# Experimental evaluation of the maximal force before debonding a part from the build platform of an AM printer

SPITAELS Laurent<fnref n="1" />,ª*, ALDEITURRIAGA OLABARRI Naiara<fnref n="1" />,<fnref n="2" />, BOSSU Julien<fnref n="1" />, MARTIC Gregory<fnref n="3" />, JUSTE Enrique<fnref n="3" />, RIVIERE-LORPHEVRE Edouard<fnref n="1" />, ARRAZOLA Pedro-José<fnref n="2" />, DUCOBU François<fnref n="1" />

**Keywords:** Additive Manufacturing, Peel Test, Hybrid Manufacturing, Material Extrusion

**Abstract.** Hybrid manufacturing relies on the combination of different processes to overcome their own limitations. Combining additive manufacturing such as material extrusion (MEX) and subtractive conventional processes such as milling enables the production of complex design parts with smooth surface and tight tolerances to be foreseen. However, finishing a part directly in the printer by milling operations brings new questions as the maximal force the part can withstand before debonding from the build platform. This paper proposes an experimental method to determine this force by studying the influence of several parameters: the adhesion strategy, the infill pattern and density, the zone of the part where the force is applied as well as the influence of the part design. The method was successfully tested with Polylactic Acid (PLA) parts on a standard MEX printer: the Ultimaker 2+. In the best configuration: no brim, cubic infill of 95% and a force applied on an entire part face, the parts resisted to a force of 50.59 N ± 1.97 N.

## Introduction

Additive Manufacturing processes are seen as game changers for the near future [1]. Indeed, from the first developments in the 1980s to 2023, the annual average growth rate of the AM market reached 26.7% [1]. This new way of producing goods begins to be adopted not only for prototyping or modelling, but also now for producing end-user parts [1]. Among the seven families defined by ISO 52900, Material Extrusion (MEX) and Binder Jetting (BJ) are gaining great interest from the industries [2]. They are seen as the two most promising processes in a ten-year horizon for the manufacturing of metallic or ceramics components [2,3]. Indeed, compared to the conventional manufacturing processes (as injection or machining, e.g.), the AM processes enable the production of complex shaped parts which were not possible to achieve before [4]. Moreover, the layer-by-layer nature of AM allows to bring complexity in the material and functions of the parts [4].

However, despite the large progresses achieved in the previous years, the AM processes still lack of capability for producing smooth surfaces with tight tolerances [3–5]. Arithmetic roughness (Ra) values from 9 µm to 40 µm are typical in MEX process [5], while widespread MEX printers produce parts with very large tolerances compared to conventional subtractive processes [6]:

- IT between 12 and 16 of ISO 286-1 for dimensional features.

- L class of ISO 2768-2 for geometrical features.

Post finishing operations (sanding or machining, e.g.) are then required to match the customer specifications [4]. Performing the finishing operations when the part is already totally produced brings new constraints and makes impossible the finishing of complex designs such as lattice, for example. Indeed, the subtractive processes require to access the surfaces to be treated and lack of

<figure><img src="image_95.png" /></figure>

<footnote n="1">¹ Machine Design and Production Engineering Lab - Research Institute for Science and Material Engineering - University of Mons, Place du Parc 20, Mons, Belgium</footnote>

<footnote n="2">² Mondragon Unibertsitatea, Loramendi 4, Arrasate-Mondragón, 20500, Spain</footnote>

<footnote n="3">³ Belgium Ceramic Research Centre, Avenue Gouverneur Cornez 4, Mons 7000, Belgium</footnote>

<footnote>ª laurent.spitaels@umons.ac.be</footnote>
flexibility [7]. Moreover, changing of machine to produce and finish the part increases the risk to generate scrap due to wrong part alignments and to the difficulty to take references on very rough surfaces. Therefore, the succession of additive and subtractive operations within the same hybrid machine is seen as a very promising solution to overcome both processes' limitations, while allowing the part production and finishing to be conducted directly on the build platform [1,7,8]. These machines will be able to produce highly customized components at lower cost than each process considered separately [7]. Commercial solutions already exist for the Powder Bed Fusion (PBF) and Directed Energy Deposition (DED) metal AM, but this kind of machines are expensive and requires specific and costly feedstock [7–9]. A more affordable alternative is to use an indirect method such as the MEX of polymers filled with particles of metal or ceramics [9]. In this case, the produced green part does not exhibit the final material properties [10]. It is, therefore, easier to machine and, then, less expensive. Nevertheless, the process is indirect and requires additional operations such as debinding and sintering to obtain a dense part [9]. When the chosen MEX process uses a screw and pellets instead of relying on filament extrusion, the cost savings can be higher. Indeed, Metal Injection Molding (MIM) and Ceramic Injection Molding (CIM) feedstocks exist since decades and can be used in these machines [11]. Furthermore, the knowledge for obtaining the final dense parts by debinding and sintering is already available, while the resulting hybrid machine can be easily integrated into an existing CIM or MIM routes [12].

When foreseeing the use of a hybrid machine, ensuring the adhesion of the part to the build platform is mandatory. Indeed, if the part is finished by milling, it is not clamped but bonded to the build platform. Standards already exist for bonded assemblies (ISO 8510 and ISO 11339), but none can be directly applied to AM parts. Besides, there is still a lack of data for hybrid machines, especially for the maximal forces which can be applied to the part without risking its debonding from the build platform. Many previous papers studied the influence of the printing parameters (infill density and pattern, layer thickness, e.g.) on the dimensional accuracy, surface topography and mechanical properties [13]. Only few were dedicated to the analysis of adhesion of the generated roads to the printer build platform [14]. Moreover, the existing studies mainly focus on the influence of the build platform temperature and not on the influence of other printing parameters [14]. This paper aims to close this gap by proposing an experimental method to evaluate the maximal force a part can withstand before debonding from its build platform. Several potential factors of influence are tested, such as the infill pattern and density, the adhesion strategy, the zone on which the force is applied and the part design.

## Method

Manufacturing of Parts: The material extrusion printer chosen for the experiments was the Ultimaker 2+. This affordable machine, widespread in many FabLabs and other collaborative workshops, allows to manufacture parts in a set of conventional thermoplastics as PLA (Polylactic Acid), ABS (Acrylonitrile Butadiene Styrene) or even TPU (Thermoplastic Polyurethane). For the tests, the Ultimaker PLA was chosen (material characteristics and properties are available in the datasheet [15]). The printer exhibits a cartesian architecture with a printing volume of 223 mm x 223 mm x 205 mm for its X, Y and Z axes, respectively. The machine was fitted with a 0.4 mm nozzle and uses a standard 2.85 mm diameter filament to produce the parts on its glass build platform.

Three main part designs were used for the experiments as depicted in Fig. 1:

- Geometry A: 20 mm x 20 mm x 20 mm cube with 1 mm radius fillets for its vertical edges;

- Geometry B: a 20 mm x 20 mm x 20 mm cube with 1 mm radius fillets for its vertical edges completed on its top by a cylinder of 15 mm of height and 15 mm of diameter. A 3 mm radius fillet was also added between the cylinder and the cube.
flexibility [7]. Moreover, changing of machine to produce and finish the part increases the risk to generate scrap due to wrong part alignments and to the difficulty to take references on very rough surfaces. Therefore, the succession of additive and subtractive operations within the same hybrid machine is seen as a very promising solution to overcome both processes' limitations, while allowing the part production and finishing to be conducted directly on the build platform [1,7,8]. These machines will be able to produce highly customized components at lower cost than each process considered separately [7]. Commercial solutions already exist for the Powder Bed Fusion (PBF) and Directed Energy Deposition (DED) metal AM, but this kind of machines are expensive and requires specific and costly feedstock [7-9]. A more affordable alternative is to use an indirect method such as the MEX of polymers filled with particles of metal or ceramics [9]. In this case, the produced green part does not exhibit the final material properties [10]. It is, therefore, easier to machine and, then, less expensive. Nevertheless, the process is indirect and requires additional operations such as debinding and sintering to obtain a dense part [9]. When the chosen MEX process uses a screw and pellets instead of relying on filament extrusion, the cost savings can be higher. Indeed, Metal Injection Molding (MIM) and Ceramic Injection Molding (CIM) feedstocks exist since decades and can be used in these machines [11]. Furthermore, the knowledge for obtaining the final dense parts by debinding and sintering is already available, while the resulting hybrid machine can be easily integrated into an existing CIM or MIM routes [12].

When foreseeing the use of a hybrid machine, ensuring the adhesion of the part to the build platform is mandatory. Indeed, if the part is finished by milling, it is not clamped but bonded to the build platform. Standards already exist for bonded assemblies (ISO 8510 and ISO 11339), but none can be directly applied to AM parts. Besides, there is still a lack of data for hybrid machines, especially for the maximal forces which can be applied to the part without risking its debonding from the build platform. Many previous papers studied the influence of the printing parameters (infill density and pattern, layer thickness, e.g.) on the dimensional accuracy, surface topography and mechanical properties [13]. Only few were dedicated to the analysis of adhesion of the generated roads to the printer build platform [14]. Moreover, the existing studies mainly focus on the influence of the build platform temperature and not on the influence of other printing parameters [14]. This paper aims to close this gap by proposing an experimental method to evaluate the maximal force a part can withstand before debonding from its build platform. Several potential factors of influence are tested, such as the infill pattern and density, the adhesion strategy, the zone on which the force is applied and the part design.

## Method

Manufacturing of Parts: The material extrusion printer chosen for the experiments was the Ultimaker 2+. This affordable machine, widespread in many FabLabs and other collaborative workshops, allows to manufacture parts in a set of conventional thermoplastics as PLA (Polylactic Acid), ABS (Acrylonitrile Butadiene Styrene) or even TPU (Thermoplastic Polyurethane). For the tests, the Ultimaker PLA was chosen (material characteristics and properties are available in the datasheet [15]). The printer exhibits a cartesian architecture with a printing volume of 223 mm x 223 mm x 205 mm for its X, Y and Z axes, respectively. The machine was fitted with a 0.4 mm nozzle and uses a standard 2.85 mm diameter filament to produce the parts on its glass build platform.

Three main part designs were used for the experiments as depicted in Fig. 1:

- Geometry A: 20 mm x 20 mm x 20 mm cube with 1 mm radius fillets for its vertical edges;

- Geometry B: a 20 mm x 20 mm x 20 mm cube with 1 mm radius fillets for its vertical edges completed on its top by a cylinder of 15 mm of height and 15 mm of diameter. A 3 mm radius fillet was also added between the cylinder and the cube.
- Geometry C: a 20 mm x 20 mm x 20 mm cube with 3 mm radius fillets for its vertical edges completed on its top by a cylinder of 15 mm of height and 15 mm of diameter. A 3 mm radius fillet was also added between the cylinder and the cube.

<figure><img src="image_95.png" /><figcaption>Fig. 1 Three considered part geometries.</figcaption></figure>

These geometries allow to see the influence of the design on the required force a part can withstand before debonding from the build platform. When printed, the cube's faces were oriented with respect to the X and Y axes of the printer. Cura ver. 5.2.1 was used to slice the parts. The printing parameters used to produce the parts are given in Table 1. Two different infill patterns were chosen: cubic and concentric, while three infill densities were tested: 20%, 95% and 100%. The printing direction (as depicted in Fig. 1) was chosen to avoid the need of support structure. No adhesive lacquer (or glue) was used for all the tests. The first layer thickness was set at 0.1 mm (instead of the 0.27 mm standard value in Cura) to ensure a uniform layer thickness throughout the part build direction. The Z offset of the printer was set to 0.1 mm using a gauge.

<table><caption>Table 1 Printing parameters used to produce the parts in PLA with a 2.85 mm filament.</caption><tr><td>Nozzle diameter</td><td>[mm]</td><td>0.4</td></tr><tr><td>Layer thickness</td><td>[mm]</td><td>0.1</td></tr><tr><td>First layer thickness</td><td>[mm]</td><td>0.1</td></tr><tr><td>Nozzle temperature</td><td>[°C]</td><td>220</td></tr><tr><td>Build platform temperature</td><td>[°C]</td><td>60</td></tr><tr><td>Infill pattern</td><td></td><td>Cubic or concentric</td></tr><tr><td>Infill density</td><td>[%]</td><td>20, 95 and 100</td></tr><tr><td>Adhesive lacquer or glue</td><td></td><td>No</td></tr></table>

Experimental plan and setup: The main goal of the experiments was to give an objective measurement of the force a part can withstand before debonding. The force direction of application was set parallel to the build platform since the finishing operations by milling generate mainly forces in this direction.

Three dedicated force application devices (from 1 to 3 as depicted in Fig. 2) were designed to apply a desired force on the part. Two different designs were produced to apply the force on an entire face (Fig. 2.1) of the part or in dedicated zones localized at the build platform level (Fig. 2.2) and part top (Fig. 2.3). The same design (rotated by 180° across its X axis) was used for applying the force at the build platform level and part top. When used, each device was installed on the build platform of the printer around the part to be tested. A nylon wire was attached to the device and linked to a mass using a pulley as depicted in Fig. 3. The pulley can freely turn around its axis, while a dedicated pulley holder linked it to the table on which the printer was installed.
- Geometry C: a 20 mm x 20 mm x 20 mm cube with 3 mm radius fillets for its vertical edges completed on its top by a cylinder of 15 mm of height and 15 mm of diameter. A 3 mm radius fillet was also added between the cylinder and the cube.

<figure><img src="image_96.png" /><figcaption>Fig. 1 Three considered part geometries.</figcaption></figure>

These geometries allow to see the influence of the design on the required force a part can withstand before debonding from the build platform. When printed, the cube's faces were oriented with respect to the X and Y axes of the printer. Cura ver. 5.2.1 was used to slice the parts. The printing parameters used to produce the parts are given in Table 1. Two different infill patterns were chosen: cubic and concentric, while three infill densities were tested: 20%, 95% and 100%. The printing direction (as depicted in Fig. 1) was chosen to avoid the need of support structure. No adhesive lacquer (or glue) was used for all the tests. The first layer thickness was set at 0.1 mm (instead of the 0.27 mm standard value in Cura) to ensure a uniform layer thickness throughout the part build direction. The Z offset of the printer was set to 0.1 mm using a gauge.

<table><caption>Table 1 Printing parameters used to produce the parts in PLA with a 2.85 mm filament.</caption><tr><td>Nozzle diameter</td><td>[mm]</td><td>0.4</td></tr><tr><td>Layer thickness</td><td>[mm]</td><td>0.1</td></tr><tr><td>First layer thickness</td><td>[mm]</td><td>0.1</td></tr><tr><td>Nozzle temperature</td><td>[°C]</td><td>220</td></tr><tr><td>Build platform temperature</td><td>[°C]</td><td>60</td></tr><tr><td>Infill pattern</td><td></td><td>Cubic or concentric</td></tr><tr><td>Infill density</td><td>[%]</td><td>20, 95 and 100</td></tr><tr><td>Adhesive lacquer or glue</td><td></td><td>No</td></tr></table>

Experimental plan and setup: The main goal of the experiments was to give an objective measurement of the force a part can withstand before debonding. The force direction of application was set parallel to the build platform since the finishing operations by milling generate mainly forces in this direction.

Three dedicated force application devices (from 1 to 3 as depicted in Fig. 2) were designed to apply a desired force on the part. Two different designs were produced to apply the force on an entire face (Fig. 2.1) of the part or in dedicated zones localized at the build platform level (Fig. 2.2) and part top (Fig. 2.3). The same design (rotated by 180° across its X axis) was used for applying the force at the build platform level and part top. When used, each device was installed on the build platform of the printer around the part to be tested. A nylon wire was attached to the device and linked to a mass using a pulley as depicted in Fig. 3. The pulley can freely turn around its axis, while a dedicated pulley holder linked it to the table on which the printer was installed.
The mass consisted of a plastic can filled of water. The quantity of water was adjustable to obtain the desired level of applied force. In this setup, the friction between the build platform and the force application device was neglected. This indirect force application is necessary due to the reduced accessibility inside the printer.

<figure><img src="image_96.png" /><figcaption>Fig. 2 Devices to apply a force on an entire part face (1), at the build platform level (2) and at the part top (3).</figcaption></figure>

<figure><img src="image_97.png" /><figcaption>Fig. 3 Application of the force parallel to the X axis via a pulley and a Nylon wire.</figcaption></figure>

Debonding the part can occur in several ways. Indeed, the part can completely peel off the build platform or it can just begin to do so. In both cases, the finishing of the part by a milling operation is compromised. Therefore, a dial gauge with a resolution of 0.01 mm was used to monitor the part debonding directly on the build platform by measuring the part displacement, as depicted in Fig. 4. The dial gauge was installed on a 10 mm slice of extruded polystyrene to insulate it from the build platform. The force applied by the dial gauge rack on the part at the contact point was neglected.

<figure><img src="image_98.png" /><figcaption>Fig. 4 Monitoring of the part debonding using a dial gauge (resolution: 0.01 mm).</figcaption></figure>

For each of the tests, three repetitions were performed with the same sequence:

- First, the desired part geometry is printed.
The mass consisted of a plastic can filled of water. The quantity of water was adjustable to obtain the desired level of applied force. In this setup, the friction between the build platform and the force application device was neglected. This indirect force application is necessary due to the reduced accessibility inside the printer.

<figure><img src="image_97.png" /><figcaption>Fig. 2 Devices to apply a force on an entire part face (1), at the build platform level (2) and at the part top (3).</figcaption></figure>

<figure><img src="image_98.png" /><figcaption>Fig. 3 Application of the force parallel to the X axis via a pulley and a Nylon wire.</figcaption></figure>

Debonding the part can occur in several ways. Indeed, the part can completely peel off the build platform or it can just begin to do so. In both cases, the finishing of the part by a milling operation is compromised. Therefore, a dial gauge with a resolution of 0.01 mm was used to monitor the part debonding directly on the build platform by measuring the part displacement, as depicted in Fig. 4. The dial gauge was installed on a 10 mm slice of extruded polystyrene to insulate it from the build platform. The force applied by the dial gauge rack on the part at the contact point was neglected.

<figure><img src="image_99.png" /><figcaption>Fig. 4 Monitoring of the part debonding using a dial gauge (resolution: 0.01 mm).</figcaption></figure>

For each of the tests, three repetitions were performed with the same sequence:

- First, the desired part geometry is printed.
- When the print is finished, the build platform temperature is set at the same value as during the printing (60°C) to guarantee having the same conditions as a part finishing performed between two printing operations.

- The force application device is positioned around the part to be tested.

- The dial gauge is then installed in contact with the force application device.

- Five minutes are waited to allow the dial gauge to reach a constant temperature.

- The can, already filled with water to allow a force of 2 N to be applied, is attached to the nylon wire of the setup.

- The force applied is then increased by 1 N every minute by pouring 101,94 mL in the can.

- A part displacement of 1 mm in the dial gauge is the criterion to consider the part debonded from the build platform.

## Results

**Influence of the Adhesion Strategy:** The first tests aimed to compare the effect of using an adhesion strategy or not. Six parts of Geometry A were produced with a cubic infill pattern and 95% of infill. The adhesion strategy selected to be tested was a brim of 8 mm (20 filaments roads). Three parts were printed with the brim and three without, while the force was applied on one entire face. These parameters are summed up in Table 2. The test results are depicted in Fig. 5 with ± σ error bars. The use of a brim was detrimental to the adhesion of the part. Indeed, a nearly 50% higher force can be withstood by the parts without a brim compared with those using this adhesion strategy. During the tests, the brim acted as a weakness point allowing a progressive debonding to occur. Nevertheless, this phenomenon could also result from the part design and the residual stresses it generates. Studying the influence of the part design on the brim efficiency can be an interesting perspective for this work. The repeatability of the tests without brim was also better with a relative standard deviation of 3.9% instead of 10.1% with the brim.

<table><caption>Table 2 Influence of the adhesion strategy - printing parameters.</caption><tr><td>Part geometry</td><td>Geometry A</td></tr><tr><td>Infill pattern</td><td>Cubic</td></tr><tr><td>Infill density [%]</td><td>95</td></tr><tr><td>Force application zone</td><td>One entire face (device 1)</td></tr><tr><td># roads in walls</td><td>3</td></tr><tr><td># layers in top/bottom</td><td>5</td></tr><tr><td>Adhesion by a brim</td><td>With and without</td></tr></table>

<caption>Fig. 5 Influence of the adhesion strategy on the debonding force.</caption>

**Influence of the Infill Density and Pattern:** The influence of the infill density was evaluated by printing three parts with 20% of infill and three others with 95% of infill. The design of the parts was Geometry A, no brim was used, while a cubic infill pattern was selected. The force was applied on one entire face. These parameters are depicted in Table 3. The maximal observed force before debonding is given in Fig. 6. As it can be seen in the graph, higher the infill density, better the
- When the print is finished, the build platform temperature is set at the same value as during the printing (60°C) to guarantee having the same conditions as a part finishing performed between two printing operations.

- The force application device is positioned around the part to be tested.

- The dial gauge is then installed in contact with the force application device.

- Five minutes are waited to allow the dial gauge to reach a constant temperature.

- The can, already filled with water to allow a force of 2 N to be applied, is attached to the nylon wire of the setup.

- The force applied is then increased by 1 N every minute by pouring 101,94 mL in the can.

- A part displacement of 1 mm in the dial gauge is the criterion to consider the part debonded from the build platform.

## Results

**Influence of the Adhesion Strategy:** The first tests aimed to compare the effect of using an adhesion strategy or not. Six parts of Geometry A were produced with a cubic infill pattern and 95% of infill. The adhesion strategy selected to be tested was a brim of 8 mm (20 filaments roads). Three parts were printed with the brim and three without, while the force was applied on one entire face. These parameters are summed up in Table 2. The test results are depicted in Fig. 5 with ± σ error bars. The use of a brim was detrimental to the adhesion of the part. Indeed, a nearly 50% higher force can be withstood by the parts without a brim compared with those using this adhesion strategy. During the tests, the brim acted as a weakness point allowing a progressive debonding to occur. Nevertheless, this phenomenon could also result from the part design and the residual stresses it generates. Studying the influence of the part design on the brim efficiency can be an interesting perspective for this work. The repeatability of the tests without brim was also better with a relative standard deviation of 3.9% instead of 10.1% with the brim.

<table><caption>Table 2 Influence of the adhesion strategy - printing parameters.</caption><thead><tr><th></th><th></th></tr></thead><tbody><tr><td>Part geometry</td><td>Geometry A</td></tr><tr><td>Infill pattern</td><td>Cubic</td></tr><tr><td>Infill density [%]</td><td>95</td></tr><tr><td>Force application zone</td><td>One entire face (device 1)</td></tr><tr><td># roads in walls</td><td>3</td></tr><tr><td># layers in top/bottom</td><td>5</td></tr><tr><td>Adhesion by a brim</td><td>With and without</td></tr></tbody></table><p>Debonding force [N]</p><table><thead><tr><th></th><th></th></tr></thead><tbody><tr><td>Without brim</td><td>50.59</td></tr><tr><td>With brim</td><td>33.73</td></tr></tbody></table><p>Total</p>

<caption>Fig. 5 Influence of the adhesion strategy on the debonding force.</caption>

**Influence of the Infill Density and Pattern:** The influence of the infill density was evaluated by printing three parts with 20% of infill and three others with 95% of infill. The design of the parts was Geometry A, no brim was used, while a cubic infill pattern was selected. The force was applied on one entire face. These parameters are depicted in Table 3. The maximal observed force before debonding is given in Fig. 6. As it can be seen in the graph, higher the infill density, better the
resistance of the part against the applied force. With 95% of infill, a nearly double force was withstood by the parts on average. However, the repeatability of the tests was slightly worse in this case with 4% of relative standard deviation compared to 0.2% in the case of the sparser infill.

<table><caption>Table 3 Influence of the infill density - printing parameters.</caption><tr><td>Part geometry</td><td>Geometry A</td></tr><tr><td>Infill pattern</td><td>Cubic</td></tr><tr><td>Infill density [%]</td><td>20 and 95</td></tr><tr><td>Force application zone</td><td>One entire face (device 1)</td></tr><tr><td># roads in walls</td><td>3</td></tr><tr><td># layers in top/bottom</td><td>5</td></tr><tr><td>Adhesion by a brim</td><td>Without</td></tr></table>

<figure><img src="image_99.png" /><figcaption>Fig. 6 Influence of the infill density on the debonding force.</figcaption></figure>

The infill pattern was also tested with three other parts using a concentric infill pattern and three others relying on the cubic pattern. All parts were produced with Geometry B, a cubic infill pattern with 100% of infill and the application force on an entire part face as given in Table 4, while the results are given in Fig. 7. As it can be seen, there is an influence of the infill pattern on the force which can be withstood by the part. Indeed, an almost 32.5% higher force can be withstood with the cubic infill pattern compared to the concentric. In terms of repeatability of tests, the cubic infill pattern exhibited slightly better results with a relative standard deviation of 9.5% instead of 14% in the case of the concentric pattern.

<table><caption>Table 4 Influence of the infill pattern - printing parameters.</caption><tr><td>Part geometry</td><td>Geometry B</td></tr><tr><td>Infill pattern</td><td>Cubic and concentric</td></tr><tr><td>Infill density [%]</td><td>100</td></tr><tr><td>Force application zone</td><td>Entire face (device 1)</td></tr><tr><td># roads in walls</td><td>0</td></tr><tr><td># layers in top/bottom</td><td>0</td></tr><tr><td>Adhesion by a brim</td><td>Without</td></tr></table>

<figure><img src="image_100.png" /><figcaption>Fig. 7 Influence of the infill pattern on the debonding force.</figcaption></figure>
resistance of the part against the applied force. With 95% of infill, a nearly double force was withstood by the parts on average. However, the repeatability of the tests was slightly worse in this case with 4% of relative standard deviation compared to 0.2% in the case of the sparser infill.

<table><caption>Table 3 Influence of the infill density - printing parameters.</caption><tr><td>Part geometry</td><td>Geometry A</td></tr><tr><td>Infill pattern</td><td>Cubic</td></tr><tr><td>Infill density [%]</td><td>20 and 95</td></tr><tr><td>Force application zone</td><td>One entire face (device 1)</td></tr><tr><td># roads in walls</td><td>3</td></tr><tr><td># layers in top/bottom</td><td>5</td></tr><tr><td>Adhesion by a brim</td><td>Without</td></tr></table>

<figure><img src="image_100.png" /><figcaption>Fig. 6 Influence of the infill density on the debonding force.</figcaption></figure>

The infill pattern was also tested with three other parts using a concentric infill pattern and three others relying on the cubic pattern. All parts were produced with Geometry B, a cubic infill pattern with 100% of infill and the application force on an entire part face as given in Table 4, while the results are given in Fig. 7. As it can be seen, there is an influence of the infill pattern on the force which can be withstood by the part. Indeed, an almost 32.5% higher force can be withstood with the cubic infill pattern compared to the concentric. In terms of repeatability of tests, the cubic infill pattern exhibited slightly better results with a relative standard deviation of 9.5% instead of 14% in the case of the concentric pattern.

<table><caption>Table 4 Influence of the infill pattern - printing parameters.</caption><tr><td>Part geometry</td><td>Geometry B</td></tr><tr><td>Infill pattern</td><td>Cubic and concentric</td></tr><tr><td>Infill density [%]</td><td>100</td></tr><tr><td>Force application zone</td><td>Entire face (device 1)</td></tr><tr><td># roads in walls</td><td>0</td></tr><tr><td># layers in top/bottom</td><td>0</td></tr><tr><td>Adhesion by a brim</td><td>Without</td></tr></table>

<figure><img src="image_101.png" /><figcaption>Fig. 7 Influence of the infill pattern on the debonding force.</figcaption></figure>
Influence of the Force Application Zone: The influence of the force application zone was also observed by applying a force on an entire face, at the build platform level and at the top of the cube. For each of the following tests, the part design was Geometry A with a cubic infill pattern and 20% of infill density. The parameters are summed up in Table 5.

Fig. 8, shows the results for each of the configurations. As it can be seen, the highest force is observed for the application point nearer the build platform, followed by the application distributed over an entire part's face. On average the part can endure 55% more force when it is applied the nearest from the build platform. This is logical since a force applied on part's top will generate a torque which will be detrimental to the adhesion of the part. However, even if the application point nearest the build platform led to better average results, it also exhibited the larger relative deviations with up to 13%. Relative deviations of 2.4% and less than 1% were observed for the application zones cube top and entire part's face, respectively. This may come from the temperature of the zone in the vicinity of the build platform. Indeed, a constant temperature of 60°C is maintained for the build platform. The layers in contact with this zone will then be closer to the glass transition temperature of the material, while other zones can reach lower temperatures. However, applying a force at the build platform level may result in unacceptable permanent deformation of the parts due to material softening as shown in Fig. 9a and b.

<table><caption>Table 5 Influence of the force application zone - printing parameters.</caption><tr><td>Part geometry</td><td>Geometry A</td></tr><tr><td>Infill pattern</td><td>Cubic</td></tr><tr><td>Infill density [%]</td><td>20</td></tr><tr><td>Force application zone</td><td>Entire face (device 1)<br>Build platform level (device 2)<br>Top of the cube (device 3)</td></tr><tr><td># roads in walls</td><td>3</td></tr><tr><td># layers in top/bottom</td><td>5</td></tr><tr><td>Adhesion by a brim</td><td>Without</td></tr></table>

<figure><img src="image_101.png" /><figcaption>Fig. 8 Influence of the force application zone on the debonding force.</figcaption></figure>

<figure><img src="image_102.png" /><figcaption>Fig. 9 Parts with a force applied on an entire face (a) and at the build platform level (b).</figcaption></figure>
Influence of the Force Application Zone: The influence of the force application zone was also observed by applying a force on an entire face, at the build platform level and at the top of the cube. For each of the following tests, the part design was Geometry A with a cubic infill pattern and 20% of infill density. The parameters are summed up in Table 5.

Fig. 8, shows the results for each of the configurations. As it can be seen, the highest force is observed for the application point nearer the build platform, followed by the application distributed over an entire part's face. On average the part can endure 55% more force when it is applied the nearest from the build platform. This is logical since a force applied on part's top will generate a torque which will be detrimental to the adhesion of the part. However, even if the application point nearest the build platform led to better average results, it also exhibited the larger relative deviations with up to 13%. Relative deviations of 2.4% and less than 1% were observed for the application zones cube top and entire part's face, respectively. This may come from the temperature of the zone in the vicinity of the build platform. Indeed, a constant temperature of 60°C is maintained for the build platform. The layers in contact with this zone will then be closer to the glass transition temperature of the material, while other zones can reach lower temperatures. However, applying a force at the build platform level may result in unacceptable permanent deformation of the parts due to material softening as shown in Fig. 9a and b.

<table><caption>Table 5 Influence of the force application zone - printing parameters.</caption><tr><td>Part geometry</td><td>Geometry A</td></tr><tr><td>Infill pattern</td><td>Cubic</td></tr><tr><td>Infill density [%]</td><td>20</td></tr><tr><td>Force application zone</td><td>Entire face (device 1)<br>Build platform level (device 2)<br>Top of the cube (device 3)</td></tr><tr><td># roads in walls</td><td>3</td></tr><tr><td># layers in top/bottom</td><td>5</td></tr><tr><td>Adhesion by a brim</td><td>Without</td></tr></table>

<figure><img src="image_102.png" /><figcaption>Fig. 8 Influence of the force application zone on the debonding force.</figcaption></figure>

<figure><img src="image_103.png" /><figcaption>Fig. 9 Parts with a force applied on an entire face (a) and at the build platform level (b).</figcaption></figure>
Influence of the Part Design: The part design can have a strong influence on the residual stresses in the parts [1]. This influence was observed by modifying the design of the parts while keeping all other parameters constant. Geometry A, B and C were considered for the tests. Table 6 gives the printing parameters, while Fig. 10 depicts the results of the experiments.

Geometry A and B can be compared together to highlight the influence of the overall part design (respectively cube or cube + cylinder). As it can be seen, Geometry B resulted in a better adhesion of the part to the build platform with a debonding force 48% higher than Geometry A. This better adhesion may originate from the difference of mass caused by the addition of the cylinder (33% increase of mass) and from the thermal history of the part in the vicinity of the build platform.

Comparing then the results of Geometry B and C allowed to see the influence of the chosen radius fillets. Geometry B exhibited fillets radii of 1 mm for its vertical edges, while Geometry C relied on 3 mm radii. The higher radii fillets of Geometry C allowed a 24% higher force to be withstood compared to Geometry B. Moreover, the repeatability was also improved with a relative standard deviation of 3%, 14% and 36% for Geometry C, B and A, respectively. Among the designs, the higher forces were achieved by Geometry C as well as the best repeatability.

The results obtained for the Geometry A in Fig. 10 can be compared to those of Fig. 6. Indeed, the infill pattern and force application zones are identical, while nearly the same infill density was used (95% in Fig. 6 instead of 100% in Fig. 10). However, the results are very different with a maximal force before debonding of 50.59 N and 15.88 N in the case of Fig. 6 and Fig. 10, respectively. This is mostly due to the printing parameters which were different. In Fig. 6, the number of roads for the walls stood at 3, while the bottom and top layers of the part were made of 5 layers. For the tests of Fig. 10, these values were set to zero. This shows a potential strong influence of these printing parameters on the resulting force the part can withstand. However, since the infill density are not identical, no direct conclusion can be drawn. Nevertheless, studying the influence of these parameters can be a perspective of prime interest.

<table><caption>Table 6 Influence of the part design (geometry) - printing parameters.</caption><tr><td>Part geometry</td><td>Geometry A, B and C</td></tr><tr><td>Infill pattern</td><td>Cubic</td></tr><tr><td>Infill density [%]</td><td>100</td></tr><tr><td>Force application zone</td><td>Entire face (device 1)</td></tr><tr><td># roads in walls</td><td>0</td></tr><tr><td># layers in top/bottom</td><td>0</td></tr><tr><td>Adhesion by a brim</td><td>Without</td></tr></table>

<figure><img src="image_103.png" /><figcaption>Fig. 10 Influence of the part design (geometry) on the debonding force.</figcaption></figure>

Best Configuration and discussion: The best tested configuration is the following:

- No adhesion strategy (brim).

- A cubic infill pattern with 95% of density.

- A part design of Geometry A.
Influence of the Part Design: The part design can have a strong influence on the residual stresses in the parts [1]. This influence was observed by modifying the design of the parts while keeping all other parameters constant. Geometry A, B and C were considered for the tests. Table 6 gives the printing parameters, while Fig. 10 depicts the results of the experiments.

Geometry A and B can be compared together to highlight the influence of the overall part design (respectively cube or cube + cylinder). As it can be seen, Geometry B resulted in a better adhesion of the part to the build platform with a debonding force 48% higher than Geometry A. This better adhesion may originate from the difference of mass caused by the addition of the cylinder (33% increase of mass) and from the thermal history of the part in the vicinity of the build platform.

Comparing then the results of Geometry B and C allowed to see the influence of the chosen radius fillets. Geometry B exhibited fillets radii of 1 mm for its vertical edges, while Geometry C relied on 3 mm radii. The higher radii fillets of Geometry C allowed a 24% higher force to be withstood compared to Geometry B. Moreover, the repeatability was also improved with a relative standard deviation of 3%, 14% and 36% for Geometry C, B and A, respectively. Among the designs, the higher forces were achieved by Geometry C as well as the best repeatability.

The results obtained for the Geometry A in Fig. 10 can be compared to those of Fig. 6. Indeed, the infill pattern and force application zones are identical, while nearly the same infill density was used (95% in Fig. 6 instead of 100% in Fig. 10). However, the results are very different with a maximal force before debonding of 50.59 N and 15.88 N in the case of Fig. 6 and Fig. 10, respectively. This is mostly due to the printing parameters which were different. In Fig. 6, the number of roads for the walls stood at 3, while the bottom and top layers of the part were made of 5 layers. For the tests of Fig. 10, these values were set to zero. This shows a potential strong influence of these printing parameters on the resulting force the part can withstand. However, since the infill density are not identical, no direct conclusion can be drawn. Nevertheless, studying the influence of these parameters can be a perspective of prime interest.

<table><caption>Table 6 Influence of the part design (geometry) - printing parameters.</caption><tr><td>Part geometry</td><td>Geometry A, B and C</td></tr><tr><td>Infill pattern</td><td>Cubic</td></tr><tr><td>Infill density [%]</td><td>100</td></tr><tr><td>Force application zone</td><td>Entire face (device 1)</td></tr><tr><td># roads in walls</td><td>0</td></tr><tr><td># layers in top/bottom</td><td>0</td></tr><tr><td>Adhesion by a brim</td><td>Without</td></tr></table>

<figure><img src="image_104.png" /><figcaption>Fig. 10 Influence of the part design (geometry) on the debonding force.</figcaption></figure>

Best Configuration and discussion: The best tested configuration is the following:

- No adhesion strategy (brim).

- A cubic infill pattern with 95% of density.

- A part design of Geometry A.
- A force applied on an entire part face or localized nearer the build platform.

- 3 roads for the walls and 5 layers for the bottom and top layers.

In the perspective of hybrid manufacturing, careful care will be required when foreseeing the finishing of parts produced by the printer directly on its build platform. Indeed, due to the apparition of a torque, applying forces far from the build platform is detrimental to the part adhesion. The process planning for the additive and subtractive operations should then take this observation into account.

Nevertheless, the cutting forces to finish a part made of a thermoplastic composite are relatively low, with level below 15 N [9,16]. Even if the material is not the same, it gives the order of magnitude of the forces during cutting operations. They are lower than the maximal force a part can withstand before debonding. Finish milling operations can then be foreseen.

## Conclusions

This paper presented an experimental method to evaluate the maximal force a part can withstand on a build platform just after its printing. This information is of prime interest when planning the finishing of the part directly on the build platform by milling, for example, as in hybrid machines. These are the key findings of the present study:

- In the best tested configuration: no brim, cubic infill of 95%, Geometry A with a force on an entire face, 3 roads for the walls and 5 layers for the bottom and top layers, the parts resisted to $50.59 \text{ N} \pm 1.97 \text{ N}$.

- The use of a brim is detrimental to the adhesion of the part to the build platform. It leads to progressively detach and deform.

- Higher the infill density (95% vs 20%), higher the part adhesion (48% of difference).

- Cubic infill pattern (vs concentric) ensures a 30% higher force to be withstood.

- A force nearer the build platform results in a better part adhesion.

- The part design and, potentially, its resulting residual stresses and thermal history may have an influence on its debonding behavior. However, further measurements (out of the scope of this study) may be required to assess this tendency.

- Higher fillet radii (3 mm instead of 1 mm) for the part edges in contact with the build platform allows 20% higher forces to be withstood.

## Perspectives

Comparing the results of the maximal force before debonding with the cutting forces during finish milling of part with the same material is a direct perspective of the work. Assessing the influence of other parameters influencing the part adhesion (build platform texture, number of part wall roads and top and bottom layers) can be also very interesting, while the addition of a thermal camera to the setup could bring some insights about the temperature influence.

## References

[1] D. Omidvarkarjan, R. Rosenbauer, C. Klahn, M. Meboldt, Implementation of additive manufacturing in industry, in: E. Pei, A. Bernard, D. Gu, C. Klahn, M. Monzón, M. Petersen, et al. (Eds.), Springer handbook of additive manufacturing, Cham: Springer International Publishing, 2023, pp. 55-68. https://doi.org/10.1007/978-3-031-20752-5_4

[2] Additive Manufacturing Research, Ceramics Additive Manufacturing Markets 2017-2028, an opportunity Analysis and ten-year market forecast, 2018.

[3] S.C. Altiparmak, V.A. Yardley, Z. Shi, J. Lin, Extrusion-based additive manufacturing technologies: State of the art and future perspectives, J. Manuf. Processes 83 (2022), pp. 607-636. https://doi.org/10.1016/j.jmapro.2022.09.032
- A force applied on an entire part face or localized nearer the build platform.

- 3 roads for the walls and 5 layers for the bottom and top layers.

In the perspective of hybrid manufacturing, careful care will be required when foreseeing the finishing of parts produced by the printer directly on its build platform. Indeed, due to the apparition of a torque, applying forces far from the build platform is detrimental to the part adhesion. The process planning for the additive and subtractive operations should then take this observation into account.

Nevertheless, the cutting forces to finish a part made of a thermoplastic composite are relatively low, with level below 15 N [9,16]. Even if the material is not the same, it gives the order of magnitude of the forces during cutting operations. They are lower than the maximal force a part can withstand before debonding. Finish milling operations can then be foreseen.

## Conclusions

This paper presented an experimental method to evaluate the maximal force a part can withstand on a build platform just after its printing. This information is of prime interest when planning the finishing of the part directly on the build platform by milling, for example, as in hybrid machines. These are the key findings of the present study:

- In the best tested configuration: no brim, cubic infill of 95%, Geometry A with a force on an entire face, 3 roads for the walls and 5 layers for the bottom and top layers, the parts resisted to $50.59 \text{ N} \pm 1.97 \text{ N}$.

- The use of a brim is detrimental to the adhesion of the part to the build platform. It leads to progressively detach and deform.

- Higher the infill density (95% vs 20%), higher the part adhesion (48% of difference).

- Cubic infill pattern (vs concentric) ensures a 30% higher force to be withstood.

- A force nearer the build platform results in a better part adhesion.

- The part design and, potentially, its resulting residual stresses and thermal history may have an influence on its debonding behavior. However, further measurements (out of the scope of this study) may be required to assess this tendency.

- Higher fillet radii (3 mm instead of 1 mm) for the part edges in contact with the build platform allows 20% higher forces to be withstood.

## Perspectives

Comparing the results of the maximal force before debonding with the cutting forces during finish milling of part with the same material is a direct perspective of the work. Assessing the influence of other parameters influencing the part adhesion (build platform texture, number of part wall roads and top and bottom layers) can be also very interesting, while the addition of a thermal camera to the setup could bring some insights about the temperature influence.

## References

[1] D. Omidvarkarjan, R. Rosenbauer, C. Klahn, M. Meboldt, Implementation of additive manufacturing in industry, in: E. Pei, A. Bernard, D. Gu, C. Klahn, M. Monzón, M. Petersen, et al. (Eds.), Springer handbook of additive manufacturing, Cham: Springer International Publishing, 2023, pp. 55-68. https://doi.org/10.1007/978-3-031-20752-5_4

[2] Additive Manufacturing Research, Ceramics Additive Manufacturing Markets 2017-2028, an opportunity Analysis and ten-year market forecast, 2018.

[3] S.C. Altiparmak, V.A. Yardley, Z. Shi, J. Lin, Extrusion-based additive manufacturing technologies: State of the art and future perspectives, J. Manuf. Processes 83 (2022), pp. 607-636. https://doi.org/10.1016/j.jmapro.2022.09.032
[4] D. Rosen, Design and Manufacturing Implications of Additive Manufacturing, in: D.L. Bourell, W. Frazier, H. Kuhn, M. Seifi (Eds.), Additive Manufacturing Processes, ASM International, 2020, 24. https://doi.org/10.31399/asm.hb.v24.a0006560

[5] B.N. Turner, S.A. Gold, A review of melt extrusion additive manufacturing processes: II. Materials, dimensional accuracy, and surface roughness, Rapid Prototyp. J. 21 (2015), pp. 250-261. https://doi.org/10.1108/RPJ-02-2013-0017

[6] L. Spitaels, E. Rivière-Lorphèvre, A. Demarbaix, F. Ducobu, Adaptive benchmarking design for additive manufacturing processes, Meas. Sci. Technol. 33 (2022), 064003. https://doi.org/10.1088/1361-6501/ac5877

[7] J. Hafenecker, D. Bartels, C.-M. Kuball, M. Kreß, R. Rothfelder, M. Schmidt, et al., Hybrid process chains combining metal additive manufacturing and forming CIRP J. Manuf. Sci. Technol. 46 (2023), pp. 98-115. https://doi.org/10.1016/j.cirpj.2023.08.002

[8] J.M. Flynn, A. Shokrani, S.T. Newman, V. Dhokia, Hybrid additive and subtractive machine tools - Research and industrial developments, Int. J. Mach. Tool. Manuf. 101 (2016), pp. 79-101. https://doi.org/10.1016/j.ijmachtools.2015.11.007

[9] P. Parenti, S. Cataldo, A. Grigis, M. Covelli, M. Annoni, Implementation of hybrid additive manufacturing based on extrusion of feedstock and milling, Procedia Manuf. 34 (2019) 738-746. https://doi.org/10.1016/j.promfg.2019.06.230

[10] A. Demarbaix, E. Rivière-Lorphèvre, F. Ducobu, E. Filippi, F. Petit, N. Preux, Behaviour of pre-sintered Y-TZP during machining operations: Determination of recommended cutting parameters, J. Manuf. Process. 32 (2018), pp. 85-92. https://doi.org/10.1016/j.jmapro.2018.01.020

[11] V. Martin, J.-F. Witz, F. Gillon, D. Najjar, P. Quaegebeur, A. Benabou, et al., Low-cost 3D printing of metals using filled polymer pellets. HardwareX 11 (2022), e00292. https://doi.org/10.1016/j.ohx.2022.e00292

[12] K. Rane, M. Strano, A comprehensive review of extrusion-based additive manufacturing processes for rapid production of metallic and ceramic parts, Adv. Manuf. 7 (2019), pp. 155-173. https://doi.org/10.1007/s40436-019-00253-6

[13] O. Bouzaglou, O. Golan, N. Lachman, Process Design and Parameters Interaction in Material Extrusion 3D Printing: A Review. Polymers 15 (2023) 2280. https://doi.org/10.3390/polym15102280

[14] M. Spoerk, J. Gonzalez-Gutierrez, J. Sapkota, S. Schuschnigg, C. Holzer, Effect of the printing bed temperature on the adhesion of parts produced by fused filament fabrication. Plast. Rubber Compos. 47 (2018) pp. 17-24. https://doi.org/10.1080/14658011.2017.1399531

[15] Ultimaker PLA Technical data sheet 2022.

[16] L. Spitaels, N. Aldeiturriaga, J. Bossu, G. Martic, E. Juste, P.-J. Arrazola, et al. Tool wear for finishing milling of green thermoplastic-ceramic composites fabricated with pellet AM. Procedia CIRP 121 (2024), pp. 97-102. https://doi.org/10.1016/j.procir.2023.09.235
[4] D. Rosen, Design and Manufacturing Implications of Additive Manufacturing, in: D.L. Bourell, W. Frazier, H. Kuhn, M. Seifi (Eds.), Additive Manufacturing Processes, ASM International, 2020, 24. https://doi.org/10.31399/asm.hb.v24.a0006560

[5] B.N. Turner, S.A. Gold, A review of melt extrusion additive manufacturing processes: II. Materials, dimensional accuracy, and surface roughness, Rapid Prototyp. J. 21 (2015), pp. 250-261. https://doi.org/10.1108/RPJ-02-2013-0017

[6] L. Spitaels, E. Rivière-Lorphèvre, A. Demarbaix, F. Ducobu, Adaptive benchmarking design for additive manufacturing processes, Meas. Sci. Technol. 33 (2022), 064003. https://doi.org/10.1088/1361-6501/ac5877

[7] J. Hafenecker, D. Bartels, C.-M. Kuball, M. Kreß, R. Rothfelder, M. Schmidt, et al., Hybrid process chains combining metal additive manufacturing and forming CIRP J. Manuf. Sci. Technol. 46 (2023), pp. 98-115. https://doi.org/10.1016/j.cirpj.2023.08.002

[8] J.M. Flynn, A. Shokrani, S.T. Newman, V. Dhokia, Hybrid additive and subtractive machine tools - Research and industrial developments, Int. J. Mach. Tool. Manuf. 101 (2016), pp. 79-101. https://doi.org/10.1016/j.ijmachtools.2015.11.007

[9] P. Parenti, S. Cataldo, A. Grigis, M. Covelli, M. Annoni, Implementation of hybrid additive manufacturing based on extrusion of feedstock and milling, Procedia Manuf. 34 (2019) 738-746. https://doi.org/10.1016/j.promfg.2019.06.230

[10] A. Demarbaix, E. Rivière-Lorphèvre, F. Ducobu, E. Filippi, F. Petit, N. Preux, Behaviour of pre-sintered Y-TZP during machining operations: Determination of recommended cutting parameters, J. Manuf. Process. 32 (2018), pp. 85-92. https://doi.org/10.1016/j.jmapro.2018.01.020

[11] V. Martin, J.-F. Witz, F. Gillon, D. Najjar, P. Quaegebeur, A. Benabou, et al., Low-cost 3D printing of metals using filled polymer pellets. HardwareX 11 (2022), e00292. https://doi.org/10.1016/j.ohx.2022.e00292

[12] K. Rane, M. Strano, A comprehensive review of extrusion-based additive manufacturing processes for rapid production of metallic and ceramic parts, Adv. Manuf. 7 (2019), pp. 155-173. https://doi.org/10.1007/s40436-019-00253-6

[13] O. Bouzaglou, O. Golan, N. Lachman, Process Design and Parameters Interaction in Material Extrusion 3D Printing: A Review. Polymers 15 (2023) 2280. https://doi.org/10.3390/polym15102280

[14] M. Spoerk, J. Gonzalez-Gutierrez, J. Sapkota, S. Schuschnigg, C. Holzer, Effect of the printing bed temperature on the adhesion of parts produced by fused filament fabrication. Plast. Rubber Compos. 47 (2018) pp. 17-24. https://doi.org/10.1080/14658011.2017.1399531

[15] Ultimaker PLA Technical data sheet 2022.

[16] L. Spitaels, N. Aldeiturriaga, J. Bossu, G. Martic, E. Juste, P.-J. Arrazola, et al. Tool wear for finishing milling of green thermoplastic-ceramic composites fabricated with pellet AM. Procedia CIRP 121 (2024), pp. 97-102. https://doi.org/10.1016/j.procir.2023.09.235
# Preliminary study on the impact of thermal processing on the performances of parts obtained by fused deposition modeling (FDM)

LORENZONI Margaux<fnref n="1" />,ª*, ODENT Jérémy<fnref n="2" />, RIVIÈRE-LORPHÈVRE Édouard<fnref n="1" />, SPITAELS Laurent<fnref n="1" />, HOMRANI Mohamed Khalil<fnref n="1" /> and DUCOBU François<fnref n="1" />

**Keywords:** Additive Manufacturing, Thermal Annealing, Tensile Test, Dimensions, Roughness

**Abstract.** Additive Manufacturing (AM) allows to build complex geometries while generating less waste than conventional processes such as machining. However, in AM processes such as Material Extrusion (MEX), workpieces tend to have low mechanical properties, low dimensional accuracy as well as rough surfaces. Indeed, several pre-processing and post-processing techniques exist to attempt to control these concerns. One post-processing method, namely thermal annealing, is already widely used with metallic parts. However, the influence of this technique has yet to be tested on mechanical properties such as tensile properties, dimensions, and surface roughness for polymer parts obtained by MEX and, particularly, by Fused Deposition Modelling (FDM). This paper aims to determine the relevance of using thermal annealing as a post-process to enhance FDM-obtained parts tensile properties while keeping in mind their dimensional and surface roughness aspects.

## Introduction

Context. In the last few years, additive manufacturing (AM) processes have been developed quickly in response to the increasing demand for personalized items [1]. One of the most well-known AM technologies is Material Extrusion (MEX) and, particularly, Fused Deposition Modeling (FDM) because of its low cost and ease of use for industries and private users. This process consists of building a part by successively adding layers of melted material (including polymer, composite, ceramics in thermoplastic matrix). Therefore, some geometries impossible to obtain by conventional manufacturing processes can be achieved thanks to this approach [1].

However, some limitations remain despite the wide use of this technology. Indeed, parts obtained by FDM usually exhibit low mechanical properties, low dimensional accuracy (around 0.1 mm in optimal conditions) as well as rough surfaces [1]. In addition, mechanical properties may vary widely from one part to another because of their dependence on printing parameters such as the printing temperature, build speed, and infill pattern and percentage [2]. If many studies have been seeking to optimize these parameters [2–4], research now tends to explore post-processing solutions to enhance the mechanical properties of polymer parts obtained by AM [5].

One of the possibilities currently studied in terms of post-processing is the thermal annealing of the parts [5]. This technique which is commonly used for metal workpieces consists of raising the specimen temperature after manufacturing before bringing the temperature back to ambient

<figure><img src="image_104.png" /></figure>

<footnote n="1">¹Machine Design and Production Engineering Lab – Research Institute for Science and material Engineering – University of Mons, Place du Parc 20, Mons, Belgium</footnote>

<footnote n="2">²Laboratory of Polymeric and Composite Materials (LPCM) – Center of Innovation and Research in Materials and Polymers (CIRMAP) – University of Mons, Place du Parc 20, Mons, Belgium</footnote>

<footnote>ªmargaux.lorenzoni@umons.ac.be</footnote>
# Preliminary study on the impact of thermal processing on the performances of parts obtained by fused deposition modeling (FDM)

LORENZONI Margaux<fnref n="1" />,ª*, ODENT Jérémy<fnref n="2" />, RIVIÈRE-LORPHÈVRE Édouard<fnref n="1" />, SPITAELS Laurent<fnref n="1" />, HOMRANI Mohamed Khalil<fnref n="1" /> and DUCOBU François<fnref n="1" />

**Keywords:** Additive Manufacturing, Thermal Annealing, Tensile Test, Dimensions, Roughness

**Abstract.** Additive Manufacturing (AM) allows to build complex geometries while generating less waste than conventional processes such as machining. However, in AM processes such as Material Extrusion (MEX), workpieces tend to have low mechanical properties, low dimensional accuracy as well as rough surfaces. Indeed, several pre-processing and post-processing techniques exist to attempt to control these concerns. One post-processing method, namely thermal annealing, is already widely used with metallic parts. However, the influence of this technique has yet to be tested on mechanical properties such as tensile properties, dimensions, and surface roughness for polymer parts obtained by MEX and, particularly, by Fused Deposition Modelling (FDM). This paper aims to determine the relevance of using thermal annealing as a post-process to enhance FDM-obtained parts tensile properties while keeping in mind their dimensional and surface roughness aspects.

## Introduction

Context. In the last few years, additive manufacturing (AM) processes have been developed quickly in response to the increasing demand for personalized items [1]. One of the most well-known AM technologies is Material Extrusion (MEX) and, particularly, Fused Deposition Modeling (FDM) because of its low cost and ease of use for industries and private users. This process consists of building a part by successively adding layers of melted material (including polymer, composite, ceramics in thermoplastic matrix). Therefore, some geometries impossible to obtain by conventional manufacturing processes can be achieved thanks to this approach [1].

However, some limitations remain despite the wide use of this technology. Indeed, parts obtained by FDM usually exhibit low mechanical properties, low dimensional accuracy (around 0.1 mm in optimal conditions) as well as rough surfaces [1]. In addition, mechanical properties may vary widely from one part to another because of their dependence on printing parameters such as the printing temperature, build speed, and infill pattern and percentage [2]. If many studies have been seeking to optimize these parameters [2–4], research now tends to explore post-processing solutions to enhance the mechanical properties of polymer parts obtained by AM [5].

One of the possibilities currently studied in terms of post-processing is the thermal annealing of the parts [5]. This technique which is commonly used for metal workpieces consists of raising the specimen temperature after manufacturing before bringing the temperature back to ambient

<figure><img src="image_105.png" /></figure>

<footnote n="1">¹Machine Design and Production Engineering Lab – Research Institute for Science and material Engineering – University of Mons, Place du Parc 20, Mons, Belgium</footnote>

<footnote n="2">²Laboratory of Polymeric and Composite Materials (LPCM) – Center of Innovation and Research in Materials and Polymers (CIRMAP) – University of Mons, Place du Parc 20, Mons, Belgium</footnote>

<footnote>ªmargaux.lorenzoni@umons.ac.be</footnote>
temperature. However, studies regarding thermal annealing applied to AM parts are only starting to emerge.

Literature review. Investigations about thermal annealing of Polylactic Acid (PLA) are still in their early stages. Annealing temperatures for PLA used in literature vary from 60°C in Behzadnasab et al. [6] to 160°C in Jo et al. [7] and in Hart et al. [8]. In fact, lower temperature ranges allow to scan effects of temperatures around the PLA glass transition temperature ($T_g \approx 60°C$). On the other hand, the highest temperatures tested provide information on effects of temperatures reaching above the melting temperature ($T_m \approx 150°C$).

Regarding annealing duration for PLA, intervals of values tested in literature are also quite broad. In Wach et al. [9], samples were annealed for as few as 15 seconds. On the other side of the spectrum, the maximum annealing time (2 hours) was reported in Hart et al. [8] and in Slavkovic et al. [10].

Overall, results provided by the different studies are mitigated. Slavkovic et al. [10] saw an increase of 19% in Young's Modulus and an increase of 30.25% in Ultimate Tensile Strength (UTS). On the contrary, Behzadnasab et al. [6] reported no significant impact on Young's Modulus for the same ranges in terms of annealing temperature and duration.

Others explore the effect of thermal annealing with other types of mechanical tests such as flexural [8, 9, 11] and torsion tests [12]. According to Hart et al. [8], semi-crystalline materials like PLA will remain in an amorphous state if they are cooled quickly. On the other hand, if they are cooled slowly (10°C/min or slower), a crystalline domain will form, leading to an increase in toughness. In Wach et al. [9], the main topic discussed is the annealing time with respect to the annealing temperature. For instance, it is demonstrated that the crystallinity degree does not change at a temperature too close to the glass transition temperature $T_g$ (in the study, at 60 and 70°C). Therefore, it leads to very few effects on the mechanical properties.

An interesting point to note is that devices are used in a couple of studies to hold the samples during annealing. For instance, a mold is used in Rane et al. [13] to avoid distortions caused by thermal processing. Similarly, the use of a custom fixture is mentioned in Hart et al. [8]. In a previous study, Hart et al. attempted to anneal one ABS sample for 18 hours at 175°C without using the annealing fixture [14]. As a result, the sample was distorted and became unusable for flexural testing. According to their analysis, the deformation resulted from the release of residual stresses from the initial manufacturing process and the polymer's own weight. The hypothesis of residual stresses is also mentioned by Butt et al. [15] and Behzadnesab et al. [6] in their respective research although they do not mention holding their samples during annealing. The custom fixtures used are different depending on the studies since they were made specifically by the researchers.

Goal and motivations of the study. This study aims to determine the relevance of using thermal annealing as a post-process to enhance FDM-obtained parts tensile properties while keeping in mind their dimensional and surface roughness aspects that have not been studied so far.

## Material and Method

Part printing. The parts were printed from 2.85 mm-diameter filaments of blue Ultimaker Polylactic Acid (PLA). The printer used was an Ultimaker S3 with AA 0.4 mm print cores. A light mist of 3DLAC was applied on the build plate before each print to avoid warping. Printing parameters were chosen according to the manufacturer recommendations, given in the Cura 4.13.1 software for the “Fine” resolution. The initial layer height was modified from the default parameters (0.2 mm) to avoid creating a shell effect. The infill pattern was changed to “Lines” and oriented in the tensile test traction direction. Other printing parameters are given in Table 1.
temperature. However, studies regarding thermal annealing applied to AM parts are only starting to emerge.

Literature review. Investigations about thermal annealing of Polylactic Acid (PLA) are still in their early stages. Annealing temperatures for PLA used in literature vary from 60°C in Behzadnasab et al. [6] to 160°C in Jo et al. [7] and in Hart et al. [8]. In fact, lower temperature ranges allow to scan effects of temperatures around the PLA glass transition temperature ($T_g \approx 60°C$). On the other hand, the highest temperatures tested provide information on effects of temperatures reaching above the melting temperature ($T_m \approx 150°C$).

Regarding annealing duration for PLA, intervals of values tested in literature are also quite broad. In Wach et al. [9], samples were annealed for as few as 15 seconds. On the other side of the spectrum, the maximum annealing time (2 hours) was reported in Hart et al. [8] and in Slavkovic et al. [10].

Overall, results provided by the different studies are mitigated. Slavkovic et al. [10] saw an increase of 19% in Young's Modulus and an increase of 30.25% in Ultimate Tensile Strength (UTS). On the contrary, Behzadnasab et al. [6] reported no significant impact on Young's Modulus for the same ranges in terms of annealing temperature and duration.

Others explore the effect of thermal annealing with other types of mechanical tests such as flexural [8, 9, 11] and torsion tests [12]. According to Hart et al. [8], semi-crystalline materials like PLA will remain in an amorphous state if they are cooled quickly. On the other hand, if they are cooled slowly (10°C/min or slower), a crystalline domain will form, leading to an increase in toughness. In Wach et al. [9], the main topic discussed is the annealing time with respect to the annealing temperature. For instance, it is demonstrated that the crystallinity degree does not change at a temperature too close to the glass transition temperature $T_g$ (in the study, at 60 and 70°C). Therefore, it leads to very few effects on the mechanical properties.

An interesting point to note is that devices are used in a couple of studies to hold the samples during annealing. For instance, a mold is used in Rane et al. [13] to avoid distortions caused by thermal processing. Similarly, the use of a custom fixture is mentioned in Hart et al. [8]. In a previous study, Hart et al. attempted to anneal one ABS sample for 18 hours at 175°C without using the annealing fixture [14]. As a result, the sample was distorted and became unusable for flexural testing. According to their analysis, the deformation resulted from the release of residual stresses from the initial manufacturing process and the polymer's own weight. The hypothesis of residual stresses is also mentioned by Butt et al. [15] and Behzadnesab et al. [6] in their respective research although they do not mention holding their samples during annealing. The custom fixtures used are different depending on the studies since they were made specifically by the researchers.

Goal and motivations of the study. This study aims to determine the relevance of using thermal annealing as a post-process to enhance FDM-obtained parts tensile properties while keeping in mind their dimensional and surface roughness aspects that have not been studied so far.

## Material and Method

Part printing. The parts were printed from 2.85 mm-diameter filaments of blue Ultimaker Polylactic Acid (PLA). The printer used was an Ultimaker S3 with AA 0.4 mm print cores. A light mist of 3DLAC was applied on the build plate before each print to avoid warping. Printing parameters were chosen according to the manufacturer recommendations, given in the Cura 4.13.1 software for the “Fine” resolution. The initial layer height was modified from the default parameters (0.2 mm) to avoid creating a shell effect. The infill pattern was changed to “Lines” and oriented in the tensile test traction direction. Other printing parameters are given in Table 1.
<table><caption>Table 1 - Printing parameters</caption><tr><td>Layer height (including Initial layer height)</td><td>0.1 mm</td></tr><tr><td>Infill density</td><td>95%</td></tr><tr><td>Infill Line Directions (including Top/Bottom Lines)</td><td>Traction direction</td></tr><tr><td>Build orientation</td><td>Flat (XY)</td></tr><tr><td>Nozzle temperature</td><td>200°C</td></tr><tr><td>Build plate temperature</td><td>85°C</td></tr><tr><td>Printing speed</td><td>70 mm/s</td></tr></table>

Surface roughness and dimensions assessment. A Wurth caliper was used to measure specimens with a scale of 0.05 mm. All dimension measurements were evaluated three times and the average value of the three measurements was kept as data. A Carl Zeiss Handysurf E-35A was employed to evaluate the arithmetic surface roughness $R_a$ of the specimens top and bottom faces (corresponding respectively to the last and first layers printed). These values were measured following the guidance of ISO 4288 [16]. Measurements of surface roughness and specimen dimensions were assessed before and after thermal annealing.

Thermal annealing. Annealing was conducted using a Binder FD115 E2 Proofer. The aim was to assess the influence of annealing temperature and duration on tensile properties, sample dimensions, and surface roughness. Therefore, a total of 10 conditions were employed. 9 conditions resulted from the combination of 3 levels of annealing temperature and duration while the last one consisted of leaving the samples unannealed as a control condition. Each condition was repeated over 5 samples. Intervals of temperature and dimensions were chosen according to literature. The thermal annealing parameters used in this study are given in Table 2. The proofer was preheated at the given temperature before placing the samples for annealing.

<table><caption>Table 2 - Thermal annealing parameters</caption><thead><tr><th colspan="4">Control specimens: Category 0</th></tr><tr><th></th><th>100°C</th><th>120°C</th><th>140°C</th></tr></thead><tbody><tr><td>10 min</td><td>Category 1</td><td>Category 2</td><td>Category 3</td></tr><tr><td>35 min</td><td>Category 4</td><td>Category 5</td><td>Category 6</td></tr><tr><td>60 min</td><td>Category 7</td><td>Category 8</td><td>Category 9</td></tr></tbody></table>

At the end of the annealing time, the specimens were left in the fixture until they reached room temperature, which lasted around an hour. Indeed, according to Hart et al. [8], for semi-crystalline materials like PLA a slow cooling rate (10°C/min or slower) has a beneficial impact on the crystallinity and therefore on mechanical properties of the material.

Tensile testing. A Zwick Roell Z2.5 bench was employed with an Xforce 2.5 kN load cell and Type 8297 pneumatic grips from the same manufacturer. Test speed was set to 3 mm/s with respect to the ASTM D638 standard [17]. Type IV of this standard geometry was employed (Figure 1). testXpert II was the software used to retrieve data in terms of stress $\sigma$ [MPa] and strain $\epsilon$ [%]. Data was later processed to compute Young's Modulus $E$ [MPa], Ultimate tensile strength UTS [MPa] and the elongation at break $A$ [%]. Tensile tests were performed on 5 samples per thermal annealing category. Therefore, mean values and standard deviation of the different tensile properties cited were also computed.
<table><caption>Table 1 - Printing parameters</caption><tr><td>Layer height (including Initial layer height)</td><td>0.1 mm</td></tr><tr><td>Infill density</td><td>95%</td></tr><tr><td>Infill Line Directions (including Top/Bottom Lines)</td><td>Traction direction</td></tr><tr><td>Build orientation</td><td>Flat (XY)</td></tr><tr><td>Nozzle temperature</td><td>200°C</td></tr><tr><td>Build plate temperature</td><td>85°C</td></tr><tr><td>Printing speed</td><td>70 mm/s</td></tr></table>

Surface roughness and dimensions assessment. A Wurth caliper was used to measure specimens with a scale of 0.05 mm. All dimension measurements were evaluated three times and the average value of the three measurements was kept as data. A Carl Zeiss Handysurf E-35A was employed to evaluate the arithmetic surface roughness $R_a$ of the specimens top and bottom faces (corresponding respectively to the last and first layers printed). These values were measured following the guidance of ISO 4288 [16]. Measurements of surface roughness and specimen dimensions were assessed before and after thermal annealing.

Thermal annealing. Annealing was conducted using a Binder FD115 E2 Proofer. The aim was to assess the influence of annealing temperature and duration on tensile properties, sample dimensions, and surface roughness. Therefore, a total of 10 conditions were employed. 9 conditions resulted from the combination of 3 levels of annealing temperature and duration while the last one consisted of leaving the samples unannealed as a control condition. Each condition was repeated over 5 samples. Intervals of temperature and dimensions were chosen according to literature. The thermal annealing parameters used in this study are given in Table 2. The proofer was preheated at the given temperature before placing the samples for annealing.

<table><caption>Table 2 - Thermal annealing parameters</caption><thead><tr><th colspan="4">Control specimens: Category 0</th></tr><tr><th></th><th>100°C</th><th>120°C</th><th>140°C</th></tr></thead><tbody><tr><th>10 min</th><td>Category 1</td><td>Category 2</td><td>Category 3</td></tr><tr><th>35 min</th><td>Category 4</td><td>Category 5</td><td>Category 6</td></tr><tr><th>60 min</th><td>Category 7</td><td>Category 8</td><td>Category 9</td></tr></tbody></table>

At the end of the annealing time, the specimens were left in the fixture until they reached room temperature, which lasted around an hour. Indeed, according to Hart et al. [8], for semi-crystalline materials like PLA a slow cooling rate (10°C/min or slower) has a beneficial impact on the crystallinity and therefore on mechanical properties of the material.

Tensile testing. A Zwick Roell Z2.5 bench was employed with an Xforce 2.5 kN load cell and Type 8297 pneumatic grips from the same manufacturer. Test speed was set to 3 mm/s with respect to the ASTM D638 standard [17]. Type IV of this standard geometry was employed (Figure 1). testXpert II was the software used to retrieve data in terms of stress $\sigma$ [MPa] and strain $\epsilon$ [%]. Data was later processed to compute Young's Modulus $E$ [MPa], Ultimate tensile strength UTS [MPa] and the elongation at break $A$ [%]. Tensile tests were performed on 5 samples per thermal annealing category. Therefore, mean values and standard deviation of the different tensile properties cited were also computed.
<figure><img src="image_105.png" /><figcaption>Figure 1 - Specimen dimensions in mm from ASTM D638 Type IV [17]</figcaption></figure>

## Results and Discussion

Thermal annealing fixture. As mentioned in the literature review, some kind of mold or fixture may be employed to hold specimens during annealing. Indeed, studies [13, 14] reported geometry distortions because of the heating process. Consequently, samples were first annealed without any kind of device to hold them to assess the distortion scope in this case. Two batches of 5 specimens were annealed without fixture and with two different placements. The first batch was placed in its build orientation (first layer printed at the bottom, last layer printed at the top, Figure 2). The second batch was placed upside-down compared to its build direction (first layer printed at the top, last layer printed at the bottom, Figure 3).

The specimens did not keep their original shape as expected from literature. Indeed, no matter the annealing orientation of the samples, the latter curled towards the first layer printed (Figure 2 and Figure 3). This phenomenon may be due to residual stresses caused by the printing process. Particularly, it can be induced by the temperature gradient between the bottom layers kept heated by the build plate and the top layers being printed.

<figure><img src="image_106.png" /><figcaption>Figure 2 - Samples annealed on their build plate side</figcaption></figure>

<figure><img src="image_107.png" /><figcaption>Figure 3 - Samples annealed on their top side</figcaption></figure>

Other than the curvature making the length of the samples not measurable, the specimen dimensions, such as the width of the ends (WO, Figure 1) and narrow part (Wc, Figure 1) as well as the thickness (T, Figure 1), remained within the tolerance interval allowed by the tensile test standard ASTM D638 [17].

Following these observations, a fixture (Figure 4) was designed to maintain the samples flat during annealing. Compared to fixtures showed in literature [13, 14], the design used for this study
<figure><img src="image_106.png" /><figcaption>Figure 1 - Specimen dimensions in mm from ASTM D638 Type IV [17]</figcaption></figure>

## Results and Discussion

Thermal annealing fixture. As mentioned in the literature review, some kind of mold or fixture may be employed to hold specimens during annealing. Indeed, studies [13, 14] reported geometry distortions because of the heating process. Consequently, samples were first annealed without any kind of device to hold them to assess the distortion scope in this case. Two batches of 5 specimens were annealed without fixture and with two different placements. The first batch was placed in its build orientation (first layer printed at the bottom, last layer printed at the top, Figure 2). The second batch was placed upside-down compared to its build direction (first layer printed at the top, last layer printed at the bottom, Figure 3).

The specimens did not keep their original shape as expected from literature. Indeed, no matter the annealing orientation of the samples, the latter curled towards the first layer printed (Figure 2 and Figure 3). This phenomenon may be due to residual stresses caused by the printing process. Particularly, it can be induced by the temperature gradient between the bottom layers kept heated by the build plate and the top layers being printed.

<figure><img src="image_107.png" /><figcaption>Figure 2 - Samples annealed on their build plate side</figcaption></figure>

<figure><img src="image_108.png" /><figcaption>Figure 3 - Samples annealed on their top side</figcaption></figure>

Other than the curvature making the length of the samples not measurable, the specimen dimensions, such as the width of the ends (WO, Figure 1) and narrow part (Wc, Figure 1) as well as the thickness (T, Figure 1), remained within the tolerance interval allowed by the tensile test standard ASTM D638 [17].

Following these observations, a fixture (Figure 4) was designed to maintain the samples flat during annealing. Compared to fixtures showed in literature [13, 14], the design used for this study
also intends to maximize the heat exchange by using thin metal plates ensuring good conductivity as well as the largest contact surface possible. As a result, this fixture is made from two aluminum Al 6082 T6 plates of 5 mm of thickness where screws were evenly distributed (spaced by 44 mm) so that each sample was held in similar conditions. Therefore, this fixture was used to perform thermal annealing for the whole study. The nuts were screwed by hand, therefore the torque applied on them was not specific. A perspective for this study could be to perform it again using a consistent torque on the nuts.

<figure><img src="image_108.png" /><figcaption>Figure 4 - Fixture used to anneal the specimens</figcaption></figure>

Thermal annealing. In terms of Young's Modulus, specimens annealed for 10 minutes (Category 1 through 3) show a decrease in property values. Overall, a trend seems to exist as Young's Modulus increases with annealing temperature for the same annealing duration (Figure 5). These results need to be put into perspective as the maximum increase in Young's Modulus is only about 6.5%.

<figure><img src="image_109.png" /><figcaption>Figure 5 - Young's Modulus comparison between categories</figcaption></figure>

Regarding the UTS, it does not seem to follow any specific trend (Figure 6). In addition, all the standard deviation intervals of annealed samples overlap the standard deviation of the control specimens.
also intends to maximize the heat exchange by using thin metal plates ensuring good conductivity as well as the largest contact surface possible. As a result, this fixture is made from two aluminum Al 6082 T6 plates of 5 mm of thickness where screws were evenly distributed (spaced by 44 mm) so that each sample was held in similar conditions. Therefore, this fixture was used to perform thermal annealing for the whole study. The nuts were screwed by hand, therefore the torque applied on them was not specific. A perspective for this study could be to perform it again using a consistent torque on the nuts.

<figure><img src="image_109.png" /><figcaption>Figure 4 - Fixture used to anneal the specimens</figcaption></figure>

Thermal annealing. In terms of Young's Modulus, specimens annealed for 10 minutes (Category 1 through 3) show a decrease in property values. Overall, a trend seems to exist as Young's Modulus increases with annealing temperature for the same annealing duration (Figure 5). These results need to be put into perspective as the maximum increase in Young's Modulus is only about 6.5%.

<figure><img src="image_110.png" /><figcaption>Figure 5 - Young's Modulus comparison between categories</figcaption></figure>

Regarding the UTS, it does not seem to follow any specific trend (Figure 6). In addition, all the standard deviation intervals of annealed samples overlap the standard deviation of the control specimens.
<figure><img src="image_110.png" /><figcaption>Figure 6 - UTS comparison between categories</figcaption></figure>

As for the elongation at break, it appears to follow the opposite trend to that of Young's Modulus (Figure 7). Indeed, 10 minutes of thermal annealing seem to have a positive effect on the elongation at break. For instance, it leads to a 34.7% increase. The graph also exhibits that the elongation at break is decreasing with the annealing temperature for the same annealing duration. This correlation between elongation at break and Young's Modulus is not surprising. Indeed, materials generally tend to have a more brittle behavior when their plastic properties increase.

<figure><img src="image_111.png" /><figcaption>Figure 7 - Elongation at break comparison between categories</figcaption></figure>

Dimensions assessment. Dimensions were compared batch per batch to avoid the influence of any disparities between batches. As a reminder, each batch contained five specimens and was subjected to the categories number from 1 to 9 (Table 2).

For each dimension type (LO, WO, Wc and T, Figure 1), the difference between the mean value for one batch before annealing and the theoretical value was compared to the mean measured value for one batch after annealing and the theoretical dimension. These differences were also compared with the ISO 2768-1 tolerances classes [18]. As a reminder, the intervals allowed by tolerances classes depend on nominal linear dimensions. Consequently, the class intervals differ for LO and WO compared to Wc and T.

Overall, the mean LO, WO and Wc were decreased by annealing according to the graphs (Figure 8) although they tend to remain in the same tolerance class due to the great standard deviation. In terms of thickness T, no specific trend was established although most samples appear to keep their initial thickness. However, the lower deviation regarding thickness could be linked to the fixture use since it does not allow samples to expand in this direction. As a reminder, samples needed to be maintained in a fixture to avoid warping and large distortions. In fact, this step is necessary to obtain usable samples with respect to the ASTM D638 tensile test standard [17]. It
<figure><img src="image_111.png" /><figcaption>Figure 6 - UTS comparison between categories</figcaption></figure>

As for the elongation at break, it appears to follow the opposite trend to that of Young's Modulus (Figure 7). Indeed, 10 minutes of thermal annealing seem to have a positive effect on the elongation at break. For instance, it leads to a 34.7% increase. The graph also exhibits that the elongation at break is decreasing with the annealing temperature for the same annealing duration. This correlation between elongation at break and Young's Modulus is not surprising. Indeed, materials generally tend to have a more brittle behavior when their plastic properties increase.

<figure><img src="image_112.png" /><figcaption>Figure 7 - Elongation at break comparison between categories</figcaption></figure>

Dimensions assessment. Dimensions were compared batch per batch to avoid the influence of any disparities between batches. As a reminder, each batch contained five specimens and was subjected to the categories number from 1 to 9 (Table 2).

For each dimension type (LO, WO, Wc and T, Figure 1), the difference between the mean value for one batch before annealing and the theoretical value was compared to the mean measured value for one batch after annealing and the theoretical dimension. These differences were also compared with the ISO 2768-1 tolerances classes [18]. As a reminder, the intervals allowed by tolerances classes depend on nominal linear dimensions. Consequently, the class intervals differ for LO and WO compared to Wc and T.

Overall, the mean LO, WO and Wc were decreased by annealing according to the graphs (Figure 8) although they tend to remain in the same tolerance class due to the great standard deviation. In terms of thickness T, no specific trend was established although most samples appear to keep their initial thickness. However, the lower deviation regarding thickness could be linked to the fixture use since it does not allow samples to expand in this direction. As a reminder, samples needed to be maintained in a fixture to avoid warping and large distortions. In fact, this step is necessary to obtain usable samples with respect to the ASTM D638 tensile test standard [17]. It
should be noted that the overall decrease in specimen dimensions could be linked to the decrease in the initial voids between layers thanks to thermal annealing. In fact, samples were printed with an infill density of 95% (Table 1) making them porous. Reduction of porosity is, indeed, generally a way of increasing mechanical properties.

<figure><img src="image_112.png" /><figcaption>Figure 8 - Example of dimensions comparison graph before and after thermal annealing, here for WO</figcaption></figure>

In terms of dimensions tolerances after annealing, the samples exhibit dimensions that fall within the medium (class m, for WO and T) or coarse (class c, for LO and Wc) tolerance classes of ISO 2768-1. This should be considered in addition to the printing tolerances although measured values stay in the same order of magnitude before and after annealing.

As for the relative standard deviations, trends were not observed for LO and WO. On the contrary, thermal annealing occurs to have a negative impact on the relative standard deviation of Wc which is enlarged at least by 6.5% and up to 23.9%.

Part surface roughness. The arithmetic surface roughness $R_a$ was considered. Values of $R_a$ measured on samples were compared to classes based on the intervals of $R_a$ mentioned in ISO 1302 [19]. For each batch of 5 samples, the mean $R_a$ over the batch was compared before and after annealing. In Figure 9, the graph highlights the fact that thermal annealing tends to decrease the surface roughness of samples' top face. For instance, the general trend is for the mean surface roughness to go down a class after annealing. The only case for which the conclusion cannot be applied is samples belonging to Category 1 which corresponds to the shortest annealing time and lowest annealing temperature.
should be noted that the overall decrease in specimen dimensions could be linked to the decrease in the initial voids between layers thanks to thermal annealing. In fact, samples were printed with an infill density of 95% (Table 1) making them porous. Reduction of porosity is, indeed, generally a way of increasing mechanical properties.

<figure><img src="image_113.png" /><figcaption>Figure 8 - Example of dimensions comparison graph before and after thermal annealing, here for WO</figcaption></figure>

In terms of dimensions tolerances after annealing, the samples exhibit dimensions that fall within the medium (class m, for WO and T) or coarse (class c, for LO and Wc) tolerance classes of ISO 2768-1. This should be considered in addition to the printing tolerances although measured values stay in the same order of magnitude before and after annealing.

As for the relative standard deviations, trends were not observed for LO and WO. On the contrary, thermal annealing occurs to have a negative impact on the relative standard deviation of Wc which is enlarged at least by 6.5% and up to 23.9%.

Part surface roughness. The arithmetic surface roughness $R_a$ was considered. Values of $R_a$ measured on samples were compared to classes based on the intervals of $R_a$ mentioned in ISO 1302 [19]. For each batch of 5 samples, the mean $R_a$ over the batch was compared before and after annealing. In Figure 9, the graph highlights the fact that thermal annealing tends to decrease the surface roughness of samples' top face. For instance, the general trend is for the mean surface roughness to go down a class after annealing. The only case for which the conclusion cannot be applied is samples belonging to Category 1 which corresponds to the shortest annealing time and lowest annealing temperature.
<figure><img src="image_113.png" /><figcaption>Figure 9 - Comparison of the top face roughness before and after thermal annealing</figcaption></figure>

For the bottom surface, the effects of thermal annealing on surface roughness are inconsistent. For some categories, surface roughness decreases while, for others, it increases marginally or significantly. However, even for the most important surface roughness changes, values for each batch stay in the same order of magnitude and therefore in the same class before and after thermal annealing.

## Summary

Conclusion. The main findings are:

- Distortions appear if the parts are annealed without being held. Therefore, a fixture needs to be designed and used. This need for a fixture during thermal annealing could limit the use of this post-process. Indeed, one of the main advantages of AM is to allow the manufacture of complex shaped parts for which it is not possible to create a mold.

- Annealing tends to increase Young's Modulus of the PLA if the annealing temperature is above its crystallization temperature, with a maximum increase of 6.5% compared to unannealed parts.

- For the elongation at break, the progression is opposite to the one showed by Young's Modulus with a maximal increase in elongation at break of 34.7%.

- In terms of dimensions, overall, the length LO and widths (of the ends Wo and narrow parts Wc of the specimens) appear to be decreasing while the thickness T remains unchanged in most cases.

- As for the surface roughness of the parts, the top and bottom surface roughness need to be addressed separately. The surface roughness of the samples' top face occurs to be reduced by annealing although staying in the same surface roughness class, while no significant change appears on the bottom face.

Perspectives. These are the main perspectives of the work:

- Finding optimized annealing parameters based on the most promising parameters found in this study.

- Research could be lead on ways to avoid the build-up of residual stresses during printing such as specific printing parameters or conditions to avoid distortions during thermal annealing.
<figure><img src="image_114.png" /><figcaption>Figure 9 - Comparison of the top face roughness before and after thermal annealing</figcaption></figure>

For the bottom surface, the effects of thermal annealing on surface roughness are inconsistent. For some categories, surface roughness decreases while, for others, it increases marginally or significantly. However, even for the most important surface roughness changes, values for each batch stay in the same order of magnitude and therefore in the same class before and after thermal annealing.

## Summary

Conclusion. The main findings are:

- Distortions appear if the parts are annealed without being held. Therefore, a fixture needs to be designed and used. This need for a fixture during thermal annealing could limit the use of this post-process. Indeed, one of the main advantages of AM is to allow the manufacture of complex shaped parts for which it is not possible to create a mold.

- Annealing tends to increase Young's Modulus of the PLA if the annealing temperature is above its crystallization temperature, with a maximum increase of 6.5% compared to unannealed parts.

- For the elongation at break, the progression is opposite to the one showed by Young's Modulus with a maximal increase in elongation at break of 34.7%.

- In terms of dimensions, overall, the length LO and widths (of the ends Wo and narrow parts Wc of the specimens) appear to be decreasing while the thickness T remains unchanged in most cases.

- As for the surface roughness of the parts, the top and bottom surface roughness need to be addressed separately. The surface roughness of the samples' top face occurs to be reduced by annealing although staying in the same surface roughness class, while no significant change appears on the bottom face.

Perspectives. These are the main perspectives of the work:

- Finding optimized annealing parameters based on the most promising parameters found in this study.

- Research could be lead on ways to avoid the build-up of residual stresses during printing such as specific printing parameters or conditions to avoid distortions during thermal annealing.
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[4] C. Oliveira, J. Rocha, and J. E. Ribeiro, "Mechanical and Physical Characterization of Parts Manufactured by 3D Printing," in *Materials Design and Applications IV*, L. F. M. da Silva, Ed., in Advanced Structured Materials. Cham: Springer International Publishing, 2023, pp. 77-88. https://doi.org/10.1007/978-3-031-18130-6_6

[5] R. V. Pazhamannil, J. N. V. N., G. P., and A. Edacherian, "Property enhancement approaches of fused filament fabrication technology: A review," *Polym. Eng. Sci.*, vol. 62, no. 5, pp. 1356-1376, 2022. https://doi.org/10.1002/pen.25948

[6] M. Behzadnasab, A. A. Yousefi, D. Ebrahimibagha, and F. Nasiri, "Effects of processing conditions on mechanical properties of PLA printed parts," *Rapid Prototyp. J.*, vol. 26, no. 2, pp. 381-389, Jan. 2019. https://doi.org/10.1108/RPJ-02-2019-0048

[7] W. Jo, O.-C. Kwon, and M.-W. Moon, "Investigation of influence of heat treatment on mechanical strength of FDM printed 3D objects," *Rapid Prototyp. J.*, vol. 24, no. 3, pp. 637-644, Jan. 2018. https://doi.org/10.1108/RPJ-06-2017-0131

[8] K. R. Hart, R. M. Dunn, and E. D. Wetzel, "Increased fracture toughness of additively manufactured semi-crystalline thermoplastics via thermal annealing," *Polymer*, vol. 211, p. 123091, Dec. 2020. https://doi.org/10.1016/j.polymer.2020.123091

[9] R. A. Wach, P. Wolszczak, and A. Adamus-Wlodarczyk, "Enhancement of Mechanical Properties of FDM-PLA Parts via Thermal Annealing," *Macromol. Mater. Eng.*, vol. 303, no. 9, p. 1800169, 2018. https://doi.org/10.1002/mame.201800169

[10] V. Slavković, N. Grujović, A. Disic, and A. Radovanović, "Influence of Annealing and Printing Directions on Mechanical Properties of PLA Shape Memory Polymer Produced by Fused Deposition Modeling," 2017.

[11] W. Yu, X. Wang, X. Yin, E. Ferraris, and J. Zhang, "The effects of thermal annealing on the performance of material extrusion 3D printed polymer parts," *Mater. Des.*, vol. 226, p. 111687, Feb. 2023. https://doi.org/10.1016/j.matdes.2023.111687

[12] J. Torres, J. Cotelo, J. Karl, and A. P. Gordon, "Mechanical Property Optimization of FDM PLA in Shear with Multiple Objectives," *JOM*, vol. 67, no. 5, pp. 1183-1193, May 2015. https://doi.org/10.1007/s11837-015-1367-y

[13] R. Rane, A. Kulkarni, H. Prajapati, R. Taylor, A. Jain, and V. Chen, "Post-Process Effects of Isothermal Annealing and Initially Applied Static Uniaxial Loading on the Ultimate Tensile Strength of Fused Filament Fabrication Parts," *Materials*, vol. 13, p. 352, Jan. 2020. https://doi.org/10.3390/ma13020352
References

[1] E. Pei et al., Eds., *Springer Handbook of Additive Manufacturing*. in Springer Handbooks. Cham: Springer International Publishing, 2023. https://doi.org/10.1007/978-3-031-20752-5

[2] S. L. Rodríguez-Reyna, C. Mata, J. H. Díaz-Aguilera, H. R. Acevedo-Parra, and F. Tapia, “Mechanical properties optimization for PLA, ABS and Nylon + CF manufactured by 3D FDM printing,” *Mater. Today Commun.*, vol. 33, p. 104774, Dec. 2022. https://doi.org/10.1016/j.mtcomm.2022.104774

[3] A. García-Domínguez, J. Claver, A. M. Camacho, and M. A. Sebastián, “Considerations on the Applicability of Test Methods for Mechanical Characterization of Materials Manufactured by FDM,” *Materials*, vol. 13, no. 1, Art. no. 1, Jan. 2020. https://doi.org/10.3390/ma13010028

[4] C. Oliveira, J. Rocha, and J. E. Ribeiro, “Mechanical and Physical Characterization of Parts Manufactured by 3D Printing,” in *Materials Design and Applications IV*, L. F. M. da Silva, Ed., in Advanced Structured Materials. Cham: Springer International Publishing, 2023, pp. 77–88. https://doi.org/10.1007/978-3-031-18130-6_6

[5] R. V. Pazhamannil, J. N. V. N., G. P., and A. Edacherian, “Property enhancement approaches of fused filament fabrication technology: A review,” *Polym. Eng. Sci.*, vol. 62, no. 5, pp. 1356–1376, 2022. https://doi.org/10.1002/pen.25948

[6] M. Behzadnasab, A. A. Yousefi, D. Ebrahimibagha, and F. Nasiri, “Effects of processing conditions on mechanical properties of PLA printed parts,” *Rapid Prototyp. J.*, vol. 26, no. 2, pp. 381–389, Jan. 2019. https://doi.org/10.1108/RPJ-02-2019-0048

[7] W. Jo, O.-C. Kwon, and M.-W. Moon, “Investigation of influence of heat treatment on mechanical strength of FDM printed 3D objects,” *Rapid Prototyp. J.*, vol. 24, no. 3, pp. 637–644, Jan. 2018. https://doi.org/10.1108/RPJ-06-2017-0131

[8] K. R. Hart, R. M. Dunn, and E. D. Wetzel, “Increased fracture toughness of additively manufactured semi-crystalline thermoplastics via thermal annealing,” *Polymer*, vol. 211, p. 123091, Dec. 2020. https://doi.org/10.1016/j.polymer.2020.123091

[9] R. A. Wach, P. Wolszczak, and A. Adamus-Wlodarczyk, “Enhancement of Mechanical Properties of FDM-PLA Parts via Thermal Annealing,” *Macromol. Mater. Eng.*, vol. 303, no. 9, p. 1800169, 2018. https://doi.org/10.1002/mame.201800169

[10] V. Slavković, N. Grujović, A. Disic, and A. Radovanović, “Influence of Annealing and Printing Directions on Mechanical Properties of PLA Shape Memory Polymer Produced by Fused Deposition Modeling, 2017.

[11] W. Yu, X. Wang, X. Yin, E. Ferraris, and J. Zhang, “The effects of thermal annealing on the performance of material extrusion 3D printed polymer parts,” *Mater. Des.*, vol. 226, p. 111687, Feb. 2023. https://doi.org/10.1016/j.matdes.2023.111687

[12] J. Torres, J. Cotelo, J. Karl, and A. P. Gordon, “Mechanical Property Optimization of FDM PLA in Shear with Multiple Objectives,” *JOM*, vol. 67, no. 5, pp. 1183–1193, May 2015. https://doi.org/10.1007/s11837-015-1367-y

[13] R. Rane, A. Kulkarni, H. Prajapati, R. Taylor, A. Jain, and V. Chen, “Post-Process Effects of Isothermal Annealing and Initially Applied Static Uniaxial Loading on the Ultimate Tensile Strength of Fused Filament Fabrication Parts,” *Materials*, vol. 13, p. 352, Jan. 2020. https://doi.org/10.3390/ma13020352
[14] K. R. Hart, R. M. Dunn, J. M. Sietins, C. M. Hofmeister Mock, M. E. Mackay, and E. D. Wetzel, “Increased fracture toughness of additively manufactured amorphous thermoplastics via thermal annealing,” *Polymer*, vol. 144, pp. 192–204, May 2018. https://doi.org/10.1016/j.polymer.2018.04.024

[15] J. Butt and R. Bhaskar, “Investigating the Effects of Annealing on the Mechanical Properties of FFF-Printed Thermoplastics,” *J. Manuf. Mater. Process.*, vol. 4, no. 2, Art. no. 2, Jun. 2020. https://doi.org/10.3390/jmmp4020038

[16] International Organization for Standardization, “Geometrical Product Specifications (GPS) - Surface texture: Profile method - Rules and procedures for the assessment of surface texture.” 1997.

[17] D20 Committee, “Test Method for Tensile Properties of Plastics.” ASTM International. https://doi.org/10.1520/D0638-14

[18] International Organization for Standardization, “General tolerances - Part 1: Tolerances for linear and angular dimensions without individual tolerance indications.” Nov. 15, 1989.

[19] International Organization for Standardization, “Geometrical Product Specifications (GPS) - Indication of surface texture in technical product documentation.” Feb. 2002.
[14] K. R. Hart, R. M. Dunn, J. M. Sietins, C. M. Hofmeister Mock, M. E. Mackay, and E. D. Wetzel, “Increased fracture toughness of additively manufactured amorphous thermoplastics via thermal annealing,” *Polymer*, vol. 144, pp. 192–204, May 2018. https://doi.org/10.1016/j.polymer.2018.04.024

[15] J. Butt and R. Bhaskar, “Investigating the Effects of Annealing on the Mechanical Properties of FFF-Printed Thermoplastics,” *J. Manuf. Mater. Process.*, vol. 4, no. 2, Art. no. 2, Jun. 2020. https://doi.org/10.3390/jmmp4020038

[16] International Organization for Standardization, “Geometrical Product Specifications (GPS) - Surface texture: Profile method - Rules and procedures for the assessment of surface texture.” 1997.

[17] D20 Committee, “Test Method for Tensile Properties of Plastics.” ASTM International. https://doi.org/10.1520/D0638-14

[18] International Organization for Standardization, “General tolerances - Part 1: Tolerances for linear and angular dimensions without individual tolerance indications.” Nov. 15, 1989.

[19] International Organization for Standardization, “Geometrical Product Specifications (GPS) - Indication of surface texture in technical product documentation.” Feb. 2002.
# Improving the structural integrity of challenging to manufacture LPBF components with toolpath correction

JENSCH Felix<fnref n="1" />,ª, EISSING Katharina<fnref n="2" />,ᵇ, WILLIAMS Richard<fnref n="2" />,ᶜ, TRAUTMAN Marcus<fnref n="2" />,ᵈ, YANG Yitong<fnref n="3" />,ᵉ, DUBININ Sergej<fnref n="1" />,>f, FERGANI Omar<fnref n="2" />,⁹, HÄRTEL Sebastian <fnref n="1" />,h

**Keywords:** LPBF-Process, Machine Learning, Microstructural Investigation

**Abstract.** This work deals with the influence of optimised exposure strategies on the distortion and microstructure of components susceptible to overheating and warpage. Therefore, different distortion-prone specimen geometries of 316L were fabricated with the standard parameters, as well as with exposure strategies optimised by machine learning, which were generated using the AMAIZE software package. The manufactured samples were analysed with regard to distortion. The results of the distortion analysis were then linked with the results of the digital tomography from AMAIZE. Furthermore, components were manufactured that tend to overheat due to their geometry and orientation on the substrate plate. The influence of overheating during the LPBF process on the microstructure and porosity was investigated along the build-up direction by means of an EBSD analysis and a porosity analysis. With the presented approach for optimising the exposure strategy with AMAIZE, it could be shown that a successful production of distortion-prone components with a porosity of less than 1 % is possible in the first trial.

**Introduction**

Due to increasing requirements in terms of lightweight construction, customisation and the performance of components in aerospace, medical technology, the automotive industry and many other sectors, additive manufacturing is becoming increasingly important. This is reflected in the increase in market volume from 3 billion USD in 2013 to almost 12 billion USD in 2019 [1]. The LPBF process has established itself as an important method in metal additive manufacturing, as the low layer thickness and the small focal diameter of the laser, often less than 100 µm, allow components with a complex geometry to be produced close to the final shape. The laser melts the powder locally, which results in a large amount of heat energy being generated during the process, depending on the process parameters [2]. As the surrounding metal powder only dissipates this thermal energy, depending on the particle shape and size, to a limited extent, it is mainly dissipated through the already solidified layers of the component generating large, non-uniform temperature gradients [3, 4]. In order to produce components with large overhang angles, support structures are used in these areas, which on the one hand support heat dissipation, but on the other hand result in higher material consumption, longer production times and time-consuming removal procedures [5, 6]. This limits the advantage of environmentally friendly near-net-shape production and reduces productivity due to the required post-processing to remove the support structures.

<figure><img src="image_114.png" /></figure>

<footnote n="1">¹Chair of Hybrid Manufacturing (FHF), Brandenburg University of Technology, 03046 Cottbus, Germany</footnote>

<footnote n="2">²1000 Kelvin GmbH, 10625 Berlin, Germany</footnote>

<footnote n="3">³Chair of Physical Metallurgy and Materials Technology, Brandenburg University of Technology, 03046 Cottbus, Germany</footnote>

<footnote>ªfelix.jensch@b-tu.de; ᵇkatharina.eissing@1000kelvin.com; ᶜrichard.williams@1000kelvin.com; ᵈmarcus.trautmann@1000kelvin.com, ᶜyang@b-tu.de, ᵉdubinser@b-tu.de, ᵉomar@1000kelvin.com, ᵇhaertel@b-tu.de</footnote>
# Improving the structural integrity of challenging to manufacture LPBF components with toolpath correction

JENSCH Felix<fnref n="1" />,ª, EISSING Katharina<fnref n="2" />,ᵇ, WILLIAMS Richard<fnref n="2" />,ᶜ, TRAUTMAN Marcus<fnref n="2" />,ᵈ, YANG Yitong<fnref n="3" />,ᵉ, DUBININ Sergej<fnref n="1" />,>f, FERGANI Omar<fnref n="2" />,⁹, HÄRTEL Sebastian <fnref n="1" />,h

**Keywords:** LPBF-Process, Machine Learning, Microstructural Investigation

**Abstract.** This work deals with the influence of optimised exposure strategies on the distortion and microstructure of components susceptible to overheating and warpage. Therefore, different distortion-prone specimen geometries of 316L were fabricated with the standard parameters, as well as with exposure strategies optimised by machine learning, which were generated using the AMAIZE software package. The manufactured samples were analysed with regard to distortion. The results of the distortion analysis were then linked with the results of the digital tomography from AMAIZE. Furthermore, components were manufactured that tend to overheat due to their geometry and orientation on the substrate plate. The influence of overheating during the LPBF process on the microstructure and porosity was investigated along the build-up direction by means of an EBSD analysis and a porosity analysis. With the presented approach for optimising the exposure strategy with AMAIZE, it could be shown that a successful production of distortion-prone components with a porosity of less than 1 % is possible in the first trial.

**Introduction**

Due to increasing requirements in terms of lightweight construction, customisation and the performance of components in aerospace, medical technology, the automotive industry and many other sectors, additive manufacturing is becoming increasingly important. This is reflected in the increase in market volume from 3 billion USD in 2013 to almost 12 billion USD in 2019 [1]. The LPBF process has established itself as an important method in metal additive manufacturing, as the low layer thickness and the small focal diameter of the laser, often less than 100 µm, allow components with a complex geometry to be produced close to the final shape. The laser melts the powder locally, which results in a large amount of heat energy being generated during the process, depending on the process parameters [2]. As the surrounding metal powder only dissipates this thermal energy, depending on the particle shape and size, to a limited extent, it is mainly dissipated through the already solidified layers of the component generating large, non-uniform temperature gradients [3, 4]. In order to produce components with large overhang angles, support structures are used in these areas, which on the one hand support heat dissipation, but on the other hand result in higher material consumption, longer production times and time-consuming removal procedures [5, 6]. This limits the advantage of environmentally friendly near-net-shape production and reduces productivity due to the required post-processing to remove the support structures.

<figure><img src="image_115.png" /></figure>

<footnote n="1">¹Chair of Hybrid Manufacturing (FHF), Brandenburg University of Technology, 03046 Cottbus, Germany</footnote>

<footnote n="2">²1000 Kelvin GmbH, 10625 Berlin, Germany</footnote>

<footnote n="3">³Chair of Physical Metallurgy and Materials Technology, Brandenburg University of Technology, 03046 Cottbus, Germany</footnote>

<footnote>ªfelix.jensch@b-tu.de; ᵇkatharina.eissing@1000kelvin.com; ᶜrichard.williams@1000kelvin.com; ᵈmarcus.trautmann@1000kelvin.com, ᶜyang@b-tu.de, ᵉdubinser@b-tu.de, ᵉomar@1000kelvin.com, ᵇhaertel@b-tu.de</footnote>
Furthermore, component areas that are connected to support structures have a poorer surface quality, which has a negative impact on fatigue strength [7].

Complex geometries also generally have areas of different thicknesses. With a continuous increase in the cross-sectional area from layer to layer, the heat introduced accumulates in the narrow areas. Due to the different temperature history in different areas of the component, a locally varying microstructure and internal stresses develop [8, 9]. Depending on the heat history, martensitic, bainitic, ferritic, pearlitic or austenitic structures can develop in steels [10]. These inhomogeneities in the microstructure cause the mechanical properties of complex components to vary locally [11]. This can only be considered with great effort when designing the components. The production of components with a homogeneous microstructure would therefore increase component reliability and speed up qualification processes in areas such as aviation.

To melt the powder, the laser is guided over the powder layer in meander-shaped paths. However, due to the usually complex geometries of the components, the length of the exposure vectors varies. With constant pause times between adjacent vectors and an unchanged scanning speed, areas with short vectors are exposed faster with the same energy input. Accordingly, the temperature is locally higher in areas with short exposure vectors, which leads to uneven heat distribution in the component [12, 13, 14, 15, 16]. This uneven application of energy leads to an increase in internal stress. If the residual component stresses exceed a certain value, that is defined by the strength of the support structures or the base metal, the component is deformed during the manufacturing process [17]. If the component is severely distorted, the recoater can collide with the component, which leads to the production process being cancelled.

A more even heat distribution in the component can be achieved, for example, through feedback control by means of monitoring and an adjustment of the process parameters based on this. Kavas et. al. have presented a variant in which the laser power is automatically adjusted within a defined range during the process by taking images with a thermographic camera in order to achieve a stable interlayer temperature [18]. In 2023, EOS published the "Smart Fusion" closed-loop control system, which uses optical tomography data to adjust the laser power layer by layer using correction factors with the aim of achieving a homogeneous temperature distribution in the component [19]. The aim of this work is to demonstrate that the AI software AMAIZE from the company 1000 Kelvin can be used to influence the process parameters using machine learning algorithms in such a way that the component distortion can be reduced and the component quality improved. This allows the development of defects to be predicted and avoided even before the process starts.

## Materials and Methods

### Part Manufacturing

The distortion-prone samples and the cubes were manufactured on an EOS M290 system from EOS GmbH, which has a focus diameter of approx. 0.08 mm and in which an argon atmosphere was set up. Stainless steel alloy 316L powder with a particle size between 20 and 45 µm was used, which was produced by m4p material solutions GmbH. The powder layer was applied using a carbon fibre brush. The process parameters shown in Table 1 were used for all LPBF processes carried out.

<table><caption>Table 1: Process parameters used in the LPBF processes</caption><thead><tr><th>Processing Parameters</th><th>Laser Power</th><th>Scanning speed</th><th>Hatching distance</th><th>Layer thickness</th></tr></thead><tbody><tr><th>Value</th><td>200 W</td><td>900 mm/s</td><td>0,1 mm</td><td>0,04 mm</td></tr></tbody></table>

As a reference, the components were initially manufactured without optimising the exposure strategy with the parameters mentioned (Figure 1 a)). Subsequently, the exposure strategy for the
Furthermore, component areas that are connected to support structures have a poorer surface quality, which has a negative impact on fatigue strength [7].

Complex geometries also generally have areas of different thicknesses. With a continuous increase in the cross-sectional area from layer to layer, the heat introduced accumulates in the narrow areas. Due to the different temperature history in different areas of the component, a locally varying microstructure and internal stresses develop [8, 9]. Depending on the heat history, martensitic, bainitic, ferritic, pearlitic or austenitic structures can develop in steels [10]. These inhomogeneities in the microstructure cause the mechanical properties of complex components to vary locally [11]. This can only be considered with great effort when designing the components. The production of components with a homogeneous microstructure would therefore increase component reliability and speed up qualification processes in areas such as aviation.

To melt the powder, the laser is guided over the powder layer in meander-shaped paths. However, due to the usually complex geometries of the components, the length of the exposure vectors varies. With constant pause times between adjacent vectors and an unchanged scanning speed, areas with short vectors are exposed faster with the same energy input. Accordingly, the temperature is locally higher in areas with short exposure vectors, which leads to uneven heat distribution in the component [12, 13, 14, 15, 16]. This uneven application of energy leads to an increase in internal stress. If the residual component stresses exceed a certain value, that is defined by the strength of the support structures or the base metal, the component is deformed during the manufacturing process [17]. If the component is severely distorted, the recoater can collide with the component, which leads to the production process being cancelled.

A more even heat distribution in the component can be achieved, for example, through feedback control by means of monitoring and an adjustment of the process parameters based on this. Kavas et. al. have presented a variant in which the laser power is automatically adjusted within a defined range during the process by taking images with a thermographic camera in order to achieve a stable interlayer temperature [18]. In 2023, EOS published the "Smart Fusion" closed-loop control system, which uses optical tomography data to adjust the laser power layer by layer using correction factors with the aim of achieving a homogeneous temperature distribution in the component [19]. The aim of this work is to demonstrate that the AI software AMAIZE from the company 1000 Kelvin can be used to influence the process parameters using machine learning algorithms in such a way that the component distortion can be reduced and the component quality improved. This allows the development of defects to be predicted and avoided even before the process starts.

## Materials and Methods

### Part Manufacturing

The distortion-prone samples and the cubes were manufactured on an EOS M290 system from EOS GmbH, which has a focus diameter of approx. 0.08 mm and in which an argon atmosphere was set up. Stainless steel alloy 316L powder with a particle size between 20 and 45 µm was used, which was produced by m4p material solutions GmbH. The powder layer was applied using a carbon fibre brush. The process parameters shown in Table 1 were used for all LPBF processes carried out.

<table><caption>Table 1: Process parameters used in the LPBF processes</caption><thead><tr><th>Processing Parameters</th><th>Laser Power</th><th>Scanning speed</th><th>Hatching distance</th><th>Layer thickness</th></tr></thead><tbody><tr><th>Value</th><td>200 W</td><td>900 mm/s</td><td>0,1 mm</td><td>0,04 mm</td></tr></tbody></table>

As a reference, the components were initially manufactured without optimising the exposure strategy with the parameters mentioned (Figure 1 a)). Subsequently, the exposure strategy for the
components of the same build job was optimised with AMAIZE. For example, the pause times between adjacent vectors were adjusted through optimisation in order to achieve uniform heat distribution across the entire component. With these optimisations, the build job was then manufactured again with the process parameters from Table 1 in an argon atmosphere. (Figure 1 b)). A 25 x 25 x 25 mm³ cube was then produced on a rod with a diameter of 2.5 mm, which overheats due to heat dissipation constriction from the cube to the rod (Figure 1). To analyse the porosity and microstructure, the cube was separated by wire EDM as shown in Figure 1 d after removing the rod.

<figure><img src="image_115.png" /><figcaption>Figure 1: a) distortion-prone samples without optimised exposure strategy; b) distortion-prone specimens with optimised exposure strategy; c) LPBF-manufactured components that tend to overheat; d) cut of the cubes for the porosity analysis and the EBSD analysis</figcaption></figure>

### 3D-Scanning

The 3D scanning of the components was carried out using the ATOS 5 - 8M digitalisation platform from GOM GmbH. The resolution of the two CMOS cameras is 3357 x 2456 pixels, whereby the measurement was carried out over a measuring field of 500 x 370 mm². The measurement was performed after separating the components from the substrate plate. The scanned geometries were then superimposed with the CAD files of the manufactured components in the GOM Inspect software package from GOM GmbH and the deviations between the two geometries were measured digitally (Figure 2).
components of the same build job was optimised with AMAIZE. For example, the pause times between adjacent vectors were adjusted through optimisation in order to achieve uniform heat distribution across the entire component. With these optimisations, the build job was then manufactured again with the process parameters from Table 1 in an argon atmosphere. (Figure 1 b)). A 25 x 25 x 25 mm³ cube was then produced on a rod with a diameter of 2.5 mm, which overheats due to heat dissipation constriction from the cube to the rod (Figure 1). To analyse the porosity and microstructure, the cube was separated by wire EDM as shown in Figure 1 d after removing the rod.

<figure><img src="image_116.png" /><figcaption>Figure 1: a) distortion-prone samples without optimised exposure strategy; b) distortion-prone specimens with optimised exposure strategy; c) LPBF-manufactured components that tend to overheat; d) cut of the cubes for the porosity analysis and the EBSD analysis</figcaption></figure>

### 3D-Scanning

The 3D scanning of the components was carried out using the ATOS 5 - 8M digitalisation platform from GOM GmbH. The resolution of the two CMOS cameras is 3357 x 2456 pixels, whereby the measurement was carried out over a measuring field of 500 x 370 mm². The measurement was performed after separating the components from the substrate plate. The scanned geometries were then superimposed with the CAD files of the manufactured components in the GOM Inspect software package from GOM GmbH and the deviations between the two geometries were measured digitally (Figure 2).
<figure><img src="image_116.png" /><figcaption>Figure 2: Determination of the distortion with the ATOS 5 - 8M system</figcaption></figure>

### Microstructure analysis

For metallographic examination, the cubes were cut along the build-up direction, from the connection to the rod to the tip. The samples were then cold-embedded, ground with 240-4000 grit sandpaper and then polished with 6-3 µm diamond suspension and followed with the 0.15 µm OPS polish. The porosity was analysed using the VHX 7000 digital microscope and software from the KEYENCE DEUTSCHLAND GmbH. The EBSD analysis was performed with the field-free ultra-high-resolution scanning electron microscope TESCAN AMBER from TESCAN GmbH, which is equipped with an energy dispersive X-ray spectrometer and a high-resolution electron backscatter diffraction detector Symmetry S3 EBSD system from Oxford instruments plc. The EBSD data with was recorded with a step size of 2.02 µm and analysed with the AZtecCrystal software from Oxford Instruments plc.

### Results and discussion

#### Distortion of test geometries manufactured with and without AMAIZE

As the optimisation of the process parameters and the exposure strategy with AMAIZE is to be tested for different geometric elements, such as material accumulations or thin walls, components that tend to warp during the LPBF process were manufactured, 3D-scanned and compared to the CAD-file. To show the effect of the adjusted parameters, the components were first manufactured with the standard parameters and then with the optimised parameters. In the angled component, heat builds up at the transition from the component to the support. This is caused by the reduced thermal conductivity of the support structure compared to solid material [20]. Furthermore, the scan vectors are shorter due to the contour boundaries at edges and corners, which, as described by Liu et al., promotes heat accumulation due to the shortened time intervals when exposing adjacent vectors [12]. The resulting overheating causes distortion here. The aim here was to minimize the local heat input at the edges and corners of the component in order to prevent the weakly designed support from tearing off the component. Parameter optimisation with AMAIZE is used to control the pause times of short, adjacent scan vectors in such a way that the heat introduced is dissipated as evenly as possible. Figure 1a shows that the support is torn off over the entire length of the overhanging component area without optimizing the parameters. When using the optimized process parameters, the support structure essentially tears in the structure itself, which can be avoided by designing the structure more densely. Figure 3 also shows that the warpage does not extend as far into the component with the optimised parameters.
<figure><img src="image_117.png" /><figcaption>Figure 2: Determination of the distortion with the ATOS 5 - 8M system</figcaption></figure>

### Microstructure analysis

For metallographic examination, the cubes were cut along the build-up direction, from the connection to the rod to the tip. The samples were then cold-embedded, ground with 240-4000 grit sandpaper and then polished with 6-3 µm diamond suspension and followed with the 0.15 µm OPS polish. The porosity was analysed using the VHX 7000 digital microscope and software from the KEYENCE DEUTSCHLAND GmbH. The EBSD analysis was performed with the field-free ultra-high-resolution scanning electron microscope TESCAN AMBER from TESCAN GmbH, which is equipped with an energy dispersive X-ray spectrometer and a high-resolution electron backscatter diffraction detector Symmetry S3 EBSD system from Oxford instruments plc. The EBSD data with was recorded with a step size of 2.02 µm and analysed with the AZtecCrystal software from Oxford Instruments plc.

### Results and discussion

#### Distortion of test geometries manufactured with and without AMAIZE

As the optimisation of the process parameters and the exposure strategy with AMAIZE is to be tested for different geometric elements, such as material accumulations or thin walls, components that tend to warp during the LPBF process were manufactured, 3D-scanned and compared to the CAD-file. To show the effect of the adjusted parameters, the components were first manufactured with the standard parameters and then with the optimised parameters. In the angled component, heat builds up at the transition from the component to the support. This is caused by the reduced thermal conductivity of the support structure compared to solid material [20]. Furthermore, the scan vectors are shorter due to the contour boundaries at edges and corners, which, as described by Liu et al., promotes heat accumulation due to the shortened time intervals when exposing adjacent vectors [12]. The resulting overheating causes distortion here. The aim here was to minimize the local heat input at the edges and corners of the component in order to prevent the weakly designed support from tearing off the component. Parameter optimisation with AMAIZE is used to control the pause times of short, adjacent scan vectors in such a way that the heat introduced is dissipated as evenly as possible. Figure 1a shows that the support is torn off over the entire length of the overhanging component area without optimizing the parameters. When using the optimized process parameters, the support structure essentially tears in the structure itself, which can be avoided by designing the structure more densely. Figure 3 also shows that the warpage does not extend as far into the component with the optimised parameters.
<figure><img src="image_117.png" /><figcaption>Figure 3: Comparison of the distortion between the manufactured winkle and the CAD file without (left) and with (right) optimisation of the manufacturing parameters</figcaption></figure>

During the production of the thin walls with the accumulation of material on the top, the components warped very strongly when the non-optimised process parameters were used. This damaged the carbon fibre brush used to apply the powder, which also resulted in major geometric deviations in other components of the build job. As a result, one wall of the structure suffered bonding failures and ultimately the complete loss of the connection between the component and the subsequently melted powder material, which also damaged the two walls behind it. This effect was avoided after optimising the exposure strategy with AMAIZE (Figure 4). The three walls on the left-hand side of the component also exhibit less warpage after optimisation than the walls produced with the standard parameters. Due to the small cross-sectional area of the walls, the thermal energy introduced cannot be completely dissipated. This leads to a heat build-up in the area of the material accumulation at the tip, which is the cause of the warpage. Contrary to expectations, the greatest distortion occurred in the wall with the smallest accumulation of material at the top. One reason for this is that the exposure vectors are shorter for the smaller accumulations of material. As a result, the exposure of neighbouring vectors is faster, which results in higher local heat generation in the short term than with longer exposure vectors, what subsequently causes warping.

<figure><img src="image_118.png" /><figcaption>Figure 4: Comparison of the distortion between the manufactured thin walls and the CAD file without (left) and with (right) optimisation of the manufacturing parameters</figcaption></figure>

## Linking warping with the AMAIZE digital tomography

For a quick assessment of the thermal energy input, a two- and three-dimensional digital tomography of the components to be manufactured in the LPBF process is created in AMAIZE. Figure 5 shows the results of the digital tomography from various layers located in the warped area shortly after the end of the support structure for the angle. It can be seen that an increased energy input is predicted in layer 1015 over a large area of the exposed surface. This result reflects the
<figure><img src="image_118.png" /><figcaption>Figure 3: Comparison of the distortion between the manufactured winkle and the CAD file without (left) and with (right) optimisation of the manufacturing parameters</figcaption></figure>

During the production of the thin walls with the accumulation of material on the top, the components warped very strongly when the non-optimised process parameters were used. This damaged the carbon fibre brush used to apply the powder, which also resulted in major geometric deviations in other components of the build job. As a result, one wall of the structure suffered bonding failures and ultimately the complete loss of the connection between the component and the subsequently melted powder material, which also damaged the two walls behind it. This effect was avoided after optimising the exposure strategy with AMAIZE (Figure 4). The three walls on the left-hand side of the component also exhibit less warpage after optimisation than the walls produced with the standard parameters. Due to the small cross-sectional area of the walls, the thermal energy introduced cannot be completely dissipated. This leads to a heat build-up in the area of the material accumulation at the tip, which is the cause of the warpage. Contrary to expectations, the greatest distortion occurred in the wall with the smallest accumulation of material at the top. One reason for this is that the exposure vectors are shorter for the smaller accumulations of material. As a result, the exposure of neighbouring vectors is faster, which results in higher local heat generation in the short term than with longer exposure vectors, what subsequently causes warping.

<figure><img src="image_119.png" /><figcaption>Figure 4: Comparison of the distortion between the manufactured thin walls and the CAD file without (left) and with (right) optimisation of the manufacturing parameters</figcaption></figure>

Linking warping with the AMAIZE digital tomography

For a quick assessment of the thermal energy input, a two- and three-dimensional digital tomography of the components to be manufactured in the LPBF process is created in AMAIZE. Figure 5 shows the results of the digital tomography from various layers located in the warped area shortly after the end of the support structure for the angle. It can be seen that an increased energy input is predicted in layer 1015 over a large area of the exposed surface. This result reflects the
poorer heat dissipation through the support structure and reflects the distortion in this area. After optimising the exposure parameters, the predicted heat development in this area is significantly lower. An increased temperature is detected at regular intervals in the other layers. This is due to the strip-like exposure strategy used to produce the components and the meandering exposure of the individual exposure vectors within these strips. After the exposure of a vector at the end of the strip has been completed, the adjacent vector starts, usually without waiting time, at a definable distance, which is often only between 50 and 200 µm. This results in a higher energy input for a short time in this area than in the centre of the strips. The machine learning algorithm in AMAIZE influences the pause times between adjacent vectors based on the thermophysical properties, which are based on the geometry and the other process parameters. This significantly reduces overheating at the ends of the strips.

<figure><img src="image_119.png" /><figcaption>Figure 5: Digital tomography on different layers of the winkle component</figcaption></figure>

As with the angle element, the heat development in the thin walls is greater with the standard parameters than with the parameters optimised by AMAIZE. Figure 6 shows three of the layers where the accumulation of material on the wall begins. It is notable that the walls with the smaller accumulations of material are predicted to have particularly high heat generation with the standard parameters. This reflects the distortion in these areas from the LPBF-test. One possible cause, as already suspected, is the short length of the scan vectors in this area, with the cross-section increasing from layer to layer. The resulting high heat input cannot be dissipated sufficiently through the wall, which is why distortion occurs. After optimising the parameters using AMAIZE, a reduction in heat development is achieved for these walls in particular, which is again due to the increased pause times between adjacent exposure vectors.
poorer heat dissipation through the support structure and reflects the distortion in this area. After optimising the exposure parameters, the predicted heat development in this area is significantly lower. An increased temperature is detected at regular intervals in the other layers. This is due to the strip-like exposure strategy used to produce the components and the meandering exposure of the individual exposure vectors within these strips. After the exposure of a vector at the end of the strip has been completed, the adjacent vector starts, usually without waiting time, at a definable distance, which is often only between 50 and 200 µm. This results in a higher energy input for a short time in this area than in the centre of the strips. The machine learning algorithm in AMAIZE influences the pause times between adjacent vectors based on the thermophysical properties, which are based on the geometry and the other process parameters. This significantly reduces overheating at the ends of the strips.

<figure><img src="image_120.png" /><figcaption>Figure 5: Digital tomography on different layers of the winkle component</figcaption></figure>

As with the angle element, the heat development in the thin walls is greater with the standard parameters than with the parameters optimised by AMAIZE. Figure 6 shows three of the layers where the accumulation of material on the wall begins. It is notable that the walls with the smaller accumulations of material are predicted to have particularly high heat generation with the standard parameters. This reflects the distortion in these areas from the LPBF-test. One possible cause, as already suspected, is the short length of the scan vectors in this area, with the cross-section increasing from layer to layer. The resulting high heat input cannot be dissipated sufficiently through the wall, which is why distortion occurs. After optimising the parameters using AMAIZE, a reduction in heat development is achieved for these walls in particular, which is again due to the increased pause times between adjacent exposure vectors.
<figure><img src="image_120.png" /><figcaption>Figure 6: Digital tomography on different layers of the thin wall component</figcaption></figure>

Influence of overheating on the porosity and the microstructure

The cubes on a rod have been successfully produced using AMAIZE optimized exposure strategy. We differentiate between the top and bottom sections of the cube due to observable variations. Specifically, the bottom part exhibits signs of overheating, such as external oxidation, indicating a distinct thermal signature compared to the top section.

To evaluate the cubes produced in Figure 1 c/d, the cubes were examined for their porosity and microstructure in the bottom region, which shows tarnishing on the surface, and the upper top area. The porosity is very low at 0.04 % in the bottom area and 0.02 % in the top area over the entire sample (Figure 7). It is assumed that due to the uniform overheating over the entire surface, the thermal gradient between the exposed area and the surrounding areas was lower, so that the melt pool flows were lower than for strongly cooled surfaces, in accordance with the Marangoni effect [21]. Figure 7 b shows an increased number of pores between 5 and 15 µm in the overheated and non-overheated areas. It is notable that the overheated area sometimes has very large pores of over 100 µm. One reason for this may be a greater melt pool depth due to the already high surface temperature, which has led to the formation of larger pores corresponding to the keyhole effect [22].

Using an EBSD analysis, an average grain size of approx. 91 µm was determined in the overheated area. In the non-overheated upper area of the cube, the average grain size of approx. 48 µm was only about half of this. As the overheating in the lower area of the cube is effective for almost the entire duration of the printing process, effects similar to those of heat treatment already occur here. The increased temperature results in greater grain growth, which leads to lower tensile strength and higher elongation at break [23]. Figure 7 c and d also show that the grains are stretched along the direction in which they are built up. This results from the temperature gradient, which is caused by the dissipation of the heat introduced in the negative build-up direction. As the heat accumulated in the overheated area, the grains could also grow at right angles to the build-up direction, although this effect was less pronounced than expected.
<figure><img src="image_121.png" /><figcaption>Figure 6: Digital tomography on different layers of the thin wall component</figcaption></figure>

Influence of overheating on the porosity and the microstructure

The cubes on a rod have been successfully produced using AMAIZE optimized exposure strategy. We differentiate between the top and bottom sections of the cube due to observable variations. Specifically, the bottom part exhibits signs of overheating, such as external oxidation, indicating a distinct thermal signature compared to the top section.

To evaluate the cubes produced in Figure 1 c/d, the cubes were examined for their porosity and microstructure in the bottom region, which shows tarnishing on the surface, and the upper top area. The porosity is very low at 0.04 % in the bottom area and 0.02 % in the top area over the entire sample (Figure 7). It is assumed that due to the uniform overheating over the entire surface, the thermal gradient between the exposed area and the surrounding areas was lower, so that the melt pool flows were lower than for strongly cooled surfaces, in accordance with the Marangoni effect [21]. Figure 7 b shows an increased number of pores between 5 and 15 µm in the overheated and non-overheated areas. It is notable that the overheated area sometimes has very large pores of over 100 µm. One reason for this may be a greater melt pool depth due to the already high surface temperature, which has led to the formation of larger pores corresponding to the keyhole effect [22].

Using an EBSD analysis, an average grain size of approx. 91 µm was determined in the overheated area. In the non-overheated upper area of the cube, the average grain size of approx. 48 µm was only about half of this. As the overheating in the lower area of the cube is effective for almost the entire duration of the printing process, effects similar to those of heat treatment already occur here. The increased temperature results in greater grain growth, which leads to lower tensile strength and higher elongation at break [23]. Figure 7 c and d also show that the grains are stretched along the direction in which they are built up. This results from the temperature gradient, which is caused by the dissipation of the heat introduced in the negative build-up direction. As the heat accumulated in the overheated area, the grains could also grow at right angles to the build-up direction, although this effect was less pronounced than expected.
<figure><img src="image_121.png" /><figcaption>Figure 7: overview of the a) porosity and b) the pore distribution as well as an EBSD overview of c) an overheated area and d) a not overheated area</figcaption></figure>

## Conclusions and outlook

In this work, various components susceptible to distortion were manufactured from 316L with standard parameters and with parameters optimised by a machine learning algorithm and investigated with regard to distortion. Furthermore, geometries that are prone to severe overheating were analysed in terms of porosity and microstructure. Based on the results, the following conclusions can be drawn:

* Components that are difficult to manufacture in the LPBF process were successfully produced in the first attempt after optimizing the exposure strategy with AMAIZE

* By utilizing advanced machine learning algorithms to adjust the exposure strategy and process parameters, we can achieve a more uniform distribution of the laser's energy throughout the component. This enhanced control helps in evenly heating the material, thereby significantly reducing the likelihood of distortion during the manufacturing process.

* Regardless of the overheating of the component, the porosity remained roughly identical in all areas of the component, whereby the size of the pores is larger in overheated areas

* As the high temperatures in overheated areas are maintained over the duration of the process, this results in increased grain growth, which leads to an inhomogeneous microstructure and therefore varying mechanical properties in the component

By optimising the process parameters with machine learning algorithms, a more homogeneous microstructure and therefore better mechanical properties can be achieved due to a more even heat distribution in the component. This demonstrates a pathway to reduce build failures in metal AM, improving the economic viability and reducing the environmental footprint of the process. Further, such toolpath corrections can enable process engineers to use fewer support structures, increasing
<figure><img src="image_122.png" /><figcaption>Figure 7: overview of the a) porosity and b) the pore distribution as well as an EBSD overview of c) an overheated area and d) a not overheated area</figcaption></figure>

## Conclusions and outlook

In this work, various components susceptible to distortion were manufactured from 316L with standard parameters and with parameters optimised by a machine learning algorithm and investigated with regard to distortion. Furthermore, geometries that are prone to severe overheating were analysed in terms of porosity and microstructure. Based on the results, the following conclusions can be drawn:

* Components that are difficult to manufacture in the LPBF process were successfully produced in the first attempt after optimizing the exposure strategy with AMAIZE

* By utilizing advanced machine learning algorithms to adjust the exposure strategy and process parameters, we can achieve a more uniform distribution of the laser's energy throughout the component. This enhanced control helps in evenly heating the material, thereby significantly reducing the likelihood of distortion during the manufacturing process.

* Regardless of the overheating of the component, the porosity remained roughly identical in all areas of the component, whereby the size of the pores is larger in overheated areas

* As the high temperatures in overheated areas are maintained over the duration of the process, this results in increased grain growth, which leads to an inhomogeneous microstructure and therefore varying mechanical properties in the component

By optimising the process parameters with machine learning algorithms, a more homogeneous microstructure and therefore better mechanical properties can be achieved due to a more even heat distribution in the component. This demonstrates a pathway to reduce build failures in metal AM, improving the economic viability and reducing the environmental footprint of the process. Further, such toolpath corrections can enable process engineers to use fewer support structures, increasing
freedom to designers and reducing the cost per part. The machine learning algorithm is to be trained further as part of further investigations. For this purpose, monitoring data will also be recorded using a thermographic camera, melt pool monitoring or optical tomography, which will map the temperature history of the component and thus check and improve the accuracy of the algorithm. The aim is to produce components with a homogeneous microstructure and therefore homogeneous mechanical properties.

Reference

[1] C. Tan, F. Weng, S. Sui, Y. Chew und G. Bi, „Progress and perspectives in laser additive manufacturing of key aeroengine materials,“ *International Journal of Machine Tools and Manufacture*, 2021. https://doi.org/10.1016/j.ijmachtools.2021.103804

[2] W. Di, Y. Yongqiang, S. Xubin und C. Yonghua, „Study on energy input and its influences on single-track, multi-track, and multi-layer in SLM,“ *International Journal of Advanced Manufacturing Technology*, 2012. https://doi.org/10.1007/s00170-011-3443-y

[3] J. Gan, Z. Zhou und A. Yu, „Effect of particle shape and size on effective thermal conductivity of packed beds,“ *Powder Technology*, 2017. https://doi.org/10.1016/j.powtec.2017.01.024

[4] L. C. Wei, L. E. Ehrlich, M. J. Powell-Palm, C. Montgomery, J. Beuth und J. A. Malen, „Thermal conductivity of metal powders for powder bed additive manufacturing,“ *Additive Manufacturing*, 2018. https://doi.org/10.1016/j.addma.2018.02.002

[5] T. Miki und S. Nishiwaki, „Topology optimization of the support structure for heat dissipation in additive manufacturing,“ *Finite Elements in Analysis & Design*, 2022. https://doi.org/10.1016/j.finel.2021.103708

[6] S. C. Subedi, A. Shahba, M. Thevamaran, D. J. Thoma und K. Suresh, „Towards the optimal design of support structures for laser powder bed fusion-based metal additive manufacturing via thermal equivalent static loads,“ *Additive Manufacturing*, 2022. https://doi.org/10.1016/j.addma.2022.102956

[7] J. Elambasseril, J. Rogers, C. Wallbrink, D. Munk, M. Leary und M. Qian, „Laser powder bed fusion additive manufacturing (LPBF-AM): the influence of design features and LPBF variables on surface topography and effect on fatigue properties,“ *Critical Reviews in Solid State and Materials Sciences*, 2023. https://doi.org/10.1080/10408436.2022.2041396

[8] G. Mohr, N. Scheuschner und K. Hilgenberg, „In situ heat accumulation by geometrical features obstructing heat flux and by reduced inter layer times in laser powder bed fusion of AISI 316L stainless steel,“ *Procedia CIRP*, 2020. https://doi.org/10.1016/j.procir.2020.09.030

[9] J. Munk, E. Breitbarth, T. Siemer, N. Pirch und C. Häfner, „Geometry Effect on Microstructure and Mechanical Properties in Laser Powder Bed Fusion of Ti-6Al-4V,“ *metals*, 2022. https://doi.org/10.3390/met12030482

[10] S. Astafurov und E. Astafurova, „Phase Composition of Austenitic Stainless Steels in Additive Manufacturing: A Review,“ *Metals*, 2021. https://doi.org/10.3390/met11071052

[11] T. A. Rodrigues, V. Duarte, J. A. Avila, T. G. Santos, R. Miranda und J. Oliveira, „Wire and arc additive manufacturing of HSLA steel: Effect of thermal cycles on microstructure and mechanical properties,“ *Additive Manufacturing*, 2019. https://doi.org/10.1016/j.addma.2019.03.029

[12] Y. Liu, J. Li, K. Xu, T. Cheng, D. Zhao, W. Li, Q. Teng und Q. Wei, „An optimized scanning strategy to mitigate excessive heat accumulation caused by short scanning lines in laser
freedom to designers and reducing the cost per part. The machine learning algorithm is to be trained further as part of further investigations. For this purpose, monitoring data will also be recorded using a thermographic camera, melt pool monitoring or optical tomography, which will map the temperature history of the component and thus check and improve the accuracy of the algorithm. The aim is to produce components with a homogeneous microstructure and therefore homogeneous mechanical properties.

Reference

[1] C. Tan, F. Weng, S. Sui, Y. Chew und G. Bi, „Progress and perspectives in laser additive manufacturing of key aeroengine materials,“ *International Journal of Machine Tools and Manufacture*, 2021. https://doi.org/10.1016/j.ijmachtools.2021.103804

[2] W. Di, Y. Yongqiang, S. Xubin und C. Yonghua, „Study on energy input and its influences on single-track, multi-track, and multi-layer in SLM,“ *International Journal of Advanced Manufacturing Technology*, 2012. https://doi.org/10.1007/s00170-011-3443-y

[3] J. Gan, Z. Zhou und A. Yu, „Effect of particle shape and size on effective thermal conductivity of packed beds,“ *Powder Technology*, 2017. https://doi.org/10.1016/j.powtec.2017.01.024

[4] L. C. Wei, L. E. Ehrlich, M. J. Powell-Palm, C. Montgomery, J. Beuth und J. A. Malen, „Thermal conductivity of metal powders for powder bed additive manufacturing,“ *Additive Manufacturing*, 2018. https://doi.org/10.1016/j.addma.2018.02.002

[5] T. Miki und S. Nishiwaki, „Topology optimization of the support structure for heat dissipation in additive manufacturing,“ *Finite Elements in Analysis & Design*, 2022. https://doi.org/10.1016/j.finel.2021.103708

[6] S. C. Subedi, A. Shahba, M. Thevamaran, D. J. Thoma und K. Suresh, „Towards the optimal design of support structures for laser powder bed fusion-based metal additive manufacturing via thermal equivalent static loads,“ *Additive Manufacturing*, 2022. https://doi.org/10.1016/j.addma.2022.102956

[7] J. Elambasseril, J. Rogers, C. Wallbrink, D. Munk, M. Leary und M. Qian, „Laser powder bed fusion additive manufacturing (LPBF-AM): the influence of design features and LPBF variables on surface topography and effect on fatigue properties,“ *Critical Reviews in Solid State and Materials Sciences*, 2023. https://doi.org/10.1080/10408436.2022.2041396

[8] G. Mohr, N. Scheuschner und K. Hilgenberg, „In situ heat accumulation by geometrical features obstructing heat flux and by reduced inter layer times in laser powder bed fusion of AISI 316L stainless steel,“ *Procedia CIRP*, 2020. https://doi.org/10.1016/j.procir.2020.09.030

[9] J. Munk, E. Breitbarth, T. Siemer, N. Pirch und C. Häfner, „Geometry Effect on Microstructure and Mechanical Properties in Laser Powder Bed Fusion of Ti-6Al-4V,“ *metals*, 2022. https://doi.org/10.3390/met12030482

[10] S. Astafurov und E. Astafurova, „Phase Composition of Austenitic Stainless Steels in Additive Manufacturing: A Review,“ *Metals*, 2021. https://doi.org/10.3390/met11071052

[11] T. A. Rodrigues, V. Duarte, J. A. Avila, T. G. Santos, R. Miranda und J. Oliveira, „Wire and arc additive manufacturing of HSLA steel: Effect of thermal cycles on microstructure and mechanical properties,“ *Additive Manufacturing*, 2019. https://doi.org/10.1016/j.addma.2019.03.029

[12] Y. Liu, J. Li, K. Xu, T. Cheng, D. Zhao, W. Li, Q. Teng und Q. Wei, „An optimized scanning strategy to mitigate excessive heat accumulation caused by short scanning lines in laser
powder bed fusion process," *Additive Manufacturing*, 2022.
https://doi.org/10.1016/j.addma.2022.103256

[13] L. Wang, X. Jiang, Y. Zhu, X. Zhu, J. Sun und B. Yan, „An approach to predict the residual stress and distortion during the selective laser melting of AlSi10Mg parts,“ *The International Journal of Advanced Manufacturing Technology*, 2018. https://doi.org/10.1007/s00170-018-2207-3

[14] D. Dai, D. Gu, Q. Ge, Y. Li, X. Shi, Y. Sun und S. Li, „Mesoscopic study of thermal behavior, fluid dynamics and surface morphology during selective laser melting of Ti-based composites,“ *Computational Materials Science*, 2020.
https://doi.org/10.1016/j.commatsci.2020.109598

[15] Z. Li, R. Xu, Z. Zhang und I. Kucukkoc, „The influence of scan length on fabricating thin-walled components in selective laser melting,“ *International Journal of Machine Tools and Manufacture*, 2018. https://doi.org/10.1016/j.ijmachtools.2017.11.012

[16] Y.-L. Lo, B.-Y. Liu und H.-C. Tran, „Optimized hatch space selection in double-scanning track selective laser melting process,“ *The International Journal of Advanced Manufacturing Technology*, 2019.

[17] N. Levkulich, S. Semiatin, J. Gockel, J. Middendorf, A. DeWald und N. Klingbeil, „The effect of process parameters on residual stress evolution and distortion in the laser powder bed fusion of Ti-6Al-4V,“ *Additive Manufacturing*, 2019.
https://doi.org/10.1016/j.addma.2019.05.015

[18] B. Kavas, E. C. Balta, M. Tucker, A. Rupenyan, J. Lygeros und M. Bambach, „Layer-to-layer closed-loop feedback control application for inter-layer temperature stabilization in laser powder bed fusion,“ *Additive Manufacturing*, 2023.
https://doi.org/10.2139/ssrn.4496744

[19] A. Yağmur, I. Pääkkönen und A. Miles, „The Hitchhiker's Guide to Smart Fusion,“ July 2023. [Online]. Available: https://www.eos.info/smart-fusion.

[20] L. White, X. Liang, G. Zhang, J. Cagan und Y. J. Zhang, „Coupling Simulated Annealing and Homogenization to Design Thermally Conductive Hybrid Lattice Support Structures for LPBF,“ *Proceedings of the ASME 2023*, 2023. https://doi.org/10.1115/DETC2023-114953

[21] M. Rombouts, J. Kruth, L. Froyen und P. Mercelis, „Fundamentals of Selective Laser Melting of alloyed steel powders,“ *CIRP Annals*, 2006. https://doi.org/10.1016/S0007-8506(07)60395-3

[22] G. Kasperovich, J. Haubrich, J. Gussone und G. Requena, „Correlation between porosity and processing parameters in TiAl6V4 produced by selective laser melting,“ *Materials and Design*, 2016. https://doi.org/10.1016/j.matdes.2016.05.070

[23] A. Yadollahi, N. Shansaei, S. M. Thompson und D. W. Seely, „Effects of process time interval and heat treatment on the mechanical and microstructural properties of direct laser deposited 316L stainless steel,“ *Materials Science & Engineering A*, 2015.
https://doi.org/10.1016/j.msea.2015.07.056
powder bed fusion process," *Additive Manufacturing*, 2022.
https://doi.org/10.1016/j.addma.2022.103256

[13] L. Wang, X. Jiang, Y. Zhu, X. Zhu, J. Sun und B. Yan, „An approach to predict the residual stress and distortion during the selective laser melting of AlSi10Mg parts,“ *The International Journal of Advanced Manufacturing Technology*, 2018. https://doi.org/10.1007/s00170-018-2207-3

[14] D. Dai, D. Gu, Q. Ge, Y. Li, X. Shi, Y. Sun und S. Li, „Mesoscopic study of thermal behavior, fluid dynamics and surface morphology during selective laser melting of Ti-based composites,“ *Computational Materials Science*, 2020.
https://doi.org/10.1016/j.commatsci.2020.109598

[15] Z. Li, R. Xu, Z. Zhang und I. Kucukkoc, „The influence of scan length on fabricating thin-walled components in selective laser melting,“ *International Journal of Machine Tools and Manufacture*, 2018. https://doi.org/10.1016/j.ijmachtools.2017.11.012

[16] Y.-L. Lo, B.-Y. Liu und H.-C. Tran, „Optimized hatch space selection in double-scanning track selective laser melting process,“ *The International Journal of Advanced Manufacturing Technology*, 2019.

[17] N. Levkulich, S. Semiatin, J. Gockel, J. Middendorf, A. DeWald und N. Klingbeil, „The effect of process parameters on residual stress evolution and distortion in the laser powder bed fusion of Ti-6Al-4V,“ *Additive Manufacturing*, 2019.
https://doi.org/10.1016/j.addma.2019.05.015

[18] B. Kavas, E. C. Balta, M. Tucker, A. Rupenyan, J. Lygeros und M. Bambach, „Layer-to-layer closed-loop feedback control application for inter-layer temperature stabilization in laser powder bed fusion,“ *Additive Manufacturing*, 2023.
https://doi.org/10.2139/ssrn.4496744

[19] A. Yağmur, I. Pääkkönen und A. Miles, „The Hitchhiker's Guide to Smart Fusion,“ July 2023. [Online]. Available: https://www.eos.info/smart-fusion.

[20] L. White, X. Liang, G. Zhang, J. Cagan und Y. J. Zhang, „Coupling Simulated Annealing and Homogenization to Design Thermally Conductive Hybrid Lattice Support Structures for LPBF,“ *Proceedings of the ASME 2023*, 2023. https://doi.org/10.1115/DETC2023-114953

[21] M. Rombouts, J. Kruth, L. Froyen und P. Mercelis, „Fundamentals of Selective Laser Melting of alloyed steel powders,“ *CIRP Annals*, 2006. https://doi.org/10.1016/S0007-8506(07)60395-3

[22] G. Kasperovich, J. Haubrich, J. Gussone und G. Requena, „Correlation between porosity and processing parameters in TiAl6V4 produced by selective laser melting,“ *Materials and Design*, 2016. https://doi.org/10.1016/j.matdes.2016.05.070

[23] A. Yadollahi, N. Shansaei, S. M. Thompson und D. W. Seely, „Effects of process time interval and heat treatment on the mechanical and microstructural properties of direct laser deposited 316L stainless steel,“ *Materials Science & Engineering A*, 2015.
https://doi.org/10.1016/j.msea.2015.07.056
# Influence of remelting lasing strategies on the fracture toughness of Hastelloy X manufactured by laser powder bed fusion

KELLER Clément<sup>1,a*</sup>, VIEILLE Benoit<sup>2,b</sup> and DUCHAUSSOY Amandine<sup>2,b</sup>

<sup>1</sup>Laboratoire Génie de Production, Université de Technologie de Tarbes Occitanie Pyrénées, Université de Toulouse, Tarbes, France

<sup>2</sup>Groupe de Physique des Matériaux, Normandie Université, UMRCNRS 6634, Saint-Etienne du Rouvray, France

<sup>a</sup>clément.keller@enit.fr, <sup>b</sup>benoit.vieille@insa-rouen.fr, <sup>c</sup>amandine.duchaussoy@insa-rouen.fr

**Keywords:** Remelting, Fracture Toughness, Microstructure, Laser Powder Bed Fusion, Hastelloy X

**Abstract.** Additive manufacturing and, in particular, Laser Powder Bed Fusion Processes (LPBF), are known to generate non-equilibrium microstructures having a strong impact on the mechanical properties such as tensile of fatigue ones. To better control this impact, optimized lasing strategies can be employed. Among them, relasing or remelting ones have been proved to reduce the porosity and the surface roughness as well as to increase the ductility [1,2]. This particular manufacturing strategy seems, hence, to be a promising tool to modify the microstructure and mechanical properties of LPBF alloys and to optimize the mechanical properties. The objective of this study is to assess the influence of such relasing and its characteristics on the fracture toughness of Hastelloy X superalloys.

**Introduction**

Laser powder bed-based processes are nowadays well employed in industry to generate parts with high degree of geometrical complexity. Those processes are then a powerful tool to generate lightweight structures and to limit the use of primary material in relation to the necessary reduction of greenhouse gaz emission. Nevertheless, due to its intrinsic characteristics, those processes generate thermal gradients and cooling rates as far as millions of K/m and K/s, respectively [1]. Consequently, the produced alloys exhibit complex microstructures from morphological and crystallographical texture to dendrite structures and nano-oxides [2]. Other defects such as lack-of-fusion, keyholes, residual stresses or high surface roughness are also observed for AM materials [3].

The mechanical properties of those LPBF materials are generally modified compared to their conventionally produced counterparts (forging, casting, rolling) [3]. In tension, if the process parameters are correctly set up, mechanical properties of many alloy families are generally higher than for others manufacturing processes [4]. In fatigue, even after surface roughness removal and residual stresses relieve, the endurance limit is still lower for AM compared to cast [5].

In order to optimize the microstructure and reduce material defects, remelting strategies have been proposed in literature. Using such remelting, a many folds microstructure modification has been observed. Such modification affects the grain size distribution [6], the surface roughness [7], the porosity [7] as well as the dendrite size [8]. Consequently, an increase in mechanical properties in tension affecting the yield stress [8, 9], the ultimate tensile strength [8, 9] and the ductility has been reported [8]. Those results reveal the great potential for such on-line treatments which can be applied in the entire part or only locally where improved microstructure and mechanical properties are needed.

<figure><img src="image_122.png" /></figure>
# Influence of remelting lasing strategies on the fracture toughness of Hastelloy X manufactured by laser powder bed fusion

KELLER Clément<sup>1,a*</sup>, VIEILLE Benoit<sup>2,b</sup> and DUCHAUSSOY Amandine<sup>2,b</sup>

<sup>1</sup>Laboratoire Génie de Production, Université de Technologie de Tarbes Occitanie Pyrénées, Université de Toulouse, Tarbes, France

<sup>2</sup>Groupe de Physique des Matériaux, Normandie Université, UMRCNRS 6634, Saint-Etienne du Rouvray, France

<sup>a</sup>clément.keller@enit.fr, <sup>b</sup>benoit.vieille@insa-rouen.fr, <sup>c</sup>amandine.duchaussoy@insa-rouen.fr

**Keywords:** Remelting, Fracture Toughness, Microstructure, Laser Powder Bed Fusion, Hastelloy X

**Abstract.** Additive manufacturing and, in particular, Laser Powder Bed Fusion Processes (LPBF), are known to generate non-equilibrium microstructures having a strong impact on the mechanical properties such as tensile of fatigue ones. To better control this impact, optimized lasing strategies can be employed. Among them, relasing or remelting ones have been proved to reduce the porosity and the surface roughness as well as to increase the ductility [1,2]. This particular manufacturing strategy seems, hence, to be a promising tool to modify the microstructure and mechanical properties of LPBF alloys and to optimize the mechanical properties. The objective of this study is to assess the influence of such relasing and its characteristics on the fracture toughness of Hastelloy X superalloys.

**Introduction**

Laser powder bed-based processes are nowadays well employed in industry to generate parts with high degree of geometrical complexity. Those processes are then a powerful tool to generate lightweight structures and to limit the use of primary material in relation to the necessary reduction of greenhouse gaz emission. Nevertheless, due to its intrinsic characteristics, those processes generate thermal gradients and cooling rates as far as millions of K/m and K/s, respectively [1]. Consequently, the produced alloys exhibit complex microstructures from morphological and crystallographical texture to dendrite structures and nano-oxides [2]. Other defects such as lack-of-fusion, keyholes, residual stresses or high surface roughness are also observed for AM materials [3].

The mechanical properties of those LPBF materials are generally modified compared to their conventionally produced counterparts (forging, casting, rolling) [3]. In tension, if the process parameters are correctly set up, mechanical properties of many alloy families are generally higher than for others manufacturing processes [4]. In fatigue, even after surface roughness removal and residual stresses relieve, the endurance limit is still lower for AM compared to cast [5].

In order to optimize the microstructure and reduce material defects, remelting strategies have been proposed in literature. Using such remelting, a many folds microstructure modification has been observed. Such modification affects the grain size distribution [6], the surface roughness [7], the porosity [7] as well as the dendrite size [8]. Consequently, an increase in mechanical properties in tension affecting the yield stress [8, 9], the ultimate tensile strength [8, 9] and the ductility has been reported [8]. Those results reveal the great potential for such on-line treatments which can be applied in the entire part or only locally where improved microstructure and mechanical properties are needed.

<figure><img src="image_123.png" /></figure>
Nevertheless, those remelting strategies have not been extensively characterized for fatigue or fracture toughness improvements which are key properties for an industrial use of many LPBF parts. The objective of this study is, hence, to focus on the effect of remelting on fracture toughness of superalloy Hastelloy X (HX) produced by laser powder bed fusion. This study is based on two previous works made by the authors: one the capacity of remelting/relasing to improve the mechanical properties in tension [8] and the second one on the fracture toughness of Hastelloy X produced by LPBF without remelting [10].

## Experimental procedures

Laser powder bed fusion manufacturing was carried out using a M400 EOS machine located in the Volum-e company in France. Hastelloy X powder was provided by Auber and Duval and is characterized by an average particle diameter of about 30 µm. The chemical composition followed the standard for this material. To investigate the role played by the laser remelting on the fracture toughness of HX, three manufacturing strategies were considered to produce mechanical samples. A first conventional strategy without remelting was employed to generate reference data on mechanical properties. This strategy is called HX\_ref in the following paragraphs and parameters related to this production are summarized in table 1. Two others manufacturing strategies with remelting/relasing (a first laser scan melts the powder and a second scan interacts with the solidified material) were also employed based on a previous work [8]. The first one, denoted HX\_R100 corresponds to a remelting with the exact same parameters as the reference one, in particular the laser power which is kept as 100% of the nominal value. The second one - HX\_R60 - is similar with, nevertheless, a laser power set up to 60% of the nominal value employed for the two previous strategies. This 60% relasing power was found to provide the best compromise between yield stress and ductility in tension for vertical samples [8]. For the two relasing/remelting strategies the laser follows the exact same path for the two scans.

<table><caption>Table 1: Laser Power Bed Fusion manufacturing parameters (SLM EOS M400-4)</caption><thead><tr><td>Laser power [W]</td><td>Laser scanning speed [mm/s]</td><td>Hatch distance [mm]</td><td>Powder layer thickness [µm]</td><td>Layer to layer angle misorientation [°]</td></tr></thead><tbody><tr><td>288</td><td>960</td><td>0.11</td><td>40</td><td>67</td></tr></tbody></table>

To estimate the material fracture toughness, ASTM E1820 standard was applied [11]. It consists in investigating the crack propagation in Single Edge Notched samples submitted to bending (SENB tests). The fracture toughness is then estimated by means of the critical strain energy release $J_{tc}$ which is computed from the crack tip opening displacement as well as from mechanical properties in tension. Consequently, in order to estimate fracture toughness values as accurately as possible, both bending samples and tensile tests ones were produced using the three manufacturing strategies described above. As the mechanical behavior of this material produced by LPBF has been proved to be anisotropic, two sample configurations, vertical (V) and horizontal (H) were manufactured. In the specific case of fracture toughness samples, the crack is initially oriented parallel and perpendicular to the lasing plane for the V and H configuration, in the order given. For the tensile samples, they were extracted from a 20x20x80 mm block using EDM. Each test was repeated three times. Figure 1 illustrates the sample geometry for the two mechanical tests. Sample thickness for tension is about 1 mm.
Nevertheless, those remelting strategies have not been extensively characterized for fatigue or fracture toughness improvements which are key properties for an industrial use of many LPBF parts. The objective of this study is, hence, to focus on the effect of remelting on fracture toughness of superalloy Hastelloy X (HX) produced by laser powder bed fusion. This study is based on two previous works made by the authors: one the capacity of remelting/relasing to improve the mechanical properties in tension [8] and the second one on the fracture toughness of Hastelloy X produced by LPBF without remelting [10].

## Experimental procedures

Laser powder bed fusion manufacturing was carried out using a M400 EOS machine located in the Volum-e company in France. Hastelloy X powder was provided by Auber and Duval and is characterized by an average particle diameter of about 30 µm. The chemical composition followed the standard for this material. To investigate the role played by the laser remelting on the fracture toughness of HX, three manufacturing strategies were considered to produce mechanical samples. A first conventional strategy without remelting was employed to generate reference data on mechanical properties. This strategy is called HX\_ref in the following paragraphs and parameters related to this production are summarized in table 1. Two others manufacturing strategies with remelting/relasing (a first laser scan melts the powder and a second scan interacts with the solidified material) were also employed based on a previous work [8]. The first one, denoted HX\_R100 corresponds to a remelting with the exact same parameters as the reference one, in particular the laser power which is kept as 100% of the nominal value. The second one - HX\_R60 - is similar with, nevertheless, a laser power set up to 60% of the nominal value employed for the two previous strategies. This 60% relasing power was found to provide the best compromise between yield stress and ductility in tension for vertical samples [8]. For the two relasing/remelting strategies the laser follows the exact same path for the two scans.

<table><caption>Table 1: Laser Power Bed Fusion manufacturing parameters (SLM EOS M400-4)</caption><thead><tr><td>Laser power [W]</td><td>Laser scanning speed [mm/s]</td><td>Hatch distance [mm]</td><td>Powder layer thickness [µm]</td><td>Layer to layer angle misorientation [°]</td></tr></thead><tbody><tr><td>288</td><td>960</td><td>0.11</td><td>40</td><td>67</td></tr></tbody></table>

To estimate the material fracture toughness, ASTM E1820 standard was applied [11]. It consists in investigating the crack propagation in Single Edge Notched samples submitted to bending (SENB tests). The fracture toughness is then estimated by means of the critical strain energy release $J_{tc}$ which is computed from the crack tip opening displacement as well as from mechanical properties in tension. Consequently, in order to estimate fracture toughness values as accurately as possible, both bending samples and tensile tests ones were produced using the three manufacturing strategies described above. As the mechanical behavior of this material produced by LPBF has been proved to be anisotropic, two sample configurations, vertical (V) and horizontal (H) were manufactured. In the specific case of fracture toughness samples, the crack is initially oriented parallel and perpendicular to the lasing plane for the V and H configuration, in the order given. For the tensile samples, they were extracted from a 20x20x80 mm block using EDM. Each test was repeated three times. Figure 1 illustrates the sample geometry for the two mechanical tests. Sample thickness for tension is about 1 mm.
<figure><img src="image_123.png" /><figcaption>Figure 1: Illustration of the sample geometry for: (a) single edge notched bending tests and (b) tensile tests. All dimensions are in mm.</figcaption></figure>

Tensile tests were conducted in strain rate control (10⁻⁴ s⁻¹) mode using a conventional extensometer whereas the SENB tests were carried out in displacement-controlled mode (1 mm/min). The digital image monitoring of the crack propagation and the data analysis method for computing *J*<sub>IC</sub> are described in a previous work [10]. Digital images were binarized and analyzed using a Scilab script enabling the crack length to be measured as well as the Crack Tip Opening Displacement (CTOD) at any time of the loading [10].

## Results and discussion

Figure 2 illustrates the tensile behavior of the H and V samples manufactured using the three LPBF strategies. Different trends can be observed. On the one hand, despite different lasing strategies, for each direction (H or V), the tensile curves remain similar, in particular in the vertical direction. It can be also observed that the HX_R100 samples exhibit lower strain hardening, as already reported for cylindrical samples in a previous study [8]. On the other hand, regardless the lasing strategy, samples manufactured horizontally exhibit larger flow stress than the vertical direction. Nevertheless, this anisotropy seems to be lower for the samples manufactured with relasing, especially the HX_R100 strategy. Due to the low thickness of the tensile samples, fracture occurs without necking when maximal stress is reached.

<figure><img src="image_124.png" /><figcaption>Figure 2. Engineering tensile curves of Hastelloy X samples manufactured with or without relasing in the horizontal or vertical configurations.</figcaption></figure>
<figure><img src="image_124.png" /><figcaption>Figure 1: Illustration of the sample geometry for: (a) single edge notched bending tests and (b) tensile tests. All dimensions are in mm.</figcaption></figure>

Tensile tests were conducted in strain rate control (10⁻⁴ s⁻¹) mode using a conventional extensometer whereas the SENB tests were carried out in displacement-controlled mode (1 mm/min). The digital image monitoring of the crack propagation and the data analysis method for computing *J*<sub>IC</sub> are described in a previous work [10]. Digital images were binarized and analyzed using a Scilab script enabling the crack length to be measured as well as the Crack Tip Opening Displacement (CTOD) at any time of the loading [10].

## Results and discussion

Figure 2 illustrates the tensile behavior of the H and V samples manufactured using the three LPBF strategies. Different trends can be observed. On the one hand, despite different lasing strategies, for each direction (H or V), the tensile curves remain similar, in particular in the vertical direction. It can be also observed that the HX_R100 samples exhibit lower strain hardening, as already reported for cylindrical samples in a previous study [8]. On the other hand, regardless the lasing strategy, samples manufactured horizontally exhibit larger flow stress than the vertical direction. Nevertheless, this anisotropy seems to be lower for the samples manufactured with relasing, especially the HX_R100 strategy. Due to the low thickness of the tensile samples, fracture occurs without necking when maximal stress is reached.

<figure><img src="image_125.png" /><figcaption>Figure 2. Engineering tensile curves of Hastelloy X samples manufactured with or without relasing in the horizontal or vertical configurations.</figcaption></figure>
The bending curves representing the flexural force versus displacement are depicted in figure 3(a) and figure 3(b) for the vertical and horizontal configuration, respectively. The different curves are similar with a first stage characterized by an increasing force related to the elastoplastic behavior and further hardening of the material. The second stage characterizes the crack propagation associated with a significant decrease in the flexural force. The crack initiation is expected to occur during the end of stage I but not necessarily at the maximal force. As for tensile conditions, the mechanical behavior in bending is more anisotropic for the horizontal direction compared to the vertical one. It can be also observed that plasticity in the HX_R60 appears for lower forces than for the other two configurations.

<figure><img src="image_125.png" /><figcaption>Figure 3: Bending curves for the single edged notch Hastelloy samples manufactured with the three LPBF configurations with or without relasing: (a) vertical and (b) horizontal configurations.</figcaption></figure>

In order to precisely estimate the fracture toughness at initiation, the Crack Tip Opening Displacement $\delta$ was computed during the test by means of a CCD camera allowing the crack propagation to be monitored. This parameter is computed according to eq. 1 [10]:

$$ \delta = \delta_{elastic} + \delta_{plastic} = \frac{K^2(1-v^2)}{2\sigma_0 E} + \frac{r_p(W-a_0)V_p}{r_p(W-a_0)+a_0} \quad (1) $$

In this equation $K$ is the stress intensity factor relative to the specimen, the initial crack geometry and load at initiation, $v$ is the Poisson ratio, $E$ is the Young modulus, and $r_p$ is the plastic rotational factor ($r_{p,SENB} \approx 0.44$) whereas $W$ and $a_0$ correspond to the sample width and the initial crack length, respectively. $\sigma_0$ represents the yield stress and $V_p$ the crack mouth displacement at initiation measured by image analysis.

The critical strain energy release can be hence computed from the value $\delta$ at initiation as well as the yield stress and ultimate tensile stress provided by the tensile tests for all manufacturing conditions as detailed in eq. 2:

$$ J_{lc} = m \left( \frac{\sigma_0 + \sigma_{UTS}}{2} \right) \delta \quad (2) $$
The bending curves representing the flexural force versus displacement are depicted in figure 3(a) and figure 3(b) for the vertical and horizontal configuration, respectively. The different curves are similar with a first stage characterized by an increasing force related to the elastoplastic behavior and further hardening of the material. The second stage characterizes the crack propagation associated with a significant decrease in the flexural force. The crack initiation is expected to occur during the end of stage I but not necessarily at the maximal force. As for tensile conditions, the mechanical behavior in bending is more anisotropic for the horizontal direction compared to the vertical one. It can be also observed that plasticity in the HX_R60 appears for lower forces than for the other two configurations.

<figure><img src="image_126.png" /><figcaption>Figure 3: Bending curves for the single edged notch Hastelloy samples manufactured with the three LPBF configurations with or without relasing: (a) vertical and (b) horizontal configurations.</figcaption></figure>

In order to precisely estimate the fracture toughness at initiation, the Crack Tip Opening Displacement $\delta$ was computed during the test by means of a CCD camera allowing the crack propagation to be monitored. This parameter is computed according to eq. 1 [10]:

$$ \delta = \delta_{elastic} + \delta_{plastic} = \frac{K^2(1-\nu^2)}{2\sigma_0 E} + \frac{r_p(W-a_0)V_p}{r_p(W-a_0)+a_0} \quad (1) $$

In this equation $K$ is the stress intensity factor relative to the specimen, the initial crack geometry and load at initiation, $\nu$ is the Poisson ratio, $E$ is the Young modulus, and $r_p$ is the plastic rotational factor ($r_{p,SENB} \approx 0.44$) whereas $W$ and $a_0$ correspond to the sample width and the initial crack length, respectively. $\sigma_0$ represents the yield stress and $V_p$ the crack mouth displacement at initiation measured by image analysis.

The critical strain energy release can be hence computed from the value $\delta$ at initiation as well as the yield stress and ultimate tensile stress provided by the tensile tests for all manufacturing conditions as detailed in eq. 2:

$$ J_{lc} = m \left( \frac{\sigma_0 + \sigma_{UTS}}{2} \right) \delta \quad (2) $$
In that equation, $\sigma_0$ and $\sigma_{UTS}$ represent the yield stress and ultimate tensile strength, respectively and $m$ is a parameter depending on the crack geometry and on the material's strain hardening which can be empirically estimated using ASTM E1290 [12] and eq. 3:

$$m = 1.221 + 0.793(a_0/W) + 2.751n - 1.418(a_0/W)n \quad (3)$$

$n$ is, in that case, the strain hardening exponent which can be computed using eq. 4 following ASTM E1290 standard. $R$ represents $R = \sigma_{UTS}/\sigma_0$ computed using tensile data.

$$n = 1.724 - 6.098/R + 8.326/R^2 - 3.965/R^3 \quad (4)$$

Using the actual values of the bending sample geometry, the tensile data as well as the Crack Tip Opening Displacement at initiation, the fracture toughness $J_{IC}$ can be estimated in all configurations (vertical and horizontal), and for the three manufacturing strategies. Figure 4 illustrates the fracture toughness values for all configurations tested in this study. From these values, it turns that, considering the experimental scattering, the manufacturing strategy does not seem to significantly affect the fracture toughness at initiation in the vertical configuration. Besides, a slightly improved value of fracture toughness is exhibited for the relasing configuration with 60% of the nominal power in agreement with the increase in ductility already observed for the vertical configuration [8]. For the horizontal configuration, the relasing strategy has a major effect with a maximal value of $J_{IC}$ of about 1300 kJ.m⁻² for the standard manufacturing strategy. The value obtained for relasing strategies are about 1000 kJ.m⁻² which are lower than for the vertical configuration but still widely larger than for the cast (230 kJ.m⁻² [13]). One can also observed that the fracture toughness anisotropy almost disappeared for the HX_R100 configuration as also observed in figure 2 for tensile conditions.

<figure><img src="image_126.png" /><figcaption>Figure 4: Fracture toughness estimation for Hastelloy X samples manufactured with different LPBF strategies and building directions (vertical and horizontal).</figcaption></figure>

Following the conclusions of previous works conducted on Inconel 718 [14] and 316L [15] printed by LPBF, the horizontal configuration is generally detrimental to the fracture toughness when the crack propagates alongside the building direction due to the preferential intergranular mode involved by the elongation of the grains with respect to this direction. Similar mechanisms could be operating in the case of the Hastelloy X for the relasing strategies. Microstructural observation must be performed to understand the origin of the lower fracture toughness in the horizontal configuration for both relasing strategies compared to the reference case.
In that equation, $\sigma_0$ and $\sigma_{UTS}$ represent the yield stress and ultimate tensile strength, respectively and $m$ is a parameter depending on the crack geometry and on the material's strain hardening which can be empirically estimated using ASTM E1290 [12] and eq. 3:

$$m = 1.221 + 0.793(a_0/W) + 2.751n - 1.418(a_0/W)n \quad (3)$$

$n$ is, in that case, the strain hardening exponent which can be computed using eq. 4 following ASTM E1290 standard. $R$ represents $R = \sigma_{UTS}/\sigma_0$ computed using tensile data.

$$n = 1.724 - 6.098/R + 8.326/R^2 - 3.965/R^3 \quad (4)$$

Using the actual values of the bending sample geometry, the tensile data as well as the Crack Tip Opening Displacement at initiation, the fracture toughness $J_{IC}$ can be estimated in all configurations (vertical and horizontal), and for the three manufacturing strategies. Figure 4 illustrates the fracture toughness values for all configurations tested in this study. From these values, it turns that, considering the experimental scattering, the manufacturing strategy does not seem to significantly affect the fracture toughness at initiation in the vertical configuration. Besides, a slightly improved value of fracture toughness is exhibited for the relasing configuration with 60% of the nominal power in agreement with the increase in ductility already observed for the vertical configuration [8]. For the horizontal configuration, the relasing strategy has a major effect with a maximal value of $J_{IC}$ of about 1300 kJ.m⁻² for the standard manufacturing strategy. The value obtained for relasing strategies are about 1000 kJ.m⁻² which are lower than for the vertical configuration but still widely larger than for the cast (230 kJ.m⁻² [13]). One can also observed that the fracture toughness anisotropy almost disappeared for the HX_R100 configuration as also observed in figure 2 for tensile conditions.

<figure><img src="image_127.png" /><figcaption>Figure 4: Fracture toughness estimation for Hastelloy X samples manufactured with different LPBF strategies and building directions (vertical and horizontal).</figcaption></figure>

Following the conclusions of previous works conducted on Inconel 718 [14] and 316L [15] printed by LPBF, the horizontal configuration is generally detrimental to the fracture toughness when the crack propagates alongside the building direction due to the preferential intergranular mode involved by the elongation of the grains with respect to this direction. Similar mechanisms could be operating in the case of the Hastelloy X for the relasing strategies. Microstructural observation must be performed to understand the origin of the lower fracture toughness in the horizontal configuration for both relasing strategies compared to the reference case.
## Conclusions

Based on the characterization of tensile and fracture behavior, results from this study show that remelting affects the overall mechanical properties with, in particular:

* A decrease in the flow stress with remelting in the horizontal configuration.

* A detrimental effect on fracture toughness for horizontal configuration (crack propagating perpendicularly to the lasing planes) independently of the relasing strategy. Values are nevertheless five times larger than the one reported for cast.

* A slightly beneficial effect on fracture toughness for low relasing power (60%) in the vertical configuration (crack lying in the lasing planes).

A detailed analysis on the microstructure is then needed to better understand the origin of such effect of the remelting strategy.

## References

[1] M.V. Pantawane, Y-H. Ho, S. S. Joshi, N. B. Dahotre, Computational Assessment of Thermokinetics and Associated Microstructural Evolution in Laser Powder Bed Fusion Manufacturing of Ti6Al4V Alloy, Scientific reports, 10 (2020) 7579. https://doi.org/10.1038/s41598-020-63281-4

[2] Y.M. Wang, T. Voisin, J.T. McKeown, J. Ye, N. P. Calta, Z. Li, Z. Zeng, Y. Zhang, W. Chen, T.T. Roehling, R.T. Ott, M.K. Santala, P.J. Depond, M.J. Matthews, A.V. Hamza, T. Zhu, Additively manufactured hierarchical stainless steels with high strength and ductility. Nature Mater 17 (2018) 63–71. https://doi.org/10.1038/nmat5021

[3] T. DebRoy, H.L. Wei, J.S. Zuback, T. Mukherjee, J. W. Elmer, J. O. Milewski, A. M. Beese, A. Wilson-Heid, A. De, W. Zhang, Additive manufacturing of metallic components – Process, structure and properties. Progress in Materials Science 92 (2018) 112–224. https://doi.org/10.1016/j.pmatsci.2017.10.001

[4] S. Gorsse, C. Hutchinson, M. Gouné, R. Banerjee, Additive manufacturing of metals: a brief review of the characteristic microstructures and properties of steels, Ti-6Al-4V and high-entropy alloys. Science and Technology of Advanced Materials 18 (2017) 584–610. https://doi.org/10.1080/14686996.2017.1361305

[5] O. Karaka, F. B. Kardes, P. Fito, F. Berto, An overview of factors affecting high cycle fatigue of additive manufacturing metals, FFEMS 46 (2023) 1649-1668. https://doi.org/10.1111/ffe.13967

[6] S. Griffiths, M. Rossell, J. Croteau, N. Vo, D. C. Dunand, C. Leinenbach, Effect of laser rescanning on the grain microstructure of a selective laser melted al-mg-zr alloy, Materials Characterization 143 (2018) 34–42. https://doi.org/10.1016/j.matchar.2018.03.033

[7] E. Yasa, J. Deckers, J.-P. Kruth, The investigation of the influence of laser re-melting on density, surface quality and microstructure of selective laser melting parts, Rapid Prototyping Journal 17 (5) (2011) 312–327. https://doi.org/10.1108/13552541111156450

[8] C. Keller, M. Mokhtari, B. Vieille, H. Briatta, P. Bernard, Influence of a rescanning strategy with different laser powers on the microstructure and mechanical properties of Hastelloy X elaborated by powder bed fusion, Mat. Sci. Eng. A803 (2021) 1404754. https://doi.org/10.1016/j.msea.2020.140474

[9] Z. Xiao, C. Chen, Z. Hu, H. Zhu, X. Zeng, Effect of rescanning cycles on the characteristics of selective laser melting of ti6al4v, Optics Laser Technology 122 (2020) 105890. https://doi.org/10.1016/j.optlastec.2019.105890
## Conclusions

Based on the characterization of tensile and fracture behavior, results from this study show that remelting affects the overall mechanical properties with, in particular:

* A decrease in the flow stress with remelting in the horizontal configuration.

* A detrimental effect on fracture toughness for horizontal configuration (crack propagating perpendicularly to the lasing planes) independently of the relasing strategy. Values are nevertheless five times larger than the one reported for cast.

* A slightly beneficial effect on fracture toughness for low relasing power (60%) in the vertical configuration (crack lying in the lasing planes).

A detailed analysis on the microstructure is then needed to better understand the origin of such effect of the remelting strategy.

## References

[1] M.V. Pantawane, Y-H. Ho, S. S. Joshi, N. B. Dahotre, Computational Assessment of Thermokinetics and Associated Microstructural Evolution in Laser Powder Bed Fusion Manufacturing of Ti6Al4V Alloy, Scientific reports, 10 (2020) 7579. https://doi.org/10.1038/s41598-020-63281-4

[2] Y.M. Wang, T. Voisin, J.T. McKeown, J. Ye, N. P. Calta, Z. Li, Z. Zeng, Y. Zhang, W. Chen, T.T. Roehling, R.T. Ott, M.K. Santala, P.J. Depond, M.J. Matthews, A.V. Hamza, T. Zhu, Additively manufactured hierarchical stainless steels with high strength and ductility. Nature Mater 17 (2018) 63–71. https://doi.org/10.1038/nmat5021

[3] T. DebRoy, H.L. Wei, J.S. Zuback, T. Mukherjee, J. W. Elmer, J. O. Milewski, A. M. Beese, A. Wilson-Heid, A. De, W. Zhang, Additive manufacturing of metallic components – Process, structure and properties. Progress in Materials Science 92 (2018) 112–224. https://doi.org/10.1016/j.pmatsci.2017.10.001

[4] S. Gorsse, C. Hutchinson, M. Gouné, R. Banerjee, Additive manufacturing of metals: a brief review of the characteristic microstructures and properties of steels, Ti-6Al-4V and high-entropy alloys. Science and Technology of Advanced Materials 18 (2017) 584–610. https://doi.org/10.1080/14686996.2017.1361305

[5] O. Karaka, F. B. Kardes, P. Fito, F. Berto, An overview of factors affecting high cycle fatigue of additive manufacturing metals, FFEMS 46 (2023) 1649-1668. https://doi.org/10.1111/ffe.13967

[6] S. Griffiths, M. Rossell, J. Croteau, N. Vo, D. C. Dunand, C. Leinenbach, Effect of laser rescanning on the grain microstructure of a selective laser melted al-mg-zr alloy, Materials Characterization 143 (2018) 34–42. https://doi.org/10.1016/j.matchar.2018.03.033

[7] E. Yasa, J. Deckers, J.-P. Kruth, The investigation of the influence of laser re-melting on density, surface quality and microstructure of selective laser melting parts, Rapid Prototyping Journal 17 (5) (2011) 312–327. https://doi.org/10.1108/13552541111156450

[8] C. Keller, M. Mokhtari, B. Vieille, H. Briatta, P. Bernard, Influence of a rescanning strategy with different laser powers on the microstructure and mechanical properties of Hastelloy X elaborated by powder bed fusion, Mat. Sci. Eng. A803 (2021) 1404754. https://doi.org/10.1016/j.msea.2020.140474

[9] Z. Xiao, C. Chen, Z. Hu, H. Zhu, X. Zeng, Effect of rescanning cycles on the characteristics of selective laser melting of ti6al4v, Optics Laser Technology 122 (2020) 105890. https://doi.org/10.1016/j.optlastec.2019.105890
[10] B. Vieille, A. Duchaussoy, S. Benmabrouk, R. Henry, C. Keller, Fracture behavior of Hastelloy X elaborated by laser powder bed fusion: Influence of microstructure and building direction, Journal of Alloys and Compounds 918 (2022) 165570. https://doi.org/10.1016/j.jallcom.2022.165570

[11] ASTM E1820-01, Standard Test Method for Measurement of Fracture Toughness, ASTM International, West Conshohocken, PA, 2001.

[12] ASTM E1290-90. Standard test method for crack-tip opening displacement (CTOD) fracture toughness measurement. American Society for Testing and Materials; 1998

[13] K. Krompholz, E.D. Grosser, K. Ewert. Determination of J-integral R-curves for Hastelloyx X and Inconel 617 up to 1223K using the potential drop technique. Materialwissenschaft Und Werkstofftechnik (1982), 13, pp. 236-244. https://doi.org/10.1002/mawe.19820130704

[14] B. Vieille, C. Keller, M. Mokhtari, H. Briatta, T. Breteau, J. Nguejio, F. Barbe, M. Ben Azzouna, E. Baustert, Investigations on the fracture behavior of Inconel 718 superalloys obtained from cast and additive manufacturing processes, Mat. Sci. Eng, A790 (2020) 139666. https://doi.org/10.1016/j.msea.2020.139666

[15] E. De Sonis, S. Depinoy, P-F. Giroux, H. Maskrot, P. Wident, O. Hercher, F. Villaret, A-F. Gourgues-Lorezon, Microstructure – Toughness relationships in 316L stainless steel produced by laser powder bed fusion, Mat. Sci. Eng, A877 (2023) 145179. https://doi.org/10.1016/j.msea.2023.145179
[10] B. Vieille, A. Duchaussoy, S. Benmabrouk, R. Henry, C. Keller, Fracture behavior of Hastelloy X elaborated by laser powder bed fusion: Influence of microstructure and building direction, Journal of Alloys and Compounds 918 (2022) 165570. https://doi.org/10.1016/j.jallcom.2022.165570

[11] ASTM E1820-01, Standard Test Method for Measurement of Fracture Toughness, ASTM International, West Conshohocken, PA, 2001.

[12] ASTM E1290-90. Standard test method for crack-tip opening displacement (CTOD) fracture toughness measurement. American Society for Testing and Materials; 1998

[13] K. Krompholz, E.D. Grosser, K. Ewert. Determination of J-integral R-curves for Hastelloyx X and Inconel 617 up to 1223K using the potential drop technique. Materialwissenschaft Und Werkstofftechnik (1982), 13, pp. 236-244. https://doi.org/10.1002/mawe.19820130704

[14] B. Vieille, C. Keller, M. Mokhtari, H. Briatta, T. Breteau, J. Nguejio, F. Barbe, M. Ben Azzouna, E. Baustert, Investigations on the fracture behavior of Inconel 718 superalloys obtained from cast and additive manufacturing processes, Mat. Sci. Eng, A790 (2020) 139666. https://doi.org/10.1016/j.msea.2020.139666

[15] E. De Sonis, S. Depinoy, P-F. Giroux, H. Maskrot, P. Wident, O. Hercher, F. Villaret, A-F. Gourgues-Lorezon, Microstructure – Toughness relationships in 316L stainless steel produced by laser powder bed fusion, Mat. Sci. Eng, A877 (2023) 145179. https://doi.org/10.1016/j.msea.2023.145179