<figure><img src="image_159.png" /><figcaption>Fig. 7: A graphical representation of how the particle size is affected by the process parameters. The content of SiC particles in AMC powder is dependent on (a) the heat source, and when utilizing induction, is dependent on (b) pressure difference, (c) nozzle size, and (d) frequency.</figcaption></figure>

The main reason for the deviating composite structure achieved by arc and plasma melting is the dissolution of SiC particles due to the high temperature of both heat sources during the melting of the feed material [10]. In the case of induction melting, the low content of SiC particles in the AMC powder after the induction melting process may result from the melting temperature, as SiC can react with molten Al alloy at temperatures ranging from 667°C to 1347°C [10]. However, it is presumed that the more likely reason for the low SiC content in the AMC powder is process-related effects, such as (1) insufficient material flow leading to segregation or (2) excessively fast material flow resulting in unprocessed material due to nozzle clogging.

<figure><img src="image_160.png" /><figcaption>Fig. 8: A graphical representation of how the content of SiC in powder matrix is affected by the process parameters. The average size of the AMC powder is dependent on (a) the heat source, and when utilizing induction, is dependent on (b) pressure difference, (c) nozzle size, and (d) frequency.</figcaption></figure>

## Summary

In the present work, AMC powder, comprising AlSi9Mg reinforced with SiC particles, was produced employing ultrasonic atomization technology under varied process conditions. The following conclusions can be drawn:

* The efficiency of the ultrasonic process for producing AMC powder depended on the selected heat source. The highest process efficiency (almost 50%) was achieved with the induction system. Further improvement of the efficiency for this heat source required overcoming determined process challenges.

* The morphology of ultrasonically atomized powders was similar for all utilized heat sources, with an aspect ratio above 0.93, indicating high powder sphericity. In the case of induction melting, a correlation between pressure difference and aspect ratio was determined.
* The average size of AMC powder produced by all heat sources was comparable, ranging from 88.2 to 120 µm, while the high frequency in the case of induction resulted in an increased powder size of up to 170 µm.

* The content of SiC particles in the produced AMC powders was very low for all applied process conditions. In the case of arc and plasma melting, the likely reason for this effect is the dissolution of the SiC particles due to the high temperature of both heat sources. In contrast, the low SiC content in the AMC powders resulted from a process challenge, including the compensation of material flow to avoid segregations of SiC from the aluminum alloy and the lack of material processing due to nozzle clogging.

Ultrasonic atomization with induction melting has the greatest potential for processing AMC powder due to its low temperature, which prevents the dissolution of SiC particles. An increase in process efficiency is possible with more trials and process optimization. However, achieving a composite structure through induction requires further development of the current setup, for example, by introducing a stirring device that could help in mixing both components.

## References

[1] P. Garg, A. Jamwal, D. Kumar, K. K. Sadasivuni, C. M. Hussain, P. Gupta, Advance research progresses in aluminium matrix composites: Manufacturing & applications. Journal of Materials Research and Technology, 2019, 8(5):4924-39. https://doi.org/10.1016/j.jmrt.2019.06.028

[2] C. Y. Liu, Q. Wang, Y. Z. Jia, B. Zhang, R. Jing, M. Z. Ma, Q. Jing, R. P. Liu, Effect of W particles on the properties of accumulatively roll-bonded Al/W composites, Mater. Sci. Eng. A 547, 2012, 120e124. https://doi.org/10.1016/j.msea.2012.03.095

[3] S. Kobayashi, N. Hosodaa, R. Takashimab, Tungsten alloys as radiation protection materials, Nucl. Instrum. Methods Phys. Res. 390, 1997, 426e430.

[4] V. V. Cherdyntsev, M. V. Gorshenkov, V. D. Danilov, S. D. Kaloshkin, V. N. Gul'bin, Metal-matrix radiation-protective composite materials based ON aluminum, Met. Sci. Heat Treat. 55, 2013, 14e18. https://doi.org/10.1007/s11041-013-9571-2

[5] X. Xi, B. Chen, C. Tan, X. Song, J. Feng. Microstructure and mechanical properties of SiC reinforced AlSi10Mg composites fabricated by laser metal deposition. Journal of Manufacturing Processes, 2020, 58(17):763-74. https://doi.org/10.1016/j.jmapro.2020.08.073

[6] https://www.amazemet.com/repowder

[7] A. Jedynak. G. Gökhan, A. Neumann, P. Robert, S. Härtel. Semi-finished powder of aluminum matrix composite for a direct energy deposition additive manufacturing. Material Research Proceedings 28, 2023, 199-206

[8] M. Graf. New casting process for the production of AMCs with high volume content of reinforced particle phase. Proceedings 5th IMTC 2021.

[9] J. Sun, H. Yu, D. Zeng, P. Shen. Wire-powder-arc additive manufacturing: A viable strategy to fabricate carbide ceramic/aluminum alloy multi-material structures, Additive Manufacturing, Volume 51, 2022, 102637, ISSN 2214-8604, https://doi.org/10.1016/j.addma.2022.102637

[10] J. Chen, R. Zhang, B.S. Amirkhiz, H. Gu, Synthesis of In Situ SiC/Graphite/Al Hybrid Composite Coating by Laser Direct Energy Deposition, Metall. Mater. Trans. A 53 (2022) 484-502. https://doi.org/10.1007/s11661-021-06508-x
* The average size of AMC powder produced by all heat sources was comparable, ranging from 88.2 to 120 µm, while the high frequency in the case of induction resulted in an increased powder size of up to 170 µm.

* The content of SiC particles in the produced AMC powders was very low for all applied process conditions. In the case of arc and plasma melting, the likely reason for this effect is the dissolution of the SiC particles due to the high temperature of both heat sources. In contrast, the low SiC content in the AMC powders resulted from a process challenge, including the compensation of material flow to avoid segregations of SiC from the aluminum alloy and the lack of material processing due to nozzle clogging.

Ultrasonic atomization with induction melting has the greatest potential for processing AMC powder due to its low temperature, which prevents the dissolution of SiC particles. An increase in process efficiency is possible with more trials and process optimization. However, achieving a composite structure through induction requires further development of the current setup, for example, by introducing a stirring device that could help in mixing both components.

## References

[1] P. Garg, A. Jamwal, D. Kumar, K. K. Sadasivuni, C. M. Hussain, P. Gupta, Advance research progresses in aluminium matrix composites: Manufacturing & applications. Journal of Materials Research and Technology, 2019, 8(5):4924-39. https://doi.org/10.1016/j.jmrt.2019.06.028

[2] C. Y. Liu, Q. Wang, Y. Z. Jia, B. Zhang, R. Jing, M. Z. Ma, Q. Jing, R. P. Liu, Effect of W particles on the properties of accumulatively roll-bonded Al/W composites, Mater. Sci. Eng. A 547, 2012, 120e124. https://doi.org/10.1016/j.msea.2012.03.095

[3] S. Kobayashi, N. Hosodaa, R. Takashimab, Tungsten alloys as radiation protection materials, Nucl. Instrum. Methods Phys. Res. 390, 1997, 426e430.

[4] V. V. Cherdyntsev, M. V. Gorshenkov, V. D. Danilov, S. D. Kaloshkin, V. N. Gul'bin, Metal-matrix radiation-protective composite materials based ON aluminum, Met. Sci. Heat Treat. 55, 2013, 14e18. https://doi.org/10.1007/s11041-013-9571-2

[5] X. Xi, B. Chen, C. Tan, X. Song, J. Feng. Microstructure and mechanical properties of SiC reinforced AlSi10Mg composites fabricated by laser metal deposition. Journal of Manufacturing Processes, 2020, 58(17):763-74. https://doi.org/10.1016/j.jmapro.2020.08.073

[6] https://www.amazemet.com/repowder

[7] A. Jedynak. G. Gökhan, A. Neumann, P. Robert, S. Härtel. Semi-finished powder of aluminum matrix composite for a direct energy deposition additive manufacturing. Material Research Proceedings 28, 2023, 199-206

[8] M. Graf. New casting process for the production of AMCs with high volume content of reinforced particle phase. Proceedings 5th IMTC 2021.

[9] J. Sun, H. Yu, D. Zeng, P. Shen. Wire-powder-arc additive manufacturing: A viable strategy to fabricate carbide ceramic/aluminum alloy multi-material structures, Additive Manufacturing, Volume 51, 2022, 102637, ISSN 2214-8604, https://doi.org/10.1016/j.addma.2022.102637

[10] J. Chen, R. Zhang, B.S. Amirkhiz, H. Gu, Synthesis of In Situ SiC/Graphite/Al Hybrid Composite Coating by Laser Direct Energy Deposition, Metall. Mater. Trans. A 53 (2022) 484-502. https://doi.org/10.1007/s11661-021-06508-x
# Assessment of wire arc additive manufacturing with respect to the repeatability of the process under uncertainties

PULICKAN Shyam<fnref n="1" />,ᵃ*, LAFON Pascal<fnref n="1" />,ᵇ, LANGLOIS Laurent<fnref n="2" />,ᶜ, CHEVRET Sandra<fnref n="2" />,ᵈ

**Keywords:** WAAM-CMT, Additive Manufacturing, Uncertainties, Process Parameters, Repeatability

**Abstract.** Energy Deposition Methods (EDM) is on the rise with its capabilities to manufacture relatively big parts, and Wire-Arc Additive Manufacturing (WAAM) has been on the forefront. Developments in this technology has led to the innovation of WAAM-Cold Metal Transfer (WAAM-CMT), which is much more efficient than its predecessors. Nevertheless, it comes with its own complexities, giving rise to uncertainties, eventually causing variations in the geometry of fabricated parts. But these uncertainties can be traced back to the input parameters involved in the process. Even though, WAAM is influenced by several phenomena and parameters, only a handful of parameters are practically controlled during the process. Among these the most important parameters identified are Wire Feed Rate (Vfeed) and Travel Speed (Vt) of the robot, which have a direct influence on the parts fabricated. To quantify the variability of these process parameters and the geometry of the bead, 140 single-layered beads are fabricated with four different set of input parameters. Quantification of the variations and the repeatability aspect of these variables, and thereby the process itself, are then studied using statistical tools like ANOVA. This gives an idea on how these vary from bead to bead. Quantifying these uncertainties and understanding the variations would help improve the process. This would enable better control of the process parameters, thus helping to make better design of the process and better predictions about the variations to expect.

**Introduction**

Wire Arc Additive Manufacturing (WAAM) is a process which uses the concept of adding layer upon layer, of material to build structures, contrary to Gas Metal Arc Welding (GMAW) which deposits molten material just for the purpose of joining two parts. This characteristic upgrades a simple metal joining technique into an additive manufacturing process. Or in other words, WAAM is an upgraded version of GMAW which comes from the class of Direct Energy Deposition (DED) processes [1]. Therefore, it is a direct metal feeding process which makes use of an electric arc as heat source to melt the material which is the feedstock [2,3]. Owing to its high deposition rate, ability to manufacture and/or repair structures, and manufacturing flexibility WAAM has received attention, relative to other conventional production techniques [1,4,5].

The process involves rapid heating, melting, and cooling of metal which determines the metallurgical properties and geometry of the parts. Different physical phenomena like the flow of metal in the weld pool, the influence of gravity, the rate of cooling etc. have a direct influence on the mechanical and geometric characteristics. But during the process, phenomena like distortion, squash down of metal due to gravity, surface undulations, inaccuracies with arc force etc. plays an important role and induces uncertainties in the process [2].

<figure><img src="image_163.png" /></figure>

<footnote n="1">¹ LASMIS, Université de Technologie de Troyes, Troyes, France</footnote>

<footnote n="2">² LCFC, École Nationale Supérieure d'Arts et Métiers, Metz, France</footnote>

<footnote>ᵃ shyam_george.pulickan@utt.fr, ᵇ pascal.lafon@utt.fr, ᶜ laurent.langlois@ensam.eu,</footnote>

<footnote>ᵈ sandra.chevret@ensam.eu</footnote>
# Assessment of wire arc additive manufacturing with respect to the repeatability of the process under uncertainties

PULICKAN Shyam<fnref n="1" />,ᵃ*, LAFON Pascal<fnref n="1" />,ᵇ, LANGLOIS Laurent<fnref n="2" />,ᶜ, CHEVRET Sandra<fnref n="2" />,ᵈ

**Keywords:** WAAM-CMT, Additive Manufacturing, Uncertainties, Process Parameters, Repeatability

**Abstract.** Energy Deposition Methods (EDM) is on the rise with its capabilities to manufacture relatively big parts, and Wire-Arc Additive Manufacturing (WAAM) has been on the forefront. Developments in this technology has led to the innovation of WAAM-Cold Metal Transfer (WAAM-CMT), which is much more efficient than its predecessors. Nevertheless, it comes with its own complexities, giving rise to uncertainties, eventually causing variations in the geometry of fabricated parts. But these uncertainties can be traced back to the input parameters involved in the process. Even though, WAAM is influenced by several phenomena and parameters, only a handful of parameters are practically controlled during the process. Among these the most important parameters identified are Wire Feed Rate (Vfeed) and Travel Speed (Vt) of the robot, which have a direct influence on the parts fabricated. To quantify the variability of these process parameters and the geometry of the bead, 140 single-layered beads are fabricated with four different set of input parameters. Quantification of the variations and the repeatability aspect of these variables, and thereby the process itself, are then studied using statistical tools like ANOVA. This gives an idea on how these vary from bead to bead. Quantifying these uncertainties and understanding the variations would help improve the process. This would enable better control of the process parameters, thus helping to make better design of the process and better predictions about the variations to expect.

**Introduction**

Wire Arc Additive Manufacturing (WAAM) is a process which uses the concept of adding layer upon layer, of material to build structures, contrary to Gas Metal Arc Welding (GMAW) which deposits molten material just for the purpose of joining two parts. This characteristic upgrades a simple metal joining technique into an additive manufacturing process. Or in other words, WAAM is an upgraded version of GMAW which comes from the class of Direct Energy Deposition (DED) processes [1]. Therefore, it is a direct metal feeding process which makes use of an electric arc as heat source to melt the material which is the feedstock [2,3]. Owing to its high deposition rate, ability to manufacture and/or repair structures, and manufacturing flexibility WAAM has received attention, relative to other conventional production techniques [1,4,5].

The process involves rapid heating, melting, and cooling of metal which determines the metallurgical properties and geometry of the parts. Different physical phenomena like the flow of metal in the weld pool, the influence of gravity, the rate of cooling etc. have a direct influence on the mechanical and geometric characteristics. But during the process, phenomena like distortion, squash down of metal due to gravity, surface undulations, inaccuracies with arc force etc. plays an important role and induces uncertainties in the process [2].

<figure><img src="image_161.png" /></figure>

<footnote n="1">¹ LASMIS, Université de Technologie de Troyes, Troyes, France</footnote>

<footnote n="2">² LCFC, École Nationale Supérieure d'Arts et Métiers, Metz, France</footnote>

<footnote>ᵃ shyam_george.pulickan@utt.fr, ᵇ pascal.lafon@utt.fr, ᶜ laurent.langlois@ensam.eu,</footnote>

<footnote>ᵈ sandra.chevret@ensam.eu</footnote>
Several overlapping beads forms a layer and several such layers forms produces complete parts [6,7]. To reduce the various fabrication issues during the process, different WAAM methods were devised depending on the necessity. Among the various innovations, the integration of Cold Metal Transfer (CMT) with WAAM is gaining popularity which provides a controlled deposition mode through its advanced wire feed and digital control system. The higher efficiency, lower heat input and lower cost makes it more feasible relative to other WAAM processes [1,2,6,8-10].

The quality of the structure depends on the weld bead geometry and dilution, which are controlled by the process parameters like Travel Speed, Wire Feed Rate, Area of Overlap, Wire diameter, Torch Angle, Temperature etc.[11]. Therefore, studying and constructing a hierarchy of parameters of WAAM is difficult as several parameters are involved and these are inter-dependent on each other, making WAAM-CMT a complex process [1,12].

As the process is aiming higher, to be established as a standard manufacturing process, it is therefore important for the engineers to be able to predict the geometric as well as mechanical characteristics of the structure. This can only be attained through understanding the process parameters, and the related variations. This understanding of the effects of the uncertainties arising during the process and thereby optimising these parameters are necessary to obtain the desired part geometries.

The present research, therefore, characterises the variation in the process by using an observation-based approach to quantify the uncertainties present in the input parameters and output geometry of the beads. Several experiments are done fabricating simple beads, and the variations in the input parameters relative to the initially set values are recorded. The beads are then scanned using a laser scanner to study the variation in its Height and Width. This basic understanding using a simple bead can be used to create models capable of predicting parameters for much more complex parts. This work is, to the best of our knowledge, the first modelling approach of WAAM-CMT from a statistical point of view, to quantify the variations due to uncertainties. Quantifying the repeatability of these parameters (*Section: Repeatability of the Process*) is therefore the first stage in developing a model, since it would provide an insight into the input and output parameters which are assumed to be constant, if not for the influence of uncertainties.

## Process Parameters

The number of parameters influencing the process is numerous and could pose difficulties while trying to optimize the process, because of the influence of one parameter on the other. The principal parameters which influence the WAAM-CMT includes[8,9,13]: -

a) Wire Feed Rate ($V_{feed}$), which is the speed at which the wire is pulled from the bobbin.

b) Travel Speed ($V_T$), which is the speed at which the torch is displaced by the robot.

c) Torch Angle ($T_a$), which the inclination of the welding torch with respect to the surface of the substrate

d) Stick-Out Distance (SoD), which is the length of the wire, protruding from the tip of the torch.

e) Wire Diameter ($d_w$), which is the diameter of the wire used for the fabrication.

Other factors like the material used, shielding gas and its proportion (82% Argon + 18% Carbon Dioxide, for the current experiments), temperature of the surface to which the bead is going to be laid, Flatness of the surface of the substrate used etc. could also have passive influence on the fabrication.

Nevertheless, in the current scope of experiments $T_a$ & $d_w$ stays constant at 90° and 1.2 mm respectively. SoD is set to 13 mm but modified and controlled by the robot during the experiments to maintain a stabilised arc. Therefore, SoD continuously varies and is uncontrollable during fabrication. The welding wire used is steel with welding standards ER100.
Several overlapping beads forms a layer and several such layers forms produces complete parts [6,7]. To reduce the various fabrication issues during the process, different WAAM methods were devised depending on the necessity. Among the various innovations, the integration of Cold Metal Transfer (CMT) with WAAM is gaining popularity which provides a controlled deposition mode through its advanced wire feed and digital control system. The higher efficiency, lower heat input and lower cost makes it more feasible relative to other WAAM processes [1,2,6,8-10].

The quality of the structure depends on the weld bead geometry and dilution, which are controlled by the process parameters like Travel Speed, Wire Feed Rate, Area of Overlap, Wire diameter, Torch Angle, Temperature etc.[11]. Therefore, studying and constructing a hierarchy of parameters of WAAM is difficult as several parameters are involved and these are inter-dependent on each other, making WAAM-CMT a complex process [1,12].

As the process is aiming higher, to be established as a standard manufacturing process, it is therefore important for the engineers to be able to predict the geometric as well as mechanical characteristics of the structure. This can only be attained through understanding the process parameters, and the related variations. This understanding of the effects of the uncertainties arising during the process and thereby optimising these parameters are necessary to obtain the desired part geometries.

The present research, therefore, characterises the variation in the process by using an observation-based approach to quantify the uncertainties present in the input parameters and output geometry of the beads. Several experiments are done fabricating simple beads, and the variations in the input parameters relative to the initially set values are recorded. The beads are then scanned using a laser scanner to study the variation in its Height and Width. This basic understanding using a simple bead can be used to create models capable of predicting parameters for much more complex parts. This work is, to the best of our knowledge, the first modelling approach of WAAM-CMT from a statistical point of view, to quantify the variations due to uncertainties. Quantifying the repeatability of these parameters (*Section: Repeatability of the Process*) is therefore the first stage in developing a model, since it would provide an insight into the input and output parameters which are assumed to be constant, if not for the influence of uncertainties.

## Process Parameters

The number of parameters influencing the process is numerous and could pose difficulties while trying to optimize the process, because of the influence of one parameter on the other. The principal parameters which influence the WAAM-CMT includes[8,9,13]: -

a) Wire Feed Rate ($V_{feed}$), which is the speed at which the wire is pulled from the bobbin.

b) Travel Speed ($V_T$), which is the speed at which the torch is displaced by the robot.

c) Torch Angle ($T_a$), which the inclination of the welding torch with respect to the surface of the substrate

d) Stick-Out Distance (SoD), which is the length of the wire, protruding from the tip of the torch.

e) Wire Diameter ($d_w$), which is the diameter of the wire used for the fabrication.

Other factors like the material used, shielding gas and its proportion (82% Argon + 18% Carbon Dioxide, for the current experiments), temperature of the surface to which the bead is going to be laid, Flatness of the surface of the substrate used etc. could also have passive influence on the fabrication.

Nevertheless, in the current scope of experiments $T_a$ & $d_w$ stays constant at 90° and 1.2 mm respectively. SoD is set to 13 mm but modified and controlled by the robot during the experiments to maintain a stabilised arc. Therefore, SoD continuously varies and is uncontrollable during fabrication. The welding wire used is steel with welding standards ER100.
This leave $V_{feed}$ & $V_T$ as the only two variables which can be controlled during the scope of present research. These two parameters play the most important role in determining the shape and size of the beads. The experiments done showed that a faster $V_T$ combined with a low $V_{feed}$ would yield in a bead which is thin and short, whereas a slow $V_T$ with high $V_{feed}$ would result in a bulkier bead. Thus, any combination of these parameters would result in a bead which is different in its geometric characteristics. Since the form of beads is primarily influenced by $V_{feed}$ & $V_T$, the uncertainties in these parameters could also have a direct influence on the form of the bead.

## Experiments

The robot and welding generator used, permits to layup the beads with a wide range of initial values, but these does not always result in a good quality of beads with good structural integrity. A study on the quality of the beads as a function of $V_{feed}$ & $V_T$, was carried out in [14], in which an acceptable domain of the functionality of the process was recommended. Four combinations of $V_{feed}$ & $V_T$ were chosen from this domain at its boundaries, thus being able to capture the variations at the limits of the process (Table 1). The parameters in Table 1 are the initial input values for the fabrication, but their true values, and variations remains uncertain (Section: Data Analysis).

<table><caption>Table 1: Set of Input Parameters used for the Experiments.</caption><thead><tr><td>Set No.</td><td>V<sub>feed</sub> (m/min)</td><td>V<sub>T</sub> (mm/s)</td></tr></thead><tbody><tr><td>P1</td><td>9</td><td>13.3</td></tr><tr><td>P2</td><td>7.5</td><td>18.3</td></tr><tr><td>P3</td><td>10</td><td>8.3</td></tr><tr><td>P4</td><td>10</td><td>18.3</td></tr></tbody></table>

The number of beads to be fabricated per set of parameters was determined based on the central limit theorem. According to the probability theorem, it states that the distribution of the means of a random sample could follow a normal distribution as the sample size reaches 30. This becomes relevant for the statistical analysis and hypothesis test discussed on later sections.

Therefore, a total of 35 beads of length 150 mm were fabricated per set of parameters. This was to avoid any misdone fabrication process, which could arise due to any other issues, including human errors, that could disrupt the statistical analysis. Thus, a total of 140 beads were fabricated.

The fabrication process was guided using an ABB IRB 8700, a 6 axes Robot, and the welding was controlled by a Fronius TP-180i generator. The commands for the Robot and the Post were programmed using the software RobotStudio, in which the axis of the substrate and the dimensions of the substrate was specified. The substrate onto which the beads were to be laid was cut from a steel plate of thickness 20mm.

**Fabrication.** The sand-blasted, smooth, and clean surface of the substrate is clamped onto the workbench. The experiments are done in such an order that the set of parameters for each bead were chosen at random. This avoids any uncertainty which arise because of the repeated fabrication using the same set of parameters.

The metallic wire is pulled from the bobbin, pass through the torch, and is guided by the robot, following a predetermined trajectory and speed ($V_T$). The wire, as it comes into contact with the metallic surface of the substrate, completes an electrical circuit. This creates an arc at the point of contact which melts the wire, and a drop of metal is laid onto the surface of the substrate. Now as the circuit is broken, the wire is pulled back-up, and the robot moves a step ahead. This process continues until the required length of bead is laid [Fig. 1].
This leave $V_{feed}$ & $V_T$ as the only two variables which can be controlled during the scope of present research. These two parameters play the most important role in determining the shape and size of the beads. The experiments done showed that a faster $V_T$ combined with a low $V_{feed}$ would yield in a bead which is thin and short, whereas a slow $V_T$ with high $V_{feed}$ would result in a bulkier bead. Thus, any combination of these parameters would result in a bead which is different in its geometric characteristics. Since the form of beads is primarily influenced by $V_{feed}$ & $V_T$, the uncertainties in these parameters could also have a direct influence on the form of the bead.

## Experiments

The robot and welding generator used, permits to layup the beads with a wide range of initial values, but these does not always result in a good quality of beads with good structural integrity. A study on the quality of the beads as a function of $V_{feed}$ & $V_T$, was carried out in [14], in which an acceptable domain of the functionality of the process was recommended. Four combinations of $V_{feed}$ & $V_T$ were chosen from this domain at its boundaries, thus being able to capture the variations at the limits of the process (Table 1). The parameters in Table 1 are the initial input values for the fabrication, but their true values, and variations remains uncertain (Section: Data Analysis).

<table><caption>Table 1: Set of Input Parameters used for the Experiments.</caption><thead><tr><td>Set No.</td><td>V<sub>feed</sub> (m/min)</td><td>V<sub>T</sub> (mm/s)</td></tr></thead><tbody><tr><td>P1</td><td>9</td><td>13.3</td></tr><tr><td>P2</td><td>7.5</td><td>18.3</td></tr><tr><td>P3</td><td>10</td><td>8.3</td></tr><tr><td>P4</td><td>10</td><td>18.3</td></tr></tbody></table>

The number of beads to be fabricated per set of parameters was determined based on the central limit theorem. According to the probability theorem, it states that the distribution of the means of a random sample could follow a normal distribution as the sample size reaches 30. This becomes relevant for the statistical analysis and hypothesis test discussed on later sections.

Therefore, a total of 35 beads of length 150 mm were fabricated per set of parameters. This was to avoid any misdone fabrication process, which could arise due to any other issues, including human errors, that could disrupt the statistical analysis. Thus, a total of 140 beads were fabricated.

The fabrication process was guided using an ABB IRB 8700, a 6 axes Robot, and the welding was controlled by a Fronius TP-180i generator. The commands for the Robot and the Post were programmed using the software RobotStudio, in which the axis of the substrate and the dimensions of the substrate was specified. The substrate onto which the beads were to be laid was cut from a steel plate of thickness 20mm.

**Fabrication.** The sand-blasted, smooth, and clean surface of the substrate is clamped onto the workbench. The experiments are done in such an order that the set of parameters for each bead were chosen at random. This avoids any uncertainty which arise because of the repeated fabrication using the same set of parameters.

The metallic wire is pulled from the bobbin, pass through the torch, and is guided by the robot, following a predetermined trajectory and speed ($V_T$). The wire, as it comes into contact with the metallic surface of the substrate, completes an electrical circuit. This creates an arc at the point of contact which melts the wire, and a drop of metal is laid onto the surface of the substrate. Now as the circuit is broken, the wire is pulled back-up, and the robot moves a step ahead. This process continues until the required length of bead is laid [Fig. 1].
The robot, to maintain a stabilised and constant arc length, controls the stand-off distance by controlling the $V_{feed}$ of the wire, which induces variations in the parameter. Even though this ensures an optimised fabrication, the change in $V_{feed}$ affects the geometry of the bead. The Fronius Station records this real time $V_{feed}$ during the fabrication.

<figure><img src="image_164.png" /><figcaption>Figure 1: (a) WAAM-CMT Process (Laser tracker reflector visible on the left side) (b) Laser scanner used for the capturing the bead profile</figcaption></figure>

The real time coordinates of the robot were tracked using a laser tracker, using which the $V_T$ was later computed. Once the fabrication is completed the surface of the bead was cleared-off any silicates formed [Fig. 2] due to the presence of silicon in the steel wire which reacts with Oxygen. This is important as these could become inclusions while fabricating more complex structures. Once the bead is cooled down, the profile along the length of the bead was scanned using a laser scanner Keyence LJ-V7080 [Fig. 1] which captured the bead as a point cloud using which the various geometrical aspects of the bead could be studied. Since the scan using light to capture the bead geometry, the clearing-off of the silicates becomes important, as the silicates could disrupt the precision of the scan due to refraction of the light from the scanner.

<figure><img src="image_165.png" /><figcaption>Figure 2: Welding defects: Bump, Dent and Silicates formed during the fabrication The splatter can also be seen around the bead as small dots</figcaption></figure>

The process also brings with it, anomalies in the fabricated beads at the beginning and end of the bead (a bump and a dent respectively) because of the way the welding generator controls the process [Fig. 2]. These are undesirable parts of the bead and does not correspond to the bead geometry expected while using a specific set of parameters. Therefore, while studying the uncertainties linked with a given set of input parameters, these parts are removed from the analysis, which accounts to 50mm of the 150mm fabricated, i.e., 25mm from either side of the bead.

## Data Analysis and Results

*Wire Feed Rate.* The Fronius Station records the $V_{feed}$ with a resolution of one entry per 0.1 seconds.
The robot, to maintain a stabilised and constant arc length, controls the stand-off distance by controlling the $V_{feed}$ of the wire, which induces variations in the parameter. Even though this ensures an optimised fabrication, the change in $V_{feed}$ affects the geometry of the bead. The Fronius Station records this real time $V_{feed}$ during the fabrication.

<figure><img src="image_162.png" /><figcaption>Figure 1: (a) WAAM-CMT Process (Laser tracker reflector visible on the left side) (b) Laser scanner used for the capturing the bead profile</figcaption></figure>

The real time coordinates of the robot were tracked using a laser tracker, using which the $V_T$ was later computed. Once the fabrication is completed the surface of the bead was cleared-off any silicates formed [Fig. 2] due to the presence of silicon in the steel wire which reacts with Oxygen. This is important as these could become inclusions while fabricating more complex structures. Once the bead is cooled down, the profile along the length of the bead was scanned using a laser scanner Keyence LJ-V7080 [Fig. 1] which captured the bead as a point cloud using which the various geometrical aspects of the bead could be studied. Since the scan using light to capture the bead geometry, the clearing-off of the silicates becomes important, as the silicates could disrupt the precision of the scan due to refraction of the light from the scanner.

<figure><img src="image_163.png" /><figcaption>Figure 2: Welding defects: Bump, Dent and Silicates formed during the fabrication The splatter can also be seen around the bead as small dots</figcaption></figure>

The process also brings with it, anomalies in the fabricated beads at the beginning and end of the bead (a bump and a dent respectively) because of the way the welding generator controls the process [Fig. 2]. These are undesirable parts of the bead and does not correspond to the bead geometry expected while using a specific set of parameters. Therefore, while studying the uncertainties linked with a given set of input parameters, these parts are removed from the analysis, which accounts to 50mm of the 150mm fabricated, i.e., 25mm from either side of the bead.

## Data Analysis and Results

*Wire Feed Rate.* The Fronius Station records the $V_{feed}$ with a resolution of one entry per 0.1 seconds.
The noise in the measurement, or the outliers, (not to be confused with the uncertainties during the fabrication) which are errors, are removed using the Inter-Quartile Range (IQR). The analysis following showed that: -

* All parameters take a few milliseconds to reach the specified value, but Parameter 3 takes more time than the others to get stabilised at a constant level as seen from the slope at the first two seconds [Fig. 3a].

* Even at the identical $V_{feed}$, P3 shows higher variation than P4, illustrating the influence of $V_T$ on $V_{feed}$ [Fig. 3a].

* All the actual fabrication $V_{feed}$ falls below their set value, except for which P2 oscillates around its set value [Fig. 3a].

* The statistical analysis shows that, even though fabricated at a specified $V_{feed}$, practically the 35 beads are fabricated at different and fluctuating $V_{feed}$ [Fig. 3b]. This could be very important while creating prediction models as, the beads cannot be assumed to be fabricated at the specified $V_{feed}$.

<figure><img src="image_166.png" /><figcaption>Figure 3: (a) Variation in $V_{feed}$ for one bead fabricated with each Parameter and (b) The statistical distribution of 35 beads fabricated with Parameter 2</figcaption></figure>

Travel Speed. $V_T$ is not a direct output like the $V_{feed}$. The position of a tracker attached to the robot is recorded using a laser tracker, which gives its the real-time coordinates [Fig. 1a]. Each output is recorded at a frequency of acquisition of 96ms. Using the coordinates, the travel speed components are calculated since the time between two measurements are known. The coordinates are smoothed to accommodate the resolution error of ±0.3 µm of the laser tracker.

Then with all the components, $V_T$ is calculated as the norm of the speed components: -

$$V_T = \sqrt{v_x^2 + v_y^2 + v_z^2} \qquad (1)$$

The analysis following showed that: -

* All the recorded $V_T$ respect the set value with very small variations [Fig. 4a].

* There is not enough evidence to say that $V_T$ is depended on $V_{feed}$, since P2 and P4 are superposed at the set value of 18.3 mm/s [Fig. 4a].

* The statistical analysis showed that, the beads were fabricated more or less at the set value of $V_T$ which gives confidence on this parameter.
The noise in the measurement, or the outliers, (not to be confused with the uncertainties during the fabrication) which are errors, are removed using the Inter-Quartile Range (IQR). The analysis following showed that: -

* All parameters take a few milliseconds to reach the specified value, but Parameter 3 takes more time than the others to get stabilised at a constant level as seen from the slope at the first two seconds [Fig. 3a].

* Even at the identical $V_{feed}$, P3 shows higher variation than P4, illustrating the influence of $V_T$ on $V_{feed}$ [Fig. 3a].

* All the actual fabrication $V_{feed}$ falls below their set value, except for which P2 oscillates around its set value [Fig. 3a].

* The statistical analysis shows that, even though fabricated at a specified $V_{feed}$, practically the 35 beads are fabricated at different and fluctuating $V_{feed}$ [Fig. 3b]. This could be very important while creating prediction models as, the beads cannot be assumed to be fabricated at the specified $V_{feed}$.

<figure><img src="image_164.png" /><figcaption>Figure 3: (a) Variation in $V_{feed}$ for one bead fabricated with each Parameter and (b) The statistical distribution of 35 beads fabricated with Parameter 2</figcaption></figure>

Travel Speed. $V_T$ is not a direct output like the $V_{feed}$. The position of a tracker attached to the robot is recorded using a laser tracker, which gives its the real-time coordinates [Fig. 1a]. Each output is recorded at a frequency of acquisition of 96ms. Using the coordinates, the travel speed components are calculated since the time between two measurements are known. The coordinates are smoothed to accommodate the resolution error of ±0.3 µm of the laser tracker.

Then with all the components, $V_T$ is calculated as the norm of the speed components: -

$$V_T = \sqrt{v_x^2 + v_y^2 + v_z^2} \qquad (1)$$

The analysis following showed that: -

* All the recorded $V_T$ respect the set value with very small variations [Fig. 4a].

* There is not enough evidence to say that $V_T$ is depended on $V_{feed}$, since P2 and P4 are superposed at the set value of 18.3 mm/s [Fig. 4a].

* The statistical analysis showed that, the beads were fabricated more or less at the set value of $V_T$ which gives confidence on this parameter.
<figure><img src="image_167.png" /><figcaption>Figure 4: (a) Variation in $V_T$ for one bead fabricated with each Parameter and (b) The statistical distribution of 13 beads fabricated with Parameter 2</figcaption></figure>

*Bead Geometry.* The bead geometry is characterized by a set of cross-sections obtained by the laser scanner. For each cross-section, elementary shape parameters can be defined, such as height, the position of the cross-section's feet and its apex along the width etc [Fig.5]. The results shown in Figure 6 & 7 illustrates the variation in the principal geometric variables, height, and width of the bead.

The Laser scanner used for scanning the bead had a good resolution of 0.05mm, which would give around 2000 sections of useful sections per bead (100 mm), after filtering off the defects. Each of these sections contains 800 points corresponding to 40mm of cross-section of bead. The section of the bead profile is later extracted from the point cloud excluding the substrate [Fig. 5].

<figure><img src="image_168.png" /><figcaption>Figure 5: Plot showing four sections of an extracted bead, clearly showing the variations in its form along its length. Both the height and width can be seen to be fluctuating</figcaption></figure>

As expected, beads fabricated at lower $V_T$ and higher $V_{feed}$ have bigger profiles compared to the ones which had faster $V_T$ and lower $V_{feed}$.

It can also be noted that the variation in Parameter 3 is more prominent than the others, possibly because of the bigger drop size, higher volume of melt pool, higher cool down time etc. This variation could have an adverse effect on the profile of the bead, with change in geometry from section to section [Fig. 6 & Fig. 7].
<figure><img src="image_165.png" /><figcaption>Figure 4: (a) Variation in $V_T$ for one bead fabricated with each Parameter and (b) The statistical distribution of 13 beads fabricated with Parameter 2</figcaption></figure>

*Bead Geometry.* The bead geometry is characterized by a set of cross-sections obtained by the laser scanner. For each cross-section, elementary shape parameters can be defined, such as height, the position of the cross-section's feet and its apex along the width etc [Fig. 5]. The results shown in Figure 6 & 7 illustrates the variation in the principal geometric variables, height, and width of the bead.

The Laser scanner used for scanning the bead had a good resolution of 0.05mm, which would give around 2000 sections of useful sections per bead (100 mm), after filtering off the defects. Each of these sections contains 800 points corresponding to 40mm of cross-section of bead. The section of the bead profile is later extracted from the point cloud excluding the substrate [Fig. 5].

<figure><img src="image_166.png" /><figcaption>Figure 5: Plot showing four sections of an extracted bead, clearly showing the variations in its form along its length. Both the height and width can be seen to be fluctuating</figcaption></figure>

As expected, beads fabricated at lower $V_T$ and higher $V_{feed}$ have bigger profiles compared to the ones which had faster $V_T$ and lower $V_{feed}$.

It can also be noted that the variation in Parameter 3 is more prominent than the others, possibly because of the bigger drop size, higher volume of melt pool, higher cool down time etc. This variation could have an adverse effect on the profile of the bead, with change in geometry from section to section [Fig. 6 & Fig. 7].
<figure><img src="image_169.png" /><figcaption>Figure 6: (a) Variation in Height for one bead fabricated with each Parameter and (b) The statistical distribution of Height of 35 beads fabricated with Parameter 2</figcaption></figure>

The statistical analysis of Height and Width shows that there is no evident relation between the distributions of beads fabricated with the same set of parameters [Fig. 6b & 7b]. Rather the distributions are widespread showing that no two beads could be reproduced with a given set of parameters.

<figure><img src="image_170.png" /><figcaption>Figure 7: (a) Variation in Width for one bead fabricated with each Parameter and (b) The statistical distribution of Width of 35 beads fabricated with Parameter 2</figcaption></figure>

*Repeatability of the Process.* The general statistical inferences made are rather qualitative and it is essential to prove quantitatively if the difference between the statistical distributions of two parameters are significant or not. The Joint Committee for Guides in Metrology (JCGM) defines repeatability as "The closeness of the agreement between the results of successive measurements of the same measured carried out under the same conditions of measurement" [15].

Mathematically, it can be quantified as the standard deviation of a mean result under the repeatability conditions [16]. In the present study it can be used to identify if the difference between the distributions of several beads, fabricated using a set of parameters are considerable or not. Thus, it can be quantified if the parameters and thereby the process itself can be repeated.

The standard test used for this is the Analysis of Variance (ANOVA) Test, which numerically precise if the difference between two distributions is significant or not. The test computes the ratio between the mean square of variability of two group means (here means of two beads for the parameter under consideration), with the variation of the distribution itself. This is called F-Statistic which is compared with a Critical F-statistic obtained from the F-distribution table at a significance level of 5%.
<figure><img src="image_167.png" /><figcaption>Figure 6: (a) Variation in Height for one bead fabricated with each Parameter and (b) The statistical distribution of Height of 35 beads fabricated with Parameter 2</figcaption></figure>

The statistical analysis of Height and Width shows that there is no evident relation between the distributions of beads fabricated with the same set of parameters [Fig. 6b & 7b]. Rather the distributions are widespread showing that no two beads could be reproduced with a given set of parameters.

<figure><img src="image_168.png" /><figcaption>Figure 7: (a) Variation in Width for one bead fabricated with each Parameter and (b) The statistical distribution of Width of 35 beads fabricated with Parameter 2</figcaption></figure>

*Repeatability of the Process.* The general statistical inferences made are rather qualitative and it is essential to prove quantitatively if the difference between the statistical distributions of two parameters are significant or not. The Joint Committee for Guides in Metrology (JCGM) defines repeatability as "The closeness of the agreement between the results of successive measurements of the same measured carried out under the same conditions of measurement" [15].

Mathematically, it can be quantified as the standard deviation of a mean result under the repeatability conditions [16]. In the present study it can be used to identify if the difference between the distributions of several beads, fabricated using a set of parameters are considerable or not. Thus, it can be quantified if the parameters and thereby the process itself can be repeated.

The standard test used for this is the Analysis of Variance (ANOVA) Test, which numerically precise if the difference between two distributions is significant or not. The test computes the ratio between the mean square of variability of two group means (here means of two beads for the parameter under consideration), with the variation of the distribution itself. This is called F-Statistic which is compared with a Critical F-statistic obtained from the F-distribution table at a significance level of 5%.
$$ 
\begin{aligned}
F_{\text{statistic}} &= f(\text{group means, overall mean, number of groups,} \\
&\quad \text{number of observations per group}) \\
&= \frac{\text{MS between the group}}{\text{MS within the group}} (2)
\end{aligned}
$$

<table><caption>Table 2. ANOVA test results of Input Parameters</caption><thead><tr><th>Parameter No.</th><th>V<sub>feed</sub> (m/min)</th><th>V<sub>T</sub> (mm/s)</th><th>F-Statistic V<sub>feed</sub></th><th>F-Critical V<sub>feed</sub></th><th>F-Statistic V<sub>T</sub></th><th>F-Critical V<sub>T</sub></th></tr></thead><tbody><tr><td>P1</td><td>9</td><td>13.3</td><td>79.90</td><td>2.32</td><td>1.83</td><td>1.59</td></tr><tr><td>P2</td><td>7.5</td><td>18.3</td><td>53.24</td><td>3.26</td><td>1.88</td><td>1.72</td></tr><tr><td>P3</td><td>10</td><td>8.3</td><td>20.63</td><td>3.24</td><td>1.81</td><td>1.51</td></tr><tr><td>P4</td><td>10</td><td>18.3</td><td>65.35</td><td>1.47</td><td>0.19</td><td>1.72</td></tr></tbody></table>

<table><caption>Table 3. ANOVA test results of Output Variables</caption><thead><tr><th>Parameter No.</th><th>V<sub>feed</sub> (m/min)</th><th>V<sub>T</sub> (mm/s)</th><th>F-Statistic Height</th><th>F-Statistic Width</th><th>F-Critical</th></tr></thead><tbody><tr><td>P1</td><td>9</td><td>13.3</td><td>1072.1</td><td>234.31</td><td>1,43</td></tr><tr><td>P2</td><td>7.5</td><td>18.3</td><td>1287.6</td><td>107.92</td><td>1.43</td></tr><tr><td>P3</td><td>10</td><td>8.3</td><td>947.1</td><td>133.40</td><td>1.44</td></tr><tr><td>P4</td><td>10</td><td>18.3</td><td>522.5</td><td>78.74</td><td>1,45</td></tr></tbody></table>

The ANOVA test returned expected results with very high values of F-Statistic for $V_{feed}$, Height and Width, implying that these have a high numerator of F-Statistic [Table 2 & 3]. Therefore, the variation from one bead to another is much more significant than the variation within a bead making these variables non-repeatable. On the other hand, $V_T$ showed promising results even though it failed ANOVA test since the ratio of numerator to the denominator of F-Statistic is very small [Table 2].

A further investigation into this data, involving bead to bead comparisons, proved that around 90% of the beads fabricated were repeatable. Only a couple of beads were going beyond the acceptable variation for $V_T$, the reason for which the ANOVA test failed. Nevertheless, it can be ascertained that the process is not repeatable, with drastic variations in geometry found even on a single bead.

## Conclusion

WAAM-CMT despite being an advanced manufacturing process used for fabrication of big structures, comes with its own defects. Today, a non-standardised factor of safety is used during the fabrication to accommodate for the bad repeatability of the process depending on the in-situ research. The present study has been able to analyse the principal process parameters and basic geometry of the bead to quantify the variation in these parameters. Important variations in the Wire Feed Rate are found in the recorded measurements, which have a direct influence on the fabricated parts. Nevertheless, the variations found in the Travel Speed of the robot were rather small. The bead profiles, and thereby, the height and the width were different from one another even on a single bead analysed.

The ANOVA test confirms the inferences that neither the input parameters nor the output geometry is repeatable. It also shows the extend of this aspect, which indicate that the Travel Speed is close to being identifiable as repeatable, with only a couple of beads going beyond the acceptable variations. Nevertheless, the Wire Feed Rate is much further from being repeatable, which could
$$ 
\begin{aligned}
F_{\text{statistic}} &= f(\text{group means, overall mean, number of groups,} \\
&\quad \text{number of observations per group}) \\
&= \frac{\text{MS between the group}}{\text{MS within the group}} (2)
\end{aligned}
 $$

<table><caption>Table 2. ANOVA test results of Input Parameters</caption><thead><tr><th>Parameter No.</th><th>V<sub>feed</sub> (m/min)</th><th>V<sub>T</sub> (mm/s)</th><th>F-Statistic V<sub>feed</sub></th><th>F-Critical V<sub>feed</sub></th><th>F-Statistic V<sub>T</sub></th><th>F-Critical V<sub>T</sub></th></tr></thead><tbody><tr><td>P1</td><td>9</td><td>13.3</td><td>79.90</td><td>2.32</td><td>1.83</td><td>1.59</td></tr><tr><td>P2</td><td>7.5</td><td>18.3</td><td>53.24</td><td>3.26</td><td>1.88</td><td>1.72</td></tr><tr><td>P3</td><td>10</td><td>8.3</td><td>20.63</td><td>3.24</td><td>1.81</td><td>1.51</td></tr><tr><td>P4</td><td>10</td><td>18.3</td><td>65.35</td><td>1.47</td><td>0.19</td><td>1.72</td></tr></tbody></table>

<table><caption>Table 3. ANOVA test results of Output Variables</caption><thead><tr><th>Parameter No.</th><th>V<sub>feed</sub> (m/min)</th><th>V<sub>T</sub> (mm/s)</th><th>F-Statistic Height</th><th>F-Statistic Width</th><th>F-Critical</th></tr></thead><tbody><tr><td>P1</td><td>9</td><td>13.3</td><td>1072.1</td><td>234.31</td><td>1,43</td></tr><tr><td>P2</td><td>7.5</td><td>18.3</td><td>1287.6</td><td>107.92</td><td>1.43</td></tr><tr><td>P3</td><td>10</td><td>8.3</td><td>947.1</td><td>133.40</td><td>1.44</td></tr><tr><td>P4</td><td>10</td><td>18.3</td><td>522.5</td><td>78.74</td><td>1,45</td></tr></tbody></table>

The ANOVA test returned expected results with very high values of F-Statistic for V<sub>feed</sub>, Height and Width, implying that these have a high numerator of F-Statistic [Table 2 & 3]. Therefore, the variation from one bead to another is much more significant than the variation within a bead making these variables non-repeatable. On the other hand, V<sub>T</sub> showed promising results even though it failed ANOVA test since the ratio of numerator to the denominator of F-Statistic is very small [Table 2].

A further investigation into this data, involving bead to bead comparisons, proved that around 90% of the beads fabricated were repeatable. Only a couple of beads were going beyond the acceptable variation for V<sub>T</sub>, the reason for which the ANOVA test failed. Nevertheless, it can be ascertained that the process is not repeatable, with drastic variations in geometry found even on a single bead.

## Conclusion

WAAM-CMT despite being an advanced manufacturing process used for fabrication of big structures, comes with its own defects. Today, a non-standardised factor of safety is used during the fabrication to accommodate for the bad repeatability of the process depending on the in-situ research. The present study has been able to analyse the principal process parameters and basic geometry of the bead to quantify the variation in these parameters. Important variations in the Wire Feed Rate are found in the recorded measurements, which have a direct influence on the fabricated parts. Nevertheless, the variations found in the Travel Speed of the robot were rather small. The bead profiles, and thereby, the height and the width were different from one another even on a single bead analysed.

The ANOVA test confirms the inferences that neither the input parameters nor the output geometry is repeatable. It also shows the extend of this aspect, which indicate that the Travel Speed is close to being identifiable as repeatable, with only a couple of beads going beyond the acceptable variations. Nevertheless, the Wire Feed Rate is much further from being repeatable, which could
be a characteristic of this fabrication technique and more difficult to optimise to reduce the variations. As a result, the process becomes non-repeatable even for the single beads fabricated. Therefore, more complex geometries involving overlapped and layered beads is certain to be non-repeatable since the physics of the fabrications are much more complex.

It is also worth pointing out that the uncertainties cannot be attributed only to these two inputs parameters and could be a result of several other thermo-chemical-mechanical phenomena. Even the unaccounted-for principal parameters like Torch Angle, and wire diameter could introduce small uncertainties into the process, because of their tolerances.

The present study throw light into the repeatability aspect of the process. But further research is required, to take into account additional uncertainties that take place while fabricating more complex geometries than just a bead. Nevertheless, these understanding of variations, with further research on the profile of the bead would enable to model the process better. The study could help future research avoid the hypothesis of constant geometric characteristics for WAAM beads while developing models. Further research on the additional bead profile characteristics like skewness, the overlapping and layered fabrication could be modelled to predict the geometry using statistical tools. Thus, understanding more of the variations in the input parameters and output geometry have the potential to optimize the errors in the modelling arising due to the uncertainties and aid the user to have more stable results.

## Acknowledgements

The authors express their gratitude towards the Agence Nationale de la Recherche (ANR), France for the financial support provided for the project ANR-21-CE10-0005.

The authors would also like to acknowledge the efforts of Mr. Stephane Mathieu of ENSAM, Metz for the planning and execution of experiments.

## References

[1] Dahat S, Hurtig K, Andersson J and Scotti A 2020 A Methodology to Parameterize Wire + Arc Additive Manufacturing: A Case Study for Wall Quality Analysis *J. Manuf. Mater. Process.* 4 14. https://doi.org/10.3390/jmmp4010014

[2] Su C and Chen X 2019 Effect of depositing torch angle on the first layer of wire arc additive manufacture using cold metal transfer (CMT) *Ind. Robot Int. J. Robot. Res.* 46 259–66. https://doi.org/10.1108/IR-11-2018-0233

[3] Zhang M, Fang X, Wang Y, Jiang X, Chang T, Xi N and Huang K 2022 High superelasticity NiTi fabricated by cold metal transfer based wire arc additive manufacturing *Mater. Sci. Eng. A* 840 143001. https://doi.org/10.1016/j.msea.2022.143001

[4] Kannan T and Yoganandh J 2010 Effect of process parameters on clad bead geometry and its shape relationships of stainless steel claddings deposited by GMAW *Int. J. Adv. Manuf. Technol.* 47 1083–95. https://doi.org/10.1007/s00170-009-2226-1

[5] Huang Y, Yang L and Xin Q 2023 Novel geometrical model and design mechanical parameters for CMT-WAAM stainless steel *J. Constr. Steel Res.* 210 108071. https://doi.org/10.1016/j.jCSR.2023.108071

[6] Ding D, Pan Z, Cuiuri D and Li H 2015 A multi-bead overlapping model for robotic wire and arc additive manufacturing (WAAM) *Robot. Comput.-Integr. Manuf.* 31 101–10. https://doi.org/10.1016/j.rcim.2014.08.008

[7] Novelino A L B, Carvalho G C and Ziberov M 2022 Influence of WAAM-CMT deposition parameters on wall geometry *Adv. Ind. Manuf. Eng.* 5 100105. https://doi.org/10.1016/j.aime.2022.100105
be a characteristic of this fabrication technique and more difficult to optimise to reduce the variations. As a result, the process becomes non-repeatable even for the single beads fabricated. Therefore, more complex geometries involving overlapped and layered beads is certain to be non-repeatable since the physics of the fabrications are much more complex.

It is also worth pointing out that the uncertainties cannot be attributed only to these two inputs parameters and could be a result of several other thermo-chemical-mechanical phenomena. Even the unaccounted-for principal parameters like Torch Angle, and wire diameter could introduce small uncertainties into the process, because of their tolerances.

The present study throw light into the repeatability aspect of the process. But further research is required, to take into account additional uncertainties that take place while fabricating more complex geometries than just a bead. Nevertheless, these understanding of variations, with further research on the profile of the bead would enable to model the process better. The study could help future research avoid the hypothesis of constant geometric characteristics for WAAM beads while developing models. Further research on the additional bead profile characteristics like skewness, the overlapping and layered fabrication could be modelled to predict the geometry using statistical tools. Thus, understanding more of the variations in the input parameters and output geometry have the potential to optimize the errors in the modelling arising due to the uncertainties and aid the user to have more stable results.

## Acknowledgements

The authors express their gratitude towards the Agence Nationale de la Recherche (ANR), France for the financial support provided for the project ANR-21-CE10-0005.

The authors would also like to acknowledge the efforts of Mr. Stephane Mathieu of ENSAM, Metz for the planning and execution of experiments.

## References

[1] Dahat S, Hurtig K, Andersson J and Scotti A 2020 A Methodology to Parameterize Wire + Arc Additive Manufacturing: A Case Study for Wall Quality Analysis J. Manuf. Mater. Process. 4 14. https://doi.org/10.3390/jmmp4010014

[2] Su C and Chen X 2019 Effect of depositing torch angle on the first layer of wire arc additive manufacture using cold metal transfer (CMT) Ind. Robot Int. J. Robot. Res. Appl. 46 259-66. https://doi.org/10.1108/IR-11-2018-0233

[3] Zhang M, Fang X, Wang Y, Jiang X, Chang T, Xi N and Huang K 2022 High superelasticity NiTi fabricated by cold metal transfer based wire arc additive manufacturing Mater. Sci. Eng. A 840 143001. https://doi.org/10.1016/j.msea.2022.143001

[4] Kannan T and Yoganandh J 2010 Effect of process parameters on clad bead geometry and its shape relationships of stainless steel claddings deposited by GMAW Int. J. Adv. Manuf. Technol. 47 1083-95. https://doi.org/10.1007/s00170-009-2226-1

[5] Huang Y, Yang L and Xin Q 2023 Novel geometrical model and design mechanical parameters for CMT-WAAM stainless steel J. Constr. Steel Res. 210 108071. https://doi.org/10.1016/j.jCSR.2023.108071

[6] Ding D, Pan Z, Cuiuri D and Li H 2015 A multi-bead overlapping model for robotic wire and arc additive manufacturing (WAAM) Robot. Comput.-Integr. Manuf. 31 101-10. https://doi.org/10.1016/j.rcim.2014.08.008

[7] Novelino A L B, Carvalho G C and Ziberov M 2022 Influence of WAAM-CMT deposition parameters on wall geometry Adv. Ind. Manuf. Eng. 5 100105. https://doi.org/10.1016/j.aime.2022.100105
[8] Wang Z, Zimmer-Chevret S, Léonard F and Abba G 2022 Improvement strategy for the geometric accuracy of bead's beginning and end parts in wire-arc additive manufacturing (WAAM) *Int. J. Adv. Manuf. Technol.* 118 2139-51. https://doi.org/10.1007/s00170-021-08037-8

[9] Pickin C G, Williams S W and Lunt M 2011 Characterisation of the cold metal transfer (CMT) process and its application for low dilution cladding *J. Mater. Process. Technol.* 211 496-502. https://doi.org/10.1016/j.jmatprotec.2010.11.005

[10] Selvi S, Vishvaksenan A and Rajasekar E 2018 Cold metal transfer (CMT) technology - An overview *Def. Technol.* 14 28-44. https://doi.org/10.1016/j.dt.2017.08.002

[11] Palani P K and Murugan N 2006 Sensitivity Analysis for Process Parameters in Cladding of Stainless Steel by Flux Cored Arc Welding *J. Manuf. Process.* 8 90-100. https://doi.org/10.1016/S1526-6125(06)80004-6

[12] Bourlet C Développement de la fabrication additive par procédé arc-fil pour les aciers: caractérisation microstructurale et mécanique des dépôts en nuances ER100 et 316L pour la validation des propriétés d'emploi de pièces industrielles

[13] Rosli N A and Alkahari M R Effects Of Wire Feeding Angle in Wire Arc Additive Manufacturing

[14] WANG Z 2022 Robotisation de la fabrication additive par procédé arc-fil : Identification et amélioration de la commande (UNIVERSITÉ DE LORRAINE)

[15] Anon Evaluation of measurement data — Guide to the expression of uncertainty in measurement

[16] Hibbert D B 2018 QUALITY ASSURANCE | Interlaboratory Studies Reference Module in Chemistry, Molecular Sciences and Chemical Engineering (Elsevier) p B9780124095472127970. https://doi.org/10.1016/B978-0-12-409547-2.12797-0
[8] Wang Z, Zimmer-Chevret S, Léonard F and Abba G 2022 Improvement strategy for the geometric accuracy of bead's beginning and end parts in wire-arc additive manufacturing (WAAM) *Int. J. Adv. Manuf. Technol.* 118 2139-51. https://doi.org/10.1007/s00170-021-08037-8

[9] Pickin C G, Williams S W and Lunt M 2011 Characterisation of the cold metal transfer (CMT) process and its application for low dilution cladding *J. Mater. Process. Technol.* 211 496-502. https://doi.org/10.1016/j.jmatprotec.2010.11.005

[10] Selvi S, Vishvaksenan A and Rajasekar E 2018 Cold metal transfer (CMT) technology - An overview *Def. Technol.* 14 28-44. https://doi.org/10.1016/j.dt.2017.08.002

[11] Palani P K and Murugan N 2006 Sensitivity Analysis for Process Parameters in Cladding of Stainless Steel by Flux Cored Arc Welding *J. Manuf. Process.* 8 90-100. https://doi.org/10.1016/S1526-6125(06)80004-6

[12] Bourlet C Développement de la fabrication additive par procédé arc-fil pour les aciers: caractérisation microstructurale et mécanique des dépôts en nuances ER100 et 316L pour la validation des propriétés d'emploi de pièces industrielles

[13] Rosli N A and Alkahari M R Effects Of Wire Feeding Angle in Wire Arc Additive Manufacturing

[14] WANG Z 2022 Robotisation de la fabrication additive par procédé arc-fil : Identification et amélioration de la commande (UNIVERSITÉ DE LORRAINE)

[15] Anon Evaluation of measurement data — Guide to the expression of uncertainty in measurement

[16] Hibbert D B 2018 QUALITY ASSURANCE | Interlaboratory Studies Reference Module in Chemistry, Molecular Sciences and Chemical Engineering (Elsevier) p B9780124095472127970. https://doi.org/10.1016/B978-0-12-409547-2.12797-0
# An experimental investigation into tribological behaviour of additively manufactured biocompatible Ti-6Al-4V alloy

MADEJ Marcin<fnref n="1" />,ª, KARKALOS Nikolaos E.<fnref n="2" />,ᵇ,
KARMIRIS-OBRATANSKI Panagiotis<fnref n="3" />,ᶜ, PAPAZOGLOU Emmanouil L.<fnref n="2" />,ᵈ*,
GEORGAKOPOULOS-SOARES Ilias<fnref n="4" />,ᵉ and MARKOPOULOS Angelos P.<fnref n="1" />,>f

**Keywords:** Friction Coefficient, Wear Rate, Tribology, Additive Manufacturing, Ti-6Al-4V

**Abstract.** Additive manufacturing techniques are increasingly being utilized in industrial-level applications due to their flexibility and ability to produce customized parts, such as various types of biomedical implants. However, the conditions during additive manufacturing fabrication and the nature of these processes can lead to implications on the properties of the produced parts, potentially requiring appropriate post-processing before real applications. The tribological behavior of printed parts not only affects their performance but also their service life, making it crucial to investigate their wear rate and friction coefficient under different lubricant environments. In this study, an experimental investigation was conducted on as-printed Ti-6Al-4V specimens to determine the effect of various lubricant environments on wear rate and friction coefficient. The results demonstrated that the reduction in wear rate in liquid environments can be significantly hindered by the accumulation of debris from the worn specimen. However, the development of a thin film of an appropriate lubricant was shown to be favorable regarding the friction behavior of printed parts.

**Introduction**

Additive Manufacturing (AM) is a nonconventional, automated freeform fabrication manufacturing process that operates on the principle of layer-by-layer construction [1,2]. AM can offer several other advantages, including the elimination of the need for tools and fixtures, the capability for nesting and parallel processing, the reduction in scrap material or waste, and the ability to economically produce customized parts and components [3]. Moreover, AM can simultaneously define both the geometry and material properties of a part during the build process [4]. Laser Powder Bed Fusion (L-PBF) is among the first commercialized AM methods, wherein the power/energy source for bonding the raw material, which is in powder form, is provided by a focused laser beam. L-PBF began as a polymer-based method but has, over the years, expanded to include metals and ceramics [5]. Although LPBF processes provide significant advantages in metal processing, they are limited by uncertainties due to complex physical phenomena that occur on multiple spatial and temporal scales [6]. Consequently, extensive research focuses on understanding post-printing material properties and exploring methods to improve them. The

<figure><img src="image_171.png" /></figure>

Content from this work may be used under the terms of the Creative Commons Attribution 3.0 license. Any further distribution of this work must maintain attribution to the author(s) and the title of the work, journal citation and DOI. Published under license by Materials Research Forum LLC.

<footnote n="1">¹Department of Physical and Powder Metallurgy, Faculty of Metals Engineering and Industrial Computer Science, AGH University of Cracow, 30-059 Cracow, Poland</footnote>

<footnote n="2">²School of Mechanical Engineering, Section of Manufacturing Technology, National Technical University of Athens, Heroon Polytechniou 9, 15780 Athens, Greece</footnote>

<footnote n="3">³Department of Manufacturing Systems, Faculty of Mechanical Engineering and Robotics, AGH University of Cracow, 30-059 Cracow, Poland</footnote>

<footnote n="4">⁴Institute for Personalized Medicine, Department of Biochemistry and Molecular Biology, The Pennsylvania State University College of Medicine, Hershey, PA, USA</footnote>

<footnote>ªmmadej@agh.edu.pl, ᵇnkark@mail.ntua.gr, ᶜkarmiris@agh.edu.pl, ᵈmlpapazoglou@gmail.com, ᶜizg5139@psu.edu, ᶠamark@mail.ntua.gr</footnote>
# An experimental investigation into tribological behaviour of additively manufactured biocompatible Ti-6Al-4V alloy

MADEJ Marcin<fnref n="1" />,ª, KARKALOS Nikolaos E.<fnref n="2" />,ᵇ,

KARMIRIS-OBRATANSKI Panagiotis<fnref n="3" />,ᶜ, PAPAZOGLOU Emmanouil L.<fnref n="2" />,ᵈ*,
GEORGAKOPOULOS-SOARES Ilias<fnref n="4" />,ᵉ and MARKOPOULOS Angelos P.<fnref n="1" />,>f

**Keywords:** Friction Coefficient, Wear Rate, Tribology, Additive Manufacturing, Ti-6Al-4V

**Abstract.** Additive manufacturing techniques are increasingly being utilized in industrial-level applications due to their flexibility and ability to produce customized parts, such as various types of biomedical implants. However, the conditions during additive manufacturing fabrication and the nature of these processes can lead to implications on the properties of the produced parts, potentially requiring appropriate post-processing before real applications. The tribological behavior of printed parts not only affects their performance but also their service life, making it crucial to investigate their wear rate and friction coefficient under different lubricant environments. In this study, an experimental investigation was conducted on as-printed Ti-6Al-4V specimens to determine the effect of various lubricant environments on wear rate and friction coefficient. The results demonstrated that the reduction in wear rate in liquid environments can be significantly hindered by the accumulation of debris from the worn specimen. However, the development of a thin film of an appropriate lubricant was shown to be favorable regarding the friction behavior of printed parts.

**Introduction**

Additive Manufacturing (AM) is a nonconventional, automated freeform fabrication manufacturing process that operates on the principle of layer-by-layer construction [1,2]. AM can offer several other advantages, including the elimination of the need for tools and fixtures, the capability for nesting and parallel processing, the reduction in scrap material or waste, and the ability to economically produce customized parts and components [3]. Moreover, AM can simultaneously define both the geometry and material properties of a part during the build process [4]. Laser Powder Bed Fusion (L-PBF) is among the first commercialized AM methods, wherein the power/energy source for bonding the raw material, which is in powder form, is provided by a focused laser beam. L-PBF began as a polymer-based method but has, over the years, expanded to include metals and ceramics [5]. Although LPBF processes provide significant advantages in metal processing, they are limited by uncertainties due to complex physical phenomena that occur on multiple spatial and temporal scales [6]. Consequently, extensive research focuses on understanding post-printing material properties and exploring methods to improve them. The

<figure><img src="image_169.png" /></figure>

Content from this work may be used under the terms of the Creative Commons Attribution 3.0 license. Any further distribution of this work must maintain attribution to the author(s) and the title of the work, journal citation and DOI. Published under license by Materials Research Forum LLC.

<footnote n="1">¹Department of Physical and Powder Metallurgy, Faculty of Metals Engineering and Industrial Computer Science, AGH University of Cracow, 30-059 Cracow, Poland</footnote>

<footnote n="2">²School of Mechanical Engineering, Section of Manufacturing Technology, National Technical University of Athens, Heroon Polytechniou 9, 15780 Athens, Greece</footnote>

<footnote n="3">³Department of Manufacturing Systems, Faculty of Mechanical Engineering and Robotics, AGH University of Cracow, 30-059 Cracow, Poland</footnote>

<footnote n="4">⁴Institute for Personalized Medicine, Department of Biochemistry and Molecular Biology, The Pennsylvania State University College of Medicine, Hershey, PA, USA</footnote>

<footnote>ªmmadej@agh.edu.pl, ᵇnkark@mail.ntua.gr, ᶜkarmiris@agh.edu.pl, ᵈmlpapazoglou@gmail.com,</footnote>

<footnote>ᵉizg5139@psu.edu, ᶜamark@mail.ntua.gr</footnote>
combination of titanium/titanium alloys and L-PBF has proven to be a feasible technology, finding applicability in the aerospace, medical, and automotive sectors. This is a result of titanium's excellent properties and AM's inherent capability to produce complex and lightweight structures and components [7, 8]. However, the outcome of the LPBF process for titanium alloys such as the resulting microstructures and mechanical properties, and consequently, the properties of the final built part can be influenced by various factors [9]. The most common research topic regarding L-PBF processed materials involves studying their mechanical properties. However, considering the demanding operating conditions of titanium alloys and their applicability across a wide range of fields, it is also scientifically and practically valuable to investigate the tribological properties and wear behavior of L-PBF fabricated titanium alloys under various conditions.

The current research's feasibility is supported by two main factors. Firstly, the tribological and wear properties of a material are directly dependent on its microstructure [10]. Secondly, the wear environment significantly influences the development of friction, thereby affecting the tribological behavior and resultant wear. Inherent to L-PBF is the development of a unique microstructure, a consequence of thermal loading cycles involving rapid heating and cooling. Consequently, it is anticipated that materials processed through L-PBF will exhibit different properties compared to those manufactured by conventional methods. For instance, Attar et al. [11] conducted a comparison of the wear properties between cPTi (commercial pure titanium) parts produced by Selective Laser Melting (SLM) and their cast-produced counterparts. Based on SEM observations, they noted a martensitic microstructure in SLM-fabricated specimens, in contrast to a plate-like microstructure in cast specimens. Furthermore, although both SLM-fabricated and cast cPTi samples exhibited similar wear mechanisms, the SLM-fabricated samples demonstrated superior wear resistance compared to the cast versions. Liang et al. [12] linked the tribological properties of SLM-fabricated Ti6Al4V and the material's machinability, focusing specifically on its tribological behavior against cemented carbide under dry conditions. Among various notable findings, they highlighted that the relatively high friction coefficient and its wide fluctuation range during the friction between SLM titanium alloy and cemented carbide suggest intense wear under dry conditions. Additionally, they observed that the average friction coefficient was largely independent of the applied load, while the influence of temperature on tribological properties was more complex. In another notable study focusing on the tribological and wear behavior of L-PBF fabricated titanium alloys, Liu et al. [13] examine the anisotropy of wear and tribo-corrosion properties in L-PBF fabricated Ti6Al4V. They concluded that this anisotropy is linked to the material's microstructure. Consequently, they found that the horizontal plane, perpendicular to the building direction, exhibits superior wear and tribo-corrosion resistance compared to the vertical plane, which is parallel to the building direction.

Kaur et al. [14] and Mahamood et al. [15] have highlighted the significance of studying the wear and tribological behavior of titanium alloys, especially given their widespread use in demanding bioengineering applications. Consequently, it is reasonably deduced that the investigation of the tribological and wear properties of titanium alloys is a crucial and continuously developing area of research. Simultaneously, it is equally important to investigate the materials' tribological behavior under various environments and conditions that closely simulate their actual application scenarios. For instance, Pandey et al. [16] investigated the tribological performance of SS316L, and Ti6Al4V in various solutions, i.e., artificial saliva (AS), phosphate buffer solution (PBS), Ringer's solution (RS), and simulated body fluid (SBF), to simulate their use in biomedical applications. Their findings indicated that wear is a combination of abrasion, adhesion, oxidation, and delamination. Furthermore, they demonstrated that the wear characteristics of both cPTi and Ti6Al4V are significantly influenced by the different solutions, simulating varied environments. Sirin et al. [17] also considered different environments in their study, where they conducted a series of friction wear tests on titanium Grade 2 under various conditions. These conditions
combination of titanium/titanium alloys and L-PBF has proven to be a feasible technology, finding applicability in the aerospace, medical, and automotive sectors. This is a result of titanium's excellent properties and AM's inherent capability to produce complex and lightweight structures and components [7, 8]. However, the outcome of the LPBF process for titanium alloys such as the resulting microstructures and mechanical properties, and consequently, the properties of the final built part can be influenced by various factors [9]. The most common research topic regarding L-PBF processed materials involves studying their mechanical properties. However, considering the demanding operating conditions of titanium alloys and their applicability across a wide range of fields, it is also scientifically and practically valuable to investigate the tribological properties and wear behavior of L-PBF fabricated titanium alloys under various conditions.

The current research's feasibility is supported by two main factors. Firstly, the tribological and wear properties of a material are directly dependent on its microstructure [10]. Secondly, the wear environment significantly influences the development of friction, thereby affecting the tribological behavior and resultant wear. Inherent to L-PBF is the development of a unique microstructure, a consequence of thermal loading cycles involving rapid heating and cooling. Consequently, it is anticipated that materials processed through L-PBF will exhibit different properties compared to those manufactured by conventional methods. For instance, Attar et al. [11] conducted a comparison of the wear properties between cPTi (commercial pure titanium) parts produced by Selective Laser Melting (SLM) and their cast-produced counterparts. Based on SEM observations, they noted a martensitic microstructure in SLM-fabricated specimens, in contrast to a plate-like microstructure in cast specimens. Furthermore, although both SLM-fabricated and cast cPTi samples exhibited similar wear mechanisms, the SLM-fabricated samples demonstrated superior wear resistance compared to the cast versions. Liang et al. [12] linked the tribological properties of SLM-fabricated Ti6Al4V and the material's machinability, focusing specifically on its tribological behavior against cemented carbide under dry conditions. Among various notable findings, they highlighted that the relatively high friction coefficient and its wide fluctuation range during the friction between SLM titanium alloy and cemented carbide suggest intense wear under dry conditions. Additionally, they observed that the average friction coefficient was largely independent of the applied load, while the influence of temperature on tribological properties was more complex. In another notable study focusing on the tribological and wear behavior of L-PBF fabricated titanium alloys, Liu et al. [13] examine the anisotropy of wear and tribo-corrosion properties in L-PBF fabricated Ti6Al4V. They concluded that this anisotropy is linked to the material's microstructure. Consequently, they found that the horizontal plane, perpendicular to the building direction, exhibits superior wear and tribo-corrosion resistance compared to the vertical plane, which is parallel to the building direction.

Kaur et al. [14] and Mahamood et al. [15] have highlighted the significance of studying the wear and tribological behavior of titanium alloys, especially given their widespread use in demanding bioengineering applications. Consequently, it is reasonably deduced that the investigation of the tribological and wear properties of titanium alloys is a crucial and continuously developing area of research. Simultaneously, it is equally important to investigate the materials' tribological behavior under various environments and conditions that closely simulate their actual application scenarios. For instance, Pandey et al. [16] investigated the tribological performance of SS316L, and Ti6Al4V in various solutions, i.e., artificial saliva (AS), phosphate buffer solution (PBS), Ringer's solution (RS), and simulated body fluid (SBF), to simulate their use in biomedical applications. Their findings indicated that wear is a combination of abrasion, adhesion, oxidation, and delamination. Furthermore, they demonstrated that the wear characteristics of both cPTi and Ti6Al4V are significantly influenced by the different solutions, simulating varied environments. Sirin et al. [17] also considered different environments in their study, where they conducted a series of friction wear tests on titanium Grade 2 under various conditions. These conditions
included dry environment, base fluid (vegetable-based oil), graphene nanoplatelets (GNPs), hexagonal boron nitride (hBN) nanofluids, and a hybrid nanofluid of GNPs and hBN. Their results indicated that the friction environment significantly impacts the material's wear characteristics.

In the current study tribological tests were conducted on as-printed Ti-6Al-4V specimens under different lubricant environments in order to determine their effect on the tribological behavior of the printed specimens by analyzing friction coefficient and wear rate. The tribological tests were carried out for two different sliding distances in order to further investigate the effect of environment on the progression of wear on the specimens.

## Materials and Methods

In the present study, tribological properties of additively manufactured Ti-6Al-4V specimens under different conditions were evaluated, in order to determine the effect of lubricants on the friction and wear behavior of the printed specimens. At first, a small square shaped Ti-6Al-4V slab was printed by a L-PBF machine tool under the recommended optimum conditions. The tribological specimens were then cut to the required dimensions in a Wire-EDM machine tool in order to eliminate possible alterations to the specimens' microstructure due to heat-induced phenomena. Then, the tribological properties were tested by using a T-05 block-on-ring tribotester (Łukasiewicz - ITeE, Radom, Poland) at ambient temperature (21 °C) under technically dry and wet friction conditions. The wet sliding contact was conducted with two different liquids, namely distilled water and SUPRACUT MQL 45 is a synthetic neat cutting oil developed for machining involving chips removal of ferrous and non-ferrous metals. During the test, in order to ensure proper contact between the specimen and the steel ring (heat-treated steel 100Cr6, 53 HRC, Ø49.5×8 mm) rotating at a constant speed, a rectangular tribological sample (4x4x20 mm) was fixed in a holder containing a hemispherical insert. The surface of the specimen in contact with the friction was perpendicular to the direction of the load (L). A double lever system was used to push the specimen in the direction of the ring with an accuracy of the load of ± 1.5%. In total, six different conditions were employed for the tribological tests, as can be seen in Table 1, by conducting the tests under three different environments for 2 different sliding distances, which were supposed to be representative of the conditions occurring under real operating cases. The other testing parameters, such as rotational speed and load were kept constant. Each test was repeated three times and standard deviation of the measurements was also reported.

<table><caption>Table 1. The wear test parameters.</caption><thead><tr><th>Sliding contact</th><th>Counter-specimen</th><th>Rotational speed</th><th>Load</th><th>Sliding distance</th></tr></thead><tbody><tr><td>Technical dry</td><td>steel 100Cr6,</td><td rowspan="2">200 [rpm]</td><td rowspan="2">100 [N]</td><td>250 [m]</td></tr><tr><td>Wet (water and oil)</td><td>heat-treated, with hardness 53 HRC</td><td>1000 [m]</td></tr></tbody></table>

The loss of mass during the test and the average coefficient of friction both in wet and technically dry conditions were measured by dedicated software, whereas the wear rate was determined according to Eq.1:

$$W_r = \frac{\Delta m}{L \cdot S_d} \qquad (1)$$

where Δm represents loss of mass in g, L represents load in N and Sd represents sliding distance in m.
included dry environment, base fluid (vegetable-based oil), graphene nanoplatelets (GNPs), hexagonal boron nitride (hBN) nanofluids, and a hybrid nanofluid of GNPs and hBN. Their results indicated that the friction environment significantly impacts the material's wear characteristics.

In the current study tribological tests were conducted on as-printed Ti-6Al-4V specimens under different lubricant environments in order to determine their effect on the tribological behavior of the printed specimens by analyzing friction coefficient and wear rate. The tribological tests were carried out for two different sliding distances in order to further investigate the effect of environment on the progression of wear on the specimens.

## Materials and Methods

In the present study, tribological properties of additively manufactured Ti-6Al-4V specimens under different conditions were evaluated, in order to determine the effect of lubricants on the friction and wear behavior of the printed specimens. At first, a small square shaped Ti-6Al-4V slab was printed by a L-PBF machine tool under the recommended optimum conditions. The tribological specimens were then cut to the required dimensions in a Wire-EDM machine tool in order to eliminate possible alterations to the specimens' microstructure due to heat-induced phenomena. Then, the tribological properties were tested by using a T-05 block-on-ring tribotester (Łukasiewicz - ITeE, Radom, Poland) at ambient temperature (21 °C) under technically dry and wet friction conditions. The wet sliding contact was conducted with two different liquids, namely distilled water and SUPRACUT MQL 45 is a synthetic neat cutting oil developed for machining involving chips removal of ferrous and non-ferrous metals. During the test, in order to ensure proper contact between the specimen and the steel ring (heat-treated steel 100Cr6, 53 HRC, Ø49.5×8 mm) rotating at a constant speed, a rectangular tribological sample (4x4x20 mm) was fixed in a holder containing a hemispherical insert. The surface of the specimen in contact with the friction was perpendicular to the direction of the load (L). A double lever system was used to push the specimen in the direction of the ring with an accuracy of the load of ± 1.5%. In total, six different conditions were employed for the tribological tests, as can be seen in Table 1, by conducting the tests under three different environments for 2 different sliding distances, which were supposed to be representative of the conditions occurring under real operating cases. The other testing parameters, such as rotational speed and load were kept constant. Each test was repeated three times and standard deviation of the measurements was also reported.

<table><caption>Table 1. The wear test parameters.</caption><thead><tr><th>Sliding contact</th><th>Counter-specimen</th><th>Rotational speed</th><th>Load</th><th>Sliding distance</th></tr></thead><tbody><tr><td>Technical dry</td><td>steel 100Cr6,</td><td rowspan="2">200 [rpm]</td><td rowspan="2">100 [N]</td><td>250 [m]</td></tr><tr><td>Wet (water and oil)</td><td>heat-treated, with hardness 53 HRC</td><td>1000 [m]</td></tr></tbody></table>

The loss of mass during the test and the average coefficient of friction both in wet and technically dry conditions were measured by dedicated software, whereas the wear rate was determined according to Eq.1:

$$W_r = \frac{\Delta m}{L \cdot S_d} \qquad (1)$$

where Δm represents loss of mass in g, L represents load in N and Sd represents sliding distance in m.
## Results and Discussion

*Mass loss, average friction coefficient and wear rate.* After the tribological tests were carried out, the results regarding mass loss, average coefficient of friction and wear rate were obtained.

<figure><img src="image_172.png" /><figcaption>Fig. 1. Effect of tribological test parameters on mass loss.</figcaption></figure>

At first, the wear resistance of the specimens was able to be studied based on mass loss measurements. It is anticipated that mass loss is strictly dependent on the friction conditions (dry or wet) and the friction distance. In Fig. 1, it can be seen that an increase in the friction distance results in an increase in weight loss, in every case, with the increase being more prominent in the case of wet environment. Moreover, the use of water and oil eventually leads to higher mass loss than the one occurring in dry environment in any case.

<figure><img src="image_173.png" /><figcaption>Fig. 2. Wear rate as a function of tribological test parameters.</figcaption></figure>

However, based on the results of Fig. 2, it can be seen that the phenomena occurring during contact stabilize over time, as the rate of mass loss is gradually reduced along the imposed sliding distance, thus leading to a decrease in the wear rate, except for the tests conducted in oil environment, for which the rate of mass loss is assumed to be slightly increasing thus causing higher levels of wear rate.

Furthermore, the results of Fig. 2 suggest that, in the case of the titanium alloy tested, both water and oil do not reduce wear but significantly increase it. These observations can be directly attributed to the fact that both water and oil became contaminated with chipped sample fragments, turning into suspensions, which also influenced the test as they could contribute to increased material removal from the specimens. Moreover, the inherently porous nature of printed parts and their relatively high roughness may also be the reason for this phenomenon, as during the early
## Results and Discussion

*Mass loss, average friction coefficient and wear rate.* After the tribological tests were carried out, the results regarding mass loss, average coefficient of friction and wear rate were obtained.

<figure><img src="image_170.png" /><figcaption>Fig. 1. Effect of tribological test parameters on mass loss.</figcaption></figure>

At first, the wear resistance of the specimens was able to be studied based on mass loss measurements. It is anticipated that mass loss is strictly dependent on the friction conditions (dry or wet) and the friction distance. In Fig. 1, it can be seen that an increase in the friction distance results in an increase in weight loss, in every case, with the increase being more prominent in the case of wet environment. Moreover, the use of water and oil eventually leads to higher mass loss than the one occurring in dry environment in any case.

<figure><img src="image_171.png" /><figcaption>Fig. 2. Wear rate as a function of tribological test parameters.</figcaption></figure>

However, based on the results of Fig. 2, it can be seen that the phenomena occurring during contact stabilize over time, as the rate of mass loss is gradually reduced along the imposed sliding distance, thus leading to a decrease in the wear rate, except for the tests conducted in oil environment, for which the rate of mass loss is assumed to be slightly increasing thus causing higher levels of wear rate.

Furthermore, the results of Fig. 2 suggest that, in the case of the titanium alloy tested, both water and oil do not reduce wear but significantly increase it. These observations can be directly attributed to the fact that both water and oil became contaminated with chipped sample fragments, turning into suspensions, which also influenced the test as they could contribute to increased material removal from the specimens. Moreover, the inherently porous nature of printed parts and their relatively high roughness may also be the reason for this phenomenon, as during the early
stages of the wear, fragments from rough protrusions on the specimen surface could have probably been removed and remain in the liquid as debris, something that can also occur in real applications.

On the other hand, regarding coefficient of friction, the use of liquid environment is proven to achieve at least a comparable friction coefficient or a significantly lower one. Based on the results of Fig. 3, it can be seen that the use of water does not reduce the coefficient of friction, even over a distance of 250 m, showing similarity to dry friction conditions. However, as expected, the application of oil causes a significant reduction in the coefficient of friction compared to dry friction due to the development of a thin film between the two contacting surfaces during the tests, although this reduction did not correspond to a decrease in the wear rate.

<figure><img src="image_174.png" /><figcaption>Fig. 3. Coefficient of friction as a function of tribological test parameters.</figcaption></figure>

Especially, regarding friction coefficient, its progression can be directly observed in the friction force plots of Fig. 4 and 5.

<figure><img src="image_175.png" /><figcaption>Fig. 4. Friction force of the investigated samples as a function of time and processing conditions, sliding distance: 250 m.</figcaption></figure>
stages of the wear, fragments from rough protrusions on the specimen surface could have probably been removed and remain in the liquid as debris, something that can also occur in real applications.

On the other hand, regarding coefficient of friction, the use of liquid environment is proven to achieve at least a comparable friction coefficient or a significantly lower one. Based on the results of Fig. 3, it can be seen that the use of water does not reduce the coefficient of friction, even over a distance of 250 m, showing similarity to dry friction conditions. However, as expected, the application of oil causes a significant reduction in the coefficient of friction compared to dry friction due to the development of a thin film between the two contacting surfaces during the tests, although this reduction did not correspond to a decrease in the wear rate.

<figure><img src="image_172.png" /><figcaption>Fig. 3. Coefficient of friction as a function of tribological test parameters.</figcaption></figure>

Especially, regarding friction coefficient, its progression can be directly observed in the friction force plots of Fig. 4 and 5.

<figure><img src="image_173.png" /><figcaption>Fig. 4. Friction force of the investigated samples as a function of time and processing conditions, sliding distance: 250 m.</figcaption></figure>
<figure><img src="image_176.png" /><figcaption>Fig. 5. Friction force of the investigated samples as a function of time and processing conditions, sliding distance: 1000 m.</figcaption></figure>

The changes in frictional force recorded over time illustrate the stability of the friction process. Analyzing the curves in Fig. 4 and 5, it can be concluded that the greatest fluctuations occur in the test performed with water. The use of oil stabilizes the friction path significantly; in the dry-start test, the force gradually decreases and stabilizes after approximately 500 s. No significant instantaneous changes in force are observed in the oil process curve, indicating the presence of an oil film between the test sample and the steel counter example.

*Wear mechanisms.* Apart from determining the fundamental tribological properties of the material under the applied test conditions, it is important to identify the phenomena taking place at the friction node. This identification is made possible by observing the surface morphology after friction, which, in turn, allows the recognition of the friction mechanisms that occurred during the test. Thus, Fig. 6 presents the surface prior to tribology tests, while Figs. 7–9 depict the surfaces after the tests, providing a comprehensive overview of typical examples of the post-friction surface morphology of the tested titanium alloy.

<figure><img src="image_177.png" /><figcaption>Fig. 6. Surface morphology prior to tribology experiments.</figcaption></figure>
<figure><img src="image_174.png" /><figcaption>Fig. 5. Friction force of the investigated samples as a function of time and processing conditions, sliding distance: 1000 m.</figcaption></figure>

The changes in frictional force recorded over time illustrate the stability of the friction process. Analyzing the curves in Fig. 4 and 5, it can be concluded that the greatest fluctuations occur in the test performed with water. The use of oil stabilizes the friction path significantly; in the dry-start test, the force gradually decreases and stabilizes after approximately 500 s. No significant instantaneous changes in force are observed in the oil process curve, indicating the presence of an oil film between the test sample and the steel counter example.

*Wear mechanisms.* Apart from determining the fundamental tribological properties of the material under the applied test conditions, it is important to identify the phenomena taking place at the friction node. This identification is made possible by observing the surface morphology after friction, which, in turn, allows the recognition of the friction mechanisms that occurred during the test. Thus, Fig. 6 presents the surface prior to tribology tests, while Figs. 7–9 depict the surfaces after the tests, providing a comprehensive overview of typical examples of the post-friction surface morphology of the tested titanium alloy.

<figure><img src="image_175.png" /><figcaption>Fig. 6. Surface morphology prior to tribology experiments.</figcaption></figure>
<figure><img src="image_178.png" /><figcaption>Fig. 7. Surface morphology after sliding contact under technically dry friction.</figcaption></figure>

<figure><img src="image_179.png" /></figure>

<figure><img src="image_180.png" /><figcaption>Fig. 8. Surface morphology after sliding contact under wet (water) friction.</figcaption></figure>

<figure><img src="image_181.png" /></figure>

<figure><img src="image_182.png" /><figcaption>Fig. 9. Surface morphology after sliding contact under wet (oil) friction.</figcaption></figure>

<figure><img src="image_183.png" /></figure>
<figure><img src="image_176.png" /><figcaption>Fig. 7. Surface morphology after sliding contact under technically dry friction.</figcaption></figure>

<figure><img src="image_177.png" /></figure>

<figure><img src="image_178.png" /><figcaption>Fig. 8. Surface morphology after sliding contact under wet (water) friction.</figcaption></figure>

<figure><img src="image_179.png" /></figure>

<figure><img src="image_180.png" /><figcaption>Fig. 9. Surface morphology after sliding contact under wet (oil) friction.</figcaption></figure>

<figure><img src="image_181.png" /></figure>
After analyzing the morphology of the worn surfaces depicted in Figs. 6 to 8, it can be concluded that the predominant friction mechanism is abrasive wear, characterized by scratching and furrowing. Additionally, chipping marks are visible on all observed surfaces, regardless of whether wet or dry friction was employed. As mentioned earlier, both oil and water became contaminated with wear products, indicating the intensity of the process despite the short test time. Analyzing the weight losses shown in Fig. 1 under different friction conditions, it is evident that chipping processes and the reaction of fluids with the resulting cracks are crucial in the friction process. Both oil and water, due to capillary forces, penetrate fatigue-induced cracks on the surface, causing rapid chipping. This phenomenon is not observed in dry friction and is described as pitting, typically occurring after an extended period of friction.

Furthermore, these observations suggest a weak particle bond easily broken by load-related frictional forces, leading to crack formation into which fluid infiltrates under high pressure. Unfortunately, in the case of oil, despite a significant reduction in the friction coefficient and the obvious stabilization of the friction force, there is no improvement in the wear resistance of the tested materials. These findings can be proven crucial for the potential applications of the printed specimens.

## Conclusions

In this study, tribological tests were carried out on biomedical grade additively manufactured Ti-6Al-4V specimens under different lubricant conditions in order to evaluate the effect of lubrication environment on the friction and wear behavior of the printed specimens for two different sliding distances.

Wear rate calculations based on mass loss measurements showed that the mass loss rate is gradually reduced leading to lower wear rate for longer sliding distances for dry and water environment whereas a slight increase was observed in the oil environment, probably due to accumulation of debris in the liquid, resulting from the removal of protrusions from the porous surface of as-printed specimens. This phenomenon also caused the wear rate being larger in the case of wet friction tests, compared to the dry friction case.

Regarding friction coefficient, it was demonstrated that the friction coefficient can be efficiently reduced by using an appropriate lubricant such as oil, as a threefold reduction in friction coefficient was observed due to the development of a thin film of lubricant between the two contacting surfaces. However, the capability of distilled water as lubricant was proven insignificant in the present case.

## References

[1] A. Gebhardt, Understanding Additive Manufacturing. München: Carl Hanser Verlag GmbH & Co. KG, 2011.

[2] S. Kumar, Additive Manufacturing Solutions. Springer, 2022.

[3] M. Srivastava, S. Rathee, S. Maheshwari, and T. K. Kundra, Additive Manufacturing Fundamentals and advancements. New York, U.S.A: CRC Press, 2020.

[4] A. Gebhardt and J.-S. Hotter, Additive Manufacturing - 3D Printing for Prototyping and Manufacturing. HANSER Publications, 2016.

[5] I. Gibson, D. W. Rosen, and B. Stucker, Additive Manufacturing Technologies, 2nd ed. Springer-Verlag New York, 2015.

[6] F. Lopez, P. Witherell, and B. Lane, “Identifying uncertainty in laser powder bed fusion additive manufacturing models,” J. Mech. Des., vol. 138, no. 11, pp. 1–4, 2016. https://doi.org/10.1115/1.4034103
After analyzing the morphology of the worn surfaces depicted in Figs. 6 to 8, it can be concluded that the predominant friction mechanism is abrasive wear, characterized by scratching and furrowing. Additionally, chipping marks are visible on all observed surfaces, regardless of whether wet or dry friction was employed. As mentioned earlier, both oil and water became contaminated with wear products, indicating the intensity of the process despite the short test time. Analyzing the weight losses shown in Fig. 1 under different friction conditions, it is evident that chipping processes and the reaction of fluids with the resulting cracks are crucial in the friction process. Both oil and water, due to capillary forces, penetrate fatigue-induced cracks on the surface, causing rapid chipping. This phenomenon is not observed in dry friction and is described as pitting, typically occurring after an extended period of friction.

Furthermore, these observations suggest a weak particle bond easily broken by load-related frictional forces, leading to crack formation into which fluid infiltrates under high pressure. Unfortunately, in the case of oil, despite a significant reduction in the friction coefficient and the obvious stabilization of the friction force, there is no improvement in the wear resistance of the tested materials. These findings can be proven crucial for the potential applications of the printed specimens.

## Conclusions

In this study, tribological tests were carried out on biomedical grade additively manufactured Ti-6Al-4V specimens under different lubricant conditions in order to evaluate the effect of lubrication environment on the friction and wear behavior of the printed specimens for two different sliding distances.

Wear rate calculations based on mass loss measurements showed that the mass loss rate is gradually reduced leading to lower wear rate for longer sliding distances for dry and water environment whereas a slight increase was observed in the oil environment, probably due to accumulation of debris in the liquid, resulting from the removal of protrusions from the porous surface of as-printed specimens. This phenomenon also caused the wear rate being larger in the case of wet friction tests, compared to the dry friction case.

Regarding friction coefficient, it was demonstrated that the friction coefficient can be efficiently reduced by using an appropriate lubricant such as oil, as a threefold reduction in friction coefficient was observed due to the development of a thin film of lubricant between the two contacting surfaces. However, the capability of distilled water as lubricant was proven insignificant in the present case.

## References

[1] A. Gebhardt, Understanding Additive Manufacturing. München: Carl Hanser Verlag GmbH & Co. KG, 2011.

[2] S. Kumar, Additive Manufacturing Solutions. Springer, 2022.

[3] M. Srivastava, S. Rathee, S. Maheshwari, and T. K. Kundra, Additive Manufacturing Fundamentals and advancements. New York, U.S.A: CRC Press, 2020.

[4] A. Gebhardt and J.-S. Hotter, Additive Manufacturing - 3D Printing for Prototyping and Manufacturing. HANSER Publications, 2016.

[5] I. Gibson, D. W. Rosen, and B. Stucker, Additive Manufacturing Technologies, 2nd ed. Springer-Verlag New York, 2015.

[6] F. Lopez, P. Witherell, and B. Lane, “Identifying uncertainty in laser powder bed fusion additive manufacturing models,” J. Mech. Des., vol. 138, no. 11, pp. 1–4, 2016. https://doi.org/10.1115/1.4034103
[7] C. Y. Yap et al., "Review of selective laser melting: Materials and applications," Appl. Phys. Rev., vol. 2, no. 4, 2015. https://doi.org/10.1063/1.4935926

[8] A. N. Aufa, M. Z. Hassan, and Z. Ismail, "Recent advances in Ti-6Al-4V additively manufactured by selective laser melting for biomedical implants: Prospect development," J. Alloys Compd., vol. 896, p. 163072, 2022. https://doi.org/10.1016/j.jallcom.2021.163072

[9] T. A. Mukalay, J. A. Trimble, K. Mpofu, and R. Muvunzi, "A systematic review of process uncertainty in Ti6Al4V-selective laser melting," CIRP J. Manuf. Sci. Technol., vol. 36, pp. 185-212, 2022. https://doi.org/10.1016/j.cirpj.2021.12.005

[10] M. Tang, L. Zhang, and N. Zhang, "Microstructural evolution, mechanical and tribological properties of TiC/Ti6Al4V composites with unique microstructure prepared by SLM," Mater. Sci. Eng. A, vol. 814, no. April 2020, p. 141187, 2021. https://doi.org/10.1016/j.msea.2021.141187

[11] H. Attar et al., "Comparison of wear properties of commercially pure titanium prepared by selective laser melting and casting processes," Mater. Lett., vol. 142, pp. 38-41, 2015. https://doi.org/10.1016/j.matlet.2014.11.156

[12] X. Liang, P. Du, S. Li, and C. Zhang, "Tribological properties of additive manufactured Ti6Al4V against cemented carbide under dry sliding conditions," Tribol. Int., vol. 167, no. September 2021, p. 107358, 2022. https://doi.org/10.1016/j.triboint.2021.107358

[13] H. Liu, X. Huang, S. Huang, L. Qiao, and Y. Yan, "Anisotropy of wear and tribocorrosion properties of L-PBF Ti6Al4V," J. Mater. Res. Technol., vol. 25, pp. 2690-2701, 2023. https://doi.org/10.1016/j.jmrt.2023.06.128

[14] S. Kaur, K. Ghadirinejad, and R. H. Oskouei, "An overview on the tribological performance of titanium alloys with surface modifications for biomedical applications," Lubricants, vol. 7, no. 8, 2019. https://doi.org/10.3390/lubricants7080065

[15] R. M. Mahamood, T.-C. Jen, S. A. Akinlabi, S. Hassan, and E. T. Akinlabi, "Tribology of additively manufactured titanium alloy for medical implant," in Tribology of Additively Manufactured Materials, Elsevier, 2022, pp. 267-288

[16] A. K. Pandey, A. Kumar, R. Kumar, R. K. Gautam, and C. K. Behera, "Tribological performance of SS 316L, commercially pure Titanium, and Ti6Al4V in different solutions for biomedical applications," Mater. Today Proc., vol. 78, pp. A1-A8, 2022. https://doi.org/10.1016/j.matpr.2023.03.736

[17] Ş. Şirin, S. Akıncıoğlu, M. K. Gupta, T. Kıvak, and N. Khanna, "A tribological performance of vegetable-based oil combined with GNPs and hBN nanoparticles on the friction-wear tests of titanium grade 2," Tribol. Int., vol. 181, no. January, 2023. https://doi.org/10.1016/j.triboint.2023.108314
[7] C. Y. Yap et al., "Review of selective laser melting: Materials and applications," Appl. Phys. Rev., vol. 2, no. 4, 2015. https://doi.org/10.1063/1.4935926

[8] A. N. Aufa, M. Z. Hassan, and Z. Ismail, "Recent advances in Ti-6Al-4V additively manufactured by selective laser melting for biomedical implants: Prospect development," J. Alloys Compd., vol. 896, p. 163072, 2022. https://doi.org/10.1016/j.jallcom.2021.163072

[9] T. A. Mukalay, J. A. Trimble, K. Mpofu, and R. Muvunzi, "A systematic review of process uncertainty in Ti6Al4V-selective laser melting," CIRP J. Manuf. Sci. Technol., vol. 36, pp. 185-212, 2022. https://doi.org/10.1016/j.cirpj.2021.12.005

[10] M. Tang, L. Zhang, and N. Zhang, "Microstructural evolution, mechanical and tribological properties of TiC/Ti6Al4V composites with unique microstructure prepared by SLM," Mater. Sci. Eng. A, vol. 814, no. April 2020, p. 141187, 2021. https://doi.org/10.1016/j.msea.2021.141187

[11] H. Attar et al., "Comparison of wear properties of commercially pure titanium prepared by selective laser melting and casting processes," Mater. Lett., vol. 142, pp. 38-41, 2015. https://doi.org/10.1016/j.matlet.2014.11.156

[12] X. Liang, P. Du, S. Li, and C. Zhang, "Tribological properties of additive manufactured Ti6Al4V against cemented carbide under dry sliding conditions," Tribol. Int., vol. 167, no. September 2021, p. 107358, 2022. https://doi.org/10.1016/j.triboint.2021.107358

[13] H. Liu, X. Huang, S. Huang, L. Qiao, and Y. Yan, "Anisotropy of wear and tribocorrosion properties of L-PBF Ti6Al4V," J. Mater. Res. Technol., vol. 25, pp. 2690-2701, 2023. https://doi.org/10.1016/j.jmrt.2023.06.128

[14] S. Kaur, K. Ghadirinejad, and R. H. Oskouei, "An overview on the tribological performance of titanium alloys with surface modifications for biomedical applications," Lubricants, vol. 7, no. 8, 2019. https://doi.org/10.3390/lubricants7080065

[15] R. M. Mahamood, T.-C. Jen, S. A. Akinlabi, S. Hassan, and E. T. Akinlabi, "Tribology of additively manufactured titanium alloy for medical implant," in Tribology of Additively Manufactured Materials, Elsevier, 2022, pp. 267-288

[16] A. K. Pandey, A. Kumar, R. Kumar, R. K. Gautam, and C. K. Behera, "Tribological performance of SS 316L, commercially pure Titanium, and Ti6Al4V in different solutions for biomedical applications," Mater. Today Proc., vol. 78, pp. A1-A8, 2022. https://doi.org/10.1016/j.matpr.2023.03.736

[17] Ş. Şirin, S. Akıncıoğlu, M. K. Gupta, T. Kıvak, and N. Khanna, "A tribological performance of vegetable-based oil combined with GNPs and hBN nanoparticles on the friction-wear tests of titanium grade 2," Tribol. Int., vol. 181, no. January, 2023. https://doi.org/10.1016/j.triboint.2023.108314
# Metallization of Vitrimers by cold spray: A preliminary study

VISCUSI Antonio<fnref n="1" />,*, PERNA Alessia Serena<fnref n="1" />,†, MARTONE Alfonso<fnref n="2" />,†, PALMIERI Barbara<fnref n="2" />,†, CILENTO Fabrizia<fnref n="2" />,‡, AMENDOLA Eugenio<fnref n="2" />,†, TUCCI Fausto<fnref n="3" />,†, BORRELLI Domenico<fnref n="4" />,†, CARAVIELLO Antonio<fnref n="4" />,†, SICIGNANO Nicola<fnref n="4" />,†

**Keywords:** Cold Spray Technology, Vitrimer Polymer, Metallization

**Abstract.** Cold spray is a promising solution for the production of metallic coatings on polymer substrates. However, the adhesion mechanism between the impacting particle and the polymer in cold spraying depends on the chemical structure of the polymer itself. In this scenario, the emerging vitrimer polymers, which were proved to combine enhanced mechanical and chemical performances with abilities to be healed, welded, reprocessed, and recycled and that show a ductile behavior when exposed to given operating temperatures, can be particularly suitable for being functionalized by cold spray. Therefore, the aim of this work is to experimentally prove the feasibility of the metallization process via cold spray of epoxy vitrimer substrates. For this purpose, an epoxy vitrimer formulation was studied and experimentally characterized. Vitrimer-based panels were manufactured and used as substrates for the surface metallization. A low-pressure cold spray facility was used for the deposition of aluminum alloy particles. Microscope analyses were carried out for the characterization of the samples.

**Introduction**

It is known from literature that the relatively new additive cold spray technique (CS) is the best solution for the manufacturing of metallic coatings on polymer-based substrates [1]. Compared to traditional thermal spray technologies, cold spray offers benefits in terms of efficiency in the use of resources which generate economic and environmental advantages contributing to sustainability [2,3]. Some special characteristics such as electrical and thermal conductivity, erosion and radiation protection, wear and poor flame resistance can be enhanced through metallization of polymers [4,5].

When cold spraying metal on polymer, the only possible bonding mechanism between the particle and the substrate is the mechanical interlocking and the adhesion depends on the chemical structure of the polymer itself [6,7]. In this regard, Ganesan et al. [8] found that the particles can bond with the surrounding polymer material only when they penetrate the thermoplastic substrate profoundly. On the contrary, the fragile behavior of thermosetting substrates makes coating formation and growth difficult, if not impossible. In the case of thermosets, the coatings process

<figure><img src="image_184.png" /></figure>

<footnote n="1">¹Department of Chemical, Materials and Industrial Production Engineering, University of Naples Federico II, Piazzale Tecchio 80, Naples, 80125, Italy</footnote>

<footnote n="2">²National Research Council of Italy (IPCB-CNR), Institute for Polymers, Composites and Biomaterials, Portici, 80055, Italy</footnote>

<footnote n="3">³Department of Industrial Engineering, University of Salerno, Via Giovanni Paolo II 132, Fisciano, Italy</footnote>

<footnote n="4">⁴Sòphia High Tech S.r.l., Via Malatesta 30 A, 80049 - Somma Vesuviana (NA) - Italy</footnote>

<footnote>ªantonio.viscusi@unina.it, ᵇalessiaserena.perna@unina.it, ᶜalfonso.martone@cnr.it,</footnote>

<footnote>dbarbara.palmieri@ipcb.cnr.it, efabrizia.cilento@ipcb.cnr.it, ᵈeugenio.amendola@cnr.it,</footnote>

<footnote>gftucci@unisa.it, hdomenico.borrelli@sophiahightech.com,</footnote>

<footnote>iantonio.caraviello@sophiahightech.com, lnicola.sicignano@sophiahightech.com</footnote>
# Metallization of Vitrimers by cold spray: A preliminary study

VISCUSI Antonio<fnref n="1" />,*, PERNA Alessia Serena<fnref n="1" />,†, MARTONE Alfonso<fnref n="2" />,†, PALMIERI Barbara<fnref n="2" />,†, CILENTO Fabrizia<fnref n="2" />,‡, AMENDOLA Eugenio<fnref n="2" />,†, TUCCI Fausto<fnref n="3" />,†, BORRELLI Domenico<fnref n="4" />,†, CARAVIELLO Antonio<fnref n="4" />,†, SICIGNANO Nicola<fnref n="4" />,†

**Keywords:** Cold Spray Technology, Vitrimer Polymer, Metallization

**Abstract.** Cold spray is a promising solution for the production of metallic coatings on polymer substrates. However, the adhesion mechanism between the impacting particle and the polymer in cold spraying depends on the chemical structure of the polymer itself. In this scenario, the emerging vitrimer polymers, which were proved to combine enhanced mechanical and chemical performances with abilities to be healed, welded, reprocessed, and recycled and that show a ductile behavior when exposed to given operating temperatures, can be particularly suitable for being functionalized by cold spray. Therefore, the aim of this work is to experimentally prove the feasibility of the metallization process via cold spray of epoxy vitrimer substrates. For this purpose, an epoxy vitrimer formulation was studied and experimentally characterized. Vitrimer-based panels were manufactured and used as substrates for the surface metallization. A low-pressure cold spray facility was used for the deposition of aluminum alloy particles. Microscope analyses were carried out for the characterization of the samples.

**Introduction**

It is known from literature that the relatively new additive cold spray technique (CS) is the best solution for the manufacturing of metallic coatings on polymer-based substrates [1]. Compared to traditional thermal spray technologies, cold spray offers benefits in terms of efficiency in the use of resources which generate economic and environmental advantages contributing to sustainability [2,3]. Some special characteristics such as electrical and thermal conductivity, erosion and radiation protection, wear and poor flame resistance can be enhanced through metallization of polymers [4,5].

When cold spraying metal on polymer, the only possible bonding mechanism between the particle and the substrate is the mechanical interlocking and the adhesion depends on the chemical structure of the polymer itself [6,7]. In this regard, Ganesan et al. [8] found that the particles can bond with the surrounding polymer material only when they penetrate the thermoplastic substrate profoundly. On the contrary, the fragile behavior of thermosetting substrates makes coating formation and growth difficult, if not impossible. In the case of thermosets, the coatings process

<figure><img src="image_182.png" /></figure>

<footnote n="1">¹Department of Chemical, Materials and Industrial Production Engineering, University of Naples Federico II, Piazzale Tecchio 80, Naples, 80125, Italy</footnote>

<footnote n="2">²National Research Council of Italy (IPCB-CNR), Institute for Polymers, Composites and Biomaterials, Portici, 80055, Italy</footnote>

<footnote n="3">³Department of Industrial Engineering, University of Salerno, Via Giovanni Paolo II 132, Fisciano, Italy</footnote>

<footnote n="4">⁴Sòphia High Tech S.r.l., Via Malatesta 30 A, 80049 - Somma Vesuviana (NA) - Italy</footnote>

<footnote>ªantonio.viscusi@unina.it, ᵇalessiaserena.perna@unina.it, ᶜalfonso.martone@cnr.it, ᵈbarbara.palmieri@ipcb.cnr.it, ᵉfabrizia.cilento@ipcb.cnr.it, ᵃeugenio.amendola@cnr.it, ᵈftucci@unisa.it, ᵇdomenico.borrelli@sophiahightech.com, ᵈantonio.caraviello@sophiahightech.com, ᵈnicola.sicignano@sophiahightech.com</footnote>
is even more challenging since degradation rather than softening happens at elevated temperatures. Several solutions were proposed in literature with the scope to overcome this issue for thermosets, the interlayer concept [9], or the PLA treatment [10] represent the promising methods; however, none of these solutions are based on new formulations for thermoset polymers fitting the cold spray requirements.

In this scenario, the emerging class of vitrimers combines the structural performances and thermal stability of thermosets with the flow ability of thermoplastics. Epoxy vitrimers and their composites have the ability to be reprocessed, welded and easily recycled [11], therefore they are ideal candidates to be functionalized by cold spraying. In fact, epoxy vitrimers exhibit different behavior at specific temperatures: below the topology freezing transition temperature ($T_v$) behave like traditional thermosets, with good mechanical and thermal properties. While above the $T_v$, they can flow like viscoelastic fluid due to the topology rearrangement induced by dynamic exchange reaction (i.e. transesterification reactions) [12,13]. The occurrence of the transesterification reaction is governed by the addition of a catalyst; several catalysts, including organocatalysts and metal-containing compounds, have been implemented to promote transesterification in epoxy vitrimer production [14]. Zinc ions ($Zn^{2+}$) significantly accelerates esterification by opening anhydride ring and yielding monoester and carboxylic acids [15].

Based on these premises, the aim of this work is to experimentally prove the feasibility of the metallization process via cold spray of epoxy vitrimer substrates. In other words, since the cold spray deposition on thermosetting material was proved to be difficult, if not impossible today, the possibility to coat epoxy vitrimer polymers was investigated. To achieve the scope, an epoxy vitrimer was formulated modifying a commercially available epoxy system by adding zinc acetate ($Zn(Ac)_2$) and the chemical-physical properties of the vitrimers were studied. Vitrimer-based panels were manufactured and used as substrates for surface metallization. A low-pressure cold spray facility was used for spraying aluminum alloy particles. Microscope analyses were carried out for the characterization of the samples.

## Materials and methods

In this section, the materials, the methods, and the technologies used to produce both the epoxy vitrimer panels, used as substrates, and the aluminum coatings, will be presented; the experimental tests used for the characterization of the samples proving the feasibility of the process developed, will be also described.

The resin used in this study consists of a mixture of Bisphenol A diglycidyl ether (DGEBA) epoxy resin, curing agent tetrahydro-methyl phthalic anhydride (THMPA), and of a 2,4,6-tris (dimethylaminomethyl)phenol as catalyst, all provided by Huntsman Corporation. The standard epoxy system has been modified by adding anhydrous zinc acetate $Zn(Ac)_2$ as transesterification catalyst, to promote the vitrimeric behavior.

The formulation has been prepared whit a stoichiometric ratio of epoxy/acyl equals to 1, and a 10% wt. of $Zn(Ac)_2$ with respect to the total acyl groups. The manufacturing process is illustrated in Fig. 1: (i) hand-mixing of epoxy resin and $Zn(Ac)_2$ fine powder; (ii) adding of cross-linking Anhydride and mixed by a planetary centrifugal mixer (THINKY mixer ARV 310) under vacuum at room temperature; (iii) addition of the accelerator (Amine) and further mixing in the THINKY to obtain a homogenous mixture; (iv) casting in Teflon molds and (v) cured in the oven at 120°C for 1 h 30 min and then post cured at 140°C for two hours.

The coating was produced by using micron sized AlSi10Mg powders with a particle mean size of 25 µm, provided by LPW South Group. The low-pressure cold spray facility (DYMET 423) fed by nitrogen as carrier gas was used for the deposition. The spray gun was attached to a cartesian robot for the control and the repeatability of the deposition process.
is even more challenging since degradation rather than softening happens at elevated temperatures. Several solutions were proposed in literature with the scope to overcome this issue for thermosets, the interlayer concept [9], or the PLA treatment [10] represent the promising methods; however, none of these solutions are based on new formulations for thermoset polymers fitting the cold spray requirements.

In this scenario, the emerging class of vitrimers combines the structural performances and thermal stability of thermosets with the flow ability of thermoplastics. Epoxy vitrimers and their composites have the ability to be reprocessed, welded and easily recycled [11], therefore they are ideal candidates to be functionalized by cold spraying. In fact, epoxy vitrimers exhibit different behavior at specific temperatures: below the topology freezing transition temperature ($T_v$) behave like traditional thermosets, with good mechanical and thermal properties. While above the $T_v$, they can flow like viscoelastic fluid due to the topology rearrangement induced by dynamic exchange reaction (i.e. transesterification reactions) [12,13]. The occurrence of the transesterification reaction is governed by the addition of a catalyst; several catalysts, including organocatalysts and metal-containing compounds, have been implemented to promote transesterification in epoxy vitrimer production [14]. Zinc ions ($Zn^{2+}$) significantly accelerates esterification by opening anhydride ring and yielding monoester and carboxylic acids [15].

Based on these premises, the aim of this work is to experimentally prove the feasibility of the metallization process via cold spray of epoxy vitrimer substrates. In other words, since the cold spray deposition on thermosetting material was proved to be difficult, if not impossible today, the possibility to coat epoxy vitrimer polymers was investigated. To achieve the scope, an epoxy vitrimer was formulated modifying a commercially available epoxy system by adding zinc acetate ($Zn(Ac)_2$) and the chemical-physical properties of the vitrimers were studied. Vitrimer-based panels were manufactured and used as substrates for surface metallization. A low-pressure cold spray facility was used for spraying aluminum alloy particles. Microscope analyses were carried out for the characterization of the samples.

## Materials and methods

In this section, the materials, the methods, and the technologies used to produce both the epoxy vitrimer panels, used as substrates, and the aluminum coatings, will be presented; the experimental tests used for the characterization of the samples proving the feasibility of the process developed, will be also described.

The resin used in this study consists of a mixture of Bisphenol A diglycidyl ether (DGEBA) epoxy resin, curing agent tetrahydro-methyl phthalic anhydride (THMPA), and of a 2,4,6-tris (dimethylaminomethyl)phenol as catalyst, all provided by Huntsman Corporation. The standard epoxy system has been modified by adding anhydrous zinc acetate $Zn(Ac)_2$ as transesterification catalyst, to promote the vitrimeric behavior.

The formulation has been prepared whit a stoichiometric ratio of epoxy/acyl equals to 1, and a 10% wt. of $Zn(Ac)_2$ with respect to the total acyl groups. The manufacturing process is illustrated in Fig. 1: (i) hand-mixing of epoxy resin and $Zn(Ac)_2$ fine powder; (ii) adding of cross-linking Anhydride and mixed by a planetary centrifugal mixer (THINKY mixer ARV 310) under vacuum at room temperature; (iii) addition of the accelerator (Amine) and further mixing in the THINKY to obtain a homogenous mixture; (iv) casting in Teflon molds and (v) cured in the oven at 120°C for 1 h 30 min and then post cured at 140°C for two hours.

The coating was produced by using micron sized AlSi10Mg powders with a particle mean size of 25 µm, provided by LPW South Group. The low-pressure cold spray facility (DYMET 423) fed by nitrogen as carrier gas was used for the deposition. The spray gun was attached to a cartesian robot for the control and the repeatability of the deposition process.
<figure><img src="image_185.png" /><figcaption>Fig. 1. Manufacturing process of vitrimeric epoxy resin [15].</figcaption></figure>

To analyze the interaction between the cold sprayed particles and the epoxy vitrimer substrate, only one layer was sprayed on the panel [16] producing a 50 mm length metallic track. Different tracks of aluminum coatings were produced on vitrimers to set the proper cold spray process parameters.

The nozzle to substrate standoff distance (SoD) was varied between 45 and 100 mm [6,17], and then set to 70 mm after a preliminary experimental test. The gas pressure and the gun traverse speed were kept constant in all the experiments and set equals to 6 bar and 1 mm/s, respectively.

The inlet gas temperature was found to be a key factor for the deposition on polymers [18]. If on one hand it affects the particle velocity, on the other it was proved to have a noticeable influence on the physical behavior of the polymeric substrate when impacted by the particles, due to the temperature reached by the gas flow on the target surface. Two values of inlet gas temperature were chosen in this activity (T₁=200°C, T₂=300°C) in agreement with previous authors' research and literature data [6], so resulting in two types of coated vitrimeric samples (here referred as V<sub>T1</sub> and V<sub>T2</sub>). The main cold spray process parameters used are reported in Table 1.

<table><caption>Table 1. Main process parameters in cold spraying AlSi10Mg particles on epoxy vitrimer substrates.</caption><thead><tr><td>Sample</td><td>Inlet gas temperature [°C]</td><td>SoD [mm]</td><td>Gas pressure [bar]</td><td>Gun traverse speed [mm/s]</td></tr></thead><tbody><tr><td>V<sub>T1</sub></td><td>200</td><td>70</td><td>6</td><td>1</td></tr><tr><td>V<sub>T2</sub></td><td>300</td><td>70</td><td>6</td><td>1</td></tr></tbody></table>

The epoxy vitrimer was characterized through creep tests performed using the DMA Q800 from TA Instrument equipped with the tension film (TF) clamp. Rectangular specimens of 10x5.5x2 mm³ were tested, at different temperatures between 70°C to 220°C, at 25°C steps to evaluate the strain rate variation from the glassy to the rubbery state.

The surface morphology of the vitrimeric panels was observed by means of confocal Olympus OLS5100 microscope in order to analyze the surface characteristics and chose the most appropriate
<figure><img src="image_183.png" /><figcaption>Fig. 1. Manufacturing process of vitrimeric epoxy resin [15].</figcaption></figure>

To analyze the interaction between the cold sprayed particles and the epoxy vitrimer substrate, only one layer was sprayed on the panel [16] producing a 50 mm length metallic track. Different tracks of aluminum coatings were produced on vitrimers to set the proper cold spray process parameters.

The nozzle to substrate standoff distance (SoD) was varied between 45 and 100 mm [6,17], and then set to 70 mm after a preliminary experimental test. The gas pressure and the gun traverse speed were kept constant in all the experiments and set equals to 6 bar and 1 mm/s, respectively.

The inlet gas temperature was found to be a key factor for the deposition on polymers [18]. If on one hand it affects the particle velocity, on the other it was proved to have a noticeable influence on the physical behavior of the polymeric substrate when impacted by the particles, due to the temperature reached by the gas flow on the target surface. Two values of inlet gas temperature were chosen in this activity (T₁=200°C, T₂=300°C) in agreement with previous authors' research and literature data [6], so resulting in two types of coated vitrimeric samples (here referred as V<sub>T1</sub> and V<sub>T2</sub>). The main cold spray process parameters used are reported in Table 1.

<table><caption>Table 1. Main process parameters in cold spraying AlSi10Mg particles on epoxy vitrimer substrates.</caption><thead><tr><td>Sample</td><td>Inlet gas temperature [°C]</td><td>SoD [mm]</td><td>Gas pressure [bar]</td><td>Gun traverse speed [mm/s]</td></tr></thead><tbody><tr><td>V<sub>T1</sub></td><td>200</td><td>70</td><td>6</td><td>1</td></tr><tr><td>V<sub>T2</sub></td><td>300</td><td>70</td><td>6</td><td>1</td></tr></tbody></table>

The epoxy vitrimer was characterized through creep tests performed using the DMA Q800 from TA Instrument equipped with the tension film (TF) clamp. Rectangular specimens of 10x5.5x2 mm³ were tested, at different temperatures between 70°C to 220°C, at 25°C steps to evaluate the strain rate variation from the glassy to the rubbery state.

The surface morphology of the vitrimeric panels was observed by means of confocal Olympus OLS5100 microscope in order to analyze the surface characteristics and chose the most appropriate
face upon which the deposition can be carried out effectively. Note that the terms back face and front face will be used here to indicate the “as-is” and “to be coated” surface of the vitrimeric panel, respectively.

The metallized samples were observed through Hirox ST-AS microscope to show the coating obtained and analyze in more details the mechanisms ruling adhesion during deposition of aluminum particles on epoxy vitrimer substrates.

## Results and discussion

The vitrimeric behavior of the resin can be highlighted by discussing the results from the creep experiments shown in Fig. 2. In contrast with the epoxy standard behavior, the vitrimeric system displayed a significant increase in the strain rate under a constant load at temperature above 170°C, which can be identified as the topological vitrimeric transition temperature (T~v~).

Inset pictures reproduce the creep test for the standard epoxy and the vitrimer at two characteristic temperatures, 95°C and 170°C. The temperature has been chosen as representative of the dynamic change in the vitrimer and as an evidence of the actual flow happened. At 95°C both the systems act as a thermoset, the deformation rate is negligible, while at 170°C the standard epoxy keeps its behavior conversely the vitrimer starts to flow (deformation rate increases).

<figure><img src="image_186.png" /><figcaption>Fig. 2. Creep curves at different temperatures of the standard epoxy (black diamond shape) and vitrimeric system (dotted red line).</figcaption></figure>

The results from the confocal microscope observations relative to back (which is the face in contact with the Teflon mold in Fig. 1) and front face of the vitrimeric panel are reported in Fig. 3 and Fig. 4, respectively. In particular, the scope of this preliminary analysis was to study the surface morphology of the target surface and identify the most suitable face to be cold sprayed.

It can be observed from Fig. 3a that the morphology of the back face is characterized by deep craters and grooves that are much larger than the mean diameter of the feedstock particles. These characteristics should not promote the sticking of the particles impacting the substrate. In fact, as known from literature, it seems to be necessary that the valleys of the surface texturing of the polymer used as substrate had a comparable dimension to the feedstock powder diameter [19], this is a key factor for the particle interlocking between the substrate and the powders, as the latter may impinge into the valleys becoming entangled. The sketch in the right side of the figure (Fig. 3b)
face upon which the deposition can be carried out effectively. Note that the terms back face and front face will be used here to indicate the “as-is” and “to be coated” surface of the vitrimeric panel, respectively.

The metallized samples were observed through Hirox ST-AS microscope to show the coating obtained and analyze in more details the mechanisms ruling adhesion during deposition of aluminum particles on epoxy vitrimer substrates.

## Results and discussion

The vitrimeric behavior of the resin can be highlighted by discussing the results from the creep experiments shown in Fig. 2. In contrast with the epoxy standard behavior, the vitrimeric system displayed a significant increase in the strain rate under a constant load at temperature above 170°C, which can be identified as the topological vitrimeric transition temperature (T~v~).

Inset pictures reproduce the creep test for the standard epoxy and the vitrimer at two characteristic temperatures, 95°C and 170°C. The temperature has been chosen as representative of the dynamic change in the vitrimer and as an evidence of the actual flow happened. At 95°C both the systems act as a thermoset, the deformation rate is negligible, while at 170°C the standard epoxy keeps its behavior conversely the vitrimer starts to flow (deformation rate increases).

<figure><img src="image_184.png" /><figcaption>Fig. 2. Creep curves at different temperatures of the standard epoxy (black diamond shape) and vitrimeric system (dotted red line).</figcaption></figure>

The results from the confocal microscope observations relative to back (which is the face in contact with the Teflon mold in Fig. 1) and front face of the vitrimeric panel are reported in Fig. 3 and Fig. 4, respectively. In particular, the scope of this preliminary analysis was to study the surface morphology of the target surface and identify the most suitable face to be cold sprayed.

It can be observed from Fig. 3a that the morphology of the back face is characterized by deep craters and grooves that are much larger than the mean diameter of the feedstock particles. These characteristics should not promote the sticking of the particles impacting the substrate. In fact, as known from literature, it seems to be necessary that the valleys of the surface texturing of the polymer used as substrate had a comparable dimension to the feedstock powder diameter [19], this is a key factor for the particle interlocking between the substrate and the powders, as the latter may impinge into the valleys becoming entangled. The sketch in the right side of the figure (Fig. 3b)
clearly show that the particle does not impact normally to the substrate under these surface conditions, without promoting the mechanical interlocking, so rebounding from the target surface. This mechanism is known from literature in cold spraying [20].

<figure><img src="image_187.png" /><figcaption>Fig. 3. a) Surface morphology of back face of vitrimeric panel. b) Illustrative sketch of particle rebounding mechanism.</figcaption></figure>

Based on the results above, the powders have been cold sprayed on the front face of the panel to benefit the effects of the smoothed surface, as shown in Fig. 4a. An exemplificative picture of the coated epoxy vitrimer sample is displayed in Fig. 4b.

The metallized samples were observed through Hirox microscope, and the results obtained are discussed below. A specimen of the coated epoxy vitrimer, labelled as V<sub>T2</sub> in Table 1, is shown in Fig. 5.

Firstly, it can be observed that the feasibility of the process was proved with the metallic particles adhered the polymeric surface. This result shows that an epoxy-based polymer can be effectively coated by cold spraying aluminum particles when using a given vitrimeric formulation. However, some considerations need to be carried out regarding the set of the proper cold spray process parameters.

When cold spraying at T=300°C, the temperature of the gas, which expands through the nozzle, reduces, and can achieve also 200°C onto the substrate (as measured via k-type thermocouple), under the set process conditions. It is clear that the temperature on the surface of the polymer remains much higher than the vitrimeric transition temperature (see the Fig. 2 above), the material tends to soften and deform like thermoplastics [21]. The result is that the material flows and the particles impacting with relatively high velocity stick with the polymeric substrate due to the mechanical interlocking mechanism [22]. However, the softened material showed an extensive groove on the area interested by the spray deposition characterized by a depth larger than 25% of the panel thickness, that may affect the mechanical properties of the panel. This is clearly visible by looking both the macrograph and the relative profile in Fig. 5a and 5b, respectively. The observed phenomenon is related to high temperature of the impinging gas that results in a significant shoot peening effect of the upcoming particles impacting at relatively high velocity on a very soft polymer, causing possible defects and erosion of the substrate.
clearly show that the particle does not impact normally to the substrate under these surface conditions, without promoting the mechanical interlocking, so rebounding from the target surface. This mechanism is known from literature in cold spraying [20].

<figure><img src="image_185.png" /><figcaption>Fig. 3. a) Surface morphology of back face of vitrimeric panel. b) Illustrative sketch of particle rebounding mechanism.</figcaption></figure>

Based on the results above, the powders have been cold sprayed on the front face of the panel to benefit the effects of the smoothed surface, as shown in Fig. 4a. An exemplificative picture of the coated epoxy vitrimer sample is displayed in Fig. 4b.

The metallized samples were observed through Hirox microscope, and the results obtained are discussed below. A specimen of the coated epoxy vitrimer, labelled as V<sub>T2</sub> in Table 1, is shown in Fig. 5.

Firstly, it can be observed that the feasibility of the process was proved with the metallic particles adhered the polymeric surface. This result shows that an epoxy-based polymer can be effectively coated by cold spraying aluminum particles when using a given vitrimeric formulation. However, some considerations need to be carried out regarding the set of the proper cold spray process parameters.

When cold spraying at T=300°C, the temperature of the gas, which expands through the nozzle, reduces, and can achieve also 200°C onto the substrate (as measured via k-type thermocouple), under the set process conditions. It is clear that the temperature on the surface of the polymer remains much higher than the vitrimeric transition temperature (see the Fig. 2 above), the material tends to soften and deform like thermoplastics [21]. The result is that the material flows and the particles impacting with relatively high velocity stick with the polymeric substrate due to the mechanical interlocking mechanism [22]. However, the softened material showed an extensive groove on the area interested by the spray deposition characterized by a depth larger than 25% of the panel thickness, that may affect the mechanical properties of the panel. This is clearly visible by looking both the macrograph and the relative profile in Fig. 5a and 5b, respectively. The observed phenomenon is related to high temperature of the impinging gas that results in a significant shoot peening effect of the upcoming particles impacting at relatively high velocity on a very soft polymer, causing possible defects and erosion of the substrate.
<figure><img src="image_188.png" /><figcaption>Fig. 4. a) Result from confocal microscope scan of front face of vitrimeric panel. b) Specimen of cold sprayed epoxy vitrimer.</figcaption></figure>

<figure><img src="image_189.png" /><figcaption>Fig. 5. Results from Hirox observations with the large groove highlighted in yellow: a) macrograph of $V_{T2}$ coated panel; b) measurement of height profile.</figcaption></figure>

The results of the deposition process at T=200°C on epoxy vitrimer sample ($V_{T1}$) are shown in Fig. 6. It can be observed the single layer of AlSi10Mg particles proving the effectiveness of the deposition process (Fig. 6a). Contrary to $V_{T2}$ sample, the groove left by the spraying process is negligible, being about 4% in depth of the panel thickness (Fig. 6b). The reason is that, under these spraying conditions, the temperature of the gas at the impact surface (which was measured to be about 140°C) is a little below the vitrimeric transition temperature, with the epoxy polymeric material that starts to behave like thermoplastics (see the Fig. 2 above). The result is that the
<figure><img src="image_186.png" /><figcaption>Fig. 4. a) Result from confocal microscope scan of front face of vitrimeric panel. b) Specimen of cold sprayed epoxy vitrimer.</figcaption></figure>

<figure><img src="image_187.png" /><figcaption>Fig. 5. Results from Hirox observations with the large groove highlighted in yellow: a) macrograph of $V_{T2}$ coated panel; b) measurement of height profile.</figcaption></figure>

The results of the deposition process at T=200°C on epoxy vitrimer sample ($V_{T1}$) are shown in Fig. 6. It can be observed the single layer of AlSi10Mg particles proving the effectiveness of the deposition process (Fig. 6a). Contrary to $V_{T2}$ sample, the groove left by the spraying process is negligible, being about 4% in depth of the panel thickness (Fig. 6b). The reason is that, under these spraying conditions, the temperature of the gas at the impact surface (which was measured to be about 140°C) is a little below the vitrimeric transition temperature, with the epoxy polymeric material that starts to behave like thermoplastics (see the Fig. 2 above). The result is that the
particles impact the relatively rigid surface and penetrate the like-thermoplastic substrate promoting adhesion, without inducing significant defects and grooves within the material.

<figure><img src="image_190.png" /><figcaption>Fig. 6. Results from Hirox observations with the negligible groove highlighted in yellow: a) macrograph of $V_{TI}$ coated panel; b) measurement of height profile.</figcaption></figure>

## Conclusions

Based on the experimental results presented and discussed in the previous sections, the following conclusions could be drawn:

* Firstly, the cold spray deposition process was proved to be an effective technique for the metallization of epoxy vitrimers if the spraying parameters are properly set (T=200°C, SoD=70 mm, Gas pressure=6 bar).

* The epoxy vitrimer was proved to be a promising polymer to be metallized via cold spray thanks to its thermoplastic-like behavior above the $T_v$, maintaining the mechanical and thermal characteristics of the thermosetting below this temperature.

* The surface morphology of the vitrimer was proved to affect the cold spray deposition process. Particularly, it is necessary that the valleys of the surface texturing of the polymer used as substrate had a comparable dimension to the feedstock powder diameter, so promoting the mechanical interlocking mechanism.

* The inlet gas temperature was found to be a key factor for an effective deposition on vitrimers. Particularly, the temperature of the gas at the impact surface of vitrimer should be close to the vitrimeric transition temperature. Relatively higher gas temperature values may lead to defects and grooves within the material.

## Acknowledgements

This study was carried out within the "FRONTIERS: Functionally cold spRayed cOatiNgs for viTrimer composItes in lEo-space hybRid Structures" (grant number: E53D23017120001). The authors would like to thank the Ministry for Universities and Research (MUR) for funding the research activities through the Research Projects of Significant National Interest (PRIN) 2022 PNRR (funded by the European Union – NextGenerationEU).
particles impact the relatively rigid surface and penetrate the like-thermoplastic substrate promoting adhesion, without inducing significant defects and grooves within the material.

<figure><img src="image_188.png" /><figcaption>Fig. 6. Results from Hirox observations with the negligible groove highlighted in yellow: a) macrograph of $V_{T1}$ coated panel; b) measurement of height profile.</figcaption></figure>

## Conclusions

Based on the experimental results presented and discussed in the previous sections, the following conclusions could be drawn:

* Firstly, the cold spray deposition process was proved to be an effective technique for the metallization of epoxy vitrimers if the spraying parameters are properly set (T=200°C, SoD=70 mm, Gas pressure=6 bar).

* The epoxy vitrimer was proved to be a promising polymer to be metallized via cold spray thanks to its thermoplastic-like behavior above the $T_v$, maintaining the mechanical and thermal characteristics of the thermosetting below this temperature.

* The surface morphology of the vitrimer was proved to affect the cold spray deposition process. Particularly, it is necessary that the valleys of the surface texturing of the polymer used as substrate had a comparable dimension to the feedstock powder diameter, so promoting the mechanical interlocking mechanism.

* The inlet gas temperature was found to be a key factor for an effective deposition on vitrimers. Particularly, the temperature of the gas at the impact surface of vitrimer should be close to the vitrimeric transition temperature. Relatively higher gas temperature values may lead to defects and grooves within the material.

## Acknowledgements

This study was carried out within the "FRONTIERS: Functionally cold spRayed cOatiNgs for viTrimer composItes in leo-space hybRid Structures" (grant number: E53D23017120001). The authors would like to thank the Ministry for Universities and Research (MUR) for funding the research activities through the Research Projects of Significant National Interest (PRIN) 2022 PNRR (funded by the European Union – NextGenerationEU).
## References

[1] R. Lupoi, W. O'Neill, Deposition of metallic coatings on polymer surfaces using cold spray, Surf. Coatings Technol. 205 (2010) 2167-2173.
https://doi.org/10.1016/j.SURFCOAT.2010.08.128

[2] A. Faheem, A. Tyagi, S.M. Pandey, F. Hasan, Q. Murtaza, A sustainable ecofriendly additive manufacturing approach of repairing and coating on the substrate: cold spray, Aust. J. Mech. Eng. 21 (2023) 1624-1641. https://doi.org/10.1080/14484846.2021.2023379

[3] M. Durante, L. Boccarusso, D. De Fazio, A. Langella, Circular cutting strategy for drilling of carbon fiber-reinforced plastics (CFRs), Mater. Manuf. Process. 34 (2019) 554-566.
https://doi.org/10.1080/10426914.2019.1566615

[4] A. Viscusi, M. Durante, A. Astarita, L. Boccarusso, L. Carrino, A.S. Perna, Experimental Evaluation of Metallic Coating on Polymer by Cold Spray, Procedia Manuf. 47 (2020) 761-765.
https://doi.org/10.1016/j.promfg.2020.04.232

[5] R. Melentiev, N. Yu, G. Lubineau, Polymer metallization via cold spray additive manufacturing: A review of process control, coating qualities, and prospective applications, Addit. Manuf. 48 (2021) 102459. https://doi.org/10.1016/j.addma.2021.102459

[6] R. Della Gatta, A.S. Perna, A. Viscusi, G. Pasquino, A. Astarita, Cold spray deposition of metallic coatings on polymers: a review, J. Mater. Sci. 57 (2022) 27-57.
https://doi.org/10.1007/s10853-021-06561-2

[7] L. Boccarusso, F. Pinto, S. Cuomo, D. De Fazio, K. Myronidis, M. Durante, M. Meo, Design, Manufacturing, and Characterization of Hybrid Carbon/Hemp Sandwich Panels, J. Mater. Eng. Perform. 31 (2022) 769-785. https://doi.org/10.1007/S11665-021-06186-1/FIGURES/22

[8] A. Ganesan, M. Yamada, M. Fukumoto, Cold Spray Coating Deposition Mechanism on the Thermoplastic and Thermosetting Polymer Substrates, J. Therm. Spray Technol. 22 (2013) 1275-1282. https://doi.org/10.1007/s11666-013-9984-x

[9] H. Parmar, R. Della Gatta, A. Viscusi, F. Tucci, A. Astarita, P. Carlone, Cold Spray Metallization of Hybrid Thermoplastic - Thermoset Fiber Reinforced Composite, Key Eng. Mater. 926 (2022) 1293-1302. https://doi.org/10.4028/p-44vhs8

[10] A. Viscusi, V. Antonucci, L. Carrino, R. Della Gatta, V. Lopresto, I. Papa, A.S. Perna, M.R. Ricciardi, A. Astarita, Manufacturing of an innovative composite structure: Design, manufacturing and impact behaviour, Compos. Struct. 250 (2020) 112637.
https://doi.org/10.1016/j.compstruct.2020.112637

[11] H. Memon, Y. Wei, C. Zhu, Recyclable and reformable epoxy resins based on dynamic covalent bonds – Present, past, and future, Polym. Test. 105 (2022) 107420.
https://doi.org/10.1016/j.polymertesting.2021.107420

[12] M. Capelot, M.M. Unterlass, F. Tournilhac, L. Leibler, Catalytic control of the vitrimer glass transition, ACS Macro Lett. 1 (2012) 789-792.
https://doi.org/10.1021/MZ300239F/SUPPL_FILE/MZ300239F_SI_001.PDF

[13] D. Montarnal, M. Capelot, F. Tournilhac, L. Leibler, Silica-like malleable materials from permanent organic networks, Science (80-.). 334 (2011) 965-968.
https://doi.org/10.1126/SCIENCE.1212648/SUPPL_FILE/MONTARNAL.SOM.PDF

[14] Y. Yang, Y. Xu, Y. Ji, Y. Wei, Functional epoxy vitrimers and composites, Prog. Mater. Sci. 120 (2021) 100710. https://doi.org/10.1016/j.pmatsci.2020.100710
## References

[1] R. Lupoi, W. O'Neill, Deposition of metallic coatings on polymer surfaces using cold spray, Surf. Coatings Technol. 205 (2010) 2167-2173.
https://doi.org/10.1016/j.SURFCOAT.2010.08.128

[2] A. Faheem, A. Tyagi, S.M. Pandey, F. Hasan, Q. Murtaza, A sustainable ecofriendly additive manufacturing approach of repairing and coating on the substrate: cold spray, Aust. J. Mech. Eng. 21 (2023) 1624-1641. https://doi.org/10.1080/14484846.2021.2023379

[3] M. Durante, L. Boccarusso, D. De Fazio, A. Langella, Circular cutting strategy for drilling of carbon fiber-reinforced plastics (CFRs), Mater. Manuf. Process. 34 (2019) 554-566.
https://doi.org/10.1080/10426914.2019.1566615

[4] A. Viscusi, M. Durante, A. Astarita, L. Boccarusso, L. Carrino, A.S. Perna, Experimental Evaluation of Metallic Coating on Polymer by Cold Spray, Procedia Manuf. 47 (2020) 761-765.
https://doi.org/10.1016/j.promfg.2020.04.232

[5] R. Melentiev, N. Yu, G. Lubineau, Polymer metallization via cold spray additive manufacturing: A review of process control, coating qualities, and prospective applications, Addit. Manuf. 48 (2021) 102459. https://doi.org/10.1016/j.addma.2021.102459

[6] R. Della Gatta, A.S. Perna, A. Viscusi, G. Pasquino, A. Astarita, Cold spray deposition of metallic coatings on polymers: a review, J. Mater. Sci. 57 (2022) 27-57.
https://doi.org/10.1007/s10853-021-06561-2

[7] L. Boccarusso, F. Pinto, S. Cuomo, D. De Fazio, K. Myronidis, M. Durante, M. Meo, Design, Manufacturing, and Characterization of Hybrid Carbon/Hemp Sandwich Panels, J. Mater. Eng. Perform. 31 (2022) 769-785. https://doi.org/10.1007/S11665-021-06186-1/FIGURES/22

[8] A. Ganesan, M. Yamada, M. Fukumoto, Cold Spray Coating Deposition Mechanism on the Thermoplastic and Thermosetting Polymer Substrates, J. Therm. Spray Technol. 22 (2013) 1275-1282. https://doi.org/10.1007/s11666-013-9984-x

[9] H. Parmar, R. Della Gatta, A. Viscusi, F. Tucci, A. Astarita, P. Carlone, Cold Spray Metallization of Hybrid Thermoplastic - Thermoset Fiber Reinforced Composite, Key Eng. Mater. 926 (2022) 1293-1302. https://doi.org/10.4028/p-44vhs8

[10] A. Viscusi, V. Antonucci, L. Carrino, R. Della Gatta, V. Lopresto, I. Papa, A.S. Perna, M.R. Ricciardi, A. Astarita, Manufacturing of an innovative composite structure: Design, manufacturing and impact behaviour, Compos. Struct. 250 (2020) 112637.
https://doi.org/10.1016/j.compstruct.2020.112637

[11] H. Memon, Y. Wei, C. Zhu, Recyclable and reformable epoxy resins based on dynamic covalent bonds – Present, past, and future, Polym. Test. 105 (2022) 107420.
https://doi.org/10.1016/j.polymertesting.2021.107420

[12] M. Capelot, M.M. Unterlass, F. Tournilhac, L. Leibler, Catalytic control of the vitrimer glass transition, ACS Macro Lett. 1 (2012) 789-792.
https://doi.org/10.1021/MZ300239F/SUPPL_FILE/MZ300239F_SI_001.PDF

[13] D. Montarnal, M. Capelot, F. Tournilhac, L. Leibler, Silica-like malleable materials from permanent organic networks, Science (80-.). 334 (2011) 965-968.
https://doi.org/10.1126/SCIENCE.1212648/SUPPL_FILE/MONTARNAL.SOM.PDF

[14] Y. Yang, Y. Xu, Y. Ji, Y. Wei, Functional epoxy vitrimers and composites, Prog. Mater. Sci. 120 (2021) 100710. https://doi.org/10.1016/j.pmatsci.2020.100710
[15] B. Palmieri, F. Cilento, E. Amendola, T. Valente, S. Dello Iacono, M. Giordano, A. Martone, Influence of Catalyst Content and Epoxy/Carboxylate Ratio on Isothermal Creep of Epoxy Vitrimers, Polym. 2023, Vol. 15, Page 3845. 15 (2023) 3845. https://doi.org/10.3390/POLYM15183845

[16] U. Prisco, A. Squillace, A. Astarita, L. Carrino, Morphology of titanium coatings deposited through single pass cold spraying, Mater. Manuf. Process. 33 (2018) 123-129. https://doi.org/10.1080/10426914.2016.1198035

[17] W.Y. Li, C. Zhang, X.P. Guo, G. Zhang, H.L. Liao, C.J. Li, C. Coddet, Effect of standoff distance on coating deposition characteristics in cold spraying, Mater. Des. 29 (2008) 297-304. https://doi.org/10.1016/j.MATDES.2007.02.005

[18] D. Giraud, F. Borit, V. Guipont, M. Jeandin, J.M. Malhaire, Metallization of a polymer using cold spray: Application to aluminum coating of polyamide 66, Proc. Int. Therm. Spray Conf. (2012) 265-270. https://doi.org/10.31399/ASM.CP.ITSC2012P0265/25442/METALLIZATION-OF-A-POLYMER-USING-COLD-SPRAY.

[19] A. Viscusi, R. Della Gatta, F. Delloro, I. Papa, A.S. Perna, A. Astarita, A novel manufacturing route for integrated 3D-printed composites and cold-sprayed metallic layer, Mater. Manuf. Process. 37 (2022) 568-581. https://doi.org/10.1080/10426914.2021.1942908

[20] C.-J. Li, W.-Y. Li, Y.-Y. Wang, H. Fukanuma, Effect of Spray Angle on Deposition Characteristics in Cold Spraying, in: B.R. Marple, C. Moreau (Eds.), Therm. Spray 2003 Proc. from Int. Therm. Spray Conf., ASM International, 2003: pp. 91-96. https://doi.org/10.31399/asm.cp.itsc2003p0091

[21] B. Krishnakumar, R.V.S.P. Sanka, W.H. Binder, V. Parthasarthy, S. Rana, N. Karak, Vitrimers: Associative dynamic covalent adaptive networks in thermoset polymers, Chem. Eng. J. 385 (2020) 123820. https://doi.org/10.1016/j.cej.2019.123820

[22] H. Che, X. Chu, P. Vo, S. Yue, Metallization of Various Polymers by Cold Spray, J. Therm. Spray Technol. 27 (2018) 169-178. https://doi.org/10.1007/s11666-017-0663-1
[15] B. Palmieri, F. Cilento, E. Amendola, T. Valente, S. Dello Iacono, M. Giordano, A. Martone, Influence of Catalyst Content and Epoxy/Carboxylate Ratio on Isothermal Creep of Epoxy Vitrimers, Polym. 2023, Vol. 15, Page 3845. 15 (2023) 3845. https://doi.org/10.3390/POLYM15183845

[16] U. Prisco, A. Squillace, A. Astarita, L. Carrino, Morphology of titanium coatings deposited through single pass cold spraying, Mater. Manuf. Process. 33 (2018) 123-129. https://doi.org/10.1080/10426914.2016.1198035

[17] W.Y. Li, C. Zhang, X.P. Guo, G. Zhang, H.L. Liao, C.J. Li, C. Coddet, Effect of standoff distance on coating deposition characteristics in cold spraying, Mater. Des. 29 (2008) 297-304. https://doi.org/10.1016/j.MATDES.2007.02.005

[18] D. Giraud, F. Borit, V. Guipont, M. Jeandin, J.M. Malhaire, Metallization of a polymer using cold spray: Application to aluminum coating of polyamide 66, Proc. Int. Therm. Spray Conf. (2012) 265-270. https://doi.org/10.31399/ASM.CP.ITSC2012P0265/25442/METALLIZATION-OF-A-POLYMER-USING-COLD-SPRAY.

[19] A. Viscusi, R. Della Gatta, F. Delloro, I. Papa, A.S. Perna, A. Astarita, A novel manufacturing route for integrated 3D-printed composites and cold-sprayed metallic layer, Mater. Manuf. Process. 37 (2022) 568-581. https://doi.org/10.1080/10426914.2021.1942908

[20] C.-J. Li, W.-Y. Li, Y.-Y. Wang, H. Fukanuma, Effect of Spray Angle on Deposition Characteristics in Cold Spraying, in: B.R. Marple, C. Moreau (Eds.), Therm. Spray 2003 Proc. from Int. Therm. Spray Conf., ASM International, 2003: pp. 91-96. https://doi.org/10.31399/asm.cp.itsc2003p0091

[21] B. Krishnakumar, R.V.S.P. Sanka, W.H. Binder, V. Parthasarthy, S. Rana, N. Karak, Vitrimers: Associative dynamic covalent adaptive networks in thermoset polymers, Chem. Eng. J. 385 (2020) 123820. https://doi.org/10.1016/j.cej.2019.123820

[22] H. Che, X. Chu, P. Vo, S. Yue, Metallization of Various Polymers by Cold Spray, J. Therm. Spray Technol. 27 (2018) 169-178. https://doi.org/10.1007/s11666-017-0663-1
# Preliminary assessment of material extrusion (MEX) for medical applications: The effect of hatch angle

SALA Francesca<fnref n="1" />,ª*, NANI Lorenzo²,ᵇ, QUARTO Mariangela<fnref n="1" />,ᶜ and D'URSO Gianluca<fnref n="1" />,ᵈ

**Keywords:** Material Extrusion, MEX, Laser Powder Bed Fusion, L-PBF, Healthcare

**Abstract.** Material extrusion (MEX) is one of the most widely used Additive Manufacturing (AM) technologies owing to its simplicity and accessible cost. The technique is based on the principle of extrusion of thermoplastic material, layer-by-layer, on a building platform through multiple head nozzles. Metal filled filaments, in combination with debinding and sintering cycles, may innovate and transform the traditional functioning of the MEX technique into a cost-effective alternative for the conventional metallic AM processes. In the present document, the optimal printing conditions characterizing LPBF technology were replicated on MEX technology, with the aim of assessing the effects of the printing parameter hatch angle over the material properties and, at the same time, providing a better understanding of the production of medical metal parts via MEX. Indeed, in this particular context, the use of Powder Bed Fusion (PBF) and Directed Energy Deposition (DED) prevails, requiring MEX-based technique extensive research for its applicability. The influence of a specific AM process parameter, the hatch angle, was assessed following a single factor Design of Experiment (DOE), varying over two levels: the optimal Laser Powder Bed Fusion (LPBF) scanning strategy (67°k) and the most common MEX deposition strategy (±45°). Specimens were manufactured, using MEX technology (Ultimaker S5) and AISI 316L filament (BASF Ultrafuse 316L) and tested. Results of the defect analysis, including closed and open porosity, and mechanical properties were collected and statistically compared to determine any difference in the two-deposition strategies. Furthermore, in the analysis, LPBF key characteristics are reported as benchmark values.

**Introduction**

The integration of metallic materials and additive manufacturing (AM) is revolutionizing the production of metal devices for medical applications. Currently, metal AM technologies, belonging to the class of Powder Bed Fusion (PBF) or Directed Energy Deposition (DED), benefit from widespread adoption in medical research. Despite its practicality and cost-effectiveness, Material Extrusion (MEX) is less adopted owing to the less control over material density and greater porosity, thereby adversely affecting the mechanical characteristics [1-3].

In the present research, Laser Powder Bed Fusion (LPBF) and MEX technologies were compared. Specifically, the optimal printing conditions characterizing LPBF technology were replicated on MEX technology, with the aim of evaluating the influence of the printing parameter hatch angle over the material properties and, at the same time, providing a better understanding of the fabrication of medical metal parts via MEX.

<figure><img src="image_191.png" /></figure>

<footnote n="1">¹University of Bergamo - Department of Management, Information and Production Engineering, via Pasubio 7/b, Dalmine (BG), 24044, Italy</footnote>

<footnote n="1">¹University of Bergamo - Department of Engineering and Applied Sciences, via Pasubio 7/b, Dalmine (BG), 24044, Italy</footnote>

<footnote>ªfrancesca.sala@unibg.it, blorenzo.nani@unibg.it,</footnote>

<footnote>ᶜmariangela.quarto@unibg.it, dgianluca.d-urso@unibg.it</footnote>
# Preliminary assessment of material extrusion (MEX) for medical applications: The effect of hatch angle

SALA Francesca<fnref n="1" />,ª*, NANI Lorenzo²,ᵇ, QUARTO Mariangela<fnref n="1" />,ᶜ and D'URSO Gianluca<fnref n="1" />,ᵈ

**Keywords:** Material Extrusion, MEX, Laser Powder Bed Fusion, L-PBF, Healthcare

**Abstract.** Material extrusion (MEX) is one of the most widely used Additive Manufacturing (AM) technologies owing to its simplicity and accessible cost. The technique is based on the principle of extrusion of thermoplastic material, layer-by-layer, on a building platform through multiple head nozzles. Metal filled filaments, in combination with debinding and sintering cycles, may innovate and transform the traditional functioning of the MEX technique into a cost-effective alternative for the conventional metallic AM processes. In the present document, the optimal printing conditions characterizing LPBF technology were replicated on MEX technology, with the aim of assessing the effects of the printing parameter hatch angle over the material properties and, at the same time, providing a better understanding of the production of medical metal parts via MEX. Indeed, in this particular context, the use of Powder Bed Fusion (PBF) and Directed Energy Deposition (DED) prevails, requiring MEX-based technique extensive research for its applicability. The influence of a specific AM process parameter, the hatch angle, was assessed following a single factor Design of Experiment (DOE), varying over two levels: the optimal Laser Powder Bed Fusion (LPBF) scanning strategy (67°k) and the most common MEX deposition strategy (±45°). Specimens were manufactured, using MEX technology (Ultimaker S5) and AISI 316L filament (BASF Ultrafuse 316L) and tested. Results of the defect analysis, including closed and open porosity, and mechanical properties were collected and statistically compared to determine any difference in the two-deposition strategies. Furthermore, in the analysis, LPBF key characteristics are reported as benchmark values.

**Introduction**

The integration of metallic materials and additive manufacturing (AM) is revolutionizing the production of metal devices for medical applications. Currently, metal AM technologies, belonging to the class of Powder Bed Fusion (PBF) or Directed Energy Deposition (DED), benefit from widespread adoption in medical research. Despite its practicality and cost-effectiveness, Material Extrusion (MEX) is less adopted owing to the less control over material density and greater porosity, thereby adversely affecting the mechanical characteristics [1-3].

In the present research, Laser Powder Bed Fusion (LPBF) and MEX technologies were compared. Specifically, the optimal printing conditions characterizing LPBF technology were replicated on MEX technology, with the aim of evaluating the influence of the printing parameter hatch angle over the material properties and, at the same time, providing a better understanding of the fabrication of medical metal parts via MEX.

<figure><img src="image_189.png" /></figure>

<footnote n="1">¹University of Bergamo - Department of Management, Information and Production Engineering, via Pasubio 7/b, Dalmine (BG), 24044, Italy</footnote>

<footnote n="1">¹University of Bergamo - Department of Engineering and Applied Sciences, via Pasubio 7/b, Dalmine (BG), 24044, Italy</footnote>

<footnote>ªfrancesca.sala@unibg.it, blorenzo.nani@unibg.it,</footnote>

<footnote>ᶜmariangela.quarto@unibg.it, dgianluca.d-urso@unibg.it</footnote>
Optimizing and adjusting LPBF parameters have significant influences on the quality of the printed parts and the required densification, microstructure and mechanical properties [4–6]. Hence, the accomplishment of satisfactory outcomes in LPBF manufacturing relies on the balance of several factors, including laser power, scanning speed, layer thickness, hatch distance and angle. Among the process parameters, the hatch angle represents the angle of rotation with respect to the element growth axis at which scanning vectors belonging to successive layers are melted. For LPBF technologies, the optimal configuration corresponds to 67°. Thus, 67°k scanning strategy, where k indicates the number of layers, involves the rotation of each layer by 67°, creating a structure that minimizes the overlap of scanning vectors belonging to alternating layers.

Generally, the most common deposition strategy adopted by MEX processes is ±45° and it involves the deposition of material lines oriented at +45° on even layers and -45° on odd layers, commonly resulting in the overlapping phenomenon [7]. Specifically, it generates an envelope of porosity, creating continuous pores that follow the intersection of the layers along the growth direction [7]. Therefore, the objective of the current article consisted in the determination of the impact of adopting the optimal scanning strategy, distinctive of the LPBF technologies, over the key properties of elements produced with a different AM printing technique, MEX.

The current research focused on the evaluation of the influence of two different deposition strategies over MEX-built elements, in terms of tensile and hardness properties and distribution of defects (porosity). The results were assessed through statistical analysis.

The evaluations were performed on AISI 316L specimens. Stainless steel alloys are well known for their low cost and good mechanical properties, heat resistance and biocompatibility [8,9]. This material is primarily used in the fabrication of surgical instrumentation. Tough, it is relevant to note that its applications extend to orthopedic implants and fixators, artificial heart valves, syringe needles and numerous other medical tools too [10].

## Methodology

*Sample preparation.* The test samples were produced by means of an Ultimaker S5 (Ultimaker B.V., Utrecht, NL, USA) MEX technology, using a filament with a diameter of 2.85mm provided by BASF (Ultrafuse 316L). The filament was made up of 80% austenitic stainless steel (AISI 316L) and 20% polymer content (polyoxymethylene (POM)), the latter forming the thermoplastic binder matrix that ensures material processing via MEX technology. A print core CC 0.6 with a hardened steel nozzle was used to extrude the filament.

The geometry of the specimen was designed according to international standards applicable to the mechanical tests to be performed, as further detailed in the subsequent sections.

Then, the sample geometries were imported to a slicing software, Ultimaker Cura, where the process parameters (Table 1) were defined following the optimized factors, with respect to density and shrinkage, identified in previous research [11]. Print speed was slightly increased to avoid the delamination effect in the first deposited layers, generated by the tensile specimen geometry.

<table><caption>Table 1. Selected process parameters for the MEX-built specimens.</caption><thead><tr><th></th><th>Layer height</th><th>Infill density</th><th>Infill pattern</th><th>Nozzle temperature</th><th>Build plate temperature</th><th>Print Speed</th></tr></thead><tbody><tr><td>± 45°</td><td>0.1 mm</td><td>100%</td><td>± 45°</td><td>240°</td><td>100°</td><td>40 mm/s</td></tr><tr><td>67° k</td><td>0.1 mm</td><td>100%</td><td>67° k</td><td>240°</td><td>100°</td><td>40 mm/s</td></tr></tbody></table>

The specimens were produced following a Design of Experiment (DOE) with a single factor, planned to study the effect of two different deposition strategies. Specifically, the hatch angle, or the angle defined between the infill lines of one layer and the subsequent, was varied over two levels: ± 45° and 67°k, where k represents the layer number. Figure 1 illustrates in a layer-by-layer
Optimizing and adjusting LPBF parameters have significant influences on the quality of the printed parts and the required densification, microstructure and mechanical properties [4–6]. Hence, the accomplishment of satisfactory outcomes in LPBF manufacturing relies on the balance of several factors, including laser power, scanning speed, layer thickness, hatch distance and angle. Among the process parameters, the hatch angle represents the angle of rotation with respect to the element growth axis at which scanning vectors belonging to successive layers are melted. For LPBF technologies, the optimal configuration corresponds to 67°. Thus, 67°k scanning strategy, where k indicates the number of layers, involves the rotation of each layer by 67°, creating a structure that minimizes the overlap of scanning vectors belonging to alternating layers.

Generally, the most common deposition strategy adopted by MEX processes is ±45° and it involves the deposition of material lines oriented at +45° on even layers and -45° on odd layers, commonly resulting in the overlapping phenomenon [7]. Specifically, it generates an envelope of porosity, creating continuous pores that follow the intersection of the layers along the growth direction [7]. Therefore, the objective of the current article consisted in the determination of the impact of adopting the optimal scanning strategy, distinctive of the LPBF technologies, over the key properties of elements produced with a different AM printing technique, MEX.

The current research focused on the evaluation of the influence of two different deposition strategies over MEX-built elements, in terms of tensile and hardness properties and distribution of defects (porosity). The results were assessed through statistical analysis.

The evaluations were performed on AISI 316L specimens. Stainless steel alloys are well known for their low cost and good mechanical properties, heat resistance and biocompatibility [8,9]. This material is primarily used in the fabrication of surgical instrumentation. Tough, it is relevant to note that its applications extend to orthopedic implants and fixators, artificial heart valves, syringe needles and numerous other medical tools too [10].

## Methodology

*Sample preparation.* The test samples were produced by means of an Ultimaker S5 (Ultimaker B.V., Utrecht, NL, USA) MEX technology, using a filament with a diameter of 2.85mm provided by BASF (Ultrafuse 316L). The filament was made up of 80% austenitic stainless steel (AISI 316L) and 20% polymer content (polyoxymethylene (POM)), the latter forming the thermoplastic binder matrix that ensures material processing via MEX technology. A print core CC 0.6 with a hardened steel nozzle was used to extrude the filament.

The geometry of the specimen was designed according to international standards applicable to the mechanical tests to be performed, as further detailed in the subsequent sections.

Then, the sample geometries were imported to a slicing software, Ultimaker Cura, where the process parameters (Table 1) were defined following the optimized factors, with respect to density and shrinkage, identified in previous research [11]. Print speed was slightly increased to avoid the delamination effect in the first deposited layers, generated by the tensile specimen geometry.

<table><caption>Table 1. Selected process parameters for the MEX-built specimens.</caption><thead><tr><th></th><th>Layer height</th><th>Infill density</th><th>Infill pattern</th><th>Nozzle temperature</th><th>Build plate temperature</th><th>Print Speed</th></tr></thead><tbody><tr><td>± 45°</td><td>0.1 mm</td><td>100%</td><td>± 45°</td><td>240°</td><td>100°</td><td>40 mm/s</td></tr><tr><td>67° k</td><td>0.1 mm</td><td>100%</td><td>67° k</td><td>240°</td><td>100°</td><td>40 mm/s</td></tr></tbody></table>

The specimens were produced following a Design of Experiment (DOE) with a single factor, planned to study the effect of two different deposition strategies. Specifically, the hatch angle, or the angle defined between the infill lines of one layer and the subsequent, was varied over two levels: ± 45° and 67°k, where k represents the layer number. Figure 1 illustrates in a layer-by-layer
comparison, the schematic representation of the two adopted deposition strategies for the core portion of the MEX-built specimens.

As anticipated, in LPBF processes, the 67°k scanning strategy is known for completely minimizing the overlap of the scanning vector of successive layers. With the adoption of the LPBF optimal strategy in MEX processes, the variation in the deposition strategy was performed by changing the infill pattern value during the set-up of the process parameters.

<figure><img src="image_192.png" /></figure>

<figure><img src="image_193.png" /></figure>

<figure><img src="image_194.png" /><figcaption>Figure 1. Layer-by-layer schematic representation of the different deposition strategies. (a) Variation of the hatch angle in four subsequent layers (+45° odd layers, -45° even layers) in ± 45° cylindrical specimen. (b) Constant variation of the hatch angle (steady rotation of 67°) in the 67°k cylindrical specimen.</figcaption></figure>

Once 3D printed, the green parts underwent debinding and sintering processes to obtain, in the first instance, the brown part and, ultimately, the fully metal component. During the debinding phase, the majority of the polymeric component is removed from the green part through a catalytic and thermal process that takes place at 120°C with HNO₃ (98% concentration). During the sintering cycle, which is performed in an argon atmosphere, the brown part is subjected to three thermal ramps (Table 2). The treatments were performed by an external company.

Two sets of samples were produced and compared to benchmark values derived from literature studies and available technical datasheet on AISI 316L specimens manufactured through LPBF powders and technologies [12–16].
comparison, the schematic representation of the two adopted deposition strategies for the core portion of the MEX-built specimens.

As anticipated, in LPBF processes, the 67°k scanning strategy is known for completely minimizing the overlap of the scanning vector of successive layers. With the adoption of the LPBF optimal strategy in MEX processes, the variation in the deposition strategy was performed by changing the infill pattern value during the set-up of the process parameters.

<figure><img src="image_190.png" /><figcaption>(a) ±45°</figcaption></figure>

<figure><img src="image_191.png" /></figure>

<figure><img src="image_192.png" /></figure>

<figure><img src="image_193.png" /></figure>

<figure><img src="image_194.png" /><figcaption>(b) 67°k</figcaption></figure>

<figure><img src="image_195.png" /></figure>

<figure><img src="image_196.png" /></figure>

<figure><img src="image_197.png" /></figure>

<figure><img src="image_198.png" /><figcaption>Figure 1. Layer-by-layer schematic representation of the different deposition strategies. (a) Variation of the hatch angle in four subsequent layers (+45° odd layers, -45° even layers) in ± 45° cylindrical specimen. (b) Constant variation of the hatch angle (steady rotation of 67°) in the 67°k cylindrical specimen.</figcaption></figure>

Once 3D printed, the green parts underwent debinding and sintering processes to obtain, in the first instance, the brown part and, ultimately, the fully metal component. During the debinding phase, the majority of the polymeric component is removed from the green part through a catalytic and thermal process that takes place at 120°C with HNO₃ (98% concentration). During the sintering cycle, which is performed in an argon atmosphere, the brown part is subjected to three thermal ramps (Table 2). The treatments were performed by an external company.

Two sets of samples were produced and compared to benchmark values derived from literature studies and available technical datasheet on AISI 316L specimens manufactured through LPBF powders and technologies [12–16].
<table><caption>Table 2. Thermal ramps of the sintering cycle.</caption><thead><tr><th></th><th>Initial temperature</th><th>Final temperature</th><th>Ramp rate</th><th>Holding time</th></tr></thead><tbody><tr><td>1<sup>st</sup> Ramp</td><td>Room temperature</td><td>600°C</td><td>5°C/min</td><td>1 h</td></tr><tr><td>2<sup>nd</sup> Ramp</td><td>600°C</td><td>1380°C</td><td>5°C/min</td><td>3 h</td></tr><tr><td>3<sup>rd</sup> Ramp</td><td>1380°C</td><td>Room temperature</td><td>Furnace cooling</td><td></td></tr></tbody></table>

Porosity analysis. Compared to LPBF deposition processes, MEX generates structures characterized by reduced bulk density due to the presence of defects (i.e., open and closed porosity). In the present study, defect analysis was performed on cylindrical specimens with dimension 15 mm diameter and 5 mm height, in three separate stages.

In the first instance, porosities were studied by considering a cross-section orthogonal to the deposition lines (Figure 2). The surface section was polished up to 1 µm with diamond suspension. The sample sections were observed with a Keyence VHX-7100 digital-optical microscope and the amount of porosity was quantified through ImageJ software.

The metallurgical sections can only provide an indication of the porosity distribution, which is specific to identified cross-section. Given the fact that MEX processes, unlike the LPBF, tend to generate envelopes owing to the continuity of the defects between different subsequent layers rather than spherical pores, additional investigation are required. For this reason, to adequately quantify the amount of closed porosity on the whole specimen, a gas pycnometer was used. Specifically, the pycnometer estimated the average value of the volume occupied by the printed part, namely the volume of the material, and the close porosity; then, knowing the density of the material, it is possible to estimate the closed porosity as reported in Eq.1 and Eq.2.

$$ \delta_{\text{bulk}} = \frac{\text{weight}}{V_{\text{pycnometer}}} \qquad (1) $$

$$ cp = \frac{\delta_{\text{material}} - \delta_{\text{bulk}}}{\delta_{\text{material}}} \qquad (2) $$

Where $\delta_{\text{bulk}}$ and $\delta_{\text{material}}$ are the mass density of the printed part and the material used for the printing process (8 g/cm³) respectively, weight is the mass of the printed part, $V_{\text{pycnometer}}$ is the volume of the part measured by the gas pycnometer and cp identifies the amount of closed porosity.

In a final step, an estimation of the open porosity was conducted considering the overall volume derived by the dimensions of the printed samples. In particular, the amount of open porosity (op) was computed as the difference between the overall volume and the volume defined by the gas pycnometer.

<figure><img src="image_195.png" /><figcaption>Figure 2. Schematic representation of the considered plane (cross-section parallel to the growth direction of the specimen) for the analysis of porosity.</figcaption></figure>
<table><caption>Table 2. Thermal ramps of the sintering cycle.</caption><thead><tr><th></th><th>Initial temperature</th><th>Final temperature</th><th>Ramp rate</th><th>Holding time</th></tr></thead><tbody><tr><td>1<sup>st</sup> Ramp</td><td>Room temperature</td><td>600°C</td><td>5°C/min</td><td>1 h</td></tr><tr><td>2<sup>nd</sup> Ramp</td><td>600°C</td><td>1380°C</td><td>5°C/min</td><td>3 h</td></tr><tr><td>3<sup>rd</sup> Ramp</td><td>1380°C</td><td>Room temperature</td><td>Furnace cooling</td><td></td></tr></tbody></table>

*Porosity analysis.* Compared to LPBF deposition processes, MEX generates structures characterized by reduced bulk density due to the presence of defects (i.e., open and closed porosity). In the present study, defect analysis was performed on cylindrical specimens with dimension 15 mm diameter and 5 mm height, in three separate stages.

In the first instance, porosities were studied by considering a cross-section orthogonal to the deposition lines (Figure 2). The surface section was polished up to 1 µm with diamond suspension. The sample sections were observed with a Keyence VHX-7100 digital-optical microscope and the amount of porosity was quantified through ImageJ software.

The metallurgical sections can only provide an indication of the porosity distribution, which is specific to identified cross-section. Given the fact that MEX processes, unlike the LPBF, tend to generate envelopes owing to the continuity of the defects between different subsequent layers rather than spherical pores, additional investigation are required. For this reason, to adequately quantify the amount of closed porosity on the whole specimen, a gas pycnometer was used. Specifically, the pycnometer estimated the average value of the volume occupied by the printed part, namely the volume of the material, and the close porosity; then, knowing the density of the material, it is possible to estimate the closed porosity as reported in Eq.1 and Eq.2.

$$ \delta_{\text{bulk}} = \frac{\text{weight}}{V_{\text{pycnometer}}} \qquad (1) $$

$$ cp = \frac{\delta_{\text{material}} - \delta_{\text{bulk}}}{\delta_{\text{material}}} \qquad (2) $$

Where $\delta_{\text{bulk}}$ and $\delta_{\text{material}}$ are the mass density of the printed part and the material used for the printing process (8 g/cm³) respectively, *weight* is the mass of the printed part, $V_{\text{pycnometer}}$ is the volume of the part measured by the gas pycnometer and *cp* identifies the amount of closed porosity.

In a final step, an estimation of the open porosity was conducted considering the overall volume derived by the dimensions of the printed samples. In particular, the amount of open porosity (*op*) was computed as the difference between the overall volume and the volume defined by the gas pycnometer.

<figure><img src="image_199.png" /><figcaption>Figure 2. Schematic representation of the considered plane (cross-section parallel to the growth direction of the specimen) for the analysis of porosity.</figcaption></figure>
*Tensile test.* According to ASTM F3122-14 [17], the procedure outlined in the test method UNI EN ISO 6892-1:2020 [18] defines the methodology for conducting a tensile test on metal materials and determining its mechanical properties, including elongation at break (A%) and ultimate tensile strengths (UTS). The procedure was performed on specimens characterized by a circular section, with a calibrated length of 30 mm and a diameter of 6 mm. Tensile tests were performed using a Galdabini testing machine with a 50kN load cell. The tests were conducted perpendicular to the growth direction of the layers, under deformation control $2.5 \times 10^{-4} \text{ s}^{-1}$ and a preload of 400N. The procedure was conducted on tensile specimens and repeated four times per deposition pattern variation.

*Hardness test.* ASTM F3122-14 specifies UNI EN ISO 6507-1:2018 [19] as a standard for the Vickers hardness (HV) of additive-manufactured parts. The top surface of cylindrical specimen was polished up to $1 \mu\text{m}$ with diamond suspension. A $5 \times 5$ grid was drawn on the internal section, obtaining 25 uniformly spaced indentations. The procedure involved applying a 1kgf load for 15s. The protocol was repeated one time per each deposition pattern variation.

## Results and discussion

*Porosity analysis.* The porosity analysis evaluated by means of the metallurgical sections is reported in Figure 3. The image analysis allowed to discriminate the porosity area, estimating an amount of porosity equal to 1.99% for ±45° infill pattern, and 1.47% for 67°k. For both deposition strategies, the observed porosity percentage values were reduced compared to the literature results. This result may be related to the specificity of the measurement, meaning that the value is specific to a single plane section and not directly attributable to the specimen in its entirety.

<figure><img src="image_196.png" /><figcaption>Figure 3. Metallurgical section for the two deposition strategies. (a) ±45° Infill pattern and (b) 67°k infill pattern.</figcaption></figure>

The image investigation was complemented by the pycnometer analysis, which provided the aggregate value of closed porosity. The information related to the extension of the porosity along the depth dimension was captured. The pycnometer confirmed the adherence of the obtained porosity values to the conservative literature results. Furthermore, comparing the volume estimated by the gas pycnometer and the nominal part volume, the percentage of open porosity were estimated too.

The results showed that samples characterized by ± 45° infill pattern had closed porosity values of 1.29% and open porosity of 12.95%; whilst 67°k infill pattern had closed porosity values of 2.79% and open porosity of 3.70%. Although, moving from the ± 45° deposition strategy to the 67°k, the values suggested an increase in the amount of closed porosity, but a reduction considering the open porosity, prompting the ability of the 67°k to interrupt the continuity of the porosity envelope. Still, the porosity values recorded in MEX-built specimens are not comparable to the LPBF ones, capable of providing densities greater than 99%.
*Tensile test.* According to ASTM F3122-14 [17], the procedure outlined in the test method UNI EN ISO 6892-1:2020 [18] defines the methodology for conducting a tensile test on metal materials and determining its mechanical properties, including elongation at break (A%) and ultimate tensile strengths (UTS). The procedure was performed on specimens characterized by a circular section, with a calibrated length of 30 mm and a diameter of 6 mm. Tensile tests were performed using a Galdabini testing machine with a 50kN load cell. The tests were conducted perpendicular to the growth direction of the layers, under deformation control $2.5 \times 10^{-4} \text{ s}^{-1}$ and a preload of 400N. The procedure was conducted on tensile specimens and repeated four times per deposition pattern variation.

*Hardness test.* ASTM F3122-14 specifies UNI EN ISO 6507-1:2018 [19] as a standard for the Vickers hardness (HV) of additive-manufactured parts. The top surface of cylindrical specimen was polished up to $1 \mu\text{m}$ with diamond suspension. A $5 \times 5$ grid was drawn on the internal section, obtaining 25 uniformly spaced indentations. The procedure involved applying a 1kgf load for 15s. The protocol was repeated one time per each deposition pattern variation.

## Results and discussion

*Porosity analysis.* The porosity analysis evaluated by means of the metallurgical sections is reported in Figure 3. The image analysis allowed to discriminate the porosity area, estimating an amount of porosity equal to 1.99% for ±45° infill pattern, and 1.47% for 67°k. For both deposition strategies, the observed porosity percentage values were reduced compared to the literature results. This result may be related to the specificity of the measurement, meaning that the value is specific to a single plane section and not directly attributable to the specimen in its entirety.

<figure><img src="image_200.png" /><figcaption>Figure 3. Metallurgical section for the two deposition strategies. (a) ±45° Infill pattern and (b) 67°k infill pattern.</figcaption></figure>

The image investigation was complemented by the pycnometer analysis, which provided the aggregate value of closed porosity. The information related to the extension of the porosity along the depth dimension was captured. The pycnometer confirmed the adherence of the obtained porosity values to the conservative literature results. Furthermore, comparing the volume estimated by the gas pycnometer and the nominal part volume, the percentage of open porosity were estimated too.

The results showed that samples characterized by ± 45° infill pattern had closed porosity values of 1.29% and open porosity of 12.95%; whilst 67°k infill pattern had closed porosity values of 2.79% and open porosity of 3.70%. Although, moving from the ± 45° deposition strategy to the 67°k, the values suggested an increase in the amount of closed porosity, but a reduction considering the open porosity, prompting the ability of the 67°k to interrupt the continuity of the porosity envelope. Still, the porosity values recorded in MEX-built specimens are not comparable to the LPBF ones, capable of providing densities greater than 99%.
*Tensile test.* To assess the impact of the infill pattern on mechanical strength, a statistical analysis was conducted on the outcomes of tensile tests (Table 4).

The Analysis of Variance (ANOVA) showed an effect of the infill pattern on the elongation (A%) and, on the contrary, it seemed to not affect the Ultimate Tensile Strength (UTS). Notably, the elongation demonstrated a *p*-value of 0.042, which is below the predetermined confidence level of 5%. Conversely, the UTS exhibited a *p*-value of 0.951, emphasizing the absence of any substantial effect.

As reported in Figure 4, the elongation resulted to be greater for 67°k samples; this suggested that the samples 3D printed applying a 67°k pattern may undergo more deformation before reaching the point of rupture. In other words, this suggested a higher ductile behavior and greater ability to withstand deformation before necking under stress. The result may be a positive indicator for certain applications, such as when the deformability of the material is important. For example, the result may be beneficial in the medical context, where the increased ductile behavior represents improved mechanical reliability.

<table><caption>Table 3. Mechanical properties of tensile specimens.</caption><thead><tr><th>Infill pattern</th><th>Repetition</th><th>A%</th><th>UTS</th></tr></thead><tbody><tr><td rowspan="5">± 45°</td><td>1</td><td>50.5%</td><td>479 MPa</td></tr><tr><td>2</td><td>47.2%</td><td>433 MPa</td></tr><tr><td>3</td><td>48.4%</td><td>393 MPa</td></tr><tr><td>4</td><td>40.1%</td><td>453 MPa</td></tr><tr><td>1</td><td>54.3%</td><td>471 MPa</td></tr><tr><td rowspan="3">67°k</td><td>2</td><td>61.9%</td><td>439 MPa</td></tr><tr><td>3</td><td>57.3%</td><td>454 MPa</td></tr><tr><td>4</td><td>49.0%</td><td>400 MPa</td></tr></tbody></table>

<figure><img src="image_197.png" /><figcaption>Figure 4. Main effect plot for (a) elongation at break (A%) and (b) Ultimate Tensile Strength (UTS).</figcaption></figure>

<figure><img src="image_198.png" /></figure>

Furthermore, the outcome could be correlated with the arrangement of the pores, particularly with the breaking of the envelopes caused by the continuous rotation of the pattern 67°k. This finding is supported by the images of the fracture surfaces reported in Figure 5, where it is possible to observe the distinct arrangement of the pores based on the infill pattern. It is evident that in the ± 45° infill pattern, the pores are placed along columns suggesting preferential direction for the fracture propagation, whereas for the 67°k pattern, the pores are uniformly distributed across the entire surface.
*Tensile test.* To assess the impact of the infill pattern on mechanical strength, a statistical analysis was conducted on the outcomes of tensile tests (Table 4).

The Analysis of Variance (ANOVA) showed an effect of the infill pattern on the elongation (A%) and, on the contrary, it seemed to not affect the Ultimate Tensile Strength (UTS). Notably, the elongation demonstrated a *p*-value of 0.042, which is below the predetermined confidence level of 5%. Conversely, the UTS exhibited a *p*-value of 0.951, emphasizing the absence of any substantial effect.

As reported in Figure 4, the elongation resulted to be greater for 67°k samples; this suggested that the samples 3D printed applying a 67°k pattern may undergo more deformation before reaching the point of rupture. In other words, this suggested a higher ductile behavior and greater ability to withstand deformation before necking under stress. The result may be a positive indicator for certain applications, such as when the deformability of the material is important. For example, the result may be beneficial in the medical context, where the increased ductile behavior represents improved mechanical reliability.

<table><caption>Table 3. Mechanical properties of tensile specimens.</caption><thead><tr><th>Infill pattern</th><th>Repetition</th><th>A%</th><th>UTS</th></tr></thead><tbody><tr><td rowspan="5">± 45°</td><td>1</td><td>50.5%</td><td>479 MPa</td></tr><tr><td>2</td><td>47.2%</td><td>433 MPa</td></tr><tr><td>3</td><td>48.4%</td><td>393 MPa</td></tr><tr><td>4</td><td>40.1%</td><td>453 MPa</td></tr><tr><td>1</td><td>54.3%</td><td>471 MPa</td></tr><tr><td rowspan="3">67°k</td><td>2</td><td>61.9%</td><td>439 MPa</td></tr><tr><td>3</td><td>57.3%</td><td>454 MPa</td></tr><tr><td>4</td><td>49.0%</td><td>400 MPa</td></tr></tbody></table>

<figure><img src="image_201.png" /><figcaption>Figure 4. Main effect plot for (a) elongation at break (A%) and (b) Ultimate Tensile Strength (UTS).</figcaption></figure>

<figure><img src="image_202.png" /></figure>

Furthermore, the outcome could be correlated with the arrangement of the pores, particularly with the breaking of the envelopes caused by the continuous rotation of the pattern 67°k. This finding is supported by the images of the fracture surfaces reported in Figure 5, where it is possible to observe the distinct arrangement of the pores based on the infill pattern. It is evident that in the ± 45° infill pattern, the pores are placed along columns suggesting preferential direction for the fracture propagation, whereas for the 67°k pattern, the pores are uniformly distributed across the entire surface.