of the fabric. As the sequence of the paths is determined according to an ascending shear angle in the fabric as projected by a simulation, shear is introduced in a determined manner.

This paper aims to adapt the draping strategy [15] to SPIF to improve the incremental forming of woven FRPs together with metal dummy sheets. Therefore, multiple novel tool path strategies taking into account both the requirements of metal forming and fabric draping are developed and tested in a setup very similar to the one realized by Al-Obaidi et al. [5,12].

## Methodology

Materials. For each forming experiment, a woven carbon fiber-reinforced organo sheet was placed between two metal sheets, separated on each side by PTFE sheets to allow the sliding of the CFRP. The fiber orientations corresponded to the edges of the sheets and were therefore also aligned with the clamping frame. The organo sheets consisted of a styrene-acrylonitrile (SAN) matrix and two layers of twill-woven carbon fibers (CF) with a basis weight of 245 g/m², resulting in a thickness of 0.6 mm and a fiber-volume content of 45 %. With a size of 210 mm x 240 mm, the CFRP was smaller than the metal sheets with 350 mm x 350 mm. On the lower side, an aluminum sheet EN AW-1050A with a thickness of 0.5 mm was used as a backing plate. On the upper side, a DC04 steel sheet of 0.8 mm thickness was used to transfer the deformation while protecting the organo sheet from direct tool contact. The PTFE sheets were of 0.05 mm thickness and 270 mm x 270 mm in size. The layup is shown in Fig. 1.

<figure><img src="image_614.png" /><figcaption>Fig. 1 Layup of organo sheet, PTFE sheets, and metal sheets, including clamping and support</figcaption></figure>

Setup. The sandwich layup of metal, PTFE, and organo sheets was clamped on top of a heating chamber made of vermiculite and calcium silicate panels. It was ensured that only the metal sheets
woven thermoset prepreg to a metal sheet, protecting the FRP surface and transferring the deformation to the flexible fabric. Xiao et al. [10] and Xu et al. [11] placed a die below the sheet layup in a two-point incremental forming process. Heated continuous fiber-reinforced organo sheets were formed between two metal dummy sheets with PTFE layers to reduce adhesion, by Al-Obaidi et al. [5,12], Emami et al. [13], and Hou et al. [8]. Conte et al. [14] proposed to use the pressure of a heated fluid to apply an FRTP onto a metal sheet during SPIF. Defects such as cracks and wrinkles were often observed, especially at higher wall angles [5,8,13]. One reason could be the tool path strategies used, which have been basic z-level or spiral contouring tool paths of metal SPIF. Hou et al. [8] were the only researchers to additionally try an inside-out tool path, starting forming in the center of the sheet in each layer, and reported on improved part quality.

While the standard tool path strategies have been proven for metal sheets, they do not account for the deformation mechanisms of woven fiber fabrics, which must be sheared to represent a 3D shape with double curvature smoothly. This shear is usually introduced by tensioning the fabric in-plane in fiber or bias direction [4], which is hard to realize with undetermined fiber movement between two metal sheets and in a dieless process. Therefore, Rath et al. [15] developed a draping strategy for the dieless forming of woven FRP and tested it in various alternative processes for forming FRP without a mold [16,17]. The strategy involves a 3D tool movement in fiber direction along the target geometry, starting at a fixed point and moving towards the edge of the fabric. As the sequence of the paths is determined according to an ascending shear angle in the fabric as projected by a simulation, shear is introduced in a determined manner.

This paper aims to adapt the draping strategy [15] to SPIF to improve the incremental forming of woven FRPs together with metal dummy sheets. Therefore, multiple novel tool path strategies taking into account both the requirements of metal forming and fabric draping are developed and tested in a setup very similar to the one realized by Al-Obaidi et al. [5,12].

## Methodology

Materials. For each forming experiment, a woven carbon fiber-reinforced organo sheet was placed between two metal sheets, separated on each side by PTFE sheets to allow the sliding of the CFRP. The fiber orientations corresponded to the edges of the sheets and were therefore also aligned with the clamping frame. The organo sheets consisted of a styrene-acrylonitrile (SAN) matrix and two layers of twill-woven carbon fibers (CF) with a basis weight of 245 g/m², resulting in a thickness of 0.6 mm and a fiber-volume content of 45 %. With a size of 210 mm x 240 mm, the CFRP was smaller than the metal sheets with 350 mm x 350 mm. On the lower side, an aluminum sheet EN AW-1050A with a thickness of 0.5 mm was used as a backing plate. On the upper side, a DC04 steel sheet of 0.8 mm thickness was used to transfer the deformation while protecting the organo sheet from direct tool contact. The PTFE sheets were of 0.05 mm thickness and 270 mm x 270 mm in size. The layup is shown in Fig. 1.

<figure><img src="image_625.png" /><figcaption>Fig. 1 Layup of organo sheet, PTFE sheets, and metal sheets, including clamping and support</figcaption></figure>

Setup. The sandwich layup of metal, PTFE, and organo sheets was clamped on top of a heating chamber made of vermiculite and calcium silicate panels. It was ensured that only the metal sheets
were clamped while the organo and PTFE sheets could slide between them. Due to the comparatively high distance between the formed area and the clamping, for some experiments, a rigid support plate with 5 mm thickness and a hole of 105 mm diameter with 5 mm edge radius was additionally placed below the aluminum sheet to act as a template, as shown in Fig. 1. Two Leister Mistral 2 Premium hot air guns, each with a power of 3.4 kW, were used to heat the layup to 180-200 °C. Thereby, the hot air was deflected downwards, as this resulted in a homogeneous temperature distribution on the sheet stack in initial tests. Temperature was measured on the top steel sheet using two thermocouples. The clamping was tightened only after the desired temperature was reached to allow thermal expansion of the metal. SPIF was performed using a steel 1.2210 tool with a hemispherical tip of 20 mm diameter guided by an ABB IRB 6660-205 industrial robot with 205 kg load capacity, as shown in Fig. 2.

After forming, the layup was cooled to room temperature, unclamped, and separated. The resulting geometries were visually inspected for defects such as cracks or wrinkles and 3D-scanned with a GOM ATOS CompactScan 5M.

<figure><img src="image_615.png" /><figcaption>Fig. 2 Robotic single point incremental forming setup with heating chamber</figcaption></figure>

Forming Strategies. Frustum cones with an upper diameter of 100 mm, a lower diameter of 25 mm, and a height of 30 mm were formed, corresponding to a wall angle of approx. 38.7°. Unless stated otherwise, a tool speed of 20 mm/s and a layer height of 0.5 mm was used. The tool orientation was kept constant throughout the process, and no lubrication was used.

First, a traditional z-level contouring strategy was used as a reference, denoted as strategy 1. Thereby, the tool travels along the outer contour of the cone in each layer, stepping down after each circle. The same is done in strategy 2, except that at the end, the lower face of the truncated cone is filled spirally from the outside to the inside, radially moving inwards 0.6 mm with every revolution. In strategy 3, each layer is formed by spiraling from the center point to the outer contour, filling each slice, again with a winding pitch of 0.6 mm.

Strategy 4 aims to resemble the path forming strategy for fabric draping developed by Rath et al. [15]. However, because forming in SPIF must originate from the initial plane, it is not possible to follow 3D paths directly. Instead, the paths running from a central starting point outwards along the fiber orientations of the fabric are projected in the cutting planes of the geometry. Therefore, just like strategy 3, strategy 4 also includes slicing the desired geometry and filling each layer with tool paths; however, the orientation of these paths depends on the fiber orientations of the woven
were clamped while the organo and PTFE sheets could slide between them. Due to the comparatively high distance between the formed area and the clamping, for some experiments, a rigid support plate with 5 mm thickness and a hole of 105 mm diameter with 5 mm edge radius was additionally placed below the aluminum sheet to act as a template, as shown in Fig. 1. Two Leister Mistral 2 Premium hot air guns, each with a power of 3.4 kW, were used to heat the layup to 180-200 °C. Thereby, the hot air was deflected downwards, as this resulted in a homogeneous temperature distribution on the sheet stack in initial tests. Temperature was measured on the top steel sheet using two thermocouples. The clamping was tightened only after the desired temperature was reached to allow thermal expansion of the metal. SPIF was performed using a steel 1.2210 tool with a hemispherical tip of 20 mm diameter guided by an ABB IRB 6660-205 industrial robot with 205 kg load capacity, as shown in Fig. 2.

After forming, the layup was cooled to room temperature, unclamped, and separated. The resulting geometries were visually inspected for defects such as cracks or wrinkles and 3D-scanned with a GOM ATOS CompactScan 5M.

<figure><img src="image_626.png" /><figcaption>Fig. 2 Robotic single point incremental forming setup with heating chamber</figcaption></figure>

Forming Strategies. Frustum cones with an upper diameter of 100 mm, a lower diameter of 25 mm, and a height of 30 mm were formed, corresponding to a wall angle of approx. 38.7°. Unless stated otherwise, a tool speed of 20 mm/s and a layer height of 0.5 mm was used. The tool orientation was kept constant throughout the process, and no lubrication was used.

First, a traditional z-level contouring strategy was used as a reference, denoted as strategy 1. Thereby, the tool travels along the outer contour of the cone in each layer, stepping down after each circle. The same is done in strategy 2, except that at the end, the lower face of the truncated cone is filled spirally from the outside to the inside, radially moving inwards 0.6 mm with every revolution. In strategy 3, each layer is formed by spiraling from the center point to the outer contour, filling each slice, again with a winding pitch of 0.6 mm.

Strategy 4 aims to resemble the path forming strategy for fabric draping developed by Rath et al. [15]. However, because forming in SPIF must originate from the initial plane, it is not possible to follow 3D paths directly. Instead, the paths running from a central starting point outwards along the fiber orientations of the fabric are projected in the cutting planes of the geometry. Therefore, just like strategy 3, strategy 4 also includes slicing the desired geometry and filling each layer with tool paths; however, the orientation of these paths depends on the fiber orientations of the woven
fabric and their sequence on the projected shear angle. First, straight lines are formed from the center point outwards in four perpendicular directions, corresponding to the directions of the perpendicular warp and weft yarns of the fabric. Subsequently, parallel paths are formed on both sides of the four lines, again traveling from the inside to the outside, maintaining a distance of 2 mm between the already created lines. This process is repeated until the whole contour of the layer is filled with parallel lines. After completing a layer, the next z-increment of 0.75 mm is formed likewise, again starting from the center point. Tool speed during the forming operations was 20 mm/s and higher during transfer movements of the robot to reach the starting point of the following path when the tool was out of contact with the surface.

In strategies 1 to 4, forming starts with the upper, larger diameter, which requires the fabric to always slide inwards through the cone's already created edge and slope. However, it would be better and more closely match the draping strategy if the cone was formed starting with the lower, smaller diameter. This approach, however, would require moving the whole already formed geometry downwards and adding on to the slope until reaching the upper diameter. As a controlled rigid body movement is hard to achieve in SPIF, the already formed geometry must be pushed downwards at every point to realize the desired deformation. This ought to be achieved with strategy 5, which incorporates actual 3D paths. First, parallel paths in perpendicular fiber directions are formed at the first z-level, just as in strategy 4; however, only up to the diameter of the lower face of the cone. Subsequently, these paths are repeated at the next z-level, whereby they are extended along the lateral surface slope until the radius of the next slice is reached at z = 0. This process is repeated until the desired height is achieved, again with a z-increment of 0.75 mm.

Fig. 3 gives an overview of the different forming strategies. For better visibility of the paths, a higher winding pitch, path distance, and z-increment were used in the plots, and layers were additionally separated.

<figure><img src="image_616.png" /><figcaption>Fig. 3 Overview of the tool path strategies used for SPIF of woven FRPT sheets</figcaption></figure>

## Results & Discussion

Visual Observations. Fig. 4 shows all organo sheets that were formed into conical shapes. None of these developed any macroscopic cracks. As desired, fibers were drawn in towards the forming
fabric and their sequence on the projected shear angle. First, straight lines are formed from the center point outwards in four perpendicular directions, corresponding to the directions of the perpendicular warp and weft yarns of the fabric. Subsequently, parallel paths are formed on both sides of the four lines, again traveling from the inside to the outside, maintaining a distance of 2 mm between the already created lines. This process is repeated until the whole contour of the layer is filled with parallel lines. After completing a layer, the next z-increment of 0.75 mm is formed likewise, again starting from the center point. Tool speed during the forming operations was 20 mm/s and higher during transfer movements of the robot to reach the starting point of the following path when the tool was out of contact with the surface.

In strategies 1 to 4, forming starts with the upper, larger diameter, which requires the fabric to always slide inwards through the cone's already created edge and slope. However, it would be better and more closely match the draping strategy if the cone was formed starting with the lower, smaller diameter. This approach, however, would require moving the whole already formed geometry downwards and adding on to the slope until reaching the upper diameter. As a controlled rigid body movement is hard to achieve in SPIF, the already formed geometry must be pushed downwards at every point to realize the desired deformation. This ought to be achieved with strategy 5, which incorporates actual 3D paths. First, parallel paths in perpendicular fiber directions are formed at the first z-level, just as in strategy 4; however, only up to the diameter of the lower face of the cone. Subsequently, these paths are repeated at the next z-level, whereby they are extended along the lateral surface slope until the radius of the next slice is reached at z = 0. This process is repeated until the desired height is achieved, again with a z-increment of 0.75 mm.

Fig. 3 gives an overview of the different forming strategies. For better visibility of the paths, a higher winding pitch, path distance, and z-increment were used in the plots, and layers were additionally separated.

<figure><img src="image_627.png" /><figcaption>Fig. 3 Overview of the tool path strategies used for SPIF of woven FRPT sheets</figcaption></figure>

## Results & Discussion

Visual Observations. Fig. 4 shows all organo sheets that were formed into conical shapes. None of these developed any macroscopic cracks. As desired, fibers were drawn in towards the forming
area and bent successfully. However, the amount of shear achieved in the fabric differs significantly among the strategies. The results of strategies 1 and 2 without support show wrinkles of high magnitude, especially in the corners, where the woven fabric should have been sheared. Towards the more advanced strategies, more shear becomes visible, and wrinkles decrease in height while increasing in width. Among the experiments without support, strategy 5 introduced the most shear and produced the smoothest and most distinguishable cone with a clearly visible base edge. This result, however, is the case for all experiments with support. While all eventual wrinkles are minor in height compared to forming without support, strategies 1 and 2 show more small wrinkles in the material surrounding the cone, which would be trimmed in practice.

On all organo sheets, deconsolidated areas are visible, where the surface shows the individual yarns instead of a smooth matrix layer. In strategies 3 and 4, this effect leaves most of the cone's surface, except for the lower part, uncovered. This might not only be due to resin accumulation because of gravity but also to the formation of a gap between the metal sheets. Comparing strategies 1 and 2, the effect of forming and thus applying pressure on the composite becomes visible. As the bottom surface of the cone was not covered by paths in strategy 1, no pressure was applied, and the organo sheet deconsolidated. The same happened on the lateral surface areas of cones formed in strategies 1 and 2 with support, where the lower aluminum sheet cracked and tore during the process, as visible in Fig. 5, thus not providing support for consolidation. A reason for the tearing could be that the commonly used contour tool paths in strategies 1 and 2 provoke rigid body motion of the area enclosed by the contour, leading to a more widely distributed pressure on the aluminum sheet instead of a relatively localized force in the other strategies. With the support preventing a global deformation of the aluminum sheet, cracks developed, and the sheets tore.

As shown in Fig. 5, cracks also appeared in both upper and lower metal sheets used in strategy 5. As the primary deformation occurs in the bottom surface of the cone, which is pushed downwards layer by layer, material thinning led to the development of cracks in the steel sheet starting at nominal depths of 28.5 mm and 23.5 mm without and with support, respectively. These cracks in the path direction grew with increasing depth and, in the case of the experiment with a support plate, led to the formation of tears perpendicular to the path direction, forcing the termination of the forming process at a depth of 26.25 mm. Without support, the sheets bent and deformed more globally in the free area within the clamping frame. Therefore, the actual depth of the cone with respect to the base plate was lower than the nominal depth of the tool and forming was possible until the end.

Unlike the other strategies, both results from strategy 5 show a spreading of the yarns in the formed area. This effect intensifies towards the bottom surface of the cone and must be due to a high and repeated compaction of the yarns under pressure from the tool. Another observation for strategy 5 that is more noticeable when support is used is that the base of the cone is no longer circular. Particularly towards the end of the forming process, the tool is deflected from the diagonals formed by the starting points of the paths on the lateral surface of the cone, towards the initial, central paths. As a result, the paths are no longer perpendicular to each other, as highlighted in Fig. 5, and the radius of the base circle decreases towards the diagonals. The reason for this result is that the paths in strategy 5, except the initial paths starting from the center, do not run radially but combined radially and tangentially along the lateral surface of the cone. However, the resistance to the deformation acts, in simple terms, normal to the surface so that a force component perpendicular to the direction of movement results. This force is higher in case the support is used, and the comparatively low stiffness of the robot thus causes a deviation from the target paths.
area and bent successfully. However, the amount of shear achieved in the fabric differs significantly among the strategies. The results of strategies 1 and 2 without support show wrinkles of high magnitude, especially in the corners, where the woven fabric should have been sheared. Towards the more advanced strategies, more shear becomes visible, and wrinkles decrease in height while increasing in width. Among the experiments without support, strategy 5 introduced the most shear and produced the smoothest and most distinguishable cone with a clearly visible base edge. This result, however, is the case for all experiments with support. While all eventual wrinkles are minor in height compared to forming without support, strategies 1 and 2 show more small wrinkles in the material surrounding the cone, which would be trimmed in practice.

On all organo sheets, deconsolidated areas are visible, where the surface shows the individual yarns instead of a smooth matrix layer. In strategies 3 and 4, this effect leaves most of the cone's surface, except for the lower part, uncovered. This might not only be due to resin accumulation because of gravity but also to the formation of a gap between the metal sheets. Comparing strategies 1 and 2, the effect of forming and thus applying pressure on the composite becomes visible. As the bottom surface of the cone was not covered by paths in strategy 1, no pressure was applied, and the organo sheet deconsolidated. The same happened on the lateral surface areas of cones formed in strategies 1 and 2 with support, where the lower aluminum sheet cracked and tore during the process, as visible in Fig. 5, thus not providing support for consolidation. A reason for the tearing could be that the commonly used contour tool paths in strategies 1 and 2 provoke rigid body motion of the area enclosed by the contour, leading to a more widely distributed pressure on the aluminum sheet instead of a relatively localized force in the other strategies. With the support preventing a global deformation of the aluminum sheet, cracks developed, and the sheets tore.

As shown in Fig. 5, cracks also appeared in both upper and lower metal sheets used in strategy 5. As the primary deformation occurs in the bottom surface of the cone, which is pushed downwards layer by layer, material thinning led to the development of cracks in the steel sheet starting at nominal depths of 28.5 mm and 23.5 mm without and with support, respectively. These cracks in the path direction grew with increasing depth and, in the case of the experiment with a support plate, led to the formation of tears perpendicular to the path direction, forcing the termination of the forming process at a depth of 26.25 mm. Without support, the sheets bent and deformed more globally in the free area within the clamping frame. Therefore, the actual depth of the cone with respect to the base plate was lower than the nominal depth of the tool and forming was possible until the end.

Unlike the other strategies, both results from strategy 5 show a spreading of the yarns in the formed area. This effect intensifies towards the bottom surface of the cone and must be due to a high and repeated compaction of the yarns under pressure from the tool. Another observation for strategy 5 that is more noticeable when support is used is that the base of the cone is no longer circular. Particularly towards the end of the forming process, the tool is deflected from the diagonals formed by the starting points of the paths on the lateral surface of the cone, towards the initial, central paths. As a result, the paths are no longer perpendicular to each other, as highlighted in Fig. 5, and the radius of the base circle decreases towards the diagonals. The reason for this result is that the paths in strategy 5, except the initial paths starting from the center, do not run radially but combined radially and tangentially along the lateral surface of the cone. However, the resistance to the deformation acts, in simple terms, normal to the surface so that a force component perpendicular to the direction of movement results. This force is higher in case the support is used, and the comparatively low stiffness of the robot thus causes a deviation from the target paths.
<figure><img src="image_617.png" /><figcaption>Fig. 4 Upper steel sheets formed without support (view on cone facing down), as well as formed organo sheets for all path strategies with and without support (view on cone facing up)</figcaption></figure>

<figure><img src="image_618.png" /><figcaption>Fig. 5 Cracks in the lower aluminum and upper steel sheets after forming with support (S)</figcaption></figure>
<figure><img src="image_628.png" /><figcaption>Fig. 4 Upper steel sheets formed without support (view on cone facing down), as well as formed organo sheets for all path strategies with and without support (view on cone facing up)</figcaption></figure>

<figure><img src="image_629.png" /><figcaption>Fig. 5 Cracks in the lower aluminum and upper steel sheets after forming with support (S)</figcaption></figure>
3D Scans. Optical 3D measurements of the formed sheets allow for a quantitative evaluation and comparison of the results. For this purpose, each scanned point cloud was meshed before an ideal target cone surface was manually aligned and then fine-registered with it using the Iterative Closest Point Algorithm (ICP). In accordance with the target geometry, the scanned surface was segmented at a radial distance of 20 mm around the upper base circle edge of the cone. Finally, absolute and signed distances between the scanned and the target meshes were calculated. The results are shown in Fig. 6, in top view on the upwards-facing cone.

All cones formed without support are below the target geometry in the tip area, denoted by blue color, and generally above the target geometry around the base circle, indicated by red color. This means that the desired height of the cones could not be reached. Furthermore, the wrinkles are clearly visible in the images. However, among the experiments without support, an improvement in shape accuracy can be observed for each more sophisticated strategy. This is also reflected in the mean absolute distances, plotted in Fig. 7, and matches the visual observations from above. Strategy 5, which builds the cone in a tip-to-base approach best corresponding to the draping requirements of woven fibers, especially sets itself apart. In every formed layer, the tool movements along the fiber directions compact the fabric, push eventual wrinkles outwards through the lateral surface of the cone, and foster the introduction of shear through pulling in fiber direction.

However, an even better match between the samples deformed with support and the target geometry becomes evident from the results, even though the color scales are different to better visualize the localization of the distances on the surfaces. Among the samples deformed with support, the mean absolute distances are very similar except for strategy 5, where the cone did not reach the target height for the abovementioned reasons. Furthermore, the frequent loading and unloading of the sheet by detaching and repositioning the tool in strategies 4 and 5 could have led to unfavorable sliding of the sheets over the edge of the support, which also would have had to grow with the current base circle diameter for strategy 5 to fully unfold its advantages.

<figure><img src="image_619.png" /><figcaption>Fig. 6 3D Scans of organo sheets formed without and with support, colors represent signed distances to target geometry (please note different scales)</figcaption></figure>
3D Scans. Optical 3D measurements of the formed sheets allow for a quantitative evaluation and comparison of the results. For this purpose, each scanned point cloud was meshed before an ideal target cone surface was manually aligned and then fine-registered with it using the Iterative Closest Point Algorithm (ICP). In accordance with the target geometry, the scanned surface was segmented at a radial distance of 20 mm around the upper base circle edge of the cone. Finally, absolute and signed distances between the scanned and the target meshes were calculated. The results are shown in Fig. 6, in top view on the upwards-facing cone.

All cones formed without support are below the target geometry in the tip area, denoted by blue color, and generally above the target geometry around the base circle, indicated by red color. This means that the desired height of the cones could not be reached. Furthermore, the wrinkles are clearly visible in the images. However, among the experiments without support, an improvement in shape accuracy can be observed for each more sophisticated strategy. This is also reflected in the mean absolute distances, plotted in Fig. 7, and matches the visual observations from above. Strategy 5, which builds the cone in a tip-to-base approach best corresponding to the draping requirements of woven fibers, especially sets itself apart. In every formed layer, the tool movements along the fiber directions compact the fabric, push eventual wrinkles outwards through the lateral surface of the cone, and foster the introduction of shear through pulling in fiber direction.

However, an even better match between the samples deformed with support and the target geometry becomes evident from the results, even though the color scales are different to better visualize the localization of the distances on the surfaces. Among the samples deformed with support, the mean absolute distances are very similar except for strategy 5, where the cone did not reach the target height for the abovementioned reasons. Furthermore, the frequent loading and unloading of the sheet by detaching and repositioning the tool in strategies 4 and 5 could have led to unfavorable sliding of the sheets over the edge of the support, which also would have had to grow with the current base circle diameter for strategy 5 to fully unfold its advantages.

<figure><img src="image_630.png" /><figcaption>Fig. 6 3D Scans of organo sheets formed without and with support, colors represent signed distances to target geometry (please note different scales)</figcaption></figure>
<figure><img src="image_620.png" /><figcaption>Fig. 7 Mean absolute distances between formed organo sheets and target geometry</figcaption></figure>

Cross Sections. The reason for the better shape accuracy of cones formed with support becomes apparent when looking at Fig. 8, showing a comparison of cross sections of the scanned surfaces of the upper steel sheets, organo sheets, and lower aluminum sheets on the example of strategy 4 experiments with and without support. As mentioned above, the unsupported deformation of the sheets includes a global elastic and plastic bending, which is significantly more pronounced for the lower aluminum sheet than for the upper steel sheet. While the tool introduces a relatively local deformation in the upper sheet, forming forces are transferred and spread within the steel, PTFE, and organo sheets, so that the pressure on the aluminum sheet is applied in a comparably more extensive area, leading to a rather global deformation. As the two metal sheets detach and voids build between them, the enclosed organo sheet obtains space for the development of wrinkles and, due to gravity, conforms to the lower, globally bent sheet. In contrast, if the lower sheet is supported, the geometries of all three sheets match relatively well, apart from deviations caused by elastic spring-back and other effects.

<figure><img src="image_621.png" /><figcaption>Fig. 8 Comparison of cross sections of formed sheets for strategy 4 without and with support, each taken through the center of the cone and manually aligned in x-z-plane for comparison</figcaption></figure>
<figure><img src="image_631.png" /><figcaption>Fig. 7 Mean absolute distances between formed organo sheets and target geometry</figcaption></figure>

Cross Sections. The reason for the better shape accuracy of cones formed with support becomes apparent when looking at Fig. 8, showing a comparison of cross sections of the scanned surfaces of the upper steel sheets, organo sheets, and lower aluminum sheets on the example of strategy 4 experiments with and without support. As mentioned above, the unsupported deformation of the sheets includes a global elastic and plastic bending, which is significantly more pronounced for the lower aluminum sheet than for the upper steel sheet. While the tool introduces a relatively local deformation in the upper sheet, forming forces are transferred and spread within the steel, PTFE, and organo sheets, so that the pressure on the aluminum sheet is applied in a comparably more extensive area, leading to a rather global deformation. As the two metal sheets detach and voids build between them, the enclosed organo sheet obtains space for the development of wrinkles and, due to gravity, conforms to the lower, globally bent sheet. In contrast, if the lower sheet is supported, the geometries of all three sheets match relatively well, apart from deviations caused by elastic spring-back and other effects.

<figure><img src="image_632.png" /><figcaption>Fig. 8 Comparison of cross sections of formed sheets for strategy 4 without and with support, each taken through the center of the cone and manually aligned in x-z-plane for comparison</figcaption></figure>
## Summary, Conclusions & Outlook

Combining the flexibility of a robot with a digital process chain in a dieless forming process, robotic ISF can open new opportunities for the economic production of FRP prototypes and individual parts. In order to optimize part accuracy and quality in SPIF of woven FRP placed between two metal sheets, different novel tool path strategies were developed, tested, and compared to a conventional contouring strategy. These include surface-filling z-level approaches, such as spiraling inside out or moving in parallel lines in the fiber directions in each slice, and a three-dimensional approach inspired by fabric draping. From the results, it can be concluded that tool paths considering the draping requirements of woven FRP can lead to better results in SPIF without support. Paths following the fiber directions apply tension in the underlying fabric in fiber direction, introducing the necessary shear. Forming the cones in a tip-to-base approach with 3D tool paths yielded the best results in unsupported SPIF. However, as the already formed area of the cone needs to be pushed downwards in every subsequent step and paths are repeatedly formed on each z-level, a higher metal sheet thickness is required to prevent the development of cracks.

An even more significant improvement in part accuracy was achieved by introducing a fixed support with the size of the cone's base diameter below the sheets, preventing global bending of the lower metal sheet due to a spreading of the forming force through the layers. However, using conventional tool paths, the lower sheet tore in the supported setup, which led to the local deconsolidation of the FRTP due to the lack of pressure being applied on the organo sheet. The improved tool path strategies prevented the tearing while not entirely preventing deconsolidation—instead, the thermoplastic matrix accumulated in the lower part of the cone.

Further experiments with different metal and organo sheet configurations and forming temperatures should be conducted to investigate their influence on part quality, taking into account microscopic analyses to determine resulting consolidation states, fiber-volume-contents, and internal defects. Forming limits for the individual strategies need to be determined by forming different geometries, varying wall angles. In addition, investigations on other fiber architectures and textiles, such as triaxial woven fabrics or non-crimp fabrics (NCF) would be of relevance. Specifically, it should be studied if fiber orientation-based tool path approaches are also applicable and beneficial in these cases, which is expected by the authors. Simulations of the thermo-mechanical deformation and draping behavior of the fabrics should be used to determine the shear angles and the fiber deflections from the initial state as inputs to the path planning. Lastly, an optimized support, such as a globally applied hydrostatic pressure or even a double-sided incremental forming process, should be tested in conjunction with the optimal tool path strategy.

## References

[1] M. Neitzel, P. Mitschang, U. Breuer, Handbuch Verbundwerkstoffe: Werkstoffe, Verarbeitung, Anwendung, 2nd ed., Hanser, München, 2014.

[2] K. Bilisik, N.S. Karaduman, N.E. Bilisik, Fiber Architectures for Composite Applications, in: R. Fangueiro, S. Rana (Eds.), Fibrous and Textile Materials for Composite Applications, 1st ed., Springer Singapore, Singapore, 2016, pp. 75–134.

[3] P. Boisse, R. Akkerman, P. Carlone, L. Kärger, S.V. Lomov, J.A. Sherwood, Advances in composite forming through 25 years of ESAFORM, Int J Mater Form 15 (2022). https://doi.org/10.1007/s12289-022-01682-8

[4] S.G. Hancock, K.D. Potter, The use of kinematic drape modelling to inform the hand lay-up of complex composite components using woven reinforcements, Compos. - A: Appl. Sci. Manuf. 37 (2006) 413–422. https://doi.org/10.1016/j.compositesa.2005.05.044
## Summary, Conclusions & Outlook

Combining the flexibility of a robot with a digital process chain in a dieless forming process, robotic ISF can open new opportunities for the economic production of FRP prototypes and individual parts. In order to optimize part accuracy and quality in SPIF of woven FRP placed between two metal sheets, different novel tool path strategies were developed, tested, and compared to a conventional contouring strategy. These include surface-filling z-level approaches, such as spiraling inside out or moving in parallel lines in the fiber directions in each slice, and a three-dimensional approach inspired by fabric draping. From the results, it can be concluded that tool paths considering the draping requirements of woven FRP can lead to better results in SPIF without support. Paths following the fiber directions apply tension in the underlying fabric in fiber direction, introducing the necessary shear. Forming the cones in a tip-to-base approach with 3D tool paths yielded the best results in unsupported SPIF. However, as the already formed area of the cone needs to be pushed downwards in every subsequent step and paths are repeatedly formed on each z-level, a higher metal sheet thickness is required to prevent the development of cracks.

An even more significant improvement in part accuracy was achieved by introducing a fixed support with the size of the cone's base diameter below the sheets, preventing global bending of the lower metal sheet due to a spreading of the forming force through the layers. However, using conventional tool paths, the lower sheet tore in the supported setup, which led to the local deconsolidation of the FRTP due to the lack of pressure being applied on the organo sheet. The improved tool path strategies prevented the tearing while not entirely preventing deconsolidation—instead, the thermoplastic matrix accumulated in the lower part of the cone.

Further experiments with different metal and organo sheet configurations and forming temperatures should be conducted to investigate their influence on part quality, taking into account microscopic analyses to determine resulting consolidation states, fiber-volume-contents, and internal defects. Forming limits for the individual strategies need to be determined by forming different geometries, varying wall angles. In addition, investigations on other fiber architectures and textiles, such as triaxial woven fabrics or non-crimp fabrics (NCF) would be of relevance. Specifically, it should be studied if fiber orientation-based tool path approaches are also applicable and beneficial in these cases, which is expected by the authors. Simulations of the thermo-mechanical deformation and draping behavior of the fabrics should be used to determine the shear angles and the fiber deflections from the initial state as inputs to the path planning. Lastly, an optimized support, such as a globally applied hydrostatic pressure or even a double-sided incremental forming process, should be tested in conjunction with the optimal tool path strategy.

## References

[1] M. Neitzel, P. Mitschang, U. Breuer, Handbuch Verbundwerkstoffe: Werkstoffe, Verarbeitung, Anwendung, 2nd ed., Hanser, München, 2014.

[2] K. Bilisik, N.S. Karaduman, N.E. Bilisik, Fiber Architectures for Composite Applications, in: R. Fangueiro, S. Rana (Eds.), Fibrous and Textile Materials for Composite Applications, 1st ed., Springer Singapore, Singapore, 2016, pp. 75–134.

[3] P. Boisse, R. Akkerman, P. Carlone, L. Kärger, S.V. Lomov, J.A. Sherwood, Advances in composite forming through 25 years of ESAFORM, Int J Mater Form 15 (2022). https://doi.org/10.1007/s12289-022-01682-8

[4] S.G. Hancock, K.D. Potter, The use of kinematic drape modelling to inform the hand lay-up of complex composite components using woven reinforcements, Compos. - A: Appl. Sci. Manuf. 37 (2006) 413–422. https://doi.org/10.1016/j.compositesa.2005.05.044
[5] A. Al-Obaidi, A. Kunke, V. Kräusel, Hot single-point incremental forming of glass-fiber-reinforced polymer (PA6GF47) supported by hot air, J. Manuf. Process. 43 (2019) 17-25.
https://doi.org/10.1016/j.jmapro.2019.04.036

[6] L. Ben Said, The incremental sheet forming; technology, modeling and formability: A brief review, Proc. Inst. Mech. Eng. E: J. Process Mech. Eng. 236 (2022) 2729-2755.
https://doi.org/10.1177/09544089221093306

[7] P. Martins, L. Kwiatkowski, V. Franzen, A.E. Tekkaya, M. Kleiner, Single point incremental forming of polymers, CIRP Annals 58 (2009) 229-232.
https://doi.org/10.1016/j.cirp.2009.03.095

[8] C. Hou, X. Su, X. Peng, X. Wu, D. Yang, Thermal-Assisted Single Point Incremental Forming of Jute Fabric Reinforced Poly(lactic acid) Biocomposites, Fibers Polym. 21 (2020) 2373-2379.
https://doi.org/10.1007/s12221-020-1016-0

[9] M. Fiorotto, M. Sorgente, G. Lucchetta, Preliminary studies on single point incremental forming for composite materials, Int. J. Mater. Form. 3 (2010) 951-954.
https://doi.org/10.1007/s12289-010-0926-6

[10] X. Xiao, J.-J. Kim, S.-H. Oh, Y.-S. Kim, Study on the incremental sheet forming of CFRP sheet, Compos. - A: Appl. Sci. Manuf. 141 (2021) 106209.
https://doi.org/10.1016/j.compositesa.2020.106209

[11] P. Xu, X. Li, F. Feng, X. Li, Y. Yang, Experimental and numerical studies on two-point incremental forming of woven fabric composite sheet, J. Manuf. Process. 85 (2023) 205-215.
https://doi.org/10.1016/j.jmapro.2022.11.049

[12] A. Al-Obaidi, A. Graf, V. Kräusel, M. Trautmann, Heat supported single point incremental forming of hybrid laminates for orthopedic applications, Procedia Manuf. 29 (2019) 21-27.
https://doi.org/10.1016/j.promfg.2019.02.101

[13] R. Emami, M.J. Mirnia, M. Elyasi, A. Zolfaghari, An experimental investigation into single point incremental forming of glass fiber-reinforced polyamide sheet with different fiber orientations and volume fractions at elevated temperatures, J. Thermoplast. Compos. Mater. 36 (2022) 1893-1917.
https://doi.org/10.1177/08927057221074266

[14] R. Conte, G. Serratore, G. Ambrogio, F. Gagliardi, Numerical analyses of long fiber-reinforced polymeric sheets processed by Single Point Incremental Forming, Int. J. Adv. Manuf. Technol. 123 (2022) 1203-1214.
https://doi.org/10.1007/s00170-022-10212-4

[15] J.-E. Rath, L.-S. Schwieger, T. Schüppstuhl, Robotic Die-Less Forming Strategy for Fiber-Reinforced Plastic Composites Production, Procedia CIRP 107 (2022) 1281-1286.
https://doi.org/10.1016/j.procir.2022.05.145

[16] J.-E. Rath, R. Graupner, T. Schüppstuhl, Processing Strategies for Dieless Forming of Fiber-Reinforced Plastic Composites, Machines 11 (2023) 365.
https://doi.org/10.3390/machines11030365

[17] J.-E. Rath, T. Schüppstuhl, Investigation of Metal Wire Mesh as Support Material for Dieless Forming of Woven Reinforcement Textiles, JMMP 7 (2023) 182.
https://doi.org/10.3390/jmmp7050182
[5] A. Al-Obaidi, A. Kunke, V. Kräusel, Hot single-point incremental forming of glass-fiber-reinforced polymer (PA6GF47) supported by hot air, J. Manuf. Process. 43 (2019) 17-25.
https://doi.org/10.1016/j.jmapro.2019.04.036

[6] L. Ben Said, The incremental sheet forming; technology, modeling and formability: A brief review, Proc. Inst. Mech. Eng. E: J. Process Mech. Eng. 236 (2022) 2729-2755.
https://doi.org/10.1177/09544089221093306

[7] P. Martins, L. Kwiatkowski, V. Franzen, A.E. Tekkaya, M. Kleiner, Single point incremental forming of polymers, CIRP Annals 58 (2009) 229-232.
https://doi.org/10.1016/j.cirp.2009.03.095

[8] C. Hou, X. Su, X. Peng, X. Wu, D. Yang, Thermal-Assisted Single Point Incremental Forming of Jute Fabric Reinforced Poly(lactic acid) Biocomposites, Fibers Polym. 21 (2020) 2373-2379.
https://doi.org/10.1007/s12221-020-1016-0

[9] M. Fiorotto, M. Sorgente, G. Lucchetta, Preliminary studies on single point incremental forming for composite materials, Int. J. Mater. Form. 3 (2010) 951-954.
https://doi.org/10.1007/s12289-010-0926-6

[10] X. Xiao, J.-J. Kim, S.-H. Oh, Y.-S. Kim, Study on the incremental sheet forming of CFRP sheet, Compos. - A: Appl. Sci. Manuf. 141 (2021) 106209.
https://doi.org/10.1016/j.compositesa.2020.106209

[11] P. Xu, X. Li, F. Feng, X. Li, Y. Yang, Experimental and numerical studies on two-point incremental forming of woven fabric composite sheet, J. Manuf. Process. 85 (2023) 205-215.
https://doi.org/10.1016/j.jmapro.2022.11.049

[12] A. Al-Obaidi, A. Graf, V. Kräusel, M. Trautmann, Heat supported single point incremental forming of hybrid laminates for orthopedic applications, Procedia Manuf. 29 (2019) 21-27.
https://doi.org/10.1016/j.promfg.2019.02.101

[13] R. Emami, M.J. Mirnia, M. Elyasi, A. Zolfaghari, An experimental investigation into single point incremental forming of glass fiber-reinforced polyamide sheet with different fiber orientations and volume fractions at elevated temperatures, J. Thermoplast. Compos. Mater. 36 (2022) 1893-1917.
https://doi.org/10.1177/08927057221074266

[14] R. Conte, G. Serratore, G. Ambrogio, F. Gagliardi, Numerical analyses of long fiber-reinforced polymeric sheets processed by Single Point Incremental Forming, Int. J. Adv. Manuf. Technol. 123 (2022) 1203-1214.
https://doi.org/10.1007/s00170-022-10212-4

[15] J.-E. Rath, L.-S. Schwieger, T. Schüppstuhl, Robotic Die-Less Forming Strategy for Fiber-Reinforced Plastic Composites Production, Procedia CIRP 107 (2022) 1281-1286.
https://doi.org/10.1016/j.procir.2022.05.145

[16] J.-E. Rath, R. Graupner, T. Schüppstuhl, Processing Strategies for Dieless Forming of Fiber-Reinforced Plastic Composites, Machines 11 (2023) 365.
https://doi.org/10.3390/machines11030365

[17] J.-E. Rath, T. Schüppstuhl, Investigation of Metal Wire Mesh as Support Material for Dieless Forming of Woven Reinforcement Textiles, JMMP 7 (2023) 182.
https://doi.org/10.3390/jmmp7050182
# On the methodological measurements of viscosity of unidirectional Flax/PP composites: Towards a benchmark

SIMAFROOKTEH Sepehr<sup>1,2,a</sup>, PERRIN Henri<sup>3,b</sup>, IVENS Jan<sup>1,2,c</sup>, LOMOV Stepan V.<sup>2,d</sup>, and BODAGHI Masoud<sup>3,e*</sup>

<sup>1</sup> Department of Materials Engineering, KU Leuven Campus De Nayer, Sint-Katelijne-Waver, Belgium

<sup>2</sup> Department of Materials Engineering, KU Leuven, Leuven, Belgium

<sup>3</sup> Luxembourg Institute of Science and Technology (LIST), 5, rue Bommel | L-4940, Hautcharage, Luxembourg

<sup>a</sup>sepehr.simaafrookhteh@kuleuven.be, <sup>b</sup>henri.perrin@list.lu, <sup>c</sup>jan.ivens@kuleuven.be, <sup>d</sup>stepan.lomov@kuleuven.be, <sup>e</sup>masoud.bodaghi@list.lu

**Keywords:** Rheology, Viscosity, Squeeze Flow, Thermoplastic Biocomposite

**Abstract** Understanding the viscosity behaviour of the pre-impregnated thermoplastic composites is crucial for optimizing manufacturing processes and ensuring the end product's performance. Reviewing prior research on this topic reveals that viscosity measurements have encountered numerous difficulties, resulting in unclear interpretations and complexities in data processing. The variations in viscosity can be influenced by factors such as cavity thickness, temperature, and compression rate. The objective of this work was to enhance the reliability of results, taking a step toward initiating a benchmark for viscosity measurements of pre-impregnated thermoplastic. The study focuses on transverse squeeze flow experiments of UD flax/polypropylene (PP) for viscosity measurements.

**Introduction**

Different flow and deformation mechanisms are involved during the processing and thermoforming of composite materials, as discussed for various composites and processing techniques in the literature [1]. During the consolidation of pre-impregnated thermoplastic tapes with continuous fibres, a major physical phenomenon is the squeeze flow of the thermoplastic polymer composite [2]. By utilizing a lubrication assumption, an incompressible fluid, and considering a simple geometry, an analytical solution for the governing equations can be derived. This model establishes a correlation between the transient change of the thermoplastic polymer (TP) composite thickness and several factors such as process conditions (pressure, temperature, and time), component geometry, and TP viscosity. A precise understanding of the TP viscosity correlation with temperature and shear rate is crucial for extracting meaningful predictions from the squeeze flow model [3].

Previous studies have investigated the rheological characteristics of TP composites, revealing that the viscosity of TP composites is one to several orders of magnitude higher than the neat TP. For instance, for polypropylene composites, the viscosity falls within the range of 10<sup>4</sup>- 10<sup>6</sup> for temperatures between 175-195 °C. Some research only considered the neat resin viscosity dominating the flow during laminate compaction, assuming that no fibres act at the tape surface [4-6]. Some authors used effective tape viscosity [7-11]. Others rely on literature data [9,11,12] instead of directly characterizing the effective viscosity of the tape under investigation.

It was shown before that tapes from different suppliers may exhibit distinct rheological behaviour, even when composed of similar fibres and polymer types [7,13]. This can be due to the presence of a resin-rich layer, different fibres distribution, and molecular weight of the polymer.

<figure><img src="image_622.png" /></figure>
# On the methodological measurements of viscosity of unidirectional Flax/PP composites: Towards a benchmark

SIMAFROOKTEH Sepehr<sup>1,2,a</sup>, PERRIN Henri<sup>3,b</sup>, IVENS Jan<sup>1,2,c</sup>, LOMOV Stepan V.<sup>2,d</sup>, and BODAGHI Masoud<sup>3,e*</sup>

<sup>1</sup> Department of Materials Engineering, KU Leuven Campus De Nayer, Sint-Katelijne-Waver, Belgium

<sup>2</sup> Department of Materials Engineering, KU Leuven, Leuven, Belgium

<sup>3</sup> Luxembourg Institute of Science and Technology (LIST), 5, rue Bommel | L-4940, Hautcharage, Luxembourg

<sup>a</sup>sepehr.simaafrookhteh@kuleuven.be, <sup>b</sup>henri.perrin@list.lu, <sup>c</sup>jan.ivens@kuleuven.be, <sup>d</sup>stepan.lomov@kuleuven.be, <sup>e</sup>masoud.bodaghi@list.lu

**Keywords:** Rheology, Viscosity, Squeeze Flow, Thermoplastic Biocomposite

**Abstract** Understanding the viscosity behaviour of the pre-impregnated thermoplastic composites is crucial for optimizing manufacturing processes and ensuring the end product's performance. Reviewing prior research on this topic reveals that viscosity measurements have encountered numerous difficulties, resulting in unclear interpretations and complexities in data processing. The variations in viscosity can be influenced by factors such as cavity thickness, temperature, and compression rate. The objective of this work was to enhance the reliability of results, taking a step toward initiating a benchmark for viscosity measurements of pre-impregnated thermoplastic. The study focuses on transverse squeeze flow experiments of UD flax/polypropylene (PP) for viscosity measurements.

**Introduction**

Different flow and deformation mechanisms are involved during the processing and thermoforming of composite materials, as discussed for various composites and processing techniques in the literature [1]. During the consolidation of pre-impregnated thermoplastic tapes with continuous fibres, a major physical phenomenon is the squeeze flow of the thermoplastic polymer composite [2]. By utilizing a lubrication assumption, an incompressible fluid, and considering a simple geometry, an analytical solution for the governing equations can be derived. This model establishes a correlation between the transient change of the thermoplastic polymer (TP) composite thickness and several factors such as process conditions (pressure, temperature, and time), component geometry, and TP viscosity. A precise understanding of the TP viscosity correlation with temperature and shear rate is crucial for extracting meaningful predictions from the squeeze flow model [3].

Previous studies have investigated the rheological characteristics of TP composites, revealing that the viscosity of TP composites is one to several orders of magnitude higher than the neat TP. For instance, for polypropylene composites, the viscosity falls within the range of 10<sup>4</sup>- 10<sup>6</sup> for temperatures between 175-195 °C. Some research only considered the neat resin viscosity dominating the flow during laminate compaction, assuming that no fibres act at the tape surface [4-6]. Some authors used effective tape viscosity [7-11]. Others rely on literature data [9,11,12] instead of directly characterizing the effective viscosity of the tape under investigation.

It was shown before that tapes from different suppliers may exhibit distinct rheological behaviour, even when composed of similar fibres and polymer types [7,13]. This can be due to the presence of a resin-rich layer, different fibres distribution, and molecular weight of the polymer.

<figure><img src="image_633.png" /></figure>
Nevertheless, the viscosity results were not consistently reproducible in consecutive replicate tests. Several studies have experimentally demonstrated a power-law model for viscosity variation with shear rate, describing the non-Newtonian behaviour of the thermoplastic at elevated temperatures. The significant role of TP viscosity in squeeze flow modelling and compression moulding has led researchers to spend several years developing squeeze flow approaches for viscosity measurement [14]. The procedure involves compressing thermoplastic composite specimens at elevated temperatures under isothermal conditions while concurrently recording displacement and applied pressure. Subsequently, the squeeze flow model is applied to compute the viscosity parameters.

Since there is no standard method of measuring the effective viscosity of UD composites, researchers and different laboratories worldwide were required to develop their in-house squeeze flow test instruments. To take a step towards standardisation, the involvement of a large part of the research community is necessary to create a viscosity database and organize a large-scale benchmark. However, even with these efforts, based on our lessons from three consecutive worldwide permeability benchmarks [15–17], certain issues need attention regarding understanding the fundamentals of variability and uncertainty in composite viscosity measurements. This approach aims to propose a unified procedure and implement more controlled experimental conditions to elucidate the factors contributing to variability, enabling a better comparison of experimental results. To achieve this goal, the ongoing collaborative work between LIST and KU Leuven has conducted several experiments to investigate the test conditions for measuring the viscosity of thermoplastic composites. The study outlines how various processing parameters, such as specimen thickness, thickness reduction rate, and temperature, can impact the reproducibility of the viscosity measurements.

While the complexities of viscosity measurement have been extensively studied in the context of continuous synthetic fibre-reinforced thermoplastic composites, the analogous investigation for natural fibres appears to be comparatively limited. The existing studies are limited to short flax fibres with low fibre volume fractions where standard rheometers are applicable [18–20]. To this end, using a squeeze flow test apparatus and a previously derived Newtonian squeeze flow model [21], this paper investigates the viscosity of the UD flax/PP composites.

## Material and methods

The material under examination is a unidirectional prepreg tape composed of flax fibres and semi-crystalline PP, EcoRein/UD50 (BPREG Composites and Textiles A.Ş., Bursa, Turkey). The tape specifications are listed in Table 1.

<table><caption>Table 1 - The technical specifications of the flax/PP unidirectional tape.</caption><thead><tr><td>Supplier</td><td>Fibre (wt %)</td><td>Polymer type</td><td>Thickness (mm)</td><td>Width (cm)</td><td>Areal density (g/m²)</td></tr></thead><tbody><tr><td>BPREG</td><td>50</td><td>Semi-crystalline PP</td><td>0.3</td><td>60</td><td>300</td></tr></tbody></table>

The samples were prepared first by cutting layers and drying them at 80°C for at least 12 hours to diminish the moisture content. The blank was prepared to the desired thickness and consolidated utilizing a hot press at 190 °C under a pressure of 5 bar for 15 minutes of dwell time inside a mould with a spring mechanism to control the final thickness. The laminate was then cooled under pressure before demoulding. The laminate production steps with its different thermal cycle steps are shown in Fig. 1.
Nevertheless, the viscosity results were not consistently reproducible in consecutive replicate tests. Several studies have experimentally demonstrated a power-law model for viscosity variation with shear rate, describing the non-Newtonian behaviour of the thermoplastic at elevated temperatures. The significant role of TP viscosity in squeeze flow modelling and compression moulding has led researchers to spend several years developing squeeze flow approaches for viscosity measurement [14]. The procedure involves compressing thermoplastic composite specimens at elevated temperatures under isothermal conditions while concurrently recording displacement and applied pressure. Subsequently, the squeeze flow model is applied to compute the viscosity parameters.

Since there is no standard method of measuring the effective viscosity of UD composites, researchers and different laboratories worldwide were required to develop their in-house squeeze flow test instruments. To take a step towards standardisation, the involvement of a large part of the research community is necessary to create a viscosity database and organize a large-scale benchmark. However, even with these efforts, based on our lessons from three consecutive worldwide permeability benchmarks [15–17], certain issues need attention regarding understanding the fundamentals of variability and uncertainty in composite viscosity measurements. This approach aims to propose a unified procedure and implement more controlled experimental conditions to elucidate the factors contributing to variability, enabling a better comparison of experimental results. To achieve this goal, the ongoing collaborative work between LIST and KU Leuven has conducted several experiments to investigate the test conditions for measuring the viscosity of thermoplastic composites. The study outlines how various processing parameters, such as specimen thickness, thickness reduction rate, and temperature, can impact the reproducibility of the viscosity measurements.

While the complexities of viscosity measurement have been extensively studied in the context of continuous synthetic fibre-reinforced thermoplastic composites, the analogous investigation for natural fibres appears to be comparatively limited. The existing studies are limited to short flax fibres with low fibre volume fractions where standard rheometers are applicable [18–20]. To this end, using a squeeze flow test apparatus and a previously derived Newtonian squeeze flow model [21], this paper investigates the viscosity of the UD flax/PP composites.

## Material and methods

The material under examination is a unidirectional prepreg tape composed of flax fibres and semi-crystalline PP, EcoRein/UD50 (BPREG Composites and Textiles A.Ş., Bursa, Turkey). The tape specifications are listed in Table 1.

<table><caption>Table 1 - The technical specifications of the flax/PP unidirectional tape.</caption><thead><tr><td>Supplier</td><td>Fibre (wt %)</td><td>Polymer type</td><td>Thickness (mm)</td><td>Width (cm)</td><td>Areal density (g/m²)</td></tr></thead><tbody><tr><td>BPREG</td><td>50</td><td>Semi-crystalline PP</td><td>0.3</td><td>60</td><td>300</td></tr></tbody></table>

The samples were prepared first by cutting layers and drying them at 80°C for at least 12 hours to diminish the moisture content. The blank was prepared to the desired thickness and consolidated utilizing a hot press at 190 °C under a pressure of 5 bar for 15 minutes of dwell time inside a mould with a spring mechanism to control the final thickness. The laminate was then cooled under pressure before demoulding. The laminate production steps with its different thermal cycle steps are shown in Fig. 1.
<figure><img src="image_623.png" /><figcaption>Figure 1 - Specimen preparation: (a) hot press moulding of the UD flax/PP prepregs, (b) recorded thermal cycle of the production process, c) a produced plate prepared for cutting the test specimens.</figcaption></figure>

Samples of L (length) 40 × W (width) 20 mm (with fibres oriented in the length direction) were cut from the plates using a band saw cutting machine and then dried in a vacuum oven for 15 hours at 80 °C before testing. The compression test setup used is very similar to the one used for the compressibility of fabrics [22]. It consists of two circular steel plates with a diameter of 80 mm, aligned parallel to each other and fixated in a universal testing machine (Instron 5985, Instron, USA) equipped with a 30 kN load cell. The setup inside a ventilated hot-air oven is shown in Fig. 2 and b. The oven was first heated to the desired temperature (T) and then opened to put the specimen inside. Samples were allowed to equilibrate with the oven temperature, checking with thermocouples embedded on the bottom plate and the specimen (Fig. 2a and b). Fig. 2c shows the force and temperature cycle after putting the specimen inside when the specimen temperature starts to rise toward the desired temperature. Two temperatures of 180 and 190 °C were used to study the temperature dependency of the measurements. Kapton films were placed between the sample and the steel plates to avoid sticking. When the temperature reached, a force ramp was applied to achieve a constant force of 100 N. Of course, it takes some time for the machine to reach the constant force value, and this transient region can affect the results (see Fig. 2c). The test starts at t₁ when the temperature on the specimen reaches the desired value (T) and continues until t₃, as plotted in Fig. 2c. t₂ minus t₁ is the ramp time that is required for the machine to build up the force to the required constant value. This time can be defined based on the compaction rate in the first stage of the machine profile. The model derived by Lin and Advani [21] is used to measure the composite viscosity (μmf):

$$ \mu_{mf} = \frac{5}{8} \frac{Ft}{L_i W_i^3} \frac{(h_i^2 h_f^5)}{(h_i^5 - h_f^5)} $$

where Lᵢ and Wᵢ are the initial length and initial half-width of the laminate, respectively. hᵢ represents the initial thickness and hᵣ is the final thickness of the laminate due to a constant force of F during the evaluation time of t. The model is derived from the continuity and motion equations for an incompressible fluid (negligible void content). Moreover, a lubrication assumption is made that is valid for a specimen width much higher than its thickness. A no-slip boundary condition is assumed at the top and bottom tool-material interface. This assumption is validated after the test by the intact initial surfaces of the compacted specimen. After the test and removing the Kapton film, there were no signs of material flow (slipping) on the initial surface area of the material [13].
<figure><img src="image_634.png" /><figcaption>Figure 1 - Specimen preparation: (a) hot press moulding of the UD flax/PP prepregs, (b) recorded thermal cycle of the production process, c) a produced plate prepared for cutting the test specimens.</figcaption></figure>

Samples of L (length) 40 × W (width) 20 mm (with fibres oriented in the length direction) were cut from the plates using a band saw cutting machine and then dried in a vacuum oven for 15 hours at 80 °C before testing. The compression test setup used is very similar to the one used for the compressibility of fabrics [22]. It consists of two circular steel plates with a diameter of 80 mm, aligned parallel to each other and fixated in a universal testing machine (Instron 5985, Instron, USA) equipped with a 30 kN load cell. The setup inside a ventilated hot-air oven is shown in Fig. 2 and b. The oven was first heated to the desired temperature (T) and then opened to put the specimen inside. Samples were allowed to equilibrate with the oven temperature, checking with thermocouples embedded on the bottom plate and the specimen (Fig. 2a and b). Fig. 2c shows the force and temperature cycle after putting the specimen inside when the specimen temperature starts to rise toward the desired temperature. Two temperatures of 180 and 190 °C were used to study the temperature dependency of the measurements. Kapton films were placed between the sample and the steel plates to avoid sticking. When the temperature reached, a force ramp was applied to achieve a constant force of 100 N. Of course, it takes some time for the machine to reach the constant force value, and this transient region can affect the results (see Fig. 2c). The test starts at t₁ when the temperature on the specimen reaches the desired value (T) and continues until t₃, as plotted in Fig. 2c. t₂ minus t₁ is the ramp time that is required for the machine to build up the force to the required constant value. This time can be defined based on the compaction rate in the first stage of the machine profile. The model derived by Lin and Advani [21] is used to measure the composite viscosity (μmf):

$$ \mu_{mf} = \frac{5}{8} \frac{Ft}{L_i W_i^3} \frac{(h_i^2 h_f^5)}{(h_i^5 - h_f^5)} $$

where Lᵢ and Wᵢ are the initial length and initial half-width of the laminate, respectively. hᵢ represents the initial thickness and hᵣ is the final thickness of the laminate due to a constant force of F during the evaluation time of t. The model is derived from the continuity and motion equations for an incompressible fluid (negligible void content). Moreover, a lubrication assumption is made that is valid for a specimen width much higher than its thickness. A no-slip boundary condition is assumed at the top and bottom tool-material interface. This assumption is validated after the test by the intact initial surfaces of the compacted specimen. After the test and removing the Kapton film, there were no signs of material flow (slipping) on the initial surface area of the material [13].
Ideally, the compaction displacement is measured using Digital Image Correlation, but this was not possible for the thin samples inside the oven. Thus, the crosshead displacement was used but corrected for system compliances based on Instron instructions. A total of eight batches were prepared, as listed in Table 2. These batches involved two distinct thicknesses of 5.3 and 8.2 mm, temperatures set at 180 and 190 °C, and three different initial compaction rates of 0.5, 5, and 20 mm/min until a constant force of 100 N was achieved. The thicknesses are referred to as t5 ($h_i$ = 5.3 mm) and t8 ($h_i$ = 8.2 mm) in the batch names for simplicity.

<figure><img src="image_624.png" /><figcaption>Figure 2 - Experimental setup for the squeeze flow test. (a) compression setup inside the oven, (b) a magnified view of the specimen mounted for the constant volume test (larger compression plates), (c) temperature and force cycle after mounting the specimen inside ($t = 0$) until the end of the test ($t_3$).</figcaption></figure>

<table><caption>Table 2 - Different testing conditions and coding for each test batch.</caption><thead><tr><th>Batch</th><th>Thickness (mm)</th><th>Test temperature (°C)</th><th>Initial rate (mm/min)</th><th>Constant force (N)</th></tr></thead><tbody><tr><td>t5_T180_r05</td><td rowspan="3">5</td><td rowspan="3">180</td><td>0.5</td><td rowspan="6">100</td></tr><tr><td>t5_T180_r5</td><td>5</td></tr><tr><td>t5_T180_r20</td><td>20</td></tr><tr><td>t5_T190_r05</td><td>190</td><td>0.5</td></tr><tr><td>t8_T180_r05</td><td rowspan="3">8</td><td rowspan="3">180</td><td>0.5</td></tr><tr><td>t8_T180_r5</td><td>5</td></tr><tr><td>t8_T180_r20</td><td>20</td></tr><tr><td>t8_T190_r05</td><td>190</td><td>0.5</td></tr></tbody></table>
Ideally, the compaction displacement is measured using Digital Image Correlation, but this was not possible for the thin samples inside the oven. Thus, the crosshead displacement was used but corrected for system compliances based on Instron instructions. A total of eight batches were prepared, as listed in Table 2. These batches involved two distinct thicknesses of 5.3 and 8.2 mm, temperatures set at 180 and 190 °C, and three different initial compaction rates of 0.5, 5, and 20 mm/min until a constant force of 100 N was achieved. The thicknesses are referred to as t5 ($h_i$ = 5.3 mm) and t8 ($h_i$ = 8.2 mm) in the batch names for simplicity.

<figure><img src="image_635.png" /><figcaption>Figure 2 - Experimental setup for the squeeze flow test. (a) compression setup inside the oven, (b) a magnified view of the specimen mounted for the constant volume test (larger compression plates), (c) temperature and force cycle after mounting the specimen inside (t = 0) until the end of the test (t3).</figcaption></figure>

<table><caption>Table 2 - Different testing conditions and coding for each test batch.</caption><thead><tr><th>Batch</th><th>Thickness (mm)</th><th>Test temperature (°C)</th><th>Initial rate (mm/min)</th><th>Constant force (N)</th></tr></thead><tbody><tr><td>t5_T180_r05</td><td rowspan="3">5</td><td rowspan="3">180</td><td>0.5</td><td rowspan="6">100</td></tr><tr><td>t5_T180_r5</td><td>5</td></tr><tr><td>t5_T180_r20</td><td>20</td></tr><tr><td>t5_T190_r05</td><td>190</td><td>0.5</td></tr><tr><td>t8_T180_r05</td><td rowspan="3">8</td><td rowspan="3">180</td><td>0.5</td></tr><tr><td>t8_T180_r5</td><td>5</td></tr><tr><td>t8_T180_r20</td><td>20</td></tr><tr><td>t8_T190_r05</td><td>190</td><td>0.5</td></tr></tbody></table>
## Results and discussion

Reduction in thickness over time is depicted in Fig. 3, along with the corresponding force changes. In both cases, it is evident that the adjusted thickness following compliance calibration is just slightly higher than the thickness reduction determined by the original displacement of the machine. This confirms the negligible compliance of the setup, allowing the direct utilisation of the original displacement of the machine for data reduction. Nevertheless, it is crucial to validate this step to guarantee accurate thickness reduction measurement, independent of machine compliance, as emphasized in [8].

It is worth mentioning that for all tested samples, significant dimensional change was observed in the transverse direction, confirming squeeze flow as the primary deformation mechanism, as there was no increase in the specimen length due to fibres inextensibility and much higher extensional viscosity in the longitudinal direction [23–25]. Additionally, as expected and shown in Fig. 3, a reduced initial rate requires a significantly longer time to reach the steady force region (100 N) essential for viscosity measurements. Fig. 4 presents the relationship between force increase and thickness reduction for both investigated thicknesses. As shown in Figs. 4a and b, a lower initial rate leads to a higher thickness reduction before achieving the constant force necessary to enter the evaluation region. This is more severe for the thicker specimens (t8).

<figure><img src="image_625.png" /><figcaption>Figure 3 – Thickness reduction curves before and after compliance calibration for specimens with the thickness of 5 mm and from batch (a) t5_T180_r05, (b) t5_T180_r5.</figcaption></figure>

The obtained viscosity values are reported in Table 3. As shown, all the values are in the range of 10 MPa·s and higher. This is at least an order of magnitude higher than the values reported for short flax/PP composite up to 30 wt% of flax fibres which were in the range of 10⁴-10⁶ Pa·s for [18,19]. Looking at the effect of the testing parameters, as expected, a higher temperature of 190 °C resulted in a lower viscosity value for both thicknesses at a constant initial compaction rate of 0.5 mm/min. However, increasing the initial rate at a constant thickness and temperature resulted in a lower viscosity. This can be explained by Fig. 4, which shows a higher deformation of the r05 specimens before reaching the constant load region. Increased compaction and reduced thickness would result in more interaction and fibre entanglement (locking) that can act as a barrier impeding material flow, ultimately contributing to elevated viscosity.
## Results and discussion

Reduction in thickness over time is depicted in Fig. 3, along with the corresponding force changes. In both cases, it is evident that the adjusted thickness following compliance calibration is just slightly higher than the thickness reduction determined by the original displacement of the machine. This confirms the negligible compliance of the setup, allowing the direct utilisation of the original displacement of the machine for data reduction. Nevertheless, it is crucial to validate this step to guarantee accurate thickness reduction measurement, independent of machine compliance, as emphasized in [8].

It is worth mentioning that for all tested samples, significant dimensional change was observed in the transverse direction, confirming squeeze flow as the primary deformation mechanism, as there was no increase in the specimen length due to fibres inextensibility and much higher extensional viscosity in the longitudinal direction [23–25]. Additionally, as expected and shown in Fig. 3, a reduced initial rate requires a significantly longer time to reach the steady force region (100 N) essential for viscosity measurements. Fig. 4 presents the relationship between force increase and thickness reduction for both investigated thicknesses. As shown in Figs. 4a and b, a lower initial rate leads to a higher thickness reduction before achieving the constant force necessary to enter the evaluation region. This is more severe for the thicker specimens (t8).

<figure><img src="image_636.png" /><figcaption>Figure 3 – Thickness reduction curves before and after compliance calibration for specimens with the thickness of 5 mm and from batch (a) t5_T180_r05, (b) t5_T180_r5.</figcaption></figure>

The obtained viscosity values are reported in Table 3. As shown, all the values are in the range of 10 MPa·s and higher. This is at least an order of magnitude higher than the values reported for short flax/PP composite up to 30 wt% of flax fibres which were in the range of 10⁴-10⁶ Pa·s for [18,19]. Looking at the effect of the testing parameters, as expected, a higher temperature of 190 °C resulted in a lower viscosity value for both thicknesses at a constant initial compaction rate of 0.5 mm/min. However, increasing the initial rate at a constant thickness and temperature resulted in a lower viscosity. This can be explained by Fig. 4, which shows a higher deformation of the r05 specimens before reaching the constant load region. Increased compaction and reduced thickness would result in more interaction and fibre entanglement (locking) that can act as a barrier impeding material flow, ultimately contributing to elevated viscosity.
<figure><img src="image_626.png" /><figcaption>Figure 4 – Thickness reduction for specimens from different batches tested at 180 °C. (a) thickness of 5 mm, (b) thickness of 8 mm.</figcaption></figure>

This could be more severe in the case of natural fibres as the fibres are not perfectly straight, and they pose variations in the fibres diameter. It is worth noting that the highest standard deviation is obtained for both batches with the lowest compaction rate of 0.5 mm/min (r05) at 180 °C, while the coefficient of variation (CoV) is higher for t5_T180_r05. A higher temperature of 190 °C (T190) shows a lower standard deviation and CoV than those tested at 180 °C (T180). However, more experiments are still needed for some batches, but the standard deviation is lowered by increasing the initial compaction rate for the cases tested at 180 °C.

It should be noted that the material is treated as a Newtonian fluid in the current work, which serves as one of the first works on the viscosity measurement of UD flax/PP laminates. However, a shear-thinning behaviour is expected by increasing the shear rate for most thermoplastic tapes, which is not considered here. More complex models can be studied using the models based on the Carreau fluid [26] or power law behaviour [27].

<table><caption>Table 3 – Measured effective viscosity of UD flax/PP batches with the standard deviation and coefficient of variation.</caption><thead><tr><th>Batch</th><th>Number of tested specimens</th><th>Min. viscosity (× 10<sup>6</sup> Pa·s)</th><th>Max. viscosity (× 10<sup>6</sup> Pa·s)</th><th>Avg. viscosity (× 10<sup>6</sup> Pa·s)</th><th>Standard deviation (× 10<sup>6</sup> Pa·s)</th><th>CoV (%)</th></tr></thead><tbody><tr><td>t5_T180_r05</td><td>5</td><td>18.5</td><td>34.9</td><td>28.2</td><td>7.5</td><td>26.8</td></tr><tr><td>t5_T180_r5</td><td>2</td><td>10.4</td><td>11.3</td><td>10.8</td><td>0.6</td><td>5.8</td></tr><tr><td>t5_T180_r20</td><td>2</td><td>6.0</td><td>10.5</td><td>8.2</td><td>3.1</td><td>38.1</td></tr><tr><td>t5_T190_r05</td><td>4</td><td>18.8</td><td>27.9</td><td>22.3</td><td>4.4</td><td>4.4</td></tr><tr><td>t8_T180_r05</td><td>4</td><td>60.5</td><td>75.5</td><td>66.6</td><td>6.5</td><td>9.8</td></tr><tr><td>t8_T180_r5</td><td>3</td><td>37.3</td><td>43.7</td><td>41.0</td><td>3.3</td><td>8.1</td></tr><tr><td>t8_T180_r20</td><td>3</td><td>34.0</td><td>37.3</td><td>35.7</td><td>1.7</td><td>4.6</td></tr><tr><td>t8_T190_r05</td><td>4</td><td>53.9</td><td>60.2</td><td>56</td><td>2.9</td><td>5.1</td></tr></tbody></table>

Summary

This paper reported the viscosity data for UD flax/PP composites tested at different conditions, which may be a base for the formulation of a benchmarking exercise. The primary objective was to identify critical aspects of the test method necessary for obtaining reliable results and to propose a recommended test procedure for viscosity measurements in reinforced thermoplastic composites. Transverse squeeze flow experiments were conducted to achieve this goal. Although several models are available for this measurement based on the assumption of fluid behaviour, a simple
<figure><img src="image_637.png" /><figcaption>Figure 4 – Thickness reduction for specimens from different batches tested at 180 °C. (a) thickness of 5 mm, (b) thickness of 8 mm.</figcaption></figure>

This could be more severe in the case of natural fibres as the fibres are not perfectly straight, and they pose variations in the fibres diameter. It is worth noting that the highest standard deviation is obtained for both batches with the lowest compaction rate of 0.5 mm/min (r05) at 180 °C, while the coefficient of variation (CoV) is higher for t5_T180_r05. A higher temperature of 190 °C (T190) shows a lower standard deviation and CoV than those tested at 180 °C (T180). However, more experiments are still needed for some batches, but the standard deviation is lowered by increasing the initial compaction rate for the cases tested at 180 °C.

It should be noted that the material is treated as a Newtonian fluid in the current work, which serves as one of the first works on the viscosity measurement of UD flax/PP laminates. However, a shear-thinning behaviour is expected by increasing the shear rate for most thermoplastic tapes, which is not considered here. More complex models can be studied using the models based on the Carreau fluid [26] or power law behaviour [27].

<table><caption>Table 3 – Measured effective viscosity of UD flax/PP batches with the standard deviation and coefficient of variation.</caption><thead><tr><th>Batch</th><th>Number of tested specimens</th><th>Min. viscosity (× 10<sup>6</sup> Pa·s)</th><th>Max. viscosity (× 10<sup>6</sup> Pa·s)</th><th>Avg. viscosity (× 10<sup>6</sup> Pa·s)</th><th>Standard deviation (× 10<sup>6</sup> Pa·s)</th><th>CoV (%)</th></tr></thead><tbody><tr><td>t5_T180_r05</td><td>5</td><td>18.5</td><td>34.9</td><td>28.2</td><td>7.5</td><td>26.8</td></tr><tr><td>t5_T180_r5</td><td>2</td><td>10.4</td><td>11.3</td><td>10.8</td><td>0.6</td><td>5.8</td></tr><tr><td>t5_T180_r20</td><td>2</td><td>6.0</td><td>10.5</td><td>8.2</td><td>3.1</td><td>38.1</td></tr><tr><td>t5_T190_r05</td><td>4</td><td>18.8</td><td>27.9</td><td>22.3</td><td>4.4</td><td>4.4</td></tr><tr><td>t8_T180_r05</td><td>4</td><td>60.5</td><td>75.5</td><td>66.6</td><td>6.5</td><td>9.8</td></tr><tr><td>t8_T180_r5</td><td>3</td><td>37.3</td><td>43.7</td><td>41.0</td><td>3.3</td><td>8.1</td></tr><tr><td>t8_T180_r20</td><td>3</td><td>34.0</td><td>37.3</td><td>35.7</td><td>1.7</td><td>4.6</td></tr><tr><td>t8_T190_r05</td><td>4</td><td>53.9</td><td>60.2</td><td>56</td><td>2.9</td><td>5.1</td></tr></tbody></table>

## Summary

This paper reported the viscosity data for UD flax/PP composites tested at different conditions, which may be a base for the formulation of a benchmarking exercise. The primary objective was to identify critical aspects of the test method necessary for obtaining reliable results and to propose a recommended test procedure for viscosity measurements in reinforced thermoplastic composites. Transverse squeeze flow experiments were conducted to achieve this goal. Although several models are available for this measurement based on the assumption of fluid behaviour, a simple
Newtonian model is considered here to initiate the exercise. The experiments involved compressing a consolidated laminate of UD flax/PP between two parallel plates. The results revealed that the viscosity data depends on the thickness, exhibiting a power law pattern as a function of initial thickness and thickness reduction rate. Altering the sample thickness from t5 mm to t8 mm and conducting the experiment at a comparable temperature and initial rate did not yield consistent results. In these cases, the viscosity can be up to three times higher than that measured in the t5 samples.

Furthermore, at a constant temperature, a higher scatter was observed at the given target thickness for lower initial thickness reduction rates. The measured viscosity shows lower values by increasing the compaction rate, as less time is available for more compaction and fibre locking before entering the evaluation region. The results show that employing a higher initial compaction rate before reaching the constant force regime could be beneficial to avoid possible fibre locking that would cause a considerably higher and erroneous viscosity measurement.

As a recommendation, it can be suggested that the viscosity measurement should be conducted across various laminate thicknesses from thin (1-2 mm) to thick laminate (10-12 mm) while employing a high compaction rate. This approach aims to minimise uncertainty associated with the measured viscosity and improve the reliability of the results. Moreover, thermal expansion at elevated temperatures may be considered as it might alter the initial dimensions of the specimens.

In addition to the mentioned influencing parameters, a comprehensive benchmark exercise could also investigate other parameters such as void content, time effect in the evaluation region, holding time inside the oven, and other models available in the literature.

## Acknowledgment

H. Perrin and M. Bodaghi acknowledge the support from the Luxembourg National Research Fund (FNR) for funding Structural composite material, NATALINA under the research grant INTER/MERA/19/13991124 in frame of M-ERA.Net.

## References

[1] Boisse P, Akkerman R, Carlone P, Kärger L, Lomov SV, Sherwood JA. Advances in composite forming through 25 years of ESAFORM. Int J Mater Form 2022;15:39. https://doi.org/10.1007/s12289-022-01682-8

[2] Picher-Martel G-P. Compression moulding of randomly-oriented strands thermoplastic composites: a study of the flow and deformation mechanisms. PhD Dissertation. McGill University, 2015.

[3] Yassin K, Hojjati M. Processing of thermoplastic matrix composites through automated fiber placement and tape laying methods: A review. Journal of Thermoplastic Composite Materials 2018;31:1676-725. https://doi.org/10.1177/0892705717738305

[4] Grouve WJB, Warnet LL, Rietman B, Visser HA, Akkerman R. Optimization of the tape placement process parameters for carbon-PPS composites. Composites Part A: Applied Science and Manufacturing 2013;50:44-53. https://doi.org/10.1016/j.compositesa.2013.03.003

[5] Schaefer P, Guglhoer T, Sause M, Drechsler K. Development of intimate contact during processing of carbon fiber reinforced Polyamide-6 tapes. Journal of Reinforced Plastics and Composites 2017;36:593-607. https://doi.org/10.1177/0731684416687041

[6] Budelmann D, Schmidt C, Steuernagel L, Meiners D. Adhesion-cohesion balance of prepreg tack in thermoset automated fiber placement. Part 2: Ply-ply cohesion through contact formation and autohesion. Composites Part C: Open Access 2023;12:100396. https://doi.org/10.1016/j.jcomc.2023.100396
Newtonian model is considered here to initiate the exercise. The experiments involved compressing a consolidated laminate of UD flax/PP between two parallel plates. The results revealed that the viscosity data depends on the thickness, exhibiting a power law pattern as a function of initial thickness and thickness reduction rate. Altering the sample thickness from t5 mm to t8 mm and conducting the experiment at a comparable temperature and initial rate did not yield consistent results. In these cases, the viscosity can be up to three times higher than that measured in the t5 samples.

Furthermore, at a constant temperature, a higher scatter was observed at the given target thickness for lower initial thickness reduction rates. The measured viscosity shows lower values by increasing the compaction rate, as less time is available for more compaction and fibre locking before entering the evaluation region. The results show that employing a higher initial compaction rate before reaching the constant force regime could be beneficial to avoid possible fibre locking that would cause a considerably higher and erroneous viscosity measurement.

As a recommendation, it can be suggested that the viscosity measurement should be conducted across various laminate thicknesses from thin (1-2 mm) to thick laminate (10-12 mm) while employing a high compaction rate. This approach aims to minimise uncertainty associated with the measured viscosity and improve the reliability of the results. Moreover, thermal expansion at elevated temperatures may be considered as it might alter the initial dimensions of the specimens.

In addition to the mentioned influencing parameters, a comprehensive benchmark exercise could also investigate other parameters such as void content, time effect in the evaluation region, holding time inside the oven, and other models available in the literature.

## Acknowledgment

H. Perrin and M. Bodaghi acknowledge the support from the Luxembourg National Research Fund (FNR) for funding Structural composite material, NATALINA under the research grant INTER/MERA/19/13991124 in frame of M-ERA.Net.

## References

[1] Boisse P, Akkerman R, Carlone P, Kärger L, Lomov SV, Sherwood JA. Advances in composite forming through 25 years of ESAFORM. Int J Mater Form 2022;15:39. https://doi.org/10.1007/s12289-022-01682-8

[2] Picher-Martel G-P. Compression moulding of randomly-oriented strands thermoplastic composites: a study of the flow and deformation mechanisms. PhD Dissertation. McGill University, 2015.

[3] Yassin K, Hojjati M. Processing of thermoplastic matrix composites through automated fiber placement and tape laying methods: A review. Journal of Thermoplastic Composite Materials 2018;31:1676-725. https://doi.org/10.1177/0892705717738305

[4] Grouve WJB, Warnet LL, Rietman B, Visser HA, Akkerman R. Optimization of the tape placement process parameters for carbon-PPS composites. Composites Part A: Applied Science and Manufacturing 2013;50:44-53. https://doi.org/10.1016/j.compositesa.2013.03.003

[5] Schaefer P, Guglhoer T, Sause M, Drechsler K. Development of intimate contact during processing of carbon fiber reinforced Polyamide-6 tapes. Journal of Reinforced Plastics and Composites 2017;36:593-607. https://doi.org/10.1177/0731684416687041

[6] Budelmann D, Schmidt C, Steuernagel L, Meiners D. Adhesion-cohesion balance of prepreg tack in thermoset automated fiber placement. Part 2: Ply-ply cohesion through contact formation and autohesion. Composites Part C: Open Access 2023;12:100396. https://doi.org/10.1016/j.jcomc.2023.100396
[7] Schaefer PM. Material characterization for determining the consolidation properties of carbon fiber tapes with PA6 matrix. ICCM20, Copenhagen: 2015.

[8] Khan MA, Mitschang P, Schledjewski R. Identification of some optimal parameters to achieve higher laminate quality through tape placement process. Adv Polym Technol 2010;29:98-111. https://doi.org/10.1002/adv.20177

[9] Khodaei A, Shadmehri F. Intimate contact development for automated fiber placement of thermoplastic composites. Composites Part C: Open Access 2022;8:100290. https://doi.org/10.1016/j.jcomc.2022.100290

[10] Yang F, Pitchumani R. A fractal Cantor set based description of interlaminar contact evolution during thermoplastic composites processing. Journal of Materials Science 2001;36:4661-4671. https://doi.org/10.1023/A:1017950215945

[11] Çelik O, Peeters D, Dransfeld C, Teuwen J. Intimate contact development during laser assisted fiber placement: Microstructure and effect of process parameters. Composites Part A: Applied Science and Manufacturing 2020;134:105888. https://doi.org/10.1016/j.compositesa.2020.105888

[12] Kok T. On the consolidation quality in laser assisted fiber placement: the role of the heating phase. PhD Dissertation. University of Twente, 2018. https://doi.org/10.3990/1.9789036546065

[13] Carlesso P. Temperature-dependent viscosity assessment of neat polyamide 6 and unidirectional carbon-fiber reinforced polyamide 6 (UD CF/PA6). MSc Thesis. KU Leuven, 2023.

[14] Engmann J, Servais C, Burbidge AS. Squeeze flow theory and applications to rheometry: A review. Journal of Non-Newtonian Fluid Mechanics 2005;132:1-27. https://doi.org/10.1016/j.jnnfm.2005.08.007

[15] Arbter R, Beraud JM, Binetruy C, Bizet L, Bréard J, Comas-Cardona S, et al. Experimental determination of the permeability of textiles: A benchmark exercise. Composites Part A: Applied Science and Manufacturing 2011;42:1157-68. https://doi.org/10.1016/j.compositesa.2011.04.021

[16] Vernet N, Ruiz E, Advani S, Alms JB, Aubert M, Barburski M, et al. Experimental determination of the permeability of engineering textiles: Benchmark II. Composites Part A: Applied Science and Manufacturing 2014;61:172-84. https://doi.org/10.1016/j.compositesa.2014.02.010

[17] May D, Aktas A, Advani SG, Berg DC, Endruweit A, Fauster E, et al. In-plane permeability characterization of engineering textiles based on radial flow experiments: A benchmark exercise. Composites Part A: Applied Science and Manufacturing 2019;121:100-14. https://doi.org/10.1016/j.compositesa.2019.03.006

[18] Sojoudiasli H, Heuzey M-C, Carreau PJ. Rheological, morphological and mechanical properties of flax fiber polypropylene composites: influence of compatibilizers. Cellulose 2014;21:3797-812. https://doi.org/10.1007/s10570-014-0375-3

[19] Doumbia AS, Castro M, Jouannet D, Kervoëlen A, Falher T, Cauret L, et al. Flax/polypropylene composites for lightened structures: Multiscale analysis of process and fibre parameters. Materials & Design 2015;87:331-41. https://doi.org/10.1016/j.matdes.2015.07.139

[20] Mihai M, Denault J. Interrelation Between Melt Processing Conditions, Formulation And Properties Of Polypropylene / Short Flax And Hemp Fiber Composites, Boston, Massachusetts, USA: 2011.

[21] Lin H-R, Advani SG. Processing models and characterization of thermoplastic composite wound parts. Polym Compos 1997;18:405-11. https://doi.org/10.1002/pc.10291
[7] Schaefer PM. Material characterization for determining the consolidation properties of carbon fiber tapes with PA6 matrix. ICCM20, Copenhagen: 2015.

[8] Khan MA, Mitschang P, Schledjewski R. Identification of some optimal parameters to achieve higher laminate quality through tape placement process. Adv Polym Technol 2010;29:98-111. https://doi.org/10.1002/adv.20177

[9] Khodaei A, Shadmehri F. Intimate contact development for automated fiber placement of thermoplastic composites. Composites Part C: Open Access 2022;8:100290. https://doi.org/10.1016/j.jcomc.2022.100290

[10] Yang F, Pitchumani R. A fractal Cantor set based description of interlaminar contact evolution during thermoplastic composites processing. Journal of Materials Science 2001;36:4661-4671. https://doi.org/10.1023/A:1017950215945

[11] Çelik O, Peeters D, Dransfeld C, Teuwen J. Intimate contact development during laser assisted fiber placement: Microstructure and effect of process parameters. Composites Part A: Applied Science and Manufacturing 2020;134:105888. https://doi.org/10.1016/j.compositesa.2020.105888

[12] Kok T. On the consolidation quality in laser assisted fiber placement: the role of the heating phase. PhD Dissertation. University of Twente, 2018. https://doi.org/10.3990/1.9789036546065

[13] Carlesso P. Temperature-dependent viscosity assessment of neat polyamide 6 and unidirectional carbon-fiber reinforced polyamide 6 (UD CF/PA6). MSc Thesis. KU Leuven, 2023.

[14] Engmann J, Servais C, Burbidge AS. Squeeze flow theory and applications to rheometry: A review. Journal of Non-Newtonian Fluid Mechanics 2005;132:1-27. https://doi.org/10.1016/j.jnnfm.2005.08.007

[15] Arbter R, Beraud JM, Binetruy C, Bizet L, Bréard J, Comas-Cardona S, et al. Experimental determination of the permeability of textiles: A benchmark exercise. Composites Part A: Applied Science and Manufacturing 2011;42:1157-68. https://doi.org/10.1016/j.compositesa.2011.04.021

[16] Vernet N, Ruiz E, Advani S, Alms JB, Aubert M, Barburski M, et al. Experimental determination of the permeability of engineering textiles: Benchmark II. Composites Part A: Applied Science and Manufacturing 2014;61:172-84. https://doi.org/10.1016/j.compositesa.2014.02.010

[17] May D, Aktas A, Advani SG, Berg DC, Endruweit A, Fauster E, et al. In-plane permeability characterization of engineering textiles based on radial flow experiments: A benchmark exercise. Composites Part A: Applied Science and Manufacturing 2019;121:100-14. https://doi.org/10.1016/j.compositesa.2019.03.006

[18] Sojoudiasli H, Heuzey M-C, Carreau PJ. Rheological, morphological and mechanical properties of flax fiber polypropylene composites: influence of compatibilizers. Cellulose 2014;21:3797-812. https://doi.org/10.1007/s10570-014-0375-3

[19] Doumbia AS, Castro M, Jouannet D, Kervoëlen A, Falher T, Cauret L, et al. Flax/polypropylene composites for lightened structures: Multiscale analysis of process and fibre parameters. Materials & Design 2015;87:331-41. https://doi.org/10.1016/j.matdes.2015.07.139

[20] Mihai M, Denault J. Interrelation Between Melt Processing Conditions, Formulation And Properties Of Polypropylene / Short Flax And Hemp Fiber Composites, Boston, Massachusetts, USA: 2011.

[21] Lin H-R, Advani SG. Processing models and characterization of thermoplastic composite wound parts. Polym Compos 1997;18:405-11. https://doi.org/10.1002/pc.10291
[22] Yong AXH, Aktas A, May D, Endruweit A, Lomov SV, Advani S, et al. Experimental characterisation of textile compaction response: A benchmark exercise. Composites Part A: Applied Science and Manufacturing 2021;142:106243. https://doi.org/10.1016/j.compositesa.2020.106243

[23] Goshawk JA, Navez VP, Jones RS. Squeezing flow of continuous fibre-reinforced composites. Journal of Non-Newtonian Fluid Mechanics 1997;73:327-42. https://doi.org/10.1016/S0377-0257(97)00049-9

[24] Picher-Martel G-P, Levy A, Hubert P. Compression molding of Carbon/Polyether ether ketone composites: Squeeze flow behavior of unidirectional and randomly oriented strands. Polymer Composites 2017;38:1828-37.

[25] Valverde MA, Belnoue JP-H, Kupfer R, Kawashita LF, Gude M, Hallett SR. Compaction behaviour of continuous fibre-reinforced thermoplastic composites under rapid processing conditions. Composites Part A: Applied Science and Manufacturing 2021;149:106549. https://doi.org/10.1016/j.compositesa.2021.106549

[26] Shuler SF, Advani SG. Transverse squeeze flow of concentrated aligned fibers in viscous fluids. Journal of Non-Newtonian Fluid Mechanics 1996;65:47-74. https://doi.org/10.1016/0377-0257(96)01440-1

[27] Arquier R, Miquelard-Garnier G, Iliopoulos I, Régnier G. Assessing the shear viscous behavior of continuous carbon fiber reinforced PEKK composites with squeeze flow measurements. Polymer Testing 2023;123:108060. https://doi.org/10.1016/j.polymertesting.2023.108060
[22] Yong AXH, Aktas A, May D, Endruweit A, Lomov SV, Advani S, et al. Experimental characterisation of textile compaction response: A benchmark exercise. Composites Part A: Applied Science and Manufacturing 2021;142:106243. https://doi.org/10.1016/j.compositesa.2020.106243

[23] Goshawk JA, Navez VP, Jones RS. Squeezing flow of continuous fibre-reinforced composites. Journal of Non-Newtonian Fluid Mechanics 1997;73:327-42. https://doi.org/10.1016/S0377-0257(97)00049-9

[24] Picher-Martel G-P, Levy A, Hubert P. Compression molding of Carbon/Polyether ether ketone composites: Squeeze flow behavior of unidirectional and randomly oriented strands. Polymer Composites 2017;38:1828-37.

[25] Valverde MA, Belnoue JP-H, Kupfer R, Kawashita LF, Gude M, Hallett SR. Compaction behaviour of continuous fibre-reinforced thermoplastic composites under rapid processing conditions. Composites Part A: Applied Science and Manufacturing 2021;149:106549. https://doi.org/10.1016/j.compositesa.2021.106549

[26] Shuler SF, Advani SG. Transverse squeeze flow of concentrated aligned fibers in viscous fluids. Journal of Non-Newtonian Fluid Mechanics 1996;65:47-74. https://doi.org/10.1016/0377-0257(96)01440-1

[27] Arquier R, Miquelard-Garnier G, Iliopoulos I, Régnier G. Assessing the shear viscous behavior of continuous carbon fiber reinforced PEKK composites with squeeze flow measurements. Polymer Testing 2023;123:108060. https://doi.org/10.1016/j.polymertesting.2023.108060
Extrusion and drawing

<figure><img src="image_627.png" /></figure>
Extrusion and drawing

<figure><img src="image_638.png" /></figure>
# Direct recycling of AA6063 chips by hot extrusion applying pseudo porthole die

FUNAZUKA Tatsuya<fnref n="1" />, a*, YAMASHITA Syun<fnref n="2" />, b, URAKAWA Takumi<fnref n="2" />, c, SHIRATORI Tomomi<fnref n="1" />, d, TAKATSUJI Norio<fnref n="1" />, e, and DOHDA Kuniaki<fnref n="3" />, f

**Keywords:** Hot Extrusion, 6063 Aluminum Alloy, Waste Chip, Pseudo Porthole Die, Recycling

**Abstract.** To obtain recycled aluminum from sash waste with A6063 alloy, which is used for aluminum sashes, it is necessary to perform hot extrusion after a casting process. Impurities introduced during casting constitute a problem in this process, and recycling without casting is necessary to control the effects of these impurities. The purpose of this study is to develop a recycling method for aluminum alloys, which does not involve the casting process, using hot extrusion from scrap materials, such as the cutting chips of A6063 alloy. In the extrusion process of waste materials such as cutting chips, the chips should be kept in close contact. To achieve this, a pseudo porthole die that provides a large shear deformation is used as the die shape for the extrusion process. The chips were compression-filled in advance to determine the filling ratio, and then extruded using a pseudo porthole die. The use of a pseudo porthole die resulted in an extrusion force 1.5 times greater than that of a solid die. A simulation of the shear strain and observation of the stress imparted to the material by the pseudo porthole die revealed that the flow-diverting and crimping mechanisms inside the pseudo porthole die resulted in a large amount of stress on the material. No defects were observed on the extruded product surface of the chip material produced using the pseudo porthole die. Four types of chip materials with different densities were used to examine the effect of the chip material density. We found that defects did not occur on the surface of the product when the density of the chip material was 2.5 g/cm<fnref n="3" /> or higher.

**Introduction**

Aluminum (Al) is one of the most widely used metal materials worldwide and is expected that its use would continue to increase in the future owing to its light weight and high workability [1]. Until now, Al-elongated materials have mainly been applied to residential building materials such as aluminum sashes; however, they are now being actively used in transportation equipment such as automobiles and aircraft to replace steel materials owing to their lightweight property [2, 3].

A major drawback of Al production is that a large amount of energy is required to extract Al from alumina [4]. However, Al is highly recyclable, and the use of recycled Al can reduce energy consumption by 95% [5]. Existing Al recycling technologies include remelting and refining. In the case of recycling by remelting, the removal and detoxification of impurities during casting are major issues, and research is being conducted on how to adjust the material composition and control casting defects [6]. In the case of recycling expanded materials with high Al purity such as aluminum sashes, there are many cases in which the materials are recycled into low-purity casting

<figure><img src="image_628.png" /></figure>

<footnote n="1">¹Academic Assembly Faculty of Engineering, University of Toyama, Toyama, Japan</footnote>

<footnote n="2">²Graduate School of Science and Engineering for Education, University of Toyama, Toyama, Japan</footnote>

<footnote n="3">³Department of Mechanical Engineering, Northwestern University, Evanston, Illinois, USA</footnote>

<footnote>a funazuka@eng.u-toyama.ac.jp; b m22c1505@ems.u-toyama.ac.jp;</footnote>

<footnote>c s2070414@ems.u-toyama.ac.jp; d shira@eng.u-toyama.ac.jp; e michinori0204@gmail.com; f</footnote>

<footnote>dohda.kuni@northwestern.edu</footnote>
# Direct recycling of AA6063 chips by hot extrusion applying pseudo porthole die

FUNAZUKA Tatsuya<fnref n="1" />, a*, YAMASHITA Syun<fnref n="2" />, b, URAKAWA Takumi<fnref n="2" />, c, SHIRATORI Tomomi<fnref n="1" />, d, TAKATSUJI Norio<fnref n="1" />, e, and DOHDA Kuniaki<fnref n="3" />, f

**Keywords:** Hot Extrusion, 6063 Aluminum Alloy, Waste Chip, Pseudo Porthole Die, Recycling

**Abstract.** To obtain recycled aluminum from sash waste with A6063 alloy, which is used for aluminum sashes, it is necessary to perform hot extrusion after a casting process. Impurities introduced during casting constitute a problem in this process, and recycling without casting is necessary to control the effects of these impurities. The purpose of this study is to develop a recycling method for aluminum alloys, which does not involve the casting process, using hot extrusion from scrap materials, such as the cutting chips of A6063 alloy. In the extrusion process of waste materials such as cutting chips, the chips should be kept in close contact. To achieve this, a pseudo porthole die that provides a large shear deformation is used as the die shape for the extrusion process. The chips were compression-filled in advance to determine the filling ratio, and then extruded using a pseudo porthole die. The use of a pseudo porthole die resulted in an extrusion force 1.5 times greater than that of a solid die. A simulation of the shear strain and observation of the stress imparted to the material by the pseudo porthole die revealed that the flow-diverting and crimping mechanisms inside the pseudo porthole die resulted in a large amount of stress on the material. No defects were observed on the extruded product surface of the chip material produced using the pseudo porthole die. Four types of chip materials with different densities were used to examine the effect of the chip material density. We found that defects did not occur on the surface of the product when the density of the chip material was 2.5 g/cm<fnref n="3" /> or higher.

**Introduction**

Aluminum (Al) is one of the most widely used metal materials worldwide and is expected that its use would continue to increase in the future owing to its light weight and high workability [1]. Until now, Al-elongated materials have mainly been applied to residential building materials such as aluminum sashes; however, they are now being actively used in transportation equipment such as automobiles and aircraft to replace steel materials owing to their lightweight property [2, 3].

A major drawback of Al production is that a large amount of energy is required to extract Al from alumina [4]. However, Al is highly recyclable, and the use of recycled Al can reduce energy consumption by 95% [5]. Existing Al recycling technologies include remelting and refining. In the case of recycling by remelting, the removal and detoxification of impurities during casting are major issues, and research is being conducted on how to adjust the material composition and control casting defects [6]. In the case of recycling expanded materials with high Al purity such as aluminum sashes, there are many cases in which the materials are recycled into low-purity casting

<figure><img src="image_639.png" /></figure>

<footnote n="1">¹Academic Assembly Faculty of Engineering, University of Toyama, Toyama, Japan</footnote>

<footnote n="2">²Graduate School of Science and Engineering for Education, University of Toyama, Toyama, Japan</footnote>

<footnote n="3">³Department of Mechanical Engineering, Northwestern University, Evanston, Illinois, USA</footnote>

<footnote>a funazuka@eng.u-toyama.ac.jp; b m22c1505@ems.u-toyama.ac.jp;</footnote>

<footnote>c s2070414@ems.u-toyama.ac.jp; d shira@eng.u-toyama.ac.jp; e michinori0204@gmail.com; f</footnote>

<footnote>dohda.kuni@northwestern.edu</footnote>
materials by remelting (cascade recycling). The total amount of horizontally recycled aluminum sash in Japan is only 10% [7]. It has been reported that a further reduction in energy consumption can be achieved if recycling is possible by extrusion or forging only without casting [8].

To eliminate the casting process, the products should be obtained from the chip material after crushing by solidification compaction only. Extrusion is a material processing method used for the solidification and formation of chip materials [9-11]. Despite the demands in the industry, this method has not yet been realized.

The author conducted research on Al product surface defects, such as pickup defects, during the extrusion of sash materials [12]. In general, a porthole die is used to extrude hollow materials with complex shapes, such as sash materials, from solid billets. To obtain a hollow material, a porthole die enables the formation of complex shapes by means of a port section that diverts the material and a chamber section that crimps the material. In this study, a pseudo porthole die is proposed to reproduce "pseudo" porthole extrusion, and the generation mechanism of blistering defects is determined after reproducing the flow-diverting and crimping behavior. This study also demonstrated that the material inside the pseudo porthole die underwent strong shear deformation owing to flow-diverting and crimping behaviors, resulting in grain size reduction. Tekkaya et al. reported that the use of a porthole die can solve the technical problems of chip extrusion, which have been considered difficult for industrial implementation due to problems of material adhesion and internal defects in the product [13-15]. This is because a porthole die applies high shear stress to the material, which destroys the oxide film of Al and increases its adhesion strength. The increase in the adhesion strength also improves the bonding strength and surface quality of the material. However, no detailed study has been conducted on the density of the chip material. Although previous studies concluded that a larger density is better, these existing studies reported different values. In addition, there were only a few cases in which the density was varied significantly and its limiting values were investigated.

In this study, AA6063 chips were hot-extruded using a pseudo porthole die to achieve direct horizontal recycling of sash material. The chip material was fabricated by turning AA6063 ingots and compressing them into billets for extrusion. Extrusion experiments were conducted by quantitatively varying the density of the chip material and the strength of compression forming. The effects of extrusion formability and chip density on the surface properties of the product were evaluated by applying a pseudo porthole die.

## Experimental Methods

In this study, AA6063 was used as the chip material. Table 1 lists the chemical composition of the AA6063 used in this study.

<table><caption>Table 1 Chemical composition of AA6063 [mass%]</caption><tr><td></td><td>Si</td><td>Fe</td><td>Cu</td><td>Mn</td><td>Mg</td><td>Cr</td><td>Zn</td><td>Ti</td><td>Al</td></tr><tr><td>AA6063</td><td>0.49</td><td>0.18</td><td>0.02</td><td>0.02</td><td>0.54</td><td>0.01</td><td>0.01</td><td>0.01</td><td>Bal.</td></tr></table>

The chips were extracted by turning a cast AA6063 billet into a billet of chip material. The cutting conditions were a depth of cut of 0.5 mm and rotation speed of 920 rpm, and chips with a width of approximately 2 mm were extracted. The chips were compacted and preformed on a 30 t Amsler-type universal testing machine with a load of 3 t, as shown in Fig. 1 (a). The chips were inserted into the container, and a preformed chip billet with a diameter of 38.5 mm and length of 80 mm was formed.
materials by remelting (cascade recycling). The total amount of horizontally recycled aluminum sash in Japan is only 10% [7]. It has been reported that a further reduction in energy consumption can be achieved if recycling is possible by extrusion or forging only without casting [8].

To eliminate the casting process, the products should be obtained from the chip material after crushing by solidification compaction only. Extrusion is a material processing method used for the solidification and formation of chip materials [9-11]. Despite the demands in the industry, this method has not yet been realized.

The author conducted research on Al product surface defects, such as pickup defects, during the extrusion of sash materials [12]. In general, a porthole die is used to extrude hollow materials with complex shapes, such as sash materials, from solid billets. To obtain a hollow material, a porthole die enables the formation of complex shapes by means of a port section that diverts the material and a chamber section that crimps the material. In this study, a pseudo porthole die is proposed to reproduce "pseudo" porthole extrusion, and the generation mechanism of blistering defects is determined after reproducing the flow-diverting and crimping behavior. This study also demonstrated that the material inside the pseudo porthole die underwent strong shear deformation owing to flow-diverting and crimping behaviors, resulting in grain size reduction. Tekkaya et al. reported that the use of a porthole die can solve the technical problems of chip extrusion, which have been considered difficult for industrial implementation due to problems of material adhesion and internal defects in the product [13-15]. This is because a porthole die applies high shear stress to the material, which destroys the oxide film of Al and increases its adhesion strength. The increase in the adhesion strength also improves the bonding strength and surface quality of the material. However, no detailed study has been conducted on the density of the chip material. Although previous studies concluded that a larger density is better, these existing studies reported different values. In addition, there were only a few cases in which the density was varied significantly and its limiting values were investigated.

In this study, AA6063 chips were hot-extruded using a pseudo porthole die to achieve direct horizontal recycling of sash material. The chip material was fabricated by turning AA6063 ingots and compressing them into billets for extrusion. Extrusion experiments were conducted by quantitatively varying the density of the chip material and the strength of compression forming. The effects of extrusion formability and chip density on the surface properties of the product were evaluated by applying a pseudo porthole die.

## Experimental Methods

In this study, AA6063 was used as the chip material. Table 1 lists the chemical composition of the AA6063 used in this study.

<table><caption>Table 1 Chemical composition of AA6063 [mass%]</caption><thead><tr><th></th><th>Si</th><th>Fe</th><th>Cu</th><th>Mn</th><th>Mg</th><th>Cr</th><th>Zn</th><th>Ti</th><th>Al</th></tr></thead><tbody><tr><td>AA6063</td><td>0.49</td><td>0.18</td><td>0.02</td><td>0.02</td><td>0.54</td><td>0.01</td><td>0.01</td><td>0.01</td><td>Bal.</td></tr></tbody></table>

The chips were extracted by turning a cast AA6063 billet into a billet of chip material. The cutting conditions were a depth of cut of 0.5 mm and rotation speed of 920 rpm, and chips with a width of approximately 2 mm were extracted. The chips were compacted and preformed on a 30 t Amsler-type universal testing machine with a load of 3 t, as shown in Fig. 1 (a). The chips were inserted into the container, and a preformed chip billet with a diameter of 38.5 mm and length of 80 mm was formed.
<figure><img src="image_629.png" /><figcaption>Fig. 1 (a) 30 t Amsler-type universal testing machine, (b) Schematic diagram of pre-solidification compacting</figcaption></figure>

After pre-solidification and formation, the billets were solidified and formed in a 200 horizontal hydraulic extrusion press, as shown in Fig. 2 (a). A cartridge heater and thermocouple were mounted inside the container, and the billet temperature and ram speed were controlled using a controller. The extrusion force during the hot extrusion was measured using a pressure sensor attached to the hydraulic controller of the press, and the ram stroke was measured using a laser displacement meter. The solidification-forming die shown in Fig. 2 (b) was used for solidification forming. The solidification forming conditions are listed in Table 2. The ram speed was 0.5 mm/s. To change the density of the solidified compact, the billet and die temperatures were set to 25 °C, 100 °C, and 200 °C, and the load values were 15 t, 50 t, and 100 t. The obtained solidified and molded billets were sampled with a diameter of 41.5 mm and a length of approximately 30 mm considering the container diameter during the extrusion experiment.

In this study, a 200 horizontal hydraulic extrusion press was also used for the hot extrusion experiments, as shown in Fig. 2 (a). A solid die, shown in Fig. 3 (a), and a pseudoporthole die, shown in Fig. 3 (b), were used in the extrusion experiments [12]. To reproduce the joining process in the chamber section in an easy-to-understand manner, the pseudo porthole die developed in this study was simplified by eliminating the mandrel portion that formed the hollow section of the general purpose porthole die and designed to reproduce only the flow-diverting and joining processes inside the die. The port length of the pseudo porthole die was 25 mm and the thickness of the port section was 6 mm. The inner diameter of the port and chamber section was φ25 mm and the length of the chamber section was 20 mm. The shape of the extruded profile was a solid plate with a width of 18.6 mm and a thickness of 1.5 mm, as shown in Fig. 3 (c).

<figure><img src="image_630.png" /><figcaption>Fig. 2 (a) 200 t horizontal extrusion machine, (b) Compacting die</figcaption></figure>
<figure><img src="image_640.png" /><figcaption>Fig. 1 (a) 30 t Amsler-type universal testing machine, (b) Schematic diagram of pre-solidification compacting</figcaption></figure>

After pre-solidification and formation, the billets were solidified and formed in a 200 horizontal hydraulic extrusion press, as shown in Fig. 2 (a). A cartridge heater and thermocouple were mounted inside the container, and the billet temperature and ram speed were controlled using a controller. The extrusion force during the hot extrusion was measured using a pressure sensor attached to the hydraulic controller of the press, and the ram stroke was measured using a laser displacement meter. The solidification-forming die shown in Fig. 2 (b) was used for solidification forming. The solidification forming conditions are listed in Table 2. The ram speed was 0.5 mm/s. To change the density of the solidified compact, the billet and die temperatures were set to 25 °C, 100 °C, and 200 °C, and the load values were 15 t, 50 t, and 100 t. The obtained solidified and molded billets were sampled with a diameter of 41.5 mm and a length of approximately 30 mm considering the container diameter during the extrusion experiment.

In this study, a 200 horizontal hydraulic extrusion press was also used for the hot extrusion experiments, as shown in Fig. 2 (a). A solid die, shown in Fig. 3 (a), and a pseudoporthole die, shown in Fig. 3 (b), were used in the extrusion experiments [12]. To reproduce the joining process in the chamber section in an easy-to-understand manner, the pseudo porthole die developed in this study was simplified by eliminating the mandrel portion that formed the hollow section of the general purpose porthole die and designed to reproduce only the flow-diverting and joining processes inside the die. The port length of the pseudo porthole die was 25 mm and the thickness of the port section was 6 mm. The inner diameter of the port and chamber section was φ25 mm and the length of the chamber section was 20 mm. The shape of the extruded profile was a solid plate with a width of 18.6 mm and a thickness of 1.5 mm, as shown in Fig. 3 (c).

<figure><img src="image_641.png" /><figcaption>Fig. 2 (a) 200 t horizontal extrusion machine, (b) Compacting die</figcaption></figure>
The experimental conditions for the extrusion are listed in Table 2. The extrusion temperature was 500 °C and the ram speed was 0.5 mm. The billet dimensions were φ41.5 mm and length of 90 mm, which was obtained by stacking three solidified and molded billets, as shown in Fig. 2. The billet material was AA6063 solid billet (virgin material) and AA6063 chip billet under four conditions, and the die configuration was a solid die (Fig. 3 (a)) and a pseudo porthole die (Fig. 3 (b)).

The surface defects at the head, middle, and end of the extrudate obtained in the extrusion experiments were observed. An optical microscope (Optical microscope, OM; Nikon ECLIPSE LV150N) was used to observe the surface of the extrudate.

<table><caption>Table 2 Compacting and extrusion conditions</caption><thead><tr><th colspan="2">Compacting condition</th></tr></thead><tbody><tr><td>Billet material</td><td>AA6063</td></tr><tr><td>Billet temperature [°C]</td><td>25, 100, 200</td></tr><tr><td>Die temperature [°C]</td><td>25, 100, 200</td></tr><tr><td>Ram speed [mm/s]</td><td>0.5</td></tr><tr><td>Die shape</td><td>Compacting die</td></tr><tr><td>Compacting force [t]</td><td>15, 50, 100</td></tr><tr><th colspan="2">Extrusion condition</th></tr><tr><td>Billet material</td><td>AA6063 solid (virgin material)<br>AA6063 chip</td></tr><tr><td>Extrusion temperature (billet and die) [°C]</td><td>500</td></tr><tr><td>Ram speed [mm/s]</td><td>0.5</td></tr><tr><td>Die structure</td><td>Solid die<br>Pseudo porthole die</td></tr></tbody></table>

<figure><img src="image_631.png" /><figcaption>(a) Solid die</figcaption></figure>

<figure><img src="image_632.png" /><figcaption>(C) Die shape of extrusion</figcaption></figure>

<figure><img src="image_633.png" /><figcaption>(b) Pseudo port-hole die</figcaption></figure>

<figure><img src="image_634.png" /><figcaption>(d) Die set</figcaption></figure>

<caption>Fig. 3 Schematic diagram of the die and die set used in the hot extrusion experiments [12]</caption>

The conditions for the hot extrusion simulations are listed in Table 3. DEFORM-3D was used as the simulation software and the Lagrangian simulation method was used. The material used was AA6063 and the shear friction coefficient was derived from a friction test developed for hot plastic
The experimental conditions for the extrusion are listed in Table 2. The extrusion temperature was 500 °C and the ram speed was 0.5 mm. The billet dimensions were φ41.5 mm and length of 90 mm, which was obtained by stacking three solidified and molded billets, as shown in Fig. 2. The billet material was AA6063 solid billet (virgin material) and AA6063 chip billet under four conditions, and the die configuration was a solid die (Fig. 3 (a)) and a pseudo porthole die (Fig. 3 (b)).

The surface defects at the head, middle, and end of the extrudate obtained in the extrusion experiments were observed. An optical microscope (Optical microscope, OM; Nikon ECLIPSE LV150N) was used to observe the surface of the extrudate.

<table><caption>Table 2 Compacting and extrusion conditions</caption><thead><tr><th colspan="2">Compacting condition</th></tr></thead><tbody><tr><td>Billet material</td><td>AA6063</td></tr><tr><td>Billet temperature [°C]</td><td>25, 100, 200</td></tr><tr><td>Die temperature [°C]</td><td>25, 100, 200</td></tr><tr><td>Ram speed [mm/s]</td><td>0.5</td></tr><tr><td>Die shape</td><td>Compacting die</td></tr><tr><td>Compacting force [t]</td><td>15, 50, 100</td></tr><tr><th colspan="2">Extrusion condition</th></tr><tr><td>Billet material</td><td>AA6063 solid (virgin material)<br>AA6063 chip</td></tr><tr><td>Extrusion temperature (billet and die) [°C]</td><td>500</td></tr><tr><td>Ram speed [mm/s]</td><td>0.5</td></tr><tr><td>Die structure</td><td>Solid die<br>Pseudo porthole die</td></tr></tbody></table>

<figure><img src="image_642.png" /><figcaption>(a) Solid die</figcaption></figure>

<figure><img src="image_643.png" /><figcaption>(C) Die shape of extrusion</figcaption></figure>

<figure><img src="image_644.png" /><figcaption>(b) Pseudo port-hole die</figcaption></figure>

<figure><img src="image_645.png" /><figcaption>(d) Die set</figcaption></figure>

<caption>Fig. 3 Schematic diagram of the die and die set used in the hot extrusion experiments [12]</caption>

The conditions for the hot extrusion simulations are listed in Table 3. DEFORM-3D was used as the simulation software and the Lagrangian simulation method was used. The material used was AA6063 and the shear friction coefficient was derived from a friction test developed for hot plastic
forming based on a previous study [16]. Because the internal structure of the die was symmetrical, the billet and die set in the simulation were modeled as one-quarter models. During the hot extrusion simulation, we observed the stresses applied to the material in the solid die and pseudo porthole die.

<table><caption>Table 3 Simulation conditions</caption><tr><td>Analysis method</td><td>Lagrangian</td></tr><tr><td>Billet material data</td><td>AA6063</td></tr><tr><td>Billet temperature [°C]</td><td>500</td></tr><tr><td>Number of elements</td><td>100000</td></tr><tr><td>Ram speed [mm/s]</td><td>0.5</td></tr><tr><td>Share friction coefficient m</td><td>1.0</td></tr><tr><td>Heat transfer coefficient [N·mm/(s·°C)]</td><td>11</td></tr><tr><td>Die structure</td><td>Solid die<br>Pseudo porthole die</td></tr></table>

## Experimental Results

Fig. 4 shows the solidified compacts compacted at room temperature (25 °C) and 15 t (density: 1.640 g/cm³), at 100 °C and 50 t (density: 2.151 g/cm³), at 100 °C and 100 t (density: 2.466 g/cm³), and at 200 °C and 100 t (density: 2.607 g/cm³). The solidified compacts were compacted at 100 °C and 100 t (density: 2.466 g/cm³) and at 200 °C and 100 t (density: 2.607 g/cm³). The density of Al (2.7 g/cm³) approached that of Al at higher compressive temperatures and forces.

<table><caption>Fig. 4 Solidified compacts at each compression condition</caption><tr><td>Compression conditions</td><td>Room Temperature (R.T.) 15 t</td><td>100 °C 50 t</td><td>100 °C 100 t</td><td>200 °C 100 t</td></tr><tr><td>Appearance</td><td></td><td></td><td></td><td></td></tr><tr><td>Density [g/cm³]</td><td>1.640</td><td>2.151</td><td>2.466</td><td>2.607</td></tr></table>

Hot extrusion experiments were conducted using AA6063 chip billets and chip billets (density: 2.607 g/cm³) to determine the extrudability of the solid die and pseudo porthole die. The extrudability was determined by examining the extrusion force and appearance of the extruded product. In the extrudate appearance investigation, we determined the presence or absence of defects on the surface of the chip material and the size of the defects.

Fig. 5 shows the extrusion force–ram stroke curve for the hot extrusion of a chip billet (density: 2.607 g/cm³) and a solid billet. The maximum extrusion force was 94.7 t using the pseudo porthole die, which is approximately 1.5 times that of the solid die. In the case of extrusion using the pseudo porthole die, the extrusion force increased in two stages, and it is considered that the billet passed through the port section at the first increase in force and passed through the forming section after the second increase in force. No large force difference was observed between the pseudo porthole die and the solid die for the chip and solid billets.
forming based on a previous study [16]. Because the internal structure of the die was symmetrical, the billet and die set in the simulation were modeled as one-quarter models. During the hot extrusion simulation, we observed the stresses applied to the material in the solid die and pseudo porthole die.

<table><caption>Table 3 Simulation conditions</caption><tr><td>Analysis method</td><td>Lagrangian</td></tr><tr><td>Billet material data</td><td>AA6063</td></tr><tr><td>Billet temperature [°C]</td><td>500</td></tr><tr><td>Number of elements</td><td>100000</td></tr><tr><td>Ram speed [mm/s]</td><td>0.5</td></tr><tr><td>Share friction coefficient m</td><td>1.0</td></tr><tr><td>Heat transfer coefficient [N·mm/(s·°C)]</td><td>11</td></tr><tr><td>Die structure</td><td>Solid die<br>Pseudo porthole die</td></tr></table>

## Experimental Results

Fig. 4 shows the solidified compacts compacted at room temperature (25 °C) and 15 t (density: 1.640 g/cm³), at 100 °C and 50 t (density: 2.151 g/cm³), at 100 °C and 100 t (density: 2.466 g/cm³), and at 200 °C and 100 t (density: 2.607 g/cm³). The solidified compacts were compacted at 100 °C and 100 t (density: 2.466 g/cm³) and at 200 °C and 100 t (density: 2.607 g/cm³). The density of Al (2.7 g/cm³) approached that of Al at higher compressive temperatures and forces.

<table><caption>Fig. 4 Solidified compacts at each compression condition</caption><tr><td>Compression conditions</td><td>Room Temperature (R.T.) 15 t</td><td>100 °C 50 t</td><td>100 °C 100 t</td><td>200 °C 100 t</td></tr><tr><td>Appearance</td><td></td><td></td><td></td><td></td></tr><tr><td>Density [g/cm³]</td><td>1.640</td><td>2.151</td><td>2.466</td><td>2.607</td></tr></table>

Hot extrusion experiments were conducted using AA6063 chip billets and chip billets (density: 2.607 g/cm³) to determine the extrudability of the solid die and pseudo porthole die. The extrudability was determined by examining the extrusion force and appearance of the extruded product. In the extrudate appearance investigation, we determined the presence or absence of defects on the surface of the chip material and the size of the defects.

Fig. 5 shows the extrusion force–ram stroke curve for the hot extrusion of a chip billet (density: 2.607 g/cm³) and a solid billet. The maximum extrusion force was 94.7 t using the pseudo porthole die, which is approximately 1.5 times that of the solid die. In the case of extrusion using the pseudo porthole die, the extrusion force increased in two stages, and it is considered that the billet passed through the port section at the first increase in force and passed through the forming section after the second increase in force. No large force difference was observed between the pseudo porthole die and the solid die for the chip and solid billets.
<figure><img src="image_635.png" /><figcaption>Fig. 5 Extrusion force–ram stroke curve
(Extrusion temperature: 500 °C, ram speed: 0.5 mm/s)</figcaption></figure>

Fig. 6 shows the appearance of the products extruded from the chip billets (density: 2.607 g/cm³) using a solid die and a pseudo porthole die. It was confirmed that the material extruded for the solid die had defects on its surface, which were similar in shape to voids, whereas no defects were observed for the pseudo porthole die. For the solid die, the bonding strength of each chip was weak, and large defects were generated on the surface because the poorly bonded parts were extruded as they were. However, the complex plastic flow (flow separation and crimping) for the pseudo porthole die increased the adhesiveness of the chip materials, and the defects caused by poor adhesiveness were suppressed.

<figure><img src="image_636.png" /><figcaption>Fig. 6 Appearance of extruded product using chip material (density: 2.607 g/cm3) in pseudo porthole die and solid die</figcaption></figure>

Fig. 7 shows the equivalent stresses obtained from the extrusion analysis using a solid die and a pseudo porthole die. The solid die exhibited high stress only near the die-forming area, similar to the study by Tekkaya et al., whereas the pseudo porthole die exhibited high strain throughout the material from the port to the chamber area, confirming the effect of complex plastic flow [13, 14]. In the case of a pseudo porthole die, strong shear deformation can be applied over a wide range inside the die structure, which increases the bond strength of each chip and prevents the extruded material surface from forming, as shown in Fig. 6.
<figure><img src="image_646.png" /><figcaption>Fig. 5 Extrusion force–ram stroke curve
(Extrusion temperature: 500 °C, ram speed: 0.5 mm/s)</figcaption></figure>

Fig. 6 shows the appearance of the products extruded from the chip billets (density: 2.607 g/cm³) using a solid die and a pseudo porthole die. It was confirmed that the material extruded for the solid die had defects on its surface, which were similar in shape to voids, whereas no defects were observed for the pseudo porthole die. For the solid die, the bonding strength of each chip was weak, and large defects were generated on the surface because the poorly bonded parts were extruded as they were. However, the complex plastic flow (flow separation and crimping) for the pseudo porthole die increased the adhesiveness of the chip materials, and the defects caused by poor adhesiveness were suppressed.

<figure><img src="image_647.png" /><figcaption>Fig. 6 Appearance of extruded product using chip material (density: 2.607 g/cm3) in pseudo porthole die and solid die</figcaption></figure>

Fig. 7 shows the equivalent stresses obtained from the extrusion analysis using a solid die and a pseudo porthole die. The solid die exhibited high stress only near the die-forming area, similar to the study by Tekkaya et al., whereas the pseudo porthole die exhibited high strain throughout the material from the port to the chamber area, confirming the effect of complex plastic flow [13, 14]. In the case of a pseudo porthole die, strong shear deformation can be applied over a wide range inside the die structure, which increases the bond strength of each chip and prevents the extruded material surface from forming, as shown in Fig. 6.
<figure><img src="image_637.png" /><figcaption>Fig. 7 Equivalent stresses inside a solid die and a pseudo porthole die</figcaption></figure>

Fig. 8 shows the extrusion force–ram stroke curve for the chip billet shown in Fig. 4 using a pseudo porthole die. The maximum extrusion force was 93.9 t for chip billets (density: 1.640 g/cm³), 79.5 t for chip billets (density: 2.151 g/cm³), 78.7 t for chip billets (density: 2.466 g/cm³), and 92.8 t for chip billets (density: 2.607 g/cm³). The maximum extrusion force was 78.7 t for the billet length. Although the maximum extrusion force is significantly affected by the length of the billet, it is approximately the same and all the extrusion force curves had a similar shape.

Fig. 9 shows the appearance of the extruded product for each chip billet. Chip billets (density: 2.466 g/cm³) and chip billets (density: 2.607 g/cm³) had no defects on the product surface, resulting in a beautiful, defect-free product. Chip billets (density: 2.151 g/cm³) had surface defects in the first half of the extrusion (Head), and chip billets (density: 1.640 g/cm³) had surface defects from the first half (Head) to the middle of the extrusion (Middle). This indicates that as the density of the chip billet approached that of Al (density: 2.7 g/cm³), no surface defects were observed on the product. In addition, for all billet densities, no product surface defects were observed in the latter half (end) of the extrusion process. For low-density chip billets, large product surface defects occurred in the first half of the extrusion (Head), and the size of the defects decreased in the middle of the extrusion (Middle). In the case of extrusion with a chip billet, the amount of strain due to shear deformation in the container and inside the pseudo porthole die increased as the ram stroke progressed, which increased the adhesion strength of the chip during forming, decreased the size of the product surface defects, and suppressed the defects toward the end of the extrusion process.

<figure><img src="image_638.png" /><figcaption>Fig. 8 Extrusion force–ram stroke curve in each chip billet
(Extrusion temperature: 500 °C, ram speed: 0.5 mm/s, die structure: pseudo porthole die)</figcaption></figure>
<figure><img src="image_648.png" /><figcaption>Fig. 7 Equivalent stresses inside a solid die and a pseudo porthole die</figcaption></figure>

Fig. 8 shows the extrusion force–ram stroke curve for the chip billet shown in Fig. 4 using a pseudo porthole die. The maximum extrusion force was 93.9 t for chip billets (density: 1.640 g/cm³), 79.5 t for chip billets (density: 2.151 g/cm³), 78.7 t for chip billets (density: 2.466 g/cm³), and 92.8 t for chip billets (density: 2.607 g/cm³). The maximum extrusion force was 78.7 t for the billet length. Although the maximum extrusion force is significantly affected by the length of the billet, it is approximately the same and all the extrusion force curves had a similar shape.

Fig. 9 shows the appearance of the extruded product for each chip billet. Chip billets (density: 2.466 g/cm³) and chip billets (density: 2.607 g/cm³) had no defects on the product surface, resulting in a beautiful, defect-free product. Chip billets (density: 2.151 g/cm³) had surface defects in the first half of the extrusion (Head), and chip billets (density: 1.640 g/cm³) had surface defects from the first half (Head) to the middle of the extrusion (Middle). This indicates that as the density of the chip billet approached that of Al (density: 2.7 g/cm³), no surface defects were observed on the product. In addition, for all billet densities, no product surface defects were observed in the latter half (end) of the extrusion process. For low-density chip billets, large product surface defects occurred in the first half of the extrusion (Head), and the size of the defects decreased in the middle of the extrusion (Middle). In the case of extrusion with a chip billet, the amount of strain due to shear deformation in the container and inside the pseudo porthole die increased as the ram stroke progressed, which increased the adhesion strength of the chip during forming, decreased the size of the product surface defects, and suppressed the defects toward the end of the extrusion process.

<figure><img src="image_649.png" /><figcaption>Fig. 8 Extrusion force–ram stroke curve in each chip billet
(Extrusion temperature: 500 °C, ram speed: 0.5 mm/s, die structure: pseudo porthole die)</figcaption></figure>
<figure><img src="image_639.png" /><figcaption>Fig. 9 Appearance of extruded product with chip billets of various densities</figcaption></figure>

## Conclusions

The following conclusions are drawn based on the hot extrusion, extrusion simulation, and surface observation of AA6063 chip material using a solid die and a pseudo porthole die.

1) Chip materials can be solidified to a density close to that of AA6063 by pressing at 200 °C and 100 t.

2) In the case of extrusion using a pseudo porthole die, the extrusion force was approximately 1.5 times higher than that using a solid die.

3) A higher strain was generated in the material using the pseudo porthole die than in the material using the solid die.

4) In the hot extrusion process of chip billets, the use of a pseudo porthole die suppressed the defects on the surface of the product, which occurs when a solid die is used.

5) When the density of chip billets approached that of AA6063, extruded materials without surface defects were obtained.

6) For a low density chip billets, product surface defects occurred until the middle stage of the extrusion; however, no product surface defects occurred in the latter half of the extrusion.

Future research will focus on evaluating the mechanical properties of the extruded chip billets for each density and internal porosity defects such as blister tests. In addition, the port shape of the pseudo porthole die will be varied to determine the relationship between the amount of shear strain applied to the interior and the billet density.

## References

[1] F. Czerwinski, Current trends in automotive lightweighting strategies and materials. Materials, 14(21), (2021), 6631. https://doi.org/10.3390/ma14216631

[2] R. Modaresi, D. B. Müller, The role of automobiles for the future of aluminum recycling, Environ. Sci. Technol., 46(16), (2012), 8587-8594. https://doi.org/10.1021/es300648w

[3] K. Martinsen, S. Gulbrandsen-Dahl, Use of post-consumer scrap in aluminium wrought alloy structural components for the transportation sector, Procedia CIRP, 29, (2015), 686-691. https://doi.org/10.1016/j.procir.2015.02.072