<figure><img src="image_2601.png" /><figcaption>Figure 1: Geometry of specimen used (a) and experimental setup for torsion fatigue tests</figcaption></figure>

Torsional fatigue strength tests. The torsional fatigue strength of the specimen was tested on a FLATO of Schenk at the Institute for Materials Applications in Mechanical Engineering IWM of RWTH Aachen University (see Figure 1 (b)). The tests were carried out in the finite life fatigue area with a torque ratio of $R_T = -1$. The test frequency was $f = 25$ Hz and the maximum number of load cycles was $N_m = 5 \cdot 10^6$.

Analysis of surface layer properties. To understand the cause effect relationship between the surface treatments and the resulting fatigue strength the roughness, the microhardness, the fracture pattern and the residual stresses within the cross bore were considered. The roughness of the cross bore surface was determined using a Hommel-Etamic nanoscan 855 tactile contour and roughness measuring device from JENOPTIK Industrial Metrology Germany GmbH. Measurements were taken in three areas of the cross bore, namely in the area of the chamfer on both sides and in the center of the cross bore. Three measurements were carried out in each area for statistical assurance. The roughness measurements were evaluated in accordance with DIN EN ISO 4288. The evaluation was based on the mean roughness depth $Rz$ and the arithmetic mean roughness $Ra$. The microhardness test was carried out in accordance with DIN EN ISO 14577-1 on a Fischerscope HM2000 from Helmut Fischer and converted into Vickers hardness. To determine the microhardness in the cross bore, metallographic sections of the analogy specimens in the area of the cross bore were used. The test force was $F_{HM} = 490$ mN, which was applied over a period of $t = 20$ s. The holding time during the hardness test was $T_{hold} = 10$ s. The subsequent reduction in force also took place over a period of $t = 20$ s. The fracture patterns were analyzed using reflected light microscope images and scanning electron microscope (SEM) images. The residual stresses were recorded using X-ray diffraction (XRD) measurements with an Xstress G2R with a Mythen detector from Stresstech GmbH. The combination of XRD measurement, electrochemical etching with a Polimat from Buehler and depth measurements using a dial gauge with an alignment using a MarSurf GD 120 tactile feed device from Mahr was used to determine residual stress depth curves.

In addition to the residual stress measurements of the specimen as manufactured, the redistribution of the residual stresses due to cyclic loading was taken into account both experimentally and numerically (see [12]). In addition to the residual stresses, the load stress distribution due to the torsional load was taken into account using finite element (FE) simulations. The simulation model presented in Uhlmann et al. was used, although no residual stresses were considered, and the calculation was simplified using a purely elastic material model [12].
## Results and Discussion

Observations. In Figure 2 the results of the torsional fatigue strength by means of S-N curves are shown. The following observations were made on terms of the results:

* Comparing the references (REF and IND), the induction hardened specimen has a higher fatigue strength, while the slope of the S-N curve is similar.

* Comparing the deep rolled (DR1, DR2 and IND + DR1) and the shot peened (SP1 and SP2) variants to the references (REF and IND) a different slope is observed, resulting in a higher fatigue strength for lower load amplitudes.

* The fatigue strength of the shot peened (SP1 and SP2) variants for a lower load amplitude is higher than for the deep rolled (DR1 and DR2) variants.

* The fatigue strength for the induction hardened and deep rolled (IND + DR1) is increased and the slope is improved compared to REF.

* The breakage of some IND + DR1 specimen is shifted out of the cross bore.

<figure><img src="image_2609.png" /></figure>

REF: Reference (not surface treated); SP: Shot peened; DR: Deep rolled; IND: Induction hardened

<caption>Figure 2: Fatigue strength results</caption>

Explanation approach. In the following, only the results of the surface layer properties were presented, based on which the explanatory approaches were derived. It is to say, that the explanatory approaches are not invalidated by the additional results not shown.

The increase of fatigue strength of the IND specimen compared to the REF specimen may be explained on the basis of the crack initiation. In Figure 3 (a) the fracture pattern of the REF and the IND specimen are juxtaposed. The crack initiation location of the hardened specimen is shifted to the inside of the bore, where the load stresses are lower than at the crack initiation location of the REF specimen resulting in the higher fatigue strength. In Figure 3 (b) the relative decrease of the load stresses is plotted over the distance to the chamfer into the bore for a shaft diameter of $d_s = 30$ mm to reinforce the explanation approach. While the load stresses at the crack initiation of
## Results and Discussion

Observations. In Figure 2 the results of the torsional fatigue strength by means of S-N curves are shown. The following observations were made on terms of the results:

* Comparing the references (REF and IND), the induction hardened specimen has a higher fatigue strength, while the slope of the S-N curve is similar.

* Comparing the deep rolled (DR1, DR2 and IND + DR1) and the shot peened (SP1 and SP2) variants to the references (REF and IND) a different slope is observed, resulting in a higher fatigue strength for lower load amplitudes.

* The fatigue strength of the shot peened (SP1 and SP2) variants for a lower load amplitude is higher than for the deep rolled (DR1 and DR2) variants.

* The fatigue strength for the induction hardened and deep rolled (IND + DR1) is increased and the slope is improved compared to REF.

* The breakage of some IND + DR1 specimen is shifted out of the cross bore.

<figure><img src="image_2602.png" /></figure>

REF: Reference (not surface treated); SP: Shot peened; DR: Deep rolled; IND: Induction hardened

<caption>Figure 2: Fatigue strength results</caption>

Explanation approach. In the following, only the results of the surface layer properties were presented, based on which the explanatory approaches were derived. It is to say, that the explanatory approaches are not invalidated by the additional results not shown.

The increase of fatigue strength of the IND specimen compared to the REF specimen may be explained on the basis of the crack initiation. In Figure 3 (a) the fracture pattern of the REF and the IND specimen are juxtaposed. The crack initiation location of the hardened specimen is shifted to the inside of the bore, where the load stresses are lower than at the crack initiation location of the REF specimen resulting in the higher fatigue strength. In Figure 3 (b) the relative decrease of the load stresses is plotted over the distance to the chamfer into the bore for a shaft diameter of $d_s$ = 30 mm to reinforce the explanation approach. While the load stresses at the crack initiation of
the REF specimen were at the maximum, the load stresses at the crack initiation of the IND specimen were about 30 % reduced.

<figure><img src="image_2610.png" /><figcaption>Figure 3: Comparison of fracture pattern of REF and IND specimen (a), numerically obtained load stresses depending on the distance to chamfer (b) and microstructure of IND specimen (c)</figcaption></figure>

The shift of the crack initiation location of the INT specimen into the cross bore is explained due to the hardened surface layer as a result of martensite transformation. In its initial state, EN-GJS-700-2 has a ferritic-pearlitic microstructure with spheroidal graphite inclusions. Due to hardening, the surface layer is transformed into martensite (see Figure 3 (c)). The higher hardness also results in reduced crack growth, which can be seen in the fracture patterns from the lower roughness in the hardened area compared to the area of the original material.

The residual stresses and especially the redistribution of the residual stresses in the area of crack initiation offer a possible explanation for the different slopes of the S-N curve comparing the deep rolled or shot peened variants with the reference variants. In Figure 4 (a) the tangential stress graphs for DR1 and SP1 after loading are compared to the residual stresses before loading. A qualitative comparison of the graphs for the different loads show that the redistribution of residual stresses is more pronounced at higher loads. Therefore, the influence of the residual stresses induced by deep rolling and shot peening on the fatigue strength is less pronounced. Accordingly, the slope of the S-N curve is improved compared to the REF variant.
the REF specimen were at the maximum, the load stresses at the crack initiation of the IND specimen were about 30 % reduced.

<figure><img src="image_2603.png" /><figcaption>Figure 3: Comparison of fracture pattern of REF and IND specimen (a), numerically obtained load stresses depending on the distance to chamfer (b) and microstructure of IND specimen (c)</figcaption></figure>

The shift of the crack initiation location of the INT specimen into the cross bore is explained due to the hardened surface layer as a result of martensite transformation. In its initial state, EN-GJS-700-2 has a ferritic-pearlitic microstructure with spheroidal graphite inclusions. Due to hardening, the surface layer is transformed into martensite (see Figure 3 (c)). The higher hardness also results in reduced crack growth, which can be seen in the fracture patterns from the lower roughness in the hardened area compared to the area of the original material.

The residual stresses and especially the redistribution of the residual stresses in the area of crack initiation offer a possible explanation for the different slopes of the S-N curve comparing the deep rolled or shot peened variants with the reference variants. In Figure 4 (a) the tangential stress graphs for DR1 and SP1 after loading are compared to the residual stresses before loading. A qualitative comparison of the graphs for the different loads show that the redistribution of residual stresses is more pronounced at higher loads. Therefore, the influence of the residual stresses induced by deep rolling and shot peening on the fatigue strength is less pronounced. Accordingly, the slope of the S-N curve is improved compared to the REF variant.
<figure><img src="image_2611.png" /><figcaption>Figure 4: *Redistributed residual stresses (a) and residual stresses at the cross bore surface after deep rolling and shot peening (b)*</figcaption></figure>

The fatigue strength of the shot peened (SP1 and SP2) variants is higher than that of the deep rolled (DR1 and DR2) variants under lower loadings due to the induced residual stresses and load stresses. Torsional loading results in tangential load stresses in the bore, and the tangential residual stresses counter the load stresses. Figure 4 (b) shows the tangential and axial residual stresses at the surface of the bore for the shot peened variant SP1 and the deep rolled variants DR1 and DR2. The residual stresses in both directions are homogeneous after shot peening. However, after deep rolling, the tangential residual stresses are lower and the axial residual stresses are higher than after shot peening, resulting in lower fatigue strength. For this case, the results indicate that residual stresses have a more significant impact on fatigue strength than roughness. This is because the deep rolled variants have a lower roughness than the shot peened variants which suggests increased fatigue strength.

The increased fatigue strength and improved slope of the S-N curve of the IND + DR1 variant can be explained by the compressive residual stresses described above as well as the shift of the crack initiation location. Induction hardening shifts the crack initiation location into the bore, resulting in reduced load stresses at the crack initiation location (compare Figure 3). Deep rolling induces compressive residual stresses in the surface area of the specimen (compare Figure 4). This is particularly important because, due to the larger volume of martensite in the hardened region compared to the ferrite-pearlite structure, it can be assumed that compressive residual stresses are present in the hardened region and, prior to deep rolling, tensile residual stresses are present in the transition area, where the crack initiation is located. The compressive residual stresses counteract the critical tensile load stresses increasing the fatigue strength. Due to a load depending redistribution of the compressive residual stresses this effect is more pronounced for lower loads resulting in the observed change of slope.

The shift of the breakage location of the IND + DR1 variant observed for some specimen is explained by a shift of the weakest specimen location. Those specimens always broke in the transition area (axial to the shaft shown in Figure 2). As described above in this area tensile residual stresses are expected and combined with possible surface defects the observed shift of the breakage location results.

The increase in fatigue strength can be used to reduce the dimensioning of the shafts. To quantify this, the maximum load stresses were first determined for a target fatigue strength of $N = 10^6$ load cycles using the torsional moment at the breakpoint (see Figure 2) and the elastic FE simulation. The geometry of the cross bore was not varied, resulting in no linear relationship between the load stresses and the shaft diameter. A Gaussian process regression model (GPR) was derived on the basis of the FE simulations in order to estimate the load stresses for intermediate
<figure><img src="image_2604.png" /><figcaption>Figure 4: *Redistributed residual stresses (a) and residual stresses at the cross bore surface after deep rolling and shot peening (b)*</figcaption></figure>

The fatigue strength of the shot peened (SP1 and SP2) variants is higher than that of the deep rolled (DR1 and DR2) variants under lower loadings due to the induced residual stresses and load stresses. Torsional loading results in tangential load stresses in the bore, and the tangential residual stresses counter the load stresses. Figure 4 (b) shows the tangential and axial residual stresses at the surface of the bore for the shot peened variant SP1 and the deep rolled variants DR1 and DR2. The residual stresses in both directions are homogeneous after shot peening. However, after deep rolling, the tangential residual stresses are lower and the axial residual stresses are higher than after shot peening, resulting in lower fatigue strength. For this case, the results indicate that residual stresses have a more significant impact on fatigue strength than roughness. This is because the deep rolled variants have a lower roughness than the shot peened variants which suggests increased fatigue strength.

The increased fatigue strength and improved slope of the S-N curve of the IND + DR1 variant can be explained by the compressive residual stresses described above as well as the shift of the crack initiation location. Induction hardening shifts the crack initiation location into the bore, resulting in reduced load stresses at the crack initiation location (compare Figure 3). Deep rolling induces compressive residual stresses in the surface area of the specimen (compare Figure 4). This is particularly important because, due to the larger volume of martensite in the hardened region compared to the ferrite-pearlite structure, it can be assumed that compressive residual stresses are present in the hardened region and, prior to deep rolling, tensile residual stresses are present in the transition area, where the crack initiation is located. The compressive residual stresses counteract the critical tensile load stresses increasing the fatigue strength. Due to a load depending redistribution of the compressive residual stresses this effect is more pronounced for lower loads resulting in the observed change of slope.

The shift of the breakage location of the IND + DR1 variant observed for some specimen is explained by a shift of the weakest specimen location. Those specimens always broke in the transition area (axial to the shaft shown in Figure 2). As described above in this area tensile residual stresses are expected and combined with possible surface defects the observed shift of the breakage location results.

The increase in fatigue strength can be used to reduce the dimensioning of the shafts. To quantify this, the maximum load stresses were first determined for a target fatigue strength of $N = 10^6$ load cycles using the torsional moment at the breakpoint (see Figure 2) and the elastic FE simulation. The geometry of the cross bore was not varied, resulting in no linear relationship between the load stresses and the shaft diameter. A Gaussian process regression model (GPR) was derived on the basis of the FE simulations in order to estimate the load stresses for intermediate
points as well and to reduce the simulation effort. Figure 5 (a) shows the validation of the GPR. For validation, an additional FE simulation was carried out in the area of maximum uncertainty of the GPR. The deviation of the GPR is less than 5 %. In Figure 5 (b) the relationship between the torsional moment, shaft diameter and the maximum Mises stresses is shown. Furthermore, the uncertainty of the GPR is presented.

<figure><img src="image_2612.png" /><figcaption>Figure 5: Validation of GPR model (a) and relationship between torsion moment, shaft diameter and maximum Mises stresses as well as the GPR uncertainty (b)</figcaption></figure>

Subsequently, the minimum diameter was determined at which a torsional moment of $T_M$ = 495.89 Nm (Torsional moment of REF variant) causes the previously determined maximum load stresses. The results are shown in Table 1. The reduced shaft diameter results in a reduced shaft cross-section and correspondingly reduced specimen weight of 18-37 %.

<table><caption>Table 1: Estimation of the savings potential in terms of the shaft diameter for a target fatigue strength of N = 10⁶ load cycles</caption><thead><tr><th>Variant</th><th>Torsional moment M<sub>T</sub></th><th>Mises stresses for D = 30 mm in Mpa</th><th>Minimum diameter D<sub>min</sub> for torsional moment M<sub>T</sub> = 495.89 Nm in mm</th></tr></thead><tbody><tr><td>REF</td><td>495.89</td><td>309.25</td><td>30</td></tr><tr><td>DR2</td><td>647.71</td><td>406.33</td><td>27.2</td></tr><tr><td>DR1</td><td>668.81</td><td>419.83</td><td>26.9</td></tr><tr><td>SP2</td><td>683.09</td><td>428.09</td><td>26.7</td></tr><tr><td>SP1</td><td>694.72</td><td>436.39</td><td>26.5</td></tr><tr><td>IND</td><td>722.46</td><td>454.13</td><td>26.2</td></tr><tr><td>IND + DR1</td><td>939.32</td><td>592.79</td><td>23.9</td></tr></tbody></table>
points as well and to reduce the simulation effort. Figure 5 (a) shows the validation of the GPR. For validation, an additional FE simulation was carried out in the area of maximum uncertainty of the GPR. The deviation of the GPR is less than 5 %. In Figure 5 (b) the relationship between the torsional moment, shaft diameter and the maximum Mises stresses is shown. Furthermore, the uncertainty of the GPR is presented.

<figure><img src="image_2605.png" /><figcaption>Figure 5: Validation of GPR model (a) and relationship between torsion moment, shaft diameter and maximum Mises stresses as well as the GPR uncertainty (b)</figcaption></figure>

Subsequently, the minimum diameter was determined at which a torsional moment of $T_M$ = 495.89 Nm (Torsional moment of REF variant) causes the previously determined maximum load stresses. The results are shown in Table 1. The reduced shaft diameter results in a reduced shaft cross-section and correspondingly reduced specimen weight of 18-37 %.

<table><caption>Table 1: Estimation of the savings potential in terms of the shaft diameter for a target fatigue strength of N = 10⁶ load cycles</caption><thead><tr><th>Variant</th><th>Torsional moment M<sub>T</sub></th><th>Mises stresses for D = 30 mm in Mpa</th><th>Minimum diameter D<sub>min</sub> for torsional moment M<sub>T</sub> = 495.89 Nm in mm</th></tr></thead><tbody><tr><td>REF</td><td>495.89</td><td>309.25</td><td>30</td></tr><tr><td>DR2</td><td>647.71</td><td>406.33</td><td>27.2</td></tr><tr><td>DR1</td><td>668.81</td><td>419.83</td><td>26.9</td></tr><tr><td>SP2</td><td>683.09</td><td>428.09</td><td>26.7</td></tr><tr><td>SP1</td><td>694.72</td><td>436.39</td><td>26.5</td></tr><tr><td>IND</td><td>722.46</td><td>454.13</td><td>26.2</td></tr><tr><td>IND + DR1</td><td>939.32</td><td>592.79</td><td>23.9</td></tr></tbody></table>
## Summary and outlook

In the presented work the cause effect relations between the surface layer properties and the resulting fatigue strength of specimen with cross bores from EN-GJS-700-2 was investigated. The surface layer properties of the cross bore were manipulated using deep rolling, shot peening and induction hardening. The results suggest that the induced tangential compressive residual stresses improve the fatigue strength. Due to a load dependent redistribution of the residual stresses the compressive residual stresses were more effective for lower loads resulting in an improved slope of the S-N curve. Furthermore, the crack initiation location was shifted into the bore due to induction hardening. As a result, the acting load stresses at the crack initiation location was reduced resulting in an increased fatigue strength. The increase in fatigue strength may be used to reduce the dimensions of the shaft resulting in a weight reduction of up to 37 %. In future work, the approach to predict the fatigue strength of the shot peened variant of Reissner et al. will be applied to the induction hardened, the deep rolled and induction hardened and then deep rolled variants [11].

## Acknowledgements

This research was carried out in the framework of the industrial collective research programme (IGF no. 20407 N). It was supported by the Federal Ministry for Economic Affairs and Climate Action (BMWK) through the AiF (German Federation of Industrial Research Associations eV) based on a decision taken by the German Bundestag. Further, the authors would like to thank the ECOROLL AG Werkzeugtechnik for their support in conducting the deep rolling experiments, the Metal Improvement Company LLC for conducting the shot peening experiments, FONDIUM Singen GmbH for manufacturing the specimen cited in this paper and the Institute for Materials Applications in Mechanical Engineering (IWM) of RWTH Aachen University especially Lennart Mirko Scholl and Withold Hildebrandt for the support and provision of the torsion test rig. The authors would also like to thank the Fraunhofer Institute for Structural Durability and System Reliability LBF, in particular Felix Reissner, for the excellent cooperation during the project. Simulations were performed with computing resources granted by RWTH Aachen University under project rwth1215.

## References

[1] Federal Climate Change Act: KSG, 2019. [Online]. Available: https://www.gesetze-im-internet.de/englisch_ksg/englisch_ksg.pdf

[2] C. Miki, K. Homma, and T. Tominaga, "High strength and high performance steels and their use in bridge structures," Journal of Constructional Steel Research, vol. 58, no. 1, pp. 3–20, 2002. https://doi.org/10.1016/S0143-974X(01)00028-1

[3] A. Ucharzewski, "Massivumformung - eine Prozesskette für den Leichtbau," (in de), Lightweight Des, vol. 3, no. 2, pp. 42–49, 2010. https://doi.org/10.1007/bf03223603

[4] J. S. Eckersley and B. Ferrelli, "Using Shot Peening to Multiply the Life of Compressor Components," International Compressor Engineering Conference, vol. 1992, Paper 885, 1992.

[5] N. Tsuji, S. Tanaka, and T. Takasugi, "Effect of combined plasma-carburizing and deep-rolling on notch fatigue property of Ti-6Al-4V alloy," Materials Science and Engineering: A, vol. 499, 1-2, pp. 482–488, 2009. https://doi.org/10.1016/j.msea.2008.09.008

[6] U. Kabasakaloglu and H. Saruhan, "Effects of induction hardened surface depth on the dynamic behavior of rotating shaft systems," Materials Testing, vol. 61, no. 3, pp. 277–281, 2019. https://doi.org/10.3139/120.111316
## Summary and outlook

In the presented work the cause effect relations between the surface layer properties and the resulting fatigue strength of specimen with cross bores from EN-GJS-700-2 was investigated. The surface layer properties of the cross bore were manipulated using deep rolling, shot peening and induction hardening. The results suggest that the induced tangential compressive residual stresses improve the fatigue strength. Due to a load dependent redistribution of the residual stresses the compressive residual stresses were more effective for lower loads resulting in an improved slope of the S-N curve. Furthermore, the crack initiation location was shifted into the bore due to induction hardening. As a result, the acting load stresses at the crack initiation location was reduced resulting in an increased fatigue strength. The increase in fatigue strength may be used to reduce the dimensions of the shaft resulting in a weight reduction of up to 37 %. In future work, the approach to predict the fatigue strength of the shot peened variant of Reissner et al. will be applied to the induction hardened, the deep rolled and induction hardened and then deep rolled variants [11].

## Acknowledgements

This research was carried out in the framework of the industrial collective research programme (IGF no. 20407 N). It was supported by the Federal Ministry for Economic Affairs and Climate Action (BMWK) through the AiF (German Federation of Industrial Research Associations eV) based on a decision taken by the German Bundestag. Further, the authors would like to thank the ECOROLL AG Werkzeugtechnik for their support in conducting the deep rolling experiments, the Metal Improvement Company LLC for conducting the shot peening experiments, FONDIUM Singen GmbH for manufacturing the specimen cited in this paper and the Institute for Materials Applications in Mechanical Engineering (IWM) of RWTH Aachen University especially Lennart Mirko Scholl and Withold Hildebrandt for the support and provision of the torsion test rig. The authors would also like to thank the Fraunhofer Institute for Structural Durability and System Reliability LBF, in particular Felix Reissner, for the excellent cooperation during the project. Simulations were performed with computing resources granted by RWTH Aachen University under project rwth1215.

## References

[1] Federal Climate Change Act: KSG, 2019. [Online]. Available: https://www.gesetze-im-internet.de/englisch_ksg/englisch_ksg.pdf

[2] C. Miki, K. Homma, and T. Tominaga, "High strength and high performance steels and their use in bridge structures," Journal of Constructional Steel Research, vol. 58, no. 1, pp. 3–20, 2002. https://doi.org/10.1016/S0143-974X(01)00028-1

[3] A. Ucharzewski, "Massivumformung - eine Prozesskette für den Leichtbau," (in de), Lightweight Des, vol. 3, no. 2, pp. 42–49, 2010. https://doi.org/10.1007/bf03223603

[4] J. S. Eckersley and B. Ferrelli, "Using Shot Peening to Multiply the Life of Compressor Components," International Compressor Engineering Conference, vol. 1992, Paper 885, 1992.

[5] N. Tsuji, S. Tanaka, and T. Takasugi, "Effect of combined plasma-carburizing and deep-rolling on notch fatigue property of Ti-6Al-4V alloy," Materials Science and Engineering: A, vol. 499, 1-2, pp. 482–488, 2009. https://doi.org/10.1016/j.msea.2008.09.008

[6] U. Kabasakaloglu and H. Saruhan, "Effects of induction hardened surface depth on the dynamic behavior of rotating shaft systems," Materials Testing, vol. 61, no. 3, pp. 277–281, 2019. https://doi.org/10.3139/120.111316
[7] W. Zielecki, M. Bucior, T. Trzepiecinski, and K. Ochał, "Effect of slide burnishing of shoulder fillets on the fatigue strength of X19NiCrMo4 steel shafts," Int J Adv Manuf Technol, vol. 106, 5-6, pp. 2583-2593, 2020. https://doi.org/10.1007/s00170-019-04815-7

[8] S. Woo, Design of mechanical systems: Accelerated lifecycle testing and reliability. Cham: Springer, 2023.

[9] M. O. Görtan and B. Yüksel, "Improvement of Fatigue Properties of EN AW 6082 Aluminum Alloy using Different Deep Rolling Directions," Int. J. Automot. Mech. Eng., vol. 20, no. 2, pp. 10351-10358, 2023. https://doi.org/10.15282/ijame.20.2.2023.02.0800

[10] M. Kobayashi, T. Matsui, and Y. Murakami, "Mechanism of creation of compressive residual stress by shot peening," International Journal of Fatigue, vol. 20, no. 5, pp. 351-357, 1998. https://doi.org/10.1016/S0142-1123(98)00002-4

[11] F. Reissner, L. Uhlmann, J. Baumgartner, T. Herrig, and T. Bergs, "Improved lifetime estimation of shot-peened shaft bores using a numerical approach," Procedia Structural Integrity, 2024.

[12] L. Uhlmann, F. Reissner, S. N. Rathnakar, T. Herrig, J. Baumgartner, and T. Bergs, "Numerical Modeling of the Redistribution of Residual Stresses in Deep Rolled Cross Bores in Shafts from GJS700-2," in Lecture Notes in Mechanical Engineering, Proceedings of the 14th International Conference on the Technology of Plasticity - Current Trends in the Technology of Plasticity: ICTP 2023 - Volume 1, K. Mocellin, P.-O. Bouchard, R. Bigot, and T. Balan, Eds., 1st ed., Cham: Springer Nature Switzerland; Imprint Springer, 2024, pp. 283-290.
[7] W. Zielecki, M. Bucior, T. Trzepiecinski, and K. Ochał, "Effect of slide burnishing of shoulder fillets on the fatigue strength of X19NiCrMo4 steel shafts," Int J Adv Manuf Technol, vol. 106, 5-6, pp. 2583-2593, 2020. https://doi.org/10.1007/s00170-019-04815-7

[8] S. Woo, Design of mechanical systems: Accelerated lifecycle testing and reliability. Cham: Springer, 2023.

[9] M. O. Görtan and B. Yüksel, "Improvement of Fatigue Properties of EN AW 6082 Aluminum Alloy using Different Deep Rolling Directions," Int. J. Automot. Mech. Eng., vol. 20, no. 2, pp. 10351-10358, 2023. https://doi.org/10.15282/ijame.20.2.2023.02.0800

[10] M. Kobayashi, T. Matsui, and Y. Murakami, "Mechanism of creation of compressive residual stress by shot peening," International Journal of Fatigue, vol. 20, no. 5, pp. 351-357, 1998. https://doi.org/10.1016/S0142-1123(98)00002-4

[11] F. Reissner, L. Uhlmann, J. Baumgartner, T. Herrig, and T. Bergs, "Improved lifetime estimation of shot-peened shaft bores using a numerical approach," Procedia Structural Integrity, 2024.

[12] L. Uhlmann, F. Reissner, S. N. Rathnakar, T. Herrig, J. Baumgartner, and T. Bergs, "Numerical Modeling of the Redistribution of Residual Stresses in Deep Rolled Cross Bores in Shafts from GJS700-2," in Lecture Notes in Mechanical Engineering, Proceedings of the 14th International Conference on the Technology of Plasticity - Current Trends in the Technology of Plasticity: ICTP 2023 - Volume 1, K. Mocellin, P.-O. Bouchard, R. Bigot, and T. Balan, Eds., 1st ed., Cham: Springer Nature Switzerland; Imprint Springer, 2024, pp. 283-290.
# Preliminary investigation of the energy consumption in friction stir welding of aluminum alloys

COZZOLINO Ersilia¹,ª*, SILVESTRI Alessia Teresa¹,ᵇ, DE ALTERIIS Giorgio²,ᶜ, ASTARITA Antonello¹,ᵈ, SCHIANO LO MORIELLO Rosaro²,ᵉ, SQUILLACE Antonino¹,ᵉ

¹Department of Chemical, Materials and Production Engineering, University of Naples Federico II, Naples, Italy

²Department of Industrial Engineering, University of Naples Federico II, Naples, Italy

ªersilia.cozzolino@unina.it, ᵇalessiateresa.silvestri@unina.it, ᶜgiorgio.dealteriis@unina.it,

ᵈantonello.astarita@unina.it, ᶜrosario.schianolomoriello@unina.it, ᵈsquillac@unina.it

**Keywords:** Friction Stir Welding, Sustainability, Energy Consumption

**Abstract.** Manufacturing is responsible for significant emissions and energy consumption. However, several manufacturing processes, especially the non-conventional ones, are still poorly investigated from a sustainable perspective. Friction Stir Welding (FSW) is a non-conventional solid-state welding process developed in the last decade which can produce joints with higher mechanical and metallurgical properties. Very few studies exist in the literature on the FSW from a sustainable perspective, so this paper aims to fill this gap of knowledge. Following the in-depth approach of CO2PE! procedure, the power profile of the machine overall the process has been investigated to carry out time and energy studies. The main factors affecting electrical energy demand in the FSW process are highlighted and guidelines for decision-making from an energy-saving perspective are released.

**Introduction**

Manufacturing has always been seen as the key factor for the prosperity and development of countries [1] [2]. Indeed, manufacturing processes transform materials and information into goods for the satisfaction of human needs. However, transforming raw materials into products is also a significant source of environmental pollution. This environmental pollution can be directly due to manufacturing processes, or indirectly through the energy needed for realizing these processes. Manufacturing waste depends on various factors such as the technology adopted and the material processed. The massive use of energy for industrial operations in Europe is responsible for significant CO2 emissions and thus global climate change. Nevertheless, several manufacturing processes, especially the non-conventional ones, are still poorly investigated from a sustainable perspective [3]. In the last decades, researchers have developed new joining technologies to radically settle all the issues linked to the melting-based welding techniques. Friction Stir Welding (FSW) is a solid-state welding process developed and patented in 1991 by The Welding Institute (TWI) of Cambridge. FSW is a promising solution to produce high-quality welds for low-melting materials, such as aluminum, copper, magnesium, and their alloys, in which the cylindrical tool having shoulder and pin segments rotates and travels along the joint line. Compared to the conventional fusion welding methods, this FSW can produce joints with higher mechanical and metallurgical properties. The main advantage of FSW over traditional fusion welding is the reduction in the heat-affected zone (HAZ), and the joints exhibit excellent mechanical and corrosion resistance properties [4] [5]. The FSW operation does not require any consumables or shielding element techniques in the process, including the electrodes, fillers, and shielding gas. For this reason, it can be thought that the FSW process is more sustainable than conventional ones.

<figure><img src="image_2613.png" /></figure>
# Preliminary investigation of the energy consumption in friction stir welding of aluminum alloys

COZZOLINO Ersilia¹,ᵃ*, SILVESTRI Alessia Teresa¹,ᵇ, DE ALTERIIS Giorgio²,ᶜ, ASTARITA Antonello¹,ᵈ, SCHIANO LO MORIELLO Rosaro²,ᵉ, SQUILLACE Antonino¹,ᵉ

¹Department of Chemical, Materials and Production Engineering, University of Naples Federico II, Naples, Italy

²Department of Industrial Engineering, University of Naples Federico II, Naples, Italy

ªersilia.cozzolino@unina.it, ᵇalessiateresa.silvestri@unina.it, ᶜgiorgio.dealteriis@unina.it,

ᵈantonello.astarita@unina.it, ᶜrosario.schianolomoriello@unina.it, ᵈsquillac@unina.it

**Keywords:** Friction Stir Welding, Sustainability, Energy Consumption

**Abstract.** Manufacturing is responsible for significant emissions and energy consumption. However, several manufacturing processes, especially the non-conventional ones, are still poorly investigated from a sustainable perspective. Friction Stir Welding (FSW) is a non-conventional solid-state welding process developed in the last decade which can produce joints with higher mechanical and metallurgical properties. Very few studies exist in the literature on the FSW from a sustainable perspective, so this paper aims to fill this gap of knowledge. Following the in-depth approach of CO2PE! procedure, the power profile of the machine overall the process has been investigated to carry out time and energy studies. The main factors affecting electrical energy demand in the FSW process are highlighted and guidelines for decision-making from an energy-saving perspective are released.

**Introduction**

Manufacturing has always been seen as the key factor for the prosperity and development of countries [1] [2]. Indeed, manufacturing processes transform materials and information into goods for the satisfaction of human needs. However, transforming raw materials into products is also a significant source of environmental pollution. This environmental pollution can be directly due to manufacturing processes, or indirectly through the energy needed for realizing these processes. Manufacturing waste depends on various factors such as the technology adopted and the material processed. The massive use of energy for industrial operations in Europe is responsible for significant CO2 emissions and thus global climate change. Nevertheless, several manufacturing processes, especially the non-conventional ones, are still poorly investigated from a sustainable perspective [3]. In the last decades, researchers have developed new joining technologies to radically settle all the issues linked to the melting-based welding techniques. Friction Stir Welding (FSW) is a solid-state welding process developed and patented in 1991 by The Welding Institute (TWI) of Cambridge. FSW is a promising solution to produce high-quality welds for low-melting materials, such as aluminum, copper, magnesium, and their alloys, in which the cylindrical tool having shoulder and pin segments rotates and travels along the joint line. Compared to the conventional fusion welding methods, this FSW can produce joints with higher mechanical and metallurgical properties. The main advantage of FSW over traditional fusion welding is the reduction in the heat-affected zone (HAZ), and the joints exhibit excellent mechanical and corrosion resistance properties [4] [5]. The FSW operation does not require any consumables or shielding element techniques in the process, including the electrodes, fillers, and shielding gas. For this reason, it can be thought that the FSW process is more sustainable than conventional ones.

<figure><img src="image_2606.png" /></figure>

Content from this work may be used under the terms of the Creative Commons Attribution 3.0 license. Any further distribution of this work must maintain attribution to the author(s) and the title of the work, journal citation and DOI. Published under license by Materials Research Forum LLC.
However, the literature lacks guidelines for performing the FSW process by minimizing energy consumption. Also, power consumption during the subphases of the process is not well investigated. This study aims at filling these gaps of knowledge.

In particular, in this paper, an electrical energy demand characterization of the FSW of AA2024-T3 and AA7075-T6 sheets in a lap-joint configuration is presented. According to the in-depth approach methodology proposed by Kellens et al.[6] [7], both time study as well as power studies are developed on a FSW dedicated machine tool. Key factors and subphases in FSW affecting electrical energy demand are highlighted and guidelines for decision-making from a sustainable perspective.

## Experimental set-up

Aluminum sheets 158x120 mm, AA2024-T3 and AA7075-T6, were lap-joint welded by FSW. These aluminum alloys are usually difficult to join using traditional welding so FSW can be considered a useful solution. The lap joint configuration is typically adopted in aeronautical industry where structural parts need to be lightweight and mechanically resistant. The final thickness of the lap-joint welded structure is 6.3 mm.

A traditional milling machine Adria Machine PBM-4EVS was used for the experimental activities and thanks to a home-made setup we performed the FSW process. The tool adopted in this study was characterized by a threaded pin having a truncated cone shape. Table 1 contains the main geometrical parameters of the tool.

<table><caption>Table 1. Geometrical parameters of the tool adopted to perform the FSW</caption><thead><tr><td>Parameter</td><td>Value</td></tr></thead><tbody><tr><td>Pin major diameter (mm)</td><td>3.96</td></tr><tr><td>Pin minor diameter (mm)</td><td>5.40</td></tr><tr><td>Shoulder diameter (mm)</td><td>13</td></tr><tr><td>Tool diameter (mm)</td><td>20</td></tr><tr><td>Tilt angle (°)</td><td>1</td></tr></tbody></table>

The sheets were previously cut in half and then AA 2024-T3 sheets was positioned on the top of the AA 7075-T6 sheets.

Technological parameters to perform the FSW are contained in Table 2. Two levels of tool rotation (1800 rpm and 3200 rpm) have been used for two different welds, by fixing the other process conditions, to analyze the differences in terms of time and energy consumption.

<table><caption>Table 2. Technological parameters of the tool adopted to perform the FSW</caption><thead><tr><td>Parameter</td><td>Value</td></tr></thead><tbody><tr><td>Tool rotation (rpm)</td><td>1800/3200</td></tr><tr><td>Tool feed rate (mm/min)</td><td>100</td></tr><tr><td>Tool plunge (mm)</td><td>4.1</td></tr><tr><td>Welded length (mm)</td><td>90</td></tr></tbody></table>

Qualistar Plus Power and Energy Quality Analyser CA8331 is the device adopted to record power consumption during the FSW process. The machine used to perform the FSW has a three-phase connection without neutral 32 A 380 V. Thus, three tension cables, three crocodile clips and three current sensors MiniFLEX MA193 were used to record the voltage and the current to measure the power. The sampling period for data acquisition was of 1 second [8]. Fig. 1 depicts the setup of the power device connected to the electrical connection of the machine.
However, the literature lacks guidelines for performing the FSW process by minimizing energy consumption. Also, power consumption during the subphases of the process is not well investigated. This study aims at filling these gaps of knowledge.

In particular, in this paper, an electrical energy demand characterization of the FSW of AA2024-T3 and AA7075-T6 sheets in a lap-joint configuration is presented. According to the in-depth approach methodology proposed by Kellens et al.[6] [7], both time study as well as power studies are developed on a FSW dedicated machine tool. Key factors and subphases in FSW affecting electrical energy demand are highlighted and guidelines for decision-making from a sustainable perspective.

## Experimental set-up

Aluminum sheets 158x120 mm, AA2024-T3 and AA7075-T6, were lap-joint welded by FSW. These aluminum alloys are usually difficult to join using traditional welding so FSW can be considered a useful solution. The lap joint configuration is typically adopted in aeronautical industry where structural parts need to be lightweight and mechanically resistant. The final thickness of the lap-joint welded structure is 6.3 mm.

A traditional milling machine Adria Machine PBM-4EVS was used for the experimental activities and thanks to a home-made setup we performed the FSW process. The tool adopted in this study was characterized by a threaded pin having a truncated cone shape. Table 1 contains the main geometrical parameters of the tool.

<table><caption>Table 1. Geometrical parameters of the tool adopted to perform the FSW</caption><thead><tr><td>Parameter</td><td>Value</td></tr></thead><tbody><tr><td>Pin major diameter (mm)</td><td>3.96</td></tr><tr><td>Pin minor diameter (mm)</td><td>5.40</td></tr><tr><td>Shoulder diameter (mm)</td><td>13</td></tr><tr><td>Tool diameter (mm)</td><td>20</td></tr><tr><td>Tilt angle (°)</td><td>1</td></tr></tbody></table>

The sheets were previously cut in half and then AA 2024-T3 sheets was positioned on the top of the AA 7075-T6 sheets.

Technological parameters to perform the FSW are contained in Table 2. Two levels of tool rotation (1800 rpm and 3200 rpm) have been used for two different welds, by fixing the other process conditions, to analyze the differences in terms of time and energy consumption.

<table><caption>Table 2. Technological parameters of the tool adopted to perform the FSW</caption><thead><tr><td>Parameter</td><td>Value</td></tr></thead><tbody><tr><td>Tool rotation (rpm)</td><td>1800/3200</td></tr><tr><td>Tool feed rate (mm/min)</td><td>100</td></tr><tr><td>Tool plunge (mm)</td><td>4.1</td></tr><tr><td>Welded length (mm)</td><td>90</td></tr></tbody></table>

Qualistar Plus Power and Energy Quality Analyser CA8331 is the device adopted to record power consumption during the FSW process. The machine used to perform the FSW has a three-phase connection without neutral 32 A 380 V. Thus, three tension cables, three crocodile clips and three current sensors MiniFLEX MA193 were used to record the voltage and the current to measure the power. The sampling period for data acquisition was of 1 second [8]. Fig. 1 depicts the setup of the power device connected to the electrical connection of the machine.
<figure><img src="image_2614.png" /><figcaption>Figure 1. Experimental setup to measure current and tension during FSW process</figcaption></figure>

The in-depth approach of the CO2PE! procedure has been adopted in this study. The objective of the CO2PE! Initiative – CO-operative effort on Process Emissions in manufacturing – is to cluster forces in different continents in order to analyze existing and emerging discrete part manufacturing processes for their ecological impact in terms of direct and indirect emissions [6]. In this investigation, energy study and time study were carried out based on the measurements to identify focus points for improvements in the FSW process from an energy-saving perspective based on the results generated [9].

## Results and Discussion

As above described, an in-depth approach has been carried out in this study to quantify the electrical energy demand of the experimental set-up to perform the FSW to join two aluminum sheets in a lap joint configuration. In particular, the time study was performed to identify the time spent by the machine tool to realize a specific operation by the switching on to the switching off of the machine at the end of the process. Energy study was carried out to highlight the contribution given by each subphase of the FSW process to the overall energy consumption. Fig. 2 shows the power recorded over time during the whole process by selecting a tool feed rate of 100 mm/min and a tool rotation of 1800 rpm.
<figure><img src="image_2607.png" /><figcaption>Figure 1. Experimental setup to measure current and tension during FSW process</figcaption></figure>

The in-depth approach of the CO2PE! procedure has been adopted in this study. The objective of the CO2PE! Initiative – CO-operative effort on Process Emissions in manufacturing – is to cluster forces in different continents in order to analyze existing and emerging discrete part manufacturing processes for their ecological impact in terms of direct and indirect emissions [6]. In this investigation, energy study and time study were carried out based on the measurements to identify focus points for improvements in the FSW process from an energy-saving perspective based on the results generated [9].

## Results and Discussion

As above described, an in-depth approach has been carried out in this study to quantify the electrical energy demand of the experimental set-up to perform the FSW to join two aluminum sheets in a lap joint configuration. In particular, the time study was performed to identify the time spent by the machine tool to realize a specific operation by the switching on to the switching off of the machine at the end of the process. Energy study was carried out to highlight the contribution given by each subphase of the FSW process to the overall energy consumption. Fig. 2 shows the power recorded over time during the whole process by selecting a tool feed rate of 100 mm/min and a tool rotation of 1800 rpm.
<figure><img src="image_2615.png" /><figcaption>Figure 2. Power consumption profile over time during the FSW process (100 mm/min, 1800 rpm)</figcaption></figure>

The power absorption of the machine during the FSW process has been investigated at different sub-phases to identify each contribution. A power around 150 W is recorded when the machine is switched on. Then, a peak of power of around 3000 W is registered when motors of the machine are turned on. When the tool plunge starts, a peak of power of about 2300 W is reached. When the tool shoulder comes in contact with the top surface of the sheets, power rapidly increases reaching the peak value of 4191 W. This result is very similar to that obtained by Buffa et al. [9]. Then, power gradually decreases because of the sheet softening induced by the heat produced by the friction (dwell time). When the welding starts, a peak value of 3589 W is reached. Then, it rapidly decreases when an equilibrium is obtained between the heat introduced by the friction forces and the heat dissipated by conduction and convection. After the completion of the tool path, the tool is lifted out resulting in a decrease of the power. Consequentially, the motors are switched off before the machine is turned off.

By following the in-depth approach, a time study was developed. Fig. 3 displays a pie chart of the time study: the percentages of the single subphases in terms of time overall process are shown from the machine start to the machine shut down. According to our results, 75% of total time is used for productive phases, including tool plunge, dwell time, welding and tool lift. In the productive phases, 31% is for welding while the remaining 44% is used for tool plunging, dwelling time and tool lifting. Dwell time required the 32% of time. The remaining time is required for non-productive phases including auxiliary operations and tool positioning. It is worthwhile to note that Welding time is function of the length welded. If this length increases, welding time obviously increases.
<figure><img src="image_2608.png" /><figcaption>Figure 2. Power consumption profile over time during the FSW process (100 mm/min, 1800 rpm)</figcaption></figure>

The power absorption of the machine during the FSW process has been investigated at different sub-phases to identify each contribution. A power around 150 W is recorded when the machine is switched on. Then, a peak of power of around 3000 W is registered when motors of the machine are turned on. When the tool plunge starts, a peak of power of about 2300 W is reached. When the tool shoulder comes in contact with the top surface of the sheets, power rapidly increases reaching the peak value of 4191 W. This result is very similar to that obtained by Buffa et al. [9]. Then, power gradually decreases because of the sheet softening induced by the heat produced by the friction (dwell time). When the welding starts, a peak value of 3589 W is reached. Then, it rapidly decreases when an equilibrium is obtained between the heat introduced by the friction forces and the heat dissipated by conduction and convection. After the completion of the tool path, the tool is lifted out resulting in a decrease of the power. Consequentially, the motors are switched off before the machine is turned off.

By following the in-depth approach, a time study was developed. Fig. 3 displays a pie chart of the time study: the percentages of the single subphases in terms of time overall process are shown from the machine start to the machine shut down. According to our results, 75% of total time is used for productive phases, including tool plunge, dwell time, welding and tool lift. In the productive phases, 31% is for welding while the remaining 44% is used for tool plunging, dwelling time and tool lifting. Dwell time required the 32% of time. The remaining time is required for non-productive phases including auxiliary operations and tool positioning. It is worthwhile to note that Welding time is function of the length welded. If this length increases, welding time obviously increases.
<figure><img src="image_2616.png" /><figcaption>Figure 3. Time study of the FSW process (100 mm/min, 1800 rpm)</figcaption></figure>

By combining power and time measurements, energy consumption has been calculated for each subphase of the whole process. Fig. 4 shows the percentages of the energy study carried out. 94% of the energy consumption is required by the productive phases, including tool plunge, dwell time, welding, and tool lift. Welding time required almost half of the overall energy consumption (45%). Dwell time also required a non-negligible amount of energy consumption due to the fact that power reached over time is higher than other subphases. Thus, in terms of both time and energy consumption, welding subphase plays a dominant role. Instead, terms of power values reached, dwell time plays a crucial role.

<figure><img src="image_2617.png" /><figcaption>Figure 4. Energy study of the FSW process (100 mm/min, 1800 rpm)</figcaption></figure>

Similarly, Fig. 5 shows the power recorded over time during the whole process by selecting a tool feed rate of 100 mm/min and a tool rotation of 3200 rpm.
<figure><img src="image_2609.png" /><figcaption>Figure 3. Time study of the FSW process (100 mm/min, 1800 rpm)</figcaption></figure>

By combining power and time measurements, energy consumption has been calculated for each subphase of the whole process. Fig. 4 shows the percentages of the energy study carried out. 94% of the energy consumption is required by the productive phases, including tool plunge, dwell time, welding, and tool lift. Welding time required almost half of the overall energy consumption (45%). Dwell time also required a non-negligible amount of energy consumption due to the fact that power reached over time is higher than other subphases. Thus, in terms of both time and energy consumption, welding subphase plays a dominant role. Instead, terms of power values reached, dwell time plays a crucial role.

<figure><img src="image_2610.png" /><figcaption>Figure 4. Energy study of the FSW process (100 mm/min, 1800 rpm)</figcaption></figure>

Similarly, Fig. 5 shows the power recorded over time during the whole process by selecting a tool feed rate of 100 mm/min and a tool rotation of 3200 rpm.
<figure><img src="image_2618.png" /><figcaption>Figure 5. Power consumption profile over time during the FSW process (100 mm/min, 3200 rpm)</figcaption></figure>

Also in this case, power around 150 W is recorded when the machine is switched on. Then, a peak of power of around 3500 W is registered when motors of the machine are turned on. A peak value of 5800 W has been reached when the wool shoulder comes in contact with the top surface of the sheets. Then, power gradually decreases because of the sheet softening induced by the heat produced by the friction (dwell time). When the welding starts, a peak value of 3900 W is reached, very similarly to the power value found by fixing 1800 rpm as tool rotation. Then, it rapidly decreases when an equilibrium is obtained between the heat introduced by the friction forces and the heat dissipated by conduction and convection.

Fig. 6 displays a pie chart of the time study by fixing 100 mm/min and 3200 rpm. According to our results, 79% of the total time is used for productive phases, including tool plunge, dwell time, welding and tool lift. Dwell time required is 34% of the time. The remaining time is required for non-productive phases including auxiliary operations and tool positioning. Very similar percentage results were obtained by observing the time study with 1800 rpm (Fig. 3).

Fig. 7 shows the percentages of the energy study by fixing 100 mm/min and 3200 rpm as process parameters. 79% of the energy consumption is required by the productive phases, including tool plunge, dwell time, welding, and tool lift. Comparing the results obtained in the energy study by using 1800 rpm (Fig. 4), it can be observed a reduction of up to 15% in the energy consumption by increasing the tool rotation by almost up to 50%. Since the duration of all the subphases was very similar among the combinations of process parameters investigated in this study, it can be stated that a lower power was recorded during the productive subphases, especially welding, thus reducing energy consumption. This result can be explained by considering that lower friction forces exist by adopting a higher tool rotation.
<figure><img src="image_2611.png" /><figcaption>Figure 5. Power consumption profile over time during the FSW process (100 mm/min, 3200 rpm)</figcaption></figure>

Also in this case, power around 150 W is recorded when the machine is switched on. Then, a peak of power of around 3500 W is registered when motors of the machine are turned on. A peak value of 5800 W has been reached when the wool shoulder comes in contact with the top surface of the sheets. Then, power gradually decreases because of the sheet softening induced by the heat produced by the friction (dwell time). When the welding starts, a peak value of 3900 W is reached, very similarly to the power value found by fixing 1800 rpm as tool rotation. Then, it rapidly decreases when an equilibrium is obtained between the heat introduced by the friction forces and the heat dissipated by conduction and convection.

Fig. 6 displays a pie chart of the time study by fixing 100 mm/min and 3200 rpm. According to our results, 79% of the total time is used for productive phases, including tool plunge, dwell time, welding and tool lift. Dwell time required is 34% of the time. The remaining time is required for non-productive phases including auxiliary operations and tool positioning. Very similar percentage results were obtained by observing the time study with 1800 rpm (Fig. 3).

Fig. 7 shows the percentages of the energy study by fixing 100 mm/min and 3200 rpm as process parameters. 79% of the energy consumption is required by the productive phases, including tool plunge, dwell time, welding, and tool lift. Comparing the results obtained in the energy study by using 1800 rpm (Fig. 4), it can be observed a reduction of up to 15% in the energy consumption by increasing the tool rotation by almost up to 50%. Since the duration of all the subphases was very similar among the combinations of process parameters investigated in this study, it can be stated that a lower power was recorded during the productive subphases, especially welding, thus reducing energy consumption. This result can be explained by considering that lower friction forces exist by adopting a higher tool rotation.
<figure><img src="image_2619.png" /><figcaption>Figure 6. Time study of the FSW process (100 mm/min, 3200 rpm)</figcaption></figure>

<figure><img src="image_2620.png" /><figcaption>Figure 7. Energy study of the FSW process (100 mm/min, 3200 rpm)</figcaption></figure>

## Conclusions

In this paper, a preliminary investigation of the electrical energy demand during the Friction Stir Welding of two aluminum sheets, AA2024-T3 and AA7075-T6 is presented. Current, voltage and power were recorded overall the process from the machine start to the machine shut down during two combination of process conditions. By following the in-depth approach of the CO2PE! Procedure, time study and energy study were carried out to understand the main contribution of each subphase of the FSW process in terms of both time and energy consumption. For the analyzed process conditions power consumption profile over time during the FSW process showed that the highest power peak was reached when the tool shoulder comes in contact with the top surface of the sheets. Productive subphases requires more time and energy than non-productive subphases.

A lower power was recorded during the productive subphases, especially welding, by using a higher tool rotation because of the lower friction forces, thus reducing energy consumption. Welding contribution in terms of time and energy consumption is a consequence of the speed adopted. Thus, according to these preliminary results, the welding phase should be shortened by
<figure><img src="image_2612.png" /><figcaption>Figure 6. Time study of the FSW process (100 mm/min, 3200 rpm)</figcaption></figure>

<figure><img src="image_2613.png" /><figcaption>Figure 7. Energy study of the FSW process (100 mm/min, 3200 rpm)</figcaption></figure>

## Conclusions

In this paper, a preliminary investigation of the electrical energy demand during the Friction Stir Welding of two aluminum sheets, AA2024-T3 and AA7075-T6 is presented. Current, voltage and power were recorded overall the process from the machine start to the machine shut down during two combination of process conditions. By following the in-depth approach of the CO2PE! Procedure, time study and energy study were carried out to understand the main contribution of each subphase of the FSW process in terms of both time and energy consumption. For the analyzed process conditions power consumption profile over time during the FSW process showed that the highest power peak was reached when the tool shoulder comes in contact with the top surface of the sheets. Productive subphases requires more time and energy than non-productive subphases.

A lower power was recorded during the productive subphases, especially welding, by using a higher tool rotation because of the lower friction forces, thus reducing energy consumption.

Welding contribution in terms of time and energy consumption is a consequence of the speed adopted. Thus, according to these preliminary results, the welding phase should be shortened by
increasing the feed rate as much as possible to reduce the energy consumption in accordance with the mechanical properties to be obtained for the lap joint. Thus, further investigation of this study may regard the variation of the process conditions to understand their effect on the mechanical properties from an energy-saving perspective. Also, an LCA should be carried out to investigate whether and eventually how each single process condition influences the emissions on the environment for a better decision-making.

## References

[1] A. Sartal, R. Bellas, A. M. Mejías, and A. García-Collado, "The sustainable manufacturing concept, evolution and opportunities within Industry 4.0: A literature review," *Adv. Mech. Eng.*, vol. 12, no. 5, 2020. https://doi.org/10.1177/1687814020925232

[2] C. Herrmann, C. Schmidt, D. Kurle, S. Blume, and S. Thiede, "Sustainability in manufacturing and factories of the future," *Int. J. Precis. Eng. Manuf. - Green Technol.*, vol. 1, no. 4, pp. 283-292, 2014. https://doi.org/10.1007/s40684-014-0034-z

[3] E. Cozzolino, V. Lopresto, D. Borrelli, and A. Caraviello, "An integrated approach to investigate the energy consumption for manufacturing and surface finishing 3D printed Inconel 718 parts," *J. Manuf. Process.*, vol. 79, no. December 2021, pp. 193-205, 2022. https://doi.org/10.1016/j.jmapro.2022.04.045

[4] T. N. C. N. A. Van, "Sustainability-based optimization of dissimilar friction stir welding parameters in terms of energy saving, product quality, and cost-effectiveness," *Neural Comput. Appl.*, vol. 35, no. 7, pp. 5221-5249, 2023. https://doi.org/10.1007/s00521-022-07898-8

[5] D. G. Mohan and C. Wu, "A Review on Friction Stir Welding of Steels," *Chinese J. Mech. Eng.*, 2021. https://doi.org/10.1186/s10033-021-00655-3

[6] K. Kellens, W. Dewulf, M. Overcash, M. Z. Hauschild, and J. R. Duflou, "Methodology for systematic analysis and improvement of manufacturing unit process life-cycle inventory (UPLCI) — CO2 PE! initiative (cooperative effort on process emissions in manufacturing). Part 1: Methodology description," pp. 69-78, 2012. https://doi.org/10.1007/s11367-011-0340-4

[7] T. Peng, K. Kellens, R. Tang, C. Chen, and G. Chen, "Sustainability of additive manufacturing: An overview on its energy demand and environmental impact," *Addit. Manuf.*, vol. 21, no. April, pp. 694-704, 2018. https://doi.org/10.1016/j.addma.2018.04.022

[8] E. Cozzolino and A. Astarita, "Energy saving in milling of electron beam - melted Ti6Al4V parts: influence of process parameters," pp. 179-194, 2023.

[9] G. Buffa, D. Campanella, R. Di, and L. Fratini, "Analysis of electrical energy demands in Friction Stir Welding of aluminum alloys," *Procedia Eng.*, vol. 183, pp. 206-212, 2017. https://doi.org/10.1016/j.proeng.2017.04.022
increasing the feed rate as much as possible to reduce the energy consumption in accordance with the mechanical properties to be obtained for the lap joint. Thus, further investigation of this study may regard the variation of the process conditions to understand their effect on the mechanical properties from an energy-saving perspective. Also, an LCA should be carried out to investigate whether and eventually how each single process condition influences the emissions on the environment for a better decision-making.

## References

[1] A. Sartal, R. Bellas, A. M. Mejías, and A. García-Collado, "The sustainable manufacturing concept, evolution and opportunities within Industry 4.0: A literature review," *Adv. Mech. Eng.*, vol. 12, no. 5, 2020. https://doi.org/10.1177/1687814020925232

[2] C. Herrmann, C. Schmidt, D. Kurle, S. Blume, and S. Thiede, "Sustainability in manufacturing and factories of the future," *Int. J. Precis. Eng. Manuf. - Green Technol.*, vol. 1, no. 4, pp. 283-292, 2014. https://doi.org/10.1007/s40684-014-0034-z

[3] E. Cozzolino, V. Lopresto, D. Borrelli, and A. Caraviello, "An integrated approach to investigate the energy consumption for manufacturing and surface finishing 3D printed Inconel 718 parts," *J. Manuf. Process.*, vol. 79, no. December 2021, pp. 193-205, 2022. https://doi.org/10.1016/j.jmapro.2022.04.045

[4] T. N. C. N. A. Van, "Sustainability-based optimization of dissimilar friction stir welding parameters in terms of energy saving, product quality, and cost-effectiveness," *Neural Comput. Appl.*, vol. 35, no. 7, pp. 5221-5249, 2023. https://doi.org/10.1007/s00521-022-07898-8

[5] D. G. Mohan and C. Wu, "A Review on Friction Stir Welding of Steels," *Chinese J. Mech. Eng.*, 2021. https://doi.org/10.1186/s10033-021-00655-3

[6] K. Kellens, W. Dewulf, M. Overcash, M. Z. Hauschild, and J. R. Duflou, "Methodology for systematic analysis and improvement of manufacturing unit process life-cycle inventory (UPLCI) — CO2 PE! initiative (cooperative effort on process emissions in manufacturing). Part 1: Methodology description," pp. 69–78, 2012. https://doi.org/10.1007/s11367-011-0340-4

[7] T. Peng, K. Kellens, R. Tang, C. Chen, and G. Chen, "Sustainability of additive manufacturing: An overview on its energy demand and environmental impact," *Addit. Manuf.*, vol. 21, no. April, pp. 694–704, 2018. https://doi.org/10.1016/j.addma.2018.04.022

[8] E. Cozzolino and A. Astarita, "Energy saving in milling of electron beam - melted Ti6Al4V parts: influence of process parameters," pp. 179-194, 2023.

[9] G. Buffa, D. Campanella, R. Di, and L. Fratini, "Analysis of electrical energy demands in Friction Stir Welding of aluminum alloys," *Procedia Eng.*, vol. 183, pp. 206-212, 2017. https://doi.org/10.1016/j.proeng.2017.04.022
# Innovative control system for straightening machines using sensor information from downstream processes

BATHELT Lukas<fnref n="1" />,ᵃ*, DJAKOW Eugen<fnref n="1" />,ᵇ, HENKE Christian<fnref n="1" />,ᶜ and TRÄCHTLER Ansgar<fnref n="2" />,ᵈ

**Keywords:** Straightening Machine, Punch-Bending Process, Feedback Control

**Abstract.** Increasing the sustainability and resource efficiency of forming processes is one of today's major goals. High-strength wire materials are usually available as strip material and are subjected to a downstream forming process such as punch-bending to produce parts for the electronics industry, for example. During the manufacturing process of the semi-finished product, residual stresses and plastic deformations are introduced into the wire by rolling and drawing processes. Straightening machines are used in the production lines to compensate for these. To increase the sustainability of these production lines, the straightening process is an essential step. Before the continuous manufacturing process starts, the straightening process must be set up and the optimal roller positions must be found. Once the process is set up, the roller position settings are usually not changed. Due to missing measurement systems for the straightening quality, it is not possible to dynamically adjust the positions of the straightening rollers to variations in the material properties. This leads to deviations in the dimensional accuracy of the components to be produced and thus to an increase in the rejection rate in the manufacturing processes. To reduce the rejection rate, a novel control system for a continuous feedback control of a straightening process is presented in this paper. This leads to a reduction of the rejection rate and unnecessary preforming operations in wire straightening process. The result is an increasing sustainability and efficiency of these production process.

**Introduction**

Nowadays, components manufactured in punch-bending machines are an essential part of everyday life. They can be found in various industrial sectors. For example, they are regularly used in the manufacture of terminal blocks in the electrical connectivity industry, where they take on crucial functions [1]. Due to their small size and relatively low unit price, the aspect of sustainability is often neglected, and a high reject rate is accepted, especially when setting up the production processes. This contradicts the concept of sustainability, which is becoming increasingly important in today's world. The aim of this work is to reduce the amount of waste and increase efficiency by applying a novel control approach to adjust straightening machines. The sub-processes straightening and punch-bending are coordinated with each other by exchanging sensor data. In this case, punch bending is an example of a downstream process. An application of the control system for straightening machines in combination with other downstream processes such as progressive processes is also conceivable.

The aim of this paper is to address research questions that deal with the interaction of forming sub-processes. In this case, the focus is specifically on the interactions between the sub-processes of straightening and punch-bending. In a further step, it will be investigated how the straightening process can be adjusted to have a positive effect on the production of punch-bending parts. These

<figure><img src="image_2621.png" /></figure>

<footnote n="1">¹Fraunhofer IEM, Zukunftsmeile 1, 33102 Paderborn Germany</footnote>

<footnote n="2">²Heinz Nixdorf Institute, Paderborn University, Fürstenallee 11, 33102 Paderborn, Germany</footnote>

<footnote>ᵃlukas.bathelt@iem.fraunhofer.de, ᵇeugen.djakow@iem.fraunhofer.de,</footnote>

<footnote>ᶜchristian.henke@iem.fraunhofer.de, ᵈansgar.traechtler@hni.upb.de</footnote>
# Innovative control system for straightening machines using sensor information from downstream processes

BATHELT Lukas<fnref n="1" />,ᵃ*, DJAKOW Eugen<fnref n="1" />,ᵇ, HENKE Christian<fnref n="1" />,ᶜ and TRÄCHTLER Ansgar<fnref n="2" />,ᵈ

**Keywords:** Straightening Machine, Punch-Bending Process, Feedback Control

**Abstract.** Increasing the sustainability and resource efficiency of forming processes is one of today's major goals. High-strength wire materials are usually available as strip material and are subjected to a downstream forming process such as punch-bending to produce parts for the electronics industry, for example. During the manufacturing process of the semi-finished product, residual stresses and plastic deformations are introduced into the wire by rolling and drawing processes. Straightening machines are used in the production lines to compensate for these. To increase the sustainability of these production lines, the straightening process is an essential step. Before the continuous manufacturing process starts, the straightening process must be set up and the optimal roller positions must be found. Once the process is set up, the roller position settings are usually not changed. Due to missing measurement systems for the straightening quality, it is not possible to dynamically adjust the positions of the straightening rollers to variations in the material properties. This leads to deviations in the dimensional accuracy of the components to be produced and thus to an increase in the rejection rate in the manufacturing processes. To reduce the rejection rate, a novel control system for a continuous feedback control of a straightening process is presented in this paper. This leads to a reduction of the rejection rate and unnecessary preforming operations in wire straightening process. The result is an increasing sustainability and efficiency of these production process.

**Introduction**

Nowadays, components manufactured in punch-bending machines are an essential part of everyday life. They can be found in various industrial sectors. For example, they are regularly used in the manufacture of terminal blocks in the electrical connectivity industry, where they take on crucial functions [1]. Due to their small size and relatively low unit price, the aspect of sustainability is often neglected, and a high reject rate is accepted, especially when setting up the production processes. This contradicts the concept of sustainability, which is becoming increasingly important in today's world. The aim of this work is to reduce the amount of waste and increase efficiency by applying a novel control approach to adjust straightening machines. The sub-processes straightening and punch-bending are coordinated with each other by exchanging sensor data. In this case, punch bending is an example of a downstream process. An application of the control system for straightening machines in combination with other downstream processes such as progressive processes is also conceivable.

The aim of this paper is to address research questions that deal with the interaction of forming sub-processes. In this case, the focus is specifically on the interactions between the sub-processes of straightening and punch-bending. In a further step, it will be investigated how the straightening process can be adjusted to have a positive effect on the production of punch-bending parts. These

<figure><img src="image_2614.png" /></figure>

<footnote n="1">¹Fraunhofer IEM, Zukunftsmeile 1, 33102 Paderborn Germany</footnote>

<footnote n="2">²Heinz Nixdorf Institute, Paderborn University, Fürstenallee 11, 33102 Paderborn, Germany</footnote>

<footnote>ᵃlukas.bathelt@iem.fraunhofer.de, ᵇeugen.djakow@iem.fraunhofer.de,</footnote>

<footnote>ᶜchristian.henke@iem.fraunhofer.de, ᵈansgar.traechtler@hni.upb.de</footnote>
findings are intended to enable adjustments in manufacturing processes when fluctuations in material properties occur without having to stop and reset the process.

This paper is structured as follows: First, the state of the art of control approaches for straightening machines is considered. Then the designed system and control structure is presented. On this basis, the dependency between the straightening process and the stamping and bending process is explained. At least, the design of the control algorithm is described and validated on a test rig.

## State of the Art

Industrially used straightening machines are usually adjusted manually or by motor and remain in a constant position during the manufacturing process of one coil. This means that fluctuations in the material parameters occurring during the running process have a negative effect on the straightening result [2]. Several approaches have already been pursued in research to determine a suitable roller setting for the straightening process. These approaches can be divided into two groups: Feedforward and feedback control systems.

Feedforward control approaches often work with simulative or empirical datasets. In [3], a method for the numerical determination of the optimal roller position is developed with the aid of an FE simulation. For a specific parameter configuration consisting of material properties, material dimensions, roller arrangement and dimensions, it was shown that the FE analysis can be used to determine suitable positions for the straightening rollers.

In [4], a concept for the automated adjustment of a five-roller straightening machine is presented. Based on a force measurement in the first bending triangle, the material parameters required to determine the roller positions are determined. The appropriate roller positions are known for many different straightening material properties based on a previously performed simulative parameter study. The parameter study is based on the previously described approach from [3].

In [5], a method for online identification of the straightening properties is presented to be able to react to fluctuations in the material properties. This makes it possible to determine changes in the yield strength and the pre-curvature of the steel strip and to readjust the roller position accordingly. The process capability was confirmed by experimental validation tests.

Further examples of simulation models that are used to set the straightening rollers of straightening machines can be found in [6] and [7], among others. However, what all these approaches have in common is that the straightening quality generated behind the straightener is not verified. This does not ensure that the selected roller positions produce the correct straightening result, especially not during the running manufacturing process.

Control approaches for straightening machines are based on this approach. The process quality is monitored, and corrective action is taken by feeding back sensor information. In [8], a control concept for non-rotating straightening machines with an integrated measuring device for determining the residual curvature is presented. The desired value is the lowest possible bending curvature in the processed material at the output of the straightening machine. The position of the wire is identified at three points behind the straightening machine by using optical sensors. With this data a representing value for the bending curvature is determined. In [9], a concept for controlling a straightening machine using a curvature measurement with capacitive sensors is shown. An experimental validation of both concepts is not yet known.

Another concept for online control of a straightening machine for strip material is presented in [2]. For this purpose, the roller positions are adjusted depending on the measured residual curvature of the straightened material after the straightening process so that the strip is as straight as possible. Validation tests have shown that the curvature measurement is the weak point of this approach.
findings are intended to enable adjustments in manufacturing processes when fluctuations in material properties occur without having to stop and reset the process.

This paper is structured as follows: First, the state of the art of control approaches for straightening machines is considered. Then the designed system and control structure is presented. On this basis, the dependency between the straightening process and the stamping and bending process is explained. At least, the design of the control algorithm is described and validated on a test rig.

## State of the Art

Industrially used straightening machines are usually adjusted manually or by motor and remain in a constant position during the manufacturing process of one coil. This means that fluctuations in the material parameters occurring during the running process have a negative effect on the straightening result [2]. Several approaches have already been pursued in research to determine a suitable roller setting for the straightening process. These approaches can be divided into two groups: Feedforward and feedback control systems.

Feedforward control approaches often work with simulative or empirical datasets. In [3], a method for the numerical determination of the optimal roller position is developed with the aid of an FE simulation. For a specific parameter configuration consisting of material properties, material dimensions, roller arrangement and dimensions, it was shown that the FE analysis can be used to determine suitable positions for the straightening rollers.

In [4], a concept for the automated adjustment of a five-roller straightening machine is presented. Based on a force measurement in the first bending triangle, the material parameters required to determine the roller positions are determined. The appropriate roller positions are known for many different straightening material properties based on a previously performed simulative parameter study. The parameter study is based on the previously described approach from [3].

In [5], a method for online identification of the straightening properties is presented to be able to react to fluctuations in the material properties. This makes it possible to determine changes in the yield strength and the pre-curvature of the steel strip and to readjust the roller position accordingly. The process capability was confirmed by experimental validation tests.

Further examples of simulation models that are used to set the straightening rollers of straightening machines can be found in [6] and [7], among others. However, what all these approaches have in common is that the straightening quality generated behind the straightener is not verified. This does not ensure that the selected roller positions produce the correct straightening result, especially not during the running manufacturing process.

Control approaches for straightening machines are based on this approach. The process quality is monitored, and corrective action is taken by feeding back sensor information. In [8], a control concept for non-rotating straightening machines with an integrated measuring device for determining the residual curvature is presented. The desired value is the lowest possible bending curvature in the processed material at the output of the straightening machine. The position of the wire is identified at three points behind the straightening machine by using optical sensors. With this data a representing value for the bending curvature is determined. In [9], a concept for controlling a straightening machine using a curvature measurement with capacitive sensors is shown. An experimental validation of both concepts is not yet known.

Another concept for online control of a straightening machine for strip material is presented in [2]. For this purpose, the roller positions are adjusted depending on the measured residual curvature of the straightened material after the straightening process so that the strip is as straight as possible. Validation tests have shown that the curvature measurement is the weak point of this approach.
In [10], a concept for controlling a straightening process based on the measurement of the magnetic properties of the material is presented. Considering the Barkhausen effect, a correlation between the magnetic properties of the material and the residual stresses existing in the material is shown. The authors found out that the residual stress distribution is not sufficiently accurate with this type of measurement, which is why the control concept was not implemented. A similar approach was also considered in [11]. Here, the change in the magnetic and electrical properties of flat wire as a function of the bending curvature could be determined by measurement. However, this approach is only applicable under ideal environmental conditions due to the large number of variables, that influence the measured variable.

An evaluation of the state of the art shows that there is a need for a reliable way of evaluating the straightening quality during the continuously running manufacturing process. Furthermore, a control approach needs to be developed that corrects the roller position during the process without pausing it.

## System description and control structure

The dynamic adjustment of the straightening rollers during the production process is desirable in order to be able to react to fluctuations in the material properties of the semi-finished product. Focus is placed on the development of a robust and efficient method that does not require time-consuming or resource-intensive preliminary tests and is suitable for industrial use. To develop such an approach, this paper considers the system shown in Figure 1.

<figure><img src="image_2622.png" /><figcaption>Figure 1: Schematic representation of the entire manufacturing process</figcaption></figure>

The semi-finished product is uncoiled from a coil and fed to the straightening process via guide rollers. This is equipped with sensors to determine the vertical roller positions $z$ and the straightening forces $F$ on the upper three rollers. The semi-finished product is pulled through the straightening machine by a feeding unit. This is followed by a material buffer to make the transition from the continuous straightening process to the discontinuous punch-bending process. The punch-bending process is used as an example for a downstream process in this paper. There are other possible processes like progressive tools which can be used instead. The punch-bending machine consists of a feeding unit and a punch unit. In this application, an L-shaped component is produced. The bending angle $\alpha$ and the straightness $\theta$ of the leg are recorded via a camera. They represent the quality parameters of the manufactured parts and serve as variables to be controlled within a closed-loop control algorithm. The straightness $\theta$ describes the magnitude of the curvature of the wire as indicated by the dotted line in Figure 1. However, the focus is on controlling the bending angle $\alpha$ in this paper. For this reason, the straightness $\theta$ is not described in more detail. A PLC controls all drives and processes the sensor data. The determining variables for the process speed are the coil speed $\omega_C$ upstream of the straightening machine, the feeding speed $\omega_F$ of the feeding unit downstream of the straightening machine, and the cam disk speed $\omega_{CD}$ of the punch-bending
In [10], a concept for controlling a straightening process based on the measurement of the magnetic properties of the material is presented. Considering the Barkhausen effect, a correlation between the magnetic properties of the material and the residual stresses existing in the material is shown. The authors found out that the residual stress distribution is not sufficiently accurate with this type of measurement, which is why the control concept was not implemented. A similar approach was also considered in [11]. Here, the change in the magnetic and electrical properties of flat wire as a function of the bending curvature could be determined by measurement. However, this approach is only applicable under ideal environmental conditions due to the large number of variables, that influence the measured variable.

An evaluation of the state of the art shows that there is a need for a reliable way of evaluating the straightening quality during the continuously running manufacturing process. Furthermore, a control approach needs to be developed that corrects the roller position during the process without pausing it.

## System description and control structure

The dynamic adjustment of the straightening rollers during the production process is desirable in order to be able to react to fluctuations in the material properties of the semi-finished product. Focus is placed on the development of a robust and efficient method that does not require time-consuming or resource-intensive preliminary tests and is suitable for industrial use. To develop such an approach, this paper considers the system shown in Figure 1.

<figure><img src="image_2615.png" /><figcaption>Figure 1: Schematic representation of the entire manufacturing process</figcaption></figure>

The semi-finished product is uncoiled from a coil and fed to the straightening process via guide rollers. This is equipped with sensors to determine the vertical roller positions $z$ and the straightening forces $F$ on the upper three rollers. The semi-finished product is pulled through the straightening machine by a feeding unit. This is followed by a material buffer to make the transition from the continuous straightening process to the discontinuous punch-bending process. The punch-bending process is used as an example for a downstream process in this paper. There are other possible processes like progressive tools which can be used instead. The punch-bending machine consists of a feeding unit and a punch unit. In this application, an L-shaped component is produced. The bending angle $\alpha$ and the straightness $\theta$ of the leg are recorded via a camera. They represent the quality parameters of the manufactured parts and serve as variables to be controlled within a closed-loop control algorithm. The straightness $\theta$ describes the magnitude of the curvature of the wire as indicated by the dotted line in Figure 1. However, the focus is on controlling the bending angle $\alpha$ in this paper. For this reason, the straightness $\theta$ is not described in more detail. A PLC controls all drives and processes the sensor data. The determining variables for the process speed are the coil speed $\omega_C$ upstream of the straightening machine, the feeding speed $\omega_F$ of the feeding unit downstream of the straightening machine, and the cam disk speed $\omega_{CD}$ of the punch-bending
machine. No further control variables are considered in the punch-bending process. The influence of the process speed remains to be investigated. In particular, the tensile force applied to the wire due to feeding can have a significant influence on the straightening process.

The aim of the control approach presented here is the dynamic adjustment of the reference positions $z_R$ for the straightening rollers. On the one hand, this should keep the quality of the straightening process reproducible at a high level. On the other hand, it should be possible to react to fluctuations in the material properties with a dynamic adjustment of the straightening rollers in the process. However, as the flat wire is fed directly into the punch bending process, it is not possible to measure the residual curvature directly after the straightening machine. For this reason, the control approach contains three subsystems: the setup assistant, disturbance compensation and feedback control, as shown in Figure 2. The initial positions $z_R(t_0)$ of the straightening rollers are determined before the start of the manufacturing process with the aid of a setup assistant. It was presented in [12]. Due to the fluctuations in the material properties, a correction of the initial positions is necessary during the running process. The change in the pre-curvature $\kappa_0(t)$ over the coil radius $r_c$ is interpreted as a deterministic disturbance variable. As part of a disturbance compensation, the trajectory $\tilde{\kappa}_0(t)$ of the pre-curvature is estimated and an adjustment $\Delta z_z$ is calculated using a plant model. A detailed description of disturbance compensation can be found in [13]. The plant model is shown in [14].

<figure><img src="image_2623.png" /><figcaption>Figure 2: Schematic representation of the overall control structure</figcaption></figure>

In order to compensate for model inaccuracies and stochastic disturbance variables such as width and thickness variations or the material composition, feedback control is used as a third subsystem. For this purpose, the geometric properties of the manufactured component are measured after the bending operation in the punch bending machine. In this paper, the component angle $\alpha$ is considered. In case of deviations from its reference value $\alpha_R$, a second adjustment $\Delta z_{SB}$ for the roller positions is determined. This control approach is based on the two-degree-of-freedom structure (cf. [15]). The reason for using disturbance compensation in parallel with feedback control is the large dead time between the straightening machine and the punch-bending machine. The final reference position $z_R$ of the straightening rollers thus results from the addition of the three components:

$$z_R(t) = z_R(t_0) + \Delta z_z(t) + \Delta z_{SB}(t) \quad (1)$$

The reference positions $z_R$ are adjusted with the aid of a subordinate position control. The straightening rollers thus act on the incoming wire in the first part of the control plant. After the straightening operation a residual curvature $\kappa_R$ remains in the wire. In the second part of the control
machine. No further control variables are considered in the punch-bending process. The influence of the process speed remains to be investigated. In particular, the tensile force applied to the wire due to feeding can have a significant influence on the straightening process.

The aim of the control approach presented here is the dynamic adjustment of the reference positions $z_R$ for the straightening rollers. On the one hand, this should keep the quality of the straightening process reproducible at a high level. On the other hand, it should be possible to react to fluctuations in the material properties with a dynamic adjustment of the straightening rollers in the process. However, as the flat wire is fed directly into the punch bending process, it is not possible to measure the residual curvature directly after the straightening machine. For this reason, the control approach contains three subsystems: the setup assistant, disturbance compensation and feedback control, as shown in Figure 2. The initial positions $z_R(t_0)$ of the straightening rollers are determined before the start of the manufacturing process with the aid of a setup assistant. It was presented in [12]. Due to the fluctuations in the material properties, a correction of the initial positions is necessary during the running process. The change in the pre-curvature $\kappa_0(t)$ over the coil radius $r_c$ is interpreted as a deterministic disturbance variable. As part of a disturbance compensation, the trajectory $\tilde{\kappa}_0(t)$ of the pre-curvature is estimated and an adjustment $\Delta z_Z$ is calculated using a plant model. A detailed description of disturbance compensation can be found in [13]. The plant model is shown in [14].

<figure><img src="image_2616.png" /><figcaption>Figure 2: Schematic representation of the overall control structure</figcaption></figure>

In order to compensate for model inaccuracies and stochastic disturbance variables such as width and thickness variations or the material composition, feedback control is used as a third subsystem. For this purpose, the geometric properties of the manufactured component are measured after the bending operation in the punch bending machine. In this paper, the component angle $\alpha$ is considered. In case of deviations from its reference value $\alpha_R$, a second adjustment $\Delta z_{SB}$ for the roller positions is determined. This control approach is based on the two-degree-of-freedom structure (cf. [15]). The reason for using disturbance compensation in parallel with feedback control is the large dead time between the straightening machine and the punch-bending machine. The final reference position $z_R$ of the straightening rollers thus results from the addition of the three components:

$$z_R(t) = z_R(t_0) + \Delta z_Z(t) + \Delta z_{SB}(t) \quad (1)$$

The reference positions $z_R$ are adjusted with the aid of a subordinate position control. The straightening rollers thus act on the incoming wire in the first part of the control plant. After the straightening operation a residual curvature $\kappa_R$ remains in the wire. In the second part of the control
plant, the flat wire is formed to components in the punch bending machine. The dimensional accuracy of the components therefore depends directly on the setting of the straightening machine.

This description of the system under test and of the control structure contains the configuration of all sensor information currently available at the test rig. The focus of this paper is on the presentation of the control algorithm. To influence the residual curvature of the wire when it leaves the straightening machine, it is sufficient to control only the position of the straightening roller in the last bending triangle. Therefore, from now on a straightening machine with only one straightening triangle and one variable straightening roller position $z_R$ is considered.

## Analysis of the punch-bending process and controller design

To develop and design a suitable control algorithm, it is first necessary to investigate the dependencies between the straightening machine and the downstream process. A single-stage punch bending machine is considered in this paper and schematically shown in Figure 3 (a). The flat wire is fed into the forming zone via a feeder. There, the downholder holds the flat wire in a fixed position while the bending punch performs the forming operation. At the same time the cutting punch separates the component from the rest of the flat wire. The bending and cutting punches are moved by a cam disk driven by an electric motor. The formed component is characterized by the bending angle $\alpha$. A camera was installed on the test bench to record this during the process. The setup is shown in Figure 3 (b). The camera with lens is positioned directly in front of the forming zone of the stamping and bending machine. A backlight LED makes it easier for the camera to identify the component. The image is captured at a specific time after the component has been formed. The camera software is used to determine the bending angle $\alpha$ and transfer it to the PLC.

<figure><img src="image_2624.png" /><figcaption>Figure 3: Representation of the Punch Bending process; (a) Schematic representation; (b) Photographic representation of the forming zone</figcaption></figure>

The effect of the straightening process on the bending angle $\alpha$ of the manufactured component is now to be investigated for the design of the control system. For this purpose, samples with different curvatures are produced in the straightening machine and fed into the punch bending machine. The investigations are carried out on flat wire made of high-strength steel (material designation 1.4310) with the dimensions 3.9 mm x 0.4 mm (width x thickness). Samples were produced with a curvature of $B_{Rest} = \pm 25$ mm / 200 mm. The gradations in the range $B_{Rest} < 0$ were selected finer. After evaluating these samples, the area $B_{Rest} > 0$ was divided more coarsely to confirm the trend. The value $B_{Rest}$ is a value to quantify the bending curvature $\kappa$ and can be calculated using the following correlation:
plant, the flat wire is formed to components in the punch bending machine. The dimensional accuracy of the components therefore depends directly on the setting of the straightening machine.

This description of the system under test and of the control structure contains the configuration of all sensor information currently available at the test rig. The focus of this paper is on the presentation of the control algorithm. To influence the residual curvature of the wire when it leaves the straightening machine, it is sufficient to control only the position of the straightening roller in the last bending triangle. Therefore, from now on a straightening machine with only one straightening triangle and one variable straightening roller position $z_R$ is considered.

## Analysis of the punch-bending process and controller design

To develop and design a suitable control algorithm, it is first necessary to investigate the dependencies between the straightening machine and the downstream process. A single-stage punch bending machine is considered in this paper and schematically shown in Figure 3 (a). The flat wire is fed into the forming zone via a feeder. There, the downholder holds the flat wire in a fixed position while the bending punch performs the forming operation. At the same time the cutting punch separates the component from the rest of the flat wire. The bending and cutting punches are moved by a cam disk driven by an electric motor. The formed component is characterized by the bending angle $\alpha$. A camera was installed on the test bench to record this during the process. The setup is shown in Figure 3 (b). The camera with lens is positioned directly in front of the forming zone of the stamping and bending machine. A backlight LED makes it easier for the camera to identify the component. The image is captured at a specific time after the component has been formed. The camera software is used to determine the bending angle $\alpha$ and transfer it to the PLC.

<figure><img src="image_2617.png" /><figcaption>Figure 3: Representation of the Punch Bending process; (a) Schematic representation; (b) Photographic representation of the forming zone</figcaption></figure>

The effect of the straightening process on the bending angle $\alpha$ of the manufactured component is now to be investigated for the design of the control system. For this purpose, samples with different curvatures are produced in the straightening machine and fed into the punch bending machine. The investigations are carried out on flat wire made of high-strength steel (material designation 1.4310) with the dimensions 3.9 mm x 0.4 mm (width x thickness). Samples were produced with a curvature of $B_{Rest} = \pm 25$ mm / 200 mm. The gradations in the range $B_{Rest} < 0$ were selected finer. After evaluating these samples, the area $B_{Rest} > 0$ was divided more coarsely to confirm the trend. The value $B_{Rest}$ is a value to quantify the bending curvature $\kappa$ and can be calculated using the following correlation:
$$R = \frac{1}{\kappa} = \frac{B_{Rest}}{2} + \frac{L^2}{8 \cdot B_{Rest}} \qquad (2)$$

R is the radius of curvature and L is the reference length for the characteristic value $B_{Rest}$. A reference length of L = 200 mm is used in this paper.

Approx. 200 components were manufactured for each curvature. The result of this experimental study is shown in Figure 4. In Plot (a), an example of the component angle for the curvature $B_{Rest} = -0.53$ mm / 200 mm over the number of components produced is shown. 80 percent of all components are within a band of $\pm 0.1$ degrees around the mean value. Except for four outliers, the remaining 20 percent of the components are close to this band. To evaluate the test, the mean value of all manufactured components was saved together with the curvature of the flat wire. After completion of all the tests, all the mean values were plotted over their curvature value. The result is shown in Figure 4 (b). There is a direct correlation between the curvature of the wire and the component angle. The component angle $\alpha$ increases when the curvature $B_{Rest}$ of the wire decreases and the other way round. This correlation can be described by a linear function:

$$\alpha(B_{Rest}) = -0.085 \cdot B_{Rest} + 55.085 \qquad (3)$$

The root mean squared error (RMSE) was calculated to determine the accuracy of the linear approximation. This is RMSE = 0.166, which indicates a clear correlation.

<figure><img src="image_2625.png" /><figcaption>Figure 4: (a) Behavior of the bending angle $\alpha$ with constant wire curvature $B_{Rest} = -0.53$ mm / 200 mm; (b) Behavior of the bending angle $\alpha$ with variable wire curvature $B_{Rest}$</figcaption></figure>

The control design is described based on the previous investigations. In Figure 5, the control algorithm designed for this purpose is shown schematically. The measurement signal of the component angle is first processed. A moving average filter is used to reduce the influence of outliers on the control variable. Therefore, a filter window size of 10 measured angles is used. After that, the averaged component angle $\bar{\alpha}$ is compared with the reference value $\alpha_R$ and the control difference $e$ is fed to a PI controller. The calculated actuator variable $u$ is limited, as the position of the straightening roller can only be adjusted within certain limits. To prevent the windup effect, an anti-windup element is also used.
$$R = \frac{1}{\kappa} = \frac{B_{Rest}}{2} + \frac{L^2}{8 \cdot B_{Rest}} \qquad (2)$$

R is the radius of curvature and L is the reference length for the characteristic value $B_{Rest}$. A reference length of L = 200 mm is used in this paper.

Approx. 200 components were manufactured for each curvature. The result of this experimental study is shown in Figure 4. In Plot (a), an example of the component angle for the curvature $B_{Rest} = -0.53$ mm / 200 mm over the number of components produced is shown. 80 percent of all components are within a band of ±0.1 degrees around the mean value. Except for four outliers, the remaining 20 percent of the components are close to this band. To evaluate the test, the mean value of all manufactured components was saved together with the curvature of the flat wire. After completion of all the tests, all the mean values were plotted over their curvature value. The result is shown in Figure 4 (b). There is a direct correlation between the curvature of the wire and the component angle. The component angle $\alpha$ increases when the curvature $B_{Rest}$ of the wire decreases and the other way round. This correlation can be described by a linear function:

$$\alpha(B_{Rest}) = -0.085 \cdot B_{Rest} + 55.085 \qquad (3)$$

The root mean squared error (RMSE) was calculated to determine the accuracy of the linear approximation. This is RMSE = 0.166, which indicates a clear correlation.

<figure><img src="image_2618.png" /><figcaption>Figure 4: (a) Behavior of the bending angle $\alpha$ with constant wire curvature $B_{Rest} = -0.53$ mm / 200 mm; (b) Behavior of the bending angle $\alpha$ with variable wire curvature $B_{Rest}$</figcaption></figure>

The control design is described based on the previous investigations. In Figure 5, the control algorithm designed for this purpose is shown schematically. The measurement signal of the component angle is first processed. A moving average filter is used to reduce the influence of outliers on the control variable. Therefore, a filter window size of 10 measured angles is used. After that, the averaged component angle $\bar{\alpha}$ is compared with the reference value $\alpha_R$ and the control difference $e$ is fed to a PI controller. The calculated actuator variable $u$ is limited, as the position of the straightening roller can only be adjusted within certain limits. To prevent the windup effect, an anti-windup element is also used.
<figure><img src="image_2626.png" /><figcaption>Figure 5: Detailed description of the control algorithm</figcaption></figure>

The values of the PI controller are chosen to $K_p = 0.01$ and $K_i = 0.0025$. The time constant of the anti-windup element is set to twice the value of $K_i$ to $K_w = 0.005$. The actuator variable limit is set depending on the current operating point, which is previously determined by the setup assistant (cf. Figure 1). In the configuration considered here, a range of $-4 < z_R < 0$ is available as the absolute position for the straightening roller. This range is limited by design, among other things.

## Validation

After the controller has been designed, proof of its functionality is to be provided as part of the validation process. For this purpose, the test bench described above in Figure 1 is fully set up. The dead time in the current configuration is 34 components due to the distance between the straightening machine and the punch bending machine at the test rig. The distance can be significantly reduced in machines on the market. To verify the function of the control system only one straightening triangle is used. Depending on the straightening strategy and application, the use of a five- or seven-roller straightening machine is suitable. However, the residual curvature in the flat wire after the straightening machine is particularly influenced by the configuration of the last straightening triangle. For this reason, only one straightening triangle will be considered in this paper. The straightening roller position is initially set to the value $z_R(t_0 = 0) = -2$ mm. This results in a component angle of $\alpha_R = 58.5$ degrees. At time $t_1 = 5$ s, a step to $\alpha_R = 59$ degrees is specified. In Figure 6, the step response and the adjustment $\Delta z_{SB}$ of the straightening roller at a process speed of 20 components per minute is shown.

Due to the dead time between the straightening machine and the punch bending machine, the reaction of the component angle $\alpha$ to the change in straightening roller position can only be recognized after a certain amount of time. Nevertheless, the control system can follow the step of $\alpha_R$. Due to the long dead time, the controller parameters were set very low, which means that in this test configuration it takes 80 components for the control error to be reduced. The tolerance band of $\pm0.3$ degrees around the reference value is already reached after 70 components. As already mentioned, the number of components will be significantly lower due to the shorter distance between the straightening machine and the punch bending machine on machines on the market.

The current work is concerned with optimizing the controller parameters and implementing a Smith predictor for dead time compensation. This is expected to result in a significant improvement in the performance of the control system.
<figure><img src="image_2619.png" /><figcaption>Figure 5: Detailed description of the control algorithm</figcaption></figure>

The values of the PI controller are chosen to $K_p = 0.01$ and $K_i = 0.0025$. The time constant of the anti-windup element is set to twice the value of $K_i$ to $K_w = 0.005$. The actuator variable limit is set depending on the current operating point, which is previously determined by the setup assistant (cf. Figure 1). In the configuration considered here, a range of $-4 < z_R < 0$ is available as the absolute position for the straightening roller. This range is limited by design, among other things.

## Validation

After the controller has been designed, proof of its functionality is to be provided as part of the validation process. For this purpose, the test bench described above in Figure 1 is fully set up. The dead time in the current configuration is 34 components due to the distance between the straightening machine and the punch bending machine at the test rig. The distance can be significantly reduced in machines on the market. To verify the function of the control system only one straightening triangle is used. Depending on the straightening strategy and application, the use of a five- or seven-roller straightening machine is suitable. However, the residual curvature in the flat wire after the straightening machine is particularly influenced by the configuration of the last straightening triangle. For this reason, only one straightening triangle will be considered in this paper. The straightening roller position is initially set to the value $z_R(t_0 = 0) = -2$ mm. This results in a component angle of $\alpha_R = 58.5$ degrees. At time $t_1 = 5$ s, a step to $\alpha_R = 59$ degrees is specified. In Figure 6, the step response and the adjustment $\Delta z_{SB}$ of the straightening roller at a process speed of 20 components per minute is shown.

Due to the dead time between the straightening machine and the punch bending machine, the reaction of the component angle $\alpha$ to the change in straightening roller position can only be recognized after a certain amount of time. Nevertheless, the control system can follow the step of $\alpha_R$. Due to the long dead time, the controller parameters were set very low, which means that in this test configuration it takes 80 components for the control error to be reduced. The tolerance band of $\pm0.3$ degrees around the reference value is already reached after 70 components. As already mentioned, the number of components will be significantly lower due to the shorter distance between the straightening machine and the punch bending machine on machines on the market.

The current work is concerned with optimizing the controller parameters and implementing a Smith predictor for dead time compensation. This is expected to result in a significant improvement in the performance of the control system.
<figure><img src="image_2627.png" /><figcaption>Figure 6: Step response of the bending angle $\alpha$ (a) and adjustment $\Delta z_{SB}$ of the straightening roller position (b) of the designed control system</figcaption></figure>

## Conclusion

In this paper a novel control algorithm is presented, that enables online adjustment for a straightening machine in combination with a downstream process. The straightening quality is evaluated during the running process and the settings of the straightening machine can be adjusted without stopping and resetting the whole manufacturing process. Adjusting the straightening process online enables the possibility of a demand-oriented straightening strategy. Also, the amount of waste caused by re-setting the machine is reduced to its minimum.

This method offers new possibilities for an efficient and resource-saving manufacturing process. It reduces the number of straightening rollers compared to conventional methods and avoids overbending by straightening the semi-finished product as needed. Also, the material's strength properties are preserved by using a minimum number of straightening rollers. This improves the sustainability in two ways: firstly, by allowing for adjustments to material fluctuations and reducing rejection rates. Secondly, the number and complexity of time-consuming set-up processes is reduced to a minimum. This also minimizes the amount of waste that would otherwise be produced when setting up the machines. The combination of these aspects in one manufacturing process promotes responsible and careful resource usage. Currently, there is an estimated potential for material savings of up to 11 % due to shorter setup processes and online adjustments, but further research is needed to validate this control concept.

## Acknowledgement

The research project IFG 22114 N / P1565 "Development of an innovative straightening machine for three-dimensional straightening and forming" from the Research Association for steel Application (FOSTA), Düsseldorf, is supported by the Federal Ministry of Economic Affairs and Climate Action the German Federation of Industrial Research Associations (AiF) as part of the programme for promoting industrial cooperative research (IGF) on the basis of a decision by the German Bundestag. The project is carried out at Fraunhofer IEM and Forming and Machining Technology (LUF), Paderborn University.

## References

[1] Hörmann, V.; Schäfer, C., Stanzbiegetechnik. Effiziente Fertigung von Stanzbiegeteilen und kompletten Baugruppen. Landsberg: Verl. Moderne Industrie, 2013

[2] Behrens, B.-A.; Krimm, R., Controlled Sheet Metal Straightening, TMS 2008, 137th Annual Meeting & Exhibition, New Orleans, 2008
<figure><img src="image_2620.png" /></figure>

<figure><img src="image_2621.png" /><figcaption>Figure 6: Step response of the bending angle $\alpha$ (a) and adjustment $\Delta z_{SB}$ of the straightening roller position (b) of the designed control system</figcaption></figure>

## Conclusion

In this paper a novel control algorithm is presented, that enables online adjustment for a straightening machine in combination with a downstream process. The straightening quality is evaluated during the running process and the settings of the straightening machine can be adjusted without stopping and resetting the whole manufacturing process. Adjusting the straightening process online enables the possibility of a demand-oriented straightening strategy. Also, the amount of waste caused by re-setting the machine is reduced to its minimum.

This method offers new possibilities for an efficient and resource-saving manufacturing process. It reduces the number of straightening rollers compared to conventional methods and avoids overbending by straightening the semi-finished product as needed. Also, the material's strength properties are preserved by using a minimum number of straightening rollers. This improves the sustainability in two ways: firstly, by allowing for adjustments to material fluctuations and reducing rejection rates. Secondly, the number and complexity of time-consuming set-up processes is reduced to a minimum. This also minimizes the amount of waste that would otherwise be produced when setting up the machines. The combination of these aspects in one manufacturing process promotes responsible and careful resource usage. Currently, there is an estimated potential for material savings of up to 11 % due to shorter setup processes and online adjustments, but further research is needed to validate this control concept.

## Acknowledgement

The research project IFG 22114 N / P1565 "Development of an innovative straightening machine for three-dimensional straightening and forming" from the Research Association for steel Application (FOSTA), Düsseldorf, is supported by the Federal Ministry of Economic Affairs and Climate Action the German Federation of Industrial Research Associations (AiF) as part of the programme for promoting industrial cooperative research (IGF) on the basis of a decision by the German Bundestag. The project is carried out at Fraunhofer IEM and Forming and Machining Technology (LUF), Paderborn University.

## References

[1] Hörmann, V.; Schäfer, C., Stanzbiegetechnik. Effiziente Fertigung von Stanzbiegeteilen und kompletten Baugruppen. Landsberg: Verl. Moderne Industrie, 2013

[2] Behrens, B.-A.; Krimm, R., Controlled Sheet Metal Straightening, TMS 2008, 137th Annual Meeting & Exhibition, New Orleans, 2008
[3] Oligschläger, M.; Hirt, G.: Implementation of Closed-loop Control Systems in Finite Element Simulations for Roller Leveling. Matériaux et techniques, Vol. 100, Iss. Special Issue, 2012, S. 1-14

[4] Grüber, M.; Oligschläger, M.; Hirt, G. Adjusting of Roller Levellers by Finite Element Simulations Including a Closed-Loop Control. Adv. Mater. Res. 2014, 1018, 207–214. https://doi.org/10.4028/www.scientific.net/AMR.1018.207

[5] Grüber, M., Hirt, G., Numerical Investigation of a Process Control for the Roller Levelling Process Based on a Force Measurement. MSF, 2016. https://doi.org/10.4028/www.scientific.net/msf.854.249

[6] Schiffner, K.; Henrich, L., A Model for Simulation and Optimization of the Metal Forming Process at Roller Levelling. In Discretization Methods and Structural Optimization—Procedures and Applications, Proceedings of the GAMM-Seminar, Siegen, FRG, 5–7 October 1988; Eschenauer, H., Thierauf, G., Eds.; Springer: Berlin/Heidelberg, Germany, 1989; pp. 317–324, ISBN 978-3-540-50784-0.

[7] Menz, R., Entwicklung eines Analytischen Simulationsmodells Als Grundlage Einer Geregelten Richtmaschine; VDI-Verlag: Düsseldorf, Germany, 2002; ISBN 3-18-360102-8.

[8] Post, F., Nichtrotierender Richtapparat für Biegemaschinen mit integrierter Messvorrichtung. 1995

[9] Haberland, R.; Lauer, G., Sensorik für die geregelte Blechrichtmaschine. Bleche Rohe Profile (1993), Iss. 40, 1993, p. 599-601

[10] Ashton, D.; Deigelmann, V.; Stolzenberg, M., Wolter, B., Closed loop automatic shape and residual stress control during levelling - Final report. Off. for Official Publ. of the European Communities, Vol. 22824, Luxembourg, 2007

[11] Trächtler, A.; Homberg, W., Development of an Intelligent Setup Assistance System for Straightening Units to Extend the Process Limits of Processing High-Strength Wire Materials: Final Report IGF Project No. 19799 N; Fosta e.V.: Düsseldorf, Germany, 2021; ISBN 978-3-96780-052-4.

[12] Bathelt, L.; Bader, F.; Djakow, E.; Henke, C.; Trächtler, A.; Homberg, W., Innovative Assistance System for Setting up a Mechatronic Straightening Machine. Key Eng. Mater. 2022, 926, 2397–2405. https://doi.org/10.4028/p-vs07w9

[13] BATHELT Lukas, DJAKOW Eugen, HENKE Christian, TRÄCHTLER Ansgar, Innovative self-learning disturbance compensation for straightening processes, Materials Research Proceedings, Vol. 28, pp 2005-2014, 2023. https://doi.org/10.21741/9781644902479-216

[14] Bathelt, L.; Scurk, M.; Djakow, E.; Henke, C.; Trächtler, A., Novel Straightening-Machine Design with Integrated Force Measurement for Straightening of High-Strength Flat Wire. Sensors 2023, 23, 9091. https://doi.org/10.3390/s23229091

[15] Horowitz, I. Synthesis of Feedback Systems; Elsevier: Amsterdam, The Netherlands, 1963; ISBN 9781483232829.
[3] Oligschläger, M.; Hirt, G.: Implementation of Closed-loop Control Systems in Finite Element Simulations for Roller Leveling. Matériaux et techniques, Vol. 100, Iss. Special Issue, 2012, S. 1-14

[4] Grüber, M.; Oligschläger, M.; Hirt, G. Adjusting of Roller Levellers by Finite Element Simulations Including a Closed-Loop Control. Adv. Mater. Res. 2014, 1018, 207–214. https://doi.org/10.4028/www.scientific.net/AMR.1018.207

[5] Grüber, M., Hirt, G., Numerical Investigation of a Process Control for the Roller Levelling Process Based on a Force Measurement. MSF, 2016. https://doi.org/10.4028/www.scientific.net/msf.854.249

[6] Schiffner, K.; Henrich, L., A Model for Simulation and Optimization of the Metal Forming Process at Roller Levelling. In Discretization Methods and Structural Optimization—Procedures and Applications, Proceedings of the GAMM-Seminar, Siegen, FRG, 5–7 October 1988; Eschenauer, H., Thierauf, G., Eds.; Springer: Berlin/Heidelberg, Germany, 1989; pp. 317–324, ISBN 978-3-540-50784-0.

[7] Menz, R., Entwicklung eines Analytischen Simulationsmodells Als Grundlage Einer Geregelten Richtmaschine; VDI-Verlag: Düsseldorf, Germany, 2002; ISBN 3-18-360102-8.

[8] Post, F., Nichtrotierender Richtapparat für Biegemaschinen mit integrierter Messvorrichtung. 1995

[9] Haberland, R.; Lauer, G., Sensorik für die geregelte Blechrichtmaschine. Bleche Rohe Profile (1993), Iss. 40, 1993, p. 599-601

[10] Ashton, D.; Deigelmann, V.; Stolzenberg, M., Wolter, B., Closed loop automatic shape and residual stress control during levelling - Final report. Off. for Official Publ. of the European Communities, Vol. 22824, Luxembourg, 2007

[11] Trächtler, A.; Homberg, W., Development of an Intelligent Setup Assistance System for Straightening Units to Extend the Process Limits of Processing High-Strength Wire Materials: Final Report IGF Project No. 19799 N; Fosta e.V.: Düsseldorf, Germany, 2021; ISBN 978-3-96780-052-4.

[12] Bathelt, L.; Bader, F.; Djakow, E.; Henke, C.; Trächtler, A.; Homberg, W., Innovative Assistance System for Setting up a Mechatronic Straightening Machine. Key Eng. Mater. 2022, 926, 2397–2405. https://doi.org/10.4028/p-vs07w9

[13] BATHELT Lukas, DJAKOW Eugen, HENKE Christian, TRÄCHTLER Ansgar, Innovative self-learning disturbance compensation for straightening processes, Materials Research Proceedings, Vol. 28, pp 2005-2014, 2023. https://doi.org/10.21741/9781644902479-216

[14] Bathelt, L.; Scurk, M.; Djakow, E.; Henke, C.; Trächtler, A., Novel Straightening-Machine Design with Integrated Force Measurement for Straightening of High-Strength Flat Wire. Sensors 2023, 23, 9091. https://doi.org/10.3390/s23229091

[15] Horowitz, I. Synthesis of Feedback Systems; Elsevier: Amsterdam, The Netherlands, 1963; ISBN 9781483232829.
# Friction stir-assisted cladding: Solid-state recycling of machine shop scrap for sustainable metal production

KAUSHIK Pankaj¹,ᵃ* and DATE Prashant P.¹,ᵇ

¹Department of Mechanical Engineering IIT Bombay, Powai, India-400076

ᵃpkaushik4@iitb.ac.in, ᵇppdate@iitb.ac.in

**Keywords:** Circular Economy, Recycling, Friction Stir-Assisted Cladding, Process Development

**Abstract.** The manufacturing industry, a cornerstone of modern economies, generates a substantial amount of waste in the form of manufacturing chips, which often end up as discarded byproducts. This waste not only poses environmental hazards but also represents a missed opportunity for resource efficiency and cost savings. This article explores the innovative use of friction stir-assisted cladding (FS-AC) as a sustainable and cost-effective solution for repurposing aluminum chips on a steel substrate without melting the material. The study investigates the mechanism of the FS-AC process, highlighting the conversion of metallic swarf into a feedstock material for cladding without entailing costly clad materials. Metallographic analysis demonstrates a uniformly clad layer with an average thickness of 0.95 mm. The study addresses challenges related to bonding at the clad layer-substrate interface and proposes further advanced characterization studies for diffusion analysis and coating-substrate bond strength. In conclusion, the FS-AC process presents a valuable method for recycling machine shop floor scrap, transforming it into useful feedstock material for cladding purposes, and that too in an energy-efficient manner. This proposed approach has promising potential for processing different material combinations for additive manufacturing and exploring functionally graded material development using metallic swarf.

**Introduction**

Industries use metals acquired from primary and secondary sources. Secondary sources encompass metals that have fulfilled their initial purpose and entered the economic cycle. The manufacturing industry is a cornerstone of modern economies, but it generates a substantial amount of waste in the form of manufacturing chips, which often end up as discarded byproducts. This waste not only represents an environmental hazard but also a missed opportunity for resource efficiency and cost savings. Emphasizing secondary metal production provides significant environmental benefits, such as conserving energy, landscapes, and natural resources while also reducing toxic and non-toxic waste streams [1]. Addressing this challenge is critical for achieving sustainability goals and maintaining the industry's competitiveness.

Diverse technologies play a vital role in recovering and processing metals from waste streams, impacting the quantity of reintroduced secondary metals into the system. Unlike polymers and plastics, metals can be recycled nearly endlessly, with their properties fully restored, albeit not always with ease, regardless of their chemical or physical form.

Despite the abundance of metallic scrap, distinguishing between new and old scrap is crucial. New scrap, generated within mills or factories, maintains a well-known and homogeneous quality, unlike older, potentially mixed scrap. This distinction is particularly relevant in industrial settings, where the quality, especially chemical composition, significantly influences the recycling process.

The extensive use of aluminum in contemporary industrial applications underscores the essential nature of recovering manufacturing scrap of aluminum due to cost implications. Producing aluminum from ore demands approximately 20 times more energy than utilizing recycled metal. In contrast to traditional scrap recovery, the primary objective in aluminum scrap

<figure><img src="image_2628.png" /></figure>
# Friction stir-assisted cladding: Solid-state recycling of machine shop scrap for sustainable metal production

KAUSHIK Pankaj¹,ᵃ* and DATE Prashant P.¹,ᵇ

¹Department of Mechanical Engineering IIT Bombay, Powai, India-400076

ᵃpkaushik4@iitb.ac.in, ᵇppdate@iitb.ac.in

**Keywords:** Circular Economy, Recycling, Friction Stir-Assisted Cladding, Process Development

**Abstract.** The manufacturing industry, a cornerstone of modern economies, generates a substantial amount of waste in the form of manufacturing chips, which often end up as discarded byproducts. This waste not only poses environmental hazards but also represents a missed opportunity for resource efficiency and cost savings. This article explores the innovative use of friction stir-assisted cladding (FS-AC) as a sustainable and cost-effective solution for repurposing aluminum chips on a steel substrate without melting the material. The study investigates the mechanism of the FS-AC process, highlighting the conversion of metallic swarf into a feedstock material for cladding without entailing costly clad materials. Metallographic analysis demonstrates a uniformly clad layer with an average thickness of 0.95 mm. The study addresses challenges related to bonding at the clad layer-substrate interface and proposes further advanced characterization studies for diffusion analysis and coating-substrate bond strength. In conclusion, the FS-AC process presents a valuable method for recycling machine shop floor scrap, transforming it into useful feedstock material for cladding purposes, and that too in an energy-efficient manner. This proposed approach has promising potential for processing different material combinations for additive manufacturing and exploring functionally graded material development using metallic swarf.

**Introduction**

Industries use metals acquired from primary and secondary sources. Secondary sources encompass metals that have fulfilled their initial purpose and entered the economic cycle. The manufacturing industry is a cornerstone of modern economies, but it generates a substantial amount of waste in the form of manufacturing chips, which often end up as discarded byproducts. This waste not only represents an environmental hazard but also a missed opportunity for resource efficiency and cost savings. Emphasizing secondary metal production provides significant environmental benefits, such as conserving energy, landscapes, and natural resources while also reducing toxic and non-toxic waste streams [1]. Addressing this challenge is critical for achieving sustainability goals and maintaining the industry's competitiveness.

Diverse technologies play a vital role in recovering and processing metals from waste streams, impacting the quantity of reintroduced secondary metals into the system. Unlike polymers and plastics, metals can be recycled nearly endlessly, with their properties fully restored, albeit not always with ease, regardless of their chemical or physical form.

Despite the abundance of metallic scrap, distinguishing between new and old scrap is crucial. New scrap, generated within mills or factories, maintains a well-known and homogeneous quality, unlike older, potentially mixed scrap. This distinction is particularly relevant in industrial settings, where the quality, especially chemical composition, significantly influences the recycling process.

The extensive use of aluminum in contemporary industrial applications underscores the essential nature of recovering manufacturing scrap of aluminum due to cost implications. Producing aluminum from ore demands approximately 20 times more energy than utilizing recycled metal. In contrast to traditional scrap recovery, the primary objective in aluminum scrap

<figure><img src="image_2622.png" /></figure>
recovery is not only to minimize disposal costs but also to evade regulatory liability by extracting metals from voluminous wastes. Aluminum scrap is typically processed in a reverberatory furnace as part of the recycling process. However, melting the scrap is an energy-intensive process. Also, the inherent properties of aluminum alloys generally get lost during melting due to the evaporation of various strength-imparting elements. The solid-state recycling of aluminum alloys will result in huge energy savings while keeping the properties intact. The solid-state recycling of metallic scrap (especially aluminum) results in a 100% yield of the material whereas fusion-based processes lead to a material wastage of a minimum of 15-20% in fumes. The energy saving in solid-state recycling is enormous in comparison to virgin metal production or fusion-based recycling. Approximately 156 MJ/Kg of energy is required for primary aluminum production whereas, a mere 10 MJ/Kg of energy is consumed in solid-state recycling of the aluminum whereas, the energy consumption in fusion-based recycling stands at 17.7 MJ/Kg [2,3]. Therefore, solid-state recycling of aluminum is a highly economical and sustainable process due to its huge material and energy-saving nature.

In summary, prioritizing secondary metal production offers multifaceted environmental advantages, including energy conservation and waste reduction. This is particularly crucial in the case of aluminum, a widely used metal in industrial applications, where the recovery of manufacturing scrap contributes not only to cost efficiency but also to sustainable and responsible waste management practices.

Friction-based processes have emerged as promising technologies for the manufacturing sector [4,5]. However, there remains a significant knowledge gap regarding the potential of friction-based additive manufacturing techniques for recycling machining waste [6,7]. Friction-based recycling processes could provide a sustainable and cost-effective solution for repurposing waste materials within the industry without melting the material [8].

Multiple researchers tried friction stir consolidation of the manufacturing waste metal [9,10] where they followed the high-pressure torsion method of the metallic chips using a friction stir tool. The cladding of one metal using a clad material of another metal is quite a renowned process [11] with various applications in industries ranging from corrosion-resisting material to high-temperature-resistant material. Still, the use of shop floor waste metal as a feedstock for cladding the substrate material is not being used. This process may solve a double problem of recycling the metallic waste as well as dissimilar material cladding for the enhanced surface properties using a singular friction stir-assisted cladding (FS-AC) process.

This article utilizes a friction stir-based approach for recycling aluminum chips and cladding them on a steel substrate. A rotating tungsten carbide tool is employed to stir the scrap material, generating sufficient heat to soften the chips for easy processing. The significant heat generation and axial force produced during the process induce the deformation and consolidation of the chips onto a substrate material. This consolidation is achieved through the continuous traverse motion of the tool after an appropriate dwell time, resulting in a uniform and uninterrupted deposition layer over the substrate's surface. The deposited metal layer is subsequently characterized using various metallographic techniques while the morphological features and potential defects are assessed. This article successfully validates the proof of concept for the FS-AC process, which can be further employed or cladding different materials too.

## Material and method

The Al6061 metallic swarf was received from the lathe machine shop. The swarf was chopped in an electric mixer jug to make the size uniform in the 1-2 mm range. The uniformed chips were used as feed material for cladding. The substrate metal plate was a 5 mm thick EN31 steel metal plate to be used for the FS-AC process. A fixture was designed and fabricated to rigidly clamp the substrate sheet and hold the Al chips during processing using a cover plate. The cover plate was movable and slid along with the traverse movement of the tool. Figures 1(a) and 1(b) show the
recovery is not only to minimize disposal costs but also to evade regulatory liability by extracting metals from voluminous wastes. Aluminum scrap is typically processed in a reverberatory furnace as part of the recycling process. However, melting the scrap is an energy-intensive process. Also, the inherent properties of aluminum alloys generally get lost during melting due to the evaporation of various strength-imparting elements. The solid-state recycling of aluminum alloys will result in huge energy savings while keeping the properties intact. The solid-state recycling of metallic scrap (especially aluminum) results in a 100 % yield of the material whereas fusion-based processes lead to a material wastage of a minimum of 15-20 % in fumes. The energy saving in solid-state recycling is enormous in comparison to virgin metal production or fusion-based recycling. Approximately 156 MJ/Kg of energy is required for primary aluminum production whereas, a mere 10 MJ/Kg of energy is consumed in solid-state recycling of the aluminum whereas, the energy consumption in fusion-based recycling stands at 17.7 MJ/Kg [2,3]. Therefore, solid-state recycling of aluminum is a highly economical and sustainable process due to its huge material and energy-saving nature.

In summary, prioritizing secondary metal production offers multifaceted environmental advantages, including energy conservation and waste reduction. This is particularly crucial in the case of aluminum, a widely used metal in industrial applications, where the recovery of manufacturing scrap contributes not only to cost efficiency but also to sustainable and responsible waste management practices.

Friction-based processes have emerged as promising technologies for the manufacturing sector [4,5]. However, there remains a significant knowledge gap regarding the potential of friction-based additive manufacturing techniques for recycling machining waste [6,7]. Friction-based recycling processes could provide a sustainable and cost-effective solution for repurposing waste materials within the industry without melting the material [8].

Multiple researchers tried friction stir consolidation of the manufacturing waste metal [9,10] where they followed the high-pressure torsion method of the metallic chips using a friction stir tool. The cladding of one metal using a clad material of another metal is quite a renowned process [11] with various applications in industries ranging from corrosion-resisting material to high-temperature-resistant material. Still, the use of shop floor waste metal as a feedstock for cladding the substrate material is not being used. This process may solve a double problem of recycling the metallic waste as well as dissimilar material cladding for the enhanced surface properties using a singular friction stir-assisted cladding (FS-AC) process.

This article utilizes a friction stir-based approach for recycling aluminum chips and cladding them on a steel substrate. A rotating tungsten carbide tool is employed to stir the scrap material, generating sufficient heat to soften the chips for easy processing. The significant heat generation and axial force produced during the process induce the deformation and consolidation of the chips onto a substrate material. This consolidation is achieved through the continuous traverse motion of the tool after an appropriate dwell time, resulting in a uniform and uninterrupted deposition layer over the substrate's surface. The deposited metal layer is subsequently characterized using various metallographic techniques while the morphological features and potential defects are assessed. This article successfully validates the proof of concept for the FS-AC process, which can be further employed or cladding different materials too.

## Material and method

The Al6061 metallic swarf was received from the lathe machine shop. The swarf was chopped in an electric mixer jug to make the size uniform in the 1-2 mm range. The uniformed chips were used as feed material for cladding. The substrate metal plate was a 5 mm thick EN31 steel metal plate to be used for the FS-AC process. A fixture was designed and fabricated to rigidly clamp the substrate sheet and hold the Al chips during processing using a cover plate. The cover plate was movable and slid along with the traverse movement of the tool. Figures 1(a) and 1(b) show the
schematic of the fixture plate assembly and the initial tool placement. The metallic swarf/chips are placed on the friction stir welding machine table over the substrate, and the side view is shown in Figure 1(c).

<figure><img src="image_2629.png" /><figcaption>Fig. 1. Schematic representation of fixture assembly and placement of the tool and metallic swarf used for friction stir-assisted cladding process</figcaption></figure>

The CNC-operated friction stir welding machine having a 10 HP motor was used for the FS-AC process. The tool used for friction stir processing was made up of tungsten carbide material to avoid wear during rubbing action. The tool axis was tilted 5 degrees backward to accommodate the chips underneath for the friction stir-assisted cladding process. A tool rotational speed of 700 RPM and traverse speed of 100 mm/minute were used. Finally, on completion of the 120 mm deposition, the cover plate slid out of the fixture, exposing the deposited layer of Al6061 metal over a mild steel substrate, as shown in Figure 1(d).

The cross-section of the deposited layer over the substrate surface was cut using a wire electric discharge machine and prepared using standard metallographic polishing and etching techniques. The Nital reagent, having 5 % nitric acid in methanol, was used for etching purposes. The micro-hardness testing of the clad layer was performed on INNOVATEST hardness machine with a load of 100 grams and a dwell time of 10 seconds. The indentations were marked at the mid-thickness of the clad layer.

# Results and Discussion

## Friction stir-assisted cladding process mechanism

The process of cladding aluminum metallic swarf on a steel substrate using FS-AC was developed, resulting in a uniform-clad layer over the substrate. The metallic swarf was utilized as a feedstock material for the cladding process, which led to the development of a layer on the substrate without the cost inclusion of the clad material. The actual photographs of the aluminum swarf and the cladding process using a friction stir machine are shown in Figure 2.
schematic of the fixture plate assembly and the initial tool placement. The metallic swarf/chips are placed on the friction stir welding machine table over the substrate, and the side view is shown in Figure 1(c).

<figure><img src="image_2623.png" /><figcaption>Fig. 1. Schematic representation of fixture assembly and placement of the tool and metallic swarf used for friction stir-assisted cladding process</figcaption></figure>

The CNC-operated friction stir welding machine having a 10 HP motor was used for the FS-AC process. The tool used for friction stir processing was made up of tungsten carbide material to avoid wear during rubbing action. The tool axis was tilted 5 degrees backward to accommodate the chips underneath for the friction stir-assisted cladding process. A tool rotational speed of 700 RPM and traverse speed of 100 mm/minute were used. Finally, on completion of the 120 mm deposition, the cover plate slid out of the fixture, exposing the deposited layer of Al6061 metal over a mild steel substrate, as shown in Figure 1(d).

The cross-section of the deposited layer over the substrate surface was cut using a wire electric discharge machine and prepared using standard metallographic polishing and etching techniques. The Nital reagent, having 5 % nitric acid in methanol, was used for etching purposes. The micro-hardness testing of the clad layer was performed on INNOVATEST hardness machine with a load of 100 grams and a dwell time of 10 seconds. The indentations were marked at the mid-thickness of the clad layer.

## Results and Discussion

### Friction stir-assisted cladding process mechanism

The process of cladding aluminum metallic swarf on a steel substrate using FS-AC was developed, resulting in a uniform-clad layer over the substrate. The metallic swarf was utilized as a feedstock material for the cladding process, which led to the development of a layer on the substrate without the cost inclusion of the clad material. The actual photographs of the aluminum swarf and the cladding process using a friction stir machine are shown in Figure 2.
<figure><img src="image_2630.png" /><figcaption>Fig. 2. Photographs of aluminum chips and the FS-AC process, along with the cladding developed during the FS-AC process</figcaption></figure>

The aluminum metal chips of approximately uniform size and shape were placed to rub against the rotating tool of hard tungsten carbide material with the fixation of the cover plate over them. The rotating action of the tool generated the heat required for softening the metallic swarf, and the tilt angle of the tool enabled the compaction of metallic chips to form a cladding layer with increased density. The advancing side of the cladding plate represents the side where the tool rotational velocity vector and tool traverse velocity vector align in the same direction. On the other hand, the side in which the tool rotational velocity vector and tool traverse velocity vectors are opposite in direction is represented by the retreating side.

The thickness of the incoming swarf layer was equal to the gap between the substrate plate's upper surface and the cover plate's bottom surface. The thickness of the cladding layer was controlled by the gap between the backward surface of the tool face and the upper surface of the substrate, as shown in the side-view representation of the FS-AC process in Figure 3.
<figure><img src="image_2624.png" /><figcaption>Fig. 2. Photographs of aluminum chips and the FS-AC process, along with the cladding developed during the FS-AC process</figcaption></figure>

The aluminum metal chips of approximately uniform size and shape were placed to rub against the rotating tool of hard tungsten carbide material with the fixation of the cover plate over them. The rotating action of the tool generated the heat required for softening the metallic swarf, and the tilt angle of the tool enabled the compaction of metallic chips to form a cladding layer with increased density. The advancing side of the cladding plate represents the side where the tool rotational velocity vector and tool traverse velocity vector align in the same direction. On the other hand, the side in which the tool rotational velocity vector and tool traverse velocity vectors are opposite in direction is represented by the retreating side.

The thickness of the incoming swarf layer was equal to the gap between the substrate plate's upper surface and the cover plate's bottom surface. The thickness of the cladding layer was controlled by the gap between the backward surface of the tool face and the upper surface of the substrate, as shown in the side-view representation of the FS-AC process in Figure 3.
**Side view representation of the**
**FS-AC process**

<figure><img src="image_2631.png" /></figure>

**Forces acting on the chip elements at**
**different locations**

<figure><img src="image_2632.png" /><figcaption>Fig. 3. Side view representation of the FS-AC process and the various forces acting on the chips placed at different locations with respect to the tool surface</figcaption></figure>

There are different forces acting on the metallic swarf during the FS-AC process because of the location of the swarf. The axial downwards force and the horizontal force due to the traverse movement of the tool are of a similar nature all over the swarf. However, the friction force experienced by the swarf is different at different locations. On the metallic swarf placed at the front part of the tool, which comes in contact with the tool initially, the frictional force is applied along the surface of the tool face towards inside the plane of the paper. As the tool moves ahead, the swarf under the centre of the tool does not experience any friction force due to the rotation of the tool because the velocity at the centre of the tool's face is zero. On further forward movement of the tool, the metallic swarf comes in contact with the backward part of the tool. The backward face of the tool is lower than the frontal face due to the backward tilt angle provided on the tool. This tilting enabled a wedge-shaped area for the metallic swarf to exit. There is extreme pressure in this area due to decreased area, and the friction force is applied towards the direction coming outwards from the plane of the paper. The heating of the swarf occurs due to the friction heat generated during the rotation of the tool surface against the metallic swarf, making it soft and easily flowable. The axial downward pressure applied by the tool surface is responsible for the compaction/consolidation of the already softened swarf. The compaction increases the density of the cladding layer with respect to the incoming swarf. The continuous narrowing of the area available for the swarf to flow during the tool movement enables substantial compaction of the soft material. The pressure applied by the backward part of the tool face controlled the thickness of the exiting coating layer making it of uniform thickness.

**Metallographic and mechanical analysis of the cladding**

The cross-section of the aluminum cladding over the steel plate is presented in Figure 4. The image depicts a uniform-clad layer with an average thickness of 0.95 mm. The thickness varied from 0.94 mm to 0.97 mm, showing a very good uniformity across the cross-section. The aluminum metal deposited as cladding was observed to be visibly sound and free from any significant voids or defects.

The clad layer and substrate interface were observed at a higher magnification to ascertain a defect-free interface between the clad and substrate. The interface region was found to be free from any discontinuities or cracks. The upper part of the steel substrate, which was exposed to frictional heat and mechanical rubbing of the swarf, showed a thermo-mechanically affected zone (TMAZ). The TMAZ region showed a very fine-grained structure up to a thickness of 25 µm to 40 µm due to the strain hardening of the region.
Side view representation of the
FS-AC process

<figure><img src="image_2625.png" /></figure>

Forces acting on the chip elements at
different locations

1. Chips under the frontal part of the tool

<figure><img src="image_2626.png" /></figure>

(Heating)
Frictional force due to tool rotational vector

Pressure due to
tool axial force
(Consolidation)

2. Chips under the central part of the tool

<figure><img src="image_2627.png" /></figure>

(Flattening)
Force due to tool
traverse motion

Pressure due to
tool axial force
(Consolidation)

3. Chips under the backward part of the tool

(Flattening)
Force due to tool
traverse motion

<figure><img src="image_2628.png" /></figure>

(Heating)
Frictional force due to tool rotational vector

Pressure due to
tool axial force
(Consolidation)

<caption>Fig. 3. Side view representation of the FS-AC process and the various forces acting on the chips placed at different locations with respect to the tool surface</caption>

There are different forces acting on the metallic swarf during the FS-AC process because of the location of the swarf. The axial downwards force and the horizontal force due to the traverse movement of the tool are of a similar nature all over the swarf. However, the friction force experienced by the swarf is different at different locations. On the metallic swarf placed at the front part of the tool, which comes in contact with the tool initially, the frictional force is applied along the surface of the tool face towards inside the plane of the paper. As the tool moves ahead, the swarf under the centre of the tool does not experience any friction force due to the rotation of the tool because the velocity at the centre of the tool's face is zero. On further forward movement of the tool, the metallic swarf comes in contact with the backward part of the tool. The backward face of the tool is lower than the frontal face due to the backward tilt angle provided on the tool. This tilting enabled a wedge-shaped area for the metallic swarf to exit. There is extreme pressure in this area due to decreased area, and the friction force is applied towards the direction coming outwards from the plane of the paper. The heating of the swarf occurs due to the friction heat generated during the rotation of the tool surface against the metallic swarf, making it soft and easily flowable. The axial downward pressure applied by the tool surface is responsible for the compaction/consolidation of the already softened swarf. The compaction increases the density of the cladding layer with respect to the incoming swarf. The continuous narrowing of the area available for the swarf to flow during the tool movement enables substantial compaction of the soft material. The pressure applied by the backward part of the tool face controlled the thickness of the exiting coating layer making it of uniform thickness.

Metallographic and mechanical analysis of the cladding

The cross-section of the aluminum cladding over the steel plate is presented in Figure 4. The image depicts a uniform-clad layer with an average thickness of 0.95 mm. The thickness varied from 0.94 mm to 0.97 mm, showing a very good uniformity across the cross-section. The aluminum metal deposited as cladding was observed to be visibly sound and free from any significant voids or defects.

The clad layer and substrate interface were observed at a higher magnification to ascertain a defect-free interface between the clad and substrate. The interface region was found to be free from any discontinuities or cracks. The upper part of the steel substrate, which was exposed to frictional heat and mechanical rubbing of the swarf, showed a thermo-mechanically affected zone (TMAZ). The TMAZ region showed a very fine-grained structure up to a thickness of 25 µm to 40 µm due to the strain hardening of the region.
<figure><img src="image_2633.png" /><figcaption>Fig. 4. Optical microscope images of the cross-sectional view of the cladding depicting a uniform thickness of clad material and distinct thermo-mechanically affected region towards the substrate side</figcaption></figure>

The Vickers hardness of the clad layer was examined at the mid-thickness of the layer and the maximum hardness of 102 HV and minimum hardness of 60 HV was observed across the whole layer. The average hardness value was 75 HV. A particular trend in the hardness data was not observed.

**The industrial implication, limitations, and further work**

Developing a uniformly thick clad layer over the substrate in a single-step process reduces the post-processing time for the finished product and, in turn, enhances the economic viability of this process. In this work, a novel process was developed to recycle the metallic waste scrap and clad it on a substrate with a controlled uniform thickness in a single step. Further advanced characterization studies are needed to confirm the diffusion of the dissimilar materials into each other and the chances of intermetallic formation. The mechanical and chemical testing of the specimen would also give clear insights regarding the coating-substrate bond strength and corrosion resistance improvement base material.

The FS-AC process could prove to be a go-to process for recycling the metallic scrap without melting. This process could easily be repeated for multiple layers and presented as a solid-state alternative to additive manufacturing processes. Furthermore, using waste material as feedstock for additive manufacturing would aid the economic aspect of this process.
<figure><img src="image_2629.png" /><figcaption>Fig. 4. Optical microscope images of the cross-sectional view of the cladding depicting a uniform thickness of clad material and distinct thermo-mechanically affected region towards the substrate side</figcaption></figure>

The Vickers hardness of the clad layer was examined at the mid-thickness of the layer and the maximum hardness of 102 HV and minimum hardness of 60 HV was observed across the whole layer. The average hardness value was 75 HV. A particular trend in the hardness data was not observed.

**The industrial implication, limitations, and further work**

Developing a uniformly thick clad layer over the substrate in a single-step process reduces the post-processing time for the finished product and, in turn, enhances the economic viability of this process. In this work, a novel process was developed to recycle the metallic waste scrap and clad it on a substrate with a controlled uniform thickness in a single step. Further advanced characterization studies are needed to confirm the diffusion of the dissimilar materials into each other and the chances of intermetallic formation. The mechanical and chemical testing of the specimen would also give clear insights regarding the coating-substrate bond strength and corrosion resistance improvement base material.

The FS-AC process could prove to be a go-to process for recycling the metallic scrap without melting. This process could easily be repeated for multiple layers and presented as a solid-state alternative to additive manufacturing processes. Furthermore, using waste material as feedstock for additive manufacturing would aid the economic aspect of this process.
## Conclusions

The current study presented a novel method to recycle the machine shop floor scrap and turn it into useful feedstock material for cladding over a substrate using friction-based processing. The friction stir-assisted cladding (FS-AC) process successfully developed a method to deposit the metallic swarf over a flat substrate, forming a uniform thickness layer. The superimposition of several layers of the metallic swarf and repetitive friction stir processing would lead to the development of additively manufactured components. The process not only solves the problem of shop floor waste but also leads to the fabrication of high-value components without being an energy-excessive method.

Many different material combinations can be processed using the FS-AC process, exploring the possibilities of functionally graded material development using metallic swarf.

## Acknowledgment

The authors acknowledge *Interface Design Pvt. Ltd. (Nashik)* for providing the experimental facility of friction stir machine for the present work.

## References

[1] I.K. Wernick, N.J. Themelis, Recycling metals for the environment, 1998. www.annualreviews.org.

[2] E. Balomenos, D. Panias, I. Paspaliaris, Energy and exergy analysis of the primary aluminum production processes: A review on current and future sustainability, Mineral Processing and Extractive Metallurgy Review 32 (2011) 69-89. https://doi.org/10.1080/08827508.2010.530721

[3] S. Shamsudin, M.A. Lajis, Z.W. Zhong, Solid-state recycling of light metals: A review, Advances in Mechanical Engineering 8 (2016) 1-23. https://doi.org/10.1177/1687814016661921.

[4] P. Kaushik, D. Kumar Dwivedi, Heat generation and steel fragment effects on friction stir welding of aluminum alloy with steel, Materials and Manufacturing Processes 39:2 (2023) 161-173. https://doi.org/10.1080/10426914.2023.2187827

[5] P. Kaushik, D. Kumar Dwivedi, Enhancing strength and metallurgical soundness of dissimilar Al-steel FSSW joints through novel exit hole refilling technique, Manuf Lett 38 (2023) 83-87. https://doi.org/10.1016/j.mfglet.2023.10.003

[6] T. Borgert, W. Homberg, Friction-induced recycling process for user-specific semi-finished product production, Metals (Basel) 11 (2021). https://doi.org/10.3390/met11040663

[7] P. Asadi, M. Akbari, O. Kohantorabi, M. Peyghami, M.R.M. Aliha, S.M. Salehi, H.R. Asiabaraki, F. Berto, Characterization of the Influence of Rotational and Traverse Speeds on the Mechanical and Microstructural Properties of Wires Produced By the FSBE Method, Strength of Materials 54 (2022) 318-330. https://doi.org/10.1007/s11223-022-00403-5

[8] M. Wędrychowicz, Ł. Wzorek, T. Tokarski, P. Noga, J. Wiewióra, Recycling without melting: An alternative approach to aluminum scrap recovery, in: Key Eng Mater, Trans Tech Publications Ltd, 2016: pp. 284-289. https://doi.org/10.4028/www.scientific.net/KEM.682.284

[9] D. Baffari, G. Buffa, G. Ingarao, A. Masnata, L. Fratini, Aluminium sheet metal scrap recycling through friction consolidation, in: Procedia Manuf, Elsevier B.V., 2019: pp. 560-566. https://doi.org/10.1016/j.promfg.2019.02.134

[10] A. Latif, G. Ingarao, G. Buffa, L. Fratini, Forgeability characterization of multi-material based functionally graded materials manufactured through friction stir consolidation, IOP Conf Ser Mater Sci Eng 1270 (2022) 012096. https://doi.org/10.1088/1757-899x/1270/1/012096

[11] M.K. Mahto, A. Kumar, A.R. Raja, M. Vashista, M.Z.K. Yusufzai, Friction stir cladding of copper on aluminium substrate, CIRP J Manuf Sci Technol 36 (2022) 23-34. https://doi.org/10.1016/j.cirpj.2021.10.004
## Conclusions

The current study presented a novel method to recycle the machine shop floor scrap and turn it into useful feedstock material for cladding over a substrate using friction-based processing. The friction stir-assisted cladding (FS-AC) process successfully developed a method to deposit the metallic swarf over a flat substrate, forming a uniform thickness layer. The superimposition of several layers of the metallic swarf and repetitive friction stir processing would lead to the development of additively manufactured components. The process not only solves the problem of shop floor waste but also leads to the fabrication of high-value components without being an energy-excessive method.

Many different material combinations can be processed using the FS-AC process, exploring the possibilities of functionally graded material development using metallic swarf.

## Acknowledgment

The authors acknowledge *Interface Design Pvt. Ltd. (Nashik)* for providing the experimental facility of friction stir machine for the present work.

## References

[1] I.K. Wernick, N.J. Themelis, Recycling metals for the environment, 1998. www.annualreviews.org.

[2] E. Balomenos, D. Panias, I. Paspaliaris, Energy and exergy analysis of the primary aluminum production processes: A review on current and future sustainability, Mineral Processing and Extractive Metallurgy Review 32 (2011) 69-89. https://doi.org/10.1080/08827508.2010.530721

[3] S. Shamsudin, M.A. Lajis, Z.W. Zhong, Solid-state recycling of light metals: A review, Advances in Mechanical Engineering 8 (2016) 1-23. https://doi.org/10.1177/1687814016661921.

[4] P. Kaushik, D. Kumar Dwivedi, Heat generation and steel fragment effects on friction stir welding of aluminum alloy with steel, Materials and Manufacturing Processes 39:2 (2023) 161-173. https://doi.org/10.1080/10426914.2023.2187827

[5] P. Kaushik, D. Kumar Dwivedi, Enhancing strength and metallurgical soundness of dissimilar Al-steel FSSW joints through novel exit hole refilling technique, Manuf Lett 38 (2023) 83-87. https://doi.org/10.1016/j.mfglet.2023.10.003

[6] T. Borgert, W. Homberg, Friction-induced recycling process for user-specific semi-finished product production, Metals (Basel) 11 (2021). https://doi.org/10.3390/met11040663

[7] P. Asadi, M. Akbari, O. Kohantorabi, M. Peyghami, M.R.M. Aliha, S.M. Salehi, H.R. Asiabaraki, F. Berto, Characterization of the Influence of Rotational and Traverse Speeds on the Mechanical and Microstructural Properties of Wires Produced By the FSBE Method, Strength of Materials 54 (2022) 318-330. https://doi.org/10.1007/s11223-022-00403-5

[8] M. Wędrychowicz, Ł. Wzorek, T. Tokarski, P. Noga, J. Wiewióra, Recycling without melting: An alternative approach to aluminum scrap recovery, in: Key Eng Mater, Trans Tech Publications Ltd, 2016: pp. 284-289. https://doi.org/10.4028/www.scientific.net/KEM.682.284

[9] D. Baffari, G. Buffa, G. Ingarao, A. Masnata, L. Fratini, Aluminium sheet metal scrap recycling through friction consolidation, in: Procedia Manuf, Elsevier B.V., 2019: pp. 560-566. https://doi.org/10.1016/j.promfg.2019.02.134

[10] A. Latif, G. Ingarao, G. Buffa, L. Fratini, Forgeability characterization of multi-material based functionally graded materials manufactured through friction stir consolidation, IOP Conf Ser Mater Sci Eng 1270 (2022) 012096. https://doi.org/10.1088/1757-899x/1270/1/012096

[11] M.K. Mahto, A. Kumar, A.R. Raja, M. Vashista, M.Z.K. Yusufzai, Friction stir cladding of copper on aluminium substrate, CIRP J Manuf Sci Technol 36 (2022) 23-34. https://doi.org/10.1016/j.cirpj.2021.10.004
# Assessing single and multi-step friction stir consolidated recycled billets through uniaxial upsetting test

LATIF Abdul<fnref n="1" />,ᵃ*, INGARAO Giuseppe<fnref n="1" />,ᵇ, FRATINI Livan<fnref n="1" />,ᶜ, HETZ Peter<fnref n="2" />,ᵈ and MERKLEIN Marion<fnref n="2" />,ᵉ

**Keywords:** Aluminum Alloys, Recycling, Friction Stir Consolidation, Upsetting Test

**Abstract.** Friction stir consolidation (FSC) represents a solid-state recycling technique that directly transforms machining waste into billets by the stirring action and friction heat of the rotating tool. This technology has demonstrated superior energy efficiency compared to remelting-based traditional recycling methods. Concerning the FSC process mechanics, they impart temperature, strain and strain rate gradients across the billet sections leading to variations in mechanical properties such as hardness values and grain size distribution. Therefore, multi-step FSC variants were introduced to eliminate this inherent inhomogeneity inside the billets of the single-step FSC approach. However, the billets manufactured by multi-step methods showed better mechanical properties, and gradation across the billet section still persisted. This property led to the evolution of friction stir consolidation as a new manufacturing method for developing functionally graded materials. Previously, the gradation was highlighted by the Vickers hardness test and grain size distribution. Therefore, in the ongoing study, an advanced characterization technique was adopted by extracting miniaturized samples from the section of FSC billets and subjected them to the upsetting test. Employing advanced characterization techniques to get insights into the gradation and quality of billets manufactured through various process variants and will allow to draw a better conclusion regarding adoption of a more suitable FSC process variant.

**Introduction**

The current global demand for aluminum is around 100 million tons per year [1] that is expected to rise further due to the growing application of aluminum in lightweight transport, packaging, construction, and electronic industries. This trend is putting immense pressure on industries to increase the production rate. Concerning the environmental impact, per ton of aluminum production generates 12-16 tons of greenhouse gas (GHG) during aluminum production from primary sources, which has severe adverse environmental impacts [2] and is responsible for almost 1% of total global energy consumption [3].

To address the rising demand for aluminum and achieve sustainability objectives, researchers are exploring the implementation of recycling practices. Interestingly, aluminum is an infinitely recyclable material [3]. In Europe, almost 35 % of the aluminum demand is met by recycling aluminum scraps [4]. The conventional recycling route involves re-melting the metal, and thus it skips many of the complex steps of aluminum extraction from the primary source. Furthermore, recycling is highly energy efficient, requiring 5% of the energy compared to obtain aluminum from the primary bauxite source. Thus, recycling can shift aluminum's energy and carbon balance sustainability towards much greater sustainability.

<figure><img src="image_2634.png" /></figure>

<footnote n="1">¹University of Palermo, Viale delle Scienze, 90128 Palermo, Italy</footnote>

<footnote n="2">²Institute of Manufacturing Technology, Friedrich-Alexander-Universität Erlangen-Nürnberg, Egerlandstraße 13, 91058 Erlangen, Germany</footnote>

<footnote>ᵃabdul.latif@unipa.it, ᵇgiuseppe.ingarao@unia.it, ᶜlivan.fratini@unipa.it, ᵈpeter.hetz@fau.de,</footnote>

<footnote>ᵉmarion.merklein@fau.de</footnote>
# Assessing single and multi-step friction stir consolidated recycled billets through uniaxial upsetting test

LATIF Abdul<fnref n="1" />,ᵃ*, INGARAO Giuseppe<fnref n="1" />,ᵇ, FRATINI Livan<fnref n="1" />,ᶜ, HETZ Peter<fnref n="2" />,ᵈ and MERKLEIN Marion<fnref n="2" />,ᵉ

**Keywords:** Aluminum Alloys, Recycling, Friction Stir Consolidation, Upsetting Test

**Abstract.** Friction stir consolidation (FSC) represents a solid-state recycling technique that directly transforms machining waste into billets by the stirring action and friction heat of the rotating tool. This technology has demonstrated superior energy efficiency compared to remelting-based traditional recycling methods. Concerning the FSC process mechanics, they impart temperature, strain and strain rate gradients across the billet sections leading to variations in mechanical properties such as hardness values and grain size distribution. Therefore, multi-step FSC variants were introduced to eliminate this inherent inhomogeneity inside the billets of the single-step FSC approach. However, the billets manufactured by multi-step methods showed better mechanical properties, and gradation across the billet section still persisted. This property led to the evolution of friction stir consolidation as a new manufacturing method for developing functionally graded materials. Previously, the gradation was highlighted by the Vickers hardness test and grain size distribution. Therefore, in the ongoing study, an advanced characterization technique was adopted by extracting miniaturized samples from the section of FSC billets and subjected them to the upsetting test. Employing advanced characterization techniques to get insights into the gradation and quality of billets manufactured through various process variants and will allow to draw a better conclusion regarding adoption of a more suitable FSC process variant.

**Introduction**

The current global demand for aluminum is around 100 million tons per year [1] that is expected to rise further due to the growing application of aluminum in lightweight transport, packaging, construction, and electronic industries. This trend is putting immense pressure on industries to increase the production rate. Concerning the environmental impact, per ton of aluminum production generates 12-16 tons of greenhouse gas (GHG) during aluminum production from primary sources, which has severe adverse environmental impacts [2] and is responsible for almost 1% of total global energy consumption [3].

To address the rising demand for aluminum and achieve sustainability objectives, researchers are exploring the implementation of recycling practices. Interestingly, aluminum is an infinitely recyclable material [3]. In Europe, almost 35 % of the aluminum demand is met by recycling aluminum scraps [4]. The conventional recycling route involves re-melting the metal, and thus it skips many of the complex steps of aluminum extraction from the primary source. Furthermore, recycling is highly energy efficient, requiring 5% of the energy compared to obtain aluminum from the primary bauxite source. Thus, recycling can shift aluminum's energy and carbon balance sustainability towards much greater sustainability.

<figure><img src="image_2630.png" /></figure>

<footnote n="1">¹University of Palermo, Viale delle Scienze, 90128 Palermo, Italy</footnote>

<footnote n="2">²Institute of Manufacturing Technology, Friedrich-Alexander-Universität Erlangen-Nürnberg, Egerlandstraße 13, 91058 Erlangen, Germany</footnote>

<footnote>ᵃabdul.latif@unipa.it, ᵇgiuseppe.ingarao@unia.it, ᶜlivan.fratini@unipa.it, ᵈpeter.hetz@fau.de, ᵉmarion.merklein@fau.de</footnote>
The conventional recycling method has further limitations when dealing with the recycling of aluminum machining chips. These chips, characterized by a high surface-to-volume ratio and susceptibility to oxidation, incur permanent material losses during the melting process. This results in adverse environmental impacts, increased costs, and substantial permanent material loss [5].

Looking at the proportion, the machining scraps constitute almost 17% of the total scraps [6]. Therefore, the researchers turned to solid-state recycling (SSR) techniques. SSR are those methods that transform metal scraps directly into finished or semi-finished billet through mechanical means such as high pressure, friction, rotational speed, and force [7]. Recently, SSR processes have been analyzed as a potential alternative for recycling machining scraps. Various SRR methods have been introduced for recycling metal chips based on the process mechanics and properties of the desired product [8-10]. In fact, by omitting the melting phase, permanent losses mainly due to oxidation are avoided.

Recently FSC has proved to be one of the environment recycling techniques compared to conventional remelting [11]. This technology has proved its potential beyond the concept of recycling toward upcycling method by manufacturing functionally graded material from metal scraps [12]. FSC has two main steps: compaction and consolidation. During compaction, chips or powder are pressed in a hollow die chamber by applying a specific load through a cylindrical tool. Then compacted materials are further pressed and stirred through the tool's downward force and rotational speed during the consolidation phase [13].

A typical FSC billet, due to process mechanics such as different temperatures, strain and strain rate distribution, is characterized by variation in mechanical properties along the radial and longitudinal direction of the billet [12, 13]. Previously, this variation was validated by Vickers's hardness test and grain size distribution analysis [14]. Exploring the potential of FSC in manufacturing advanced Functionally Graded Materials (FGMs) requires highlighting gradation through advanced characterization techniques. The concept of miniaturized upsetting specimens for local sheet metal characterization was introduced by Hetz et al. [15]. The test method can also be applied to analyze the local flow properties of bulk metal parts like the billets in this contribution. Previously, the authors of the current research, Latif et al. [16], extracted miniaturized samples for single-step FSC AA2011 to emphasize gradation in FSCed samples.

In specific, this research is the extended investigation of the previous study [16] aiming to highlight gradation for billets of different aluminum alloys: AA2011 and AA7075 manufactured through various process variants, namely, single-step and multi-step. In addition, this characterization method provides valuable insights into the local mechanical properties of the billets, enabling meaningful comparisons for the selection of the most suitable FSC process variants that are otherwise not possible using conventional characterization methods. The broader objective is to find a more accurate solution for the numerical modelling of the FSC process and find the potential industrial application.

## Material and methods

### Material and process set-up

The employed experimental setup consisted of ESAB LEGIO, a dedicated friction welding machine with 35 kN maximum capacity. A tool of H13 steel with 25 mm nominal diameter was integrated to the head of ESAB machine and a H13 steel die cylinder die with 25mm was fixed onto the working table as shown in Fig. 1.

As-received materials were a 20x20 mm squared bar of AA 2011-T3 and a 3 mm thick rolled sheet of AA7075. Both materials were reduced into chips by a milling operation with a rotational speed of 1250 rpm, a feed rate of 28 mm/min and a cutting depth of 1 mm. The chips were kept submerged in acetone for effective cleaning for 30 minutes which has proved as a potential washing solvent for cleaning metal scraps both at laboratory and industrial scales.
The conventional recycling method has further limitations when dealing with the recycling of aluminum machining chips. These chips, characterized by a high surface-to-volume ratio and susceptibility to oxidation, incur permanent material losses during the melting process. This results in adverse environmental impacts, increased costs, and substantial permanent material loss [5].

Looking at the proportion, the machining scraps constitute almost 17% of the total scraps [6]. Therefore, the researchers turned to solid-state recycling (SSR) techniques. SSR are those methods that transform metal scraps directly into finished or semi-finished billet through mechanical means such as high pressure, friction, rotational speed, and force [7]. Recently, SSR processes have been analyzed as a potential alternative for recycling machining scraps. Various SRR methods have been introduced for recycling metal chips based on the process mechanics and properties of the desired product [8-10]. In fact, by omitting the melting phase, permanent losses mainly due to oxidation are avoided.

Recently FSC has proved to be one of the environment recycling techniques compared to conventional remelting [11]. This technology has proved its potential beyond the concept of recycling toward upcycling method by manufacturing functionally graded material from metal scraps [12]. FSC has two main steps: compaction and consolidation. During compaction, chips or powder are pressed in a hollow die chamber by applying a specific load through a cylindrical tool. Then compacted materials are further pressed and stirred through the tool's downward force and rotational speed during the consolidation phase [13].

A typical FSC billet, due to process mechanics such as different temperatures, strain and strain rate distribution, is characterized by variation in mechanical properties along the radial and longitudinal direction of the billet [12, 13]. Previously, this variation was validated by Vickers's hardness test and grain size distribution analysis [14]. Exploring the potential of FSC in manufacturing advanced Functionally Graded Materials (FGMs) requires highlighting gradation through advanced characterization techniques. The concept of miniaturized upsetting specimens for local sheet metal characterization was introduced by Hetz et al. [15]. The test method can also be applied to analyze the local flow properties of bulk metal parts like the billets in this contribution. Previously, the authors of the current research, Latif et al. [16], extracted miniaturized samples for single-step FSC AA2011 to emphasize gradation in FSCed samples.

In specific, this research is the extended investigation of the previous study [16] aiming to highlight gradation for billets of different aluminum alloys: AA2011 and AA7075 manufactured through various process variants, namely, single-step and multi-step. In addition, this characterization method provides valuable insights into the local mechanical properties of the billets, enabling meaningful comparisons for the selection of the most suitable FSC process variants that are otherwise not possible using conventional characterization methods. The broader objective is to find a more accurate solution for the numerical modelling of the FSC process and find the potential industrial application.

## Material and methods

### Material and process set-up

The employed experimental setup consisted of ESAB LEGIO, a dedicated friction welding machine with 35 kN maximum capacity. A tool of H13 steel with 25 mm nominal diameter was integrated to the head of ESAB machine and a H13 steel die cylinder die with 25mm was fixed onto the working table as shown in Fig. 1.

As-received materials were a 20x20 mm squared bar of AA 2011-T3 and a 3 mm thick rolled sheet of AA7075. Both materials were reduced into chips by a milling operation with a rotational speed of 1250 rpm, a feed rate of 28 mm/min and a cutting depth of 1 mm. The chips were kept submerged in acetone for effective cleaning for 30 minutes which has proved as a potential washing solvent for cleaning metal scraps both at laboratory and industrial scales.
After drying, a constant mass of 15g cleaned chips was loaded inside the die and compacted with 5 kN load. The step of chips compaction was necessary in order to eliminate the possibility of sprinkling out chips during the following consolidation step. Finally, chips were consolidated with load of 20 kN pressing, 30s of processing and 1500 rpm through different FSC process variants according to the experimental plan as shown in Table 1. The choice of the process condition was driven by previous studies [12, 14].

<figure><img src="image_2635.png" /><figcaption>Fig. 1 Process chain of FSC process from scraps to consolidated billet</figcaption></figure>

In particular three process variants: single step, double sided and mass addition methods were used for the billet manufacturing. During single step method, all input mass (15g) were compacted and consolidated in a single run for the total processing time of 30s. In double step method, likewise single step all input mass was consolidated for half of the total process time (15s), then billet was removed from the die and reinserted up-side down by applying consolidation step for the half of the processing time. During mass addition approach, half of the total input mass (7.5 g) was consolidated for half of the total processing time while remaining half input mass was added and then total mass (15g) was consolidated for the remaining half of the processing time. The purpose of flipping side and step wise consolidation during double sided and mass addition methods respectively was to achieve billets with improved mechanical properties compared to single step FSC billet.

<table><caption>Table.1 Experimental plan for developing billets using different variants of FSC process</caption><thead><tr><td>Exp ID</td><td>Material</td><td>FSC process variants</td><td>Process conditions<br/>RPM-mass [g]-time [s]</td></tr></thead><tbody><tr><td>Exp 1</td><td>AA2011</td><td>Single step</td><td>1500-15-30</td></tr><tr><td>Exp 2</td><td>AA7075</td><td>Single step</td><td>1500-15-30</td></tr><tr><td>Exp 3</td><td>AA2011</td><td>Double sided</td><td>1500-15-30</td></tr><tr><td>Exp 4</td><td>AA7075</td><td>Double sided</td><td>1500-15-30</td></tr><tr><td>Exp 5</td><td>AA2011</td><td>Mass addition</td><td>1500-15-30</td></tr><tr><td>Exp 6</td><td>AA7075</td><td>Mass addition</td><td>1500-15-30</td></tr></tbody></table>

## Measured outputs

After FSC consolidated step, the obtained billets had an average height and a diameter of 10.0 mm and 25.0 mm, respectively. The top and bottom surfaces of the FSC cylindrical billet were well polished to get a very flat surface for exact alignment of the billet in the following specimen extraction process as shown in Fig. 2. Then slices were cut (perpendicular to the cylindrical axis of the billet) from the top, middle, and bottom sections with a thickness of 1.5 mm and a diameter equal to the original billet diameter (25 mm) using die-sinking EDM. From each disc slice, five miniaturized cylindrical upsetting samples were extracted (Fig. 2) that had a height (h) of 1.5 mm
After drying, a constant mass of 15g cleaned chips was loaded inside the die and compacted with 5 kN load. The step of chips compaction was necessary in order to eliminate the possibility of sprinkling out chips during the following consolidation step. Finally, chips were consolidated with load of 20 kN pressing, 30s of processing and 1500 rpm through different FSC process variants according to the experimental plan as shown in Table 1. The choice of the process condition was driven by previous studies [12, 14].

<figure><img src="image_2631.png" /><figcaption>Fig. 1 Process chain of FSC process from scraps to consolidated billet</figcaption></figure>

In particular three process variants: single step, double sided and mass addition methods were used for the billet manufacturing. During single step method, all input mass (15g) were compacted and consolidated in a single run for the total processing time of 30s. In double step method, likewise single step all input mass was consolidated for half of the total process time (15s), then billet was removed from the die and reinserted up-side down by applying consolidation step for the half of the processing time. During mass addition approach, half of the total input mass (7.5 g) was consolidated for half of the total processing time while remaining half input mass was added and then total mass (15g) was consolidated for the remaining half of the processing time. The purpose of flipping side and step wise consolidation during double sided and mass addition methods respectively was to achieve billets with improved mechanical properties compared to single step FSC billet.

<table><caption>Table.1 Experimental plan for developing billets using different variants of FSC process</caption><thead><tr><td>Exp ID</td><td>Material</td><td>FSC process variants</td><td>Process conditions<br/>RPM-mass [g]-time [s]</td></tr></thead><tbody><tr><td>Exp 1</td><td>AA2011</td><td>Single step</td><td>1500-15-30</td></tr><tr><td>Exp 2</td><td>AA7075</td><td>Single step</td><td>1500-15-30</td></tr><tr><td>Exp 3</td><td>AA2011</td><td>Double sided</td><td>1500-15-30</td></tr><tr><td>Exp 4</td><td>AA7075</td><td>Double sided</td><td>1500-15-30</td></tr><tr><td>Exp 5</td><td>AA2011</td><td>Mass addition</td><td>1500-15-30</td></tr><tr><td>Exp 6</td><td>AA7075</td><td>Mass addition</td><td>1500-15-30</td></tr></tbody></table>

## Measured outputs

After FSC consolidated step, the obtained billets had an average height and a diameter of 10.0 mm and 25.0 mm, respectively. The top and bottom surfaces of the FSC cylindrical billet were well polished to get a very flat surface for exact alignment of the billet in the following specimen extraction process as shown in Fig. 2. Then slices were cut (perpendicular to the cylindrical axis of the billet) from the top, middle, and bottom sections with a thickness of 1.5 mm and a diameter equal to the original billet diameter (25 mm) using die-sinking EDM. From each disc slice, five miniaturized cylindrical upsetting samples were extracted (Fig. 2) that had a height (h) of 1.5 mm
(slice thickness) and a diameter (d) of 1 mm through micro-EDM on an SX-200-HPM (Sarix SA) as shown in Fig. 3. One sample (x1) was extracted from the center (r=0) and the remaining four samples (x2, x3, x4, and x5) were extracted at r=10 mm from the center of the billet. Upsetting tests were performed at a strain rate of 0.5%/s with upsetting sample dimensions given in Fig. 3. The height-to-diameter ratio (1.5) of the upsetting was selected based on the following equation [15].

$$1 \le \frac{h}{d} \le 2. \qquad (1)$$

The selection of the height-to-diameter ratio was based on findings from a previous study [16]. Further, according to this equation (1), a higher ratio exceeding the maximum limit leads to the buckling of the upsetting sample, which is considered a premature failure. On the other hand, a lower ratio increases the effect of the frictional influence. Therefore, Dionol ST V 1725-2 (MKU-Chemie GmbH) lubricant was applied to minimize friction.

In addition, hardness was measured across the section of the FSC billet at constant pitch of 1.5 mm in radial and cylindrical directions. In particular, Vickers hardness test was performed by applying 49 N load for dwell time of 15s. The results of both upsetting test and hardness measurements were compared to get better insights into quality and gradation.

<figure><img src="image_2636.png" /><figcaption>Fig. 2 Process sketch miniaturized upsetting sample extraction</figcaption></figure>

<figure><img src="image_2637.png" /><figcaption>Fig. 3 Upsetting test setup and sample dimensions</figcaption></figure>
(slice thickness) and a diameter (d) of 1 mm through micro-EDM on an SX-200-HPM (Sarix SA) as shown in Fig. 3. One sample (x1) was extracted from the center (r=0) and the remaining four samples (x2, x3, x4, and x5) were extracted at r=10 mm from the center of the billet. Upsetting tests were performed at a strain rate of 0.5%/s with upsetting sample dimensions given in Fig. 3. The height-to-diameter ratio (1.5) of the upsetting was selected based on the following equation [15].

$$1 \le \frac{h}{d} \le 2. \qquad (1)$$

The selection of the height-to-diameter ratio was based on findings from a previous study [16]. Further, according to this equation (1), a higher ratio exceeding the maximum limit leads to the buckling of the upsetting sample, which is considered a premature failure. On the other hand, a lower ratio increases the effect of the frictional influence. Therefore, Dionol ST V 1725-2 (MKU-Chemie GmbH) lubricant was applied to minimize friction.

In addition, hardness was measured across the section of the FSC billet at constant pitch of 1.5 mm in radial and cylindrical directions. In particular, Vickers hardness test was performed by applying 49 N load for dwell time of 15s. The results of both upsetting test and hardness measurements were compared to get better insights into quality and gradation.

<figure><img src="image_2632.png" /><figcaption>Fig. 2 Process sketch miniaturized upsetting sample extraction</figcaption></figure>

<figure><img src="image_2633.png" /><figcaption>Fig. 3 Upsetting test setup and sample dimensions</figcaption></figure>
## Results and Discussion

This section provides details information about the true strain-strain curves of the FSC billet of AA2011 and AA 7075 manufactured through single step and multi-step methods. AA2011 single billet that has been already analyzed by authors of the current study [16], is taken as reference for detailed discussion and then influence of change in material and process variants are analyzed. Further, FSC billets plastic material behavior is then compared to as-received material ones to understand the changes made by FSC process in the mechanical properties of the materials. In addition, a correlation is also established between strength and hardness of FSC billet by analyzing the ultimate strength and hardness values for the zones of the interest particularly in top, middle and bottom layers of the FSC billet.

### True-stress strain analysis of AA2011 single step FSC billet

This section gives details information about the true strain-strain curves of the FSC 2011 billet with evidence of how far true stress-strain graphs agree with the hardness distribution profile across the FSC billet section. Further, a comparison was also made between the true stress-strain curve FSC recycled billet and as-received material AA2011 bar.

On the top slice (Fig. 4 a and b), the central sample (x1) showed higher strength compared to the remaining upsetting samples (x2 to x5). The flow curves of the remaining four were found almost aligned with each other. The reason was samples x2 to x5 were extracted from the same height (top slice) and same radial distance (r=10 mm). Interestingly, the variation in properties is enabled by FSC process mechanics, i.e., different strain rate and temperature distributions across the billet section, as reported by Tang et al. [14]. In addition, a preliminary analysis has been provided by Puleo et al. [17] but it is quite evident that the die wall and back plate serve as heat sinks, reducing the temperature. Due to this, chips close to the die wall and back plate don't receive sufficient heat input and deformation to be fully recrystallized, but some bonding is, however, achieved.

Likewise, the top slice flow curves, almost the same trend was found for the middle slice. The central specimen (y1) showed higher strength than the other samples y2 and y3 (Fig. 4c).

For all samples of the bottom slice, almost all flow curves were found aligned with each other indicating less variation in mechanical properties. It is evident that the strength of the material varies both in the radial and longitudinal direction of the billet. Further, the top part of the billet shows higher yield strength than the bottom part, and along the radial direction the central part is stronger than the region near the external surface. Also, along the radial direction, the difference in true stress-strain curves between the samples at the center (x1) and the other samples (x2 and x3) is higher for the top slice, then difference diminishes at the middle slices, and finally, all curves are aligned with each other for the bottom slice (Fig. 4c).

Interestingly, the trend of true stress-strain profile was found consistent with the hardness distribution trend (Fig. 4c and d). For example, the zone of the billet section that exhibited higher hardness value also showed higher strength. The hardness measurements were performed on the same material, AA2011, manufactured under the same process conditions.

Again in Fig. 4d the middle of the billet section shows a minor variation in hardness value along the radial direction. A similar relationship can be found for the flow curves of middle slice y1 and y2/y3 in Fig. 4c.

For the bottom section, a lower but uniform hardness value can be noticed. Likewise, lower strengths but alignment of curves were noticed for the true stress-strain graphs of the bottom section (z1, z2, and z3).

It is evident from true stress-strain curves that ultimate strength followed exactly the same trend as found in the case of hardness distribution. Overall, the central, and top zones show higher
## Results and Discussion

This section provides details information about the true strain-strain curves of the FSC billet of AA2011 and AA 7075 manufactured through single step and multi-step methods. AA2011 single billet that has been already analyzed by authors of the current study [16], is taken as reference for detailed discussion and then influence of change in material and process variants are analyzed. Further, FSC billets plastic material behavior is then compared to as-received material ones to understand the changes made by FSC process in the mechanical properties of the materials. In addition, a correlation is also established between strength and hardness of FSC billet by analyzing the ultimate strength and hardness values for the zones of the interest particularly in top, middle and bottom layers of the FSC billet.

### True-stress strain analysis of AA2011 single step FSC billet

This section gives details information about the true strain-strain curves of the FSC 2011 billet with evidence of how far true stress-strain graphs agree with the hardness distribution profile across the FSC billet section. Further, a comparison was also made between the true stress-strain curve FSC recycled billet and as-received material AA2011 bar.

On the top slice (Fig. 4 a and b), the central sample (x1) showed higher strength compared to the remaining upsetting samples (x2 to x5). The flow curves of the remaining four were found almost aligned with each other. The reason was samples x2 to x5 were extracted from the same height (top slice) and same radial distance (r=10 mm). Interestingly, the variation in properties is enabled by FSC process mechanics, i.e., different strain rate and temperature distributions across the billet section, as reported by Tang et al. [14]. In addition, a preliminary analysis has been provided by Puleo et al. [17] but it is quite evident that the die wall and back plate serve as heat sinks, reducing the temperature. Due to this, chips close to the die wall and back plate don't receive sufficient heat input and deformation to be fully recrystallized, but some bonding is, however, achieved.

Likewise, the top slice flow curves, almost the same trend was found for the middle slice. The central specimen (y1) showed higher strength than the other samples y2 and y3 (Fig. 4c).

For all samples of the bottom slice, almost all flow curves were found aligned with each other indicating less variation in mechanical properties. It is evident that the strength of the material varies both in the radial and longitudinal direction of the billet. Further, the top part of the billet shows higher yield strength than the bottom part, and along the radial direction the central part is stronger than the region near the external surface. Also, along the radial direction, the difference in true stress-strain curves between the samples at the center (x1) and the other samples (x2 and x3) is higher for the top slice, then difference diminishes at the middle slices, and finally, all curves are aligned with each other for the bottom slice (Fig. 4c).

Interestingly, the trend of true stress-strain profile was found consistent with the hardness distribution trend (Fig. 4c and d). For example, the zone of the billet section that exhibited higher hardness value also showed higher strength. The hardness measurements were performed on the same material, AA2011, manufactured under the same process conditions.

Again in Fig. 4d the middle of the billet section shows a minor variation in hardness value along the radial direction. A similar relationship can be found for the flow curves of middle slice y1 and y2/y3 in Fig. 4c.

For the bottom section, a lower but uniform hardness value can be noticed. Likewise, lower strengths but alignment of curves were noticed for the true stress-strain graphs of the bottom section (z1, z2, and z3).

It is evident from true stress-strain curves that ultimate strength followed exactly the same trend as found in the case of hardness distribution. Overall, the central, and top zones show higher