where r is the reduction ratio. The last method considered was proposed by Avitzur [5], applying the Upper-Bound approach to the drawing problem, and the formulation is shown in Eq. 9. It involved the assumption that the wire can be divided into three zones, in each of which the velocity field is assumed to be continuous: the inlet and exit zones of the die are characterized by the lack of deformation in the wire and the axial component of the velocity is uniform, while in the central zone the plastic flow occurs and the velocity field is directed toward the cone's apex with axial cylindrical symmetry.

$$ \sigma_d = \bar{\sigma}_m \frac{2f(\alpha) \ln\left(\frac{R_i}{R_f}\right) + \frac{2}{\sqrt{3}} \left(\frac{\alpha}{\sin^2(\alpha)} - \cot(\alpha)\right) + 2\mu \left[\cot(\alpha)(1 - \ln\left(\frac{R_i}{R_f}\right))\ln\left(\frac{R_i}{R_f}\right) + \frac{L}{R_f}\right]}{\left(1 + 2\mu \frac{L}{R_f}\right)} \quad (9) $$

where f(α) is a function of the die's semi-angle calculated as in Eq.7,

$$ f(\alpha) = \frac{1}{\sin^2(\alpha)} \left[ 1 - \cos(\alpha) \sqrt{1 - \frac{11}{12} \sin^2(\alpha)} + \frac{1}{\sqrt{11,12}} \ln \left( \frac{1 + \frac{11}{12}}{\frac{11}{12} \cos(\alpha) + \sqrt{1 - \frac{11}{12} \sin^2(\alpha)}} \right) \right] \quad (10) $$

All three analytical models involve the effects on the wire caused by uniform work due to the reduction ratio, non-uniform or redundant work due to the wire distortion, and friction work modeled according to Coulomb model. However, the SLAB and Wistreich-Wright models don't account for the additional friction work due to the presence of the bearing length downstream of the drawing cone.

## Numerical model

The simulation of the wire drawing process was performed using the commercial FE code Qform UK® and was represented as axisymmetric Lagrangian model. Accordingly, the following boundary conditions were set: symmetry along z axis and a constant velocity applied at the wire's end.

<figure><img src="image_723.png" /><figcaption>Fig. 5 Numerical simulation of the wire drawing process. Example of the simulation on the fourth drawing pass. On the left the distribution of stress on the wire [MPa]. On the right the strain rate field in the drawing die [1/s].</figcaption></figure>

The geometry of the dies was modeled as described in the experimental procedures and a wire length of 600 mm was simulated to be drawn in four consecutive operations corresponding to the four passes of the drawing experimentation. The flow stress of the copper wire was modeled with the equation regressed from the torsion tests, while the other material parameters were listed in Table 2, and the Coulomb model was used for friction.
<table><caption>Table 2 Material parameters for the ETP Copper and the polycrystalline diamond.</caption><thead><tr><td>Material Properties</td><td>Wire ETP Copper</td><td>Die PCD</td></tr></thead><tbody><tr><td>Density [Kg/m³]</td><td>8940</td><td>4000</td></tr><tr><td>Specific heat [J/(kg·K)]</td><td>390</td><td>502</td></tr><tr><td>Thermal conductivity [W/(m·K)]</td><td>390</td><td>600</td></tr><tr><td>Thermal expansion [1/°C]</td><td>1.60635e-5</td><td>5.2e-7</td></tr><tr><td>Young's modulus [MPa]</td><td>120000</td><td>900000</td></tr><tr><td>Poisson's ratio</td><td>0.35</td><td>0.22</td></tr><tr><td>Yield stress [MPa]</td><td></td><td>6900</td></tr></tbody></table>

## Results and discussion

The wire drawing forces were calculated applying the analytical models, considering the drawing stress on the wire multiplied by the wire section at the exit of each die. The flow stress of the material deriving from the experimental tests was used in the calculation, considering the average equivalent strain as the natural logarithm of the ratio between inlet and outlet diameter at each step, and the average strain rate as the product of the strain and drawing speed divided by the length of the deformation zone. Then, by comparing the analytically calculated forces with the experimental results of the drawing tests, a range of friction coefficient values was calculated for the unlubricated and lubricated drawing condition, for each model. The range of values in which the Coulomb friction coefficient can be included was estimated considering the maximum value and the minimum value calculated from all the tests in unlubricated conditions at speeds of 60 and 1200 mm/min and from all tests in the lubricated condition, considering that the average absolute relative error between experimental and calculated values was less than 2%. Fig. 6 shows the comparison between the drawing forces calculated by the SLAB method and the experimental data. The experimental data are indicated with the solid thick lines, while the analytical values calculated with the mean coefficient friction are indicated with dashed lines and the bar error represent the values calculated with the relative estimated range of friction coefficient. For the SLAB method it takes values between 0.1-0.16 for the unlubricated condition and between 0.055-0.1 for the lubricated condition, as indicated in Fig. 11.

<figure><img src="image_716.png" /><figcaption>Fig. 6 Comparison between the drawing forces calculated by the SLAB method and the experimental (unlubricated and lubricated conditions, speeds of 60 and 1200 mm/min. The experimental data are indicated with the solid thick lines, and the analytical values calculated with the mean coefficient friction are indicated with dashed lines. The red lines represent the first pass of reduction ration, the green lines the second step, the blue ones the third pass and the yellow the fourth step.</figcaption></figure>
<table><caption>Table 2 Material parameters for the ETP Copper and the polycrystalline diamond.</caption><thead><tr><td>Material Properties</td><td>Wire ETP Copper</td><td>Die PCD</td></tr></thead><tbody><tr><td>Density [Kg/m³]</td><td>8940</td><td>4000</td></tr><tr><td>Specific heat [J/(kg·K)]</td><td>390</td><td>502</td></tr><tr><td>Thermal conductivity [W/(m·K)]</td><td>390</td><td>600</td></tr><tr><td>Thermal expansion [1/°C]</td><td>1.60635e-5</td><td>5.2e-7</td></tr><tr><td>Young's modulus [MPa]</td><td>120000</td><td>900000</td></tr><tr><td>Poisson's ratio</td><td>0.35</td><td>0.22</td></tr><tr><td>Yield stress [MPa]</td><td></td><td>6900</td></tr></tbody></table>

## Results and discussion

The wire drawing forces were calculated applying the analytical models, considering the drawing stress on the wire multiplied by the wire section at the exit of each die. The flow stress of the material deriving from the experimental tests was used in the calculation, considering the average equivalent strain as the natural logarithm of the ratio between inlet and outlet diameter at each step, and the average strain rate as the product of the strain and drawing speed divided by the length of the deformation zone. Then, by comparing the analytically calculated forces with the experimental results of the drawing tests, a range of friction coefficient values was calculated for the unlubricated and lubricated drawing condition, for each model. The range of values in which the Coulomb friction coefficient can be included was estimated considering the maximum value and the minimum value calculated from all the tests in unlubricated conditions at speeds of 60 and 1200 mm/min and from all tests in the lubricated condition, considering that the average absolute relative error between experimental and calculated values was less than 2%. Fig. 6 shows the comparison between the drawing forces calculated by the SLAB method and the experimental data. The experimental data are indicated with the solid thick lines, while the analytical values calculated with the mean coefficient friction are indicated with dashed lines and the bar error represent the values calculated with the relative estimated range of friction coefficient. For the SLAB method it takes values between 0.1-0.16 for the unlubricated condition and between 0.055-0.1 for the lubricated condition, as indicated in Fig. 11.

<figure><img src="image_724.png" /><figcaption>Fig. 6 Comparison between the drawing forces calculated by the SLAB method and the experimental (unlubricated and lubricated conditions, speeds of 60 and 1200 mm/min. The experimental data are indicated with the solid thick lines, and the analytical values calculated with the mean coefficient friction are indicated with dashed lines. The red lines represent the first pass of reduction ration, the green lines the second step, the blue ones the third pass and the yellow the fourth step.</figcaption></figure>
<figure><img src="image_717.png" /><figcaption>Fig. 7 Comparison between the drawing forces calculated by the Wistreich-Wright model and experimental data.</figcaption></figure>

In Fig. 7 and Fig. 8 the values of the wire drawing forces calculated using respectively the Wistreich-Wright model and the Avitzur model are shown. In the Wistreich-Wright the friction coefficient assumes the values between 0.07-0.1 for the unlubricated condition and between 0.048-0.75 for the lubricated condition. While it takes the values between 0.055-0.095 for the unlubricated condition and 0.02-0.052 for the lubricated condition in the Avitzur model. The comparison between the values assumed by the friction coefficient in the different analytical and numerical models is shown in Fig. 11. The same method used for calculating the friction coefficient in the analytical models was applied in the numerical simulation and it was found that the friction coefficient is within a range between 0.14-0.17 for the unlubricated condition and between 0.11 and 0.13 for the lubricated condition. The results are shown in Fig. 9 and Fig. 10 where the drawing force values obtained from the numerical simulations with the maximum value of the friction coefficient are shown with dashed lines and the values obtained with the minimum friction coefficient with dash-dot lines.
<figure><img src="image_725.png" /><figcaption>Fig. 7 Comparison between the drawing forces calculated by the Wistreich-Wright model and experimental data.</figcaption></figure>

In Fig. 7 and Fig. 8 the values of the wire drawing forces calculated using respectively the Wistreich-Wright model and the Avitzur model are shown. In the Wistreich-Wright the friction coefficient assumes the values between 0.07-0.1 for the unlubricated condition and between 0.048-0.75 for the lubricated condition. While it takes the values between 0.055-0.095 for the unlubricated condition and 0.02-0.052 for the lubricated condition in the Avitzur model. The comparison between the values assumed by the friction coefficient in the different analytical and numerical models is shown in Fig. 11. The same method used for calculating the friction coefficient in the analytical models was applied in the numerical simulation and it was found that the friction coefficient is within a range between 0.14-0.17 for the unlubricated condition and between 0.11 and 0.13 for the lubricated condition. The results are shown in Fig. 9 and Fig. 10 where the drawing force values obtained from the numerical simulations with the maximum value of the friction coefficient are shown with dashed lines and the values obtained with the minimum friction coefficient with dash-dot lines.
<figure><img src="image_718.png" /><figcaption>Fig. 8 Comparison between the drawing forces calculated by Avitzur model and the experimental data.</figcaption></figure>

<figure><img src="image_719.png" /><figcaption>Fig. 9 Comparison between forces predicted by simulation and experimental data in unlubricated conditions.</figcaption></figure>

It was found that in the Avitzur and Wistreich-Wright modeling the friction coefficient takes on low values, since the Avitzur model also involves the effect of the friction between the wire and the die in the bearing length, while in the Wright model, the average pressure inside the die is calculated through the semi-empirical approximation of the redundant work factor. Therefore, in both models, increasing the friction coefficient value leads to an overestimation of the drawing forces. On the contrary, in the SLAB method the friction coefficient takes on higher values and its variation has less influence on the calculation of the forces, falling within a wider range of values. Finally, it was found that the friction coefficient assumes higher values in the numerical simulation and varies in small ranges. In conclusion, the numerical model provides good reliability, inserting values of the friction coefficient higher than those of the analytical models, since in the analytical models simplified conditions of average strain and strain rates in the die are considered. It is worth noting that further studies are necessary to achieve greater reliability, involving hydrodynamic effects due to high drawing speeds and temperature increases.
<figure><img src="image_726.png" /><figcaption>Fig. 8 Comparison between the drawing forces calculated by Avitzur model and the experimental data.</figcaption></figure>

<figure><img src="image_727.png" /><figcaption>Fig. 9 Comparison between forces predicted by simulation and experimental data in unlubricated conditions.</figcaption></figure>

It was found that in the Avitzur and Wistreich-Wright modeling the friction coefficient takes on low values, since the Avitzur model also involves the effect of the friction between the wire and the die in the bearing length, while in the Wright model, the average pressure inside the die is calculated through the semi-empirical approximation of the redundant work factor. Therefore, in both models, increasing the friction coefficient value leads to an overestimation of the drawing forces. On the contrary, in the SLAB method the friction coefficient takes on higher values and its variation has less influence on the calculation of the forces, falling within a wider range of values. Finally, it was found that the friction coefficient assumes higher values in the numerical simulation and varies in small ranges. In conclusion, the numerical model provides good reliability, inserting values of the friction coefficient higher than those of the analytical models, since in the analytical models simplified conditions of average strain and strain rates in the die are considered. It is worth noting that further studies are necessary to achieve greater reliability, involving hydrodynamic effects due to high drawing speeds and temperature increases.
<figure><img src="image_720.png" /><figcaption>Fig. 10 Comparison between forces predicted by simulation and experimental data in lubricated conditions.</figcaption></figure>

<figure><img src="image_721.png" /><figcaption>Fig. 11 Comparison between the range of the friction coefficient values calculated from the three analytical models and the numerical method, for the unlubricated and lubricated conditions.</figcaption></figure>

## Summary

In this paper the ETP Copper wire multi-pass drawing process was investigated, comparing the forces applied to the wire in experimental drawing tests with the values obtained from the analytical calculation according to three different models and from the numerical simulation. The experimental procedures involved wire drawing tests carried out on four steps of the process, measuring the drawing force applied to the wire by varying two speed values in lubricated and non-lubricated conditions. Moreover, the material characterization was performed through torsion tests, and a material flow stress law dependent on the strain and the strain rate was provided. The drawing forces were calculated with the SLAB analytical method, the Wistreich-Wright and the Avitzur model. By comparing the experimental data and the calculated values, a range of friction coefficient values in the lubricated and unlubricated condition was provided for each analytical models considering a maximum calculation error of 2%. A reliable model for the numerical simulation of the multi-pass drawing process was provided using the commercial FE code Qform UK®, and it was shown that the analytical models underestimate the value of the friction coefficient, due to the assumed simplified conditions of average strain and strain rate of the material within the die. Summarizing, it was found a good accuracy between experimental data and predicted values of the drawing forces acting on the wire as the reduction ratio and speed vary, but further studies are necessary to include in the model the hydrodynamic effects due to the industrial high drawing speeds.
<figure><img src="image_728.png" /><figcaption>Fig. 10 Comparison between forces predicted by simulation and experimental data in lubricated conditions.</figcaption></figure>

<figure><img src="image_729.png" /><figcaption>Fig. 11 Comparison between the range of the friction coefficient values calculated from the three analytical models and the numerical method, for the unlubricated and lubricated conditions.</figcaption></figure>

## Summary

In this paper the ETP Copper wire multi-pass drawing process was investigated, comparing the forces applied to the wire in experimental drawing tests with the values obtained from the analytical calculation according to three different models and from the numerical simulation. The experimental procedures involved wire drawing tests carried out on four steps of the process, measuring the drawing force applied to the wire by varying two speed values in lubricated and non-lubricated conditions. Moreover, the material characterization was performed through torsion tests, and a material flow stress law dependent on the strain and the strain rate was provided. The drawing forces were calculated with the SLAB analytical method, the Wistreich-Wright and the Avitzur model. By comparing the experimental data and the calculated values, a range of friction coefficient values in the lubricated and unlubricated condition was provided for each analytical models considering a maximum calculation error of 2%. A reliable model for the numerical simulation of the multi-pass drawing process was provided using the commercial FE code Qform UK®, and it was shown that the analytical models underestimate the value of the friction coefficient, due to the assumed simplified conditions of average strain and strain rate of the material within the die. Summarizing, it was found a good accuracy between experimental data and predicted values of the drawing forces acting on the wire as the reduction ratio and speed vary, but further studies are necessary to include in the model the hydrodynamic effects due to the industrial high drawing speeds.
## References

[1] G.E. Dieter, S.L. Semiatin, H.A. Kuhn, Handbook of Workability and Process Design, ASM International publications, United States of America 2003.

[2] K. Lange, Fundamentals of technical plasticity theory, in: Handbook of Metal Forming, first ed., Kurt Lange, Society of Manufacturing Engineers, 1985, pp: 99-132.

[3] J.G. Wistreich, The fundamentals of wire drawing, Metall Rev 3(10) (1958) 97-142.

[4] R. N. Wright, Basic Drawing Mechanics, in: Wire Technology-Process engineering and metallurgy, second ed., Butterworth-Heinemann, 2016, pp: 28-42.

[5] B. Avitzur, C. Narayan, Y.T. Chou, Upper-bound solutions for flow through conical converging dies, International Journal of Machine Tool Design and Research 22(3) (1982) 197-214.

[6] L. Donati, B. Reggiani, R. Pelaccia, M. Negozio and S. Di Donato, Advancements in extrusion and drawing: a review of the contributes by the ESAFORM community, International Journal of Material 15, 41 (2022). https://doi.org/10.1007/s12289-022-01664-w

[7] G. Vega, A. Haddi, A. Imad, Investigation of process parameters effect on the copper-wiredrawing, Materials & Design 30(8) (2009) 3308-3312. https://doi.org/10.1016/j.matdes.2008.12.006

[8] O. Rodríguez-Alabanda, E. Molero, M. Tintelecan, G. Guerrero-Vaca, PE. Romero, GA. Santana Martinez, Fine Electrolytic Tough Pitch Copper Multistage Wiredrawing Pass Schedule Design by Analytical and Numerical Methods, Proceedings. (2020); 63(1):12. https://doi.org/10.3390/proceedings2020063012

[9] L. Lazzarotto, L. Dubar, A. Dubois, P. Ravassard, J. Oudin, Identification of Coulomb's friction coefficient in real contact conditions applied to a wire drawing process, Wear 211(1) (1997) 54-63. https://doi.org/10.1016/S0043-1648(97)00080-X

[10] K.K. Tang, Z.X. Li, J. Wang, Numerical simulation of damage evolution in multi-pass wire drawing process and its applications, Materials & Design, 32, 6, (2011) 3299-3311, https://doi.org/10.1016/j.matdes.2011.02.026

[11] A. Panteghini, F. Genna, Effects of the strain-hardening law in the numerical simulation of wire drawing processes, Computational Materials Science, 49, 2, (2010), 236-242, https://doi.org/10.1016/j.commatsci.2010.05.002

[12] H. M. Baek, Y. G. Jin, S. K. Hwang, Y. T. Im, I. H. Son, D. L. Lee, Numerical study on the evolution of surface defects in wire drawing, Journal of Materials Processing Technology, 212, 4, (2012), 776-785, https://doi.org/10.1016/j.jmatprotec.2011.10.028

[13] G.R. Canova, U.F. Kocks, J.J. Jonas, Theory of torsion texture development, Acta metall., 32,2 (1984) 211-226. https://doi.org/10.1016/0001-6160(84)90050-6

[14] A. Gräber, K. Pöhlandt, State of the art of the torsion test for determining flow curves, Steel Res., 61 (1990) 212-218.

[15] S. Di Donato, L. Donati, M. Negozio, Copper Wire Multi-Pass Drawing: Process Modeling and Optimization, KEM 926 (2022) 499-510. https://doi.org/10.4028/p-g4wbpz

[16] D.S. Fields, W.A. Backofen, Determination of strain-hardening characteristics by torsion testing, Proc. Am. Soc. Test. Mater. 57 (1957) 1259-1272.

[17] M. El Mehtedi, S. Spigarelli, F. Gabrielli, L. Donati, Comparison Study of Constitutive Models in Predicting the Hot Deformation Behavior of AA6060 and AA6063 Aluminium Alloys, Materials Today: Proceedings, part A 2,10, (2015) 4732-4739. https://doi.org/10.1016/j.matpr.2015.10.006

[18] S. Di Donato, R. Pelaccia, M. Negozio, M. El Mehtedi, B. Reggiani, L. Donati, Hot torsion tests of AA6082 alloy, KEM, under publication.
## References

[1] G.E. Dieter, S.L. Semiatin, H.A. Kuhn, Handbook of Workability and Process Design, ASM International publications, United States of America 2003.

[2] K. Lange, Fundamentals of technical plasticity theory, in: Handbook of Metal Forming, first ed., Kurt Lange, Society of Manufacturing Engineers, 1985, pp: 99-132.

[3] J.G. Wistreich, The fundamentals of wire drawing, Metall Rev 3(10) (1958) 97-142.

[4] R. N. Wright, Basic Drawing Mechanics, in: Wire Technology-Process engineering and metallurgy, second ed., Butterworth-Heinemann, 2016, pp: 28-42.

[5] B. Avitzur, C. Narayan, Y.T. Chou, Upper-bound solutions for flow through conical converging dies, International Journal of Machine Tool Design and Research 22(3) (1982) 197-214.

[6] L. Donati, B. Reggiani, R. Pelaccia, M. Negozio and S. Di Donato, Advancements in extrusion and drawing: a review of the contributes by the ESAFORM community, International Journal of Material 15, 41 (2022). https://doi.org/10.1007/s12289-022-01664-w

[7] G. Vega, A. Haddi, A. Imad, Investigation of process parameters effect on the copper-wiredrawing, Materials & Design 30(8) (2009) 3308-3312. https://doi.org/10.1016/j.matdes.2008.12.006

[8] O. Rodríguez-Alabanda, E. Molero, M. Tintelecan, G. Guerrero-Vaca, PE. Romero, GA. Santana Martinez, Fine Electrolytic Tough Pitch Copper Multistage Wiredrawing Pass Schedule Design by Analytical and Numerical Methods, Proceedings. (2020); 63(1):12. https://doi.org/10.3390/proceedings2020063012

[9] L. Lazzarotto, L. Dubar, A. Dubois, P. Ravassard, J. Oudin, Identification of Coulomb's friction coefficient in real contact conditions applied to a wire drawing process, Wear 211(1) (1997) 54-63. https://doi.org/10.1016/S0043-1648(97)00080-X

[10] K.K. Tang, Z.X. Li, J. Wang, Numerical simulation of damage evolution in multi-pass wire drawing process and its applications, Materials & Design, 32, 6, (2011) 3299-3311, https://doi.org/10.1016/j.matdes.2011.02.026

[11] A. Panteghini, F. Genna, Effects of the strain-hardening law in the numerical simulation of wire drawing processes, Computational Materials Science, 49, 2, (2010), 236-242, https://doi.org/10.1016/j.commatsci.2010.05.002

[12] H. M. Baek, Y. G. Jin, S. K. Hwang, Y. T. Im, I. H. Son, D. L. Lee, Numerical study on the evolution of surface defects in wire drawing, Journal of Materials Processing Technology, 212, 4, (2012), 776-785, https://doi.org/10.1016/j.jmatprotec.2011.10.028

[13] G.R. Canova, U.F. Kocks, J.J. Jonas, Theory of torsion texture development, Acta metall., 32,2 (1984) 211-226. https://doi.org/10.1016/0001-6160(84)90050-6

[14] A. Gräber, K. Pöhlandt, State of the art of the torsion test for determining flow curves, Steel Res., 61 (1990) 212-218.

[15] S. Di Donato, L. Donati, M. Negozio, Copper Wire Multi-Pass Drawing: Process Modeling and Optimization, KEM 926 (2022) 499-510. https://doi.org/10.4028/p-g4wbpz

[16] D.S. Fields, W.A. Backofen, Determination of strain-hardening characteristics by torsion testing, Proc. Am. Soc. Test. Mater. 57 (1957) 1259-1272.

[17] M. El Mehtedi, S. Spigarelli, F. Gabrielli, L. Donati, Comparison Study of Constitutive Models in Predicting the Hot Deformation Behavior of AA6060 and AA6063 Aluminium Alloys, Materials Today: Proceedings, part A 2,10, (2015) 4732-4739. https://doi.org/10.1016/j.matpr.2015.10.006

[18] S. Di Donato, R. Pelaccia, M. Negozio, M. El Mehtedi, B. Reggiani, L. Donati, Hot torsion tests of AA6082 alloy, KEM, under publication.
# Prediction of longitudinal seam weld quality in multi-chamber aluminium extrusion

SARIYARLIOGLU Eren Can¹,ª, WELO Torgeir¹,ᵇ and MA Jun¹,ᶜ*

¹Department of Mechanical and Industrial Engineering, Norwegian University of Science and Technology, Trondheim 7491, Norway

ªeren.c.sariyarlioglu@ntnu.no, ᵇtorgeir.welo@ntnu.no, ᵈjun.ma@ntnu.no

**Keywords:** Aluminium Extrusion, Longitudinal Seam Weld, Multi-Chamber Profile

**Abstract.** In the extrusion of complex multi-chamber profiles, the longitudinal seam (L-seam) weld quality is significantly affected by the thermo-mechanical history and the 'local' formation conditions under which the weld is produced during the extrusion process. The L-seam welds can lead to non-uniform mechanical properties and variations with regard to the mechanical integrity of extruded products. Therefore, effective prediction and analysis of seam weld quality is crucial to quality insurance of extrusion products for advanced applications. This study deals with the prediction of L-seam weld quality and associated mechanism analysis in the extrusion of multi-chamber aluminium profiles to advance the understanding of how the 'local' thermo-mechanical history and conditions influence the weld quality. To this end, carefully controlled industrial-scale extrusion experiments were conducted to analyze the L-seam formation behaviour. Additionally, a finite element (FE) model was developed to investigate the thermo-mechanical history during the formation of L-seam welds. The model was also employed to predict the L-seam weld quality, incorporating several welding quality criteria, namely Q, K, and J criteria. Examination of the thermo-mechanical history within the welding chamber reveals a compromise in the welding quality at the centre of the L-seam compared to near the surfaces. Moreover, the findings highlight that the J-criterion can provide the most consistent predictions for the L-seam quality, providing a more accurate reflection of the thermo-mechanical history during the solid-bonding process than the Q and K criteria. The findings can contribute to the development of process guidelines and quality assurance strategies for producing multi-chamber aluminium extrusion profiles.

**Introduction**

Extrusion provides opportunities for producing complex multi-chamber sections, offering high functionality in terms of lightweight, structural integrity, etc [1]. However, meeting demanding industrial quality standards necessitates a thorough understanding and effective control of the mechanisms influencing longitudinal seam welds. In the extrusion of multi-chamber sections, the material undergoes division into multiple streams by die bridges, subsequently rejoining in the welding chamber to form longitudinal seam (L-seam) welds extending in the extrusion direction [2]. The weld quality of different L-seams can vary significantly particularly for complex-shaped cross-sections, due to the unique and inhomogeneous material flow characteristics. The welding-induced non-uniform mechanical properties on extruded profiles pose significant challenges, especially for structural applications such as the crashworthiness of automotive battery protection systems.

At present, experimental techniques are widely used to evaluate the quality of L-seam welds in extrusions, including tensile tests and fatigue tests [2–6], as well as wedge and bulge expansion tests [8]. While these methods provide valuable initial insights, their limitations include significant cost, time consumption, and inability to trace formation history, particularly for complex profile shapes. Consequently, both analytical and numerical investigations are crucial for establishing a

<figure><img src="image_722.png" /></figure>
# Prediction of longitudinal seam weld quality in multi-chamber aluminium extrusion

SARIYARLIOGLU Eren Can¹,ª, WELO Torgeir¹,ᵇ and MA Jun¹,ᶜ*

¹Department of Mechanical and Industrial Engineering, Norwegian University of Science and Technology, Trondheim 7491, Norway

ªeren.c.sariyarlioglu@ntnu.no, ᵇtorgeir.welo@ntnu.no, ᶜjun.ma@ntnu.no

**Keywords:** Aluminium Extrusion, Longitudinal Seam Weld, Multi-Chamber Profile

**Abstract.** In the extrusion of complex multi-chamber profiles, the longitudinal seam (L-seam) weld quality is significantly affected by the thermo-mechanical history and the 'local' formation conditions under which the weld is produced during the extrusion process. The L-seam welds can lead to non-uniform mechanical properties and variations with regard to the mechanical integrity of extruded products. Therefore, effective prediction and analysis of seam weld quality is crucial to quality insurance of extrusion products for advanced applications. This study deals with the prediction of L-seam weld quality and associated mechanism analysis in the extrusion of multi-chamber aluminium profiles to advance the understanding of how the 'local' thermo-mechanical history and conditions influence the weld quality. To this end, carefully controlled industrial-scale extrusion experiments were conducted to analyze the L-seam formation behaviour. Additionally, a finite element (FE) model was developed to investigate the thermo-mechanical history during the formation of L-seam welds. The model was also employed to predict the L-seam weld quality, incorporating several welding quality criteria, namely Q, K, and J criteria. Examination of the thermo-mechanical history within the welding chamber reveals a compromise in the welding quality at the centre of the L-seam compared to near the surfaces. Moreover, the findings highlight that the J-criterion can provide the most consistent predictions for the L-seam quality, providing a more accurate reflection of the thermo-mechanical history during the solid-bonding process than the Q and K criteria. The findings can contribute to the development of process guidelines and quality assurance strategies for producing multi-chamber aluminium extrusion profiles.

**Introduction**

Extrusion provides opportunities for producing complex multi-chamber sections, offering high functionality in terms of lightweight, structural integrity, etc [1]. However, meeting demanding industrial quality standards necessitates a thorough understanding and effective control of the mechanisms influencing longitudinal seam welds. In the extrusion of multi-chamber sections, the material undergoes division into multiple streams by die bridges, subsequently rejoining in the welding chamber to form longitudinal seam (L-seam) welds extending in the extrusion direction [2]. The weld quality of different L-seams can vary significantly particularly for complex-shaped cross-sections, due to the unique and inhomogeneous material flow characteristics. The welding-induced non-uniform mechanical properties on extruded profiles pose significant challenges, especially for structural applications such as the crashworthiness of automotive battery protection systems.

At present, experimental techniques are widely used to evaluate the quality of L-seam welds in extrusions, including tensile tests and fatigue tests [2–6], as well as wedge and bulge expansion tests [8]. While these methods provide valuable initial insights, their limitations include significant cost, time consumption, and inability to trace formation history, particularly for complex profile shapes. Consequently, both analytical and numerical investigations are crucial for establishing a

<figure><img src="image_730.png" /></figure>
more comprehensive understanding of the factors influencing L-seam quality, particularly the impact of thermo-mechanical extrusion history.

For theoretical analysis of L-seam welds, several criteria have been proposed to evaluate L-seam quality, including pressure-dependent criteria [8, 9], the pressure-time criterion (Q-criterion) [11], the pressure-time-flow criterion (K-criterion) [12], and the pressure-material surface-strain rate criterion (J-criterion) [7].

The Q, K and J criteria are briefly described as follows:

Q-criterion is defined by Eq. (1) [11]:

$$Q = \int_{0}^{t} \frac{\sigma_{m}}{\bar{\sigma}} dt \qquad (Eq. 1)$$

where t represents time, $\sigma_m$ is the mean stress, and $\bar{\sigma}$ is the effective stress.

K-criterion is expressed by Eq. (2) [12]:

$$K = \int_{0}^{t} \frac{\sigma_{m}}{\bar{\sigma}} v dt \qquad (Eq. 2)$$

where, v represents the material flow speed.

J-criterion is expressed by Eq. (3) [7]:

$$J = \int_{0}^{t} k_{0} \frac{\sigma_{m}}{\bar{\sigma}} \dot{\bar{\epsilon}} \exp \left( \frac{RT}{Q_{D}} \right) dt \qquad (Eq. 3)$$

where $\dot{\bar{\epsilon}}$ is the effective strain rate, $Q_D$ is the activation energy, R is the universal gas constant, and $k_0$ is the material and surface condition constant.

From the provided equations, it is noticed that the J-criterion differs from the Q and K criteria because of its dependence on strain rate and temperature, suggesting that a higher temperature and strain rate can improve the L-seam quality. However, Donati and Tomesani [13] reported that higher extrusion velocity can reduce L-seam quality due to the reduced welding time. Therefore, there is still a need for more experimental and theoretical studies to assess the J-criterion. Furthermore, Donati and Tomesani [12] reported that the K-criterion predicts L-seam quality better than the Q-criterion. Additionally, Kniazkin and Vlasov [14] proposed a modified Q-criterion (combined with the FE method) that offers improved accuracy compared to the K-criterion. Nevertheless, the exploration of the modified Q-criterion remains limited, necessitating experimental validation.

The understanding of the influence of thermo-mechanical history on the quality of L-seam welds during the aluminium extrusion process remains limited, despite the research available in the literature on L-seam weld quality. Furthermore, there is a need for additional analyses to enhance the accuracy of predicting L-seam weld quality. Consequently, this paper focuses on the modelling of L-seam welds and the prediction of their quality in multi-chamber aluminium profiles. The findings presented in this paper aim to provide practical guidance for both researchers and extruders seeking a better comprehension of the L-seam welding mechanism.

## Extrusion Experiments and Modelling

Industrial-scale extrusion experiments. A series of full-scale industrial extrusion experiments were undertaken at Benteler Automotive Raufoss AS (Raufoss, Norway) using a 55 MN extrusion press for producing aluminium alloy multi-chamber profiles. Table 1 shows the process parameters used in the extrusion experiments. The geometric dimension of the extruded profile is illustrated in Fig. 1. In this figure, the die and mandrel design employed in the extrusion experiments are also illustrated, featuring a die with eight portholes. The billet material is AA6005, prepared through the DC-casting method and homogenized before the extrusion process.
more comprehensive understanding of the factors influencing L-seam quality, particularly the impact of thermo-mechanical extrusion history.

For theoretical analysis of L-seam welds, several criteria have been proposed to evaluate L-seam quality, including pressure-dependent criteria [8, 9], the pressure-time criterion (Q-criterion) [11], the pressure-time-flow criterion (K-criterion) [12], and the pressure-material surface-strain rate criterion (J-criterion) [7].

The Q, K and J criteria are briefly described as follows:

Q-criterion is defined by Eq. (1) [11]:

$$Q = \int_{0}^{t} \frac{\sigma_{m}}{\bar{\sigma}} dt \qquad (Eq. 1)$$

where t represents time, $\sigma_m$ is the mean stress, and $\bar{\sigma}$ is the effective stress.

K-criterion is expressed by Eq. (2) [12]:

$$K = \int_{0}^{t} \frac{\sigma_{m}}{\bar{\sigma}} v dt \qquad (Eq. 2)$$

where, v represents the material flow speed.

J-criterion is expressed by Eq. (3) [7]:

$$J = \int_{0}^{t} k_{0} \frac{\sigma_{m}}{\bar{\sigma}} \dot{\bar{\epsilon}} \exp\left(\frac{RT}{Q_{D}}\right) dt \qquad (Eq. 3)$$

where $\dot{\bar{\epsilon}}$ is the effective strain rate, $Q_D$ is the activation energy, R is the universal gas constant, and $k_0$ is the material and surface condition constant.

From the provided equations, it is noticed that the J-criterion differs from the Q and K criteria because of its dependence on strain rate and temperature, suggesting that a higher temperature and strain rate can improve the L-seam quality. However, Donati and Tomesani [13] reported that higher extrusion velocity can reduce L-seam quality due to the reduced welding time. Therefore, there is still a need for more experimental and theoretical studies to assess the J-criterion. Furthermore, Donati and Tomesani [12] reported that the K-criterion predicts L-seam quality better than the Q-criterion. Additionally, Kniazkin and Vlasov [14] proposed a modified Q-criterion (combined with the FE method) that offers improved accuracy compared to the K-criterion. Nevertheless, the exploration of the modified Q-criterion remains limited, necessitating experimental validation.

The understanding of the influence of thermo-mechanical history on the quality of L-seam welds during the aluminium extrusion process remains limited, despite the research available in the literature on L-seam weld quality. Furthermore, there is a need for additional analyses to enhance the accuracy of predicting L-seam weld quality. Consequently, this paper focuses on the modelling of L-seam welds and the prediction of their quality in multi-chamber aluminium profiles. The findings presented in this paper aim to provide practical guidance for both researchers and extruders seeking a better comprehension of the L-seam welding mechanism.

## Extrusion Experiments and Modelling

Industrial-scale extrusion experiments. A series of full-scale industrial extrusion experiments were undertaken at Benteler Automotive Raufoss AS (Raufoss, Norway) using a 55 MN extrusion press for producing aluminium alloy multi-chamber profiles. Table 1 shows the process parameters used in the extrusion experiments. The geometric dimension of the extruded profile is illustrated in Fig. 1. In this figure, the die and mandrel design employed in the extrusion experiments are also illustrated, featuring a die with eight portholes. The billet material is AA6005, prepared through the DC-casting method and homogenized before the extrusion process.
To characterize and examine the quality of L-seam welds, two representative regions, marked as Region 1 and Region 2 as shown in Fig. 1, were carefully removed from the multi-chamber profile. Subsequently, the extracted samples underwent a series of preparation processes, including grinding with abrasive papers and etching with a sodium hydroxide solution. Ultimately, the L-seam welds became visible to the naked eye.

<figure><img src="image_723.png" /><figcaption>Fig. 1 Produced multi-chamber profile geometry and die design used in the production.</figcaption></figure>

<table><caption>Table 1. Parameters in extrusion experiments.</caption><thead><tr><td>Parameter</td><td>Value</td><td>Unit</td></tr></thead><tbody><tr><td>Billet diameter</td><td>280</td><td>mm</td></tr><tr><td>Billet length</td><td>1400</td><td>mm</td></tr><tr><td>Billet temperature</td><td>480</td><td>°C</td></tr><tr><td>Temperature taper</td><td>40</td><td>°C/m</td></tr><tr><td>Ram speed</td><td>4.7</td><td>mm/s</td></tr><tr><td>Extrusion ratio</td><td>28</td><td>/</td></tr><tr><td>Container temperature</td><td>420</td><td>°C</td></tr><tr><td>Die temperature</td><td>460</td><td>°C</td></tr></tbody></table>

Numerical modelling. To enable effective analysis of the longitudinal seam weld mechanisms in extruded profiles, an FE model for the extrusion process was developed based on QForm-Extrusion 10.2.1 to reproduce the material flow behaviour in the extrusion process [15]. QForm-
To characterize and examine the quality of L-seam welds, two representative regions, marked as Region 1 and Region 2 as shown in Fig. 1, were carefully removed from the multi-chamber profile. Subsequently, the extracted samples underwent a series of preparation processes, including grinding with abrasive papers and etching with a sodium hydroxide solution. Ultimately, the L-seam welds became visible to the naked eye.

<figure><img src="image_731.png" /><figcaption>Fig. 1 Produced multi-chamber profile geometry and die design used in the production.</figcaption></figure>

<table><caption>Table 1. Parameters in extrusion experiments.</caption><thead><tr><td>Parameter</td><td>Value</td><td>Unit</td></tr></thead><tbody><tr><td>Billet diameter</td><td>280</td><td>mm</td></tr><tr><td>Billet length</td><td>1400</td><td>mm</td></tr><tr><td>Billet temperature</td><td>480</td><td>°C</td></tr><tr><td>Temperature taper</td><td>40</td><td>°C/m</td></tr><tr><td>Ram speed</td><td>4.7</td><td>mm/s</td></tr><tr><td>Extrusion ratio</td><td>28</td><td>/</td></tr><tr><td>Container temperature</td><td>420</td><td>°C</td></tr><tr><td>Die temperature</td><td>460</td><td>°C</td></tr></tbody></table>

Numerical modelling. To enable effective analysis of the longitudinal seam weld mechanisms in extruded profiles, an FE model for the extrusion process was developed based on QForm-Extrusion 10.2.1 to reproduce the material flow behaviour in the extrusion process [15]. QForm-
Extrusion, a commercially available FE package, enables the coupling of die deformation and accurate prediction of weld position [16]. The model comprises four groups representing the main components: billet, porthole and welding chamber, bearing, and profile (see Fig. 2). Both the extrusion tools and billet were discretized using 3D tetrahedral 4-noded elements. The material behaviour of AA6005 was represented by an isotropic viscoplastic model using the Hensel-Spittel constitutive equation (Eq. 4) [17]:

$$ \sigma = A e^{m_1 T} T^{m_2} \varepsilon^{m_3} e^{m_4 / \varepsilon} (1 + \varepsilon)^{m_5 T} e^{m_7 \varepsilon} \dot{\varepsilon}^{m_3} \dot{\varepsilon}^{m_8 T} \quad (\text{Eq. 4}) $$

where $\sigma$ is the flow stress, $\varepsilon$ is the strain, $\dot{\varepsilon}$ is the strain rate, $T$ is the temperature, and $A$ and $m_1-m_9$ are the material coefficients (see Table 2).

<figure><img src="image_724.png" /><figcaption>Fig. 2 Numerical model.</figcaption></figure>

<table><caption>Table 2. Hensel-Spittel parameters [15].</caption><thead><tr><th>Coeff.</th><th>A</th><th>m<sub>1</sub></th><th>m<sub>2</sub></th><th>m<sub>3</sub></th><th>m<sub>4</sub></th><th>m<sub>5</sub></th><th>m<sub>7</sub></th><th>m<sub>8</sub></th><th>m<sub>9</sub></th></tr></thead><tbody><tr><th>Value</th><td>320</td><td>-0.00511</td><td>-0.12422</td><td>0.1</td><td>-0.01364</td><td>0.00024</td><td>0</td><td>0</td><td>0</td></tr></tbody></table>

Other physical and thermal properties were incorporated into the model using default parameters available in the software. The computational time required for the model was 7 hours on a 112-core Xeon processor utilizing parallel processing and 256 GB of RAM.

The accuracy of the numerical model was evaluated by comparing the predicted extrusion force to the force measured in physical experiments. The comparison, shown in Fig. 3, indicates that the numerical model accurately predicted the maximum extrusion force with an error of slightly less than 6%, and the disparity in minimum forces was slightly over 4%. These results are deemed accurate and acceptable within the context of this study.
Extrusion, a commercially available FE package, enables the coupling of die deformation and accurate prediction of weld position [16]. The model comprises four groups representing the main components: billet, porthole and welding chamber, bearing, and profile (see Fig. 2). Both the extrusion tools and billet were discretized using 3D tetrahedral 4-noded elements. The material behaviour of AA6005 was represented by an isotropic viscoplastic model using the Hensel-Spittel constitutive equation (Eq. 4) [17]:

$$ \sigma = A e^{m_1 T} T^{m_2} \varepsilon^{m_3} e^{m_4 / \varepsilon} (1 + \varepsilon)^{m_5 T} e^{m_7 \varepsilon} \dot{\varepsilon}^{m_3} \dot{\varepsilon}^{m_8 T} \quad (\text{Eq. 4}) $$

where $\sigma$ is the flow stress, $\varepsilon$ is the strain, $\dot{\varepsilon}$ is the strain rate, $T$ is the temperature, and $A$ and $m_1-m_9$ are the material coefficients (see Table 2).

<figure><img src="image_732.png" /><figcaption>Fig. 2 Numerical model.</figcaption></figure>

<table><caption>Table 2. Hensel-Spittel parameters [15].</caption><thead><tr><th>Coeff.</th><th>A</th><th>m<sub>1</sub></th><th>m<sub>2</sub></th><th>m<sub>3</sub></th><th>m<sub>4</sub></th><th>m<sub>5</sub></th><th>m<sub>7</sub></th><th>m<sub>8</sub></th><th>m<sub>9</sub></th></tr></thead><tbody><tr><th>Value</th><td>320</td><td>-0.00511</td><td>-0.12422</td><td>0.1</td><td>-0.01364</td><td>0.00024</td><td>0</td><td>0</td><td>0</td></tr></tbody></table>

Other physical and thermal properties were incorporated into the model using default parameters available in the software. The computational time required for the model was 7 hours on a 112-core Xeon processor utilizing parallel processing and 256 GB of RAM.

The accuracy of the numerical model was evaluated by comparing the predicted extrusion force to the force measured in physical experiments. The comparison, shown in Fig. 3, indicates that the numerical model accurately predicted the maximum extrusion force with an error of slightly less than 6%, and the disparity in minimum forces was slightly over 4%. These results are deemed accurate and acceptable within the context of this study.
<figure><img src="image_725.png" /><figcaption>Fig. 3 Comparison of extrusion load between extrusion experiment and numerical model.</figcaption></figure>

## Results and Discussion

In hot extrusion of multi-chamber profiles, the material undergoes separation into multiple streams through die webs and subsequent rejoining within the welding chamber, thereby creating L-seam welds along the entire length of the profile. In the case profile in this research, the utilization of eight distinct portholes in the die design results in the formation of thirteen L-seam welds in total, distributing across the entire section of the profile, as illustrated in Fig. 4.

<figure><img src="image_726.png" /><figcaption>Fig. 4 Material flows through portholes and their corresponding L-seam welds.</figcaption></figure>

In this study, the quality of welding was analyzed, with a specific focus on L-seam welds formed through material flows originating from portholes 2-3 and 6-7 (refer to the white arrow in the right figure of Fig. 4). By employing the velocity vectors derived from the extrusion simulation, the flow path of the material was identified, thereby determining the positions of the L-seam welds. The comparison of L-seam weld locations between experimental characterization and numerical simulation is depicted in Fig. 5. It can be observed that the numerical model can effectively capture the positions of L-seam welds in the cross-section of the profile.
<figure><img src="image_733.png" /><figcaption>Fig. 3 Comparison of extrusion load between extrusion experiment and numerical model.</figcaption></figure>

## Results and Discussion

In hot extrusion of multi-chamber profiles, the material undergoes separation into multiple streams through die webs and subsequent rejoining within the welding chamber, thereby creating L-seam welds along the entire length of the profile. In the case profile in this research, the utilization of eight distinct portholes in the die design results in the formation of thirteen L-seam welds in total, distributing across the entire section of the profile, as illustrated in Fig. 4.

<figure><img src="image_734.png" /><figcaption>Fig. 4 Material flows through portholes and their corresponding L-seam welds.</figcaption></figure>

In this study, the quality of welding was analyzed, with a specific focus on L-seam welds formed through material flows originating from portholes 2-3 and 6-7 (refer to the white arrow in the right figure of Fig. 4). By employing the velocity vectors derived from the extrusion simulation, the flow path of the material was identified, thereby determining the positions of the L-seam welds. The comparison of L-seam weld locations between experimental characterization and numerical simulation is depicted in Fig. 5. It can be observed that the numerical model can effectively capture the positions of L-seam welds in the cross-section of the profile.
<figure><img src="image_727.png" /><figcaption>Fig. 5 Comparison of L-seam welds positions on the profile's cross section.</figcaption></figure>

<figure><img src="image_728.png" /><figcaption>Fig. 6 Position of tracing points.</figcaption></figure>

To explore the solid-state bonding process in the selected areas during extrusion, an examination was undertaken by using tracing points positioned along each L-seam weld. Six tracing points were placed on each L-seam weld, and their positions are illustrated in Fig. 6. The examination of the welding path, traced through the backtracking of each point, reveals key thermo-mechanical variables in the solid-state welding process, such as mean stress, effective strain (rate), and temperature. In Fig. 7, the evolutions of thermo-mechanical variables within the welding chamber during extrusion are shown by P1-P6 located in Region 1. Variations in the welding time were noted for tracing points P1 to P6, with durations of 16.7 s, 5.4 s, 3.4 s, 3.2 s, 2.6 s, and 5 s, respectively. These durations represent the time that each tracing point was subjected to the solid-state welding process within the welding chamber, with the reference point (0 s) being the initiation of the welding process. Throughout the solid-state bonding process, the hydrostatic pressure exhibits overall downward trends with increasing welding time. Simultaneously, the
<figure><img src="image_735.png" /><figcaption>Fig. 5 Comparison of L-seam welds positions on the profile's cross section.</figcaption></figure>

<figure><img src="image_736.png" /><figcaption>Fig. 6 Position of tracing points.</figcaption></figure>

To explore the solid-state bonding process in the selected areas during extrusion, an examination was undertaken by using tracing points positioned along each L-seam weld. Six tracing points were placed on each L-seam weld, and their positions are illustrated in Fig. 6. The examination of the welding path, traced through the backtracking of each point, reveals key thermo-mechanical variables in the solid-state welding process, such as mean stress, effective strain (rate), and temperature. In Fig. 7, the evolutions of thermo-mechanical variables within the welding chamber during extrusion are shown by P1-P6 located in Region 1. Variations in the welding time were noted for tracing points P1 to P6, with durations of 16.7 s, 5.4 s, 3.4 s, 3.2 s, 2.6 s, and 5 s, respectively. These durations represent the time that each tracing point was subjected to the solid-state welding process within the welding chamber, with the reference point (0 s) being the initiation of the welding process. Throughout the solid-state bonding process, the hydrostatic pressure exhibits overall downward trends with increasing welding time. Simultaneously, the
effective stresses and effective strain rates initially increase and then decrease, and the temperatures rise to 548°C. Furthermore, a comparison between the metal flows in the welding chambers of P1 and P6 reveals distinct characteristics from those of P2-P5. The latter experiences higher mean stress, effective stress, and effective strain rate. Consequently, the quality of solid-state welding at the centre of the L-seam in the extruded profile is compromised, whereas the quality near the inner and outer surfaces remains relatively better. Additionally, these findings are in accordance with the study by Yu et al. [7].

<figure><img src="image_729.png" /><figcaption>Fig. 7 Examination of welding paths in terms of thermo-mechanical conditions within the welding chamber.</figcaption></figure>

Furthermore, the welding quality indices were calculated by applying three typical welding criteria reported in the literature (Q, K and J criteria), considering the thermo-mechanical history of the tracing points along the welding path. For the calculation of J-criterion, the value of $k_0$ was set to 1, following the investigation conducted by Yu et al. [7], where Q was set at 142 kJ/mol and R at 8.314 J/(K·mol).
effective stresses and effective strain rates initially increase and then decrease, and the temperatures rise to 548°C. Furthermore, a comparison between the metal flows in the welding chambers of P1 and P6 reveals distinct characteristics from those of P2-P5. The latter experiences higher mean stress, effective stress, and effective strain rate. Consequently, the quality of solid-state welding at the centre of the L-seam in the extruded profile is compromised, whereas the quality near the inner and outer surfaces remains relatively better. Additionally, these findings are in accordance with the study by Yu et al. [7].

<figure><img src="image_737.png" /><figcaption>Fig. 7 Examination of welding paths in terms of thermo-mechanical conditions within the welding chamber.</figcaption></figure>

Furthermore, the welding quality indices were calculated by applying three typical welding criteria reported in the literature (Q, K and J criteria), considering the thermo-mechanical history of the tracing points along the welding path. For the calculation of J-criterion, the value of $k_0$ was set to 1, following the investigation conducted by Yu et al. [7], where Q was set at 142 kJ/mol and R at 8.314 J/(K·mol).
<figure><img src="image_730.png" /><figcaption>Fig. 8 Computation of L-seam weld quality indices calculated using Q, K and J criteria.</figcaption></figure>

In Fig. 8, the welding quality indices for both Region 1 (P1-P6) and Region 2 (P7-P9) are presented. In Fig. 8a, the quality indices based on the Q-criterion reveal that the Q values of P1-P3 are generally higher than those of P4-P6, while the Q values of P7-P12 exhibit consistency. Conversely, when comparing Regions 1 and 2, the Q values of P1-P3 appear greater than those of P7-P9, potentially suggesting an overall superior welding quality in Region 1 compared to Region 2. Examining quality index calculations based on the K-criterion in Fig. 8b, the K-criterion mirrors the behaviour observed with the Q-criterion in Region 1. However, in Region 2, P4 and P5 exhibit smaller K values compared to the other points. Regarding the J-criterion (Fig. 8c), points near the inner and outer surfaces display greater J values than those in the centre. These findings align with the observations from Fig. 7. It can be asserted that the primary reason behind these observations is that the J-criterion takes into account several critical factors, such as mean stress, effective stress, effective strain (rate), and temperature. In contrast, the Q and K criteria only consider the effects of mean stress, effective stress, and material flow speed on the welding quality.

## Conclusions and Outlook

This paper examines the L-seam weld quality and mechanism within multi-chamber aluminium profiles to investigate how the thermo-mechanical history during extrusion influences the weld quality, employing both experimental, analytical and numerical approaches. The analysis of the thermo-mechanical history of the material within the welding chamber reveals a compromise in the quality of solid-state welding at the centre of the L-seam in the extruded profile. However, the quality near the inner and outer surfaces remains comparatively higher. Moreover, upon a comparative analysis of Q, K, and J criteria, it is observed that the J-criterion more accurately
<figure><img src="image_738.png" /><figcaption>Fig. 8 Computation of L-seam weld quality indices calculated using Q, K and J criteria.</figcaption></figure>

In Fig. 8, the welding quality indices for both Region 1 (P1-P6) and Region 2 (P7-P9) are presented. In Fig. 8a, the quality indices based on the Q-criterion reveal that the Q values of P1-P3 are generally higher than those of P4-P6, while the Q values of P7-P12 exhibit consistency. Conversely, when comparing Regions 1 and 2, the Q values of P1-P3 appear greater than those of P7-P9, potentially suggesting an overall superior welding quality in Region 1 compared to Region 2. Examining quality index calculations based on the K-criterion in Fig. 8b, the K-criterion mirrors the behaviour observed with the Q-criterion in Region 1. However, in Region 2, P4 and P5 exhibit smaller K values compared to the other points. Regarding the J-criterion (Fig. 8c), points near the inner and outer surfaces display greater J values than those in the centre. These findings align with the observations from Fig. 7. It can be asserted that the primary reason behind these observations is that the J-criterion takes into account several critical factors, such as mean stress, effective stress, effective strain (rate), and temperature. In contrast, the Q and K criteria only consider the effects of mean stress, effective stress, and material flow speed on the welding quality.

## Conclusions and Outlook

This paper examines the L-seam weld quality and mechanism within multi-chamber aluminium profiles to investigate how the thermo-mechanical history during extrusion influences the weld quality, employing both experimental, analytical and numerical approaches. The analysis of the thermo-mechanical history of the material within the welding chamber reveals a compromise in the quality of solid-state welding at the centre of the L-seam in the extruded profile. However, the quality near the inner and outer surfaces remains comparatively higher. Moreover, upon a comparative analysis of Q, K, and J criteria, it is observed that the J-criterion more accurately
captures the thermo-mechanical observations, particularly with greater values near the surfaces
than at the centre. However, further experimental validation is necessary to confirm these findings
and further assess the prediction capabilities of the quality indices.

The findings of this research gained insights into the understanding of the impact of 'local' thermo-mechanical conditions on L-seam weld quality. Future studies will concentrate on a more quantitative characterization and modelling of microstructural and mechanical properties in L-seam welds within multi-chamber profiles, facilitating to validation of the prediction capabilities of quality indices as well as enhancing the process control in aluminium extrusion with a particular focus on L-seam weld quality.

References

[1] W. Z. Misiolek, L. University, R. M. Kelly, and W. Co, “Extrusion of Aluminum Alloys,” 2005. https://doi.org/10.1361/asmhba0004015.

[2] H. Valberg, "Extrusion welding in aluminium extrusion," *International Journal of Materials and Product Technology*, vol. 17, no. 7, p. 497, 2002. https://doi.org/10.1504/IJMPT.2002.001317.

[3] S. W. Bai, G. Fang, and J. Zhou, “Analysis of the bonding strength and microstructure of AA6082 extrusion weld seams formed during physical simulation,” *J Mater Process Technol*, vol. 250, pp. 109–120, Dec. 2017. https://doi.org/10.1016/j.jmatprotec.2017.07.012.

[4] A. J. Den Bakker, R. J. Werkhoven, W. H. Sillekens, and L. Katgerman, “The origin of weld seam defects related to metal flow in the hot extrusion of aluminium alloys en AW-6060 and en AW-6082,” *J Mater Process Technol*, vol. 214, no. 11, pp. 2349–2358, 2014. https://doi.org/10.1016/j.jmatprotec.2014.05.001.

[5] A. Loukus, G. Subhash, and M. Imaninejad, “Mechanical properties and microstructural characterization of extrusion welds in AA6082-T4,” *J Mater Sci*, vol. 39, pp. 6561–6569, 2004.

[6] J. Yu, G. Zhao, and L. Chen, “Investigation of interface evolution, microstructure and mechanical properties of solid-state bonding seams in hot extrusion process of aluminum alloy profiles,” *J Mater Process Technol*, vol. 230, pp. 153–166, Apr. 2016. https://doi.org/10.1016/j.jmatprotec.2015.11.020.

[7] J. Yu, G. Zhao, and L. Chen, “Analysis of longitudinal weld seam defects and investigation of solid-state bonding criteria in porthole die extrusion process of aluminum alloy profiles,” *J Mater Process Technol*, vol. 237, pp. 31–47, Nov. 2016. https://doi.org/10.1016/j.jmatprotec.2016.05.024.

[8] B. REGGIANI and L. DONATI, “Comparison of experimental methods to evaluate seam welds quality in extruded profiles,” *Transactions of Nonferrous Metals Society of China*, vol. 30, no. 3, pp. 619–634, Mar. 2020. https://doi.org/10.1016/S1003-6326(20)65241-4.

[9] Akeret R., “Properties of pressure welds in extruded aluminum alloy sections,” *The Journal of the Institute of Metals*, vol. 10, no. 202, 1972.

[10] Bourqui B., Huber A., Moulin C., and Bunetti A., “Improved weld seam quality using 3D FEM simulation in correlation with practice,” in *Proceedings of First European Aluminum Association-Extruders Division*), Montichiari, 2002.

[11] M. Plata and J. Piwnik, “Theoretical and experimental analysis of seam weld formation in hot extrusion of aluminum alloys,” in *Proceedings of Seventh International Aluminum Extrusion Technology Seminar ET*, 2000, pp. 205–211.
captures the thermo-mechanical observations, particularly with greater values near the surfaces
than at the centre. However, further experimental validation is necessary to confirm these findings
and further assess the prediction capabilities of the quality indices.

The findings of this research gained insights into the understanding of the impact of 'local' thermo-mechanical conditions on L-seam weld quality. Future studies will concentrate on a more quantitative characterization and modelling of microstructural and mechanical properties in L-seam welds within multi-chamber profiles, facilitating to validation of the prediction capabilities of quality indices as well as enhancing the process control in aluminium extrusion with a particular focus on L-seam weld quality.

References

[1] W. Z. Misiolek, L. University, R. M. Kelly, and W. Co, “Extrusion of Aluminum Alloys,” 2005. https://doi.org/10.1361/asmhba0004015.

[2] H. Valberg, "Extrusion welding in aluminium extrusion," *International Journal of Materials and Product Technology*, vol. 17, no. 7, p. 497, 2002. https://doi.org/10.1504/IJMPT.2002.001317.

[3] S. W. Bai, G. Fang, and J. Zhou, “Analysis of the bonding strength and microstructure of AA6082 extrusion weld seams formed during physical simulation,” *J Mater Process Technol*, vol. 250, pp. 109–120, Dec. 2017. https://doi.org/10.1016/j.jmatprotec.2017.07.012.

[4] A. J. Den Bakker, R. J. Werkhoven, W. H. Sillekens, and L. Katgerman, “The origin of weld seam defects related to metal flow in the hot extrusion of aluminium alloys en AW-6060 and en AW-6082,” *J Mater Process Technol*, vol. 214, no. 11, pp. 2349–2358, 2014. https://doi.org/10.1016/j.jmatprotec.2014.05.001.

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[15] QForm UK, “QForm UK Extrusion 10,” Mar. 2022.

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# Behavior of volumetric core defects in friction extrusion of wire from Al-Cu alloy

RATH Lars¹,ᵃ*, SUHUDDIN Uceu F. H.¹,ᵇ and KLUSEMANN Benjamin¹,²,ᶜ

¹Helmholtz-Zentrum Hereon, Institute of Materials Mechanics, Solid State Materials Processing, Max-Planck Straße 1, 21502 Geesthacht, Germany

²Leuphana University Lüneburg, Institute of Product and Process Innovation, Universitätsallee 1, 21335 Lüneburg, Germany

ᵃlars.rath@hereon.de, ᵇuceu.suhuddin@hereon.de, ᶜbenjamin.klusemann@hereon.de

**Keywords:** Friction Extrusion, Solid-State Processing, Aluminum-Copper

**Abstract.** Friction extrusion describes the processing of metallic materials by inducing severe plastic deformation via frictional heating and shear strain. Rotational motion between die and feedstock is the key feature defining the potential of the process to generate consolidated extrudates with refined, homogenized microstructure. In this study, the effect of volumetric core defects on the material flow and properties of the extrudate is investigated, by processing from Al-Cu billets with a centric bore. Optical as well as scanning electron microscopy, X-ray microtomography and micro-hardness measurements are applied. Different material flow patterns and defect closure mechanisms are identified in correlation with the defect volume and the potential of controlling material flow via geometrical feedstock modification is discussed.

**Introduction**

First patented in 1993 by Thomas et al. [1], friction extrusion is characterized as the forming of metals under pressure and shear deformation. The basic principle involves the exertion of axial force by a die onto contained feedstock. Applied differential rotation between die and container results in frictional heating and shear introduction into the material, which consequently plasticizes and is being extruded, undergoing severe plastic deformation. The benefits of the process include improved energy efficiency compared to conventional extrusion and suitability for recycling purposes by using chips as feedstock, as systematically investigated by Baffari et al. [2]. The imposed material deformation furthermore can lead to significant microstructural changes with benefits for mechanical or chemical application, mainly based on the effective consolidation, refinement and homogenization of the material.

Friction extrusion of wires from machining chips of Al-Cu alloys has been investigated by Tang and Reynolds [3], observing an equiaxed and recrystallized microstructure with homogeneous ductility and hardness. Typical defects occurring on the wire surface were suggested to be in correlation with the extrusion temperature, setting the processing limits. Hosseini et al [4] performed closer investigation of the bonding between chips of AA 2025, identifying an influence of process parameters on void or defect behavior. Insufficient bonding between AlCu10 chips during friction extrusion was also reported in dependency from the extrusion ratio by Rath et al. [5] as a result of local pressure and thermal conditions in front of the die. Friction extrusion of square bars from scrap AA 1050 by Gelaw et al. [6], and the recycling of AA 7277 to produce wire via friction-stir extrusion by Behnagh et al. [7] further demonstrate the potential of this solid-state process. Due to largely different behavior of aluminum alloy groups in terms of e.g. oxide formation, warm strength, plasticizing behavior and phase formation related effects, transfer of results has to be considered carefully.

<figure><img src="image_731.png" /></figure>
# Behavior of volumetric core defects in friction extrusion of wire from Al-Cu alloy

RATH Lars¹,ª*, SUHUDDIN Uceu F. H.¹,ᵇ and KLUSEMANN Benjamin¹,²,ᶜ

¹Helmholtz-Zentrum Hereon, Institute of Materials Mechanics, Solid State Materials Processing, Max-Planck Straße 1, 21502 Geesthacht, Germany

²Leuphana University Lüneburg, Institute of Product and Process Innovation, Universitätsallee 1, 21335 Lüneburg, Germany

ᵃlars.rath@hereon.de, ᵇuceu.suhuddin@hereon.de, ᶜbenjamin.klusemann@hereon.de

**Keywords:** Friction Extrusion, Solid-State Processing, Aluminum-Copper

**Abstract.** Friction extrusion describes the processing of metallic materials by inducing severe plastic deformation via frictional heating and shear strain. Rotational motion between die and feedstock is the key feature defining the potential of the process to generate consolidated extrudates with refined, homogenized microstructure. In this study, the effect of volumetric core defects on the material flow and properties of the extrudate is investigated, by processing from Al-Cu billets with a centric bore. Optical as well as scanning electron microscopy, X-ray microtomography and micro-hardness measurements are applied. Different material flow patterns and defect closure mechanisms are identified in correlation with the defect volume and the potential of controlling material flow via geometrical feedstock modification is discussed.

**Introduction**

First patented in 1993 by Thomas et al. [1], friction extrusion is characterized as the forming of metals under pressure and shear deformation. The basic principle involves the exertion of axial force by a die onto contained feedstock. Applied differential rotation between die and container results in frictional heating and shear introduction into the material, which consequently plasticizes and is being extruded, undergoing severe plastic deformation. The benefits of the process include improved energy efficiency compared to conventional extrusion and suitability for recycling purposes by using chips as feedstock, as systematically investigated by Baffari et al. [2]. The imposed material deformation furthermore can lead to significant microstructural changes with benefits for mechanical or chemical application, mainly based on the effective consolidation, refinement and homogenization of the material.

Friction extrusion of wires from machining chips of Al-Cu alloys has been investigated by Tang and Reynolds [3], observing an equiaxed and recrystallized microstructure with homogeneous ductility and hardness. Typical defects occurring on the wire surface were suggested to be in correlation with the extrusion temperature, setting the processing limits. Hosseini et al [4] performed closer investigation of the bonding between chips of AA 2025, identifying an influence of process parameters on void or defect behavior. Insufficient bonding between AlCu10 chips during friction extrusion was also reported in dependency from the extrusion ratio by Rath et al. [5] as a result of local pressure and thermal conditions in front of the die. Friction extrusion of square bars from scrap AA 1050 by Gelaw et al. [6], and the recycling of AA 7277 to produce wire via friction-stir extrusion by Behnagh et al. [7] further demonstrate the potential of this solid-state process. Due to largely different behavior of aluminum alloy groups in terms of e.g. oxide formation, warm strength, plasticizing behavior and phase formation related effects, transfer of results has to be considered carefully.

<figure><img src="image_739.png" /></figure>
Approaches to build a general understanding of the underlying mechanism in friction extrusion have been made. Li et al. [8] applied a marker-insert-technique to visualize material flow, showing a severe influence of changing thermal and shear conditions over process time on the homogenization in the extrudate. Zhao et al. [9] employed digital image correlation to derive an experimentally validated model using viscous fluids. By correlating strain and texture with local conditions during friction extrusion, another advance was made by Li et al. [10] in 2016.

The identified objective for this study therefore is to gain further inside on the material flow and its interaction with volumetric defects as typically occurring in feedstock from chips. Further interest lies in investigating a potential technical solution to control this material flow and enhance defect closure during processing. In rotation-symmetrical wire or bar extrusions, the shear introduction to the core of the billet is severely hindered by the distance to the die face and generally deteriorating shearing conditions as investigated by Halak et al. [11]. The result is a core area with low grain refinement and, in case of chips reduced defect elimination, whereas the outer regions of the wire take full benefit from the high deformation introduced [12].

Typical conditions, which help to reduce core inhomogeneity, are often not applicable as they involve a high extrusion ratio, the use of fine powder feedstock [13] or necessitate strict control of the time and temperature variable shearing conditions in front of the die [11]. Consequently, this work investigates the deformation mechanisms and microstructural changes in the core region by means of analyzing introduced defects throughout processing.

<figure><img src="image_732.png" /><figcaption>Fig. 1: a) Vertical friction extrusion setup with rotating die, b) AA 2219 billet clamped to the backing and c) billets with 6 and 3 mm centric bore as well as solid feedstock.</figcaption></figure>

## Experimental Methods

Feedstock Material. A wrought aluminum-copper alloy, AA 2219, suitable for structural components due to its high fracture toughness and resistance to stress corrosion cracking is used in this study. Billets of 15 mm diameter with 15 mm length plus a 12 mm diameter clamping section, see Fig. 1c, are machined from AA 2219 bar material in solution heat-treated, cold worked and artificially aged T8511 temper. Some of the billets are modified by introducing centric bores of 3 and 6 mm. The depth of the bore amounts to 10 mm plus a 118° drill tip angle.
Approaches to build a general understanding of the underlying mechanism in friction extrusion have been made. Li et al. [8] applied a marker-insert-technique to visualize material flow, showing a severe influence of changing thermal and shear conditions over process time on the homogenization in the extrudate. Zhao et al. [9] employed digital image correlation to derive an experimentally validated model using viscous fluids. By correlating strain and texture with local conditions during friction extrusion, another advance was made by Li et al. [10] in 2016.

The identified objective for this study therefore is to gain further inside on the material flow and its interaction with volumetric defects as typically occurring in feedstock from chips. Further interest lies in investigating a potential technical solution to control this material flow and enhance defect closure during processing. In rotation-symmetrical wire or bar extrusions, the shear introduction to the core of the billet is severely hindered by the distance to the die face and generally deteriorating shearing conditions as investigated by Halak et al. [11]. The result is a core area with low grain refinement and, in case of chips reduced defect elimination, whereas the outer regions of the wire take full benefit from the high deformation introduced [12].

Typical conditions, which help to reduce core inhomogeneity, are often not applicable as they involve a high extrusion ratio, the use of fine powder feedstock [13] or necessitate strict control of the time and temperature variable shearing conditions in front of the die [11]. Consequently, this work investigates the deformation mechanisms and microstructural changes in the core region by means of analyzing introduced defects throughout processing.

<figure><img src="image_740.png" /><figcaption>Fig. 1: a) Vertical friction extrusion setup with rotating die, b) AA 2219 billet clamped to the backing and c) billets with 6 and 3 mm centric bore as well as solid feedstock.</figcaption></figure>

## Experimental Methods

Feedstock Material. A wrought aluminum-copper alloy, AA 2219, suitable for structural components due to its high fracture toughness and resistance to stress corrosion cracking is used in this study. Billets of 15 mm diameter with 15 mm length plus a 12 mm diameter clamping section, see Fig. 1c, are machined from AA 2219 bar material in solution heat-treated, cold worked and artificially aged T8511 temper. Some of the billets are modified by introducing centric bores of 3 and 6 mm. The depth of the bore amounts to 10 mm plus a 118° drill tip angle.
Friction Extrusion Processing. The friction extrusion of 6 mm wires is performed on a vertical spindle friction surfacing system, Henry Loitz Robotik. The billets are clamped to a steel backing, see Fig. 1b, in order to lock rotation and a hollow cylinder from gunmetal CC493K is clamped to the backing acting as a container for the feedstock material, while providing low friction in contact with the die. The rotating die plunging into the contained billet, see Fig. 1a, has a 120° opening angle with a featureless face, a 6 mm die orifice and a 4 mm bearing length. The die and the container setup are from 42CrMo4+QT steel. For the extrusions, the die is advanced in force control at 14 kN while rotating at 1000 rpm. No pre-heating step or ramping is applied and the extrusion is stopped by die retraction under rotation at a targeted 7 mm plunge depth. At least two processes for each billet geometry were performed under identical processing conditions.

Analysis. For assessment of processing behavior, the machine feedback is recorded: axial force and rotational speed as control as well as spindle torque and die displacement as response parameters. The extruded wires are scanned in X-ray microtomography, Y.Cougar SMT from Xylon International, to detect void volume and progression of the introduced core defect. Scan settings of 115 kV and 60 µA were found to yield optimal contrast for defect detection. Further investigation of the extrudates involves the metallographic preparation by grinding to grit size 4000 with SiC-foil and subsequent polishing with 1 µm diamond suspension. Electrolytic etching was applied with Barker's reagent for 90 s at 15 V. Optical microscopy is performed with a VHX-6000 digital microscopy from Keyence. Hereby, the transitions zone from feedstock to extrudate is of particular interest. For scanning electron microscopy a Quanta 650, FEI Company, field emission gun is used with focus on the immediate grain structure and topology of the defect closure area. Additionally, microhardness of cross-sections along the extrudate as well as over the processing zone are analyzed. Therefore, Vickers indentation maps (HV 0.2, 0.3 mm x 0.3 mm grid) are measured on a Durascan G5, Struers, after allowing for at least 60 days of natural aging.

## Results and Discussion

Processing Behavior. While both the billets with and without introduced bore are processed with same control parameters, the void volume or the bore effectively lowers the extrusion ratio (ER). Depending on the individual material flow behavior, the resulting ER, relevant for estimating local pressure and strain conditions, lies between 6.25 and the respectively lower values given in Table 1. This value is not necessarily constant over the process time, but gives a first explanation for the observed change in average plunge speed during the extrusion phase, Fig. 2, ranging from 0.4 mm/s for the solid billets to 0.5 mm/s for billets with 6 mm bore.

Another factor affecting the process response besides the change in feedstock volume lies in the change of extrusion resistance. This shows as a combined effect of the die angle and the removed core material, facilitating radial material flow towards the introduced void volume. However, when looking at the actual axial force as well as the torque feedback, the extrusion resistance is still depending on material flow, as shown by the inverse effect of bore diameter on axial force and torque for 3 and 6 mm bores, Table 1.

A different pressure distribution across the die face is a possible explanation, although the axial force feedback has to be considered carefully, since in force-control mode an offset to the set value (14 kN) indicates accelerations or vibrations exceeding the capability of the employed controller.
Friction Extrusion Processing. The friction extrusion of 6 mm wires is performed on a vertical spindle friction surfacing system, Henry Loitz Robotik. The billets are clamped to a steel backing, see Fig. 1b, in order to lock rotation and a hollow cylinder from gunmetal CC493K is clamped to the backing acting as a container for the feedstock material, while providing low friction in contact with the die. The rotating die plunging into the contained billet, see Fig. 1a, has a 120° opening angle with a featureless face, a 6 mm die orifice and a 4 mm bearing length. The die and the container setup are from 42CrMo4+QT steel. For the extrusions, the die is advanced in force control at 14 kN while rotating at 1000 rpm. No pre-heating step or ramping is applied and the extrusion is stopped by die retraction under rotation at a targeted 7 mm plunge depth. At least two processes for each billet geometry were performed under identical processing conditions.

Analysis. For assessment of processing behavior, the machine feedback is recorded: axial force and rotational speed as control as well as spindle torque and die displacement as response parameters. The extruded wires are scanned in X-ray microtomography, Y.Cougar SMT from Xylon International, to detect void volume and progression of the introduced core defect. Scan settings of 115 kV and 60 µA were found to yield optimal contrast for defect detection. Further investigation of the extrudates involves the metallographic preparation by grinding to grit size 4000 with SiC-foil and subsequent polishing with 1 µm diamond suspension. Electrolytic etching was applied with Barker's reagent for 90 s at 15 V. Optical microscopy is performed with a VHX-6000 digital microscopy from Keyence. Hereby, the transitions zone from feedstock to extrudate is of particular interest. For scanning electron microscopy a Quanta 650, FEI Company, field emission gun is used with focus on the immediate grain structure and topology of the defect closure area. Additionally, microhardness of cross-sections along the extrudate as well as over the processing zone are analyzed. Therefore, Vickers indentation maps (HV 0.2, 0.3 mm x 0.3 mm grid) are measured on a Durascan G5, Struers, after allowing for at least 60 days of natural aging.

## Results and Discussion

Processing Behavior. While both the billets with and without introduced bore are processed with same control parameters, the void volume or the bore effectively lowers the extrusion ratio (ER). Depending on the individual material flow behavior, the resulting ER, relevant for estimating local pressure and strain conditions, lies between 6.25 and the respectively lower values given in Table 1. This value is not necessarily constant over the process time, but gives a first explanation for the observed change in average plunge speed during the extrusion phase, Fig. 2, ranging from 0.4 mm/s for the solid billets to 0.5 mm/s for billets with 6 mm bore.

Another factor affecting the process response besides the change in feedstock volume lies in the change of extrusion resistance. This shows as a combined effect of the die angle and the removed core material, facilitating radial material flow towards the introduced void volume. However, when looking at the actual axial force as well as the torque feedback, the extrusion resistance is still depending on material flow, as shown by the inverse effect of bore diameter on axial force and torque for 3 and 6 mm bores, Table 1.

A different pressure distribution across the die face is a possible explanation, although the axial force feedback has to be considered carefully, since in force-control mode an offset to the set value (14 kN) indicates accelerations or vibrations exceeding the capability of the employed controller.
<table><caption>Table 1: Average process response parameters for each billet geometry during steady die plunge phase in force-controlled friction extrusion as well as the nominal extrusion ratio.</caption><thead><tr><th></th><th>Avg. axial force [kN]</th><th>Avg. torque [Nm]</th><th>Avg. plunge speed [mm/s]</th><th>Nominal extrusion ratio</th></tr></thead><tbody><tr><td>Solid billet</td><td>13.0</td><td>40.9</td><td>0.407</td><td>6.25</td></tr><tr><td>3 mm bore</td><td>13.1</td><td>36.8</td><td>0.440</td><td>6.00</td></tr><tr><td>6 mm bore</td><td>12.3</td><td>39.8</td><td>0.500</td><td>5.25</td></tr></tbody></table>

The die advance over plunge time presents a nearly constant progression, see Fig. 2, after the initial plunge. In this 1-1.5 s phase the 120° conical inflow region of the die is filled with material and near-constant pressure conditions are established. A slight acceleration over process time in the following extrusion phase is observed for all processes and indicates an overall rising processing temperature. All processes resulted in virtually solid wires with 5.8-5.9 mm outer diameter and a length of 30-40 mm.

<figure><img src="image_733.png" /><figcaption>Fig. 2: Die plunge depth over time for the extrusion phase in friction extrusions from solid billets as well as billets with 3 and 6 mm centric bore. Identical process control parameters applied.</figcaption></figure>

Core Defect Analysis. The X-ray microtomography revealed a significant difference in distance from the wire tip for which a volumetric defect could be detected. A direct comparison of 3D-scans from representative samples of all processed billet geometries is shown in Fig. 3a. Here the length difference of the extruded wires as a result of the altered feedstock volume is clearly visible, as well as the remainders of the introduced bores in the feedstock as well as the core defects present in the wire tips.

Upon closer inspection of the wire tips, Fig. 3b, a defect length of 10-15 mm and 3-4 mm for wires from billets with 3 and 6 mm bore, respectively, is measured in the central longitudinal plane of the wires. Due to a resolution limit of approximately 50 µm, the actual defect length is potentially longer. Removing more of the billet's core material not only facilitated radial material flow in combination with the die angle as confirmed by process response parameters, but also removed the material that typically undergoes lower shear deformation or simply elongation, as discussed by Halak et al. [03]. This material would be expected to exert higher resistance to deformation and its removal leaves feedstock with larger but in process less persistent defects. The removal of the feedstock material that typically shows low grain refinement during extrusion by dynamic recrystallization leaves a more or less thick-walled tube-like core structure with higher resistance to radial compression. This model could explain the higher persistence of the core defect in billets with initially lower defect volume.
<table><caption>Table 1: Average process response parameters for each billet geometry during steady die plunge phase in force-controlled friction extrusion as well as the nominal extrusion ratio.</caption><thead><tr><th></th><th>Avg. axial force [kN]</th><th>Avg. torque [Nm]</th><th>Avg. plunge speed [mm/s]</th><th>Nominal extrusion ratio</th></tr></thead><tbody><tr><td>Solid billet</td><td>13.0</td><td>40.9</td><td>0.407</td><td>6.25</td></tr><tr><td>3 mm bore</td><td>13.1</td><td>36.8</td><td>0.440</td><td>6.00</td></tr><tr><td>6 mm bore</td><td>12.3</td><td>39.8</td><td>0.500</td><td>5.25</td></tr></tbody></table>

The die advance over plunge time presents a nearly constant progression, see Fig. 2, after the initial plunge. In this 1-1.5 s phase the 120° conical inflow region of the die is filled with material and near-constant pressure conditions are established. A slight acceleration over process time in the following extrusion phase is observed for all processes and indicates an overall rising processing temperature. All processes resulted in virtually solid wires with 5.8-5.9 mm outer diameter and a length of 30-40 mm.

<figure><img src="image_741.png" /><figcaption>Fig. 2: Die plunge depth over time for the extrusion phase in friction extrusions from solid billets as well as billets with 3 and 6 mm centric bore. Identical process control parameters applied.</figcaption></figure>

Core Defect Analysis. The X-ray microtomography revealed a significant difference in distance from the wire tip for which a volumetric defect could be detected. A direct comparison of 3D-scans from representative samples of all processed billet geometries is shown in Fig. 3a. Here the length difference of the extruded wires as a result of the altered feedstock volume is clearly visible, as well as the remainders of the introduced bores in the feedstock as well as the core defects present in the wire tips.

Upon closer inspection of the wire tips, Fig. 3b, a defect length of 10-15 mm and 3-4 mm for wires from billets with 3 and 6 mm bore, respectively, is measured in the central longitudinal plane of the wires. Due to a resolution limit of approximately 50 µm, the actual defect length is potentially longer. Removing more of the billet's core material not only facilitated radial material flow in combination with the die angle as confirmed by process response parameters, but also removed the material that typically undergoes lower shear deformation or simply elongation, as discussed by Halak et al. [03]. This material would be expected to exert higher resistance to deformation and its removal leaves feedstock with larger but in process less persistent defects. The removal of the feedstock material that typically shows low grain refinement during extrusion by dynamic recrystallization leaves a more or less thick-walled tube-like core structure with higher resistance to radial compression. This model could explain the higher persistence of the core defect in billets with initially lower defect volume.
<figure><img src="image_734.png" /><figcaption>Fig. 3: X-ray microtomography for detection of void volume: a) 3D-scans of wires and residual feedstock friction extruded from different billet geometry and b) central longitudinal plane through wire tips showing closure of the introduced core defects.</figcaption></figure>

Microstructure Analysis. Metallographic analysis of the transition zone from feedstock to wire allowed for more information about the material flow to be obtained. All samples show a highly comparable material flow at the die-feedstock interface, where grain refinement is induced, shaping the outer shell of the extruded wire. The shear-affected zones reaching from the die face into the billets have a similar geometry.

<figure><img src="image_735.png" /><figcaption>Fig. 4: Optical microscopy: Etched cross-sections of the feedstock-wire transition zone for friction extrusion processes from a) solid billet, b) billets with 3 mm and c) 6 mm bore.</figcaption></figure>

In the billets with 3 mm bore, Fig. 3b, the material flow is closing just before significant microstructural changes are introduced, bypassing or going around the defect without large impact on the process zone. In contrast, in billets with 6 mm bore, Fig. 3c, a flow of refined material away from the die into the introduced defect is observed. The material flow pattern around the defect joins in this case closer to the die orifice in a region where high deformation and partial recrystallization is already present. This difference conforms to the observed force and torque response, Table 1, and shows in the core structure of the wires as well: While in the first case (3 mm bore) a macrograph similar to extrusion from a solid billet is obtained, in the latter case
<figure><img src="image_742.png" /><figcaption>Fig. 3: X-ray microtomography for detection of void volume: a) 3D-scans of wires and residual feedstock friction extruded from different billet geometry and b) central longitudinal plane through wire tips showing closure of the introduced core defects.</figcaption></figure>

Microstructure Analysis. Metallographic analysis of the transition zone from feedstock to wire allowed for more information about the material flow to be obtained. All samples show a highly comparable material flow at the die-feedstock interface, where grain refinement is induced, shaping the outer shell of the extruded wire. The shear-affected zones reaching from the die face into the billets have a similar geometry.

<figure><img src="image_743.png" /><figcaption>Fig. 4: Optical microscopy: Etched cross-sections of the feedstock-wire transition zone for friction extrusion processes from a) solid billet, b) billets with 3 mm and c) 6 mm bore.</figcaption></figure>

In the billets with 3 mm bore, Fig. 3b, the material flow is closing just before significant microstructural changes are introduced, bypassing or going around the defect without large impact on the process zone. In contrast, in billets with 6 mm bore, Fig. 3c, a flow of refined material away from the die into the introduced defect is observed. The material flow pattern around the defect joins in this case closer to the die orifice in a region where high deformation and partial recrystallization is already present. This difference conforms to the observed force and torque response, Table 1, and shows in the core structure of the wires as well: While in the first case (3 mm bore) a macrograph similar to extrusion from a solid billet is obtained, in the latter case
(6 mm bore) a split in grain orientation is observed in the wire core. Despite this, the grain elongation and torsion around the extrusion axis are comparable.

Scanning Electron Microscopy. The tip of the defect in wire extruded from a billet with 3 mm bore is shown in Fig. 5. The elongated grain structure observed before in the wire core, Fig. 4, is visible here by alignment of precipitates. The grains are radially compressed towards the defect without noticeable deflection from the extrusion axis. In direct proximity to the crack tip, intergranular cracks appear to occur. The defect closure mechanism in this case can be clearly differentiated from the backflow phenomena observed in billets with 6 mm bore, as only radial compression and simultaneous elongation are present, further explaining the longer presence of the defect. This gradually reduces in diameter with increasing processing temperature and decreasing material resistance. It should be noted, that the actual defect tip is difficult to analyze, due to mismatch in imaging plane and angle with the defect or wire axis. The possible continuous presence of a core defect in the wires extruded from 3 mm billets cannot be excluded.

<figure><img src="image_736.png" /><figcaption>Fig. 5: Backscattered electron images of the defect tip in wire extruded from billet with 3 mm bore. Precipitate orientation and intergranular defects in extrusion direction (left to right) are visible.</figcaption></figure>

Mechanical Properties. Another correlation with the change in material flow is seen in the hardness distribution in the extrudates. While for all samples the hardness over wire length slowly increases from around 80 HV at the tip to around 90 HV after 25 mm, for extrusions from billets with 6 mm bore, see Fig. 6c, additionally a pronounced hardness gradient across the wire diameter is observed. While compared to the core, the edges of the wires typically show an increase of up to 5 HV towards the end of the process, Fig. 6a and b, this difference amounts to more than 15 HV in the wires from billets with 6 mm bore.

<figure><img src="image_737.png" /><figcaption>Fig. 6: Vickers micro-hardness mapping of the cross-sections through the processing zone for friction extrusions from a) solid billet, b) billets with 3 mm and c) with 6 mm bore.</figcaption></figure>

The hardness distribution in the feedstock in front of the die shows overall good comparability for all billet geometries. Notable is the significantly lower hardness in the material flowing opposed to the extrusion direction in the extrusion from billets with 6 mm bore. The material filling up the drill tip cavity presents a hardness of 80 HV, comparable to the tip of the wire. Since strain and thermal history largely define the hardness for these samples exposed to otherwise identical conditions, this observation supports the hypothesis that the back-flowing material forms early on
(6 mm bore) a split in grain orientation is observed in the wire core. Despite this, the grain elongation and torsion around the extrusion axis are comparable.

Scanning Electron Microscopy. The tip of the defect in wire extruded from a billet with 3 mm bore is shown in Fig. 5. The elongated grain structure observed before in the wire core, Fig. 4, is visible here by alignment of precipitates. The grains are radially compressed towards the defect without noticeable deflection from the extrusion axis. In direct proximity to the crack tip, intergranular cracks appear to occur. The defect closure mechanism in this case can be clearly differentiated from the backflow phenomena observed in billets with 6 mm bore, as only radial compression and simultaneous elongation are present, further explaining the longer presence of the defect. This gradually reduces in diameter with increasing processing temperature and decreasing material resistance. It should be noted, that the actual defect tip is difficult to analyze, due to mismatch in imaging plane and angle with the defect or wire axis. The possible continuous presence of a core defect in the wires extruded from 3 mm billets cannot be excluded.

<figure><img src="image_744.png" /><figcaption>Fig. 5: Backscattered electron images of the defect tip in wire extruded from billet with 3 mm bore. Precipitate orientation and intergranular defects in extrusion direction (left to right) are visible.</figcaption></figure>

Mechanical Properties. Another correlation with the change in material flow is seen in the hardness distribution in the extrudates. While for all samples the hardness over wire length slowly increases from around 80 HV at the tip to around 90 HV after 25 mm, for extrusions from billets with 6 mm bore, see Fig. 6c, additionally a pronounced hardness gradient across the wire diameter is observed. While compared to the core, the edges of the wires typically show an increase of up to 5 HV towards the end of the process, Fig. 6a and b, this difference amounts to more than 15 HV in the wires from billets with 6 mm bore.

<figure><img src="image_745.png" /><figcaption>Fig. 6: Vickers micro-hardness mapping of the cross-sections through the processing zone for friction extrusions from a) solid billet, b) billets with 3 mm and c) with 6 mm bore.</figcaption></figure>

The hardness distribution in the feedstock in front of the die shows overall good comparability for all billet geometries. Notable is the significantly lower hardness in the material flowing opposed to the extrusion direction in the extrusion from billets with 6 mm bore. The material filling up the drill tip cavity presents a hardness of 80 HV, comparable to the tip of the wire. Since strain and thermal history largely define the hardness for these samples exposed to otherwise identical conditions, this observation supports the hypothesis that the back-flowing material forms early on
in the friction extrusion process and moves with the plunging die into the feedstock, enabling
defect free extrusion.

Conclusions

Solid wires from AA 2219 have been produced by friction extrusion. The behavior of defects through friction extrusion processing, introduced by centric bores in the feedstock, was analyzed with regard to interaction between shear introduction and material flow. Furthermore, efficiency of defect closure and contributing effects were discussed. The main findings summarize as follows:

* Removal of core material can have significant impact on material flow.

* Distinct defect closure mechanisms are observed for different dimensions of the introduced feedstock defect.

* Larger centric defects experienced faster closure and complete consolidation.

* Increased radial and especially backflow of feedstock material benefits defect closure.

* Potential approach for improving defect closure in low shear influenced process zones presented, e.g. for friction extrusion or wires from chips.

To further investigate the defect closure mechanism it is suggested to perform stop-action experiments, extend the defect size range as well as introducing off-center defects. A consequent utilization of the gained knowledge poses the extension to non-bulk feedstock e.g. chips for recycling purposes.

Funding

This project has received funding from the European Research Council (ERC) under the European Union's Horizon 2020 research and innovation programme (grant agreement No 101001567).

<figure><img src="image_738.png" /></figure>

<figure><img src="image_739.png" /></figure>

European Research Council

Established by the European Commission

Data Availability

The data related to this research is available online (https://doi.org/10.5281/zenodo.10456380).
in the friction extrusion process and moves with the plunging die into the feedstock, enabling
defect free extrusion.

Conclusions

Solid wires from AA 2219 have been produced by friction extrusion. The behavior of defects through friction extrusion processing, introduced by centric bores in the feedstock, was analyzed with regard to interaction between shear introduction and material flow. Furthermore, efficiency of defect closure and contributing effects were discussed. The main findings summarize as follows:

* Removal of core material can have significant impact on material flow.

* Distinct defect closure mechanisms are observed for different dimensions of the introduced feedstock defect.

* Larger centric defects experienced faster closure and complete consolidation.

* Increased radial and especially backflow of feedstock material benefits defect closure.

* Potential approach for improving defect closure in low shear influenced process zones presented, e.g. for friction extrusion or wires from chips.

To further investigate the defect closure mechanism it is suggested to perform stop-action experiments, extend the defect size range as well as introducing off-center defects. A consequent utilization of the gained knowledge poses the extension to non-bulk feedstock e.g. chips for recycling purposes.

Funding

This project has received funding from the European Research Council (ERC) under the European Union's Horizon 2020 research and innovation programme (grant agreement No 101001567).

<figure><img src="image_746.png" /></figure>

<figure><img src="image_747.png" /></figure>

European Research Council

Established by the European Commission

Data Availability

The data related to this research is available online (https://doi.org/10.5281/zenodo.10456380).
## References

[1] W.M. Thomas, E.D. Nicholas and S.B. Jones, U.S. Patent 5,262,123. (1993).

[2] D. Baffari, A.P. Reynolds, A. Masnata, L. Fratini and G. Ingarao, Friction stir extrusion to recycle aluminum alloys scraps: Energy efficiency characterization, J. Manuf. Process. 43 (2019) 63-69. https://doi.org/10.1016/j.jmapro.2019.03.049

[3] W. Tang and A.P. Reynolds, Production of wires via friction extrusion of aluminum alloy machining chips, J. Mater. Process. Technol. 210(15) (2010) 2231-2237. https://doi.org/10.1016/j.jmatprotec.2010.08.010

[4] A. Hosseini Tazehkandi, E. Azarsa, B. Davoodi and Y. Ardahani, Effect of process parameters on the physical properties of wires produced by friction extrusion method, Int. J. Adv. Eng. Technol. 3 (1) (2012) 592-597.

[5] L. Rath, U. Suhuddin and B. Klusemann, Comparison of friction extrusion processing from bulk and chips of aluminum-copper alloys, Key Engineering Materials 926 (2022) 471-480. https://doi.org/10.4028/p-vw04z5

[6] M. Gelaw, P.J. Ramulu, D. Hailu and T. Desta, Manufacturing and mechanical characterization of square bar made of aluminium scraps through friction stir back extrusion process, J. Eng. Des. Technol. (2018) 16:4. https://doi.org/10.1108/JEDT-02-2018-0030

[7] R.A. Behnagh, R. Mahdavinejad, A. Yavari, M. Abdollahi and M. Narvan, Production of wire from AA7277 aluminum chips via friction-stir extrusion (FSE), Metall. Mater. Trans. B 45 (2014) 1484-1489. https://doi.org/10.1007/s11663-014-0067-2

[8] X. Li, W. Tang and A.P. Reynolds, Visualization of material flow in friction extrusion, in H. Weiland, A.D. Rollett and W.A. Cassada (Eds.), ICAA13: 13th International Conference on Aluminum Alloys, TMS (The Minerals, Metals and Materials Society), Pittsburgh, 2012, pp. 1659-1664. https://doi.org/10.1007/978-3-319-48761-8_248

[9] X. Zhao, M.A. Sutton, H. Zhang, X. Deng, A.P. Reynolds, X. Ke and H.W. Schreier, Stereo image based motion measurements in fluids: Experimental validation and application in friction extrusion, Exp. Mech. 55 (2015) 177-200. https://doi.org/10.1007/s11340-014-9907-x

[10] X. Li, W. Tang, A.P. Reynolds, W.A. Tayon and C.A. Brice, Strain and Texture in friction extrusion of aluminum wire, J. Mater. Process. Technol. 229 (2016) 191-198. https://doi.org/10.1016/j.jmatprotec.2015.09.012

[11] R.M. Halak, L. Rath, U.F.H. Suhuddin, J.F. dos Santos and B. Klusemann, Changes in processing characteristics and microstructural evolution during friction extrusion of aluminum, Int. J. Mater. Form. (2022) 15:24. https://doi.org/10.1007/s12289-022-01670-y

[12] D. Baffari, A.P. Reynolds, X. Li and L. Fratini, Influence of processing parameters and initial temper on Friction Extrusion of 2050 aluminum alloy, J. Manuf. Process. 28 (2017) 319-325. https://doi.org/10.1016/j.jmapro.2017.06.013

[13] S. Whalen, M. Olszta, C. Roach, J. Darsell, D. Graff, Md. Reza-E-Rabby, T. Roosendaal, W. Daye, T. Pelletiers, S. Mathaudhu and N. Overman, High ductility aluminum alloy made from powder by friction extrusion, Materialia 6 (2019) 100260. https://doi.org/10.1016/j.mtla.2019.100260
## References

[1] W.M. Thomas, E.D. Nicholas and S.B. Jones, U.S. Patent 5,262,123. (1993).

[2] D. Baffari, A.P. Reynolds, A. Masnata, L. Fratini and G. Ingarao, Friction stir extrusion to recycle aluminum alloys scraps: Energy efficiency characterization, J. Manuf. Process. 43 (2019) 63-69. https://doi.org/10.1016/j.jmapro.2019.03.049

[3] W. Tang and A.P. Reynolds, Production of wires via friction extrusion of aluminum alloy machining chips, J. Mater. Process. Technol. 210(15) (2010) 2231-2237. https://doi.org/10.1016/j.jmatprotec.2010.08.010

[4] A. Hosseini Tazehkandi, E. Azarsa, B. Davoodi and Y. Ardahani, Effect of process parameters on the physical properties of wires produced by friction extrusion method, Int. J. Adv. Eng. Technol. 3 (1) (2012) 592-597.

[5] L. Rath, U. Suhuddin and B. Klusemann, Comparison of friction extrusion processing from bulk and chips of aluminum-copper alloys, Key Engineering Materials 926 (2022) 471-480. https://doi.org/10.4028/p-vw04z5

[6] M. Gelaw, P.J. Ramulu, D. Hailu and T. Desta, Manufacturing and mechanical characterization of square bar made of aluminium scraps through friction stir back extrusion process, J. Eng. Des. Technol. (2018) 16:4. https://doi.org/10.1108/JEDT-02-2018-0030

[7] R.A. Behnagh, R. Mahdavinejad, A. Yavari, M. Abdollahi and M. Narvan, Production of wire from AA7277 aluminum chips via friction-stir extrusion (FSE), Metall. Mater. Trans. B 45 (2014) 1484-1489. https://doi.org/10.1007/s11663-014-0067-2

[8] X. Li, W. Tang and A.P. Reynolds, Visualization of material flow in friction extrusion, in H. Weiland, A.D. Rollett and W.A. Cassada (Eds.), ICAA13: 13th International Conference on Aluminum Alloys, TMS (The Minerals, Metals and Materials Society), Pittsburgh, 2012, pp. 1659-1664. https://doi.org/10.1007/978-3-319-48761-8_248

[9] X. Zhao, M.A. Sutton, H. Zhang, X. Deng, A.P. Reynolds, X. Ke and H.W. Schreier, Stereo image based motion measurements in fluids: Experimental validation and application in friction extrusion, Exp. Mech. 55 (2015) 177-200. https://doi.org/10.1007/s11340-014-9907-x

[10] X. Li, W. Tang, A.P. Reynolds, W.A. Tayon and C.A. Brice, Strain and Texture in friction extrusion of aluminum wire, J. Mater. Process. Technol. 229 (2016) 191-198. https://doi.org/10.1016/j.jmatprotec.2015.09.012

[11] R.M. Halak, L. Rath, U.F.H. Suhuddin, J.F. dos Santos and B. Klusemann, Changes in processing characteristics and microstructural evolution during friction extrusion of aluminum, Int. J. Mater. Form. (2022) 15:24. https://doi.org/10.1007/s12289-022-01670-y

[12] D. Baffari, A.P. Reynolds, X. Li and L. Fratini, Influence of processing parameters and initial temper on Friction Extrusion of 2050 aluminum alloy, J. Manuf. Process. 28 (2017) 319-325. https://doi.org/10.1016/j.jmapro.2017.06.013

[13] S. Whalen, M. Olszta, C. Roach, J. Darsell, D. Graff, Md. Reza-E-Rabby, T. Roosendaal, W. Daye, T. Pelletiers, S. Mathaudhu and N. Overman, High ductility aluminum alloy made from powder by friction extrusion, Materialia 6 (2019) 100260. https://doi.org/10.1016/j.mtla.2019.100260
# Numerical simulation of the extrusion process with different FEM code approaches: analysis of thermal field, profile speed, defects evolution, and microstructure of hollow tubes

PELACCIA Riccardo<fnref n="1" />,ª*, NEGOZIO Marco<fnref n="2" />,ᵇ, DI DONATO Sara<fnref n="3" />,ᶜ, REGGIANI Barbara <fnref n="1" />,<fnref n="4" />,ᵈ, DONATI Lorenzo<fnref n="3" />,ᵉ

**Keywords:** Extrusion, Process Monitoring, Defect Analysis, Microstructure, FEM Validation

**Abstract.** Over the years, the utilization of FEM numerical simulation has become a valuable tool for both die makers and extruders. To assess the evolution of FEM codes in terms of problem-solving potentiality, accuracy, as well as reduction of computational time, the 'ICEB - International Conference on Extrusion and Benchmark' conference series was established. The participants, who are unaware of the experimental outputs in advance, share with the organizers their numerical results to be compared with the actual experimental data. Analysis and discussion of the results during the conference allows users to check if their simulation settings are adequate to replicate the problems, and software houses to verify the sensitivity of their solving methods. For the ICEB 2023, hollow tube-shaped profiles in AA6082 were extruded by varying the nitrogen flow rate and the quenching conditions. The experimental-numerical comparison was assessed in terms of press load, profile speeds, profile and die temperatures, defects evolution, nitrogen cooling efficiency, and the profiles' microstructure.

## Introduction

The 'ICEB - International Conference on Extrusion and Benchmark' conference series have been organized since 2007, not only to test the potentiality of the numerical tools, but also to propose interesting challenges and to improve the community's knowledge in the field of the extrusion process [1]. Over the years, the involvement of experts and not from different areas and backgrounds (FEM users, extruders, die makers, researchers, software houses...) has helped to create a specific conference on the extrusion process where discuss captivating challenges, understand the needs of the industrial framework, and verify the progress of FEM simulations for the process optimization. Notably, the organizers prepare and carry out the experimental campaign in collaboration with selected extruders and die makers: the die and profile geometry are selected to evidence problematics and challenges in terms of profile quality, die resistances, and defects as well as in terms of reliability of the numerical models. All necessary input data are shared with the participants, who have to simulate the case study by using the software they prefer, but without knowing the experimental results.

<figure><img src="image_740.png" /></figure>

<footnote n="1">¹DISMI Department of Sciences and Methods for Engineering, University of Modena and Reggio Emilia, Via Amendola 2, 42122, Reggio Emilia, Italy</footnote>

<footnote n="2">²Department of Engineering and Architecture, University of Parma, Parco Area delle Scienze, 181/A, 43124 Parma, Italy</footnote>

<footnote n="3">³DIN Department of Industrial Engineering, University of Bologna, Viale Risorgimento 2, 40136, Bologna, Italy</footnote>

<footnote n="4">⁴INTERMECH, University of Modena and Reggio Emilia, Italy</footnote>

<footnote>ªriccardo.pelaccia@unimore.it, ᵇmarco.negozio@unipr.it, ᶜsara.didonato2@unibo.it,</footnote>

<footnote>ᵈbarbara.reggiani@unimore.it, ᶜl.donati@unibo.it</footnote>
# Numerical simulation of the extrusion process with different FEM code approaches: analysis of thermal field, profile speed, defects evolution, and microstructure of hollow tubes

PELACCIA Riccardo<fnref n="1" />,ª*, NEGOZIO Marco<fnref n="2" />,ᵇ, DI DONATO Sara<fnref n="3" />,ᶜ, REGGIANI Barbara <fnref n="1" />,<fnref n="4" />,ᵈ, DONATI Lorenzo<fnref n="3" />,ᵉ

**Keywords:** Extrusion, Process Monitoring, Defect Analysis, Microstructure, FEM Validation

**Abstract.** Over the years, the utilization of FEM numerical simulation has become a valuable tool for both die makers and extruders. To assess the evolution of FEM codes in terms of problem-solving potentiality, accuracy, as well as reduction of computational time, the 'ICEB - International Conference on Extrusion and Benchmark' conference series was established. The participants, who are unaware of the experimental outputs in advance, share with the organizers their numerical results to be compared with the actual experimental data. Analysis and discussion of the results during the conference allows users to check if their simulation settings are adequate to replicate the problems, and software houses to verify the sensitivity of their solving methods. For the ICEB 2023, hollow tube-shaped profiles in AA6082 were extruded by varying the nitrogen flow rate and the quenching conditions. The experimental-numerical comparison was assessed in terms of press load, profile speeds, profile and die temperatures, defects evolution, nitrogen cooling efficiency, and the profiles' microstructure.

## Introduction

The 'ICEB - International Conference on Extrusion and Benchmark' conference series have been organized since 2007, not only to test the potentiality of the numerical tools, but also to propose interesting challenges and to improve the community's knowledge in the field of the extrusion process [1]. Over the years, the involvement of experts and not from different areas and backgrounds (FEM users, extruders, die makers, researchers, software houses...) has helped to create a specific conference on the extrusion process where discuss captivating challenges, understand the needs of the industrial framework, and verify the progress of FEM simulations for the process optimization. Notably, the organizers prepare and carry out the experimental campaign in collaboration with selected extruders and die makers: the die and profile geometry are selected to evidence problematics and challenges in terms of profile quality, die resistances, and defects as well as in terms of reliability of the numerical models. All necessary input data are shared with the participants, who have to simulate the case study by using the software they prefer, but without knowing the experimental results.

<figure><img src="image_748.png" /></figure>

<footnote n="1">¹DISMI Department of Sciences and Methods for Engineering, University of Modena and Reggio Emilia, Via Amendola 2, 42122, Reggio Emilia, Italy</footnote>

<footnote n="2">²Department of Engineering and Architecture, University of Parma, Parco Area delle Scienze, 181/A, 43124 Parma, Italy</footnote>

<footnote n="3">³DIN Department of Industrial Engineering, University of Bologna, Viale Risorgimento 2, 40136, Bologna, Italy</footnote>

<footnote n="4">⁴INTERMECH, University of Modena and Reggio Emilia, Italy</footnote>

<footnote>ªriccardo.pelaccia@unimore.it, ᵇmarco.negozio@unipr.it, ᶜsara.didonato2@unibo.it,</footnote>

<footnote>ᵈbarbara.reggiani@unimore.it, ᶜl.donati@unibo.it</footnote>
During the conference, the organizers compare the numerical results of the participants with the experimental ones, showing the achievements and the limits/challenges of the state of art [1-2]. In this context, the present work summarized the main outcomes obtained during the Extrusion Benchmark 2023, discussing the numerical results of the participants, and comparing them with the experimental outputs. Within the first six editions, in order to provide an even point of comparison between experiments and simulations, experimental trials were focused on critical aspects of the process: in 2007 different pocket design and profile thickness were tested finding that the slowing effect of the conical pocket can be more determining respect the profile thickness in terms of profile exiting velocities [3]; in 2009 two U-shape profiles were produced using different supporting legs to assess the die deformations induced by the design [4]; in 2011 different strategies for port-hole balancing were used for hollow profiles extrusion, evaluating the die deflections and the seam weld quality for the scientific case study [5], while a remarkable outcome of the industrial one was the decrease in profile temperature of 20°C using the die nitrogen cooling [6]; in 2013 the investigations were aimed at predicting the effects of mandrel deflection [7]; in 2015 the effect of bearing length and shape (straight, chocked or relieved) was examined [8-9]; while in 2017 the die design allowed mandrels shift during the process in order to verify profile distortions respect to nominal dimensions in addition to position of seam welds, microstructure size and charge welds evolution [10].

Following the same trend, for the Extrusion Benchmark 2023 a porthole die for hollow profile extrusion was designed with three tube-shaped openings (Fig. 1). The material flow within the tooling set was different for the 3 profiles (Fig. 2) with the aim to generate high velocity deviations at the exit. The experimental campaign involved the extrusion of fifteen AA6082 billets varying the nitrogen flow rate (0% or 100% valve opened) and the quenching conditions (air or water quenching) intending to evaluate the trend of the thermal field in both uncooled and cooled conditions as well as to assess the impact of quenching on the grain size evolution. The experimental outputs collected to provide a reliable comparison of FEM results were: press load, profile speeds, profile temperatures, die temperatures in eleven different locations, seam welds quality, charge welds and skin defect evolution, nitrogen cooling efficiency, and the microstructure size for both air and water quenched profiles.

## Experimental setting

Fig. 3 schematizes the experimental trials: the first five billets were extruded without the nitrogen cooling and using only the air quenching; the subsequent four billets using the nitrogen cooling with the valve fully opened; from the tenth to the twelfth extrusion the water quenching was enabled by keeping the nitrogen cooling system active; the last three extrusion were performed maintaining only the water quenching active. The extrusion speed was set constant at a value of 4.4 mm/s as suggested by the extruder (Metra SpA, plant in Rodengo Saiano, Brescia, IT). Fig. 3 also reports indications about the profile selected for the defects analysis (charge welds and billet skin) and for the microstructure assessment. The tooling set and the profile geometries were designed in collaboration with Almax Mori, Rovereto, TN. Three identical tube-shaped profiles (external diameter of 40 mm, thickness of 4 mm) were generated from the die set (Fig. 1), however, the three openings had differences (Fig. 2) in the ports' geometry (i.e. cylindrical or conical ports) and the bearings' surface (bearing lengths and radii) to induce at the exit high velocity deviations and consequently different strains, strain rates and temperatures in the tubes. The cooling channel was realized on the backer surface in contact with the die, and the cooling path, that surrounds the profile exits, is shown in Fig. 1c. Fig. 4 reports the location of the eleven thermocouples used to monitor the thermal field of the tooling set, while a contactless pyrometer (1500 mm far from the bolster exit face) was used to register the temperature of profile 1. Table 1 summarizes the main input data shared with the participants, while all additional information (e.g. geometry CAD files) can be asked to benchmark organizers (iceb.din@unibo.it).
During the conference, the organizers compare the numerical results of the participants with the experimental ones, showing the achievements and the limits/challenges of the state of art [1-2]. In this context, the present work summarized the main outcomes obtained during the Extrusion Benchmark 2023, discussing the numerical results of the participants, and comparing them with the experimental outputs. Within the first six editions, in order to provide an even point of comparison between experiments and simulations, experimental trials were focused on critical aspects of the process: in 2007 different pocket design and profile thickness were tested finding that the slowing effect of the conical pocket can be more determining respect the profile thickness in terms of profile exiting velocities [3]; in 2009 two U-shape profiles were produced using different supporting legs to assess the die deformations induced by the design [4]; in 2011 different strategies for port-hole balancing were used for hollow profiles extrusion, evaluating the die deflections and the seam weld quality for the scientific case study [5], while a remarkable outcome of the industrial one was the decrease in profile temperature of 20°C using the die nitrogen cooling [6]; in 2013 the investigations were aimed at predicting the effects of mandrel deflection [7]; in 2015 the effect of bearing length and shape (straight, chocked or relieved) was examined [8-9]; while in 2017 the die design allowed mandrels shift during the process in order to verify profile distortions respect to nominal dimensions in addition to position of seam welds, microstructure size and charge welds evolution [10].

Following the same trend, for the Extrusion Benchmark 2023 a porthole die for hollow profile extrusion was designed with three tube-shaped openings (Fig. 1). The material flow within the tooling set was different for the 3 profiles (Fig. 2) with the aim to generate high velocity deviations at the exit. The experimental campaign involved the extrusion of fifteen AA6082 billets varying the nitrogen flow rate (0% or 100% valve opened) and the quenching conditions (air or water quenching) intending to evaluate the trend of the thermal field in both uncooled and cooled conditions as well as to assess the impact of quenching on the grain size evolution. The experimental outputs collected to provide a reliable comparison of FEM results were: press load, profile speeds, profile temperatures, die temperatures in eleven different locations, seam welds quality, charge welds and skin defect evolution, nitrogen cooling efficiency, and the microstructure size for both air and water quenched profiles.

## Experimental setting

Fig. 3 schematizes the experimental trials: the first five billets were extruded without the nitrogen cooling and using only the air quenching; the subsequent four billets using the nitrogen cooling with the valve fully opened; from the tenth to the twelfth extrusion the water quenching was enabled by keeping the nitrogen cooling system active; the last three extrusion were performed maintaining only the water quenching active. The extrusion speed was set constant at a value of 4.4 mm/s as suggested by the extruder (Metra SpA, plant in Rodengo Saiano, Brescia, IT). Fig. 3 also reports indications about the profile selected for the defects analysis (charge welds and billet skin) and for the microstructure assessment. The tooling set and the profile geometries were designed in collaboration with Almax Mori, Rovereto, TN. Three identical tube-shaped profiles (external diameter of 40 mm, thickness of 4 mm) were generated from the die set (Fig. 1), however, the three openings had differences (Fig. 2) in the ports' geometry (i.e. cylindrical or conical ports) and the bearings' surface (bearing lengths and radii) to induce at the exit high velocity deviations and consequently different strains, strain rates and temperatures in the tubes. The cooling channel was realized on the backer surface in contact with the die, and the cooling path, that surrounds the profile exits, is shown in Fig. 1c. Fig. 4 reports the location of the eleven thermocouples used to monitor the thermal field of the tooling set, while a contactless pyrometer (1500 mm far from the bolster exit face) was used to register the temperature of profile 1. Table 1 summarizes the main input data shared with the participants, while all additional information (e.g. geometry CAD files) can be asked to benchmark organizers (iceb.din@unibo.it).
<figure><img src="image_741.png" /><figcaption>Fig. 1 ICEB 2023 case study: a) Rendering of the tools set-up, b) Tooling set composition, c) Cooling path.</figcaption></figure>

<figure><img src="image_742.png" /><figcaption>Fig. 2 Details of ports and bearings geometry for the three openings.</figcaption></figure>

<figure><img src="image_743.png" /><figcaption>Fig. 3 Selected Design of Experiment (DOE) for Extrusion Benchmark 2023.</figcaption></figure>

The requests for participants, which will be partly analyzed and discussed in the subsequent section, are reported:

Mandatory (on billet 4):

* Predicted press load; Profiles exit speeds; Profiles exit temperatures; Die temperatures in the eleven locations; Position of seam welds within the profiles.

Optional 1:
<figure><img src="image_749.png" /><figcaption>Fig. 1 ICEB 2023 case study: a) Rendering of the tools set-up, b) Tooling set composition, c) Cooling path.</figcaption></figure>

<figure><img src="image_750.png" /><figcaption>Fig. 2 Details of ports and bearings geometry for the three openings.</figcaption></figure>

<figure><img src="image_751.png" /><figcaption>Fig. 3 Selected Design of Experiment (DOE) for Extrusion Benchmark 2023.</figcaption></figure>

The requests for participants, which will be partly analyzed and discussed in the subsequent section, are reported:

Mandatory (on billet 4):

* Predicted press load; Profiles exit speeds; Profiles exit temperatures; Die temperatures in the eleven locations; Position of seam welds within the profiles.

Optional 1:
* Charge welds and skin defect evolution (billets 3-4 transition); Grain size with air quenching (billet 4 or 5); Grain size with water quenching (billet 14); Quality of the seam welds (billet 4).

Optional 2 (billet 11 with nitrogen cooling):

* Profiles exit temperature; Die temperatures in the eleven locations.

<figure><img src="image_744.png" /><figcaption>Fig. 4 Thermocouples' location within the tooling set.</figcaption></figure>

<table><caption>Table 1 Process parameters and input data for participants</caption><thead><tr><td>Process Parameter</td><td>Value</td></tr></thead><tbody><tr><td>Profile alloy</td><td>AA6082 flow stress data in [11]</td></tr><tr><td>Billet/Container diameter</td><td>279 / 286 [mm]</td></tr><tr><td>Billet length</td><td>680 [mm]</td></tr><tr><td>Billet rest</td><td>30 [mm]</td></tr><tr><td>Skin layer thickness</td><td>0.3 [mm]</td></tr><tr><td>Ram speed -all tested conditions-</td><td>4.4 [mm/s]</td></tr><tr><td>Max press load</td><td>4000 [ton] / 40 [MN]</td></tr><tr><td>Ram/Container Temperature</td><td>350 / (390-425) (ram-die) [°C]</td></tr><tr><td>Pre-heating Billet/Die Temperature</td><td>440 (no taper) / 480 [°C]</td></tr><tr><td>Pyrometer distance from bolster surface</td><td>1500 [mm]</td></tr><tr><td>Water/air spray quenching box distance from bolster surface</td><td>2500 [mm]</td></tr><tr><td>Water/air spray quenching box length</td><td>4800 [mm]</td></tr><tr><td>Decrease of profile temperature (air/water quenching)</td><td>1 / 100 [°C/sec]</td></tr><tr><td>Initial average grain size</td><td>115 [µm]</td></tr><tr><td>Nitrogen Inlet Temperature</td><td>-196 [°C]</td></tr><tr><td>Nitrogen Inlet Pressure</td><td>3 [bars]</td></tr><tr><td>Nitrogen Inlet Tube Diameter (internal)</td><td>5,5 [mm]</td></tr></tbody></table>

## Participants results: Experimental-numerical comparison

Twelve companies attended the Extrusion Benchmark 2023, but only four of them as active participants with their simulations: Hydro Innovation and Technology as research group, extruder and user of Qform UK software, Micas Simulations Limited as software house of Qform UK; Altair as software house of HyperXtrude; Almax Mori as die maker and user of HyperXtrude software. During the presentation and the discussion of the results, the four participants are
* Charge welds and skin defect evolution (billets 3-4 transition); Grain size with air quenching (billet 4 or 5); Grain size with water quenching (billet 14); Quality of the seam welds (billet 4).

Optional 2 (billet 11 with nitrogen cooling):

* Profiles exit temperature; Die temperatures in the eleven locations.

<figure><img src="image_752.png" /><figcaption>Fig. 4 Thermocouples' location within the tooling set.</figcaption></figure>

<table><caption>Table 1 Process parameters and input data for participants</caption><thead><tr><td>Process Parameter</td><td>Value</td></tr></thead><tbody><tr><td>Profile alloy</td><td>AA6082 flow stress data in [11]</td></tr><tr><td>Billet/Container diameter</td><td>279 / 286 [mm]</td></tr><tr><td>Billet length</td><td>680 [mm]</td></tr><tr><td>Billet rest</td><td>30 [mm]</td></tr><tr><td>Skin layer thickness</td><td>0.3 [mm]</td></tr><tr><td>Ram speed -all tested conditions-</td><td>4.4 [mm/s]</td></tr><tr><td>Max press load</td><td>4000 [ton] / 40 [MN]</td></tr><tr><td>Ram/Container Temperature</td><td>350 / (390-425) (ram-die) [°C]</td></tr><tr><td>Pre-heating Billet/Die Temperature</td><td>440 (no taper) / 480 [°C]</td></tr><tr><td>Pyrometer distance from bolster surface</td><td>1500 [mm]</td></tr><tr><td>Water/air spray quenching box distance from bolster surface</td><td>2500 [mm]</td></tr><tr><td>Water/air spray quenching box length</td><td>4800 [mm]</td></tr><tr><td>Decrease of profile temperature (air/water quenching)</td><td>1 / 100 [°C/sec]</td></tr><tr><td>Initial average grain size</td><td>115 [µm]</td></tr><tr><td>Nitrogen Inlet Temperature</td><td>-196 [°C]</td></tr><tr><td>Nitrogen Inlet Pressure</td><td>3 [bars]</td></tr><tr><td>Nitrogen Inlet Tube Diameter (internal)</td><td>5,5 [mm]</td></tr></tbody></table>

## Participants results: Experimental-numerical comparison

Twelve companies attended the Extrusion Benchmark 2023, but only four of them as active participants with their simulations: Hydro Innovation and Technology as research group, extruder and user of Qform UK software, Micas Simulations Limited as software house of Qform UK; Altair as software house of HyperXtrude; Almax Mori as die maker and user of HyperXtrude software. During the presentation and the discussion of the results, the four participants are
indicated randomly with the code from P1 to P4 as requested by themselves. However, the code always identifies the same participant.

Predicted Press Load. The experimental extrusion load was compared with those predicted by the participants for the fourth extrusion (steady-state uncooled condition). High accuracy of the numerical predictions was obtained with an average error below 10% with respect to the experimental maximum value of 36.1 MN.

Profile Exit Speeds. During the first run, before the use of the puller, the velocity of profile 2 was qualitatively 50% higher than the other two profiles. Then, despite the application of a tensile force of 3500 N, profile 2 continued to flow through the puller at a higher speed. In Table 2 the experimental velocity of profiles 1 and 3 was the same as the puller (avg. speed of 195 mm/s), while for profile 2, the average speed of 225 mm/s (15% higher than the puller one) was obtained by calculating the difference in terms of extrudates length at the end of the extrusion. In this case, the bearings optimization could be not enough to easily compensate for the speed differences generated by a conical port design (Fig. 2). Indeed, the experimental evidence demonstrated that the speed disparity between profile 3 and profile 1 (both extruded with conical ports) was relatively minimal and negligible without and with the puller, respectively, despite variations in bearing configuration (choke for profile 3, 1 mm higher bearing length for profile 1). Furthermore, despite profile 2 having a 1 mm greater bearing length than profile 1, the former exhibited a significant faster flow due to the presence of cylindrical ports. The numerical results were aligned with the experimental evidence of the first run (no one simulated the presence of the puller), where profile 1 was faster than profile 3, remaining profile 2 the fastest one.

<table><caption>Table 2 Profile exit speeds: Experimental-Numerical comparison (billet 4)</caption><thead><tr><th>Profile exit speed / Billet 4</th><th>Profile 1 [mm/s]</th><th>Profile 2 [mm/s]</th><th>Profile 3 [mm/s]</th></tr></thead><tbody><tr><td>Experimental</td><td>195</td><td>225</td><td>195</td></tr><tr><td>P1</td><td>213 (Num Err. +9.2%)</td><td>254 (Num Err. +12.8%)</td><td>141 (Num Err. -27.7%)</td></tr><tr><td>P2</td><td>183 (Num Err. -6.2%)</td><td>251 (Num Err. +11.5%)</td><td>176 (Num Err. -9.7%)</td></tr><tr><td>P3</td><td>183 (Num Err. -6.2%)</td><td>261 (Num Err. +16%)</td><td>161 (Num Err. -17.4%)</td></tr><tr><td>P4</td><td>184 (Num Err. -5.7%)</td><td>245 (Num Err. +8.9%)</td><td>160 (Num Err. -17.9%)</td></tr></tbody></table>

Profiles exit temperatures. The contactless pyrometer only registered the exit profile temperature of profile 1 that, during the fourth and the fifth runs, had stabilized at around 560 °C. In this case, the numerical results of the participants differed from each other due to the selection of dissimilar heat exchange with air. Indeed, despite all participants obtaining profile temperatures in the range of 550-570 °C near the bearings, the differences were more marked at the location of the pyrometer also among users of the same software: 570°C, 536°C, 495°C, and 520°C was predicted by P1, P2, P3, and P4, respectively. Therefore, P1 better predicted the low heat exchange with air (num. err. +1.8%), while other participants overestimated its cooling efficiency (max err. of -12.3% for P3). The exit profile temperature plays an important role in the prediction of the microstructure [12-14], therefore also the proper calibration of the thermal boundary conditions “downstream of the process” is not to be underestimated.

Die temperatures in the eleven locations. Fig. 5 reports the thermocouples' temperature recorded during the steady state uncooled condition (billet 4). The thermocouples T4-T6 nearby the profiles' exit recorded a maximum value of about 444 °C. T1 and T2 nearby the inlet channel marked 361°C and 406°C, respectively. T9 (bearings exit 2) indicated the value of 490 °C; T8 (bearings exit 3) only 445 °C, while T7 (bearings exit 1) recorded an intermediate value of 473 °C. As expected, the highest temperature of 535 °C was observed in T10 (mandrel) and the lowest one of about 295°C in T11 (die ring). Unfortunately, T7-T9 positions were not close enough to capture the real bearings temperature which should be very similar to the profile temperature (560 °C) [2].
indicated randomly with the code from P1 to P4 as requested by themselves. However, the code always identifies the same participant.

Predicted Press Load. The experimental extrusion load was compared with those predicted by the participants for the fourth extrusion (steady-state uncooled condition). High accuracy of the numerical predictions was obtained with an average error below 10% with respect to the experimental maximum value of 36.1 MN.

Profile Exit Speeds. During the first run, before the use of the puller, the velocity of profile 2 was qualitatively 50% higher than the other two profiles. Then, despite the application of a tensile force of 3500 N, profile 2 continued to flow through the puller at a higher speed. In Table 2 the experimental velocity of profiles 1 and 3 was the same as the puller (avg. speed of 195 mm/s), while for profile 2, the average speed of 225 mm/s (15% higher than the puller one) was obtained by calculating the difference in terms of extrudates length at the end of the extrusion. In this case, the bearings optimization could be not enough to easily compensate for the speed differences generated by a conical port design (Fig. 2). Indeed, the experimental evidence demonstrated that the speed disparity between profile 3 and profile 1 (both extruded with conical ports) was relatively minimal and negligible without and with the puller, respectively, despite variations in bearing configuration (choke for profile 3, 1 mm higher bearing length for profile 1). Furthermore, despite profile 2 having a 1 mm greater bearing length than profile 1, the former exhibited a significant faster flow due to the presence of cylindrical ports. The numerical results were aligned with the experimental evidence of the first run (no one simulated the presence of the puller), where profile 1 was faster than profile 3, remaining profile 2 the fastest one.

<table><caption>Table 2 Profile exit speeds: Experimental-Numerical comparison (billet 4)</caption><thead><tr><th>Profile exit speed / Billet 4</th><th>Profile 1 [mm/s]</th><th>Profile 2 [mm/s]</th><th>Profile 3 [mm/s]</th></tr></thead><tbody><tr><td>Experimental</td><td>195</td><td>225</td><td>195</td></tr><tr><td>P1</td><td>213 (Num Err. +9.2%)</td><td>254 (Num Err. +12.8%)</td><td>141 (Num Err. -27.7%)</td></tr><tr><td>P2</td><td>183 (Num Err. -6.2%)</td><td>251 (Num Err. +11.5%)</td><td>176 (Num Err. -9.7%)</td></tr><tr><td>P3</td><td>183 (Num Err. -6.2%)</td><td>261 (Num Err. +16%)</td><td>161 (Num Err. -17.4%)</td></tr><tr><td>P4</td><td>184 (Num Err. -5.7%)</td><td>245 (Num Err. +8.9%)</td><td>160 (Num Err. -17.9%)</td></tr></tbody></table>

Profiles exit temperatures. The contactless pyrometer only registered the exit profile temperature of profile 1 that, during the fourth and the fifth runs, had stabilized at around 560 °C. In this case, the numerical results of the participants differed from each other due to the selection of dissimilar heat exchange with air. Indeed, despite all participants obtaining profile temperatures in the range of 550-570 °C near the bearings, the differences were more marked at the location of the pyrometer also among users of the same software: 570°C, 536°C, 495°C, and 520°C was predicted by P1, P2, P3, and P4, respectively. Therefore, P1 better predicted the low heat exchange with air (num. err. +1.8%), while other participants overestimated its cooling efficiency (max err. of -12.3% for P3). The exit profile temperature plays an important role in the prediction of the microstructure [12-14], therefore also the proper calibration of the thermal boundary conditions “downstream of the process” is not to be underestimated.

Die temperatures in the eleven locations. Fig. 5 reports the thermocouples' temperature recorded during the steady state uncooled condition (billet 4). The thermocouples T4-T6 nearby the profiles' exit recorded a maximum value of about 444 °C. T1 and T2 nearby the inlet channel marked 361°C and 406°C, respectively. T9 (bearings exit 2) indicated the value of 490 °C; T8 (bearings exit 3) only 445 °C, while T7 (bearings exit 1) recorded an intermediate value of 473 °C. As expected, the highest temperature of 535 °C was observed in T10 (mandrel) and the lowest one of about 295°C in T11 (die ring). Unfortunately, T7-T9 positions were not close enough to capture the real bearings temperature which should be very similar to the profile temperature (560 °C) [2].
Moreover, T8 temperature was unexpectedly low if compared to T7 since profiles 1 and 3 had the same velocity (Table 2). Therefore, during the die loading, T8 had probably moved out of place.

<figure><img src="image_745.png" /><figcaption>Fig. 5 Thermal history of the fourth extrusion</figcaption></figure>

<table><caption>Table 3 Thermocouples' recordings: Experimental-Numerical comparison (billet 4)</caption><thead><tr><th colspan="11">Thermocouples' Temperature [°C]</th></tr><tr><th>Billet 4</th><th>T1</th><th>T2</th><th>T3</th><th>T4</th><th>T5</th><th>T6</th><th>T7</th><th>T8</th><th>T9</th><th>T10</th><th>T11</th></tr></thead><tbody><tr><td>Experimental</td><td>361</td><td>406</td><td>437</td><td>437</td><td>441</td><td>444</td><td>473</td><td>445</td><td>490</td><td>535</td><td>295</td></tr><tr><td>P1</td><td>365</td><td>390</td><td>405</td><td>405</td><td>407</td><td>409</td><td>496</td><td>497</td><td>499</td><td>492</td><td>334</td></tr><tr><td>P2</td><td>434</td><td>437</td><td>433</td><td>431</td><td>427</td><td>430</td><td>466</td><td>464</td><td>472</td><td>456</td><td>368</td></tr><tr><td>P3</td><td>374</td><td>396</td><td>408</td><td>406</td><td>408</td><td>412</td><td>482</td><td>461</td><td>494</td><td>452</td><td>310</td></tr><tr><td>P4</td><td>373</td><td>410</td><td>413</td><td>416</td><td>416</td><td>420</td><td>466</td><td>465</td><td>478</td><td>459</td><td>290</td></tr></tbody></table>

The numerical results (Table 3) were aligned with the experimental ones nearby the bearings (errors below 5% except in T8 for the above-mentioned reason), while average errors in the range of 5-10% were detected in the backer and the die ring (except for P2 with +25% in T11). A general temperature underestimation, in the range of 50-80 °C, was found only at T10 location (mandrel), however, the obtained results confirmed the improvement achieved over the years in terms of numerical thermal field prediction [2].

**Extrusion Defects.** The assessment of charge welds and billet skin defects involved the etching of the profile, the acquisition and the analysis of the images. All the analyzed profiles showed no evidence of billet skin contamination. Fig. 6b reports the evolution of charge welds for profile 1 in the transition between billets 3-4. The experimental trend for profile 2 was not analyzed due to the presence of the PCG (Peripheral Coarse Grain, Fig. 6c), which avoided precisely delineating the start and end points of the defect. For profiles 1 and 3, defects appeared within the range of +750-800 mm from the stopmark, followed by a complete replacement of the old billet material in the range of +2100-2200 mm (Fig. 6b). The numerical predictions underline the concrete possibility to optimize the scrap during the die design phase. Indeed, even if P4 greatly overestimated the start point of the charge welds with respect to the other participants, the end of the defect, and consequently, the distance from the stopmark to be scrapped was well predicted by everyone (max. err. of +10% for P4).
Moreover, T8 temperature was unexpectedly low if compared to T7 since profiles 1 and 3 had the same velocity (Table 2). Therefore, during the die loading, T8 had probably moved out of place.

<figure><img src="image_753.png" /><figcaption>Fig. 5 Thermal history of the fourth extrusion</figcaption></figure>

<table><caption>Table 3 Thermocouples' recordings: Experimental-Numerical comparison (billet 4)</caption><thead><tr><th colspan="11">Thermocouples' Temperature [°C]</th></tr><tr><th>Billet 4</th><th>T1</th><th>T2</th><th>T3</th><th>T4</th><th>T5</th><th>T6</th><th>T7</th><th>T8</th><th>T9</th><th>T10</th><th>T11</th></tr></thead><tbody><tr><td>Experimental</td><td>361</td><td>406</td><td>437</td><td>437</td><td>441</td><td>444</td><td>473</td><td>445</td><td>490</td><td>535</td><td>295</td></tr><tr><td>P1</td><td>365</td><td>390</td><td>405</td><td>405</td><td>407</td><td>409</td><td>496</td><td>497</td><td>499</td><td>492</td><td>334</td></tr><tr><td>P2</td><td>434</td><td>437</td><td>433</td><td>431</td><td>427</td><td>430</td><td>466</td><td>464</td><td>472</td><td>456</td><td>368</td></tr><tr><td>P3</td><td>374</td><td>396</td><td>408</td><td>406</td><td>408</td><td>412</td><td>482</td><td>461</td><td>494</td><td>452</td><td>310</td></tr><tr><td>P4</td><td>373</td><td>410</td><td>413</td><td>416</td><td>416</td><td>420</td><td>466</td><td>465</td><td>478</td><td>459</td><td>290</td></tr></tbody></table>

The numerical results (Table 3) were aligned with the experimental ones nearby the bearings (errors below 5% except in T8 for the above-mentioned reason), while average errors in the range of 5-10% were detected in the backer and the die ring (except for P2 with +25% in T11). A general temperature underestimation, in the range of 50-80 °C, was found only at T10 location (mandrel), however, the obtained results confirmed the improvement achieved over the years in terms of numerical thermal field prediction [2].

**Extrusion Defects.** The assessment of charge welds and billet skin defects involved the etching of the profile, the acquisition and the analysis of the images. All the analyzed profiles showed no evidence of billet skin contamination. Fig. 6b reports the evolution of charge welds for profile 1 in the transition between billets 3-4. The experimental trend for profile 2 was not analyzed due to the presence of the PCG (Peripheral Coarse Grain, Fig. 6c), which avoided precisely delineating the start and end points of the defect. For profiles 1 and 3, defects appeared within the range of +750-800 mm from the stopmark, followed by a complete replacement of the old billet material in the range of +2100-2200 mm (Fig. 6b). The numerical predictions underline the concrete possibility to optimize the scrap during the die design phase. Indeed, even if P4 greatly overestimated the start point of the charge welds with respect to the other participants, the end of the defect, and consequently, the distance from the stopmark to be scrapped was well predicted by everyone (max. err. of +10% for P4).
<figure><img src="image_746.png" /><figcaption>Fig. 6 Charge welds evolution for transitions between billets 3-4: a) Example of image analysis after etching, b) Experimental-Numerical comparison for the Profile 1, c) PCG layer in Profile 2 avoids evaluating the charge evolution</figcaption></figure>

Seam Weld Quality. The numerical prediction of the seam weld quality established if a good welding is obtained, evaluating the material conditions (strain, strain rate, and temperature) within the welding chamber during the joint formation. The criteria proposed within the software differ from each other, however, all numerical results predicted qualitatively welding with good strength. For the characterization of the seam weld quality, bulge test was selected [15] where, thanks to the pressure of rubber plug on the inside wall of the profile, a biaxial tensile stress is generated leading the tube to fracture in the proximity of the seam welds. Therefore, the quantitative comparison between the numerical welding criteria and the experimental outputs in terms of load and hoop strain at fracture is not possible. The hoop strain at fracture was estimated as the natural logarithmic ratio between the final and the initial tube circumferences. Profile 1, 2, and 3 presented average values of 0.13, 0.08, and 0.14, respectively. Moreover, the peak load at fracture follows the same trend with average values of 56.1 kN, 51.5 kN, and 57.6 kN, respectively. The results of the experimental investigation clearly indicate that profile 3 demonstrates the highest level of deformability, as evidenced by both the hoop strain and peak load measurements. Profile 1 follows closely in terms of deformability, whereas profile 2 exhibits lower values due to the presence of coarser grain structure, as discussed in the microstructure section.

<figure><img src="image_747.png" /><figcaption>Fig. 7 Comparison between air and water quenching in terms of obtained grain size</figcaption></figure>
<figure><img src="image_754.png" /><figcaption>Fig. 6 Charge welds evolution for transitions between billets 3-4: a) Example of image analysis after etching, b) Experimental-Numerical comparison for the Profile 1, c) PCG layer in Profile 2 avoids evaluating the charge evolution</figcaption></figure>

Seam Weld Quality. The numerical prediction of the seam weld quality established if a good welding is obtained, evaluating the material conditions (strain, strain rate, and temperature) within the welding chamber during the joint formation. The criteria proposed within the software differ from each other, however, all numerical results predicted qualitatively welding with good strength. For the characterization of the seam weld quality, bulge test was selected [15] where, thanks to the pressure of rubber plug on the inside wall of the profile, a biaxial tensile stress is generated leading the tube to fracture in the proximity of the seam welds. Therefore, the quantitative comparison between the numerical welding criteria and the experimental outputs in terms of load and hoop strain at fracture is not possible. The hoop strain at fracture was estimated as the natural logarithmic ratio between the final and the initial tube circumferences. Profile 1, 2, and 3 presented average values of 0.13, 0.08, and 0.14, respectively. Moreover, the peak load at fracture follows the same trend with average values of 56.1 kN, 51.5 kN, and 57.6 kN, respectively. The results of the experimental investigation clearly indicate that profile 3 demonstrates the highest level of deformability, as evidenced by both the hoop strain and peak load measurements. Profile 1 follows closely in terms of deformability, whereas profile 2 exhibits lower values due to the presence of coarser grain structure, as discussed in the microstructure section.

<figure><img src="image_755.png" /><figcaption>Fig. 7 Comparison between air and water quenching in terms of obtained grain size</figcaption></figure>
Microstructure. Specimens extracted from the middle of the extruded profiles were analyzed in terms of microstructure evolution. The image was acquired by using polarized light microscopy of the samples anodized (40 V dc, 4 min) with Barker's reagent (15 mL HBF4, 750 mL H2O). In order to measure the recrystallization layer on both the internal and external surfaces, measurements every 30° of the circumference were acquired (average values are summarized in Table 4). In Fig. 7, the comparison of microstructures obtained with air quenching and with water quenching are reported for profiles 1 and 3. The microstructure of profile 1 exhibits a fibrous grain structure with a thin PCG layer on the internal and external surfaces, while profile 3 shows a completely fibrous microstructure. In contrast, as previously shown in Fig. 6c, profile 2 presents a very large PCG thickness, both internally and externally, caused by the higher exit speed of the profile. As depicted in Fig. 7, the impact of water quenching on the evolution of grain size was found to be negligible. This suggests that the formation of PCG occurred within a shorter timeframe than the interval between the profile exit from the bearings and its arrival at the water quenching box. Furthermore, the obtained results emphasized the strong dependency between PCG formation and material conditions such as strain, strain rate, and temperature, as proved by the significant variations in microstructure observed between profiles 2 (with thick PCG layers) and 3 (with an absence of PCG). All participants (except for P4 that did not perform this simulation) used models for dynamic recrystallization that do not take into account the PCG formation [12-14]. Therefore, the numerical outputs showed large discrepancies with the experimental results, predicting (all participants in a similar way) a grain size within the range 5-80 [µm] with the absence of PCG. Numerical findings underline that the prediction of the microstructure evolution, especially for complex case studies, still remains a challenge for the industrial scene.

<table><caption>Table 4 Experimental-Numerical comparison in terms of obtained grain size (billet 5)</caption><thead><tr><th>Billet 5</th><th>Avg Int PCG layer [µm]</th><th>Avg Ext PCG Layer [µm]</th><th colspan="2">Grain size range [µm]</th></tr></thead><tbody><tr><td>Exp Exit 1</td><td>201</td><td>523</td><td>5-30 Fibrous</td><td>131-832 PCG</td></tr><tr><td>Exp Exit 2</td><td>485</td><td>1154</td><td>5-30 Fibrous</td><td>142-1885 PCG</td></tr><tr><td>Exp Exit 3</td><td>0</td><td>0</td><td>5-30 Fibrous</td><td>0 PCG</td></tr></tbody></table>

Thermal field with nitrogen cooling. Fig. 8 exhibits the effect of nitrogen cooling on the exit profile temperature, the bearings, and the mandrel, comparing the results with the uncooled condition (billet 4 vs billet 11). In detail, the cooling efficiency on the profile surfaces was limited, causing a sole 10°C of change (from 560°C to 550°C for profile 1). The channel design promoted the nitrogen flow towards the exit of profile 3, a trend confirmed by the bearings temperatures, where the higher cooling effect was captured by T8 (exit 3, from 445 °C to 417 °C), while the lower one by T7 (exit 1, from 475°C to 463°C). The effect of nitrogen cooling was remarkable within the backer nearby the inlet channel and along the first part of the cooling path, registering in T1-T4 locations a general decrease of temperature in the range of 100-120 °C (table 5). In this case, the distance of the cooling channels from the bearings and the pressure drops induced by the design, which promoted an unbalanced nitrogen flow along the path, limited the effect of the cooling on the profile surface. Therefore, according to the state of art [2, 16-17], the experimental campaign showed the potentiality of the nitrogen cooling, but also the wide margins of improvement of the channels if designed with the support of advanced numerical tools. As previously given for the microstructure prediction, the numerical results of the participants can be improved. P4 did not simulate the process with cooling, both P2 and P3 used equivalent thermal boundary conditions along the path, while only P1 implemented the study of the nitrogen flow along the channel. All results showed a general overestimation of the cooling efficiency within the tooling set, especially along the cooling path (T1-T6), not properly catching the differences in
Microstructure. Specimens extracted from the middle of the extruded profiles were analyzed in terms of microstructure evolution. The image was acquired by using polarized light microscopy of the samples anodized (40 V dc, 4 min) with Barker's reagent (15 mL HBF4, 750 mL H2O). In order to measure the recrystallization layer on both the internal and external surfaces, measurements every 30° of the circumference were acquired (average values are summarized in Table 4). In Fig. 7, the comparison of microstructures obtained with air quenching and with water quenching are reported for profiles 1 and 3. The microstructure of profile 1 exhibits a fibrous grain structure with a thin PCG layer on the internal and external surfaces, while profile 3 shows a completely fibrous microstructure. In contrast, as previously shown in Fig. 6c, profile 2 presents a very large PCG thickness, both internally and externally, caused by the higher exit speed of the profile. As depicted in Fig. 7, the impact of water quenching on the evolution of grain size was found to be negligible. This suggests that the formation of PCG occurred within a shorter timeframe than the interval between the profile exit from the bearings and its arrival at the water quenching box. Furthermore, the obtained results emphasized the strong dependency between PCG formation and material conditions such as strain, strain rate, and temperature, as proved by the significant variations in microstructure observed between profiles 2 (with thick PCG layers) and 3 (with an absence of PCG). All participants (except for P4 that did not perform this simulation) used models for dynamic recrystallization that do not take into account the PCG formation [12-14]. Therefore, the numerical outputs showed large discrepancies with the experimental results, predicting (all participants in a similar way) a grain size within the range 5-80 [µm] with the absence of PCG. Numerical findings underline that the prediction of the microstructure evolution, especially for complex case studies, still remains a challenge for the industrial scene.

<table><caption>Table 4 Experimental-Numerical comparison in terms of obtained grain size (billet 5)</caption><thead><tr><th>Billet 5</th><th>Avg Int PCG layer [µm]</th><th>Avg Ext PCG Layer [µm]</th><th colspan="2">Grain size range [µm]</th></tr></thead><tbody><tr><td>Exp Exit 1</td><td>201</td><td>523</td><td>5-30 Fibrous</td><td>131-832 PCG</td></tr><tr><td>Exp Exit 2</td><td>485</td><td>1154</td><td>5-30 Fibrous</td><td>142-1885 PCG</td></tr><tr><td>Exp Exit 3</td><td>0</td><td>0</td><td>5-30 Fibrous</td><td>0 PCG</td></tr></tbody></table>

Thermal field with nitrogen cooling. Fig. 8 exhibits the effect of nitrogen cooling on the exit profile temperature, the bearings, and the mandrel, comparing the results with the uncooled condition (billet 4 vs billet 11). In detail, the cooling efficiency on the profile surfaces was limited, causing a sole 10°C of change (from 560°C to 550°C for profile 1). The channel design promoted the nitrogen flow towards the exit of profile 3, a trend confirmed by the bearings temperatures, where the higher cooling effect was captured by T8 (exit 3, from 445 °C to 417 °C), while the lower one by T7 (exit 1, from 475°C to 463°C). The effect of nitrogen cooling was remarkable within the backer nearby the inlet channel and along the first part of the cooling path, registering in T1-T4 locations a general decrease of temperature in the range of 100-120 °C (table 5). In this case, the distance of the cooling channels from the bearings and the pressure drops induced by the design, which promoted an unbalanced nitrogen flow along the path, limited the effect of the cooling on the profile surface. Therefore, according to the state of art [2, 16-17], the experimental campaign showed the potentiality of the nitrogen cooling, but also the wide margins of improvement of the channels if designed with the support of advanced numerical tools. As previously given for the microstructure prediction, the numerical results of the participants can be improved. P4 did not simulate the process with cooling, both P2 and P3 used equivalent thermal boundary conditions along the path, while only P1 implemented the study of the nitrogen flow along the channel. All results showed a general overestimation of the cooling efficiency within the tooling set, especially along the cooling path (T1-T6), not properly catching the differences in
nitrogen flow rate towards the three exits of the profiles. The usage of equivalent heat transfer coefficients, if not properly calibrated, does not allow capturing the change in the heat transfer induced by the process settings and the channel design (pressure drops, nitrogen phase-change...). On the other hand, the FEM simulation of nitrogen flow can be inaccurate if the nitrogen gas formation generated by poorly designed solutions is not taken into account [16-17].

<figure><img src="image_748.png" /><figcaption>Fig. 8 Effect of nitrogen cooling on the profile temperature and the die thermal field</figcaption></figure>

<table><caption>Table 5 Thermocouples' recordings: Experimental-Numerical comparison (billet 11)</caption><thead><tr><th colspan="11">Thermocouples' Temperature [°C]</th></tr><tr><th>Billet 11</th><th>T1</th><th>T2</th><th>T3</th><th>T4</th><th>T5</th><th>T6</th><th>T7</th><th>T8</th><th>T9</th><th>T10</th><th>T11</th></tr></thead><tbody><tr><td>Experimental</td><td>234</td><td>323</td><td>320</td><td>320</td><td>330</td><td>360</td><td>463</td><td>417</td><td>471</td><td>520</td><td>286</td></tr><tr><td>P1</td><td>148</td><td>192</td><td>117</td><td>129</td><td>138</td><td>196</td><td>476</td><td>468</td><td>477</td><td>491</td><td>334</td></tr><tr><td>P2</td><td>121</td><td>120</td><td>36</td><td>53</td><td>70</td><td>64</td><td>376</td><td>376</td><td>380</td><td>456</td><td>355</td></tr><tr><td>P3</td><td>113</td><td>189</td><td>208</td><td>226</td><td>236</td><td>271</td><td>467</td><td>440</td><td>474</td><td>451</td><td>301</td></tr></tbody></table>

## Conclusions

In the Extrusion Benchmark 2023, the processing conditions for extruding three AA 6082 tubes were monitored and analyzed. The tests were performed on air and water quenched conditions and with the use or absence of liquid nitrogen cooling. In detail, the following experimental outputs were acquired for the comparison with the FEM results of the participants: press load, profile speed, profile temperatures, die temperatures in eleven different locations, microstructure size, charge welds and billet skin evolution, and seam weld quality with bulge test. Good accuracy of the numerical results was demonstrated in terms of press load, profile speed, and thermal field in uncooled conditions as well as in terms of evaluation of the extrusion scraps. A reliable forecast of the microstructure size and the effect of nitrogen cooling still remains a challenge.

## Funding

National Recovery and Resilience Plan (NRRP), Mission 04 Component 2 Investment 1.5-NextGenerationEU, Call for tender n.3277 dated 30/12/2021 (Award Number: 0001052 dated 23/06/2022).

## References

[1] B. Reggiani, L. Donati, L. Tomesani, ICEB - International conference on extrusion and benchmark, Light Metal Age 75 (5) (2017) 52-60.

[2] L. Donati, B. Reggiani, R. Pelaccia, M. Negozio, S. Di Donato, Advancements in extrusion and drawing: a review of the contributes by the ESAFORM community, Int. J. Mater. Form. 15 (2022) 41. https://doi.org/10.1007/s12289-022-01664-w
nitrogen flow rate towards the three exits of the profiles. The usage of equivalent heat transfer coefficients, if not properly calibrated, does not allow capturing the change in the heat transfer induced by the process settings and the channel design (pressure drops, nitrogen phase-change...). On the other hand, the FEM simulation of nitrogen flow can be inaccurate if the nitrogen gas formation generated by poorly designed solutions is not taken into account [16-17].

<figure><img src="image_756.png" /><figcaption>Fig. 8 Effect of nitrogen cooling on the profile temperature and the die thermal field</figcaption></figure>

<table><caption>Table 5 Thermocouples' recordings: Experimental-Numerical comparison (billet 11)</caption><thead><tr><th colspan="11">Thermocouples' Temperature [°C]</th></tr><tr><th>Billet 11</th><th>T1</th><th>T2</th><th>T3</th><th>T4</th><th>T5</th><th>T6</th><th>T7</th><th>T8</th><th>T9</th><th>T10</th><th>T11</th></tr></thead><tbody><tr><td>Experimental</td><td>234</td><td>323</td><td>320</td><td>320</td><td>330</td><td>360</td><td>463</td><td>417</td><td>471</td><td>520</td><td>286</td></tr><tr><td>P1</td><td>148</td><td>192</td><td>117</td><td>129</td><td>138</td><td>196</td><td>476</td><td>468</td><td>477</td><td>491</td><td>334</td></tr><tr><td>P2</td><td>121</td><td>120</td><td>36</td><td>53</td><td>70</td><td>64</td><td>376</td><td>376</td><td>380</td><td>456</td><td>355</td></tr><tr><td>P3</td><td>113</td><td>189</td><td>208</td><td>226</td><td>236</td><td>271</td><td>467</td><td>440</td><td>474</td><td>451</td><td>301</td></tr></tbody></table>

## Conclusions

In the Extrusion Benchmark 2023, the processing conditions for extruding three AA 6082 tubes were monitored and analyzed. The tests were performed on air and water quenched conditions and with the use or absence of liquid nitrogen cooling. In detail, the following experimental outputs were acquired for the comparison with the FEM results of the participants: press load, profile speed, profile temperatures, die temperatures in eleven different locations, microstructure size, charge welds and billet skin evolution, and seam weld quality with bulge test. Good accuracy of the numerical results was demonstrated in terms of press load, profile speed, and thermal field in uncooled conditions as well as in terms of evaluation of the extrusion scraps. A reliable forecast of the microstructure size and the effect of nitrogen cooling still remains a challenge.

## Funding

National Recovery and Resilience Plan (NRRP), Mission 04 Component 2 Investment 1.5-NextGenerationEU, Call for tender n.3277 dated 30/12/2021 (Award Number: 0001052 dated 23/06/2022).

## References

[1] B. Reggiani, L. Donati, L. Tomesani, ICEB - International conference on extrusion and benchmark, Light Metal Age 75 (5) (2017) 52-60.

[2] L. Donati, B. Reggiani, R. Pelaccia, M. Negozio, S. Di Donato, Advancements in extrusion and drawing: a review of the contributes by the ESAFORM community, Int. J. Mater. Form. 15 (2022) 41. https://doi.org/10.1007/s12289-022-01664-w
[3] M. Schikorra, L. Donati, L. Tomesani, A. E. Tekkaya, Extrusion Benchmark 2007 – Benchmark Experiments: Study on Material Flow Extrusion of a Flat Die, Key Eng. Mater. 367 (2008) 1-8. https://doi.org/10.4028/www.scientific.net/KEM.367.1

[4] D. Pietzka, et al., Extrusion Benchmark 2009 - Experimental analysis of deflection in extrusion dies. Key Eng. Mater. 424 (2010) 19 - 26. https://doi.org/10.4028/www.scientific.net/KEM.424.19

[5] A. Selvaggio, et al., Extrusion Benchmark 2011: Evaluation of different design strategies on process conditions, die deflection and seam weld quality in hollow profiles, Key Eng. Mater. 491 (2012) 1 – 10. https://doi.org/10.4028/www.scientific.net/KEM.491.1

[6] L. Donati, et al., Effect of liquid nitrogen die cooling on extrusion process conditions Key Eng. Mater. 491 (2012) 215 - 222. https://doi.org/10.4028/www.scientific.net/KEM.491.215

[7] A. Selvaggio, T. Kloppenborg, M. Schwane, et al., Extrusion Benchmark 2013 - Experimental Analysis of Mandrel Deflection, Local Temperature and Pressure in Extrusion Dies, Key Eng. Mater. 585 (2013) 13 – 22. https://doi.org/10.4028/www.scientific.net/KEM.585.13

[8] A. Selvaggio, L. Donati, et al., Scientific Benchmark 2015: Effect of choking and bearing length on metal flow balancing in extrusion dies, Mater. Today: Proc 2 (2015) 4704-4713. https://doi.org/10.1016/j.matpr.2015.10.003

[9] A. Gamberoni, L. Donati, et al., Industrial Benchmark 2015: process monitoring and analysis of hollow EN AW-6063 extruded profile, Mater. Today: Proc 2 (2015) 4714-4725. https://doi.org/10.1016/j.matpr.2015.10.004

[10] L. Donati L., A. Segatori et al., Extrusion Benchmark 2017: Effect of Die Design on Profile Quality and Distortions of Thin C-Shaped Hollow Profiles, Mater. Today: Proc 10 (2019) 171–184. https://doi.org/10.1016/j.matpr.2018.10.394

[11] R. Pelaccia, M. Negozio, et al. Extrusion Benchmark 2023: Effect of die design on profile speed, seam weld quality and microstructure of hollow tubes, Key Eng. Mater, under press.

[12] M. Negozio, et al., Numerical investigation of the surface recrystallization during the extrusion of a AA6082 aluminum alloy under different process conditions Int. J. Adv. Manuf. Technol. 129 (3-4) (2023) 1585-1599. https://doi.org/10.1007/s00170-023-12397-8

[13] T. Sheppard, Prediction of structure during shaped extrusion and subsequent static recrystallisation during the solution soaking operation, J. Mater. Process. Technol. 177 (1-3) (2006) 26-35. https://doi.org/10.1016/j.jmatprotec.2006.04.099

[14] M. Negozio, R. Pelaccia, L. Donati, et al., Simulation of the microstructure evolution during the extrusion of two industrial-scale AA6063 profiles, J. Manuf. Process. 99 (2023) 501-512. https://doi.org/10.1016/j.jmapro.2023.05.075

[15] B. Reggiani, L. Donati, Comparison of experimental methods to evaluate the seam welds quality in extruded profiles, Trans. Nonferrous Met. Soc China, 30 (3) (2020) 619-634. https://doi.org/10.1016/S1003-6326(20)65241-4

[16] R. Pelaccia, B. Reggiani, et al., Liquid nitrogen in the industrial practice of hot aluminium extrusion: experimental and numerical investigation, Int J Adv Manuf Technol. 119 (2022) 3141–3155. https://doi.org/10.1007/s00170-021-08422-3

[17] R. Pelaccia, P. E. Santangelo, A Homogeneous Flow Model for nitrogen cooling in the aluminum-alloy extrusion process, Int. J. Heat Mass Transf. 195 (2022) 123202. https://doi.org/10.1016/j.ijheatmasstransfer.2022.123202
[3] M. Schikorra, L. Donati, L. Tomesani, A. E. Tekkaya, Extrusion Benchmark 2007 – Benchmark Experiments: Study on Material Flow Extrusion of a Flat Die, Key Eng. Mater. 367 (2008) 1-8. https://doi.org/10.4028/www.scientific.net/KEM.367.1

[4] D. Pietzka, et al., Extrusion Benchmark 2009 - Experimental analysis of deflection in extrusion dies. Key Eng. Mater. 424 (2010) 19 - 26. https://doi.org/10.4028/www.scientific.net/KEM.424.19

[5] A. Selvaggio, et al., Extrusion Benchmark 2011: Evaluation of different design strategies on process conditions, die deflection and seam weld quality in hollow profiles, Key Eng. Mater. 491 (2012) 1 – 10. https://doi.org/10.4028/www.scientific.net/KEM.491.1

[6] L. Donati, et al., Effect of liquid nitrogen die cooling on extrusion process conditions Key Eng. Mater. 491 (2012) 215 - 222. https://doi.org/10.4028/www.scientific.net/KEM.491.215

[7] A. Selvaggio, T. Kloppenborg, M. Schwane, et al., Extrusion Benchmark 2013 - Experimental Analysis of Mandrel Deflection, Local Temperature and Pressure in Extrusion Dies, Key Eng. Mater. 585 (2013) 13 – 22. https://doi.org/10.4028/www.scientific.net/KEM.585.13

[8] A. Selvaggio, L. Donati, et al., Scientific Benchmark 2015: Effect of choking and bearing length on metal flow balancing in extrusion dies, Mater. Today: Proc 2 (2015) 4704-4713. https://doi.org/10.1016/j.matpr.2015.10.003

[9] A. Gamberoni, L. Donati, et al., Industrial Benchmark 2015: process monitoring and analysis of hollow EN AW-6063 extruded profile, Mater. Today: Proc 2 (2015) 4714-4725. https://doi.org/10.1016/j.matpr.2015.10.004

[10] L. Donati L., A. Segatori et al., Extrusion Benchmark 2017: Effect of Die Design on Profile Quality and Distortions of Thin C-Shaped Hollow Profiles, Mater. Today: Proc 10 (2019) 171–184. https://doi.org/10.1016/j.matpr.2018.10.394

[11] R. Pelaccia, M. Negozio, et al. Extrusion Benchmark 2023: Effect of die design on profile speed, seam weld quality and microstructure of hollow tubes, Key Eng. Mater, under press.

[12] M. Negozio, et al., Numerical investigation of the surface recrystallization during the extrusion of a AA6082 aluminum alloy under different process conditions Int. J. Adv. Manuf. Technol. 129 (3-4) (2023) 1585-1599. https://doi.org/10.1007/s00170-023-12397-8

[13] T. Sheppard, Prediction of structure during shaped extrusion and subsequent static recrystallisation during the solution soaking operation, J. Mater. Process. Technol. 177 (1-3) (2006) 26-35. https://doi.org/10.1016/j.jmatprotec.2006.04.099

[14] M. Negozio, R. Pelaccia, L. Donati, et al., Simulation of the microstructure evolution during the extrusion of two industrial-scale AA6063 profiles, J. Manuf. Process. 99 (2023) 501-512. https://doi.org/10.1016/j.jmapro.2023.05.075

[15] B. Reggiani, L. Donati, Comparison of experimental methods to evaluate the seam welds quality in extruded profiles, Trans. Nonferrous Met. Soc China, 30 (3) (2020) 619-634. https://doi.org/10.1016/S1003-6326(20)65241-4

[16] R. Pelaccia, B. Reggiani, et al., Liquid nitrogen in the industrial practice of hot aluminium extrusion: experimental and numerical investigation, Int J Adv Manuf Technol. 119 (2022) 3141–3155. https://doi.org/10.1007/s00170-021-08422-3

[17] R. Pelaccia, P. E. Santangelo, A Homogeneous Flow Model for nitrogen cooling in the aluminum-alloy extrusion process, Int. J. Heat Mass Transf. 195 (2022) 123202. https://doi.org/10.1016/j.ijheatmasstransfer.2022.123202
# Optimization of working conditions and increase of productivity of aluminum hot extrusion press based on finite element analysis

LLORCA-SCHENK Juan¹, a*, GAYASHAN Lasindu², b and HEWAGE Praveen², c

¹ University of Alicante. Carretera San Vicente del Raspeig s/n. 03690 San Vicente del Raspeig. Spain

² Alumex Group. Pattiwila Road, Sapugaskanda, Makola. Sri Lanka

ajuan.llorca@ua.es, b lasindu.handaragama@alumexgroup.com,

c praveen.hewage@alumexgroup.com

**Keywords:** Aluminum Extrusion, Productivity Increase, FE Analysis

**Abstract.** The finite element simulation of extrusion allows a precise analysis of the metal flow, but also a detailed evaluation of the various parameters involved in the process: temperatures, pressures, velocities and stresses. The press productivity is directly related to the extrusion velocity. In turn, the maximum achievable extrusion velocity is limited by the exit temperature of the formed profile. Traditionally, most presses set a general billet temperature conservatively to ensure that extrusion does not exceed the pressure and stress limits for most dies. This study aims to provide an additional and non-exclusive methodology to increase productivity based on the correct choice of the billet temperature at the press with the help of a previous analysis by means of finite element simulation. A controlled reduction of the billet temperature makes it possible to increase the extrusion velocity while keeping the outlet temperature, stresses and pressure within the admissible range. The use of finite element simulation avoids iterative processes with several press trials and avoids the risks of excessive press pressure or tool stresses. Thus, it allows a quick determination of the optimum temperature and performs risk-free prediction. To demonstrate the use of the methodology, a billet temperature optimisation for an example die has been developed using the simulation software Qform UK. Together with the theoretical analysis, the results of the real die in the press in collaboration with Alumex PLC (Sri Lanka) are shown. This methodology is shown to be effective and may be of particular interest to companies already using extrusion simulation for other uses, such as die design.

**Introduction**

The cost-effectiveness and quality grade of extruded profiles are their main commercial factors. Both factors are directly related to the quality and performance of the extrusion die. In addition, there are other factors such as the characteristics of the extrusion press, the quality of the billet material, the capabilities of the auxiliary equipment and the downstream operations such as age hardening, painting, anodising or other possible. Due to its very fine tolerances, special material and high thermo-mechanical fatigue performance requirements, the die is probably the most critical extrusion element [1].

The use of finite element (FE) simulation for the optimization of hot extrusion dies design has become widespread among most die makers [2]. The simulation provides a very precise analysis of the behaviour of the metal during the deformation process. At the same time, it allows a detailed evaluation of the different parameters that are involved and determine the production process: temperatures, pressures, stresses...

The production level obtained by the press is directly related to the output speed of the extruded profile, called extrusion velocity. In addition, the maximum achievable extrusion velocity is limited by the temperature of the formed profile at the exit of the die. Therefore, the thermal

<figure><img src="image_749.png" /></figure>
# Optimization of working conditions and increase of productivity of aluminum hot extrusion press based on finite element analysis

LLORCA-SCHENK Juan<fnref n="1" />, a*, GAYASHAN Lasindu<fnref n="2" />, b and HEWAGE Praveen<fnref n="2" />, c

**Keywords:** Aluminum Extrusion, Productivity Increase, FE Analysis

**Abstract.** The finite element simulation of extrusion allows a precise analysis of the metal flow, but also a detailed evaluation of the various parameters involved in the process: temperatures, pressures, velocities and stresses. The press productivity is directly related to the extrusion velocity. In turn, the maximum achievable extrusion velocity is limited by the exit temperature of the formed profile. Traditionally, most presses set a general billet temperature conservatively to ensure that extrusion does not exceed the pressure and stress limits for most dies. This study aims to provide an additional and non-exclusive methodology to increase productivity based on the correct choice of the billet temperature at the press with the help of a previous analysis by means of finite element simulation. A controlled reduction of the billet temperature makes it possible to increase the extrusion velocity while keeping the outlet temperature, stresses and pressure within the admissible range. The use of finite element simulation avoids iterative processes with several press trials and avoids the risks of excessive press pressure or tool stresses. Thus, it allows a quick determination of the optimum temperature and performs risk-free prediction. To demonstrate the use of the methodology, a billet temperature optimisation for an example die has been developed using the simulation software Qform UK. Together with the theoretical analysis, the results of the real die in the press in collaboration with Alumex PLC (Sri Lanka) are shown. This methodology is shown to be effective and may be of particular interest to companies already using extrusion simulation for other uses, such as die design.

**Introduction**

The cost-effectiveness and quality grade of extruded profiles are their main commercial factors. Both factors are directly related to the quality and performance of the extrusion die. In addition, there are other factors such as the characteristics of the extrusion press, the quality of the billet material, the capabilities of the auxiliary equipment and the downstream operations such as age hardening, painting, anodising or other possible. Due to its very fine tolerances, special material and high thermo-mechanical fatigue performance requirements, the die is probably the most critical extrusion element [1].

The use of finite element (FE) simulation for the optimization of hot extrusion dies design has become widespread among most die makers [2]. The simulation provides a very precise analysis of the behaviour of the metal during the deformation process. At the same time, it allows a detailed evaluation of the different parameters that are involved and determine the production process: temperatures, pressures, stresses...

The production level obtained by the press is directly related to the output speed of the extruded profile, called extrusion velocity. In addition, the maximum achievable extrusion velocity is limited by the temperature of the formed profile at the exit of the die. Therefore, the thermal

<figure><img src="image_757.png" /></figure>

<footnote n="1">¹ University of Alicante. Carretera San Vicente del Raspeig s/n. 03690 San Vicente del Raspeig. Spain</footnote>

<footnote n="2">² Alumex Group. Pattiwila Road, Sapugaskanda, Makola. Sri Lanka</footnote>

<footnote>ajuan.llorca@ua.es, b lasindu.handaragama@alumexgroup.com,</footnote>

<footnote>c praveen.hewage@alumexgroup.com</footnote>
gradients involved during the process affect its efficiency. For this reason, strict temperature control is required to optimise the extrusion process. In aluminium hot extrusion, the billet needs to be preheated to properly extrude the alloy and during the process the deformation energy and frictional forces between the billet and the tools cause a temperature increase in both the die and the profile. An excessive profile outlet temperature can cause various aesthetic defects or cracks on the profile surface, which can lead to scrapped products or reduced tool life. The exit profile temperature also increases considerably with extrusion speed, which creates a significant limitation in terms of maximum achievable productivity [3-6].

An effective and widespread solution to increase productivity through temperature control is the direct cooling with nitrogen channels immediately at the outlet of the formed profile. Nitrogen injection during the process has proven to be effective in achieving optimum temperature control and improved surface finish of the extruded profile. However, it requires investment in equipment and has an operational cost [7].

Usually, the choice of the billet temperature follows a criterion often linked to a traditional criterion that tried to define a single temperature valid for all dies. This includes those dies that require strict control of the maximum forces exerted by the press due to their resistance characteristics. The temperature set according to this criterion is usually in the range 460-480 °C [7-8]. In most presses, the same temperature is used for all dies. It should be noted that for the initial billets the temperatures used are lower and sometimes conical heating is used. Changing the billet temperature without any risk is possible for the vast majority of dies, especially if the maximum force exerted during the process is controlled through simulation and the die components are designed to withstand the maximum press force.

The present study aims to provide an additional and non-exclusive methodology to increase productivity based on the optimal choice of billet temperature in the press with the help of a previous analysis by means of FE simulation. A lower billet temperature at the inlet of the press translates into a lower outlet temperature of the extruded profile [9]. As the outlet temperature is the limiting factor of the process speed, lowering the outlet temperature opens the possibility to increase the process velocity [10]. It is important to note that the reduction of the inlet temperature is associated with an increase of the maximum force in the press and, therefore, of the stresses supported by the tools [11]. These parameters need to be controlled through simulation in order not to exceed the limits.

This simulation-based methodology for increasing process speed offers the possibility of achieving higher productivity on the press without investment in equipment and without additional operational costs.

## Materials and methods

The methodology used for this study focuses on the use of hot aluminium extrusion process simulation to show the possibility of increasing the process speed and productivity through a reduction of the billet temperature. It focuses on steady state of the process.

It is important to note that it is vitally important to make use of simulation to analyse the most suitable temperature for the billet. Lowering the temperature of the billet increases the force required to deform the aluminium during extrusion. By using FE simulation, it is possible to properly control the maximum force required in order not to exceed the limits of the press. It is also possible to control the stresses in the die parts. Although most dies are designed to support the maximum press force, in some cases there is not so much allowance to ensure a long life of the die.
gradients involved during the process affect its efficiency. For this reason, strict temperature control is required to optimise the extrusion process. In aluminium hot extrusion, the billet needs to be preheated to properly extrude the alloy and during the process the deformation energy and frictional forces between the billet and the tools cause a temperature increase in both the die and the profile. An excessive profile outlet temperature can cause various aesthetic defects or cracks on the profile surface, which can lead to scrapped products or reduced tool life. The exit profile temperature also increases considerably with extrusion speed, which creates a significant limitation in terms of maximum achievable productivity [3–6].

An effective and widespread solution to increase productivity through temperature control is the direct cooling with nitrogen channels immediately at the outlet of the formed profile. Nitrogen injection during the process has proven to be effective in achieving optimum temperature control and improved surface finish of the extruded profile. However, it requires investment in equipment and has an operational cost [7].

Usually, the choice of the billet temperature follows a criterion often linked to a traditional criterion that tried to define a single temperature valid for all dies. This includes those dies that require strict control of the maximum forces exerted by the press due to their resistance characteristics. The temperature set according to this criterion is usually in the range 460-480 °C [7-8]. In most presses, the same temperature is used for all dies. It should be noted that for the initial billets the temperatures used are lower and sometimes conical heating is used. Changing the billet temperature without any risk is possible for the vast majority of dies, especially if the maximum force exerted during the process is controlled through simulation and the die components are designed to withstand the maximum press force.

The present study aims to provide an additional and non-exclusive methodology to increase productivity based on the optimal choice of billet temperature in the press with the help of a previous analysis by means of FE simulation. A lower billet temperature at the inlet of the press translates into a lower outlet temperature of the extruded profile [9]. As the outlet temperature is the limiting factor of the process speed, lowering the outlet temperature opens the possibility to increase the process velocity [10]. It is important to note that the reduction of the inlet temperature is associated with an increase of the maximum force in the press and, therefore, of the stresses supported by the tools [11]. These parameters need to be controlled through simulation in order not to exceed the limits.

This simulation-based methodology for increasing process speed offers the possibility of achieving higher productivity on the press without investment in equipment and without additional operational costs.

## Materials and methods

The methodology used for this study focuses on the use of hot aluminium extrusion process simulation to show the possibility of increasing the process speed and productivity through a reduction of the billet temperature. It focuses on steady state of the process.

It is important to note that it is vitally important to make use of simulation to analyse the most suitable temperature for the billet. Lowering the temperature of the billet increases the force required to deform the aluminium during extrusion. By using FE simulation, it is possible to properly control the maximum force required in order not to exceed the limits of the press. It is also possible to control the stresses in the die parts. Although most dies are designed to support the maximum press force, in some cases there is not so much allowance to ensure a long life of the die.
<figure><img src="image_750.png" /><figcaption>Fig. 1. Drawing of the die used for the experience.</figcaption></figure>

An example die is used to illustrate the effects of the temperature variations proposed by the methodology. In addition, trial results are shown for the same die in the 23 MN and 178 mm container press of Alumex PLC (Sri Lanka). All simulations and trials have been carried out with the 6XXX group of aluminium alloys, in particular with the EN AW-6063-O alloy.

The die chosen for the application example is a 6-cavity solid die with a diameter of 230 mm and a total height of 130 mm. It consists of two parts: the die plate (50 mm) and the backer (80 mm). In the die plate there is only a single 10 mm deep feeder and the profile geometry with small bearing variations at the tips and due to the distance to the centre (Fig. 1). The die was produced by Compes, Spa.

<figure><img src="image_751.png" /><figcaption>Fig. 2. Drawing of the profile used for the experience.</figcaption></figure>

The extruded profile is a small profile (Fig. 2) with a weight of 0.151 kg/m. Therefore, the 6-cavity die has an extrusion ratio of around 75. This is a die that has already worked on several occasions prior to this research and Alumex PLC wishes to increase its productivity (Fig. 3).
<figure><img src="image_758.png" /><figcaption>Fig. 1. Drawing of the die used for the experience.</figcaption></figure>

An example die is used to illustrate the effects of the temperature variations proposed by the methodology. In addition, trial results are shown for the same die in the 23 MN and 178 mm container press of Alumex PLC (Sri Lanka). All simulations and trials have been carried out with the 6XXX group of aluminium alloys, in particular with the EN AW-6063-O alloy.

The die chosen for the application example is a 6-cavity solid die with a diameter of 230 mm and a total height of 130 mm. It consists of two parts: the die plate (50 mm) and the backer (80 mm). In the die plate there is only a single 10 mm deep feeder and the profile geometry with small bearing variations at the tips and due to the distance to the centre (Fig. 1). The die was produced by Compes, Spa.

<figure><img src="image_759.png" /><figcaption>Fig. 2. Drawing of the profile used for the experience.</figcaption></figure>

The extruded profile is a small profile (Fig. 2) with a weight of 0.151 kg/m. Therefore, the 6-cavity die has an extrusion ratio of around 75. This is a die that has already worked on several occasions prior to this research and Alumex PLC wishes to increase its productivity (Fig. 3).
In order to evaluate the benefits of controlled reduction of the billet temperature, several simulations will be carried out by changing the temperature and speed parameters. After analysing the best way to show the possibilities of the methodology, 3 sets of conditions of billet temperature, extrusion velocity and billet length have been defined. Table 1 shows the conditions chosen for the different trials.

<figure><img src="image_752.png" /><figcaption>Fig. 3. Image of the die used throughout the experiment.</figcaption></figure>

The first trial uses the conditions under which this die used to extrude in Alumex PLC. There, usually all dies extrude with a billet temperature in the range 460-470°C.

The second trial uses a billet temperature 20°C lower and doubles the extrusion speed. These are the conditions that have been optimised with the help of FE simulation to significantly increase the speed while keeping the outlet temperature and maximum press force at admissible values. The condition set for determining the temperature was not to exceed 550°C at the exit of the profile and not to approach the maximum press force (23MN) during the process.

The conditions of the third trial are chosen to show that, in the event of a possible excess of force in the press due to the temperature reduction, it is possible to reduce the billet length to reduce the maximum force required.

<table><caption>Table 1. Conditions chosen for the different research trials.</caption><thead><tr><td>Trial number</td><td>Billet temperature [°C]</td><td>Ram / extrusion velocity [mm/s]</td><td>Billet length [mm]</td></tr></thead><tbody><tr><td>1</td><td>467</td><td>3</td><td>700</td></tr><tr><td>2</td><td>447</td><td>6</td><td>700</td></tr><tr><td>3</td><td>442</td><td>6</td><td>520</td></tr></tbody></table>

## Numerical modelling results

Definition of the FE model. To verify the model predictions, they are compared with numerical simulation predictions employing the software Qform UK. It is a special-purpose software for the extrusion simulation based on the Lagrange–Euler approach. This software provides material flow analysis coupled with the mechanical problem in the tooling set, considering shape changes produced by the die deformation on the material flow through the die, which helps to ensure high accuracy of the numerical results [12].

All meshes are created with tetrahedral elements. A highly adaptive meshing is applied in order to optimise mesh quality and minimise the resources required to achieve accurate calculation