# Manufacturing of shape memory foams in hypergravity and simulated microgravity

SANTO Loredana<fnref n="1" />,ª*, IORIO Leandro<fnref n="1" />,ᵇ*, PROIETTI Alice<fnref n="1" />,ᶜ,
BELLISARIO Denise<fnref n="2" />,ᵈ, QUADRINI Fabrizio<fnref n="1" />,ᵉ, IOPPOLO Pietro<fnref n="3" />,ᵉ,
PECCI Raffaella<fnref n="3" />,⁹

**Keywords:** Shape Memory Foams, Hypergravity, Suborbital Flights, Epoxy Foams

**Abstract.** Shape memory polymer (SMP) foams have the potential to be used in space application for self-deployable structures and actuators. For this reason, shape recovery tests have been performed in microgravity, but these foams have been never manufactured in space, because of experimentation costs and flight opportunities. Some foams have been manufactured on-Earth in a random positioning machine (RPM) and a large diameter centrifuge (LDC) under different conditions with the aim of understating if the gravity can play a role during foaming. Foams have been characterized by 3D micro-Computed Tomography (micro-CT) and the effect of the foaming condition on the foam structure is discussed in this study. For the first time, it has been observed that SMP epoxy foams produced by solid state foaming present open porosity with a negligible contribution of closed pores. Moreover, it is confirmed that SSF foams is minimally affected by the gravity conditions apart the contribution of the gravity vector on the heat transfer conditions during foaming.

**Introduction**

Microgravity experiments allow studying physical mechanisms at their basis and preparing future human life in space environment. Space structures have always been manufactured on earth but, in the future, they could be produced in orbit. For this aim, some research studies have evaluated the possibility to move such technologies in space, such as friction stir welding [1] and additive manufacturing [2]. On this line, the current study wants to put the basis for further exploitation of foaming technologies in the space environment. In the past, polyurethane foams with near perfect structure have been produced under microgravity with the Consort series research rockets [3], and other experiments about foams were made in the context of the ESA FOAM project [4,5]. Metal foaming was also replicated in parabolic flights [6]. Nevertheless, epoxy resin foams have never been investigated in microgravity.

Research and technological innovation on epoxy foams are continuous, from extrusion foaming [7-8] to microcellular foaming [9]. Epoxy foams also have additional functional properties as shape memory polymers (SMPs). SMPs have the ability to restore an equilibrium shape after freezing a non-equilibrium configuration. Typically, the non-equilibrium shape is stored in the end of a thermo-mechanical cycle.

<figure><img src="image_2303.png" /></figure>

<footnote n="1">¹University of Rome Tor Vergata, via del Politecnico 1, Rome, Italy</footnote>

<footnote n="2">²Mercatorum University, p.zza Mattei 10, Rome, Italy</footnote>

<footnote n="3">³Italian National Institute of Health, Viale Regina Elena 299, Rome, Italy</footnote>

<footnote>ᵃloredana.santo@uniroma2.it, ᵇleandro.iorio@uniroma2.it, ᶜalice.proietti@uniroma2.it,</footnote>

<footnote>ᵈdenise.bellisario@unimercatorum.it, ᶜfabrizio.quadrini@uniroma2.it,</footnote>

<footnote>ᵉpietro.ioppolo@iss.it, ᵈraffaella.pecci@iss.it</footnote>
# Manufacturing of shape memory foams in hypergravity and simulated microgravity

SANTO Loredana<fnref n="1" />,ª*, IORIO Leandro<fnref n="1" />,ᵇ*, PROIETTI Alice<fnref n="1" />,ᶜ,
BELLISARIO Denise<fnref n="2" />,ᵈ, QUADRINI Fabrizio<fnref n="1" />,ᵉ, IOPPOLO Pietro<fnref n="3" />,ᶜ,
PECCI Raffaella<fnref n="3" />,⁹

**Keywords:** Shape Memory Foams, Hypergravity, Suborbital Flights, Epoxy Foams

**Abstract.** Shape memory polymer (SMP) foams have the potential to be used in space application for self-deployable structures and actuators. For this reason, shape recovery tests have been performed in microgravity, but these foams have been never manufactured in space, because of experimentation costs and flight opportunities. Some foams have been manufactured on-Earth in a random positioning machine (RPM) and a large diameter centrifuge (LDC) under different conditions with the aim of understating if the gravity can play a role during foaming. Foams have been characterized by 3D micro-Computed Tomography (micro-CT) and the effect of the foaming condition on the foam structure is discussed in this study. For the first time, it has been observed that SMP epoxy foams produced by solid state foaming present open porosity with a negligible contribution of closed pores. Moreover, it is confirmed that SSF foams is minimally affected by the gravity conditions apart the contribution of the gravity vector on the heat transfer conditions during foaming.

**Introduction**

Microgravity experiments allow studying physical mechanisms at their basis and preparing future human life in space environment. Space structures have always been manufactured on earth but, in the future, they could be produced in orbit. For this aim, some research studies have evaluated the possibility to move such technologies in space, such as friction stir welding [1] and additive manufacturing [2]. On this line, the current study wants to put the basis for further exploitation of foaming technologies in the space environment. In the past, polyurethane foams with near perfect structure have been produced under microgravity with the Consort series research rockets [3], and other experiments about foams were made in the context of the ESA FOAM project [4,5]. Metal foaming was also replicated in parabolic flights [6]. Nevertheless, epoxy resin foams have never been investigated in microgravity.

Research and technological innovation on epoxy foams are continuous, from extrusion foaming [7-8] to microcellular foaming [9]. Epoxy foams also have additional functional properties as shape memory polymers (SMPs). SMPs have the ability to restore an equilibrium shape after freezing a non-equilibrium configuration. Typically, the non-equilibrium shape is stored in the end of a thermo-mechanical cycle.

<figure><img src="image_2294.png" /></figure>

<footnote n="1">¹University of Rome Tor Vergata, via del Politecnico 1, Rome, Italy</footnote>

<footnote n="2">²Mercatorum University, p.zza Mattei 10, Rome, Italy</footnote>

<footnote n="3">³Italian National Institute of Health, Viale Regina Elena 299, Rome, Italy</footnote>

<footnote>ᵃloredana.santo@uniroma2.it, ᵇleandro.iorio@uniroma2.it, ᶜalice.proietti@uniroma2.it,</footnote>

<footnote>ᵈdenise.bellisario@unimercatorum.it, ᶜfabrizio.quadrini@uniroma2.it,</footnote>

<footnote>fpietro.ioppolo@iss.it, ᵈraffaella.pecci@iss.it</footnote>
<figure><img src="image_2304.png" /><figcaption>Figure 1. RPM (a) and LDC (b) machines at ESTEC-ESA.</figcaption></figure>

The initial shape is recovered by heating in absence of constraints. In the “I-FOAM” experiment, on the International Space Station (ISS), shape recovery of epoxy foams was achieved in microgravity [10] with 3 different configurations [11]. Actually, microgravity does not modify foam recovery, but it affects sample heating [12]. In the following experiment, “FOAM2”, the same SMPC epoxy foams were used [13], and higher recovery ratios were achieved [14]. I-FOAM and FOAM2 experiments showed the feasibility of technical applications of SMP epoxy foams in space, but those foams were manufactured on Earth. In particular, they were foamed by means of an innovative process, called “solid-state foaming” (SSF) without the use of any agent [15-16]. In order to move the production phase in orbit, the first aspect to consider is the effect of gravity on the foaming process. Theoretically, SSF could be performed in the experimental facilities on the ISS, but costs are very high. Therefore, some tests should be carried out on Earth to understand the possible role of gravity on SSF. Unfortunately, the microgravity time of parabolic flights and drop towers are too low. An alternative is using hyper-gravity or random positioning (also known as “simulated microgravity”). In fact, if no effect is measured under hyper-gravity or simulated microgravity, it is possible to infer that gravity is not able to alter the foaming process at all. For this aim, the European Space Research & Technology Centre of the European Space Agency (ESTEC-ESA) gives the possibility to use a random positioning machine (RPM) and a large diameter centrifuge (LDC), shown in Figure 1 a) and b) respectively. The RPM consists of two independent rotating frames, one inside the other [17]. Samples are fixed to the inner frame; therefore, the gravity vector is averaged in all the positions. Most of the experiments were carried out with biological samples [18]. The LDC consists of a 4-arm machine where several experiments can be accommodated [19].

In 2018, SSF of epoxy resins was performed by placing the foaming chamber into the RPM and the LDC [20]. More than 50 foams were manufactured in different conditions. It was found that the gravity vector interacts with the flow front and affects the contact between the foam precursor and the oven. However, the effect of those gravity conditions on the structure of the foamed samples have never been discussed. In this work, for the first time, the cell structure of SMP epoxy foams produced under hyper-gravity and simulated microgravity is discussed. 3D micro-Computed Tomography has been used for this aim; micro-CT is a nondestructive technique that generates cross-sectional images of a sample using an X-ray source. Using these 2D representations, it is possible to generate a 3D reconstruction of the sample and to perform analysis on the material internal structure.

## Materials and Methods

A small foaming chamber was designed to be mounted on the RPM and into a cabinet of a LDC. Foaming tests were carried out by using the same epoxy resin (3M ScotchkoteTM 206 N) of
<figure><img src="image_2295.png" /><figcaption>Figure 1. RPM (a) and LDC (b) machines at ESTEC-ESA.</figcaption></figure>

The initial shape is recovered by heating in absence of constraints. In the “I-FOAM” experiment, on the International Space Station (ISS), shape recovery of epoxy foams was achieved in microgravity [10] with 3 different configurations [11]. Actually, microgravity does not modify foam recovery, but it affects sample heating [12]. In the following experiment, “FOAM2”, the same SMPC epoxy foams were used [13], and higher recovery ratios were achieved [14]. I-FOAM and FOAM2 experiments showed the feasibility of technical applications of SMP epoxy foams in space, but those foams were manufactured on Earth. In particular, they were foamed by means of an innovative process, called “solid-state foaming” (SSF) without the use of any agent [15-16]. In order to move the production phase in orbit, the first aspect to consider is the effect of gravity on the foaming process. Theoretically, SSF could be performed in the experimental facilities on the ISS, but costs are very high. Therefore, some tests should be carried out on Earth to understand the possible role of gravity on SSF. Unfortunately, the microgravity time of parabolic flights and drop towers are too low. An alternative is using hyper-gravity or random positioning (also known as “simulated microgravity”). In fact, if no effect is measured under hyper-gravity or simulated microgravity, it is possible to infer that gravity is not able to alter the foaming process at all. For this aim, the European Space Research & Technology Centre of the European Space Agency (ESTEC-ESA) gives the possibility to use a random positioning machine (RPM) and a large diameter centrifuge (LDC), shown in Figure 1 a) and b) respectively. The RPM consists of two independent rotating frames, one inside the other [17]. Samples are fixed to the inner frame; therefore, the gravity vector is averaged in all the positions. Most of the experiments were carried out with biological samples [18]. The LDC consists of a 4-arm machine where several experiments can be accommodated [19].

In 2018, SSF of epoxy resins was performed by placing the foaming chamber into the RPM and the LDC [20]. More than 50 foams were manufactured in different conditions. It was found that the gravity vector interacts with the flow front and affects the contact between the foam precursor and the oven. However, the effect of those gravity conditions on the structure of the foamed samples have never been discussed. In this work, for the first time, the cell structure of SMP epoxy foams produced under hyper-gravity and simulated microgravity is discussed. 3D micro-Computed Tomography has been used for this aim; micro-CT is a nondestructive technique that generates cross-sectional images of a sample using an X-ray source. Using these 2D representations, it is possible to generate a 3D reconstruction of the sample and to perform analysis on the material internal structure.

## Materials and Methods

A small foaming chamber was designed to be mounted on the RPM and into a cabinet of a LDC. Foaming tests were carried out by using the same epoxy resin (3M ScotchkoteTM 206 N) of
previous experiments in Space [10,12,13,14], and on Earth [11,15,16]. This resin had already shown remarkable shape memory properties [10-16].

<figure><img src="image_2305.png" /><figcaption>Figure 2. SSF chamber (a) and foamed sample (b) with its precursor.</figcaption></figure>

<figure><img src="image_2306.png" /><figcaption>Figure 3. Extraction of a sample for the computer micro-tomography (a), typical cross-section (b) and longitudinal image (c).</figcaption></figure>

The SSF chamber is shown in Figure 2 together with a foamed sample and its tablet precursor. The foaming chamber consisted of a steel cylinder and a 470 W linear heater wound on it. An external frame protected the heater. The foaming chamber had an internal diameter of 15 mm and allowed the insertion of another tube (made of copper, 12 mm in diameter, 105 mm long) which acted as sample holder. The maximum allowable temperature for foaming was 400 °C.

In foaming tests, a tablet precursor was put in the sample holder which was put in turn into the foaming chamber. A thin metallic Al sheet was placed in the inner wall of the sample holder to avoid foam sticking. Precursors were made by cold compaction of the uncured epoxy powder. The tablets were cylindrical with diameter and height of 10 mm. Over 50 tablets were compressed and foamed, 10 of which have been characterized for the current study: 4 foams were fabricated in the RPM, 5 foams in the LDC, and 1 foam in static conditions for comparison.

Tablet foaming was carried out at the temperature of 320 °C for a minimum time of 8 min. In the case of the RPM, the random mode was used, where rotational speeds and angular displacement are randomly changed during time. The maximum value for the rotational speed of both axes was fixed to 30, 60, 89, and 120 deg/s. In the case of the LDC, the centrifugal acceleration, which was perpendicular to the gravity vector, and was fixed to 4 g, 10 g, and 20 g. For 10 g and 20 g, the
previous experiments in Space [10,12,13,14], and on Earth [11,15,16]. This resin had already shown remarkable shape memory properties [10-16].

<figure><img src="image_2296.png" /><figcaption>Figure 2. SSF chamber (a) and foamed sample (b) with its precursor.</figcaption></figure>

<figure><img src="image_2297.png" /><figcaption>Figure 3. Extraction of a sample for the computer micro-tomography (a), typical cross-section (b) and longitudinal image (c).</figcaption></figure>

The SSF chamber is shown in Figure 2 together with a foamed sample and its tablet precursor. The foaming chamber consisted of a steel cylinder and a 470 W linear heater wound on it. An external frame protected the heater. The foaming chamber had an internal diameter of 15 mm and allowed the insertion of another tube (made of copper, 12 mm in diameter, 105 mm long) which acted as sample holder. The maximum allowable temperature for foaming was 400 °C.

In foaming tests, a tablet precursor was put in the sample holder which was put in turn into the foaming chamber. A thin metallic Al sheet was placed in the inner wall of the sample holder to avoid foam sticking. Precursors were made by cold compaction of the uncured epoxy powder. The tablets were cylindrical with diameter and height of 10 mm. Over 50 tablets were compressed and foamed, 10 of which have been characterized for the current study: 4 foams were fabricated in the RPM, 5 foams in the LDC, and 1 foam in static conditions for comparison.

Tablet foaming was carried out at the temperature of 320 °C for a minimum time of 8 min. In the case of the RPM, the random mode was used, where rotational speeds and angular displacement are randomly changed during time. The maximum value for the rotational speed of both axes was fixed to 30, 60, 89, and 120 deg/s. In the case of the LDC, the centrifugal acceleration, which was perpendicular to the gravity vector, and was fixed to 4 g, 10 g, and 20 g. For 10 g and 20 g, the
foaming chamber was mounted in 2 different ways, with the foaming direction according or opposite to the centrifugal acceleration.

After foaming, the foams were cut to have a final height of 15 mm, according to Figure 3, and samples were characterized by 3D microcomputed tomography (SkyScan 1072; Bruker micro-CT, Kontich, Belgium). The acquisition stage was carried out setting a pixel size of 14.65 x 14.65 µm, which corresponds to 20X magnification with 1.3 s exposure time and 0.45° rotation step (for a total rotation of 180°) applying a voltage of 100 kV with a current of 98 µA. The obtained acquisition data were subsequently elaborated and reconstructed (NRecon V1.7.0; Bruker microCT, Kontich, Belgium), providing axial cross sections with a pixel size of 14.65 x 14.65 µm for the total height of each foamed sample. The distance between each cross section was 14 µm.

The foamed samples were then analyzed by means of CT-Analyser software (Version 1.16.9.0; Bruker micro-CT, Kontich, Belgium) to calculate the histomorphometric parameters, in particular the Structure thickness, the Object Surface/Volume ratio (Surface/volume ratio), the Total porosity (%) and the Open and Closed Porosity (%) in 10 sections of each sample at a vertical distance of 1 mm. For the evaluation of these features, the grey level images were binarized implementing a global threshold method which is based on only one parameter (i.e., the grey level value) to be set by the user.

## Results and discussion

The average bulk density of all the foam was 0.38±0.02 g/cm³, comparable with foams produced in other studies. In static condition, the density of the reference sample approaches the highest value (0.40 g/cm³) which corresponds to the lowest porosity. However, the same value has also been measured in the RPM samples at high acceleration speeds (90 and 120 deg/s) and in the LDC sample at low centrifugal acceleration (4 g). In all the other cases, a density reduction of 10% is found on average. The lowest density has been measured on the 20 g LDC sample with the foaming direction aligned with the centrifugal acceleration vector (0.33 g/cm³). It is important to observe that the foam lateral walls, in contact with the Al liner during foaming, are not porous and may have a variable thickness which can influence the density measurement. For this reason, data from acquired sections of the micro-CT provide a better comparison between the different foaming conditions. Density data are reported in Table 1 together with data extracted from tomographic sections.

<table><caption>Table 1. Measured data on foamed samples.</caption><thead><tr><th>Sample name</th><th>Bulk density [g/cm³]</th><th>Object S/V* ratio [1/mm]</th><th>Structure thickness [mm]</th><th>Closed porosity [%]</th><th>Open porosity [%]</th><th>Total porosity [%]</th></tr></thead><tbody><tr><td>30 RPM</td><td>0.35</td><td>38.8</td><td>0.101</td><td>0.030</td><td>65.2</td><td>65.2</td></tr><tr><td>60 RPM</td><td>0.38</td><td>38.9</td><td>0.100</td><td>0.027</td><td>64.4</td><td>64.4</td></tr><tr><td>90 RPM</td><td>0.40</td><td>38.6</td><td>0.103</td><td>0.038</td><td>60.3</td><td>60.3</td></tr><tr><td>120 RPM</td><td>0.40</td><td>39.0</td><td>0.102</td><td>0.034</td><td>62.2</td><td>62.2</td></tr><tr><td>TEST</td><td>0.40</td><td>44.5</td><td>0.090</td><td>0.023</td><td>69.5</td><td>69.5</td></tr><tr><td>4g LDC</td><td>0.40</td><td>49.8</td><td>0.083</td><td>0.020</td><td>72.2</td><td>72.2</td></tr><tr><td>10g LDC</td><td>0.37</td><td>42.4</td><td>0.093</td><td>0.023</td><td>66.0</td><td>66.0</td></tr><tr><td>20g LDC</td><td>0.38</td><td>40.6</td><td>0.099</td><td>0.035</td><td>61.0</td><td>61.0</td></tr><tr><td>10g LDC (aligned)</td><td>0.37</td><td>37.5</td><td>0.102</td><td>0.031</td><td>65.0</td><td>65.0</td></tr><tr><td>20g LDC (aligned)</td><td>0.34</td><td>48.9</td><td>0.086</td><td>0.011</td><td>76.3</td><td>76.3</td></tr></tbody></table>

<footnote>*Surface/Volume ratio</footnote>
foaming chamber was mounted in 2 different ways, with the foaming direction according or opposite to the centrifugal acceleration.

After foaming, the foams were cut to have a final height of 15 mm, according to Figure 3, and samples were characterized by 3D microcomputed tomography (SkyScan 1072; Bruker micro-CT, Kontich, Belgium). The acquisition stage was carried out setting a pixel size of 14.65 x 14.65 µm, which corresponds to 20X magnification with 1.3 s exposure time and 0.45° rotation step (for a total rotation of 180°) applying a voltage of 100 kV with a current of 98 µA. The obtained acquisition data were subsequently elaborated and reconstructed (NRecon V1.7.0; Bruker microCT, Kontich, Belgium), providing axial cross sections with a pixel size of 14.65 x 14.65 µm for the total height of each foamed sample. The distance between each cross section was 14 µm.

The foamed samples were then analyzed by means of CT-Analyser software (Version 1.16.9.0; Bruker micro-CT, Kontich, Belgium) to calculate the histomorphometric parameters, in particular the Structure thickness, the Object Surface/Volume ratio (Surface/volume ratio), the Total porosity (%) and the Open and Closed Porosity (%) in 10 sections of each sample at a vertical distance of 1 mm. For the evaluation of these features, the grey level images were binarized implementing a global threshold method which is based on only one parameter (i.e., the grey level value) to be set by the user.

## Results and discussion

The average bulk density of all the foam was 0.38±0.02 g/cm³, comparable with foams produced in other studies. In static condition, the density of the reference sample approaches the highest value (0.40 g/cm³) which corresponds to the lowest porosity. However, the same value has also been measured in the RPM samples at high acceleration speeds (90 and 120 deg/s) and in the LDC sample at low centrifugal acceleration (4 g). In all the other cases, a density reduction of 10% is found on average. The lowest density has been measured on the 20 g LDC sample with the foaming direction aligned with the centrifugal acceleration vector (0.33 g/cm³). It is important to observe that the foam lateral walls, in contact with the Al liner during foaming, are not porous and may have a variable thickness which can influence the density measurement. For this reason, data from acquired sections of the micro-CT provide a better comparison between the different foaming conditions. Density data are reported in Table 1 together with data extracted from tomographic sections.

<table><caption>Table 1. Measured data on foamed samples.</caption><thead><tr><th>Sample name</th><th>Bulk density [g/cm³]</th><th>Object S/V* ratio [1/mm]</th><th>Structure thickness [mm]</th><th>Closed porosity [%]</th><th>Open porosity [%]</th><th>Total porosity [%]</th></tr></thead><tbody><tr><td>30 RPM</td><td>0.35</td><td>38.8</td><td>0.101</td><td>0.030</td><td>65.2</td><td>65.2</td></tr><tr><td>60 RPM</td><td>0.38</td><td>38.9</td><td>0.100</td><td>0.027</td><td>64.4</td><td>64.4</td></tr><tr><td>90 RPM</td><td>0.40</td><td>38.6</td><td>0.103</td><td>0.038</td><td>60.3</td><td>60.3</td></tr><tr><td>120 RPM</td><td>0.40</td><td>39.0</td><td>0.102</td><td>0.034</td><td>62.2</td><td>62.2</td></tr><tr><td>TEST</td><td>0.40</td><td>44.5</td><td>0.090</td><td>0.023</td><td>69.5</td><td>69.5</td></tr><tr><td>4g LDC</td><td>0.40</td><td>49.8</td><td>0.083</td><td>0.020</td><td>72.2</td><td>72.2</td></tr><tr><td>10g LDC</td><td>0.37</td><td>42.4</td><td>0.093</td><td>0.023</td><td>66.0</td><td>66.0</td></tr><tr><td>20g LDC</td><td>0.38</td><td>40.6</td><td>0.099</td><td>0.035</td><td>61.0</td><td>61.0</td></tr><tr><td>10g LDC (aligned)</td><td>0.37</td><td>37.5</td><td>0.102</td><td>0.031</td><td>65.0</td><td>65.0</td></tr><tr><td>20g LDC (aligned)</td><td>0.34</td><td>48.9</td><td>0.086</td><td>0.011</td><td>76.3</td><td>76.3</td></tr></tbody></table>

<footnote>*Surface/Volume ratio</footnote>
The main result is that all the foams, comprising the reference one, show a structure with open pores. In fact, open porosity is coincident with the total porosity (66% on average) for all the sample and the closed porosity is always lower than 0.04%. This is the first time that this aspect is observed. This finding is very important also for understanding the mechanism of the shape memory behavior of this kind of epoxy foams. In terms of comparison between the different foaming conditions, data seem to confirm the general conclusions of the previous study [21]. The foaming process is dominated by the heat transfer and results depend on the contact between the sample holder and the foaming chamber, as well as between the Al liner and the sample holder. Therefore, some effects can be discussed but strong trends are not observed. In RPM machines, because of the random orientation, similar Surface/Volume (S/V) ratios of the cavities have been observed, lower than the reference and the LDC, where the acceleration vector is constantly oriented in one direction. The thickness of the cavity walls is comparable in all the foams, being slightly higher in RPM samples. The highest porosity is measured on the same LDC sample where the lowest density was measured, thus confirming the positive effect of the alignment between the foaming direction and the centrifugal force. However, this positive effect is visible clearly only at very high g values. Data of bulk density and porosity of Table 1 are also reported in Figure 4. These values seem to be very poorly correlated because of the discussed contribution of the sample cylindrical walls.

<figure><img src="image_2307.png" /><figcaption>Figure 4. Density and porosity data of foamed samples.</figcaption></figure>

## Conclusions

Results show that gravity minimally affects the SSF process, apart from its contribution on the heat transfer condition. Because of the resin viscosity during foaming, and the effect of the mold wall on resin sticking, the structure of the foam is not very influenced by the different foaming conditions in RPM and LDC, and some differences may be discussed only at very high g values. The adopted architecture of the foaming chamber is very close to one which could be used in space for foam manufacturing. In that case, the microgravity should not modify the foam structure as well.

For the first time it has been observed that SSF of epoxy resin precursors produces foams with open porosity and a negligible contribution of closed porosity. Nevertheless, additional studies are necessary to understand if gravity could affect shape memory properties of the foam, which depend on other factors such as cell shape and orientation.
The main result is that all the foams, comprising the reference one, show a structure with open pores. In fact, open porosity is coincident with the total porosity (66% on average) for all the sample and the closed porosity is always lower than 0.04%. This is the first time that this aspect is observed. This finding is very important also for understanding the mechanism of the shape memory behavior of this kind of epoxy foams. In terms of comparison between the different foaming conditions, data seem to confirm the general conclusions of the previous study [21]. The foaming process is dominated by the heat transfer and results depend on the contact between the sample holder and the foaming chamber, as well as between the Al liner and the sample holder. Therefore, some effects can be discussed but strong trends are not observed. In RPM machines, because of the random orientation, similar Surface/Volume (S/V) ratios of the cavities have been observed, lower than the reference and the LDC, where the acceleration vector is constantly oriented in one direction. The thickness of the cavity walls is comparable in all the foams, being slightly higher in RPM samples. The highest porosity is measured on the same LDC sample where the lowest density was measured, thus confirming the positive effect of the alignment between the foaming direction and the centrifugal force. However, this positive effect is visible clearly only at very high g values. Data of bulk density and porosity of Table 1 are also reported in Figure 4. These values seem to be very poorly correlated because of the discussed contribution of the sample cylindrical walls.

<figure><img src="image_2298.png" /><figcaption>Figure 4. Density and porosity data of foamed samples.</figcaption></figure>

## Conclusions

Results show that gravity minimally affects the SSF process, apart from its contribution on the heat transfer condition. Because of the resin viscosity during foaming, and the effect of the mold wall on resin sticking, the structure of the foam is not very influenced by the different foaming conditions in RPM and LDC, and some differences may be discussed only at very high g values. The adopted architecture of the foaming chamber is very close to one which could be used in space for foam manufacturing. In that case, the microgravity should not modify the foam structure as well.

For the first time it has been observed that SSF of epoxy resin precursors produces foams with open porosity and a negligible contribution of closed porosity. Nevertheless, additional studies are necessary to understand if gravity could affect shape memory properties of the foam, which depend on other factors such as cell shape and orientation.
## References

[1] W.R. Longhurst, C.D. Cox, B.T. Gibson, G.E Cook, A.M. Strauss, I.C. Wilbur, B.E. Osborne, Development of friction stir welding technologies for in-space manufacturing, Int. J. Adv. Manuf. Technol. 90 (2017) 81-91. https://doi.org/10.1007/s00170-016-9362-1

[2] T. Ghidini, L. Pambaguian, S. Blair, Joining the Third Industrial Revolution: 3D Printing for Space, European Space Agency Bulletin 2015, 163, 24.

[3] S.P. McManus, F.C. Wessling, J.T. Matthews, D.N. Patek, K. Emoto, D.T. Howard, T.S. Price, Production of polyurethane foams in space: Gravitational and vacuum effects on foam formation, ACS Symp. Ser. 793 (2001) 78-96. https://doi.org/10.1021/bk-2001-0793.ch006

[4] N. Vandewalle, H. Caps, G. Delon, A. Saint-Jalmes, E. Rio, L. Saulnier, M. Adler, A.L. Biance, O. Pitois, S.C. Addad, R. Hohler, D. Weaire, S. Hutzler, D. Langevin, Foam stability in microgravity, J. Phys.: Conf. Ser. 327 (2011) no.012024. https://doi.org/10.1051/epn/2014303

[5] D. Langevin, M. Vignes-Adler, Microgravity studies of aqueous wet foams, Eur. Phys. J. E 37 (2014) no.16. https://doi.org/10.1140/epje/i2014-14016-3

[6] F. García-Moreno, S.T. Tobin, M. Mukherjee, C. Jiménez, E. Solórzano, G.S. Vinod Kumar, S. Hutzler, J. Banhart, Analysis of liquid metal foams through X-ray radioscopy and microgravity experiments, Soft Matter 10 (2014) 6955-6962. https://doi.org/10.1039/C4SM00467A

[7] W.N.F. Wan Hamad, P.L. Teh, C.K. Yeoh, Effect of acetic acid as catalyst on the properties of epoxy foam, Polym.-Plas. Technol. Eng. 52 (2013) 754-760. https://doi.org/10.1080/03602559.2012.762375

[8] A. Gandhi, N. Bhatnagar, Physical Blowing Agent Residence Conditions Stimulated Morphological Transformations in Extrusion Foaming, Polym.-Plas. Technol. Eng. 54 (2015) 1812-1818. https://doi.org/10.1080/03602559.2015.1050518

[9] R. Li, D. Zeng, Q. Liu, Z. Jiang, T. Fang, Glass Transition Temperature in Microcellular Foaming Process with Supercritical Carbon Dioxide: A Review, Polym.-Plas. Technol. Eng. 54 (2015) 119-127. https://doi.org/10.1080/03602559.2014.955203

[10] L. Santo, F. Quadrini, E.A. Squeo, F. Dolce, G. Mascetti, D. Bertolotto, W. Villadei, P.L. Ganga, V. Zolesi, Behavior of shape memory epoxy foams in microgravity: experimental results of STS-134 mission, Microgravity Sci. Technol. 24 (2012) 287-296. https://doi.org/10.1007/s12217-012-9313-x

[11] F. Quadrini, L. Santo, E.A. Squeo, Shape memory epoxy foams for space applications, Mater. Lett. 69 (2012) 20-23. https://doi.org/10.1016/j.matlet.2011.11.050

[12] L. Santo, F. Quadrini, G. Mascetti, F. Dolce, V. Zolesi, Mission STS-134: Results of Shape Memory Foam Experiment, Acta Astronaut. 91 (2013) 333-340. https://doi.org/10.1016/j.actaastro.2013.06.017

[13] L. Santo, F. Quadrini, W. Villadei, G. Mascetti, V. Zolesi, Shape Memory Epoxy Foams and Composites: Ribes_FOAM2 Experiment on Spacecraft "BION-M1" and Future Perspective, Procedia Eng. 104 (2015) 50-56. https://doi.org/10.1016/j.proeng.2015.04.096

[14] L. Santo, F. Quadrini, P.L. Ganga, V. Zolesi, Mission BION-M1: results of Ribes/Foam2 experiment on shape memory polymer foams and composites, Aerosp. Sci. Technol 40 (2015) 109-114. https://doi.org/10.1016/j.ast.2014.11.008
## References

[1] W.R. Longhurst, C.D. Cox, B.T. Gibson, G.E Cook, A.M. Strauss, I.C. Wilbur, B.E. Osborne, Development of friction stir welding technologies for in-space manufacturing, Int. J. Adv. Manuf. Technol. 90 (2017) 81-91. https://doi.org/10.1007/s00170-016-9362-1

[2] T. Ghidini, L. Pambaguian, S. Blair, Joining the Third Industrial Revolution: 3D Printing for Space, European Space Agency Bulletin 2015, 163, 24.

[3] S.P. McManus, F.C. Wessling, J.T. Matthews, D.N. Patek, K. Emoto, D.T. Howard, T.S. Price, Production of polyurethane foams in space: Gravitational and vacuum effects on foam formation, ACS Symp. Ser. 793 (2001) 78-96. https://doi.org/10.1021/bk-2001-0793.ch006

[4] N. Vandewalle, H. Caps, G. Delon, A. Saint-Jalmes, E. Rio, L. Saulnier, M. Adler, A.L. Biance, O. Pitois, S.C. Addad, R. Hohler, D. Weaire, S. Hutzler, D. Langevin, Foam stability in microgravity, J. Phys.: Conf. Ser. 327 (2011) no.012024. https://doi.org/10.1051/epn/2014303

[5] D. Langevin, M. Vignes-Adler, Microgravity studies of aqueous wet foams, Eur. Phys. J. E 37 (2014) no.16. https://doi.org/10.1140/epje/i2014-14016-3

[6] F. García-Moreno, S.T. Tobin, M. Mukherjee, C. Jiménez, E. Solórzano, G.S. Vinod Kumar, S. Hutzler, J. Banhart, Analysis of liquid metal foams through X-ray radioscopy and microgravity experiments, Soft Matter 10 (2014) 6955-6962. https://doi.org/10.1039/C4SM00467A

[7] W.N.F. Wan Hamad, P.L. Teh, C.K. Yeoh, Effect of acetic acid as catalyst on the properties of epoxy foam, Polym.-Plas. Technol. Eng. 52 (2013) 754-760. https://doi.org/10.1080/03602559.2012.762375

[8] A. Gandhi, N. Bhatnagar, Physical Blowing Agent Residence Conditions Stimulated Morphological Transformations in Extrusion Foaming, Polym.-Plas. Technol. Eng. 54 (2015) 1812-1818. https://doi.org/10.1080/03602559.2015.1050518

[9] R. Li, D. Zeng, Q. Liu, Z. Jiang, T. Fang, Glass Transition Temperature in Microcellular Foaming Process with Supercritical Carbon Dioxide: A Review, Polym.-Plas. Technol. Eng. 54 (2015) 119-127. https://doi.org/10.1080/03602559.2014.955203

[10] L. Santo, F. Quadrini, E.A. Squeo, F. Dolce, G. Mascetti, D. Bertolotto, W. Villadei, P.L. Ganga, V. Zolesi, Behavior of shape memory epoxy foams in microgravity: experimental results of STS-134 mission, Microgravity Sci. Technol. 24 (2012) 287-296. https://doi.org/10.1007/s12217-012-9313-x

[11] F. Quadrini, L. Santo, E.A. Squeo, Shape memory epoxy foams for space applications, Mater. Lett. 69 (2012) 20-23. https://doi.org/10.1016/j.matlet.2011.11.050

[12] L. Santo, F. Quadrini, G. Mascetti, F. Dolce, V. Zolesi, Mission STS-134: Results of Shape Memory Foam Experiment, Acta Astronaut. 91 (2013) 333-340. https://doi.org/10.1016/j.actaastro.2013.06.017

[13] L. Santo, F. Quadrini, W. Villadei, G. Mascetti, V. Zolesi, Shape Memory Epoxy Foams and Composites: Ribes_FOAM2 Experiment on Spacecraft "BION-M1" and Future Perspective, Procedia Eng. 104 (2015) 50-56. https://doi.org/10.1016/j.proeng.2015.04.096

[14] L. Santo, F. Quadrini, P.L. Ganga, V. Zolesi, Mission BION-M1: results of Ribes/Foam2 experiment on shape memory polymer foams and composites, Aerosp. Sci. Technol 40 (2015) 109-114. https://doi.org/10.1016/j.ast.2014.11.008
[15] E.A. Squeo, F. Quadrini, Shape memory epoxy foams by solid-state foaming, Smart Mater. Struct. 19 (2010) 105002. DOI 10.1088/0964-1726/19/10/105002

[16] F. Quadrini, E.A. Squeo, Solid-state foaming of epoxy resin, J. Cell. Plast. 44 (2008) 161-173. https://doi.org/10.1177/0021955X07082486

[17] J.J.W.A. van Loon, Some history and use of the Random Positioning Machine, RPM, in gravity related research, Adv. Space Res. 39 (2007) 1161-1165. https://doi.org/10.1016/j.asr.2007.02.016

[18] A.G. Borst, J.J.W.A. van Loon, Technology and Developments for the Random Positioning Machine, RPM, Microgravity Sci. Technol. 21 (2009) no.287. https://doi.org/10.1007/s12217-008-9043-2

[19] J.J.W.A. van Loon, J. Krausse, H. Cunha, J. Goncalves, H. Almeida, P. Schiller, The Large Diameter Centrifuge, LDC, for life and physical sciences and technology. In: L. Ouwehand (eds) Life in Space for Life on Earth, ESA-SP Vol. 553. Proceedings of the conference held 22-27 June, 2008 at Angers, France (2008).

[20] F. Quadrini, G.M. Tedde, L. Santo, J.J.W.A. van Loon, Solid-state foaming of epoxy resin under hypergravity and simulated microgravity, Adv. Polym. Technol. 37 (2018) 2316-2624. https://doi.org/10.1002/adv.21937
[15] E.A. Squeo, F. Quadrini, Shape memory epoxy foams by solid-state foaming, Smart Mater. Struct. 19 (2010) 105002. DOI 10.1088/0964-1726/19/10/105002

[16] F. Quadrini, E.A. Squeo, Solid-state foaming of epoxy resin, J. Cell. Plast. 44 (2008) 161-173. https://doi.org/10.1177/0021955X07082486

[17] J.J.W.A. van Loon, Some history and use of the Random Positioning Machine, RPM, in gravity related research, Adv. Space Res. 39 (2007) 1161-1165. https://doi.org/10.1016/j.asr.2007.02.016

[18] A.G. Borst, J.J.W.A. van Loon, Technology and Developments for the Random Positioning Machine, RPM, Microgravity Sci. Technol. 21 (2009) no.287. https://doi.org/10.1007/s12217-008-9043-2

[19] J.J.W.A. van Loon, J. Krausse, H. Cunha, J. Goncalves, H. Almeida, P. Schiller, The Large Diameter Centrifuge, LDC, for life and physical sciences and technology. In: L. Ouwehand (eds) Life in Space for Life on Earth, ESA-SP Vol. 553. Proceedings of the conference held 22-27 June, 2008 at Angers, France (2008).

[20] F. Quadrini, G.M. Tedde, L. Santo, J.J.W.A. van Loon, Solid-state foaming of epoxy resin under hypergravity and simulated microgravity, Adv. Polym. Technol. 37 (2018) 2316-2624. https://doi.org/10.1002/adv.21937
# One-shot drilling of unconventional thin metal hybrid stacks for aerospace applications

PANICO Martina<fnref n="1" />,ª*, D'AGOSTINO Emmanuele<fnref n="2" />,ᵇ, DE ROSA Vincenzo<fnref n="2" />,ᶜ, DURANTE Massimo<fnref n="1" />,ᵈ, MESSERE Serena<fnref n="2" />,ᵉ, LANGELLA Antonio<fnref n="1" />,>f and BOCCARUSSO Luca<fnref n="1" />,⁹

**Keywords:** One-Shot Drilling, Thin Stacked Materials, Drilling Strategy, Burr Height

**Abstract.** The aerospace industry's relentless pursuit of lighter, stronger, and more fuel-efficient aircraft has led to the widespread use of lightweight materials, particularly aluminum alloys, in aerospace engineering. This study focuses on the challenges associated with one-shot drilling without back-up forces of thin parts, a strategy aimed at reducing cycle, avoiding part separation for cleaning after drilling, and minimizing part assembly problems. Despite its effectiveness, problems persist with defects developing at the interface due to the interlayer gap phenomenon. This gap causes metal chips to accumulate and is influenced by both process parameters and the clamping conditions and methodologies. While existing literature primarily associates process parameters with burr measurements, this study delves into the often-overlooked influence of clamping strategies on the occurrence of interface defects. Through experimental drilling on a thin stack of Al7075-T6 and Al2024-T3, various clamping schemes were analyzed to assess their impact on burr height. The results were also correlated with the drill verse, thrust force, and torque. The objective is to increase know-how about these aspects and optimize clamping strategies for effective one-shot drilling, addressing gaps in current research.

**Introduction**

The aerospace industry has continually led advancements in technology and innovation, consistently striving to design aircraft that are not just lighter and more robust but also exhibit enhanced fuel efficiency. A crucial milestone in this effort has been the extensive adoption of lightweight materials, revolutionizing the field of aerospace engineering and opening up new possibilities. This perfectly fits with the worldwide movement towards adopting manufacturing practices that are more environmentally sustainable. Moreover, there is a persistent endeavour to enhance the efficiency of these materials, with a specific focus on minimizing thickness while preserving optimal performance levels.

Aluminum alloys have been a focal point in the ongoing research for weight reduction and improved performance in the aerospace industry. They demonstrate enhanced structural integrity and increased durability. While Carbon Fiber-Reinforced Polymers (CFRPs) offer a remarkable strength-to-weight ratio, they are accompanied by notable limitations, including susceptibility to impact damage and the necessity for intricate repair processes. As a result, aluminium alloys continue to be utilized, not only for their favourable mechanical properties but also due to their cost-effectiveness and well-established manufacturing processes [1,2].

<figure><img src="image_2308.png" /></figure>

<footnote n="1">¹Department of Chemical, Materials and Production Engineering, University of Naples “Federico II” P.le Tecchio 80, 80125, Naples, Italy</footnote>

<footnote n="2">²Leonardo S.p.A, 80038 Pomigliano d'Arco NA, Italy</footnote>

<footnote>ªmartina.panico@unina.it, ᵇemmanuele.dagostino@leonardo.com,</footnote>

<footnote>ᶜvincenzo.derosa01@leonardo.com, ᵈmdurante@unina.it, ᶜserena.messere@leonardo.com,</footnote>

<footnote>f'antgella@unina.it, ⁹luca.boccarusso@unina.it</footnote>
# One-shot drilling of unconventional thin metal hybrid stacks for aerospace applications

PANICO Martina<fnref n="1" />,ª*, D'AGOSTINO Emmanuele<fnref n="2" />,ᵇ, DE ROSA Vincenzo<fnref n="2" />,ᶜ, DURANTE Massimo<fnref n="1" />,ᵈ, MESSERE Serena<fnref n="2" />,ᵉ, LANGELLA Antonio<fnref n="1" />,ᵉ and BOCCARUSSO Luca<fnref n="1" />,ғ

**Keywords:** One-Shot Drilling, Thin Stacked Materials, Drilling Strategy, Burr Height

**Abstract.** The aerospace industry's relentless pursuit of lighter, stronger, and more fuel-efficient aircraft has led to the widespread use of lightweight materials, particularly aluminum alloys, in aerospace engineering. This study focuses on the challenges associated with one-shot drilling without back-up forces of thin parts, a strategy aimed at reducing cycle, avoiding part separation for cleaning after drilling, and minimizing part assembly problems. Despite its effectiveness, problems persist with defects developing at the interface due to the interlayer gap phenomenon. This gap causes metal chips to accumulate and is influenced by both process parameters and the clamping conditions and methodologies. While existing literature primarily associates process parameters with burr measurements, this study delves into the often-overlooked influence of clamping strategies on the occurrence of interface defects. Through experimental drilling on a thin stack of Al7075-T6 and Al2024-T3, various clamping schemes were analyzed to assess their impact on burr height. The results were also correlated with the drill verse, thrust force, and torque. The objective is to increase know-how about these aspects and optimize clamping strategies for effective one-shot drilling, addressing gaps in current research.

**Introduction**

The aerospace industry has continually led advancements in technology and innovation, consistently striving to design aircraft that are not just lighter and more robust but also exhibit enhanced fuel efficiency. A crucial milestone in this effort has been the extensive adoption of lightweight materials, revolutionizing the field of aerospace engineering and opening up new possibilities. This perfectly fits with the worldwide movement towards adopting manufacturing practices that are more environmentally sustainable. Moreover, there is a persistent endeavour to enhance the efficiency of these materials, with a specific focus on minimizing thickness while preserving optimal performance levels.

Aluminum alloys have been a focal point in the ongoing research for weight reduction and improved performance in the aerospace industry. They demonstrate enhanced structural integrity and increased durability. While Carbon Fiber-Reinforced Polymers (CFRPs) offer a remarkable strength-to-weight ratio, they are accompanied by notable limitations, including susceptibility to impact damage and the necessity for intricate repair processes. As a result, aluminium alloys continue to be utilized, not only for their favourable mechanical properties but also due to their cost-effectiveness and well-established manufacturing processes [1,2].

<figure><img src="image_2299.png" /></figure>

<footnote n="1">¹Department of Chemical, Materials and Production Engineering, University of Naples “Federico II” P.le Tecchio 80, 80125, Naples, Italy</footnote>

<footnote n="2">²Leonardo S.p.A, 80038 Pomigliano d'Arco NA, Italy</footnote>

<footnote>ªmartina.panico@unina.it, ᵇemmanuele.dagostino@leonardo.com,</footnote>

<footnote>ᶜvincenzo.derosa01@leonardo.com, ᵈmdurante@unina.it, ᶜserena.messere@leonardo.com,</footnote>

<footnote>f'antgella@unina.it, ᵉluca.boccarusso@unina.it</footnote>
Different types of aluminum alloys find applications in the aerospace sector, with the most prevalent classes mainly derived from the 2xxx and 7xxx series. Notable examples include the 2024-T3 alloy, employed in applications such as fuselage skin, wing skins, and cowls, and the 7075-T6 alloy, used in fuselage bulkheads, wing skins and stringers [3].

Despite one-shot drilling of non-glued parts being an effective drilling strategy for reducing cycle time and avoiding non-coaxiality issues between holes, it still presents several open points when thin sheets are considered. One of the primary issues involves the gradual increase in thrust force during the drilling process, which can lead to the separation of sheets (interlayer gap), resulting in the accumulation of metallic chips between the parts. The interlayer gap occurs when the upper material undergoes elastic spring back, reverting to its initial shape after drilling. Simultaneously, the lower material bends under the influence of the applied thrust force. These phenomena become more prominent and noticeable for thin materials, resulting in a corresponding reduction in their flexural stiffness. Additionally, Luo et al. [4] and Panico et al. [5] highlighted that the interlayer gap influences the formation of burrs. To maintain structural integrity and assembly quality, it is necessary to remove the accumulated material and interface defects through operations such as disassembly, cleaning/deburring, and reassembly. However, these procedures are time-consuming, do not add value, and can be costly [6], nullifying the benefits offered by one-shot drilling. Although process parameters (rotational speed, feed rate, tool geometry) play a key role in studying the interlayer gap phenomenon, equally crucial is the role of the temporary part-fixing conditions adopted during drilling.

To date, scientific literature has primarily focused on the investigation of a direct correlation between process parameters and the measurement of burr [6,7], leaving largely unexplored the contribution that the clamping strategy of parts offers to the development of defects at the interface of two thin materials drilled in a one-shot configuration. In this study, an experimental drilling investigation was conducted on thin stacks composed of Al7075-T6 and Al2024-T3. For a fixed set of process parameters (i.e., rotational speed and feed rate), three different clamping pattern percentages were investigated (40%, 60% and 80%), also changing the drilling verse (i.e., Al2024-T3 or Al7075-T6). During drilling, the thrust force and torque were acquired. The height of the burr was then correlated with the clamping pattern, the drilling side material, and the thrust force and torque.

## Materials and Methods

The experimental one-shot drilling activities were conducted on stacks composed of Al7075-T6 sheet (300 mm x 300 mm x 1 mm) and Al2024-T3 sheet (300 mm x 300 mm x 0.8 mm). These materials were chosen due to their extensive use in the aerospace industry [9].

The stack was investigated in both the Al7075-T6/Al2024-T3 and Al2024-T3/Al7075-T6 configurations (these configurations will be referred as A/B and B/A, respectively) to analyze the influence of the flexural contribution of each material when positioned as the lower reinforcement on the hole quality. In the present experimentation, a coated carbide twist drill with a diameter of 4.76 mm, a cutting-edge height of 1 mm, a point angle of 130° and a helix angle of 25° was used.

The drilling tests were performed in dry condition by using a semi-automated electric drilling unit for industrial applications. The semi-automatic drilling unit is equipped with two independent electric motors, one dedicated to feed motion and the other to rotation (max. power 3 kW and max. rotational speed 8000 rpm). The hole-to-hole handling is performed by the operator, as well as the initiation of the drilling cycle. The drill remains in the drilling position thanks to a collar concentric to the drill bit, which, upon expansion, ensures a tight coupling within the rigid drilling jig (depicted in Fig. 1). The drilling machine is accompanied by specialized software that allows for the management of process parameters and real-time acquisition of current consumption for both motors, thrust force and torque curves over time. The upper and lower edges of each specimen were fixed to the rigid clamping fixture, while the lateral edges were left free to reproduce a typical
Different types of aluminum alloys find applications in the aerospace sector, with the most prevalent classes mainly derived from the 2xxx and 7xxx series. Notable examples include the 2024-T3 alloy, employed in applications such as fuselage skin, wing skins, and cowls, and the 7075-T6 alloy, used in fuselage bulkheads, wing skins and stringers [3].

Despite one-shot drilling of non-glued parts being an effective drilling strategy for reducing cycle time and avoiding non-coaxiality issues between holes, it still presents several open points when thin sheets are considered. One of the primary issues involves the gradual increase in thrust force during the drilling process, which can lead to the separation of sheets (interlayer gap), resulting in the accumulation of metallic chips between the parts. The interlayer gap occurs when the upper material undergoes elastic spring back, reverting to its initial shape after drilling. Simultaneously, the lower material bends under the influence of the applied thrust force. These phenomena become more prominent and noticeable for thin materials, resulting in a corresponding reduction in their flexural stiffness. Additionally, Luo et al. [4] and Panico et al. [5] highlighted that the interlayer gap influences the formation of burrs. To maintain structural integrity and assembly quality, it is necessary to remove the accumulated material and interface defects through operations such as disassembly, cleaning/deburring, and reassembly. However, these procedures are time-consuming, do not add value, and can be costly [6], nullifying the benefits offered by one-shot drilling. Although process parameters (rotational speed, feed rate, tool geometry) play a key role in studying the interlayer gap phenomenon, equally crucial is the role of the temporary part-fixing conditions adopted during drilling.

To date, scientific literature has primarily focused on the investigation of a direct correlation between process parameters and the measurement of burr [6,7], leaving largely unexplored the contribution that the clamping strategy of parts offers to the development of defects at the interface of two thin materials drilled in a one-shot configuration. In this study, an experimental drilling investigation was conducted on thin stacks composed of Al7075-T6 and Al2024-T3. For a fixed set of process parameters (i.e., rotational speed and feed rate), three different clamping pattern percentages were investigated (40%, 60% and 80%), also changing the drilling verse (i.e., Al2024-T3 or Al7075-T6). During drilling, the thrust force and torque were acquired. The height of the burr was then correlated with the clamping pattern, the drilling side material, and the thrust force and torque.

## Materials and Methods

The experimental one-shot drilling activities were conducted on stacks composed of Al7075-T6 sheet (300 mm x 300 mm x 1 mm) and Al2024-T3 sheet (300 mm x 300 mm x 0.8 mm). These materials were chosen due to their extensive use in the aerospace industry [9].

The stack was investigated in both the Al7075-T6/Al2024-T3 and Al2024-T3/Al7075-T6 configurations (these configurations will be referred as A/B and B/A, respectively) to analyze the influence of the flexural contribution of each material when positioned as the lower reinforcement on the hole quality. In the present experimentation, a coated carbide twist drill with a diameter of 4.76 mm, a cutting-edge height of 1 mm, a point angle of 130° and a helix angle of 25° was used.

The drilling tests were performed in dry condition by using a semi-automated electric drilling unit for industrial applications. The semi-automatic drilling unit is equipped with two independent electric motors, one dedicated to feed motion and the other to rotation (max. power 3 kW and max. rotational speed 8000 rpm). The hole-to-hole handling is performed by the operator, as well as the initiation of the drilling cycle. The drill remains in the drilling position thanks to a collar concentric to the drill bit, which, upon expansion, ensures a tight coupling within the rigid drilling jig (depicted in Fig. 1). The drilling machine is accompanied by specialized software that allows for the management of process parameters and real-time acquisition of current consumption for both motors, thrust force and torque curves over time. The upper and lower edges of each specimen were fixed to the rigid clamping fixture, while the lateral edges were left free to reproduce a typical
single structural unit of the fuselage. The experimental set-up is shown in Fig. 1. According to the literature, the drilling of aluminum alloys prefers low feed values and moderate rotational speeds to favor low thrust force and torque values and better hole quality [3]. For this reason, the experimental drilling tests were conducted with fixed process parameters: a rotational speed of 4000 rpm and a feed rate of 0.03 mm/rev. For both stack configurations, three different clamping strategies were tested by varying the fill percentage achieved using a traditional temporary fastener (Cleco fastener). The clamping strategies were chosen from those typically employed during the assembly phase of aeronautical components. The positioning of the holes is in accordance with the 5x5 matrix of holes prepared on the drilling jig, of which only the central row was utilized (see Table 1). The central row was used to mitigate edge effects and to analyse the influence of the maximum flexural contribution of the stack. The experimented clamping percentages were 40%, 60% and 80%, and for each of them, 5 holes were drilled.

<figure><img src="image_2309.png" /><figcaption>Fig. 1 a) Experimental set-up, front-view – b) Experimental set-up, lateral-view</figcaption></figure>

<table><caption>Table 1 Description of clamping strategies in line with the hole matrix on the drilling jig</caption><thead><tr><th>Filled 40%</th><th>Filled 60%</th><th>Filled 80%</th></tr></thead><tbody><tr><td><div>Temporary fastener position</div><div>Drilling point</div><div>Free position</div></td><td></td><td></td></tr></tbody></table>

One observable characteristic that influences the quality of holes in aluminum is the presence of a burr. In specific applications such as aerospace, there is a critical need to reduce or eliminate burrs that exceed the specified tolerance range. The burr at the entry side of the hole is usually
single structural unit of the fuselage. The experimental set-up is shown in Fig. 1. According to the literature, the drilling of aluminum alloys prefers low feed values and moderate rotational speeds to favor low thrust force and torque values and better hole quality [3]. For this reason, the experimental drilling tests were conducted with fixed process parameters: a rotational speed of 4000 rpm and a feed rate of 0.03 mm/rev. For both stack configurations, three different clamping strategies were tested by varying the fill percentage achieved using a traditional temporary fastener (Cleco fastener). The clamping strategies were chosen from those typically employed during the assembly phase of aeronautical components. The positioning of the holes is in accordance with the 5x5 matrix of holes prepared on the drilling jig, of which only the central row was utilized (see Table 1). The central row was used to mitigate edge effects and to analyse the influence of the maximum flexural contribution of the stack. The experimented clamping percentages were 40%, 60% and 80%, and for each of them, 5 holes were drilled.

<figure><img src="image_2300.png" /><figcaption>Fig. 1 a) Experimental set-up, front-view – b) Experimental set-up, lateral-view</figcaption></figure>

<table><caption>Table 1 Description of clamping strategies in line with the hole matrix on the drilling jig</caption><thead><tr><th>Filled 40%</th><th>Filled 60%</th><th>Filled 80%</th></tr></thead><tbody><tr><td><div>Temporary fastener position</div><div>Drilling point</div><div>Free position</div></td><td></td><td></td></tr></tbody></table>

One observable characteristic that influences the quality of holes in aluminum is the presence of a burr. In specific applications such as aerospace, there is a critical need to reduce or eliminate burrs that exceed the specified tolerance range. The burr at the entry side of the hole is usually
minor and can be easily removed by chamfering in the process of drilling a single hole. However, in a one-hole drilling configuration, it serves as an indicator of defects at the interface of the two materials in direct contact. This is particularly evident in thin configurations, where the spring back of the first drilled layer and bending of the subsequent layer, induced by thrust force, can lead to a significantly interlayer gap, thus favoring the development of defects at the interface. In contrast, the burr at the exit side is more consistent and may not always be removable due to the configuration of the assembly. To this end, a confocal microscope (Sensofar S Neox) was used to measure the height of the burr on both the top and bottom sides of each hole in both A/B and B/A drilling configurations.

## Results and Discussion

Preliminary analysis of thrust force and torque. A typical trend of thrust force and torque curves during one-shot drilling of thin Al7075-T6/Al2024-T3 (i.e., A/B configuration) is shown in Fig. 2.

<figure><img src="image_2310.png" /><figcaption>Fig. 2 Time-evolution curve of thrust force and torque during one-shot drilling</figcaption></figure>

For both stack configurations, the qualitative trend is the same, and for the sake of brevity, the curves of B/A configuration are not reported. In detail, the main stages are defined as follows:
(a) The drill bit turns and moves forward in contact with the surface.

Already from this stage, it is possible to observe a peculiarity of the one-shot drilling process of thin materials compared to what is observed for thicker materials. In fact, the increase in thrust force is not concurrent with that of torque, resulting in a pure flexion for a duration of 0.5 seconds.

(b) The increase in thrust force and torque undergoes a change in slope, representative of the onset of the actual cutting phase. Both thrust force and torque reach their respective maxima in this phase. Phase b concludes when the material at the top of the stack has been drilled through its entire thickness.

The distinctive aspect of drilling thin materials, compared to what is observed for thick ones, lies in the duration of this phase, which is closely related to the flexural behavior of the stack. The reduced flexural stiffness offered by the stack and the absence of an external element to counteract the thrust force result in a significantly prolonged tool-material engagement phase, highlighting a phenomenology more akin to forming than cutting.
minor and can be easily removed by chamfering in the process of drilling a single hole. However, in a one-hole drilling configuration, it serves as an indicator of defects at the interface of the two materials in direct contact. This is particularly evident in thin configurations, where the spring back of the first drilled layer and bending of the subsequent layer, induced by thrust force, can lead to a significantly interlayer gap, thus favoring the development of defects at the interface. In contrast, the burr at the exit side is more consistent and may not always be removable due to the configuration of the assembly. To this end, a confocal microscope (Sensofar S Neox) was used to measure the height of the burr on both the top and bottom sides of each hole in both A/B and B/A drilling configurations.

## Results and Discussion

Preliminary analysis of thrust force and torque. A typical trend of thrust force and torque curves during one-shot drilling of thin Al7075-T6/Al2024-T3 (i.e., A/B configuration) is shown in Fig. 2.

<figure><img src="image_2301.png" /><figcaption>Fig. 2 Time-evolution curve of thrust force and torque during one-shot drilling</figcaption></figure>

For both stack configurations, the qualitative trend is the same, and for the sake of brevity, the curves of B/A configuration are not reported. In detail, the main stages are defined as follows:
(a) The drill bit turns and moves forward in contact with the surface.

Already from this stage, it is possible to observe a peculiarity of the one-shot drilling process of thin materials compared to what is observed for thicker materials. In fact, the increase in thrust force is not concurrent with that of torque, resulting in a pure flexion for a duration of 0.5 seconds.

(b) The increase in thrust force and torque undergoes a change in slope, representative of the onset of the actual cutting phase. Both thrust force and torque reach their respective maxima in this phase. Phase b concludes when the material at the top of the stack has been drilled through its entire thickness.

The distinctive aspect of drilling thin materials, compared to what is observed for thick ones, lies in the duration of this phase, which is closely related to the flexural behavior of the stack. The reduced flexural stiffness offered by the stack and the absence of an external element to counteract the thrust force result in a significantly prolonged tool-material engagement phase, highlighting a phenomenology more akin to forming than cutting.
(c) In this phase, the phenomenon of the interlayer gap develops, which is evident in the abrupt decline in torque. The material at the top of the stack has been drilled, and it elastically returns to its initial configuration. Before the tool engages the next layer, there is a significant deflection caused by the thrust force.

(d) The drill bit engages and advances into the thickness of the material. The torque exhibits a second distinct peak, while there are no substantial variations in the thrust force.

(e) The thrust force and torque return to zero when the tool has completed the drilling process.

The experimental results of the maximum thrust force and torque as a function of the clamping percentage for both drilling configurations are shown in Fig. 3. The thrust force does not exhibit a clear trend in both drilling configurations. This can be anticipated, as under equivalent global boundary conditions, the percentage utilization of temporary fasteners does not provide a substantial increase in the flexural stiffness of the stack. Therefore, a significant variation in the axial force developed during drilling is not expected. As for the torque, for both drilling configurations, there is an increasing trend with the rise in the clamping percentage. Again, it is important to highlight that in numerical terms, there is no substantial variation, as for both stacks, the difference between the minimum and maximum torque values is on the order of a thousandth of N·m.

<figure><img src="image_2311.png" /><figcaption>Fig. 3 a) Thrust force values for both Al2024-T3/Al7075-T6 stacks – b) Torque values for both Al7075-T6/Al2024-T3 stacks</figcaption></figure>

Influence of drilling strategy on burr height. Based on a 3D profilometer inspection, the examined aluminum alloys manifested distinct burr morphologies attributable to their different mechanical properties. The 7075-T6 aluminum alloy revealed intermittent burrs, predominantly manifesting as discrete points. This particular variant, identified as a corona burr, is indicative of metallic materials known for their marked hardness and reduced ductility. Conversely, the 2024-T3 aluminum alloy exhibited a continuous, homogeneous, and viscous burr. This type of burr is characteristic of metals that exhibit significant ductility without excessive hardness.

As defined in the previous section, the burr that develops on the inlet side of the stack is typically smaller than that observed on the outlet side and is also automatically removable during the countersinking phase of the hole. For this reason, the maximum burr height was measured on the exit side of the stack and at the interface between the two materials (i.e. burr on the exit side for the upper material and burr on the entry side for the lower material).
(c) In this phase, the phenomenon of the interlayer gap develops, which is evident in the abrupt decline in torque. The material at the top of the stack has been drilled, and it elastically returns to its initial configuration. Before the tool engages the next layer, there is a significant deflection caused by the thrust force.

(d) The drill bit engages and advances into the thickness of the material. The torque exhibits a second distinct peak, while there are no substantial variations in the thrust force.

(e) The thrust force and torque return to zero when the tool has completed the drilling process.

The experimental results of the maximum thrust force and torque as a function of the clamping percentage for both drilling configurations are shown in Fig. 3. The thrust force does not exhibit a clear trend in both drilling configurations. This can be anticipated, as under equivalent global boundary conditions, the percentage utilization of temporary fasteners does not provide a substantial increase in the flexural stiffness of the stack. Therefore, a significant variation in the axial force developed during drilling is not expected. As for the torque, for both drilling configurations, there is an increasing trend with the rise in the clamping percentage. Again, it is important to highlight that in numerical terms, there is no substantial variation, as for both stacks, the difference between the minimum and maximum torque values is on the order of a thousandth of N·m.

<figure><img src="image_2302.png" /><figcaption>Fig. 3 a) Thrust force values for both Al2024-T3/Al7075-T6 stacks – b) Torque values for both Al7075-T6/Al2024-T3 stacks</figcaption></figure>

Influence of drilling strategy on burr height. Based on a 3D profilometer inspection, the examined aluminum alloys manifested distinct burr morphologies attributable to their different mechanical properties. The 7075-T6 aluminum alloy revealed intermittent burrs, predominantly manifesting as discrete points. This particular variant, identified as a corona burr, is indicative of metallic materials known for their marked hardness and reduced ductility. Conversely, the 2024-T3 aluminum alloy exhibited a continuous, homogeneous, and viscous burr. This type of burr is characteristic of metals that exhibit significant ductility without excessive hardness.

As defined in the previous section, the burr that develops on the inlet side of the stack is typically smaller than that observed on the outlet side and is also automatically removable during the countersinking phase of the hole. For this reason, the maximum burr height was measured on the exit side of the stack and at the interface between the two materials (i.e. burr on the exit side for the upper material and burr on the entry side for the lower material).
<figure><img src="image_2312.png" /><figcaption>Fig. 4 Maximum burr height value for Al2024-T3/Al7075-T6 stack configuration: a) Al7075-T6 exit side – b) Al7075-T6 entry side – c) Al2024-T3 exit side</figcaption></figure>

In Fig. 4, the mean value of the maximum burr height value acquired for each hole drilled in the one-shot drilling configuration of Al2024-T3/Al7075-T6 is plotted as a function of the applied clamping percentage. In particular, Fig. 4-a refers to the burr on the exit side of the stack, while Fig. 4-b and Fig. 4-c refer to the burr on the entry side of Al7075-T6 sheet and the burr on the exit side of Al2024-T3 sheet, respectively. It can be stated that the burr on the exit side of the stack decreases almost linearly as the adopted clamping percentage increases. As for the burr development on the entry side of the Al7075-T6 sheet (interface region), it exhibits a decreasing trend with an increase in the clamping percentage. This occurs most prominently during the transition from a clamping percentage of 40% to 60%, thereafter remaining relatively constant between 60% and 80%. Essentially, a low percentage of clamping elements and, consequently, a greater distance between two consecutive temporary fasteners results in a higher elastic spring back of the upper sheet, as well as a greater deflection of the lower sheet. This produces a larger interlayer gap compared to cases with a higher clamping percentage. It is important to note that the burr developed at the interface of the two materials can be influenced not only by the process itself but also by contact with the unevacuated chip. With regard to the burr developed on the exit side of the Al2024-T3 sheet, it also shows a decreasing trend with an increase in the clamping percentage, evident from the 40% to 60% clamping and then remaining almost unchanged from 60% to 80%. Similar to the analysis of the exit burr of the Al7075-T6 sheet, the reason is of a structural nature and involves the flexural stiffness of the sheets that make up the stack which will be discussed in detail later.
<figure><img src="image_2303.png" /><figcaption>Fig. 4 Maximum burr height value for Al2024-T3/Al7075-T6 stack configuration: a) Al7075-T6 exit side – b) Al7075-T6 entry side – c) Al2024-T3 exit side</figcaption></figure>

In Fig. 4, the mean value of the maximum burr height value acquired for each hole drilled in the one-shot drilling configuration of Al2024-T3/Al7075-T6 is plotted as a function of the applied clamping percentage. In particular, Fig. 4-a refers to the burr on the exit side of the stack, while Fig. 4-b and Fig. 4-c refer to the burr on the entry side of Al7075-T6 sheet and the burr on the exit side of Al2024-T3 sheet, respectively. It can be stated that the burr on the exit side of the stack decreases almost linearly as the adopted clamping percentage increases. As for the burr development on the entry side of the Al7075-T6 sheet (interface region), it exhibits a decreasing trend with an increase in the clamping percentage. This occurs most prominently during the transition from a clamping percentage of 40% to 60%, thereafter remaining relatively constant between 60% and 80%. Essentially, a low percentage of clamping elements and, consequently, a greater distance between two consecutive temporary fasteners results in a higher elastic spring back of the upper sheet, as well as a greater deflection of the lower sheet. This produces a larger interlayer gap compared to cases with a higher clamping percentage. It is important to note that the burr developed at the interface of the two materials can be influenced not only by the process itself but also by contact with the unevacuated chip. With regard to the burr developed on the exit side of the Al2024-T3 sheet, it also shows a decreasing trend with an increase in the clamping percentage, evident from the 40% to 60% clamping and then remaining almost unchanged from 60% to 80%. Similar to the analysis of the exit burr of the Al7075-T6 sheet, the reason is of a structural nature and involves the flexural stiffness of the sheets that make up the stack which will be discussed in detail later.
Similar analyses were conducted for the stack Al7075-T6/Al2024-T3. In Fig. 5, the mean value of the maximum burr height value acquired for each hole drilled in the one-shot drilling configuration of Al7075-T6/Al2024-T3 is plotted as a function of the applied clamping percentage. In particular, Fig. 5-a refers to the burr on the exit side of the stack, while Fig. 5-b and Fig. 5-c refer to the burr on the entry side of Al2024-T3 sheet and the burr on the exit side of Al7075-T6 sheet, respectively. Each trend observed with the Al2024-T3/Al7075-T6 stack is replicated in the same manner with the Al7075-T6/Al2024-T3 stacks.

<figure><img src="image_2313.png" /><figcaption>Fig. 5 Maximum burr height value for Al7075-T6/Al2024-T3 stack configuration: a) Al2024-T3 exit side – b) Al2024-T3 entry side – c) Al7075-T6 exit side</figcaption></figure>

For this reason, it becomes interesting to perform a qualitative and quantitative comparison between the maximum burr height values obtained for each stack. Since the measured burr on the entry side (interface region), as known from the literature and confirmed by the current measurements, does not reach significant values, the comparison will be conducted exclusively among the burr values acquired on the exit side of the stack (see Fig. 6-a) and on the exit side of the interface region (Fig. 6-b).
Similar analyses were conducted for the stack Al7075-T6/Al2024-T3. In Fig. 5, the mean value of the maximum burr height value acquired for each hole drilled in the one-shot drilling configuration of Al7075-T6/Al2024-T3 is plotted as a function of the applied clamping percentage. In particular, Fig. 5-a refers to the burr on the exit side of the stack, while Fig. 5-b and Fig. 5-c refer to the burr on the entry side of Al2024-T3 sheet and the burr on the exit side of Al7075-T6 sheet, respectively. Each trend observed with the Al2024-T3/Al7075-T6 stack is replicated in the same manner with the Al7075-T6/Al2024-T3 stacks.

<figure><img src="image_2304.png" /><figcaption>Fig. 5 Maximum burr height value for Al7075-T6/Al2024-T3 stack configuration: a) Al2024-T3 exit side – b) Al2024-T3 entry side – c) Al7075-T6 exit side</figcaption></figure>

For this reason, it becomes interesting to perform a qualitative and quantitative comparison between the maximum burr height values obtained for each stack. Since the measured burr on the entry side (interface region), as known from the literature and confirmed by the current measurements, does not reach significant values, the comparison will be conducted exclusively among the burr values acquired on the exit side of the stack (see Fig. 6-a) and on the exit side of the interface region (Fig. 6-b).
<figure><img src="image_2314.png" /><figcaption>Fig. 6 a) Comparison of maximum exit-side burr height values between Al2024-T3/Al7075-T6 and Al7075-T6/Al2024-T3 stacks – b) Comparison of maximum exit-side burr height values between the materials at the top of hybrid stacks (interface region)</figcaption></figure>

Starting from Fig. 6-a, it can be observed that the burr measured on the exit side of Al2024-T3 in the Al7075-T6/Al2024-T3 stack is consistently greater than that obtained for Al7075-T6 in the Al2024-T3/Al7075-T6 stack. To understand the reason, it is possible to utilize the calculation of the maximum displacement, $y_{max}$, that occurs in a simply supported beam when loaded at the mid-span:

$$y_{max} = \frac{Fl^3}{48EI} \qquad (1)$$

where $F$ is the maximum value of the thrust force; $l$ is the distance between two consecutive temporary fasteners; $E$ is the Young's Modulus, and $I$ is the Moment of Inertia of the cross-sectional area of the beam. For rectangular section, $I$ is equal to $bh^3/12$, where $b$ is the width and $h$ is the height (or thickness) of the section.

Considering the drilling schematic depicted in Fig. 2, the upper material in “phase d” no longer contributes with its flexural stiffness. Therefore, the analytical calculation exclusively considers the second layers (as well as those involved in the burr formation on the exit side of the stack).

As an illustrative case for analytical computation, the drilling performed with a clamping percentage of 40% can be considered. Utilizing the thrust force values depicted in Fig. 3-a, along with the Young's Modulus and the thickness values listed in Table 1, can be used to calculate the maximum displacement that occurs for Al2024-T3 (Al7075-T6/Al2024-T3 stack) and Al7075-T6 (Al2024-T3/Al7075-T6) at the midpoint.

As shown in the following Table 2, it measures 0.8734 mm and 0.5106 mm, respectively.

<table><caption>Table 2 Evaluation of maximum theoretical displacement for Al2024-T3 and Al7075-T6 with a clamping percentage of 40%</caption><thead><tr><th>Material</th><th>Young's Modulus, E [MPa]</th><th>Thickness, h [mm]</th><th>Width, b [mm]</th><th>Thrust force, F [N]</th><th>Span length, l [mm]</th><th>Evaluated displacement, y<sub>max</sub> [mm]</th></tr></thead><tbody><tr><td>Al2024-T3</td><td>73100</td><td>0.8</td><td>20</td><td>96.8569</td><td>30</td><td>0.8734</td></tr><tr><td>Al7075-T6</td><td>71700</td><td>1.0</td><td>20</td><td>108.4820</td><td>30</td><td>0.5106</td></tr></tbody></table>

What has been uncovered highlights a completely new phenomenon compared to what has been achieved by the scientific community thus far. For very thin materials, such as those examined in this study, boundary conditions are not negligible in determining the final quality of the hole. This
<figure><img src="image_2305.png" /><figcaption>Fig. 6 a) Comparison of maximum exit-side burr height values between Al2024-T3/Al7075-T6 and Al7075-T6/Al2024-T3 stacks – b) Comparison of maximum exit-side burr height values between the materials at the top of hybrid stacks (interface region)</figcaption></figure>

Starting from Fig. 6-a, it can be observed that the burr measured on the exit side of Al2024-T3 in the Al7075-T6/Al2024-T3 stack is consistently greater than that obtained for Al7075-T6 in the Al2024-T3/Al7075-T6 stack. To understand the reason, it is possible to utilize the calculation of the maximum displacement, $y_{max}$, that occurs in a simply supported beam when loaded at the mid-span:

$$y_{max} = \frac{Fl^3}{48EI} \qquad (1)$$

where $F$ is the maximum value of the thrust force; $l$ is the distance between two consecutive temporary fasteners; $E$ is the Young's Modulus, and $I$ is the Moment of Inertia of the cross-sectional area of the beam. For rectangular section, $I$ is equal to $bh^3/12$, where $b$ is the width and $h$ is the height (or thickness) of the section.

Considering the drilling schematic depicted in Fig. 2, the upper material in “phase d” no longer contributes with its flexural stiffness. Therefore, the analytical calculation exclusively considers the second layers (as well as those involved in the burr formation on the exit side of the stack).

As an illustrative case for analytical computation, the drilling performed with a clamping percentage of 40% can be considered. Utilizing the thrust force values depicted in Fig. 3-a, along with the Young's Modulus and the thickness values listed in Table 1, can be used to calculate the maximum displacement that occurs for Al2024-T3 (Al7075-T6/Al2024-T3 stack) and Al7075-T6 (Al2024-T3/Al7075-T6) at the midpoint.

As shown in the following Table 2, it measures 0.8734 mm and 0.5106 mm, respectively.

<table><caption>Table 2 Evaluation of maximum theoretical displacement for Al2024-T3 and Al7075-T6 with a clamping percentage of 40%</caption><thead><tr><th>Material</th><th>Young's Modulus, E [MPa]</th><th>Thickness, h [mm]</th><th>Width, b [mm]</th><th>Thrust force, F [N]</th><th>Span length, l [mm]</th><th>Evaluated displacement, y<sub>max</sub> [mm]</th></tr></thead><tbody><tr><td>Al2024-T3</td><td>73100</td><td>0.8</td><td>20</td><td>96.8569</td><td>30</td><td>0.8734</td></tr><tr><td>Al7075-T6</td><td>71700</td><td>1.0</td><td>20</td><td>108.4820</td><td>30</td><td>0.5106</td></tr></tbody></table>

What has been uncovered highlights a completely new phenomenon compared to what has been achieved by the scientific community thus far. For very thin materials, such as those examined in this study, boundary conditions are not negligible in determining the final quality of the hole. This
is because, even though the Al2024-T3/Al7075-T6 stack is subjected to a higher thrust force compared to the Al7075-T6/Al2024-T3 stack (Fig. 3-a), the greater thickness of Al7075-T6 sheet (1 mm versus 0.8 mm for the Al2024-T3 sheet) contributes to a reduction in the deflection of the sheet under the drilling thrust force, resulting in a smaller burr height (Fig. 6-b). This is because the sheet essentially undergoes a kind of deep drawing, and therefore, a greater deflection contributes to a higher burr height.

As for Fig. 6-b, it can be stated that Al7075-T6 (in the Al7075-T6/Al2024-T3) develops a larger burr compared to Al2024-T3 (in the Al2024-T37Al7075-T6 stack). The reason again lies in the flexural stiffness properties of the material constituting the stack. Specifically, when the material at the top of the stack is progressively drilled, the still intact underlaying material acts as a counterforce to the thrust force. In the case of the Al7075-T6/Al2024-T3 stack, considering the consideration related to the maximum displacement calculated with Eq.1, the Al2024-T3 sheet represents a less rigid counterforce compared to the Al7075-T6 sheet in the Al2024-T3/Al7075-T6 stack.

## Summary

The study of the hole quality was conducted considering the clamping percentage of the parts and the drilling direction for both Al2024-T3/Al7075-T6 and Al7075-T6/Al2024-T3 stacks. This study led to the following conclusions:

1. The increase in clamping percentage and drilling direction does not significantly influence the magnitude of the maximum thrust force and maximum torque. Thus, when taking into account the power consumption aspect of the semi-automatic drilling unit, significant variations are not observed. It is evident that the increase in the number of clamping elements and, consequently, the number of the pre-holes required for pre-assembly results in an increase in the drilling cycle time.

2. In a thin stack, the quality of holes obtained through one-shot drilling must necessarily be associated with structural considerations related to the flexural stiffness intrinsic to the materials and that conferred by boundary conditions. Unlike studies on one-shot drilling of thick stacks, it is no longer possible to define a direct correlation between the thrust force and the burr height that excludes significant variables for the computation of the flexural stiffness of the constituent materials of the stack.

## References

[1] A. Gloria, R. Montanari, M. Richetta, A. Varone, Alloys for Aeronautic Applications: State of the Art and Perspectives, Metals (Basel) 9 (2019) 662. https://doi.org/10.3390/met9060662

[2] R. Soni, R. Verma, R. Kumar Garg, V. Sharma, A critical review of recent advances in the aerospace materials, Mater Today Proc (2023). https://doi.org/10.1016/j.matpr.2023.08.108

[3] M. Aamir, K. Giasin, M. Tolouei-Rad, A. Vafadar, A review: drilling performance and hole quality of aluminium alloys for aerospace applications, J. Mater. Res. Technol. 9 (2020) 12484-12500. https://doi.org/10.1016/j.jmrt.2020.09.003

[4] B. Luo, K. Zhang, S. Liu, H. Cheng, R. Wang, Investigation on the interface damage in drilling low-stiffness CFRP/Ti stacks, Chinese J. Aeronaut. 32 (2019) 2211-2221. https://doi.org/10.1016/j.cja.2019.04.017

[5] Panico M, Durante M, Langella A, Boccarusso L, One-shot Drilling Process for Thin CFRP/Aluminium Alloys Stacks, Mater. Manuf. Process. (2024). https://doi.org/10.1080/10426914.2024.2311383
is because, even though the Al2024-T3/Al7075-T6 stack is subjected to a higher thrust force compared to the Al7075-T6/Al2024-T3 stack (Fig. 3-a), the greater thickness of Al7075-T6 sheet (1 mm versus 0.8 mm for the Al2024-T3 sheet) contributes to a reduction in the deflection of the sheet under the drilling thrust force, resulting in a smaller burr height (Fig. 6-b). This is because the sheet essentially undergoes a kind of deep drawing, and therefore, a greater deflection contributes to a higher burr height.

As for Fig. 6-b, it can be stated that Al7075-T6 (in the Al7075-T6/Al2024-T3) develops a larger burr compared to Al2024-T3 (in the Al2024-T37Al7075-T6 stack). The reason again lies in the flexural stiffness properties of the material constituting the stack. Specifically, when the material at the top of the stack is progressively drilled, the still intact underlaying material acts as a counterforce to the thrust force. In the case of the Al7075-T6/Al2024-T3 stack, considering the consideration related to the maximum displacement calculated with Eq.1, the Al2024-T3 sheet represents a less rigid counterforce compared to the Al7075-T6 sheet in the Al2024-T3/Al7075-T6 stack.

## Summary

The study of the hole quality was conducted considering the clamping percentage of the parts and the drilling direction for both Al2024-T3/Al7075-T6 and Al7075-T6/Al2024-T3 stacks. This study led to the following conclusions:

1. The increase in clamping percentage and drilling direction does not significantly influence the magnitude of the maximum thrust force and maximum torque. Thus, when taking into account the power consumption aspect of the semi-automatic drilling unit, significant variations are not observed. It is evident that the increase in the number of clamping elements and, consequently, the number of the pre-holes required for pre-assembly results in an increase in the drilling cycle time.

2. In a thin stack, the quality of holes obtained through one-shot drilling must necessarily be associated with structural considerations related to the flexural stiffness intrinsic to the materials and that conferred by boundary conditions. Unlike studies on one-shot drilling of thick stacks, it is no longer possible to define a direct correlation between the thrust force and the burr height that excludes significant variables for the computation of the flexural stiffness of the constituent materials of the stack.

## References

[1] A. Gloria, R. Montanari, M. Richetta, A. Varone, Alloys for Aeronautic Applications: State of the Art and Perspectives, Metals (Basel) 9 (2019) 662. https://doi.org/10.3390/met9060662

[2] R. Soni, R. Verma, R. Kumar Garg, V. Sharma, A critical review of recent advances in the aerospace materials, Mater Today Proc (2023). https://doi.org/10.1016/j.matpr.2023.08.108

[3] M. Aamir, K. Giasin, M. Tolouei-Rad, A. Vafadar, A review: drilling performance and hole quality of aluminium alloys for aerospace applications, J. Mater. Res. Technol. 9 (2020) 12484-12500. https://doi.org/10.1016/j.jmrt.2020.09.003

[4] B. Luo, K. Zhang, S. Liu, H. Cheng, R. Wang, Investigation on the interface damage in drilling low-stiffness CFRP/Ti stacks, Chinese J. Aeronaut. 32 (2019) 2211-2221. https://doi.org/10.1016/j.cja.2019.04.017

[5] Panico M, Durante M, Langella A, Boccarusso L, One-shot Drilling Process for Thin CFRP/Aluminium Alloys Stacks, Mater. Manuf. Process. (2024). https://doi.org/10.1080/10426914.2024.2311383
[6] J.M.G. de Mello, L.G. Trabasso, A.V.S. Silva, W.R. de Oliveira, Clamping force model application on the aircraft structural assembly, Int. J. Adv. Manuf. Technol. 124 (2023) 1951–1969. https://doi.org/10.1007/s00170-022-10555-y

[7] S. Bi, J. Liang, Experimental studies and optimization of process parameters for burrs in dry drilling of stacked metal materials, Int. J. Adv. Manuf. Technol 53 (2011) 867–876. https://doi.org/10.1007/s00170-010-2877-y

[8] S.N. Melkote, T.R. Newton, C. Hellstern, J.B. Morehouse, S. Turner, Interfacial Burr Formation in Drilling of Stacked Aerospace Materials, in: J.C. Aurich and D. Dornfeld (Eds.), Burrs - Analysis, Control and Removal, Springer, Berlin, Heidelberg, 2010, pp. 89–98. https://doi.org/10.1007/978-3-642-00568-8_10

[9] R.R. Boyer, J D. Cotton, M. Mohaghegh, R.E. Schafrik, Materials considerations for aerospace applications, MRS Bull, 40 (2015) 1055–1066. https://doi.org/10.1557/mrs.2015.278
[6] J.M.G. de Mello, L.G. Trabasso, A.V.S. Silva, W.R. de Oliveira, Clamping force model application on the aircraft structural assembly, Int. J. Adv. Manuf. Technol. 124 (2023) 1951–1969. https://doi.org/10.1007/s00170-022-10555-y

[7] S. Bi, J. Liang, Experimental studies and optimization of process parameters for burrs in dry drilling of stacked metal materials, Int. J. Adv. Manuf. Technol 53 (2011) 867–876. https://doi.org/10.1007/s00170-010-2877-y

[8] S.N. Melkote, T.R. Newton, C. Hellstern, J.B. Morehouse, S. Turner, Interfacial Burr Formation in Drilling of Stacked Aerospace Materials, in: J.C. Aurich and D. Dornfeld (Eds.), Burrs - Analysis, Control and Removal, Springer, Berlin, Heidelberg, 2010, pp. 89–98. https://doi.org/10.1007/978-3-642-00568-8_10

[9] R.R. Boyer, J D. Cotton, M. Mohaghegh, R.E. Schafrik, Materials considerations for aerospace applications, MRS Bull, 40 (2015) 1055–1066. https://doi.org/10.1557/mrs.2015.278
# Investigations on the production and forming of thermoplastic ceramic green tapes

TRÂN Ricardo¹,ᵃ*, MANNSCHATZ Anne²,ᵇ, PSYK Verena¹,ᶜ, GÜNTHER Anne²,ᵈ, DIX Martin¹,ᵉ

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

²Fraunhofer Institute for Ceramic Technologies and Systems IKTS, Winterbergstraße 28, 01277 Dresden, Germany

ᵃricardo.tran@iwu.fraunhofer.de, ᵇanne.mannschatz@ikts.fraunhofer.de,

ᶜverena.psyk@iwu.fraunhofer.de, ᵈanne.guenther@ikts.fraunhofer.de,

ᵉmartin.dix@iwu.fraunhofer.de

**Keywords:** Ceramic Tapes, Hydroforming, Hot Gas Forming, Ceramic Forming

**Abstract.** Technical ceramics such as alumina with its temperature stability, high mechanical stiffness, and good dielectric strength at low density, meet the requirements for highly resilient components for promising markets of high-performance electronics and the electrification of mobility. Ceramic components are usually manufactured using powder technology processes since forming of sintered ceramics is not possible due to the lack of plasticity. In this work, we use hydroforming to shape thermoplastic ceramic green tapes prior to sintering to add a third dimension to flat substrates. We developed alumina feedstocks based on a polyamide binder system that were extruded to 1 mm thick tapes, hydroformed, debinded and sintered. Depending on the binder composition, forming temperatures of 45-60 °C were sufficient, whereby precise temperature control was crucial for success. As a result, components with forming depths of up to 5 mm were produced without defects. This process offers the potential to revolutionize this market segment, not only in terms of geometric design freedom and low material waste, but also in terms of profitability of mass production.

**Introduction**

Technical ceramics are typically produced using powder technology. Starting from fine powder, a so-called green body is formed using various manufacturing processes like pressing, casting, extrusion, and injection molding [1]. Which process is most suited depends on the dimensions, complexity and number of parts to be manufactured. The fragile green body is then sintered at 1400-2000 °C depending on the ceramic material. During this heat treatment, the part densifies without melting due to material transport processes and the final polycrystalline microstructure is formed. Depending on green density and final density, the body shrinks by 45-60 vol.% [1].

The covalent or ionic bonding type of ceramics prevents dislocation movement. For example, alumina shows linear elastic behavior with fracture strains in the range of 0.1 % [2]. This makes it impossible to use typical metal forming processes on sintered ceramic parts, so that the shape has to be created prior to sintering in green state. While pressing/green machining and injection molding are state of the art for 3D-components, tape casting is used for planar products [3]. As semi-finished product, ceramic tapes are used in microelectronics as substrates on which electrical circuits or sensors are printed. Such tapes can be stacked in multilayer technology and cofired to form a 3D-structured multifunctional component [4]. Due to the drying process in the production of green tapes, the tape thickness is limited [5] and many tapes have to be laminated for 3D-components [6], which also means material waste through cut-outs. When aiming for light-weight-

<figure><img src="image_2315.png" /></figure>
# Investigations on the production and forming of thermoplastic ceramic green tapes

TRÂN Ricardo¹,ᵃ*, MANNSCHATZ Anne²,ᵇ, PSYK Verena¹,ᶜ, GÜNTHER Anne²,ᵈ, DIX Martin¹,ᵉ

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

²Fraunhofer Institute for Ceramic Technologies and Systems IKTS, Winterbergstraße 28, 01277 Dresden, Germany

ᵃricardo.tran@iwu.fraunhofer.de, ᵇanne.mannschatz@ikts.fraunhofer.de,

ᶜverena.psyk@iwu.fraunhofer.de, ᵈanne.guenther@ikts.fraunhofer.de,

ᵉmartin.dix@iwu.fraunhofer.de

**Keywords:** Ceramic Tapes, Hydroforming, Hot Gas Forming, Ceramic Forming

**Abstract.** Technical ceramics such as alumina with its temperature stability, high mechanical stiffness, and good dielectric strength at low density, meet the requirements for highly resilient components for promising markets of high-performance electronics and the electrification of mobility. Ceramic components are usually manufactured using powder technology processes since forming of sintered ceramics is not possible due to the lack of plasticity. In this work, we use hydroforming to shape thermoplastic ceramic green tapes prior to sintering to add a third dimension to flat substrates. We developed alumina feedstocks based on a polyamide binder system that were extruded to 1 mm thick tapes, hydroformed, debinded and sintered. Depending on the binder composition, forming temperatures of 45-60 °C were sufficient, whereby precise temperature control was crucial for success. As a result, components with forming depths of up to 5 mm were produced without defects. This process offers the potential to revolutionize this market segment, not only in terms of geometric design freedom and low material waste, but also in terms of profitability of mass production.

**Introduction**

Technical ceramics are typically produced using powder technology. Starting from fine powder, a so-called green body is formed using various manufacturing processes like pressing, casting, extrusion, and injection molding [1]. Which process is most suited depends on the dimensions, complexity and number of parts to be manufactured. The fragile green body is then sintered at 1400-2000 °C depending on the ceramic material. During this heat treatment, the part densifies without melting due to material transport processes and the final polycrystalline microstructure is formed. Depending on green density and final density, the body shrinks by 45-60 vol.% [1].

The covalent or ionic bonding type of ceramics prevents dislocation movement. For example, alumina shows linear elastic behavior with fracture strains in the range of 0.1 % [2]. This makes it impossible to use typical metal forming processes on sintered ceramic parts, so that the shape has to be created prior to sintering in green state. While pressing/green machining and injection molding are state of the art for 3D-components, tape casting is used for planar products [3]. As semi-finished product, ceramic tapes are used in microelectronics as substrates on which electrical circuits or sensors are printed. Such tapes can be stacked in multilayer technology and cofired to form a 3D-structured multifunctional component [4]. Due to the drying process in the production of green tapes, the tape thickness is limited [5] and many tapes have to be laminated for 3D-components [6], which also means material waste through cut-outs. When aiming for light-weight-

<figure><img src="image_2306.png" /></figure>
constructions, processing of green tapes like stamping and folding has been developed [7, 8, 9]. These processes are only possible because the green tapes contain a certain amount of polymer, which allows the tapes to be formed. Schwarzer et al. [10, 11] have developed a water-based green tape that can be formed by deep drawing. The binder components act as plasticizer, and in this case, the dimensional relaxation had to be controlled by carefully controlled humidity. One challenge in powder technology in general is the production of a homogeneous green body with the same distance between the particles to ensure isotropic shrinkage. During the tape casting process particles might orient due to shear forces which leads to differing shrinkages in lateral and z-direction and can already be observed during debinding [12, 13]. The shaping of green tapes can also influence the particle spacing as stretching in the shaping zone can lead to particle movement.

In this work, the authors chose an alternative approach using thermoplastics as binder to allow for better formability in green state. The reversible solid-viscous transition is temperature-controlled and shall be used during forming, so that less relaxation for improved dimensional stability was expected. Thermoforming is a well-established process in plastics [14] and has also been carried out for particle-filled systems such as PLA/PBS + calcium carbonate or PE, PP + talc [15, 16]. While the filler content in these cases is usually in the range of 10-30 vol. %, for ceramic processing solids loadings of 50-60 vol.% are required which is known from ceramic injection molding [17]. In forming thermoplastics, vacuum forming is well established, whereby a suction effect is created by under pressure and the heated plastic plate is drawn into a cavity [14]. However, due to the higher thickness and stiffness of the ceramic green tapes, this process seems not suitable. In this work, the sheet metal forming process hydroforming [18] is adapted to shaping the green tapes. The conventional approach uses cold tools and a water-oil emulsion as forming medium. However, for thermoplastic forming, the tool is heated above glass transition temperature of the polymer and the tape is pushed to the tool walls by the pessurized gas media (nitrogen or argon). In this way, the flat ceramic green tapes are shaped into a 3D-structure that enables complex geometries, which can currently only be produced with great effort. Potentially, the proposed process is suitable for serial production.

The aim of this first work is to evaluate the feasibility of this process for forming ceramic green tapes in order to enable easy-to-process 2D semi-finished products into 3D-components.

## Process chain for manufacturing 3D-shell components

For this study the alumina powder CT1200SG (Almatis) was used. It has irregular shaped primary particles with a mean particle size of 1.7 µm, its specific surface is 3.1 m²/g. The binder system bases on polyamide, waxes and surfactants. It was supplied in various compositions, which vary in polyamide-wax ratio and additions of modifiers. The exact composition is confidential. They are denoted with A20, A22, A26, A38, and A44, and were delivered by the company Inmatec Technologies, Rheinbach, Germany.

<figure><img src="image_2316.png" /><figcaption>Figure 1: Process route for hydroforming of ceramics.</figcaption></figure>
constructions, processing of green tapes like stamping and folding has been developed [7, 8, 9]. These processes are only possible because the green tapes contain a certain amount of polymer, which allows the tapes to be formed. Schwarzer et al. [10, 11] have developed a water-based green tape that can be formed by deep drawing. The binder components act as plasticizer, and in this case, the dimensional relaxation had to be controlled by carefully controlled humidity. One challenge in powder technology in general is the production of a homogeneous green body with the same distance between the particles to ensure isotropic shrinkage. During the tape casting process particles might orient due to shear forces which leads to differing shrinkages in lateral and z-direction and can already be observed during debinding [12, 13]. The shaping of green tapes can also influence the particle spacing as stretching in the shaping zone can lead to particle movement.

In this work, the authors chose an alternative approach using thermoplastics as binder to allow for better formability in green state. The reversible solid-viscous transition is temperature-controlled and shall be used during forming, so that less relaxation for improved dimensional stability was expected. Thermoforming is a well-established process in plastics [14] and has also been carried out for particle-filled systems such as PLA/PBS + calcium carbonate or PE, PP + talc [15, 16]. While the filler content in these cases is usually in the range of 10-30 vol. %, for ceramic processing solids loadings of 50-60 vol.% are required which is known from ceramic injection molding [17]. In forming thermoplastics, vacuum forming is well established, whereby a suction effect is created by under pressure and the heated plastic plate is drawn into a cavity [14]. However, due to the higher thickness and stiffness of the ceramic green tapes, this process seems not suitable. In this work, the sheet metal forming process hydroforming [18] is adapted to shaping the green tapes. The conventional approach uses cold tools and a water-oil emulsion as forming medium. However, for thermoplastic forming, the tool is heated above glass transition temperature of the polymer and the tape is pushed to the tool walls by the pessurized gas media (nitrogen or argon). In this way, the flat ceramic green tapes are shaped into a 3D-structure that enables complex geometries, which can currently only be produced with great effort. Potentially, the proposed process is suitable for serial production.

The aim of this first work is to evaluate the feasibility of this process for forming ceramic green tapes in order to enable easy-to-process 2D semi-finished products into 3D-components.

## Process chain for manufacturing 3D-shell components

For this study the alumina powder CT1200SG (Almatis) was used. It has irregular shaped primary particles with a mean particle size of 1.7 µm, its specific surface is 3.1 m²/g. The binder system bases on polyamide, waxes and surfactants. It was supplied in various compositions, which vary in polyamide-wax ratio and additions of modifiers. The exact composition is confidential. They are denoted with A20, A22, A26, A38, and A44, and were delivered by the company Inmatec Technologies, Rheinbach, Germany.

<figure><img src="image_2307.png" /><figcaption>Figure 1: Process route for hydroforming of ceramics.</figcaption></figure>
The production process is shown in Fig. 1. First, powder and binder were mixed and kneaded at 150°C to form a homogeneous feedstock in which particle agglomerates are broken and particles are coated by a polymer layer. The solid loading was 57 vol.%. This compounding step was carried out in three passes using a twin screw extruder KETSE 20/40 (Brabender, Germany). The feedstock was extruded at 150 °C through a 3.5 mm round die and granulated by cutting. The granulates enabled a continuous feed into the extruder and a homogenous mass flow during tape extrusion. Tapes were extruded using the same device equipped with a slot die with a width of 150 mm and a thickness of 1.0 mm. The endless tapes were cut to sheets of a size of 150 mm x 150 mm and placed into the hydroforming tool. The forming process is discussed in detail in the following section Hydroforming of Green Tapes. After shaping the green part, the samples were debinded and sintered. Debinding was a two-step procedure starting with extraction debinding in acetone for 48 h (MDU-30, Debatec) where 72 % of the binder was extracted. The remaining polymer was decomposed by thermal treatment in air at temperatures of up to 600 °C. Sintering was performed with a heating rate of 3 K/min up to a temperature of 1650 °C and then kept at that temperature for 2 h. Shrinkage of the tapes was measured using samples cut to a size of 20 mm x 40 mm. They were aligned in parallel, perpendicular, and in 45 ° to extrusion direction.

## Characterization of green tapes

The thermomechanical behavior of the green tapes was tested by two methods: DMA (dynamic mechanical analysis) and tensile testing at elevated temperatures. In the green state, the feedstock behaves like a filled polymer instead of a ceramic, as the particles are dispersed in a polymer matrix. In DMA, the visco-elastic behavior is characterized by oscillating rheological measurements. Polymers have elastic and viscous components in their mechanical response, which can be described by the complex dynamic modulus G, where G' is the storage modulus and G'' is the loss modulus. The amplitudes of stress σ₀ and strain ε₀ have a phase shift δ (loss factor tan δ) between them. (Eq. 1-3)

$$G = G' + iG'' \quad (1)$$

$$G' = \frac{\sigma_0}{\varepsilon_0} \cos \delta \quad (2)$$

$$G'' = \frac{\sigma_0}{\varepsilon_0} \sin \delta \quad (3)$$

For the oscillating rheological measurements, rectangular, solid specimens were cut in extrusion direction with a size of 55 mm x 10 mm x 1 mm. The specimens were mounted on a SRF (solid rectangular fixture) sample holder of a rheometer MCR302 by Anton Paar. A sinusoidal force was applied in torsion on the specimen and the response was measured. Prior to temperature dependent measurements an amplitude sweep was carried out between 0.001% and 1% to determine the linear-visco-elastic region. The amplitude was set to 0.01 %. The oscillation measurements were carried out with 1 Hz in a temperature range of 20-100 °C with a heating rate of 2 K/min. Two specimens were tested per sample and the results were in good agreement. These parameters comply with ASTM standard D7028-07.
The production process is shown in Fig. 1. First, powder and binder were mixed and kneaded at 150°C to form a homogeneous feedstock in which particle agglomerates are broken and particles are coated by a polymer layer. The solid loading was 57 vol.%. This compounding step was carried out in three passes using a twin screw extruder KETSE 20/40 (Brabender, Germany). The feedstock was extruded at 150 °C through a 3.5 mm round die and granulated by cutting. The granulates enabled a continuous feed into the extruder and a homogenous mass flow during tape extrusion. Tapes were extruded using the same device equipped with a slot die with a width of 150 mm and a thickness of 1.0 mm. The endless tapes were cut to sheets of a size of 150 mm x 150 mm and placed into the hydroforming tool. The forming process is discussed in detail in the following section Hydroforming of Green Tapes. After shaping the green part, the samples were debinded and sintered. Debinding was a two-step procedure starting with extraction debinding in acetone for 48 h (MDU-30, Debatec) where 72 % of the binder was extracted. The remaining polymer was decomposed by thermal treatment in air at temperatures of up to 600 °C. Sintering was performed with a heating rate of 3 K/min up to a temperature of 1650 °C and then kept at that temperature for 2 h. Shrinkage of the tapes was measured using samples cut to a size of 20 mm x 40 mm. They were aligned in parallel, perpendicular, and in 45 ° to extrusion direction.

## Characterization of green tapes

The thermomechanical behavior of the green tapes was tested by two methods: DMA (dynamic mechanical analysis) and tensile testing at elevated temperatures. In the green state, the feedstock behaves like a filled polymer instead of a ceramic, as the particles are dispersed in a polymer matrix. In DMA, the visco-elastic behavior is characterized by oscillating rheological measurements. Polymers have elastic and viscous components in their mechanical response, which can be described by the complex dynamic modulus G, where G' is the storage modulus and G'' is the loss modulus. The amplitudes of stress σ₀ and strain ε₀ have a phase shift δ (loss factor tan δ) between them. (Eq. 1-3)

$$G = G' + iG'' \qquad (1)$$

$$G' = \frac{\sigma_0}{\varepsilon_0} \cos \delta \qquad (2)$$

$$G'' = \frac{\sigma_0}{\varepsilon_0} \sin \delta \qquad (3)$$

For the oscillating rheological measurements, rectangular, solid specimens were cut in extrusion direction with a size of 55 mm x 10 mm x 1 mm. The specimens were mounted on a SRF (solid rectangular fixture) sample holder of a rheometer MCR302 by Anton Paar. A sinusoidal force was applied in torsion on the specimen and the response was measured. Prior to temperature dependent measurements an amplitude sweep was carried out between 0.001% and 1% to determine the linear-visco-elastic region. The amplitude was set to 0.01 %. The oscillation measurements were carried out with 1 Hz in a temperature range of 20-100 °C with a heating rate of 2 K/min. Two specimens were tested per sample and the results were in good agreement. These parameters comply with ASTM standard D7028-07.
<figure><img src="image_2317.png" /><figcaption>Figure 2: Storage (SM) and loss modulus (LM) of green tapes measured in oscillating rheometry.</figcaption></figure>

Fig. 2 shows storage and loss moduli for all samples. They have similar behavior since they base on the same binder system. In the low temperature region, both moduli are high. The stored elastic energy, that leads to relaxation after stresses are removed, is one magnitude higher than the loss energy which is consumed by molecule movement and friction and leads to viscous flow. The tapes appear as a solid but flexible body. At higher temperatures both moduli have dropped down by almost two magnitudes which marks the transition to a polymer melt and marks also the glass-transition temperature of polyamide. The transition area between 45 °C and 60 °C is the proposed process window for the following hydroforming experiments.

Tensile tests inspired by DIN EN ISO 527-1 were carried out on a Zwick Z020 using a strain gauge-based load cell. The test set-up including details about the shape and size of the samples is shown in Fig. 3. A total of 60 tests were performed. Hot air flow was used from one side of the specimens to heat them up and to control the temperature, which was measured using an infrared camera. The representative thermal image in Fig. 3 (d) demonstrates the homogeneity of the temperature, which was checked for both sides of the specimen in pre-tests. The specimen was a shoulder bar (Fig. 3 (b)) with a testing length shortened to 20 mm. This deviation from the standard specimen geometry was made to ensure a more even heat distribution. Based on the results of the DMA measurements, 40 °C, 50 °C and 60 °C were defined as test temperatures. The elongation of the tensile specimen is determined via the crosshead extension. The test speed was 1.0 mm/min.
<figure><img src="image_2308.png" /><figcaption>Figure 2: Storage (SM) and loss modulus (LM) of green tapes measured in oscillating rheometry.</figcaption></figure>

Fig. 2 shows storage and loss moduli for all samples. They have similar behavior since they base on the same binder system. In the low temperature region, both moduli are high. The stored elastic energy, that leads to relaxation after stresses are removed, is one magnitude higher than the loss energy which is consumed by molecule movement and friction and leads to viscous flow. The tapes appear as a solid but flexible body. At higher temperatures both moduli have dropped down by almost two magnitudes which marks the transition to a polymer melt and marks also the glass-transition temperature of polyamide. The transition area between 45 °C and 60 °C is the proposed process window for the following hydroforming experiments.

Tensile tests inspired by DIN EN ISO 527-1 were carried out on a Zwick Z020 using a strain gauge-based load cell. The test set-up including details about the shape and size of the samples is shown in Fig. 3. A total of 60 tests were performed. Hot air flow was used from one side of the specimens to heat them up and to control the temperature, which was measured using an infrared camera. The representative thermal image in Fig. 3 (d) demonstrates the homogeneity of the temperature, which was checked for both sides of the specimen in pre-tests. The specimen was a shoulder bar (Fig. 3 (b)) with a testing length shortened to 20 mm. This deviation from the standard specimen geometry was made to ensure a more even heat distribution. Based on the results of the DMA measurements, 40 °C, 50 °C and 60 °C were defined as test temperatures. The elongation of the tensile specimen is determined via the crosshead extension. The test speed was 1.0 mm/min.
<figure><img src="image_2318.png" /><figcaption>Figure 3: Tensile tests for ceramic green tape (a) stress-strain-curve for sample A22, (b) specimen geometry, (c) experimental setup, (d) heat distribution of heated sample</figcaption></figure>

The binder components are semicrystalline polymers. Since the ceramic particles act as crystallization nuclei, it is assumed that the polymer matrix is mostly crystalline with molecules coiled and folded. At the lowest temperature considered (40 °C) the recorded stress strain behavior features a linear rise of the curve indicating linear elastic deformation followed by a non-linear curve section typically indicating plastic flow. In this section of the curve, molecules unfold and are stretched until finally fracture occurs at approximately 12 % strain. At the highest considered temperature (60 °C), which is within the range of the glass transition temperatures of polyamides, the mobility of the polymer chains is considerably higher and crystalline structures gradually dissolve. Consequently, the stresses necessary to achieve a defined elongation are significantly lower and the strain at fracture is higher compared to the material behavior at 40 °C. The intermediate temperature of 50 °C marks a transitional state in which stress build-ups and releases alternate. On molecular scale, crystalline areas unfold when a certain yield stress is reached followed by the next build-up. This behavior is represented in the curve via a wavy shape. Generally, the tensile tests have confirmed the process window identified via DMA (45 - 60 °C).

## Hydroforming of green tapes

Based on the results of the characterization of the thermoplastic ceramic green tapes and especially considering the identified process window, the hydroforming process and the corresponding hydroforming tool was designed. As shown in Fig. 4 (a) the forming tool consists of an upper and lower half, which include two heated segments. The latter clamp and heat up the green tapes. The tools are made of an aluminum alloy for better heat transfer. The lower half of the tool has a heated cavity in which the green tape is formed. The upper half of the tool is flat and has an opening through which the forming medium (nitrogen) is injected. The ceramic green tape is formed into the cavity by building up an internal pressure of 2.0 - 3.5 MPa. When the forming process is complete, the tool is opened. The part is cooled by compressed air to retain its shape and removed. However, as with other forming processes at elevated temperatures, temperature control throughout the process is a critical aspect.
<figure><img src="image_2309.png" /><figcaption>Figure 3: Tensile tests for ceramic green tape (a) stress-strain-curve for sample A22, (b) specimen geometry, (c) experimental setup, (d) heat distribution of heated sample</figcaption></figure>

The binder components are semicrystalline polymers. Since the ceramic particles act as crystallization nuclei, it is assumed that the polymer matrix is mostly crystalline with molecules coiled and folded. At the lowest temperature considered (40 °C) the recorded stress strain behavior features a linear rise of the curve indicating linear elastic deformation followed by a non-linear curve section typically indicating plastic flow. In this section of the curve, molecules unfold and are stretched until finally fracture occurs at approximately 12 % strain. At the highest considered temperature (60 °C), which is within the range of the glass transition temperatures of polyamides, the mobility of the polymer chains is considerably higher and crystalline structures gradually dissolve. Consequently, the stresses necessary to achieve a defined elongation are significantly lower and the strain at fracture is higher compared to the material behavior at 40 °C. The intermediate temperature of 50 °C marks a transitional state in which stress build-ups and releases alternate. On molecular scale, crystalline areas unfold when a certain yield stress is reached followed by the next build-up. This behavior is represented in the curve via a wavy shape. Generally, the tensile tests have confirmed the process window identified via DMA (45 - 60 °C).

## Hydroforming of green tapes

Based on the results of the characterization of the thermoplastic ceramic green tapes and especially considering the identified process window, the hydroforming process and the corresponding hydroforming tool was designed. As shown in Fig. 4 (a) the forming tool consists of an upper and lower half, which include two heated segments. The latter clamp and heat up the green tapes. The tools are made of an aluminum alloy for better heat transfer. The lower half of the tool has a heated cavity in which the green tape is formed. The upper half of the tool is flat and has an opening through which the forming medium (nitrogen) is injected. The ceramic green tape is formed into the cavity by building up an internal pressure of 2.0 - 3.5 MPa. When the forming process is complete, the tool is opened. The part is cooled by compressed air to retain its shape and removed. However, as with other forming processes at elevated temperatures, temperature control throughout the process is a critical aspect.
<figure><img src="image_2319.png" /><figcaption>Figure 4: Developed experimental setup (a) and part geometry (b) for hydroforming trials.</figcaption></figure>

In the context of the process design, the time required for the 1.0 mm thick tape to completely heat up to 60°C was determined using a thermocouple inside the green tape. It was 120 s. Consequently, the forming experiments were conducted at a certain temperature in the range of 45°C-60°C after a heating time of 120 s as shown in the process flow diagram in Fig. 5 (b). This is followed by the hydroforming process in which the clamping force is applied to seal the tape and the gaseous forming medium is introduced. After a linear pressure build-up of up to 2.5 MPa, the internal pressure is briefly retained before it is released, and the tool is opened. Cooling to ambient air is sufficient to remove the form-stable part from the tool.

Finding the right process temperature proved to be the most important task. Experiments were conducted at 45 °C, 50 °C, 55 °C, and 60 °C for the tapes with different binder systems. Fig. 5 (c) summarizes the characteristic defect patterns due to inappropriate process parameters. They can be categorized in brittle cracks at too low temperatures and collapsed structures with wrinkles at too high temperatures. These effects can be correlated to the visco-elastic behavior measured by DMA as shown in Fig.5 (a). If the storage modulus is at 100 MPa and below the strength seems too low and the samples collapse. This becomes obvious when comparing the results of A22 with A20, A26, A44 which had good results at 50 °C while A22 already collapsed. At even higher temperatures and thereby lower strength all binder compositions showed deformations. For 45 °C, all compositions, except A22, were too brittle and cracks formed at edges and corners. For sample A38, it was possible to achieve defect-free samples at 50 °C despite its high storage modulus of 250 MPa. These findings show that the temperature window for defect-free forming covers a range of approximately 5 K and is therefore extremely narrow. Both lower and higher temperatures lead to defects. While at too low temperatures the higher strength is associated with brittleness, resulting in cracks at radii transitions and corners, at too high temperatures the material tends to wrinkle and corrugate. In extreme cases the entire green tape collapses. The hydroforming tests showed that already low internal pressures were sufficient for complete forming. However, higher internal pressures could compensate for wrinkling at higher temperatures.
<figure><img src="image_2310.png" /><figcaption>Figure 4: Developed experimental setup (a) and part geometry (b) for hydroforming trials.</figcaption></figure>

In the context of the process design, the time required for the 1.0 mm thick tape to completely heat up to 60°C was determined using a thermocouple inside the green tape. It was 120 s. Consequently, the forming experiments were conducted at a certain temperature in the range of 45°C-60°C after a heating time of 120 s as shown in the process flow diagram in Fig. 5 (b). This is followed by the hydroforming process in which the clamping force is applied to seal the tape and the gaseous forming medium is introduced. After a linear pressure build-up of up to 2.5 MPa, the internal pressure is briefly retained before it is released, and the tool is opened. Cooling to ambient air is sufficient to remove the form-stable part from the tool.

Finding the right process temperature proved to be the most important task. Experiments were conducted at 45 °C, 50 °C, 55 °C, and 60 °C for the tapes with different binder systems. Fig. 5 (c) summarizes the characteristic defect patterns due to inappropriate process parameters. They can be categorized in brittle cracks at too low temperatures and collapsed structures with wrinkles at too high temperatures. These effects can be correlated to the visco-elastic behavior measured by DMA as shown in Fig.5 (a). If the storage modulus is at 100 MPa and below the strength seems too low and the samples collapse. This becomes obvious when comparing the results of A22 with A20, A26, A44 which had good results at 50 °C while A22 already collapsed. At even higher temperatures and thereby lower strength all binder compositions showed deformations. For 45 °C, all compositions, except A22, were too brittle and cracks formed at edges and corners. For sample A38, it was possible to achieve defect-free samples at 50 °C despite its high storage modulus of 250 MPa. These findings show that the temperature window for defect-free forming covers a range of approximately 5 K and is therefore extremely narrow. Both lower and higher temperatures lead to defects. While at too low temperatures the higher strength is associated with brittleness, resulting in cracks at radii transitions and corners, at too high temperatures the material tends to wrinkle and corrugate. In extreme cases the entire green tape collapses. The hydroforming tests showed that already low internal pressures were sufficient for complete forming. However, higher internal pressures could compensate for wrinkling at higher temperatures.
<figure><img src="image_2320.png" /><figcaption>(a) Storage modulus of green tapes measured in DMA and defect types, detail from Fig. 3</figcaption></figure>

<figure><img src="image_2321.png" /><figcaption>(c) Characteristic defects in dependence of temperature range</figcaption></figure>

**Figure 2:** Close-up of DMA measurements (a) in linear scale within the temperature process window. Process flow diagram for temperature and forming pressure (b). Defect types for non-optimal heating conditions (c).

Depending on the binder composition, forming temperatures between 50 °C and 55 °C have been identified to provide the best compromise of strength and formability. Green parts could be reproducibly formed in this temperature range, as shown in Fig. 6. Visual inspection and 3D coordinate measurements confirmed that the target geometry could be reached without relaxation effects or visible defects.

<figure><img src="image_2322.png" /><figcaption>**Figure 3:** Images of completely formed ceramic green tapes without visible defects.</figcaption></figure>

After forming, the green parts were debinded and sintered. As a result of sintering, the components shrank considerably, as shown in Fig. 7. Due to the extrusion process, the polymer chains are stretched which leads to directional shrinkage. Precisely, shrinkage parallel to the extrusion direction was 17.3%, while shrinkages at angles of 90° and 45° to the extrusion direction were 14.4% and 15.5%, respectively. These data were obtained using cut outs of a size of 20 mm x 40 mm. This effect was also visible in the produced parts, so that the dimensions differed by
<figure><img src="image_2311.png" /><figcaption>(a) Storage modulus of green tapes measured in DMA and defect types, detail from Fig. 3</figcaption></figure>

<figure><img src="image_2312.png" /><figcaption>(c) Characteristic defects in dependence of temperature range</figcaption></figure>

**Figure 2:** Close-up of DMA measurements (a) in linear scale within the temperature process window. Process flow diagram for temperature and forming pressure (b). Defect types for non-optimal heating conditions (c).

Depending on the binder composition, forming temperatures between 50 °C and 55 °C have been identified to provide the best compromise of strength and formability. Green parts could be reproducibly formed in this temperature range, as shown in Fig. 6. Visual inspection and 3D coordinate measurements confirmed that the target geometry could be reached without relaxation effects or visible defects.

<figure><img src="image_2313.png" /><figcaption>**Figure 3:** Images of completely formed ceramic green tapes without visible defects.</figcaption></figure>

After forming, the green parts were debinded and sintered. As a result of sintering, the components shrank considerably, as shown in Fig. 7. Due to the extrusion process, the polymer chains are stretched which leads to directional shrinkage. Precisely, shrinkage parallel to the extrusion direction was 17.3%, while shrinkages at angles of 90° and 45° to the extrusion direction were 14.4% and 15.5%, respectively. These data were obtained using cut outs of a size of 20 mm x 40 mm. This effect was also visible in the produced parts, so that the dimensions differed by
almost 2 mm parallel and perpendicular to the extrusion direction. This direction-dependent shrinkage must be compensated for in the design for future ceramic components when calculating the oversize factors.

<figure><img src="image_2323.png" /><figcaption>Figure 4: Green part (left) produced within the processing window and sintered part (right).</figcaption></figure>

The sintering process itself also faces a number of challenges. For example, the manufactured samples tend to collapse during sintering if flat areas are unsupported which is caused by creep processes at 1650 °C sintering temperature. This can be remedied by both design adjustments and modifications during the sintering process, such as the inclusion support structures. In general, however, it has been found that no defects are introduced during the forming process that would subsequently lead to component failure during the sintering process.

## Summary

This study presents basic investigations on the production and processing of ceramic green tapes using thermoplastic binders. As part of the investigations, highly filled tapes with 57 vol.% Al₂O₃ particles were extruded using polyamide-based thermoplastic binders, providing ceramic flat material for a subsequent hydroforming process. The PA binder system provides the necessary formability at elevated forming temperatures to produce complex 3D ceramic green parts. The thermoplastic properties could be utilized as part of the forming process. The temperature-controlled softening and solidification process of thermoplastics is suitable for producing dimensionally stable parts without relaxation phenomena. Subsequent process steps, in particular debinding and sintering, are well established in industry and are known from ceramic injection moulding (CIM). DMA studies and tensile tests have shown that the properties of the tapes are highly temperature dependent, which is later reflected in the temperature sensitivity during forming. Defects in the production of green parts are caused by temperatures that are either too low, resulting in increased brittleness, or too high, resulting in insufficient material stability. Even and complete heating of the part before forming and accurate temperature control is therefore essential to meet the process window.

Overall, the following key results were obtained:

- Ceramic particles in thermoplastic binders can be processed into green tapes by extrusion, which can be used as flat material in further process steps.

- Preliminary mechanical tests show that plastic strains of 12 % to 18 % are possible at elevated temperatures, which clearly demonstrates the potential for forming processes.

- The developed hydroforming process at 50°C is suitable for forming green tapes into complex shaped ceramic green parts. The temperature window is highly dependent on the binder composition and must be very precisely controlled during the process.
almost 2 mm parallel and perpendicular to the extrusion direction. This direction-dependent shrinkage must be compensated for in the design for future ceramic components when calculating the oversize factors.

<figure><img src="image_2314.png" /><figcaption>Figure 4: Green part (left) produced within the processing window and sintered part (right).</figcaption></figure>

The sintering process itself also faces a number of challenges. For example, the manufactured samples tend to collapse during sintering if flat areas are unsupported which is caused by creep processes at 1650 °C sintering temperature. This can be remedied by both design adjustments and modifications during the sintering process, such as the inclusion support structures. In general, however, it has been found that no defects are introduced during the forming process that would subsequently lead to component failure during the sintering process.

## Summary

This study presents basic investigations on the production and processing of ceramic green tapes using thermoplastic binders. As part of the investigations, highly filled tapes with 57 vol.% Al₂O₃ particles were extruded using polyamide-based thermoplastic binders, providing ceramic flat material for a subsequent hydroforming process. The PA binder system provides the necessary formability at elevated forming temperatures to produce complex 3D ceramic green parts. The thermoplastic properties could be utilized as part of the forming process. The temperature-controlled softening and solidification process of thermoplastics is suitable for producing dimensionally stable parts without relaxation phenomena. Subsequent process steps, in particular debinding and sintering, are well established in industry and are known from ceramic injection moulding (CIM). DMA studies and tensile tests have shown that the properties of the tapes are highly temperature dependent, which is later reflected in the temperature sensitivity during forming. Defects in the production of green parts are caused by temperatures that are either too low, resulting in increased brittleness, or too high, resulting in insufficient material stability. Even and complete heating of the part before forming and accurate temperature control is therefore essential to meet the process window.

Overall, the following key results were obtained:

- Ceramic particles in thermoplastic binders can be processed into green tapes by extrusion, which can be used as flat material in further process steps.

- Preliminary mechanical tests show that plastic strains of 12 % to 18 % are possible at elevated temperatures, which clearly demonstrates the potential for forming processes.

- The developed hydroforming process at 50°C is suitable for forming green tapes into complex shaped ceramic green parts. The temperature window is highly dependent on the binder composition and must be very precisely controlled during the process.
- Subsequent debinding and sintering processes have enabled the production of ceramic parts based on the green parts. However, the shrinkage of the parts as a result of sintering must be taken into account in the part design.

The results presented demonstrate that ceramic green tapes with thermoplastic binders can be produced by extrusion and formed into 3D-geometries by hydroforming. The process is suitable for series production and has the potential to produce temperature-stable housings or curved substrates for microelectronics.

## Acknowledgements

The research project was funded as part of the Fraunhofer-Gesellschaft's SME-oriented research program. The authors are thankful to Inmatec Technologies for supplying the thermoplastic binder.

## References

[1] P. Boch, J.-P. Nièpce, Ceramic Materials: Processes, Properties, and Applications, Hoboken: Wiley-ISTE, 2010.

[2] E. H. Lutz, M. V. Swain, Stress-Strain Behavior of Alumina, Magnesia-Partially-Stabilized Zirconia, and Duplex Ceramics and Its Relevance for Flaw Resistance, KR-Curve Behavior, and Thermal Shock Behavior, J.Am. Ceram. Soc, 75 (1992) 3058-64. https://doi.org/10.1111/j.1151-2916.1992.tb04386.x

[3] A. Roosen, Foliengießen: Verfahren zur Herstellung planarer und dreidimensionaler keramischer Strukturen, Sonderdruck, in H. Reh (Ed.) Das Keramiker Jahrbuch 1998 (1998), Wiesbaden, p 42-53

[4] L. Khoong, Y. Tan, Y. Lam, Overview on fabrication of three-dimensional structures in multi-layer ceramic substrate, J. Europ. Ceram. Soc., 30 (2010) 1973-1987. https://doi.org/10.1016/j.jeurceramsoc.2010.03.011

[5] R. E. Mistler, Tape casting: past, present, potential, Am.Ceram.Soc. Bull. 77 (1998), 82-86

[6] K. Schindler, A. Roosen, Manufacture of 3D structures by cold low pressure lamination of ceramic green tapes, J.Europ.Ceram.Soc., 29 (2009) 899-904. https://doi.org/10.1016/j.jeurceramsoc.2008.07.041.

[7] A. Härtel, H.-J. Richter, T. Moritz, Cold embossing of ceramic green tapes, J. Microelectron. Electron. Packag. 9 (2012) 138-143. https://doi.org/10.4071/imaps.337

[8] U. Scheithauer, M. Dannowki, E. Schwarzer, T. Slawik, Innovative kiln furniture, their influence on temperature distribution within the kiln, and a new production technology, Interceram-Int. Ceram. Rev. 63 (2014) 312-316.

[9] A. Grimm, S. Bast, R. Tillmanns, M. Schumacher, Keramik Leichtbauteile Teil 1, Keram. Zeitschrift 58 (2006) 8-10

[10] E. Schwarzer, u. Scheithauer, M. Brosche, H.-J. Richter, T. Moritz, Deep Drawing of Ceramic Green Tapes – A Promising Technology for Producing New Lightweight Non-Planar Kiln Furniture, J. Ceram. Sci. Tech. 08 (2017), 169-176. DOI: 10.4416/JCST2016-00106

[11] E. Schwarzer, L. Olawsky, F. Hagen, U. Scheithauer, T. Moritz, H.-J. Richter, A. Michaelis, 3D-customized thin-walled ceramic mouldings produced by deep drawing of green tapes, J. Europ. Ceram. Soc. (2019) 4911-4920. https://doi.org/10.1016/j.jeurceramsoc.2019.07.001

[12] Z. Fu, A. Roosen, Shrinkage of tape cast products during binder burnout, J. Am. Ceram. Soc. 98 (2015) 20-29. https://doi.org/10.1111/jace.13270
- Subsequent debinding and sintering processes have enabled the production of ceramic parts based on the green parts. However, the shrinkage of the parts as a result of sintering must be taken into account in the part design.

The results presented demonstrate that ceramic green tapes with thermoplastic binders can be produced by extrusion and formed into 3D-geometries by hydroforming. The process is suitable for series production and has the potential to produce temperature-stable housings or curved substrates for microelectronics.

## Acknowledgements

The research project was funded as part of the Fraunhofer-Gesellschaft's SME-oriented research program. The authors are thankful to Inmatec Technologies for supplying the thermoplastic binder.

## References

[1] P. Boch, J.-P. Nièpce, Ceramic Materials: Processes, Properties, and Applications, Hoboken: Wiley-ISTE, 2010.

[2] E. H. Lutz, M. V. Swain, Stress-Strain Behavior of Alumina, Magnesia-Partially-Stabilized Zirconia, and Duplex Ceramics and Its Relevance for Flaw Resistance, KR-Curve Behavior, and Thermal Shock Behavior, J.Am. Ceram. Soc, 75 (1992) 3058-64. https://doi.org/10.1111/j.1151-2916.1992.tb04386.x

[3] A. Roosen, Foliengießen: Verfahren zur Herstellung planarer und dreidimensionaler keramischer Strukturen, Sonderdruck, in H. Reh (Ed.) Das Keramiker Jahrbuch 1998 (1998), Wiesbaden, p 42-53

[4] L. Khoong, Y. Tan, Y. Lam, Overview on fabrication of three-dimensional structures in multi-layer ceramic substrate, J. Europ. Ceram. Soc., 30 (2010) 1973-1987. https://doi.org/10.1016/j.jeurceramsoc.2010.03.011

[5] R. E. Mistler, Tape casting: past, present, potential, Am.Ceram.Soc. Bull. 77 (1998), 82-86

[6] K. Schindler, A. Roosen, Manufacture of 3D structures by cold low pressure lamination of ceramic green tapes, J.Europ.Ceram.Soc., 29 (2009) 899-904. https://doi.org/10.1016/j.jeurceramsoc.2008.07.041.

[7] A. Härtel, H.-J. Richter, T. Moritz, Cold embossing of ceramic green tapes, J. Microelectron. Electron. Packag. 9 (2012) 138-143. https://doi.org/10.4071/imaps.337

[8] U. Scheithauer, M. Dannowki, E. Schwarzer, T. Slawik, Innovative kiln furniture, their influence on temperature distribution within the kiln, and a new production technology, Interceram-Int. Ceram. Rev. 63 (2014) 312-316.

[9] A. Grimm, S. Bast, R. Tillmanns, M. Schumacher, Keramik Leichtbauteile Teil 1, Keram. Zeitschrift 58 (2006) 8-10

[10] E. Schwarzer, u. Scheithauer, M. Brosche, H.-J. Richter, T. Moritz, Deep Drawing of Ceramic Green Tapes – A Promising Technology for Producing New Lightweight Non-Planar Kiln Furniture, J. Ceram. Sci. Tech. 08 (2017), 169-176. DOI: 10.4416/JCST2016-00106

[11] E. Schwarzer, L. Olawsky, F. Hagen, U. Scheithauer, T. Moritz, H.-J. Richter, A. Michaelis, 3D-customized thin-walled ceramic mouldings produced by deep drawing of green tapes, J. Europ. Ceram. Soc. (2019) 4911-4920. https://doi.org/10.1016/j.jeurceramsoc.2019.07.001

[12] Z. Fu, A. Roosen, Shrinkage of tape cast products during binder burnout, J. Am. Ceram. Soc. 98 (2015) 20-29. https://doi.org/10.1111/jace.13270
[13] Z. Fu, P. Polfer, T. Kraft, A. Roosen, Correlation between anisotropic green microstructure of spherical-shaped alumina particles and their shrinkage behavior, J. Am. Ceram. Soc. 98 (2015) 3438-3444. https://doi.org/10.1111/jace.13567

[14] P. Klein, Fundamentals of Plastics Thermoforming, Morgan & Claypool Publishers, 2009.

[15] M. Barletta, M. Puopolo, Thermoforming of compostable PLA/PBS blends reinforced with highly hygroscopic calcium carbonate, J. Manuf. Proc. 56 (2020) 1185-1192, https://doi.org/10.1016/j.jmapro.2020.06.008

[16] C. H. Suh, J. L. White, Particle and polymer chain orientation development in thermoformed talc-filled thermoplastics, Polym. Eng. Sci 36 (1996) 2188-2197. https://doi.org/10.1002/pen.10616

[17] B. C. Mutsuddy, R. G. Ford, Ceramic Injection Molding, London: Chapman & Hall, 1995.

[18] R. Tran, F. Reuther, S. Winter, V. Pysk, Process Development for a Superplastic Hot Tube Gas Forming Process of Titanium (Ti-3Al-2.5V) Hollow Profiles, Metals 10, (2020). https://doi.org/10.3390/met10091150

[19] A. Franck, K. Biederbrick, Kunststoff-Kompendium, Würzburg: Vogel-Fachbuch: Technik, 1990.
[13] Z. Fu, P. Polfer, T. Kraft, A. Roosen, Correlation between anisotropic green microstructure of spherical-shaped alumina particles and their shrinkage behavior, J. Am. Ceram. Soc. 98 (2015) 3438-3444. https://doi.org/full/10.1111/jace.13567

[14] P. Klein, Fundamentals of Plastics Thermoforming, Morgan & Claypool Publishers, 2009.

[15] M. Barletta, M. Puopolo, Thermoforming of compostable PLA/PBS blends reinforced with highly hygroscopic calcium carbonate, J. Manuf. Proc. 56 (2020) 1185-1192, https://doi.org/10.1016/j.jmapro.2020.06.008

[16] C. H. Suh, J. L. White, Particle and polymer chain orientation development in thermoformed talc-filled thermoplastics, Polym. Eng. Sci 36 (1996) 2188-2197. https://doi.org/10.1002/pen.10616

[17] B. C. Mutsuddy, R. G. Ford, Ceramic Injection Molding, London: Chapman & Hall, 1995.

[18] R. Tran, F. Reuther, S. Winter, V. Pysk, Process Development for a Superplastic Hot Tube Gas Forming Process of Titanium (Ti-3Al-2.5V) Hollow Profiles, Metals 10, (2020). https://doi.org/10.3390/met10091150

[19] A. Franck, K. Biederbrick, Kunststoff-Kompendium, Würzburg: Vogel-Fachbuch: Technik, 1990.
# Virtual inline compensation by single point-tracking in free-form bending

SCANDOLA Lorenzo¹,ᵃ*, BÖHM Viktor¹,ᵇ, MAIER Daniel¹,ᶜ and VOLK Wolfram¹,ᵈ

¹Chair of Metal Forming and Casting, TUM School of Engineering and Design, Technical University of Munich, Walther-Meissner-Str. 4, 85748 Garching, Germany

ᵃlorenzo.scandola@tum.de, ᵇviktor.boehm@tum.de, ᶜdaniel.maier@tum.de, ᵈwolfram.volk@tum.de

**Keywords:** Point-Tracking, Process Monitoring, Inline Compensation

**Abstract.** Free-form bending represents an attractive process for the manufacturing of structural components. Due to its kinematics-based working principle, a variety of geometries can be realised in the 3D-space featuring different radii, angles and bending planes. Nevertheless, the kinematic design of the bending head movement represents a complex task, and in order to obtain a part in the desired tolerance range a time-consuming trial and error procedure is the actual state of the art. The part is measured offline and compared to the target geometry after the bending operation, and the analysis of deviations is carried out to derive the compensated bending program. This is not only inefficient, but does not allow for compensating geometrical errors while the process runs. In this paper, a virtual inline compensation strategy based on single point-tracking is proposed. First, the process is modelled virtually and a single point on the component is tracked in space. Subsequently, the obtained tracking signals are presented and explained. Successively a kinematics adaption concept is shown and verified. On a virtual level the results show that the tracking of a single point on the geometry during the process can be used to compensate geometrical deviations in free-form bending.

**Introduction**

In the framework of bending processes free-form bending offers a unmatched flexibility regarding the complexity of shapes and geometry that can be obtained. This is of particular interest especially for the automotive sector, where the challenges arising in the e-mobility are setting always more constraints on the design of the chassis to accommodate batteries and additional equipment [1]. In particular concerning the structural components, non-standard parts are required, which can be the result of a lengthy prototypical phase. In such a scenario free-form bending can be the process of choice, as it allows for a fast and relatively low-impact investment manufacturing of components with varying bending features.

The first free-form bending machine has been patented by Murata [2] and belongs to the class of the kinematics-based processes [3]. This confers to the process a virtually infinite freedom of design, as the final geometry is not imposed by the contour of the tool [4]. Independently on the target radius in fact, only a single tool setup is required, designed on the cross-section of the target semi-finished part [5]. This represents a noticeable difference with respect to conventional bending processes, such as rotary-draw-bending [6], which nevertheless introduces a higher complexity in the design of the tool kinematics.

The actual workflow for obtaining a part in the required tolerance range with free-form bending is characterised by a lengthy trial-and-error procedure. In order to react to a deviation from the target geometry, a part must be produced with a trial kinematics, the deviations must be assessed, and basing on this information the new kinematics can be derived. This requires to stop the process and perform an offline measurement of the part, which can eventually lead to long dead times. In order to ensure a robust and defects-free manufacturing of free-form bent parts, an inline

<figure><img src="image_2324.png" /></figure>
# Virtual inline compensation by single point-tracking in free-form bending

SCANDOLA Lorenzo¹,ᵃ*, BÖHM Viktor¹,ᵇ, MAIER Daniel¹,ᶜ and VOLK Wolfram¹,ᵈ

¹Chair of Metal Forming and Casting, TUM School of Engineering and Design, Technical University of Munich, Walther-Meissner-Str. 4, 85748 Garching, Germany

ᵃlorenzo.scandola@tum.de, ᵇviktor.boehm@tum.de, ᶜdaniel.maier@tum.de, ᵈwolfram.volk@tum.de

**Keywords:** Point-Tracking, Process Monitoring, Inline Compensation

**Abstract.** Free-form bending represents an attractive process for the manufacturing of structural components. Due to its kinematics-based working principle, a variety of geometries can be realised in the 3D-space featuring different radii, angles and bending planes. Nevertheless, the kinematic design of the bending head movement represents a complex task, and in order to obtain a part in the desired tolerance range a time-consuming trial and error procedure is the actual state of the art. The part is measured offline and compared to the target geometry after the bending operation, and the analysis of deviations is carried out to derive the compensated bending program. This is not only inefficient, but does not allow for compensating geometrical errors while the process runs. In this paper, a virtual inline compensation strategy based on single point-tracking is proposed. First, the process is modelled virtually and a single point on the component is tracked in space. Subsequently, the obtained tracking signals are presented and explained. Successively a kinematics adaption concept is shown and verified. On a virtual level the results show that the tracking of a single point on the geometry during the process can be used to compensate geometrical deviations in free-form bending.

**Introduction**

In the framework of bending processes free-form bending offers a unmatched flexibility regarding the complexity of shapes and geometry that can be obtained. This is of particular interest especially for the automotive sector, where the challenges arising in the e-mobility are setting always more constraints on the design of the chassis to accommodate batteries and additional equipment [1]. In particular concerning the structural components, non-standard parts are required, which can be the result of a lengthy prototypical phase. In such a scenario free-form bending can be the process of choice, as it allows for a fast and relatively low-impact investment manufacturing of components with varying bending features.

The first free-form bending machine has been patented by Murata [2] and belongs to the class of the kinematics-based processes [3]. This confers to the process a virtually infinite freedom of design, as the final geometry is not imposed by the contour of the tool [4]. Independently on the target radius in fact, only a single tool setup is required, designed on the cross-section of the target semi-finished part [5]. This represents a noticeable difference with respect to conventional bending processes, such as rotary-draw-bending [6], which nevertheless introduces a higher complexity in the design of the tool kinematics.

The actual workflow for obtaining a part in the required tolerance range with free-form bending is characterised by a lengthy trial-and-error procedure. In order to react to a deviation from the target geometry, a part must be produced with a trial kinematics, the deviations must be assessed, and basing on this information the new kinematics can be derived. This requires to stop the process and perform an offline measurement of the part, which can eventually lead to long dead times. In order to ensure a robust and defects-free manufacturing of free-form bent parts, an inline

<figure><img src="image_2315.png" /></figure>
measurement system as well as compensation strategy is clearly identified as the most promising approach, as already demonstrated for other processes such as three-roll push bending [7].

In this study, an inline measurement concept as well as a corresponding compensation strategy is proposed. The investigations are performed virtually on a virtual model of the process, and serve as basis for the future experimental verification of the developed framework. First of all, the digital model of the process is presented and validated, by performing a sensitivity analysis of numerical parameters. Successively, the inline measurement system based on a single point-tracking is described, and the obtained signals are commented. In addition, the sensitivity of the results to the variation of kinematics parameters, such as the target position of the bending head and the feed time, is investigated. Finally, the inline compensation strategy adapting the kinematics based on the inline signals is proposed, and the results are shown.

The investigations aim at showing the feasibility of an inline compensation method for free-form bending, based on a simple and easily implementable inline geometry assessment.

## Virtual representation of the free-form bending process

Free-form bending numerical model. The proposed investigations are carried out virtually, making use of a finite element model of the process. The equivalent reduced assembly, consisting of the semi-finished part and of the surfaces of the tools for which contact is established, is shown in Figure 1. The tube is modeled making use of 4-nodes shell elements, featuring reduced integration and hourglass control. In order to reduce the computational time, 5 integration points to the thickness are used, as suggested by Gantner [8]. Due to the small ratio between thickness and diameter, $t/D \approx 0.06$, shell elements are preferred over continuum elements. This allows for lower computational time without loss of accuracy. The bending head as well as the fixed holder and the guides are discretised with 4-nodes rigid elements and set up as rigid bodies. The mandrel is composed of 8 balls and a shaft. The elements are connected through MPC with link behaviour [9]. To reproduce the condition of a highly lubricated process the friction coefficient for every contact is set to 0.04 [10]. For the material, the mild steel P235 is taken, and the Young's modulus as well as Poisson's coefficient are set respectively to 210 GPa and 0.3. An experimental flow curve and isotropic hardening are used.

<figure><img src="image_2325.png" /><figcaption>Figure 1: Virtual model of the free-form bending process</figcaption></figure>

Validation of the numerical model. In order to validate the numerical model, different simulations are performed with the aim of assessing the impact on the results. The analysis is performed in Abaqus [11] with the explicit solver. In order to reduce the simulation time, the time scaling strategy is applied. This is preferred over the mass scaling for the free-form bending simulation, as the latter could modify the dynamic behaviour of the part. For this reason, the chosen
measurement system as well as compensation strategy is clearly identified as the most promising approach, as already demonstrated for other processes such as three-roll push bending [7].

In this study, an inline measurement concept as well as a corresponding compensation strategy is proposed. The investigations are performed virtually on a virtual model of the process, and serve as basis for the future experimental verification of the developed framework. First of all, the digital model of the process is presented and validated, by performing a sensitivity analysis of numerical parameters. Successively, the inline measurement system based on a single point-tracking is described, and the obtained signals are commented. In addition, the sensitivity of the results to the variation of kinematics parameters, such as the target position of the bending head and the feed time, is investigated. Finally, the inline compensation strategy adapting the kinematics based on the inline signals is proposed, and the results are shown.

The investigations aim at showing the feasibility of an inline compensation method for free-form bending, based on a simple and easily implementable inline geometry assessment.

## Virtual representation of the free-form bending process

Free-form bending numerical model. The proposed investigations are carried out virtually, making use of a finite element model of the process. The equivalent reduced assembly, consisting of the semi-finished part and of the surfaces of the tools for which contact is established, is shown in Figure 1. The tube is modeled making use of 4-nodes shell elements, featuring reduced integration and hourglass control. In order to reduce the computational time, 5 integration points to the thickness are used, as suggested by Gantner [8]. Due to the small ratio between thickness and diameter, $t/D \approx 0.06$, shell elements are preferred over continuum elements. This allows for lower computational time without loss of accuracy. The bending head as well as the fixed holder and the guides are discretised with 4-nodes rigid elements and set up as rigid bodies. The mandrel is composed of 8 balls and a shaft. The elements are connected through MPC with link behaviour [9]. To reproduce the condition of a highly lubricated process the friction coefficient for every contact is set to 0.04 [10]. For the material, the mild steel P235 is taken, and the Young's modulus as well as Poisson's coefficient are set respectively to 210 GPa and 0.3. An experimental flow curve and isotropic hardening are used.

<figure><img src="image_2316.png" /><figcaption>Figure 1: Virtual model of the free-form bending process</figcaption></figure>

Validation of the numerical model. In order to validate the numerical model, different simulations are performed with the aim of assessing the impact on the results. The analysis is performed in Abaqus [11] with the explicit solver. In order to reduce the simulation time, the time scaling strategy is applied. This is preferred over the mass scaling for the free-form bending simulation, as the latter could modify the dynamic behaviour of the part. For this reason, the chosen
numerical parameters for the validation of the model are the time scaling factor and the element size, expressed as number of elements along the tube's circumference. As validation quantity is taken the resulting bending angle. The parameters and results are listed in Table 1.

<table><caption>Table 1: Investigated parameter sets for the validation of the model</caption><thead><tr><td>Time scaling factor</td><td>Elements along circumf.</td><td>Bending angle in deg</td></tr></thead><tbody><tr><td>300</td><td>30</td><td>96.6</td></tr><tr><td>190</td><td>30</td><td>88.4</td></tr><tr><td>50</td><td>30</td><td>87.9</td></tr><tr><td>5</td><td>30</td><td>88.6</td></tr><tr><td>190</td><td>25</td><td>86.6</td></tr><tr><td>190</td><td>35</td><td>86.9</td></tr><tr><td>190</td><td>40</td><td>86.4</td></tr></tbody></table>

The resulting computational times are plotted against the numerical parameters in Figure 2. Regarding the time scaling factor and a fixed number of elements along the circumference of 30, it can be seen that except for the case of very high scaling factors, such as 300, the difference in the bending angle is between 1 deg. On the other side, the computational time varies substantially, from about 10 minutes for the case 300 to 14 hours for the case 5. With respect to the number of elements along the circumference with fixed time scaling factor of 190, the angle oscillates around 0.1 deg, whereas the resulting computational time also increases exponentially, from 7 minutes for the case of 25 to almost 1.5 hours for the case of 40.

<figure><img src="image_2326.png" /><figcaption>Figure 2: Elapsed computational time for a free-form bending FEM simulation as a function of the time scaling factor (left) and of the number of elements along the circumference</figcaption></figure>

For this reason, as standard value for all the following virtual tests, a time scaling factor of 190 and a number of element along the circumference of 30 have been chosen. This allows a computational time lower than 30 minutes, represented as dashed line in Figure 2. Although the results sensitivity to the time scaling is one order of magnitude bigger than that of the mesh size, the benefit in the computational time is such, that the achieved accuracy has been reputed sufficient for the scope of this work. For a more precise simulation, other factors such as the elasticity of the tools, the inertia of the bending head and advanced material modelling for the welding joint and the core material should be addressed.
numerical parameters for the validation of the model are the time scaling factor and the element size, expressed as number of elements along the tube's circumference. As validation quantity is taken the resulting bending angle. The parameters and results are listed in Table 1.

<table><caption>Table 1: Investigated parameter sets for the validation of the model</caption><thead><tr><td>Time scaling factor</td><td>Elements along circumf.</td><td>Bending angle in deg</td></tr></thead><tbody><tr><td>300</td><td>30</td><td>96.6</td></tr><tr><td>190</td><td>30</td><td>88.4</td></tr><tr><td>50</td><td>30</td><td>87.9</td></tr><tr><td>5</td><td>30</td><td>88.6</td></tr><tr><td>190</td><td>25</td><td>86.6</td></tr><tr><td>190</td><td>35</td><td>86.9</td></tr><tr><td>190</td><td>40</td><td>86.4</td></tr></tbody></table>

The resulting computational times are plotted against the numerical parameters in Figure 2. Regarding the time scaling factor and a fixed number of elements along the circumference of 30, it can be seen that except for the case of very high scaling factors, such as 300, the difference in the bending angle is between 1 deg. On the other side, the computational time varies substantially, from about 10 minutes for the case 300 to 14 hours for the case 5. With respect to the number of elements along the circumference with fixed time scaling factor of 190, the angle oscillates around 0.1 deg, whereas the resulting computational time also increases exponentially, from 7 minutes for the case of 25 to almost 1.5 hours for the case of 40.

<figure><img src="image_2317.png" /><figcaption>Figure 2: Elapsed computational time for a free-form bending FEM simulation as a function of the time scaling factor (left) and of the number of elements along the circumference</figcaption></figure>

For this reason, as standard value for all the following virtual tests, a time scaling factor of 190 and a number of element along the circumference of 30 have been chosen. This allows a computational time lower than 30 minutes, represented as dashed line in Figure 2. Although the results sensitivity to the time scaling is one order of magnitude bigger than that of the mesh size, the benefit in the computational time is such, that the achieved accuracy has been reputed sufficient for the scope of this work. For a more precise simulation, other factors such as the elasticity of the tools, the inertia of the bending head and advanced material modelling for the welding joint and the core material should be addressed.
Concept of the single point-tracking inline measurement method

Resulting ground signals. The concept for the inline measurement technique is based on the tracking of a single point on the semi-finished part. The tracked point is chosen to be at the starting edge of the tube in the middle of the diameter. The given kinematics of the bending head is shown along with the position of the point during the bending in Figure 3. Four characteristic points can be identified during the bending. The starting configuration can be observed at the time $t_0$. From there, the bending head reaches linearly the target position in the time $t_1$, where the cantilever beam effect can be observed [12]. Here the constant radius is realised, and this position is held constantly until the time $t_2$, where the bending head starts to return to the initial position, which is reached at the time $t_3$.

<figure><img src="image_2327.png" /><figcaption>Figure 3: Kinematic path of the bending head and relative position of the tracked point</figcaption></figure>

The inline signals of the tracked point acquired during the process are expressed as the variation of its coordinates during time, with respect to the fixed reference system lying on the point at the beginning of the process, represented in red in Figure 3. These include the x, y and z coordinates of the point as well as its trajectory, which are shown in Figure 4.

It can be seen, that the x coordinate (top-right) remains constant as expected and does not yield any information about the bending.

On the other side, the y coordinate shows an evident bi-linear evolution in time, (top-left). After ≈ 1 s the slope of the curve diminishes clearly. This confirms the *cantilever-beam effect* behaviour of the tube at the beginning of the bending [12]. In addition, after reaching a maximum at around ≈ 5 s, the y coordinate of the tracked point decreases, which can be interpreted as a measure of the springback or of an additional bending effect imposed by the bending head to the part when returning to its initial position.
Concept of the single point-tracking inline measurement method

Resulting ground signals. The concept for the inline measurement technique is based on the tracking of a single point on the semi-finished part. The tracked point is chosen to be at the starting edge of the tube in the middle of the diameter. The given kinematics of the bending head is shown along with the position of the point during the bending in Figure 3. Four characteristic points can be identified during the bending. The starting configuration can be observed at the time $t_0$. From there, the bending head reaches linearly the target position in the time $t_1$, where the cantilever beam effect can be observed [12]. Here the constant radius is realised, and this position is held constantly until the time $t_2$, where the bending head starts to return to the initial position, which is reached at the time $t_3$.

<figure><img src="image_2318.png" /><figcaption>Figure 3: Kinematic path of the bending head and relative position of the tracked point</figcaption></figure>

The inline signals of the tracked point acquired during the process are expressed as the variation of its coordinates during time, with respect to the fixed reference system lying on the point at the beginning of the process, represented in red in Figure 3. These include the x, y and z coordinates of the point as well as its trajectory, which are shown in Figure 4.

It can be seen, that the x coordinate (top-right) remains constant as expected and does not yield any information about the bending.

On the other side, the y coordinate shows an evident bi-linear evolution in time, (top-left). After ≈ 1 s the slope of the curve diminishes clearly. This confirms the *cantilever-beam effect* behaviour of the tube at the beginning of the bending [12]. In addition, after reaching a maximum at around ≈ 5 s, the y coordinate of the tracked point decreases, which can be interpreted as a measure of the springback or of an additional bending effect imposed by the bending head to the part when returning to its initial position.
<figure><img src="image_2328.png" /><figcaption>Figure 4: Resulting signals of the tracked point: y coordinate (top-left), x coordinate (top-right), z coordinate (bottom-left), y-z trajectory (bottom-right)</figcaption></figure>

A characteristic signal shape can be also recognised in the z coordinate, (bottom-left). The signal increases first linearly. After reaching a maximum, an inversion in the z coordinate is recorded, since during the bending the tip of the tube is displaced to the opposite direction with respect to the feed. Finally, when the bending head gets back to the initial position the z coordinate increases again.

In addition to the analysis of the evolution of the coordinates of the tracked point in time, its y-z trajectory can also be reconstructed (bottom-right). Again, a clear bi-linear behaviour can be identified when the bending head reaches its target position. Moreover, the final linear decreasing part of the signal can be used as an indication regarding the decrease of the target bending angle.

Sensitivity analysis to variation of kinematics parameters. In order to investigate the sensitivity of the described tracking signals with respect to the free-form bending process, different tests are performed. As kinematics parameters the target position of the bending head *u* as well as the total bending time *t* are chosen. The tested configurations are summarised in Table 2.

<table><caption>Table 2: Target position and total time configurations for the sensitivity analysis</caption><thead><tr><th>Label</th><th>u 7</th><th>u 9</th><th>u 11</th><th>u 13</th><th>u 15</th><th>t 3</th><th>t 4</th><th>t 5</th><th>t 6</th><th>t 7</th></tr></thead><tbody><tr><td>Total time <i>t</i> in s</td><td>5</td><td>5</td><td>5</td><td>5</td><td>5</td><td>3</td><td>4</td><td>5</td><td>6</td><td>7</td></tr><tr><td>Target position <i>u</i> in mm</td><td>7</td><td>9</td><td>11</td><td>13</td><td>15</td><td>11</td><td>11</td><td>11</td><td>11</td><td>11</td></tr></tbody></table>

For the interpretation of the results, the final free-form bent geometry is compared to the obtained tracking signals. The former is described using the resulting bending radius and angle,
<figure><img src="image_2319.png" /><figcaption>Figure 4: Resulting signals of the tracked point: y coordinate (top-left), x coordinate (top-right), z coordinate (bottom-left), y-z trajectory (bottom-right)</figcaption></figure>

A characteristic signal shape can be also recognised in the z coordinate, (bottom-left). The signal increases first linearly. After reaching a maximum, an inversion in the z coordinate is recorded, since during the bending the tip of the tube is displaced to the opposite direction with respect to the feed. Finally, when the bending head gets back to the initial position the z coordinate increases again.

In addition to the analysis of the evolution of the coordinates of the tracked point in time, its y-z trajectory can also be reconstructed (bottom-right). Again, a clear bi-linear behaviour can be identified when the bending head reaches its target position. Moreover, the final linear decreasing part of the signal can be used as an indication regarding the decrease of the target bending angle.

Sensitivity analysis to variation of kinematics parameters. In order to investigate the sensitivity of the described tracking signals with respect to the free-form bending process, different tests are performed. As kinematics parameters the target position of the bending head *u* as well as the total bending time *t* are chosen. The tested configurations are summarised in Table 2.

<table><caption>Table 2: Target position and total time configurations for the sensitivity analysis</caption><thead><tr><th>Label</th><th>u 7</th><th>u 9</th><th>u 11</th><th>u 13</th><th>u 15</th><th>t 3</th><th>t 4</th><th>t 5</th><th>t 6</th><th>t 7</th></tr></thead><tbody><tr><td>Total time <i>t</i> in s</td><td>5</td><td>5</td><td>5</td><td>5</td><td>5</td><td>3</td><td>4</td><td>5</td><td>6</td><td>7</td></tr><tr><td>Target position <i>u</i> in mm</td><td>7</td><td>9</td><td>11</td><td>13</td><td>15</td><td>11</td><td>11</td><td>11</td><td>11</td><td>11</td></tr></tbody></table>

For the interpretation of the results, the final free-form bent geometry is compared to the obtained tracking signals. The former is described using the resulting bending radius and angle,
while for the latter only the y coordinate is analysed. Although also the z coordinate and the trajectory yield similar information, see Figure 4, for a 2D-bending in the vertical direction the y coordinate allows for the simplest data interpretation. The obtained tracking signals of the y coordinate for different target positions and total time are shown in Figure 5.

<figure><img src="image_2329.png" /><figcaption>Figure 5: Y coordinate of the tracked point for different total time (left) and target position (right)</figcaption></figure>

It can be seen that an increase in the total time leads to an increase in the value of the y coordinate of the tracked point, as it can be recognized in Figure 5 left, where the total time is shown in black dashed lines. In addition, bigger values of the target position of the bending head induce a steeper signal of the y coordinate, which is represented by the linear fit on the curves in black dashed lines in Figure 5 right. To investigate if the inline signals are suitable to analyse the free-form bending process, the resulting bending parameters are plotted against the inline data in Figure 6.

<figure><img src="image_2330.png" /><figcaption>Figure 6: Correlation between inline y coordinate signal with bending angle (left) and bending radius (right)</figcaption></figure>

The resulting bending angle is shown against the value of the y coordinate at the end of the bending in Figure 6 left. It can be seen that a linear correlation between the inline signal and the bending angle arises. This is reasonable, since increasing the total time of the bending results in a proportional increase of the achieved angle. The investigations confirm that the phenomenon is linear, with a Pearson Correlation Coefficient of $R^2=0.994$.

To investigate the impact of the inline measurement on the obtained bending radius, a line fit is first applied to the linear part of the y coordinate signals in Figure 5 right (dashed black lines), resulting in the first derivative parameter $dY/dt$. Plotting the bending radius against this parameter also yield a linear correlation with $R^2=0.997$, shown in Figure 6 right. This means that the slope of
while for the latter only the y coordinate is analysed. Although also the z coordinate and the trajectory yield similar information, see Figure 4, for a 2D-bending in the vertical direction the y coordinate allows for the simplest data interpretation. The obtained tracking signals of the y coordinate for different target positions and total time are shown in Figure 5.

<figure><img src="image_2320.png" /><figcaption>Figure 5: Y coordinate of the tracked point for different total time (left) and target position (right)</figcaption></figure>

It can be seen that an increase in the total time leads to an increase in the value of the y coordinate of the tracked point, as it can be recognized in Figure 5 left, where the total time is shown in black dashed lines. In addition, bigger values of the target position of the bending head induce a steeper signal of the y coordinate, which is represented by the linear fit on the curves in black dashed lines in Figure 5 right. To investigate if the inline signals are suitable to analyse the free-form bending process, the resulting bending parameters are plotted against the inline data in Figure 6.

<figure><img src="image_2321.png" /><figcaption>Figure 6: Correlation between inline y coordinate signal with bending angle (left) and bending radius (right)</figcaption></figure>

The resulting bending angle is shown against the value of the y coordinate at the end of the bending in Figure 6 left. It can be seen that a linear correlation between the inline signal and the bending angle arises. This is reasonable, since increasing the total time of the bending results in a proportional increase of the achieved angle. The investigations confirm that the phenomenon is linear, with a Pearson Correlation Coefficient of $R^2=0.994$.

To investigate the impact of the inline measurement on the obtained bending radius, a line fit is first applied to the linear part of the y coordinate signals in Figure 5 right (dashed black lines), resulting in the first derivative parameter $dY/dt$. Plotting the bending radius against this parameter also yield a linear correlation with $R^2=0.997$, shown in Figure 6 right. This means that the slope of
the y coordinate signal in time correlates linearly with the obtained bending radius, and can hence be used as an indirect measurement of the achieved radius in real time.

The investigations confirmed that the inline tracking of a single point on the bending part can be exploited to assess the values of bending features such as the bending radius and bending angle during the process. This can be used to realise an inline compensation for free-form bent parts.

**Inline compensation strategy**

In order to show the potential of the inline tracking measurement for process monitoring and compensation, a starting geometry obtained by a first-guess kinematics, *init* in Figure 7, is used as a starting point to achieve the target geometry, *tgt*. Two approaches are presented, namely the adaption of the process time only, as well as the correction of target position and total time simultaneously.

<figure><img src="image_2331.png" /><figcaption>Figure 7: Initial and target y coordinate tracking signal (left) and resulting geometries (right)</figcaption></figure>

Compensation by time only. The first approach to compensate the part is to modify the total process time only. Since the initial guess results in a much lower bending angle than the target, Figure 7 right, the total time is increased in order to reach a higher angle. In particular, three strategies are tested. In S1A the time when the matrix begins to return to the initial position, $t_2$ in Figure 3, is set as the time where the maximum of the y coordinate signal of the target part is achieved, *A* in Figure 7 right. On the other hand the kinematics S1B is achieved by setting $t_2$ as the time at which the y coordinate signal ends, *B* in Figure 7 right. The results are shown in Figure 8, where S1C represents the configuration showing the same bending angle as the target, obtained a-posteriori.
the y coordinate signal in time correlates linearly with the obtained bending radius, and can hence be used as an indirect measurement of the achieved radius in real time.

The investigations confirmed that the inline tracking of a single point on the bending part can be exploited to assess the values of bending features such as the bending radius and bending angle during the process. This can be used to realise an inline compensation for free-form bent parts.

**Inline compensation strategy**

In order to show the potential of the inline tracking measurement for process monitoring and compensation, a starting geometry obtained by a first-guess kinematics, *init* in Figure 7, is used as a starting point to achieve the target geometry, *tgt*. Two approaches are presented, namely the adaption of the process time only, as well as the correction of target position and total time simultaneously.

<figure><img src="image_2322.png" /><figcaption>Figure 7: Initial and target y coordinate tracking signal (left) and resulting geometries (right)</figcaption></figure>

Compensation by time only. The first approach to compensate the part is to modify the total process time only. Since the initial guess results in a much lower bending angle than the target, Figure 7 right, the total time is increased in order to reach a higher angle. In particular, three strategies are tested. In S1A the time when the matrix begins to return to the initial position, $t_2$ in Figure 3, is set as the time where the maximum of the y coordinate signal of the target part is achieved, *A* in Figure 7 right. On the other hand the kinematics S1B is achieved by setting $t_2$ as the time at which the y coordinate signal ends, *B* in Figure 7 right. The results are shown in Figure 8, where S1C represents the configuration showing the same bending angle as the target, obtained a-posteriori.
<figure><img src="image_2332.png" /><figcaption>Figure 8: Results of the compensation by adaption of total time: y coordinate tracking signals (left) and resulting geometries (right)</figcaption></figure>

It can be observed, that the modification of the total time of the process cannot yield a proper compensated geometry if the target position of the bending head is not correct. This is particularly clear in Figure 8 left. Although the y coordinate of S1A is very close to the target T=88.4 deg, the resulting bending angle is reaching only T=78.1 deg. In addition, the closest angle of T=88.9 deg can be achieved with the strategy S1C, which results in a much higher y coordinate. This demonstrates that this method can only be employed successfully when the target bending head position is already correct, and the only geometrical error is represented by the bending angle, making use of linear correlation as shown in Figure 5 left.

Compensation by time and target position. In order to identify the correct kinematics able to yield the target geometry, the inline compensation strategy can be used to first modify the target position of the bending head and then to adapt the total process time. To demonstrate this, two additional kinematics starting from an incorrect target position are tested and modified until the signal of the correct target position is visualised in the y coordinate plot. In the kinematic profile S2inc the bending head reaches a too little target position, and is further displaced after 0.8 s until the y coordinate signal value matches the target tgt. The test S2dec represents the opposite path, namely the bending head reaching a too high target position, before being displaced to the lower values until the y coordinate signals match after 0.8 s. The results are shown in Figure 9, referring to the y coordinate signals (left) and the resulting geometries (right). It can be seen that the initial starting geometry could be compensated inline, only comparing the target and obtained tracking signals of a single point. First, the target position of the bending head can be successfully achieved by analyzing the y coordinate signal, Figure 9. Both S2inc, yellow, and S2dec, red, show a y coordinate value respectively lower and higher than the target at the beginning, which converge to the target after displacing the bending head in the proper direction. This results in a radius compensation, with a remaining deviation of 12.0 mm for S2inc and 1.4 mm for S2dec, clearly lower than the one of the initial geometry of 41.3 mm.

At this point, the bending can be held until the actual y coordinate signal matches the target one in time, as discussed previously. In such a way, the angle compensation based on the total process time can be carried out, resulting in deviations of 4.4 deg for S2inc and 0.01 deg for S2dec, and representing a relevant enhancement with respect to the initial geometry deviation of 30.5 deg.
<figure><img src="image_2323.png" /><figcaption>Figure 8: Results of the compensation by adaption of total time: y coordinate tracking signals (left) and resulting geometries (right)</figcaption></figure>

It can be observed, that the modification of the total time of the process cannot yield a proper compensated geometry if the target position of the bending head is not correct. This is particularly clear in Figure 8 left. Although the y coordinate of S1A is very close to the target T=88.4 deg, the resulting bending angle is reaching only T=78.1 deg. In addition, the closest angle of T=88.9 deg can be achieved with the strategy S1C, which results in a much higher y coordinate. This demonstrates that this method can only be employed successfully when the target bending head position is already correct, and the only geometrical error is represented by the bending angle, making use of linear correlation as shown in Figure 5 left.

Compensation by time and target position. In order to identify the correct kinematics able to yield the target geometry, the inline compensation strategy can be used to first modify the target position of the bending head and then to adapt the total process time. To demonstrate this, two additional kinematics starting from an incorrect target position are tested and modified until the signal of the correct target position is visualised in the y coordinate plot. In the kinematic profile S2inc the bending head reaches a too little target position, and is further displaced after 0.8 s until the y coordinate signal value matches the target tgt. The test S2dec represents the opposite path, namely the bending head reaching a too high target position, before being displaced to the lower values until the y coordinate signals match after 0.8 s. The results are shown in Figure 9, referring to the y coordinate signals (left) and the resulting geometries (right). It can be seen that the initial starting geometry could be compensated inline, only comparing the target and obtained tracking signals of a single point. First, the target position of the bending head can be successfully achieved by analyzing the y coordinate signal, Figure 9. Both S2inc, yellow, and S2dec, red, show a y coordinate value respectively lower and higher than the target at the beginning, which converge to the target after displacing the bending head in the proper direction. This results in a radius compensation, with a remaining deviation of 12.0 mm for S2inc and 1.4 mm for S2dec, clearly lower than the one of the initial geometry of 41.3 mm.

At this point, the bending can be held until the actual y coordinate signal matches the target one in time, as discussed previously. In such a way, the angle compensation based on the total process time can be carried out, resulting in deviations of 4.4 deg for S2inc and 0.01 deg for S2dec, and representing a relevant enhancement with respect to the initial geometry deviation of 30.5 deg.
<figure><img src="image_2333.png" /><figcaption>Figure 9: Results of the compensation by adaption of target position and total time: y coordinate tracking signals (left) and resulting geometries (right)</figcaption></figure>

## Conclusions

This paper addressed the topic of inline geometry acquisition and compensation in free-form bending and proposed a novel strategy, based on single point-tracking. The cumbersome trial-and-error approach to obtain parts in the desired tolerance range represents the bottle-neck for the diffusion of the process and is mainly based on offline geometry measurements. Although other approaches, such as inline bending features monitoring and laser tracking have been tested, the problem is still unresolved. In this scenario, the single point-tracking has been developed here and validated by exploiting the virtual model of the process. First, the FE model of the process has been implemented and verified by assessing the impact of the time scaling factor parameter and the mesh size. Successively, the concept of the single point-tracking has been presented, and the ground inline signals referring to the *y*, *x*, *z* coordinate as well as the point trajectory have been described. The *y* coordinate signal has been chosen as the most representative, and a sensitivity analysis concerning the variation of the total process time and of the target position of the bending head has been performed. The results showed a linear correlation between the bending angle and the absolute *y* coordinate value, as well as between the bending radius and the first derivative of the *y* coordinate in time, with Pearson coefficients higher than 0.99. Finally, this information has been used to propose an inline compensation strategy. It could be demonstrated, that in order to perform a full geometry compensation both the target position of the bending head as well as the total process time must be adapted. The *y* coordinate signal can be used for a 2D-bending to identify the correct value of the target position of bending head, and the total process time can be exploited to correct the angle, provided that the target geometry and associated signals are available. The results allowed to enhance the starting radius and angle deviations of 41.3 *mm* and 30.5 *deg* to the values of 1.4 *mm* and 0.01 *deg*.

Future work will concern the experimental validation of these investigations, and the comparison with other approaches, such as laser-based geometry acquisitions methods an inline assessment of bending features.

## References

[1] P. Gantner, H. Bauer, D.K. Harrison, and A.K. de Silva, Free-Bending-A new bending technique in the hydroforming process chain, J. Mater. Process. Technol. 167 (2005) 302–308.
https://doi.org/10.1016/j.jmatprotec.2005.05.052
<figure><img src="image_2324.png" /><figcaption>Figure 9: Results of the compensation by adaption of target position and total time: y coordinate tracking signals (left) and resulting geometries (right)</figcaption></figure>

## Conclusions

This paper addressed the topic of inline geometry acquisition and compensation in free-form bending and proposed a novel strategy, based on single point-tracking. The cumbersome trial-and-error approach to obtain parts in the desired tolerance range represents the bottle-neck for the diffusion of the process and is mainly based on offline geometry measurements. Although other approaches, such as inline bending features monitoring and laser tracking have been tested, the problem is still unresolved. In this scenario, the single point-tracking has been developed here and validated by exploiting the virtual model of the process. First, the FE model of the process has been implemented and verified by assessing the impact of the time scaling factor parameter and the mesh size. Successively, the concept of the single point-tracking has been presented, and the ground inline signals referring to the *y*, *x*, *z* coordinate as well as the point trajectory have been described. The *y* coordinate signal has been chosen as the most representative, and a sensitivity analysis concerning the variation of the total process time and of the target position of the bending head has been performed. The results showed a linear correlation between the bending angle and the absolute *y* coordinate value, as well as between the bending radius and the first derivative of the *y* coordinate in time, with Pearson coefficients higher than 0.99. Finally, this information has been used to propose an inline compensation strategy. It could be demonstrated, that in order to perform a full geometry compensation both the target position of the bending head as well as the total process time must be adapted. The *y* coordinate signal can be used for a 2D-bending to identify the correct value of the target position of bending head, and the total process time can be exploited to correct the angle, provided that the target geometry and associated signals are available. The results allowed to enhance the starting radius and angle deviations of 41.3 *mm* and 30.5 *deg* to the values of 1.4 *mm* and 0.01 *deg*.

Future work will concern the experimental validation of these investigations, and the comparison with other approaches, such as laser-based geometry acquisitions methods an inline assessment of bending features.

## References

[1] P. Gantner, H. Bauer, D.K. Harrison, and A.K. de Silva, Free-Bending-A new bending technique in the hydroforming process chain, J. Mater. Process. Technol. 167 (2005) 302–308.
https://doi.org/10.1016/j.jmatprotec.2005.05.052
[2] M. Murata, Y. Uemura, and H. Suzuki, Adv. Tech. Plast. 3 (1990) 1573-1578.
https://doi.org/10.1299/kikaic.55.2488

[3] DIN Deutsches Institut für Normung e. V., DIN 8586: Fertigungsverfahren Biegeumformen, Berlin: Beuth Verlag GmbH (2003)

[4] J. Neu, 6-Achs Freiformbiegemaschine NSB on https://www.neu-gmbh.de/de/freiformbiegen-nsb

[5] M.K. Werner, L. Scandola, D. Maier, and W. Volk, Freeform Bending Tool Design for Rectangular Profiles and Its Influence on the Process, J. Manuf. Sci. Eng. 145 (2023).
https://doi.org/10.1115/1.4062811

[6] B. Engel and M. Hinkel, Analytisch unterstützte Vorauslegung des Rotationszugbiegeprozesses, Tagungsband / XXX. Verformungskundliches Kolloquium: Leoben, (2011) 97-102

[7] A. Ghiotti, E. Simonetto, S. Bruschi, and P.F. Bariani, Springback measurement in three roll push bending process of hollow structural sections, CIRP Ann. 66 (2017) 289-292.
https://doi.org/10.1016/j.cirp.2017.04.119

[8] P. Gantner, H. Bauer, D. K. Harrison, and A. K. M. de Silva, FEA - Simulation of Bending Processes with LS-DYNA, 8th International LS-DYNA Conference (2004)

[9] X. Guo et al., Free-bending process characteristics and forming process design of copper tubular components, J. Adv. Manuf. Technol. 96 (2018) 3585-3601.
https://doi.org/10.1007/s00170-018-1788-1

[10] N. Beulich, P. Craighero and W. Volk, FEA Simulation of Free-Bending - a Preforming Step in the Hydroforming Process Chain, J. Phys. Conf. Ser. 896 (2017),
https://doi.org/10.1088/1742-6596/896/1/012063

[11] Dassault Systèmes Simulia Corp, ABAQUS/Standard User's Manual (2023)

[12] L. Scandola, J. Tschannerl, D. Maier, M.K. Werner, C. Hartmann, and W. Volk, Impact of the process velocities on the quality of free-form bent parts, IOP Conf. Ser. Mater. Sci. Eng., 1284 (2023). https://doi.org/10.1088/1757-899x/1284/1/012061
[2] M. Murata, Y. Uemura, and H. Suzuki, Adv. Tech. Plast. 3 (1990) 1573-1578.
https://doi.org/10.1299/kikaic.55.2488

[3] DIN Deutsches Institut für Normung e. V., DIN 8586: Fertigungsverfahren Biegeumformen, Berlin: Beuth Verlag GmbH (2003)

[4] J. Neu, 6-Achs Freiformbiegemaschine NSB on https://www.neu-gmbh.de/de/freiformbiegen-nsb

[5] M.K. Werner, L. Scandola, D. Maier, and W. Volk, Freeform Bending Tool Design for Rectangular Profiles and Its Influence on the Process, J. Manuf. Sci. Eng. 145 (2023).
https://doi.org/10.1115/1.4062811

[6] B. Engel and M. Hinkel, Analytisch unterstützte Vorauslegung des Rotationszugbiegeprozesses, Tagungsband / XXX. Verformungskundliches Kolloquium: Leoben, (2011) 97-102

[7] A. Ghiotti, E. Simonetto, S. Bruschi, and P.F. Bariani, Springback measurement in three roll push bending process of hollow structural sections, CIRP Ann. 66 (2017) 289-292.
https://doi.org/10.1016/j.cirp.2017.04.119

[8] P. Gantner, H. Bauer, D. K. Harrison, and A. K. M. de Silva, FEA - Simulation of Bending Processes with LS-DYNA, 8th International LS-DYNA Conference (2004)

[9] X. Guo et al., Free-bending process characteristics and forming process design of copper tubular components, J. Adv. Manuf. Technol. 96 (2018) 3585-3601.
https://doi.org/10.1007/s00170-018-1788-1

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# Nano-texturing orientation effect on the piercing behavior of five stacked amorphous electrical steel sheets

SHIRATORI Tomomi<fnref n="1" />,ᵃ*, ITO Kentaro<fnref n="2" />,ᵇ, FURUHATA Suguri<fnref n="2" />,ᶜ and
AIZAWA Tatsuhiko<fnref n="3" />,ᵈ

**Keywords:** Punching, Nano-Texturing, Amorphous Electrical Steel Sheets

**Abstract.** Higher efficiency motor becomes an expecting keyword in the sustainable society. The iron core in these motors is fabricated by piercing the amorphous electrical steel sheet with sufficient dimensional accuracy and without defects. However, this material has 2.2 GPa in the tensile strength, no ductility, and over 900 HV in hardness. Furthermore, tool life was limited to be short. In recent years, some new tooling technologies for piercing amorphous electrical steel sheets have been developed to reduce the induced damages into punched sheets. Especially, when using nano-textured tool, the process affected width was more reduced than using the non-textured tools. In the present study, nano-texturing orientation effect on the piercing behavior is investigated and its mechanism is discussed.

**Introduction**

The iron core of a motor has been processed by punching an electrical steel sheet. Non-oriented electrical steel sheets have been adapted to efficiently motor. An amorphous electrical steel sheet has been expected to the high efficiency motors. An amorphous electrical steel sheet has a 2.2 GPa in the tensile strength, no ductility, and over 900 HV in hardness. Thus, when punching amorphous electrical steel sheets, tool life was limited to be short. There are some piercing technologies has been developed in recent years. In the field of amorphous electrical steel sheets piercing, Chemical Vapor Deposition (CVD) diamond coated and adding nano-texturing new tool was developed [1] and the effect of reducing the process affected width was confirmed. A tool adding nano-texturing on SKD11 material has been developed [2]. It has been reported that changing the angle at which nano-texturing is applied reduces piercing load and process affected width. The effects of nanometer and micrometer textures on sheared surface formation were investigated [3]. This study showed that even if the direction of microgroove formation was changed by 90 degrees, the process affected width in surface of pierced hole in amorphous electrical steel sheets did not change much, and nano-texture became main factor of reduced the process affected width [3]. In the field of non-oriented electrical steel sheet piercing, reduction of adhesion was observed with CVD diamond and nano-textured punch [4]. These results indicate that nano-texture has the potential to improve the sliding properties during piercing process with the workpiece, contributing to increased tool durability. In addition to these studies, ion sharpening tools are also being developed [5]. It has been confirmed that the homogeneity of nanometer-sized tool tips has the effect of lowering the fracture surface ratio.

It has been found that nanotextures and tools with nanometer precision can suppress the process affected width during sheared surface formation and piercing process of amorphous electrical steel sheets. However, the optimization and mechanism of machining conditions such as

<figure><img src="image_2334.png" /></figure>

<footnote n="1">¹Faculty of engineering, University of Toyama, Toyama-city, Toyama 930-8555, Japan</footnote>

<footnote n="2">²Graduate School of Engineering, University of Toyama 930-8555, Japan</footnote>

<footnote n="3">³Surface Engineering Design Laboratory, SIT; Tokyo 144-0045, Japan</footnote>

<footnote>ᵃshira@eng.u-toyama.Fuac.jp, ᵇm23c1411 @ems.u-toyama.ac.jp,</footnote>

<footnote>ᶜs2279404@ems.u-toyama.ac.jp, ᵈtaizawa@sic.shibaura-it.ac.jp</footnote>