<table><caption>Table 1: Main printing parameter</caption><tr><td>Nozzle diameter</td><td>1 [mm]</td></tr><tr><td>Layer thickness</td><td>0.35 [mm]</td></tr><tr><td>First layer thickness</td><td>0.3 [mm]</td></tr><tr><td>Extrusion temperature</td><td>250 [°C]</td></tr><tr><td>Build platform temperature</td><td>70 [°C]</td></tr><tr><td>Infill strategy</td><td>Lines</td></tr><tr><td>Infill percentage</td><td>100 [%]</td></tr><tr><td>Printing time</td><td>28 [min]</td></tr></table>

<figure><img src="image_1791.png" /><figcaption>Figure 1: Part geometry and main dimensions.</figcaption></figure>

Method

A robotic machining cell composed of a Stäubli TX200 robot fitted with a Teknomotor ATC71 electrospindle was used for the milling operations. The spindle can deliver up to 7.8 kW with a maximal rotational speed of 24000 rpm. A three-jaw chuck was used to clamp the part from the cylinder (Fig. 3) which was attached to a Kistler 9256C2 dynamometer.

The used cutting tool showed in Fig. 2 is originally dedicated to the milling of thermoplastics as PA66. The supplier company is Hoffmann and it's sold with the reference 209425-6. It has a 6 mm diameter with three cutting edges having a double helix angle and a maximal cutting length of 19 mm. It is made in tungsten carbide without any coating.

As presented in a previous study [8], this tool showed great capability to generate smooth and shiny surfaces with $R_a < 1.6 \ \mu m$ in finishing operations [8-9] for composite ceramic/polymer. This study was based on the tool-material couple standard AFNOR NF E66-520-6 (1999) [10] as the present study. Four cutting parameters were used: the cutting speed ($v_c$) corresponding to the speed at which the cutting tool or the workpiece moves relative to each other during machining, the feed per tooth ($f_z$) corresponding to the distance a cutting tool advances per tooth for each revolution of the spindle, the axial depth of cut ($a_p$) representing the depth of the tool's penetration into the workpiece material during a cutting operation, and the radial depth of cut ($a_e$) representing the lateral width of the material removed by the cutting tool. The standard proposes to perform six different experiments [10]: the selection of a stable operating point, the determination of $v_{c,min}$, the
determination of minimal and maximal chip thickness ($h_{min} – h_{max}$), the determination of limiting data, the wear tests and the tests to determine the auxiliary parameters. The experiments of the standard allow the determination of the working range for $v_c, f_z, a_p$, and $a_e$ but its full method cannot be applied since the material milled (green metal) as well as the cutting conditions (finishing) are not covered by the standard.

<figure><img src="image_1794.png" /><figcaption>Figure 2: Microscope image of the milling tool selected for the study.</figcaption></figure>

The initial cutting conditions were selected according to the previous study and are given in Table 2. No lubrication was used for the tests, while compressed air was used between each test to blow the part, the tool and evacuate the chips.

<figure><img src="image_1795.png" /><figcaption>Figure 3: View of the building part.</figcaption></figure>

A Kistler 9256C2 force sensor was used to record the cutting forces for each pass at a 40 kHz sampling frequency without any filters. The total cutting force (Eq. 1) was considered as a representative value to compare each set of cutting parameters.

The total resulting force (in N) was computed as follows in Eq. 1:
determination of minimal and maximal chip thickness ($h_{min} – h_{max}$), the determination of limiting data, the wear tests and the tests to determine the auxiliary parameters. The experiments of the standard allow the determination of the working range for $v_c, f_z, a_p$, and $a_e$ but its full method cannot be applied since the material milled (green metal) as well as the cutting conditions (finishing) are not covered by the standard.

<figure><img src="image_1792.png" /><figcaption>Figure 2: Microscope image of the milling tool selected for the study.</figcaption></figure>

The initial cutting conditions were selected according to the previous study and are given in Table 2. No lubrication was used for the tests, while compressed air was used between each test to blow the part, the tool and evacuate the chips.

<figure><img src="image_1793.png" /><figcaption>Figure 3: View of the building part.</figcaption></figure>

A Kistler 9256C2 force sensor was used to record the cutting forces for each pass at a 40 kHz sampling frequency without any filters. The total cutting force (Eq. 1) was considered as a representative value to compare each set of cutting parameters.

The total resulting force (in N) was computed as follows in Eq. 1:
$$F_{tot} = \sqrt{(F_{x}^{2} + F_{y}^{2} + F_{z}^{2})} \qquad (1)$$

The first step consists in realizing 3 passes (with radial and axial depths of cut according to Table 2) from the top of the cube (Fig. 1) along the X-axis of the parts.

The arithmetic roughness and the total roughness are then measured, and a microscope view is realized after each pass.

After each step, the generated vertical surface was quantitatively and qualitatively assessed. The cutting forces were recorded for each of the passes. The surface topography was evaluated quantitatively using a Diavite DH6 contact rugosimeter with parameters following the ISO 4288 standard. The arithmetic roughness $R_a$ was obtained with an evaluation length of 4.8 mm and a cut-off length of 0.8 mm. For all measurements, the probe was displaced according to the parts X-axis (feed direction of the cutting tool). Each vertical surface was characterized by three $R_a$ and $R_t$ (corresponding to the vertical distance between the highest and lowest peak within the measured length of a surface profile) measurements, one at the beginning, middle and end of each pass. In total, 36 measurements of roughness were made. The initial surface roughness of the as-built part was estimated using a rugosimeter ZEISS Surfcom 1400. The general aspect of the generated surfaces (porosities, material pull-out) was observed qualitatively using a digital microscope AM7013MZT and the DinoCapture software from DinoLite.

The measurement of the $R_a$ and $R_t$ parameters is carried out in order to know what the minimum amount of material to remove in $a_e$ and to determine if it is possible to be below a $R_a$ of 3.2 µm which represents the class of mechanical assembly parts, which is very interesting from an industrial point of view.

<table><caption>Table 2: Set of cutting parameters used during the study.</caption><thead><tr><th>Configuration</th><th><i>v</i><sub>c</sub><br>[m/min]</th><th><i>f</i><sub>z</sub><br>[mm/tooth]</th><th><i>a</i><sub>p</sub><br>[mm]</th><th><i>a</i><sub>e</sub><br>[mm]</th></tr></thead><tbody><tr><td>1</td><td>188.4</td><td>0.03</td><td>3</td><td>0.5</td></tr><tr><td>2</td><td>235.5</td><td>0.03</td><td>3</td><td>0.5</td></tr><tr><td>3</td><td>282.6</td><td>0.03</td><td>3</td><td>0.5</td></tr><tr><td>4</td><td>329.7</td><td>0.03</td><td>3</td><td>0.5</td></tr><tr><td>5</td><td>383</td><td>0.03</td><td>3</td><td>0.5</td></tr><tr><td>6</td><td>423.9</td><td>0.03</td><td>3</td><td>0.5</td></tr><tr><td>7</td><td>383</td><td>0.026</td><td>3</td><td>0.5</td></tr><tr><td>8</td><td>383</td><td>0.035</td><td>3</td><td>0.5</td></tr><tr><td>9</td><td>383</td><td>0.03</td><td>3</td><td>0.43</td></tr><tr><td>10</td><td>383</td><td>0.03</td><td>3</td><td>0.58</td></tr><tr><td>11</td><td>383</td><td>0.03</td><td>2.6</td><td>0.5</td></tr><tr><td>12</td><td>383</td><td>0.03</td><td>3.5</td><td>0.5</td></tr></tbody></table>

A first set of parameters was tried with $v_c$= 383 m/min based on [8] then a variation of 15% was applied to know if it is a stable point as in the previous study or not. The same was tried on $f_z$, $a_p$ and $a_e$ to evaluate their impact on the surface roughness and in cutting forces.

## Results

Fig. 4 gives the view of the zones before and after the machining operation of the as-built part. The surface before cutting (blue line) can be seen on the bottom and the surface after cutting (red line) on the top.
$$F_{tot} = \sqrt{(F_{x}^{2} + F_{y}^{2} + F_{z}^{2})} \qquad (1)$$

The first step consists in realizing 3 passes (with radial and axial depths of cut according to Table 2) from the top of the cube (Fig. 1) along the X-axis of the parts.

The arithmetic roughness and the total roughness are then measured, and a microscope view is realized after each pass.

After each step, the generated vertical surface was quantitatively and qualitatively assessed. The cutting forces were recorded for each of the passes. The surface topography was evaluated quantitatively using a Diavite DH6 contact rugosimeter with parameters following the ISO 4288 standard. The arithmetic roughness $R_a$ was obtained with an evaluation length of 4.8 mm and a cut-off length of 0.8 mm. For all measurements, the probe was displaced according to the parts X-axis (feed direction of the cutting tool). Each vertical surface was characterized by three $R_a$ and $R_t$ (corresponding to the vertical distance between the highest and lowest peak within the measured length of a surface profile) measurements, one at the beginning, middle and end of each pass. In total, 36 measurements of roughness were made. The initial surface roughness of the as-built part was estimated using a rugosimeter ZEISS Surfcom 1400. The general aspect of the generated surfaces (porosities, material pull-out) was observed qualitatively using a digital microscope AM7013MZT and the DinoCapture software from DinoLite.

The measurement of the $R_a$ and $R_t$ parameters is carried out in order to know what the minimum amount of material to remove in $a_e$ and to determine if it is possible to be below a $R_a$ of 3.2 µm which represents the class of mechanical assembly parts, which is very interesting from an industrial point of view.

<table><caption>Table 2: Set of cutting parameters used during the study.</caption><thead><tr><th>Configuration</th><th><i>v</i><sub>c</sub><br>[m/min]</th><th><i>f</i><sub>z</sub><br>[mm/tooth]</th><th><i>a</i><sub>p</sub><br>[mm]</th><th><i>a</i><sub>e</sub><br>[mm]</th></tr></thead><tbody><tr><td>1</td><td>188.4</td><td>0.03</td><td>3</td><td>0.5</td></tr><tr><td>2</td><td>235.5</td><td>0.03</td><td>3</td><td>0.5</td></tr><tr><td>3</td><td>282.6</td><td>0.03</td><td>3</td><td>0.5</td></tr><tr><td>4</td><td>329.7</td><td>0.03</td><td>3</td><td>0.5</td></tr><tr><td>5</td><td>383</td><td>0.03</td><td>3</td><td>0.5</td></tr><tr><td>6</td><td>423.9</td><td>0.03</td><td>3</td><td>0.5</td></tr><tr><td>7</td><td>383</td><td>0.026</td><td>3</td><td>0.5</td></tr><tr><td>8</td><td>383</td><td>0.035</td><td>3</td><td>0.5</td></tr><tr><td>9</td><td>383</td><td>0.03</td><td>3</td><td>0.43</td></tr><tr><td>10</td><td>383</td><td>0.03</td><td>3</td><td>0.58</td></tr><tr><td>11</td><td>383</td><td>0.03</td><td>2.6</td><td>0.5</td></tr><tr><td>12</td><td>383</td><td>0.03</td><td>3.5</td><td>0.5</td></tr></tbody></table>

A first set of parameters was tried with $v_c$= 383 m/min based on [8] then a variation of 15% was applied to know if it is a stable point as in the previous study or not. The same was tried on $f_z$, $a_p$ and $a_e$ to evaluate their impact on the surface roughness and in cutting forces.

## Results

Fig. 4 gives the view of the zones before and after the machining operation of the as-built part. The surface before cutting (blue line) can be seen on the bottom and the surface after cutting (red line) on the top.
<figure><img src="image_1796.png" /><figcaption>Figure 4: Microscope view of the zone after cutting operations.</figcaption></figure>

The as-built surface topography is very rough on its sides with a $R_a$ > 38 µm, $R_t$ > 50 µm and the zone after machining exhibits a smooth surface. The line presents in this zone machined (surrounded) is due to the geometry of the cutting tool, this line of non-cut material is removed when we go down in the part.

<figure><img src="image_1797.png" /><figcaption>Figure 5: Arithmetic surface roughness $R_a$ (µm), total surface roughness $R_t$ (µm) and cutting forces of the building part after cutting tests depending on the cutting speed $v_c$ (m/min) corresponding to tests n°1, 2, 3, 4, 5 and 6 (cf. table n°2).</figcaption></figure>

Influence of the cutting speed: Fig. 5 shows the arithmetic and total surface roughness of the machined zone after each test from test n°1 to test n°6 with a constant variation of the cutting speed ranging from 188.4 m/min to 423.9 m/min. The tool showed a great capability to generate a good surface roughness (< 3.6 µm). The as-built part has a surface topography with an average $R_a$ estimated to be between 2.89 µm and 3.28 µm corresponding to a difference of 13.5% and the total roughness $R_t$ ranges vary from 19.44 µm to 24.55 µm corresponding to a difference of 26%. The best cutting speed range seems to be between 282.6 and 383 m/min with a $R_a$ difference around
<figure><img src="image_1794.png" /><figcaption>Figure 4: Microscope view of the zone after cutting operations.</figcaption></figure>

The as-built surface topography is very rough on its sides with a $R_a$ > 38 µm, $R_t$ > 50 µm and the zone after machining exhibits a smooth surface. The line presents in this zone machined (surrounded) is due to the geometry of the cutting tool, this line of non-cut material is removed when we go down in the part.

<figure><img src="image_1795.png" /><figcaption>Figure 5: Arithmetic surface roughness $R_a$ (µm), total surface roughness $R_t$ (µm) and cutting forces of the building part after cutting tests depending on the cutting speed $v_c$ (m/min) corresponding to tests n°1, 2, 3, 4, 5 and 6 (cf. table n°2).</figcaption></figure>

Influence of the cutting speed: Fig. 5 shows the arithmetic and total surface roughness of the machined zone after each test from test n°1 to test n°6 with a constant variation of the cutting speed ranging from 188.4 m/min to 423.9 m/min. The tool showed a great capability to generate a good surface roughness (< 3.6 µm). The as-built part has a surface topography with an average $R_a$ estimated to be between 2.89 µm and 3.28 µm corresponding to a difference of 13.5% and the total roughness $R_t$ ranges vary from 19.44 µm to 24.55 µm corresponding to a difference of 26%. The best cutting speed range seems to be between 282.6 and 383 m/min with a $R_a$ difference around
4% and a difference of R<sub>t</sub> around 11%. The cutting forces are low with a minimum of 1.62 N and a maximum of 4.46 N. The cutting forces are < 5 N for each test and there is no material pull out.

<figure><img src="image_1798.png" /><figcaption>Figure 6: Arithmetic surface roughness R<sub>a</sub> (µm) and total surface roughness R<sub>t</sub> (µm) of the building part after cutting tests depending on the feed per tooth f<sub>z</sub> (mm/tooth) corresponding to tests n°5, 7 and 8 (cf. table n°2).</figcaption></figure>

Influence of the feed per tooth: the surface roughness of the part is affected by the feed per tooth in the case of the PolyMIM Ti6Al4V as shown in Fig. 6. In fact, the results exhibit a minimum corresponding to the better and more reproducible R<sub>a</sub> and R<sub>t</sub> measurements for the test n°5 with a feed per tooth of 0.03 mm/tooth (variation of 18% for R<sub>a</sub> and 12% for R<sub>t</sub>). The variability of the results is evolving for R<sub>t</sub>, more the feed per tooth increases, more the standard deviation in the results increases, reflecting a decrease in the stability of the results as well as an output of the desired arithmetic surface roughness R<sub>a</sub> < 3.2 µm. The cutting forces are stable and between 2.81 and 3.3 N.

<figure><img src="image_1799.png" /><figcaption>Figure 7: Arithmetic surface roughness R<sub>a</sub> (µm) and total surface roughness R<sub>t</sub> (µm) of the building part after cutting tests depending on the radial depth of cut a<sub>e</sub> (mm) corresponding to tests n°5, 9 and 10 (cf. table n°2).</figcaption></figure>
4% and a difference of R<sub>t</sub> around 11%. The cutting forces are low with a minimum of 1.62 N and a maximum of 4.46 N. The cutting forces are < 5 N for each test and there is no material pull out.

<figure><img src="image_1796.png" /><figcaption>Figure 6: Arithmetic surface roughness R<sub>a</sub> (μm) and total surface roughness R<sub>t</sub> (μm) of the building part after cutting tests depending on the feed per tooth f<sub>z</sub> (mm/tooth) corresponding to tests n°5, 7 and 8 (cf. table n°2).</figcaption></figure>

Influence of the feed per tooth: the surface roughness of the part is affected by the feed per tooth in the case of the PolyMIM Ti6Al4V as shown in Fig. 6. In fact, the results exhibit a minimum corresponding to the better and more reproducible R<sub>a</sub> and R<sub>t</sub> measurements for the test n°5 with a feed per tooth of 0.03 mm/tooth (variation of 18% for R<sub>a</sub> and 12% for R<sub>t</sub>). The variability of the results is evolving for R<sub>t</sub>, more the feed per tooth increases, more the standard deviation in the results increases, reflecting a decrease in the stability of the results as well as an output of the desired arithmetic surface roughness R<sub>a</sub> < 3.2 μm. The cutting forces are stable and between 2.81 and 3.3 N.

<figure><img src="image_1797.png" /><figcaption>Figure 7: Arithmetic surface roughness R<sub>a</sub> (μm) and total surface roughness R<sub>t</sub> (μm) of the building part after cutting tests depending on the radial depth of cut a<sub>e</sub> (mm) corresponding to tests n°5, 9 and 10 (cf. table n°2).</figcaption></figure>
Influence of the radial depth of cut $a_e$: the observations of the Fig. 7 show that the radial depth of cut $a_e$ have a strong impact on the surface roughness of the part. The arithmetic roughness $R_a$ and the total roughness $R_t$ increase (around 20% for $R_a$ and 9% for $R_t$) when $a_e$ increases between 0.5 and 0.58 mm. The reproducibility of the results is affected too, in fact, the variation grows up between results for a same test with the increase of $a_e$. This result shows that optimal $a_e$ for the used cutting tool is equal or below 0.5 mm. The cutting forces are stable and remains under 3.28 and 3.3 N.

<figure><img src="image_1800.png" /><figcaption>Figure 8: Arithmetic surface roughness $R_a$ (µm) and total surface roughness $R_t$ (µm) of the building part after cutting tests depending on the axial depth of cut $a_p$ (mm) corresponding to tests n°5, 11 and 12 (cf. table n°2).</figcaption></figure>

Influence of the axial depth of cut $a_p$: the variation of the axial depth of cut $a_p$ seems to have a slight impact on the surface roughness of the part as shown in Fig. 8. In fact, a maximal variation around 12% for $R_a$ and 5% for $R_t$ is present between the minimum and maximum points. All configurations are in the same surface roughness class but the best range of $a_p$ is from 3 mm.

It can be noted that the cutting forces reach a minimum of 1.89 N and a maximum of 3.3 N, which is similar to the other parameters configurations.

The trends given by the $R_a$ and $R_t$ results are globally similar and relatively stable for most of the tested parameters. This is of industrial interest thanks to the wide range of use which allows the operator to choose parameters more easily according to the needs inherent to the equipment used. A set of parameters seems to stand out from the other with better results which correspond to a feed per tooth $f_z$ of 0.03 mm/tooth, a radial depth of cut $a_e \le 0.5$ mm and an axial depth of cut $a_p = 3$ mm. The feed per tooth, the radial depth of cut and the cutting speed are the parameters with the highest impact on surface roughness in the case of additive manufactured Ti6Al4V part.
Influence of the radial depth of cut $a_e$: the observations of the Fig. 7 show that the radial depth of cut $a_e$ have a strong impact on the surface roughness of the part. The arithmetic roughness $R_a$ and the total roughness $R_t$ increase (around 20% for $R_a$ and 9% for $R_t$) when $a_e$ increases between 0.5 and 0.58 mm. The reproducibility of the results is affected too, in fact, the variation grows up between results for a same test with the increase of $a_e$. This result shows that optimal $a_e$ for the used cutting tool is equal or below 0.5 mm. The cutting forces are stable and remains under 3.28 and 3.3 N.

<figure><img src="image_1798.png" /><figcaption>Figure 8: Arithmetic surface roughness $R_a$ (µm) and total surface roughness $R_t$ (µm) of the building part after cutting tests depending on the axial depth of cut $a_p$ (mm) corresponding to tests n°5, 11 and 12 (cf. table n°2).</figcaption></figure>

Influence of the axial depth of cut $a_p$: the variation of the axial depth of cut $a_p$ seems to have a slight impact on the surface roughness of the part as shown in Fig. 8. In fact, a maximal variation around 12% for $R_a$ and 5% for $R_t$ is present between the minimum and maximum points. All configurations are in the same surface roughness class but the best range of $a_p$ is from 3 mm.

It can be noted that the cutting forces reach a minimum of 1.89 N and a maximum of 3.3 N, which is similar to the other parameters configurations.

The trends given by the $R_a$ and $R_t$ results are globally similar and relatively stable for most of the tested parameters. This is of industrial interest thanks to the wide range of use which allows the operator to choose parameters more easily according to the needs inherent to the equipment used. A set of parameters seems to stand out from the other with better results which correspond to a feed per tooth $f_z$ of 0.03 mm/tooth, a radial depth of cut $a_e \le 0.5$ mm and an axial depth of cut $a_p = 3$ mm. The feed per tooth, the radial depth of cut and the cutting speed are the parameters with the highest impact on surface roughness in the case of additive manufactured Ti6Al4V part.
<figure><img src="image_1801.png" /><figcaption>Figure 9: SEM images of the Ti6Al4V powder of PolyMIM pellets after debinding.</figcaption></figure>

Yoshida [11] did a study to evaluate the impact of grain size and heterogeneity on the surface roughness of sheet metals subjected to plane-strain stretching. According to this work, the surface roughness of parts made by the PolyMIM feedstock does not seem to be able to be significantly improved because of its particle size (average particle size approximatively equal to 20 µm) and because of the heterogeneity of the particle's sizes according to the SEM images presented on Fig. 9 which represents the particles of PolyMIM Ti6Al4V granule after debinding. The size of the particles is very heterogeneous between approx. 5 and 35 µm and this heterogeneity induces inequality on the surface of the printing and machining part.

The cutting forces were recorded for each cutting test. It's possible to see that all measured values remain under 5 N for each set of cutting parameters. The higher value is for the cutting test n°2 sequence 3 with a value of 4.9 N. The material seems to behave as a polymer more than metal which makes easier the machining operation as confirmed by [8] for ceramic loading.

## Conclusions

This study proves the feasibility of machining composite Ti6Al4V/Polymers parts made by Pellets Additive Manufacturing.

The tool dedicated to thermoplastics polymers seems to be adapted to machine the composite Ti6Al4V / Binder from PolyMIM GmbH.

Total roughness $R_t$ ranges from 18.49 µm to 28.04 µm and arithmetic roughness $R_a$ ranges from 2.89 µm to 3.49 µm. The axial depth of cut $a_p$ has no significant impact on the surface roughness but the feed per tooth $f_z$, the radial depth of cut $a_e$ and the cutting speed $v_c$ needs to range respectively around 0.03 mm/tooth, inferior or equal to 0.5 mm and between 282.6 and 383 m/min to achieve the objective and not have a change of surface roughness class.

A stable operating point has been found exhibiting a $R_a < 2.89$ µm below the objective of 3.2 µm without surface roughness class change and providing an improved generated surface visual appearance.

The cutting forces are below 5 N for each set of cutting parameters which proves that there is no risk of detaching the part from the build plate during machining in the case of a hybrid process.

The proposed approach offers a promising solution for enhancing the quality of additively manufactured titanium parts, reducing post-processing requirements, and potentially optimizing production time. This study makes it possible to quickly obtain qualitatively and quantitatively (visual appearance of the generated machined surface) results since it makes it possible to find machining parameters providing parts with a surface roughness for industrial interest ($R_a < 3.2$ µm) superior to that obtained by additive manufacturing only.
<figure><img src="image_1799.png" /><figcaption>Figure 9: SEM images of the Ti6Al4V powder of PolyMIM pellets after debinding.</figcaption></figure>

Yoshida [11] did a study to evaluate the impact of grain size and heterogeneity on the surface roughness of sheet metals subjected to plane-strain stretching. According to this work, the surface roughness of parts made by the PolyMIM feedstock does not seem to be able to be significantly improved because of its particle size (average particle size approximatively equal to 20 µm) and because of the heterogeneity of the particle's sizes according to the SEM images presented on Fig. 9 which represents the particles of PolyMIM Ti6Al4V granule after debinding. The size of the particles is very heterogeneous between approx. 5 and 35 µm and this heterogeneity induces inequality on the surface of the printing and machining part.

The cutting forces were recorded for each cutting test. It's possible to see that all measured values remain under 5 N for each set of cutting parameters. The higher value is for the cutting test n°2 sequence 3 with a value of 4.9 N. The material seems to behave as a polymer more than metal which makes easier the machining operation as confirmed by [8] for ceramic loading.

## Conclusions

This study proves the feasibility of machining composite Ti6Al4V/Polymers parts made by Pellets Additive Manufacturing.

The tool dedicated to thermoplastics polymers seems to be adapted to machine the composite Ti6Al4V / Binder from PolyMIM GmbH.

Total roughness R~t~ ranges from 18.49 µm to 28.04 µm and arithmetic roughness R~a~ ranges from 2.89 µm to 3.49 µm. The axial depth of cut a~p~ has no significant impact on the surface roughness but the feed per tooth f~z~, the radial depth of cut a~e~ and the cutting speed v~c~ needs to range respectively around 0.03 mm/tooth, inferior or equal to 0.5 mm and between 282.6 and 383 m/min to achieve the objective and not have a change of surface roughness class.

A stable operating point has been found exhibiting a R~a~ < 2.89 µm below the objective of 3.2 µm without surface roughness class change and providing an improved generated surface visual appearance.

The cutting forces are below 5 N for each set of cutting parameters which proves that there is no risk of detaching the part from the build plate during machining in the case of a hybrid process.

The proposed approach offers a promising solution for enhancing the quality of additively manufactured titanium parts, reducing post-processing requirements, and potentially optimizing production time. This study makes it possible to quickly obtain qualitatively and quantitatively (visual appearance of the generated machined surface) results since it makes it possible to find machining parameters providing parts with a surface roughness for industrial interest (R~a~ < 3.2 µm) superior to that obtained by additive manufacturing only.
## Perspectives

Further research can follow this one by increase the number of tested set of parameters. Other tools can be used, as tool made especially for the titanium and its alloys or the broader applicability of this method to other materials.

## References

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[2] S.C. Altiparmak, V.A. Yardley, Z. Shi, J. Lin, Extrusion-based additive manufacturing technologies: State of the art and future perspectives. J. Manuf. Process. 83 (2022) 607-636. https://doi.org/10.1016/j.jmapro.2022.09.032

[3] R. Eickhoff, S. Antusch, D. Nötzel, T. Hanemann, New Partially Water-Soluble Feedstocks for Additive Manufacturing of Ti6Al4V Parts by Material Extrusion. Materials 16 (2023). https://doi.org/10.3390/ma16083162

[4] E.O. Ezugwu, Z. Wang, M. Materials Processing Technology Titanium alloys and their machinability a review. J. Mater. Process. Technol. 68 (1997) 262-274. https://doi.org/10.1016/S0924-0136(96)00030-1

[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. 21/3 (2015) 250-261. https://doi.org/10.1108/RPJ-02-2013-0017

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

[7] A. Demarbaix, M. Mulliez, E. Rivière-Lorphèvre, L. Spitaels, C. Duterte, N. Preux, F. Petit, F. Ducobu, Green Ceramic Machining: Determination of the Recommended Feed Rate for Y-TZP Milling. J. Compos. Sci. 5 (2021) 231. https://doi.org/10.3390/jcs5090231

[8] L. Spitaels, H. Dantinne, J. Bossu, E. Rivière-Lorphèvre, F. Ducobu, A Systematic Approach to Determine the Cutting Parameters of AM Green Zirconia in Finish Milling. J. Compos. Sci. 7 (2023) 112. https://doi.org/10.3390/jcs7030112

[9] L. Spitaels, E. Rivière-Lorphèvre, G. Martic, E. Juste, F. Ducobu, Machining of PAM green Y-TZP: Influence of build and in-plane directions on cutting forces and surface topography. Materials Research Proceedings 28 (2023) 1245-12. https://doi.org/10.21741/9781644902479-135
[10] NF E66-520-6; Working Zones of Cutting Tools. Couple Tool-Material. Association Française de Normalisation AFNOR : Paris, France, 1999.

[11] K. Yoshida, Effects of grain-scale heterogeneity on surface roughness and sheet metal necking. Int. J. Mech. Sci. 83 (2014) 48-56. https://doi.org/10.1016/j.ijmecsci.2014.03.018
## Perspectives

Further research can follow this one by increase the number of tested set of parameters. Other tools can be used, as tool made especially for the titanium and its alloys or the broader applicability of this method to other materials.

## References

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

[2] S.C. Altiparmak, V.A. Yardley, Z. Shi, J. Lin, Extrusion-based additive manufacturing technologies: State of the art and future perspectives. J. Manuf. Process. 83 (2022) 607-636. https://doi.org/10.1016/j.jmapro.2022.09.032

[3] R. Eickhoff, S. Antusch, D. Nötzel, T. Hanemann, New Partially Water-Soluble Feedstocks for Additive Manufacturing of Ti6Al4V Parts by Material Extrusion. Materials 16 (2023). https://doi.org/10.3390/ma16083162

[4] E.O. Ezugwu, Z. Wang, M. Materials Processing Technology Titanium alloys and their machinability a review. J. Mater. Process. Technol. 68 (1997) 262-274. https://doi.org/10.1016/S0924-0136(96)00030-1

[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. 21/3 (2015) 250-261. https://doi.org/10.1108/RPJ-02-2013-0017

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

[7] A. Demarbaix, M. Mulliez, E. Rivière-Lorphèvre, L. Spitaels, C. Duterte, N. Preux, F. Petit, F. Ducobu, Green Ceramic Machining: Determination of the Recommended Feed Rate for Y-TZP Milling. J. Compos. Sci. 5 (2021) 231. https://doi.org/10.3390/jcs5090231

[8] L. Spitaels, H. Dantinne, J. Bossu, E. Rivière-Lorphèvre, F. Ducobu, A Systematic Approach to Determine the Cutting Parameters of AM Green Zirconia in Finish Milling. J. Compos. Sci. 7 (2023) 112. https://doi.org/10.3390/jcs7030112

[9] L. Spitaels, E. Rivière-Lorphèvre, G. Martic, E. Juste, F. Ducobu, Machining of PAM green Y-TZP: Influence of build and in-plane directions on cutting forces and surface topography. Materials Research Proceedings 28 (2023) 1245-12. https://doi.org/10.21741/9781644902479-135
[10] NF E66-520-6; Working Zones of Cutting Tools. Couple Tool-Material. Association Française de Normalisation AFNOR : Paris, France, 1999.

[11] K. Yoshida, Effects of grain-scale heterogeneity on surface roughness and sheet metal necking. Int. J. Mech. Sci. 83 (2014) 48-56. https://doi.org/10.1016/j.ijmecsci.2014.03.018
# Effect of the hatch spacing in laser powder bed fusion on the AlSi<sub>7</sub>Mg aluminum alloy cutting forces and surface finish after turning

GHINATTI Edoardo<sup>1,a*</sup>, BERTOLINI Rachele<sup>1,b</sup>, GHIOTTI Andrea<sup>1,c</sup>, BRUSCHI Stefania<sup>1,d*</sup>

<sup>1</sup>Via Venezia 1, 35151, Padova, Italy

<sup>a</sup>edoardo.ghinatti@phd.unipd.it, <sup>b</sup>rachele.bertolini@unipd.it, <sup>c</sup>andrea.ghiotti@unipd.it, <sup>d</sup>stefania.bruschi@unipd.it

**Keywords:** Aluminum Alloy, Laser Powder Bed Fusion, Machining

**Abstract.** Applying additive manufacturing (AM) technologies in fabricating aluminum alloy automotive components can significantly enhance vehicle light-weighting through optimized design. By varying the AM process parameters, the resultant microstructure can be substantially altered, even post-heat treatment. Given that machining operations remain essential in many applications, understanding the impact of the AM parameters on the machining performances is vital. This paper explores how hatch spacing in laser powder bed fusion (LPBF) process affects the machinability of the AlSi<sub>7</sub>Mg aluminum alloy, in terms of cutting forces and machined surface roughness, providing a correlation with the microstructure induced by both the AM and heat treatment steps. In particular, the research reveals that the larger the hatch spacing the smoother the machined surface, underlining the critical role of this LPBF parameter in determining the final surface quality.

**Introduction**

Laser powder bed fusion (LPBF) is a widely utilized additive manufacturing (AM) technology used to 3D print complex-geometry shapes that are unachievable through conventional manufacturing methods [1]. This technique is particularly employed for printing aluminum alloys, especially AlSiMg alloys, which present an optimal blend of strength, corrosion resistance, and affordability [2].

The properties of components fabricated through LPBF are determined by the complex melting and solidification stages they undergo, which, in turn, are significantly affected by the chosen process parameters. To achieve a better quality in these components, it is crucial to understand how the various AM parameters, such as the hatch spacing, influence the melting and solidification patterns, which, finally, determine the part microstructural features.

In [3], the Authors developed an improved three-dimensional transient heat transfer model to study the effects of both the hatch spacing and laser spot overlapping on the temperature distribution and molten pool behavior when printing AlSi<sub>10</sub>Mg. It was observed that an increase in the hatch spacing leads to a decrease of the temperature on the powder bed. Notably, the hatch spacing influences not only the temperature of the melting pool, but also its distribution, as well as the molten pool's measured length, width, and depth. Finite element simulations showed that increasing the hatch spacing from 0.4 to 0.6 mm reduces the melt pool's width from 0.409 mm to 0.356 mm as well as its depth from 0.273 mm to 0.240 mm. In [4], the outcomes of the LPBF process carried out on the AlMgScZr alloy using different hatch distances were extensively studied, focusing on various aspects such as densification behavior, melting mode, crystallographic orientation, evolution of the grain boundary angle misorientation, and part tensile properties. The findings indicate that the densification improved at decreasing hatch distance, while the grain size increased due to the elevated cooling rate.

<figure><img src="image_1802.png" /></figure>
# Effect of the hatch spacing in laser powder bed fusion on the AlSi<sub>7</sub>Mg aluminum alloy cutting forces and surface finish after turning

GHINATTI Edoardo<sup>1,a*</sup>, BERTOLINI Rachele<sup>1,b</sup>, GHIOTTI Andrea<sup>1,c</sup>, BRUSCHI Stefania<sup>1,d*</sup>

<sup>1</sup>Via Venezia 1, 35151, Padova, Italy

<sup>a</sup>edoardo.ghinatti@phd.unipd.it, <sup>b</sup>rachele.bertolini@unipd.it, <sup>c</sup>andrea.ghiotti@unipd.it, <sup>d</sup>stefania.bruschi@unipd.it

**Keywords:** Aluminum Alloy, Laser Powder Bed Fusion, Machining

**Abstract.** Applying additive manufacturing (AM) technologies in fabricating aluminum alloy automotive components can significantly enhance vehicle light-weighting through optimized design. By varying the AM process parameters, the resultant microstructure can be substantially altered, even post-heat treatment. Given that machining operations remain essential in many applications, understanding the impact of the AM parameters on the machining performances is vital. This paper explores how hatch spacing in laser powder bed fusion (LPBF) process affects the machinability of the AlSi<sub>7</sub>Mg aluminum alloy, in terms of cutting forces and machined surface roughness, providing a correlation with the microstructure induced by both the AM and heat treatment steps. In particular, the research reveals that the larger the hatch spacing the smoother the machined surface, underlining the critical role of this LPBF parameter in determining the final surface quality.

**Introduction**

Laser powder bed fusion (LPBF) is a widely utilized additive manufacturing (AM) technology used to 3D print complex-geometry shapes that are unachievable through conventional manufacturing methods [1]. This technique is particularly employed for printing aluminum alloys, especially AlSiMg alloys, which present an optimal blend of strength, corrosion resistance, and affordability [2].

The properties of components fabricated through LPBF are determined by the complex melting and solidification stages they undergo, which, in turn, are significantly affected by the chosen process parameters. To achieve a better quality in these components, it is crucial to understand how the various AM parameters, such as the hatch spacing, influence the melting and solidification patterns, which, finally, determine the part microstructural features.

In [3], the Authors developed an improved three-dimensional transient heat transfer model to study the effects of both the hatch spacing and laser spot overlapping on the temperature distribution and molten pool behavior when printing AlSi<sub>10</sub>Mg. It was observed that an increase in the hatch spacing leads to a decrease of the temperature on the powder bed. Notably, the hatch spacing influences not only the temperature of the melting pool, but also its distribution, as well as the molten pool's measured length, width, and depth. Finite element simulations showed that increasing the hatch spacing from 0.4 to 0.6 mm reduces the melt pool's width from 0.409 mm to 0.356 mm as well as its depth from 0.273 mm to 0.240 mm. In [4], the outcomes of the LPBF process carried out on the AlMgScZr alloy using different hatch distances were extensively studied, focusing on various aspects such as densification behavior, melting mode, crystallographic orientation, evolution of the grain boundary angle misorientation, and part tensile properties. The findings indicate that the densification improved at decreasing hatch distance, while the grain size increased due to the elevated cooling rate.

<figure><img src="image_1800.png" /></figure>
Since first introduction of the LPFB technology, its process parameters have been extensively varied to optimize the printed part properties and thus increase the feasibility of its employment in different sectors. However, there remains a limited number of studies on how the AM-induced microstructure can affect the AM part machinability.

In a previous work by the Authors [5], the effect of the microstructural anisotropy induced by LPBF carried out using various scanning strategies was evaluated when turning Ti6Al4V LPBF parts. Specifically, under constant laser power and bed fusion process parameters, different scanning strategies, namely stripes and chessboard patterns, were employed. The results indicated that altering the scanning strategy significantly affects the part mean grain size. For instance, the chessboard strategy resulted in a 32% reduction in grain size compared to the stripes one. Consequently, the micro-hardness of parts produced with the chessboard strategy was 17% higher than those produced with the stripes one. These variations in scanning strategy also influenced the surface finish of the machined samples: the stripes scanning strategy led to rougher and more uneven surfaces compared to the chessboard one.

The influence on tool wear and surface topography during milling LPBF Ti6Al4V parts at various growth inclination angles was examined in [6]. Tool wear and surface roughness were observed to decrease with a reduction in the inclination angle. This decrease was attributed to the influence of weak discontinuity zones along the build direction (BD), which facilitate the material removal when favorably oriented relative to the cutting tool's edge.

In [7], the BD significant impact on the machinability of AM stainless steels was demonstrated. AISI 316L stainless steel was fabricated using a high-power laser with a large spot size along various build directions. The machinability during milling was evaluated based on the cutting forces, surface finish, and tool wear at different cutting speeds. Post-additive manufacturing characterization revealed that the cross-section at 0° with respect to the BD had higher hardness than the vertical section at 90°. This resulted in higher cutting forces of the samples at 0° with respect to the BD compared to those at 90°, highlighting the anisotropy in machinability due to the high-power direct laser deposition process. This anisotropy also affected the machined surface finish, being the samples fabricated at 0° with respect to the BD rougher than those at 90°.

As demonstrated in the previously mentioned publications, while some researchers have explored the machinability of AM parts, none have specifically addressed the impact that AM parameters, which can induce microstructural anisotropy, can have on turning outcomes. To this aim, the present paper evaluates the impact of hatch spacing in LPBF on the machinability of the AlSi7Mg aluminum alloy. For this purpose, a series of samples were fabricated at varying hatch spacing using LPBF and then subjected to heat treatment. Subsequent machining tests were conducted using fixed cutting parameters. The forces exerted during these tests were recorded, and the surface finish was evaluated. The results revealed a correlation between the alloy machinability and hatch spacing, indicating that smoother surfaces are achievable when using larger hatch spacings.

## Materials and methods

LPBF of AlSi7Mg Aluminum Alloy. In this study, AlSi7Mg is the addressed aluminum alloy, in form of powders produced by gas atomization and subsequently printed using the LPBF process. The spherical powders, ranging in size from 20-63 µm, were supplied by Eckart™, Altana. The nominal chemical composition of the AlSi7Mg powders is detailed in Table 1.
Since first introduction of the LPFB technology, its process parameters have been extensively varied to optimize the printed part properties and thus increase the feasibility of its employment in different sectors. However, there remains a limited number of studies on how the AM-induced microstructure can affect the AM part machinability.

In a previous work by the Authors [5], the effect of the microstructural anisotropy induced by LPBF carried out using various scanning strategies was evaluated when turning Ti6Al4V LPBF parts. Specifically, under constant laser power and bed fusion process parameters, different scanning strategies, namely stripes and chessboard patterns, were employed. The results indicated that altering the scanning strategy significantly affects the part mean grain size. For instance, the chessboard strategy resulted in a 32% reduction in grain size compared to the stripes one. Consequently, the micro-hardness of parts produced with the chessboard strategy was 17% higher than those produced with the stripes one. These variations in scanning strategy also influenced the surface finish of the machined samples: the stripes scanning strategy led to rougher and more uneven surfaces compared to the chessboard one.

The influence on tool wear and surface topography during milling LPBF Ti6Al4V parts at various growth inclination angles was examined in [6]. Tool wear and surface roughness were observed to decrease with a reduction in the inclination angle. This decrease was attributed to the influence of weak discontinuity zones along the build direction (BD), which facilitate the material removal when favorably oriented relative to the cutting tool's edge.

In [7], the BD significant impact on the machinability of AM stainless steels was demonstrated. AISI 316L stainless steel was fabricated using a high-power laser with a large spot size along various build directions. The machinability during milling was evaluated based on the cutting forces, surface finish, and tool wear at different cutting speeds. Post-additive manufacturing characterization revealed that the cross-section at 0° with respect to the BD had higher hardness than the vertical section at 90°. This resulted in higher cutting forces of the samples at 0° with respect to the BD compared to those at 90°, highlighting the anisotropy in machinability due to the high-power direct laser deposition process. This anisotropy also affected the machined surface finish, being the samples fabricated at 0° with respect to the BD rougher than those at 90°.

As demonstrated in the previously mentioned publications, while some researchers have explored the machinability of AM parts, none have specifically addressed the impact that AM parameters, which can induce microstructural anisotropy, can have on turning outcomes. To this aim, the present paper evaluates the impact of hatch spacing in LPBF on the machinability of the AlSi7Mg aluminum alloy. For this purpose, a series of samples were fabricated at varying hatch spacing using LPBF and then subjected to heat treatment. Subsequent machining tests were conducted using fixed cutting parameters. The forces exerted during these tests were recorded, and the surface finish was evaluated. The results revealed a correlation between the alloy machinability and hatch spacing, indicating that smoother surfaces are achievable when using larger hatch spacings.

## Materials and methods

LPBF of AlSi7Mg Aluminum Alloy. In this study, AlSi7Mg is the addressed aluminum alloy, in form of powders produced by gas atomization and subsequently printed using the LPBF process. The spherical powders, ranging in size from 20-63 µm, were supplied by Eckart™, Altana. The nominal chemical composition of the AlSi7Mg powders is detailed in Table 1.
<table><caption>Table 1: AlSi<sub>7</sub>Mg powder chemical composition.</caption><thead><tr><th>Element</th><th>Si</th><th>Mg</th><th>Fe</th><th>Cu</th><th>Zn</th><th>Ti</th><th>Mn</th><th>Al</th></tr></thead><tbody><tr><th>Value [%]</th><td>6.86</td><td>0.58</td><td>0.14</td><td>0.01</td><td>0.003</td><td>0.14</td><td>0.0004</td><td>Bal.</td></tr></tbody></table>

The AlSi<sub>7</sub>Mg samples were printed using a Sisma™ MYSINT100 LPBF machine, equipped with a 200 W Yb fiber laser and a 55 µm spot diameter. The LPBF process was conducted in an inert atmosphere, specifically by insufflating argon, ensuring an oxygen level below 100 ppm in the working chamber. A preliminary experimental campaign was carried out to optimize the LPBF process parameters. For doing that, cubic samples of 10 mm side were printed at varying laser power, scanning speed, and hatch spacing, while energy density and layer thickness were set constant. The employed scanning strategy was a 3 mm x 3 mm chessboard pattern with layers inclined at 67° to each other, as per [8]. For each parameter combination, the samples density and porosities were measured. The density was assessed using the Archimedes' principle with a high-precision KERN™ ABT 120-5DM balance, while the porosities were evaluated through optical microscopy. The samples were cold-mounted, ground with up to 4000 grit-size abrasive papers, and polished using a synthetic cloth with 6, 3, and 1 µm diamond suspensions. After being cleaned with ethanol and dried, the polished surfaces were examined using a Keyence™ VHX-7000 optical microscope at high magnification. The porosity analysis was performed using a dedicated MATLAB™ script. The primary defects that were identified were lacks of fusion and gas entrapment porosities. The former are due to inadequate energy inputs, while the latter result from excessive energy leading to moisture and gas evaporation.

The optimized process parameters that lead to maximum densification and induce the minimum percentage of porosities were established as power (P) = 170 W, scanning speed (V) = 1056 mm/s, hatch spacing (h) = 150 µm, and layer thickness (L<sub>t</sub>) = 30 µm. By using these parameters, three cylindrical samples (20 mm diameter and 70 mm height) were printed with varying hatch spacings, namely h = 125 µm, h = 150 µm, h = 175 µm. Afterwards, these cylinders underwent a T6 heat treatment, which includes solubilization at 535°C for 3 hours and artificial aging at 150°C for 6 hours, in order to relieve the LPBF-induced residual stresses and promote material strengthening through precipitate formation. The microstructure of the LPBF samples, both pre- and post-heat treatment, was analyzed using the aforementioned preparation method and Keller's etching solution. Differential scanning calorimetry (DSC) analyses with a DSC Q200™ calorimeter were conducted to assess differences among the samples, heating them from room temperature to 400°C at 10°C/min.

<figure><img src="image_1803.png" /><figcaption>Fig. 1: a) As-built LPBF AlSi<sub>7</sub>Mg cylinder and b) machining setup with the AlSi<sub>7</sub>Mg cylinder mounted.</figcaption></figure>
<table><caption>Table 1: AlSi<sub>7</sub>Mg powder chemical composition.</caption><thead><tr><th>Element</th><th>Si</th><th>Mg</th><th>Fe</th><th>Cu</th><th>Zn</th><th>Ti</th><th>Mn</th><th>Al</th></tr></thead><tbody><tr><th>Value [%]</th><td>6.86</td><td>0.58</td><td>0.14</td><td>0.01</td><td>0.003</td><td>0.14</td><td>0.0004</td><td>Bal.</td></tr></tbody></table>

The AlSi<sub>7</sub>Mg samples were printed using a Sisma™ MYSINT100 LPBF machine, equipped with a 200 W Yb fiber laser and a 55 µm spot diameter. The LPBF process was conducted in an inert atmosphere, specifically by insufflating argon, ensuring an oxygen level below 100 ppm in the working chamber. A preliminary experimental campaign was carried out to optimize the LPBF process parameters. For doing that, cubic samples of 10 mm side were printed at varying laser power, scanning speed, and hatch spacing, while energy density and layer thickness were set constant. The employed scanning strategy was a 3 mm x 3 mm chessboard pattern with layers inclined at 67° to each other, as per [8]. For each parameter combination, the samples density and porosities were measured. The density was assessed using the Archimedes' principle with a high-precision KERN™ ABT 120-5DM balance, while the porosities were evaluated through optical microscopy. The samples were cold-mounted, ground with up to 4000 grit-size abrasive papers, and polished using a synthetic cloth with 6, 3, and 1 µm diamond suspensions. After being cleaned with ethanol and dried, the polished surfaces were examined using a Keyence™ VHX-7000 optical microscope at high magnification. The porosity analysis was performed using a dedicated MATLAB™ script. The primary defects that were identified were lacks of fusion and gas entrapment porosities. The former are due to inadequate energy inputs, while the latter result from excessive energy leading to moisture and gas evaporation.

The optimized process parameters that lead to maximum densification and induce the minimum percentage of porosities were established as power (P) = 170 W, scanning speed (V) = 1056 mm/s, hatch spacing (h) = 150 µm, and layer thickness (L<sub>t</sub>) = 30 µm. By using these parameters, three cylindrical samples (20 mm diameter and 70 mm height) were printed with varying hatch spacings, namely h = 125 µm, h = 150 µm, h = 175 µm. Afterwards, these cylinders underwent a T6 heat treatment, which includes solubilization at 535°C for 3 hours and artificial aging at 150°C for 6 hours, in order to relieve the LPBF-induced residual stresses and promote material strengthening through precipitate formation. The microstructure of the LPBF samples, both pre- and post-heat treatment, was analyzed using the aforementioned preparation method and Keller's etching solution. Differential scanning calorimetry (DSC) analyses with a DSC Q200™ calorimeter were conducted to assess differences among the samples, heating them from room temperature to 400°C at 10°C/min.

<figure><img src="image_1801.png" /><figcaption>Fig. 1: a) As-built LPBF AlSi<sub>7</sub>Mg cylinder and b) machining setup with the AlSi<sub>7</sub>Mg cylinder mounted.</figcaption></figure>
Machining of LPBF AlSi<sub>7</sub>Mg Samples. The LPBF cylinders were utilized as workpieces in the turning trials, conducted on a Mori Seiki™ NL1500TM CNC lathe using conventional flooding lubrication. The employed cutting tool was the VCEX 11 03 01L-F 1125 turning inserts, procured from Sandvik Coromant™, whose main characteristics are detailed in Table 2.

<table><caption>Table 2. Turning tool specifications.</caption><thead><tr><td>Parameter</td><td>Value</td></tr></thead><tbody><tr><td>Substrate</td><td>Tungsten carbide</td></tr><tr><td>Coating</td><td>PVD TiAlN+TiAlN</td></tr><tr><td>Nose radius [mm]</td><td>0.1</td></tr><tr><td>Rake angle [°]</td><td>5°30′</td></tr><tr><td>Clearance angle [°]</td><td>7°</td></tr><tr><td>Cutting edge radius [mm]</td><td>0.01</td></tr></tbody></table>

The insert was attached to a SVJBL 2020 K11 - B1 tool holder, which was then secured to a 9129AA Kistler™ three-component piezoelectric dynamometer. As shown in Fig. 1, the cutting force (F<sub>c</sub>), thrust force (F<sub>t</sub>), and feed force (F<sub>f</sub>) were measured during the tests. The force signals were captured at a sampling rate of 2500 Hz and processed using a 2nd-order low-pass filter with a 25 Hz cutoff frequency. The turning test parameters were consistent with the tool manufacturer's guidelines for finishing aluminum alloys, namely depth of cut (d) = 0.25 mm, cutting speed (V<sub>c</sub>) = 200 m/min, and feed (f) = 0.03 mm/rev. Each experiment was conducted three times over a cutting length of 5 mm to ensure the reproducibility of the results.

Surface Finish Of The Machined LPBF AlSi<sub>7</sub>Mg Samples. The topographies of the machined surfaces were acquired by using the Sensofar Plu Neox™ 3D optical profiler with a 20x Nikon™ confocal objective. The roughness parameters were computed by using the SensoVIEW™ software using the filtering procedure indicated in [9], namely a form removal process to remove the cylindrical form after turning, followed by a λs filter equal to 2.5 µm, cut-off length λc equal to 0.8 mm, and sampling length of 4 mm. The arithmetic mean height (S<sub>a</sub>), root mean square height (S<sub>q</sub>), and total height (S<sub>z</sub>) were the surface roughness parameters considered in the present study. While S<sub>a</sub> is the conventional surface roughness parameter, S<sub>q</sub> offers additional insight by quantifying the variance in surface height deviations, thereby providing a more comprehensive understanding of the surface texture and quality. S<sub>z</sub>, on the other hand, measures the total extent of surface irregularities or deviations from a reference plane over a specific length.

## Results and discussion

Microstructure of The LPBF AlSi<sub>7</sub>Mg Samples. Fig. 2 displays the microstructure of the samples along the build direction and the scanning direction, also referred to as the x-z plane in Fig. 1, in their as-built state. After LPBF, macroscopic melting pools are clearly discernible in the low-magnification images. Each pool exhibits a fine microstructure in its central part and a coarser one at its boundaries. Notably, columnar grains along the build direction arise from localized heat dissipation by the laser. A notable variation in the melting pool size, related to hatch spacing, is evident. Reduced hatch spacing results in denser melting pools along the x direction, but fewer vertically, namely along the build direction, while the inverse is true for larger hatch spacings. This variation induces a microstructural anisotropy in the material, with the pool depths decreasing and widths increasing at increasing hatch spacings. Fig. 2 uses black and white arrows to distinguish between the pool depths and widths, respectively. This microstructural distribution is consistent with that found in [10], and is attributed to the varying energy distribution of the laser at different hatch spacings.
Machining of LPBF AlSi<sub>7</sub>Mg Samples. The LPBF cylinders were utilized as workpieces in the turning trials, conducted on a Mori Seiki™ NL1500TM CNC lathe using conventional flooding lubrication. The employed cutting tool was the VCEX 11 03 01L-F 1125 turning inserts, procured from Sandvik Coromant™, whose main characteristics are detailed in Table 2.

<table><caption>Table 2. Turning tool specifications.</caption><thead><tr><td>Parameter</td><td>Value</td></tr></thead><tbody><tr><td>Substrate</td><td>Tungsten carbide</td></tr><tr><td>Coating</td><td>PVD TiAlN+TiAlN</td></tr><tr><td>Nose radius [mm]</td><td>0.1</td></tr><tr><td>Rake angle [°]</td><td>5°30′</td></tr><tr><td>Clearance angle [°]</td><td>7°</td></tr><tr><td>Cutting edge radius [mm]</td><td>0.01</td></tr></tbody></table>

The insert was attached to a SVJBL 2020 K11 - B1 tool holder, which was then secured to a 9129AA Kistler™ three-component piezoelectric dynamometer. As shown in Fig. 1, the cutting force (F<sub>c</sub>), thrust force (F<sub>t</sub>), and feed force (F<sub>f</sub>) were measured during the tests. The force signals were captured at a sampling rate of 2500 Hz and processed using a 2nd-order low-pass filter with a 25 Hz cutoff frequency. The turning test parameters were consistent with the tool manufacturer's guidelines for finishing aluminum alloys, namely depth of cut (d) = 0.25 mm, cutting speed (V<sub>c</sub>) = 200 m/min, and feed (f) = 0.03 mm/rev. Each experiment was conducted three times over a cutting length of 5 mm to ensure the reproducibility of the results.

Surface Finish Of The Machined LPBF AlSi<sub>7</sub>Mg Samples. The topographies of the machined surfaces were acquired by using the Sensofar Plu Neox™ 3D optical profiler with a 20x Nikon™ confocal objective. The roughness parameters were computed by using the SensoVIEW™ software using the filtering procedure indicated in [9], namely a form removal process to remove the cylindrical form after turning, followed by a λs filter equal to 2.5 µm, cut-off length λc equal to 0.8 mm, and sampling length of 4 mm. The arithmetic mean height (S<sub>a</sub>), root mean square height (S<sub>q</sub>), and total height (S<sub>z</sub>) were the surface roughness parameters considered in the present study. While S<sub>a</sub> is the conventional surface roughness parameter, S<sub>q</sub> offers additional insight by quantifying the variance in surface height deviations, thereby providing a more comprehensive understanding of the surface texture and quality. S<sub>z</sub>, on the other hand, measures the total extent of surface irregularities or deviations from a reference plane over a specific length.

## Results and discussion

Microstructure of The LPBF AlSi<sub>7</sub>Mg Samples. Fig. 2 displays the microstructure of the samples along the build direction and the scanning direction, also referred to as the x-z plane in Fig. 1, in their as-built state. After LPBF, macroscopic melting pools are clearly discernible in the low-magnification images. Each pool exhibits a fine microstructure in its central part and a coarser one at its boundaries. Notably, columnar grains along the build direction arise from localized heat dissipation by the laser. A notable variation in the melting pool size, related to hatch spacing, is evident. Reduced hatch spacing results in denser melting pools along the x direction, but fewer vertically, namely along the build direction, while the inverse is true for larger hatch spacings. This variation induces a microstructural anisotropy in the material, with the pool depths decreasing and widths increasing at increasing hatch spacings. Fig. 2 uses black and white arrows to distinguish between the pool depths and widths, respectively. This microstructural distribution is consistent with that found in [10], and is attributed to the varying energy distribution of the laser at different hatch spacings.
<figure><img src="image_1804.png" /><figcaption>Fig. 2: Microstructure of the AlSi7Mg LPBF samples along the building direction before heat treatment at varying hatch spacing.</figcaption></figure>

As anticipated, along the x direction, a lower hatch spacing results in narrower melting pools due to increased overlap between adjacent scan lines. In the build direction, a lower hatch spacing delivers more energy to the material, raising the local temperature and reducing viscosity. This allows the material to spread more evenly downwards before solidifying. Conversely, at higher hatch spacings, as noted in [11], the opposite effect is observed.

The heat distribution variations caused by different hatch spacings also impact the cooling rates and solidification patterns, subsequently influencing the microstructure of the AlSi7Mg LPBF samples. Observing a single melting pool in the magnified image of Fig. 2, it is evident that samples printed with the widest hatch spacing exhibit the coarsest grains. This is likely due to the slowest cooling rate at the highest hatch spacing, which facilitates grain growth.

After heat treatment, the microstructure becomes more homogeneous and less anisotropic, regardless of the hatch spacing, as shown in Fig. 3. The solubilization step enhances silicon precipitate formation, while aging promotes the development of nanometric-sized silicon particles [12], making the melting pools less distinguishable.

DSC analyses were conducted on the LPBF AlSi7Mg samples at h = 150 µm in the as-built condition, and after heat treatment for every value of hatch spacing to assess microstructural differences related to the precipitates, with results displayed in Fig. 4. All the DSC curves show an exothermic peak around 200°C-240°C, more pronounced for the sample in the as-built condition. The high enthalpy of this peak in the as-built alloy, attributed to Si precipitation, is due to the supersaturated state from the LPBF process and is drastically reduced after heat treatment [12].

This study notes a lower specific peak temperature compared to previous reports [12], possibly due to AM factors like residual stresses, porosity, and microstructural anisotropy, influencing the
<figure><img src="image_1802.png" /><figcaption>Fig. 2: Microstructure of the AlSi7Mg LPBF samples along the building direction before heat treatment at varying hatch spacing.</figcaption></figure>

As anticipated, along the x direction, a lower hatch spacing results in narrower melting pools due to increased overlap between adjacent scan lines. In the build direction, a lower hatch spacing delivers more energy to the material, raising the local temperature and reducing viscosity. This allows the material to spread more evenly downwards before solidifying. Conversely, at higher hatch spacings, as noted in [11], the opposite effect is observed.

The heat distribution variations caused by different hatch spacings also impact the cooling rates and solidification patterns, subsequently influencing the microstructure of the AlSi7Mg LPBF samples. Observing a single melting pool in the magnified image of Fig. 2, it is evident that samples printed with the widest hatch spacing exhibit the coarsest grains. This is likely due to the slowest cooling rate at the highest hatch spacing, which facilitates grain growth.

After heat treatment, the microstructure becomes more homogeneous and less anisotropic, regardless of the hatch spacing, as shown in Fig. 3. The solubilization step enhances silicon precipitate formation, while aging promotes the development of nanometric-sized silicon particles [12], making the melting pools less distinguishable.

DSC analyses were conducted on the LPBF AlSi7Mg samples at h = 150 µm in the as-built condition, and after heat treatment for every value of hatch spacing to assess microstructural differences related to the precipitates, with results displayed in Fig. 4. All the DSC curves show an exothermic peak around 200°C-240°C, more pronounced for the sample in the as-built condition. The high enthalpy of this peak in the as-built alloy, attributed to Si precipitation, is due to the supersaturated state from the LPBF process and is drastically reduced after heat treatment [12].

This study notes a lower specific peak temperature compared to previous reports [12], possibly due to AM factors like residual stresses, porosity, and microstructural anisotropy, influencing the
DSC response and causing a peak shift after the heat treatment. Both the peak position and area vary depending on the hatch spacing used in LPBF, as indicated in Table 3. The peak temperature represents one of the maximum rates of formation or dissolution reactions, while the area under the peak correlates with the reaction heat and the phase volume fraction formed or dissolved, determined qualitatively by its enthalpy value [13]. Furthermore, the precipitate formation is closely linked to grain boundary density, as boundaries are nucleation sites due to their higher energy state. Thus, more intense precipitation occurs in samples with lower hatch spacing, evident in their finer microstructure and larger peak areas compared to those with $h = 175$ µm, indicating more material transforming.

<figure><img src="image_1805.png" /><figcaption>Fig. 3: Microstructure of the LPBF AlSi7Mg samples fabricated with $h = 150$ µm along the building direction after heat treatment.</figcaption></figure>

<figure><img src="image_1806.png" /><figcaption>Fig. 4: DSC curves of the LPBF AlSi7Mg samples before and after heat treatment at varying hatch spacing.</figcaption></figure>

<table><caption>Table 3: Peak temperature and phase transformation enthalpy derived from Fig. 4.</caption><thead><tr><th>Condition</th><th>As-built</th><th colspan="3">Heat treated</th></tr></thead><tbody><tr><td>Hatch spacing [µm]</td><td>150</td><td>125</td><td>150</td><td>175</td></tr><tr><td>T<sub>peak</sub> [°C]</td><td>216</td><td>238</td><td>234</td><td>230</td></tr><tr><td>Phase transformation enthalpy [J/g]</td><td>15.23</td><td>2.22</td><td>2.19</td><td>2.14</td></tr></tbody></table>
DSC response and causing a peak shift after the heat treatment. Both the peak position and area vary depending on the hatch spacing used in LPBF, as indicated in Table 3. The peak temperature represents one of the maximum rates of formation or dissolution reactions, while the area under the peak correlates with the reaction heat and the phase volume fraction formed or dissolved, determined qualitatively by its enthalpy value [13]. Furthermore, the precipitate formation is closely linked to grain boundary density, as boundaries are nucleation sites due to their higher energy state. Thus, more intense precipitation occurs in samples with lower hatch spacing, evident in their finer microstructure and larger peak areas compared to those with h = 175 µm, indicating more material transforming.

<figure><img src="image_1803.png" /><figcaption>Fig. 3: Microstructure of the LPBF AlSi7Mg samples fabricated with h = 150 µm along the building direction after heat treatment.</figcaption></figure>

<figure><img src="image_1804.png" /><figcaption>Fig. 4: DSC curves of the LPBF AlSi7Mg samples before and after heat treatment at varying hatch spacing.</figcaption></figure>

<table><caption>Table 3: Peak temperature and phase transformation enthalpy derived from Fig. 4.</caption><thead><tr><th>Condition</th><th>As-built</th><th colspan="3">Heat treated</th></tr></thead><tbody><tr><td>Hatch spacing [µm]</td><td>150</td><td>125</td><td>150</td><td>175</td></tr><tr><td>T<sub>peak</sub> [°C]</td><td>216</td><td>238</td><td>234</td><td>230</td></tr><tr><td>Phase transformation enthalpy [J/g]</td><td>15.23</td><td>2.22</td><td>2.19</td><td>2.14</td></tr></tbody></table>
Turning Forces. Fig. 5a displays an example of forces acquired when turning the LPBF AlSi7Mg samples manufactured with $h = 125$ µm, while Fig. 5b shows the mean values derived from Fig. 5a at varying hatch spacing. Considering the standard deviation, hatch spacing does not significantly impact the average force values. However, a higher variability in the cutting force values can be noted at decreasing hatch spacing. Given that the turning process is carried out at fixed parameters, this can be ascribed to the more heterogeneous microstructure obtained when using a lower hatch spacing. Large precipitates create indeed discontinuities for the cutting tool, resulting in increased force variability. In addition, higher precipitate density at lower hatch spacings contributes to increased variability in forces during cutting.

<figure><img src="image_1807.png" /><figcaption>Fig. 5: (a) Forces recorded during turning of the LPBF AlSi7Mg samples fabricated with $h = 125$ µm, and (b) mean values computed from (a).</figcaption></figure>

Surface Finish of The Machined LPBF AlSi7Mg Samples. Fig. 6a presents a comparison of the machined surface topography of the heat-treated LPBF AlSi7Mg samples before and after filtering fabricated using the intermediate hatch spacing. Fig. 6b shows the dependence of the surface roughness parameters on the hatch spacing. Generally, the larger the hatch spacing the better the surface quality. Specifically, the $S_a$, $S_q$, and $S_z$ roughness values decrease by 32%, 34%, and 63% from the smallest to the largest hatch spacings, respectively. Consistent with the force results of Fig. 5, the samples fabricated with the smallest hatch spacing exhibit the greatest variability. A correlation with the microstructure is again evident: microstructures characterized by coarser grains and fewer precipitates are more easily machined, facilitating the attainment of smoother surfaces.
Turning Forces. Fig. 5a displays an example of forces acquired when turning the LPBF AlSi7Mg samples manufactured with $h = 125$ µm, while Fig. 5b shows the mean values derived from Fig. 5a at varying hatch spacing. Considering the standard deviation, hatch spacing does not significantly impact the average force values. However, a higher variability in the cutting force values can be noted at decreasing hatch spacing. Given that the turning process is carried out at fixed parameters, this can be ascribed to the more heterogeneous microstructure obtained when using a lower hatch spacing. Large precipitates create indeed discontinuities for the cutting tool, resulting in increased force variability. In addition, higher precipitate density at lower hatch spacings contributes to increased variability in forces during cutting.

<figure><img src="image_1805.png" /><figcaption>Fig. 5: (a) Forces recorded during turning of the LPBF AlSi7Mg samples fabricated with $h = 125$ µm, and (b) mean values computed from (a).</figcaption></figure>

Surface Finish of The Machined LPBF AlSi7Mg Samples. Fig. 6a presents a comparison of the machined surface topography of the heat-treated LPBF AlSi7Mg samples before and after filtering fabricated using the intermediate hatch spacing. Fig. 6b shows the dependence of the surface roughness parameters on the hatch spacing. Generally, the larger the hatch spacing the better the surface quality. Specifically, the $S_a$, $S_q$, and $S_z$ roughness values decrease by 32%, 34%, and 63% from the smallest to the largest hatch spacings, respectively. Consistent with the force results of Fig. 5, the samples fabricated with the smallest hatch spacing exhibit the greatest variability. A correlation with the microstructure is again evident: microstructures characterized by coarser grains and fewer precipitates are more easily machined, facilitating the attainment of smoother surfaces.
<figure><img src="image_1808.png" /><figcaption>Fig. 6: (a) Surface topography of the LPBF AlSi7Mg samples fabricated with $h = 150$ µm before and after application of the filtering procedure, and b) surface roughness parameters at varying hatch spacing.</figcaption></figure>

## Conclusions

The present work focuses on the influence of hatch spacing varied during the LPBF process on the machinability of the AlSi7Mg alloy, analyzing its effects on the cutting forces and machined surface roughness. The main results can be summarized as follows:

* A reduced hatch spacing results in denser melting pools along the x-direction, but fewer along the build direction; the opposite occurs with the largest hatch spacing.

* Larger hatch spacings result in samples with a coarser microstructure.

* Heat treatment significantly reduces anisotropy in the microstructure induced by the LPBF process.

* The density of the silicon precipitate density after heat treatment varies with the hatch spacing.

* Hatch spacing does not affect the average cutting force values, but influences their variability.

* Larger hatch spacings lead to smoother and more consistent surfaces, which is due to the different microstructure of the alloy, having a direct effect on the machining performances.

Future research endeavors will focus on assessing how the thickness of layers affects the machinability of AlSi7Mg fabricated through LPBF. A thorough examination of the microstructure and distribution of precipitates will be conducted and correlated with the material's characteristics during cutting. It is anticipated that this will yield an optimal window of additive manufacturing process parameters, facilitating the production of dependable components that are easily machinable.

## Acknowledgments

This work was supported by the European Union - NextGenerationEU (National Sustainable Mobility Center CN00000023, Italian Ministry of University and Research Decree n. 1033 - 17/06/2022, Spoke 11 - Innovative Materials & Lightweighting).
<figure><img src="image_1806.png" /><figcaption>Fig. 6: (a) Surface topography of the LPBF AlSi7Mg samples fabricated with $h = 150$ µm before and after application of the filtering procedure, and (b) surface roughness parameters at varying hatch spacing.</figcaption></figure>

## Conclusions

The present work focuses on the influence of hatch spacing varied during the LPBF process on the machinability of the AlSi7Mg alloy, analyzing its effects on the cutting forces and machined surface roughness. The main results can be summarized as follows:

* A reduced hatch spacing results in denser melting pools along the x-direction, but fewer along the build direction; the opposite occurs with the largest hatch spacing.

* Larger hatch spacings result in samples with a coarser microstructure.

* Heat treatment significantly reduces anisotropy in the microstructure induced by the LPBF process.

* The density of the silicon precipitate density after heat treatment varies with the hatch spacing.

* Hatch spacing does not affect the average cutting force values, but influences their variability.

* Larger hatch spacings lead to smoother and more consistent surfaces, which is due to the different microstructure of the alloy, having a direct effect on the machining performances.

Future research endeavors will focus on assessing how the thickness of layers affects the machinability of AlSi7Mg fabricated through LPBF. A thorough examination of the microstructure and distribution of precipitates will be conducted and correlated with the material's characteristics during cutting. It is anticipated that this will yield an optimal window of additive manufacturing process parameters, facilitating the production of dependable components that are easily machinable.

## Acknowledgments

This work was supported by the European Union - NextGenerationEU (National Sustainable Mobility Center CN00000023, Italian Ministry of University and Research Decree n. 1033 - 17/06/2022, Spoke 11 - Innovative Materials & Lightweighting).
## References

[1] Wa. Abd-Elaziem, S. Elkatatny, A.E. Abd-Elaziem, M. Khedr, M. A. A. El-baky, M. A. Hassan, M. Abu-Okail, M. Mohammed, A. Järvenpää, T. Allam, A. Hamada, On the current research progress of metallic materials fabricated by laser powder bed fusion process: a review, Journal of Material Research and Technology 20 (2022) 681-707. https://doi.org/10.1016/j.jmrt.2022.07.085

[2] H. Hyer et al., H. Hyer, L. Zhou, S. Park, G. Gottsfritz, G. Benson, B. Tolentino, B. McWilliams, K. Cho, Y. Sohn, Understanding the Laser Powder Bed Fusion of AlSi10Mg Alloy, Metallography, Microstructure, and Analysis 9 (2020) 484-502. https://doi.org/10.1007/s13632-020-00659-w

[3] G. Nazami, S. Sahoo, Influence of hatch spacing and laser spot overlapping on heat transfer during laser powder bed fusion of aluminum alloy, Journal of Laser Applications 32 (2020) 042007. https://doi.org/04200710.2351/7.0000157

[4] X. Li, Y. Liu, Z. Zhou, Influence of hatch distance on processing, microstructure and mechanical properties of AlMgScZr alloy fabricated by laser powder bed fusion, Journal of Manufacturing Processes 81 (2022) 78-91. https://doi.org/10.1016/j.jmapro.2022.06.036

[5] L. Lizzul, R. Bertolini, A. Ghiotti, and S. Bruschi, Turning of Additively Manufactured Ti6Al4V: Effect of the Highly Oriented Microstructure on the Surface Integrity, Materials 14 (2021) 2842. https://doi.org/10.3390/ma14112842

[6] L. Lizzul, M. Sorgato, R. Bertolini, A. Ghiotti, and S. Bruschi, Influence of additive manufacturing-induced anisotropy on tool wear in end milling of Ti6Al4V, Tribology International 146 (2020) 106200. https://doi.org/10.1016/j.triboint.2020.106200

[7] P. Guo, B. Zou, C. Huang, and H. Gao, Study on microstructure, mechanical properties and machinability of efficiently additive manufactured AISI 316L stainless steel by high-power direct laser deposition, Journal of Material Research and Technology 240 (2017) 12-22. https://doi.org/10.1016/j.jmatprotec.2016.09.005

[8] M. Wang, B. Song, Q. Wei, Y. Zhang, Y. Shi, Effects of annealing on the microstructure and mechanical properties of selective laser melted AlSi7Mg alloy, Materials Science and Engineering A 739 (2019) 463-472. https://doi.org/10.1016/j.msea.2018.10.047

[9] BS EN ISO 25178-2:2022: Geometrical product specifications (GPS). Surface texture: Areal. Terms, definitions and surface texture parameters.

[10] Z. Dong, Y. Liu, W. Wen, J. Ge, J. Liang, Effect of Hatch Spacing on Melt Pool and As-built Quality During Selective Laser Melting of Stainless Steel: Modeling and Experimental Approaches, Materials 12 (2019) 50. https://doi.org/10.3390/ma12010050

[11] M. Xia, D. Gu, G. Yu, D. Dai, H. Chen, Q. Shi, Influence of hatch spacing on heat and mass transfer, thermodynamics and laser processability during additive manufacturing of Inconel 718 alloy, International Journal of Machine Tools and Manufacture 109 (2016) 147-157. https://doi.org/10.1016/j.ijmachtools.2016.07.010

[12] L. Tonelli, E. Liverani, A. Morri, L. Ceschini, Role of Direct Aging and Solution Treatment on Hardness, Microstructure and Residual Stress of the A357 (AlSi7Mg0.6) Alloy Produced by Powder Bed Fusion, Metallurgical and Materials Transactions B 52 (2021) 2484-2496. https://doi.org/10.1007/s11663-021-02179-6

[13] E. Simonetto, R. Bertolini, A. Ghiotti, S. Bruschi, Mechanical and microstructural behaviour of AA7075 aluminium alloy for sub-zero temperature sheet stamping process, International Journal of Mechanical Sciences 187 (2020) 105919. https://doi.org/10.1016/j.ijmecsci.2020.105919
## References

[1] Wa. Abd-Elaziem, S. Elkatatny, A.E. Abd-Elaziem, M. Khedr, M. A. A. El-baky, M. A. Hassan, M. Abu-Okail, M. Mohammed, A. Järvenpää, T. Allam, A. Hamada, On the current research progress of metallic materials fabricated by laser powder bed fusion process: a review, Journal of Material Research and Technology 20 (2022) 681-707. https://doi.org/10.1016/j.jmrt.2022.07.085

[2] H. Hyer et al., H. Hyer, L. Zhou, S. Park, G. Gottsfritz, G. Benson, B. Tolentino, B. McWilliams, K. Cho, Y. Sohn, Understanding the Laser Powder Bed Fusion of AlSi10Mg Alloy, Metallography, Microstructure, and Analysis 9 (2020) 484-502. https://doi.org/10.1007/s13632-020-00659-w

[3] G. Nazami, S. Sahoo, Influence of hatch spacing and laser spot overlapping on heat transfer during laser powder bed fusion of aluminum alloy, Journal of Laser Applications 32 (2020) 042007. https://doi.org/04200710.2351/7.0000157

[4] X. Li, Y. Liu, Z. Zhou, Influence of hatch distance on processing, microstructure and mechanical properties of AlMgScZr alloy fabricated by laser powder bed fusion, Journal of Manufacturing Processes 81 (2022) 78-91. https://doi.org/10.1016/j.jmapro.2022.06.036

[5] L. Lizzul, R. Bertolini, A. Ghiotti, and S. Bruschi, Turning of Additively Manufactured Ti6Al4V: Effect of the Highly Oriented Microstructure on the Surface Integrity, Materials 14 (2021) 2842. https://doi.org/10.3390/ma14112842

[6] L. Lizzul, M. Sorgato, R. Bertolini, A. Ghiotti, and S. Bruschi, Influence of additive manufacturing-induced anisotropy on tool wear in end milling of Ti6Al4V, Tribology International 146 (2020) 106200. https://doi.org/10.1016/j.triboint.2020.106200

[7] P. Guo, B. Zou, C. Huang, and H. Gao, Study on microstructure, mechanical properties and machinability of efficiently additive manufactured AISI 316L stainless steel by high-power direct laser deposition, Journal of Material Research and Technology 240 (2017) 12-22. https://doi.org/10.1016/j.jmatprotec.2016.09.005

[8] M. Wang, B. Song, Q. Wei, Y. Zhang, Y. Shi, Effects of annealing on the microstructure and mechanical properties of selective laser melted AlSi7Mg alloy, Materials Science and Engineering A 739 (2019) 463-472. https://doi.org/10.1016/j.msea.2018.10.047

[9] BS EN ISO 25178-2:2022: Geometrical product specifications (GPS). Surface texture: Areal. Terms, definitions and surface texture parameters.

[10] Z. Dong, Y. Liu, W. Wen, J. Ge, J. Liang, Effect of Hatch Spacing on Melt Pool and As-built Quality During Selective Laser Melting of Stainless Steel: Modeling and Experimental Approaches, Materials 12 (2019) 50. https://doi.org/10.3390/ma12010050

[11] M. Xia, D. Gu, G. Yu, D. Dai, H. Chen, Q. Shi, Influence of hatch spacing on heat and mass transfer, thermodynamics and laser processability during additive manufacturing of Inconel 718 alloy, International Journal of Machine Tools and Manufacture 109 (2016) 147-157. https://doi.org/10.1016/j.ijmachtools.2016.07.010

[12] L. Tonelli, E. Liverani, A. Morri, L. Ceschini, Role of Direct Aging and Solution Treatment on Hardness, Microstructure and Residual Stress of the A357 (AlSi7Mg0.6) Alloy Produced by Powder Bed Fusion, Metallurgical and Materials Transactions B 52 (2021) 2484-2496. https://doi.org/10.1007/s11663-021-02179-6

[13] E. Simonetto, R. Bertolini, A. Ghiotti, S. Bruschi, Mechanical and microstructural behaviour of AA7075 aluminium alloy for sub-zero temperature sheet stamping process, International Journal of Mechanical Sciences 187 (2020) 105919. https://doi.org/10.1016/j.ijmecsci.2020.105919
# Analysis of the influence of the initial surface integrity on the result of burnished high torque splined shafts

DAMM Philipp<fnref n="1" />,ᵃ*, ADINEHVAND Mohsen<fnref n="1" />,ᵇ, BEUTNER Martin<fnref n="1" />,ᶜ,
MEICHSNER Gunnar<fnref n="1" />,ᵈ, HAHN Georg<fnref n="2" />,ᵉ, ZIMMERMANN Dirk<fnref n="2" />,ᵉ,
STEFFENS Matthias<fnref n="2" />,<fnref n="9" /> and HACKERT-OSCHÄTZCHEN Matthias<fnref n="1" />,ʰ

**Keywords:** Ball Burnishing, Compressive Residual Stresses, Surface Integrity

**Abstract.** The increasing demand for low surface roughness and long service life in highly loaded components, such as extruder shafts, requires high strength, functional reliability and long-term durability. Extruder shafts are applied in plastics technology and in the food industry for processing plastic melts, mixing and compounding. Screw and kneading elements are required for this purpose. The screw and kneading elements are pushed form-fittingly onto a splined shaft. The torque is transmitted by the splined shaft. [1] Main causes of wear and damage of extruder shafts are corrosion and mechanical load. High surface qualities, increased residual compressive stresses and increased surface hardness are possibilities to increase the durability of the splined shafts. These desired properties can be achieved by applying the ball burnishing technology. [2] By utilizing cold working processes, such as ball burnishing, it is possible to enhance the surface and subsurface properties to increase the lifespan of rotary components intended for transmitting high rotational torques. Due to the importance of considering the roughness of the initial surface, this study aimed to generate various roughness and surface profiles through the turning process to investigate their influence on the burnishing process results. For this purpose, shaft sections made of AISI H11 steel were pre-machined with different cutting inserts in order to set different surface roughnesses. Subsequently, a statistical experimental design was created by using the Taguchi method by varying feed rate, pressure and speed. As a main result, roughness and residual stresses on the surface were measured and compared with the initial conditions.

**Introduction**

During machining processes, the surface of the workpiece is exposed to thermomechanical stresses that lead to changes in the surface properties. Surface properties include surface characteristics such as roughness, microstructure, residual stress and hardness. These characteristics are summarized under the term surface integrity and are shown graphically in Fig. 1 [3]. Fig. 1 shows high tensile stresses occur near to the surface of the workpiece, for example as a result of machining. To counteract this effect, both mechanical strength and service life can be improved by inducing compressive residual stresses through ball burnishing [4]. Ball burnishing is a cost-effective surface finishing process for a wide range of components, e.g., for turbine parts. This process involves the application of pressure and rolling of a hydraulically supported ball along the surface of the workpiece to plastically deform surface irregularities [5].

<figure><img src="image_1809.png" /></figure>

<footnote n="1">¹Chair of Manufacturing Technology with Focus Machining, Faculty of Mechanical Engineering, Otto von Guericke University Magdeburg, Universitätsplatz 2, 39106 Magdeburg, Germany</footnote>

<footnote n="2">²Extruder Experts GmbH & Co. KG, Am Windrad 23, 52156 Monschau, Germany</footnote>

<footnote>ᵃphilipp.damm@ovgu.de, ᵇmohsen.adinehvand@ovgu.de, ᶜmartin.beutner@ovgu.de,</footnote>

<footnote>ᵈgunnar.meichsner@ovgu.de, ᶜhahn@extruder-experts.com, ᵈzimmermann@extruder-experts.com,</footnote>

<footnote n="9">⁹steffens@extruder-experts.com, ᵇmatthias.hackert-oschaetzchen@ovgu.de</footnote>
# Analysis of the influence of the initial surface integrity on the result of burnished high torque splined shafts

DAMM Philipp<fnref n="1" />,ᵃ*, ADINEHVAND Mohsen<fnref n="1" />,ᵇ, BEUTNER Martin<fnref n="1" />,ᶜ,
MEICHSNER Gunnar<fnref n="1" />,ᵈ, HAHN Georg<fnref n="2" />,ᵉ, ZIMMERMANN Dirk<fnref n="2" />,ᵉ,
STEFFENS Matthias<fnref n="2" />,<fnref n="9" /> and HACKERT-OSCHÄTZCHEN Matthias<fnref n="1" />,ʰ

**Keywords:** Ball Burnishing, Compressive Residual Stresses, Surface Integrity

**Abstract.** The increasing demand for low surface roughness and long service life in highly loaded components, such as extruder shafts, requires high strength, functional reliability and long-term durability. Extruder shafts are applied in plastics technology and in the food industry for processing plastic melts, mixing and compounding. Screw and kneading elements are required for this purpose. The screw and kneading elements are pushed form-fittingly onto a splined shaft. The torque is transmitted by the splined shaft. [1] Main causes of wear and damage of extruder shafts are corrosion and mechanical load. High surface qualities, increased residual compressive stresses and increased surface hardness are possibilities to increase the durability of the splined shafts. These desired properties can be achieved by applying the ball burnishing technology. [2] By utilizing cold working processes, such as ball burnishing, it is possible to enhance the surface and subsurface properties to increase the lifespan of rotary components intended for transmitting high rotational torques. Due to the importance of considering the roughness of the initial surface, this study aimed to generate various roughness and surface profiles through the turning process to investigate their influence on the burnishing process results. For this purpose, shaft sections made of AISI H11 steel were pre-machined with different cutting inserts in order to set different surface roughnesses. Subsequently, a statistical experimental design was created by using the Taguchi method by varying feed rate, pressure and speed. As a main result, roughness and residual stresses on the surface were measured and compared with the initial conditions.

**Introduction**

During machining processes, the surface of the workpiece is exposed to thermomechanical stresses that lead to changes in the surface properties. Surface properties include surface characteristics such as roughness, microstructure, residual stress and hardness. These characteristics are summarized under the term surface integrity and are shown graphically in Fig. 1 [3]. Fig. 1 shows high tensile stresses occur near to the surface of the workpiece, for example as a result of machining. To counteract this effect, both mechanical strength and service life can be improved by inducing compressive residual stresses through ball burnishing [4]. Ball burnishing is a cost-effective surface finishing process for a wide range of components, e.g., for turbine parts. This process involves the application of pressure and rolling of a hydraulically supported ball along the surface of the workpiece to plastically deform surface irregularities [5].

<figure><img src="image_1807.png" /></figure>

<footnote n="1">¹Chair of Manufacturing Technology with Focus Machining, Faculty of Mechanical Engineering, Otto von Guericke University Magdeburg, Universitätsplatz 2, 39106 Magdeburg, Germany</footnote>

<footnote n="2">²Extruder Experts GmbH & Co. KG, Am Windrad 23, 52156 Monschau, Germany</footnote>

<footnote>ᵃphilipp.damm@ovgu.de, ᵇmohsen.adinehvand@ovgu.de, ᶜmartin.beutner@ovgu.de,</footnote>

<footnote>ᵈgunnar.meichsner@ovgu.de, ᶜhahn@extruder-experts.com, ᵈzimmermann@extruder-experts.com,</footnote>

<footnote n="9">⁹steffens@extruder-experts.com, ᵇmatthias.hackert-oschaetzchen@ovgu.de</footnote>
<figure><img src="image_1810.png" /><figcaption>Figure 1 - Surface integrity properties [3]</figcaption></figure>

In this process, machined workpieces made of formable materials are finished using hardened rolling elements. The goal is to enhance surface quality and mechanical strength properties. The process of ball burnishing is shown in Fig. 2 a). The ball rolling element is supplied to a hydrostatic pressure of up to 400 bar. The ball is pressed onto the surface by the hydraulic pressure and plastically deforms the roughness peaks. The hydrostatic bearing allows the ball to rotate in any direction. Ball burnishing induces deep compressive residual stresses (up to 1.0 mm) and high-quality surface finish, which is shown in Fig. 2 b). Due to the induced compressive residual stresses and high-quality surface, the fatigue strength and wear resistance is improved [3, 4, 6].

<figure><img src="image_1811.png" /><figcaption>Figure 2 a) - Principle of hydrostatic ball burnishing [6] b) Influence on the surface layer [7]</figcaption></figure>

In this study, the influence of the initial surface integrity of high-torque splined shafts after ball burnishing is to be analysed. For this purpose, shaft sections made of AISI H11 are pre-machined with different indexable inserts in order to set different surface roughnesses. By varying the process parameters of ball burnishing such as feed rate, pressure and speed, the surface roughness and residual stresses are to be measured and compared with the initial conditions.

## Experimental setup

The initial surface configuration was achieved through turning to generate various roughness and surface profiles. The experiments were conducted using a conventional NC lathe, specifically the Colchester Triumph 2500 model. Turning of the samples was performed applying CNMG type indexable inserts with a corner radius of 0.4 mm, 0.8 mm and 1.2 mm. For further investigations, a wiper insert with a corner radius of 0.8 mm was applied. The CNC lathe was equipped with an ECOROLL® hydraulic unit HGP4.5. The hydraulic unit contains a supply pump with a maximal hydraulic pressure of 400 bar. The hydraulic unit is connected to a ceramic ball labelled HG6 with a ball diameter of $d_k = 6$ mm.
<figure><img src="image_1808.png" /><figcaption>Figure 1 - Surface integrity properties [3]</figcaption></figure>

In this process, machined workpieces made of formable materials are finished using hardened rolling elements. The goal is to enhance surface quality and mechanical strength properties. The process of ball burnishing is shown in Fig. 2 a). The ball rolling element is supplied to a hydrostatic pressure of up to 400 bar. The ball is pressed onto the surface by the hydraulic pressure and plastically deforms the roughness peaks. The hydrostatic bearing allows the ball to rotate in any direction. Ball burnishing induces deep compressive residual stresses (up to 1.0 mm) and high-quality surface finish, which is shown in Fig. 2 b). Due to the induced compressive residual stresses and high-quality surface, the fatigue strength and wear resistance is improved [3, 4, 6].

<figure><img src="image_1809.png" /><figcaption>Figure 2 a) - Principle of hydrostatic ball burnishing [6] b) Influence on the surface layer [7]</figcaption></figure>

In this study, the influence of the initial surface integrity of high-torque splined shafts after ball burnishing is to be analysed. For this purpose, shaft sections made of AISI H11 are pre-machined with different indexable inserts in order to set different surface roughnesses. By varying the process parameters of ball burnishing such as feed rate, pressure and speed, the surface roughness and residual stresses are to be measured and compared with the initial conditions.

## Experimental setup

The initial surface configuration was achieved through turning to generate various roughness and surface profiles. The experiments were conducted using a conventional NC lathe, specifically the Colchester Triumph 2500 model. Turning of the samples was performed applying CNMG type indexable inserts with a corner radius of 0.4 mm, 0.8 mm and 1.2 mm. For further investigations, a wiper insert with a corner radius of 0.8 mm was applied. The CNC lathe was equipped with an ECOROLL® hydraulic unit HGP4.5. The hydraulic unit contains a supply pump with a maximal hydraulic pressure of 400 bar. The hydraulic unit is connected to a ceramic ball labelled HG6 with a ball diameter of $d_k = 6$ mm.
To determine the chemical composition of the hot work tool steel AISI H11, atomic emission spectrometry was used. The chemical composition can be seen in Table 1. The hot work tool steel was provided in the form of cylindrical shape with a diameter of 40 mm and a length of 200 mm.

<table><caption>Table 1 - Chemical composition of AISI H11</caption><thead><tr><th>C</th><th>Si</th><th>Mn</th><th>P</th><th>S</th><th>Cr</th><th>Mo</th><th>Ni</th><th>Al</th></tr></thead><tbody><tr><td>%</td><td>%</td><td>%</td><td>%</td><td>%</td><td>%</td><td>%</td><td>%</td><td>%</td></tr><tr><td>0.388</td><td>1.053</td><td>0.428</td><td>0.013</td><td>0.004</td><td>4.782</td><td>1.337</td><td>0.112</td><td>0.011</td></tr><tr><th>Co</th><th>Cu</th><th>Nb</th><th>Ti</th><th>V</th><th>W</th><th>Sn</th><th>N</th><th>Fe</th></tr><tr><td>%</td><td>%</td><td>%</td><td>%</td><td>%</td><td>%</td><td>%</td><td>%</td><td>%</td></tr><tr><td>0.03</td><td>0.079</td><td>0.008</td><td>0.019</td><td>0.867</td><td>0.047</td><td>0.006</td><td>0.026</td><td>90.78</td></tr></tbody></table>

Before executing the ball burnishing experiments, different surface roughnesses were set by using inserts with different corner radii with the same process parameters, such as cutting speed $v_c$ = 100 m/min, feed rate $f_a$ = 0.2 mm, RPM = 795 1/min and cutting depth $a_p$ = 1.5 mm. Based on the surface profile and the ball burnishing process parameters burnishing pressure $p_w$, burnishing speed $v_w$ and burnishing feed rate $f_w$, the experiments were designed according to the Taguchi method. Therefore, for each surface profile the process parameters were generated by using the Minitab software. The selection of all different process parameters is listed in the Table 2.

<table><caption>Table 2 - Selection of all process parameters</caption><thead><tr><th>Corner radius [mm]</th><th>Burnishing pressure [bar]</th><th>Burnishing speed [m/min]</th><th>Burnishing feed [mm]</th></tr></thead><tbody><tr><td>0.4</td><td>200</td><td>50</td><td>0.04</td></tr><tr><td>0.8</td><td>300</td><td>100</td><td>0.08</td></tr><tr><td>0.8 + Wiper</td><td>400</td><td>150</td><td>0.12</td></tr><tr><td>1.2</td><td>-</td><td>-</td><td>-</td></tr></tbody></table>

For the experiments, the workpiece samples were machined in such a way, that 10 sections are recognizable. A detailed view of the machined sample with a selection of the burnishing process parameters is shown in Fig. 3. Each section was machined with the same turning process parameters and the same corner radius. As seen in Fig. 3, each section was burnished with different process parameters. The last section, which is framed in blue, was kept as reference and was therefore not burnished.
To determine the chemical composition of the hot work tool steel AISI H11, atomic emission spectrometry was used. The chemical composition can be seen in Table 1. The hot work tool steel was provided in the form of cylindrical shape with a diameter of 40 mm and a length of 200 mm.

<table><caption>Table 1 - Chemical composition of AISI H11</caption><thead><tr><th>C</th><th>Si</th><th>Mn</th><th>P</th><th>S</th><th>Cr</th><th>Mo</th><th>Ni</th><th>Al</th></tr></thead><tbody><tr><td>%</td><td>%</td><td>%</td><td>%</td><td>%</td><td>%</td><td>%</td><td>%</td><td>%</td></tr><tr><td>0.388</td><td>1.053</td><td>0.428</td><td>0.013</td><td>0.004</td><td>4.782</td><td>1.337</td><td>0.112</td><td>0.011</td></tr><tr><th>Co</th><th>Cu</th><th>Nb</th><th>Ti</th><th>V</th><th>W</th><th>Sn</th><th>N</th><th>Fe</th></tr><tr><td>%</td><td>%</td><td>%</td><td>%</td><td>%</td><td>%</td><td>%</td><td>%</td><td>%</td></tr><tr><td>0.03</td><td>0.079</td><td>0.008</td><td>0.019</td><td>0.867</td><td>0.047</td><td>0.006</td><td>0.026</td><td>90.78</td></tr></tbody></table>

Before executing the ball burnishing experiments, different surface roughnesses were set by using inserts with different corner radii with the same process parameters, such as cutting speed $v_c$ = 100 m/min, feed rate $f_a$ = 0.2 mm, RPM = 795 1/min and cutting depth $a_p$ = 1.5 mm. Based on the surface profile and the ball burnishing process parameters burnishing pressure $p_w$, burnishing speed $v_w$ and burnishing feed rate $f_w$, the experiments were designed according to the Taguchi method. Therefore, for each surface profile the process parameters were generated by using the Minitab software. The selection of all different process parameters is listed in the Table 2.

<table><caption>Table 2 - Selection of all process parameters</caption><thead><tr><th>Corner radius [mm]</th><th>Burnishing pressure [bar]</th><th>Burnishing speed [m/min]</th><th>Burnishing feed [mm]</th></tr></thead><tbody><tr><td>0.4</td><td>200</td><td>50</td><td>0.04</td></tr><tr><td>0.8</td><td>300</td><td>100</td><td>0.08</td></tr><tr><td>0.8 + Wiper</td><td>400</td><td>150</td><td>0.12</td></tr><tr><td>1.2</td><td>-</td><td>-</td><td>-</td></tr></tbody></table>

For the experiments, the workpiece samples were machined in such a way, that 10 sections are recognizable. A detailed view of the machined sample with a selection of the burnishing process parameters is shown in Fig. 3. Each section was machined with the same turning process parameters and the same corner radius. As seen in Fig. 3, each section was burnished with different process parameters. The last section, which is framed in blue, was kept as reference and was therefore not burnished.
<figure><img src="image_1812.png" /><figcaption>Figure 3 - Detailed view of a machined sample with a selection of the burnishing process parameters</figcaption></figure>

## Results and evaluation

The influence of the turning parameters on the surface roughness is presented in Fig. 4. Fig. 4 a) provides a comparison between the calculated theoretical surface roughness and the measured surface roughness. The theoretical surface roughness was calculated in Eq. 1 by using corner radius $r_ε$ and feed rate $f$ [8].

$$R_{t,th} = r_{\epsilon} - \sqrt{r_{\epsilon}^2 - \frac{f^2}{4}} \quad (1)$$

By applying a wiper-modified insert, the feed rate can be doubled while maintaining the same surface roughness [9]. Therefore, the theoretical surface roughness was halved for the insert 0.8 + wiper.

<figure><img src="image_1813.png" /><figcaption>Figure 4 - a) Comparison of the theoretical and the measured surface roughness b) The formation of burrs at a corner radius of 1.2 mm</figcaption></figure>

In Fig. 4 a) the light blue bars represent the theoretical surface roughness and the dark blue bars represent the measured surface roughness after turning. The corner radius indicates a significant influence on the surface roughness.

Increasing the corner radius from 0.4 mm to 1.2 mm results in a reduction of the theoretical surface roughness from 12.5 µm to 4.17 µm. A larger corner radius $r_ε$, improves the stability of the cutting edge through favourable distribution of cutting forces and enhanced heat dissipation. It
<figure><img src="image_1810.png" /><figcaption>Figure 3 - Detailed view of a machined sample with a selection of the burnishing process parameters</figcaption></figure>

## Results and evaluation

The influence of the turning parameters on the surface roughness is presented in Fig. 4. Fig. 4 a) provides a comparison between the calculated theoretical surface roughness and the measured surface roughness. The theoretical surface roughness was calculated in Eq. 1 by using corner radius $r_ε$ and feed rate $f$ [8].

$$R_{t,th} = r_{\epsilon} - \sqrt{r_{\epsilon}^2 - \frac{f^2}{4}} \quad (1)$$

By applying a wiper-modified insert, the feed rate can be doubled while maintaining the same surface roughness [9]. Therefore, the theoretical surface roughness was halved for the insert 0.8 + wiper.

<figure><img src="image_1811.png" /><figcaption>Figure 4 - a) Comparison of the theoretical and the measured surface roughness b) The formation of burrs at a corner radius of 1.2 mm</figcaption></figure>

In Fig. 4 a) the light blue bars represent the theoretical surface roughness and the dark blue bars represent the measured surface roughness after turning. The corner radius indicates a significant influence on the surface roughness.

Increasing the corner radius from 0.4 mm to 1.2 mm results in a reduction of the theoretical surface roughness from 12.5 µm to 4.17 µm. A larger corner radius $r_ε$, improves the stability of the cutting edge through favourable distribution of cutting forces and enhanced heat dissipation. It
can be observed that a further improvement of the surface quality can be achieved by using a drag cutting edge (wiper). However, it is evident that the corner radius of 1.2 mm shows unusual results. The measured roughness depth contradicts its theoretical roughness depth. The formation of burrs could potentially be the underlying cause, as seen in Fig. 4 b).

After generating the initial surface profile with different corner radii, the burnishing process was executed. The comparison of turned and burnished surface roughness $R_z$ is shown in Fig. 5 a) and highlights the direct influence of the initial surface roughness on the resulting roughness outcomes after burnishing. The lower the initial surface roughness during turning, the better the roughness results that can be achieved after ball burnishing. The light blue bars represent a significant reduction in surface roughness following the burnishing process. After burnishing, the surface roughness $R_z$ significantly decreases in range from 83 % to 93 %. The largest improvement is seen at a 0.4 mm corner radius. From these values, it is evident that the surface roughness of the burnished surfaces varies in the range of 0.375 µm to 0.8 µm, compared to the turned surfaces whose roughness ranges from 2.33 µm to 9.36 µm. Notably, corner radii of 0.8 mm and 0.8 mm + Wiper provide the lowest roughness results for both turning and burnishing.

<figure><img src="image_1814.png" /><figcaption>Figure 5 a) - Comparison of turned and burnished surface roughness $R_z$ b) Selection of diagrams for average surface roughness $R_z$ dependent on burnishing pressure, speed and feed rate</figcaption></figure>

Fig. 5 b) presents a selection of diagrams created in Minitab software, allowing for a comprehensive analysis of the individual effects of the influencing factors during the burnishing process. According to the diagrams, it is evident that the corner radius of 0.8 mm + Wiper leads to the lowest surface roughness values. As shown, the surface roughness decreases noticeably with an increase in burnishing pressure from 200 bar to 300 bar. However, further increasing the burnishing pressure from 300 bar to 400 bar results in an increase in surface roughness. With increasing burnishing force, the tool penetrates deeper into the surface layer, enlarging the deformation area and resulting in a smoother surface. However, excessive force can lead to surface waviness and chipping. Therefore, further pressure increase causes surface degradation due to plastic flow and chipping.

The change in surface roughness as a function of the burnishing speed $v_w$ in Fig. 5 b) indicates that the surface roughness depends on the initial surface roughness. With an increase of the burnishing speed from 50 m/min to 150 m/min, the surface roughness increases significantly at a corner radius of 0.4 mm and 1.2 mm. However, the surface roughness decreases slightly with an increase of the burnishing feed from 50 m/min to 150 m/min with a corner radius of 0.8 mm and remain nearly constant at a corner radius of 0.8 mm + wiper.

The change in surface roughness as a function of the burnishing feed rate $f_w$ shows that the roughness increases slightly when the feed rate is increased from 0.04 mm to 0.08 mm at a corner radius of 0.4 mm and 0.8 mm + wiper and decrease at a corner radius of 0.8 mm. However, with a further increase in the feed rate to 0.12 mm, the surface roughness significantly increases at all corner radii. This can be attributed to the fact that at a low burnishing feed rate and burnishing
can be observed that a further improvement of the surface quality can be achieved by using a drag cutting edge (wiper). However, it is evident that the corner radius of 1.2 mm shows unusual results. The measured roughness depth contradicts its theoretical roughness depth. The formation of burrs could potentially be the underlying cause, as seen in Fig. 4 b).

After generating the initial surface profile with different corner radii, the burnishing process was executed. The comparison of turned and burnished surface roughness $R_z$ is shown in Fig. 5 a) and highlights the direct influence of the initial surface roughness on the resulting roughness outcomes after burnishing. The lower the initial surface roughness during turning, the better the roughness results that can be achieved after ball burnishing. The light blue bars represent a significant reduction in surface roughness following the burnishing process. After burnishing, the surface roughness $R_z$ significantly decreases in range from 83 % to 93 %. The largest improvement is seen at a 0.4 mm corner radius. From these values, it is evident that the surface roughness of the burnished surfaces varies in the range of 0.375 µm to 0.8 µm, compared to the turned surfaces whose roughness ranges from 2.33 µm to 9.36 µm. Notably, corner radii of 0.8 mm and 0.8 mm + Wiper provide the lowest roughness results for both turning and burnishing.

<figure><img src="image_1812.png" /><figcaption>Figure 5 a) - Comparison of turned and burnished surface roughness $R_z$ b) Selection of diagrams for average surface roughness $R_z$ dependent on burnishing pressure, speed and feed rate</figcaption></figure>

Fig. 5 b) presents a selection of diagrams created in Minitab software, allowing for a comprehensive analysis of the individual effects of the influencing factors during the burnishing process. According to the diagrams, it is evident that the corner radius of 0.8 mm + Wiper leads to the lowest surface roughness values. As shown, the surface roughness decreases noticeably with an increase in burnishing pressure from 200 bar to 300 bar. However, further increasing the burnishing pressure from 300 bar to 400 bar results in an increase in surface roughness. With increasing burnishing force, the tool penetrates deeper into the surface layer, enlarging the deformation area and resulting in a smoother surface. However, excessive force can lead to surface waviness and chipping. Therefore, further pressure increase causes surface degradation due to plastic flow and chipping.

The change in surface roughness as a function of the burnishing speed $v_w$ in Fig. 5 b) indicates that the surface roughness depends on the initial surface roughness. With an increase of the burnishing speed from 50 m/min to 150 m/min, the surface roughness increases significantly at a corner radius of 0.4 mm and 1.2 mm. However, the surface roughness decreases slightly with an increase of the burnishing feed from 50 m/min to 150 m/min with a corner radius of 0.8 mm and remain nearly constant at a corner radius of 0.8 mm + wiper.

The change in surface roughness as a function of the burnishing feed rate $f_w$ shows that the roughness increases slightly when the feed rate is increased from 0.04 mm to 0.08 mm at a corner radius of 0.4 mm and 0.8 mm + wiper and decrease at a corner radius of 0.8 mm. However, with a further increase in the feed rate to 0.12 mm, the surface roughness significantly increases at all corner radii. This can be attributed to the fact that at a low burnishing feed rate and burnishing
speed, the distance between successive rolling passes of the burnishing tool is minimal. Because of the minimal distance between the rolling passes, the ball makes more revolutions and smooth out the peaks and edges of the previous turning marks.

Fig. 6 illustrates a turned surface with a corner radius of 0.8 mm + Wiper and the corresponding surfaces after burnishing with three different pressures. The surface profile was conducted by applying a confocal 3D microscope µsurf expert. The surface roughness is greatly reduced compared to turning when using a burnishing pressure of 200 bar. However, the turning marks are not completely eliminated. By increasing the burnishing pressure to the value of 300 bar, the penetration depth of the ball burnishing tool increases, resulting in a roughness value of 0.275 µm, similar to grinding. However, applying a burnishing pressure of 400 bar leads to more pronounced plastic flow, along with chipping. As a result, the surface roughness increases due to the increased plastic deformations.

<figure><img src="image_1815.png" /><figcaption>Figure 6 - The topographies of the resulting surface roughness after a) Turning $r_e = 0.8$ mm + wiper b) Burnishing with $p_w = 200$ bar c) $p_w = 300$ bar d) $p_w = 400$ bar</figcaption></figure>

To determine the residual stress state of the samples, measurements in axial ($\phi = 0°$) and circumferential ($\phi = 90°$) directions are conducted using the Xstress G2R X-ray diffractometer from Stresstech. The measurements are based on the d-sin²χ method. A CrKα radiation source with acceleration voltages of approximately 30 keV was used. The measurement setup is shown in Fig. 7 b) for the measurement in axial direction.

Fig. 7 a) illustrates the residual stress results following the turning process on surfaces with varying surface roughness.
speed, the distance between successive rolling passes of the burnishing tool is minimal. Because of the minimal distance between the rolling passes, the ball makes more revolutions and smooth out the peaks and edges of the previous turning marks.

Fig. 6 illustrates a turned surface with a corner radius of 0.8 mm + Wiper and the corresponding surfaces after burnishing with three different pressures. The surface profile was conducted by applying a confocal 3D microscope µsurf expert. The surface roughness is greatly reduced compared to turning when using a burnishing pressure of 200 bar. However, the turning marks are not completely eliminated. By increasing the burnishing pressure to the value of 300 bar, the penetration depth of the ball burnishing tool increases, resulting in a roughness value of 0.275 µm, similar to grinding. However, applying a burnishing pressure of 400 bar leads to more pronounced plastic flow, along with chipping. As a result, the surface roughness increases due to the increased plastic deformations.

<figure><img src="image_1813.png" /><figcaption>Figure 6 - The topographies of the resulting surface roughness after a) Turning $r_e = 0.8$ mm + wiper b) Burnishing with $p_w = 200$ bar c) $p_w = 300$ bar d) $p_w = 400$ bar</figcaption></figure>

To determine the residual stress state of the samples, measurements in axial ($\phi = 0°$) and circumferential ($\phi = 90°$) directions are conducted using the Xstress G2R X-ray diffractometer from Stresstech. The measurements are based on the d-sin²χ method. A CrKα radiation source with acceleration voltages of approximately 30 keV was used. The measurement setup is shown in Fig. 7 b) for the measurement in axial direction.

Fig. 7 a) illustrates the residual stress results following the turning process on surfaces with varying surface roughness.
<figure><img src="image_1816.png" /><figcaption>Figure 7 - a) The influence of the corner radius $r_e$ on the residual stress $\sigma_{rs}$ in the axial and circumferential directions after turning b) Measurement setup for measurement in axial direction</figcaption></figure>

Notably, in the near-surface regions, both axial and circumferential tensile residual stresses have developed after the turning process. In the comparison of corner radii, a higher tensile residual stress is observed in the circumferential direction for a corner radius of $r_e = 0.4$ mm. Increasing the corner radius from 0.4 mm to 0.8 mm reduces surface roughness $R_z$ and tensile residual stress proportionally. The larger contact area leads to more even cutting force distribution and improved heat transfer, reducing thermal loads and tensile stresses on the workpiece surface. However, using the insert with a corner radius of 0.8 mm + Wiper and further expanding the contact area generates higher temperatures due to increased friction and tool wear, contributing to the formation of tensile residual stresses.

Fig. 8 a) shows the conversion of tensile residual stresses caused by turning into compressive residual stresses caused by ball burnishing.

<figure><img src="image_1817.png" /><figcaption>Figure 8 a) - Comparing the influence of the corner radius on the axial residual stress after turning and ball burnishing b) Selection of diagrams for axial compressive residual stress as a function of corner radii</figcaption></figure>

The variation of axial compressive residual stress respective to the corner radius $r_e$ in this diagram indicates that the corner radius $r_e$ does not have a significant influence on the formation of compressive residual stresses.

Fig. 8 b) shows diagrams that illustrate the dependence of compressive residual stresses on the burnishing pressure, burnishing speed and burnishing feed rate. These diagrams enable a detailed analysis of the impact of influencing factors on the resulting axial compressive residual stresses during ball burnishing. Fig. 8 b) shows that the corner radius $r_e$ of 0.4 mm caused the highest
<figure><img src="image_1814.png" /><figcaption>Figure 7 - a) The influence of the corner radius $r_e$ on the residual stress $\sigma_{rs}$ in the axial and circumferential directions after turning b) Measurement setup for measurement in axial direction</figcaption></figure>

Notably, in the near-surface regions, both axial and circumferential tensile residual stresses have developed after the turning process. In the comparison of corner radii, a higher tensile residual stress is observed in the circumferential direction for a corner radius of $r_e = 0.4$ mm. Increasing the corner radius from 0.4 mm to 0.8 mm reduces surface roughness $R_z$ and tensile residual stress proportionally. The larger contact area leads to more even cutting force distribution and improved heat transfer, reducing thermal loads and tensile stresses on the workpiece surface. However, using the insert with a corner radius of 0.8 mm + Wiper and further expanding the contact area generates higher temperatures due to increased friction and tool wear, contributing to the formation of tensile residual stresses.

Fig. 8 a) shows the conversion of tensile residual stresses caused by turning into compressive residual stresses caused by ball burnishing.

<figure><img src="image_1815.png" /><figcaption>Figure 8 a) - Comparing the influence of the corner radius on the axial residual stress after turning and ball burnishing b) Selection of diagrams for axial compressive residual stress as a function of corner radii</figcaption></figure>

The variation of axial compressive residual stress respective to the corner radius $r_e$ in this diagram indicates that the corner radius $r_e$ does not have a significant influence on the formation of compressive residual stresses.

Fig. 8 b) shows diagrams that illustrate the dependence of compressive residual stresses on the burnishing pressure, burnishing speed and burnishing feed rate. These diagrams enable a detailed analysis of the impact of influencing factors on the resulting axial compressive residual stresses during ball burnishing. Fig. 8 b) shows that the corner radius $r_e$ of 0.4 mm caused the highest
compressive residual stresses. Therefore, these values will be used for the evaluation of the results. An increase in the burnishing pressure $p_w$ from 200 bar to 400 bar results in an elevation of the compressive residual stresses from 890 MPa to 1609 MPa. Hence, it is important to note that doubling the burnishing pressure nearly doubles the compressive residual stresses.

The variation of compressive residual stresses respective to the burnishing speed $v_w$ indicates that the burnishing speed has a minor influence on the resulting compressive residual stresses. Low burnishing speeds allow more time for cold deformations and material flow, leading to increased compressive residual stresses. However, higher burnishing speeds can shorten the process cycle. Based on these findings, a burnishing speed of 50 m/min yields higher compressive residual stresses. Similar to the influence of burnishing speed $v_w$ on the generation of compressive residual stresses, the burnishing feed rate $f_w$ also shows no significant effect on the formation of compressive residual stresses in the near-surface regions. Only at low feed rates during ball burnishing, higher compressive residual stresses occur due to the longer exposure time. From a residual stress perspective, it is possible to burnish a surface with the machine-supported maximum feed rate and the specified surface roughness to enhance process productivity.

In order to analyse the effects of ball burnishing on the microstructure of the workpieces, the samples were examined using a scanning electron microscope (SEM). Fig. 9 shows SEM images of the microstructure in the respective edge zone after turning and after ball burnishing for a depth range of approx. 15 µm.

<figure><img src="image_1818.png" /><figcaption>Figure 9 - Visualisation of the microstructure by SEM images a) after turning and b) after ball burnishing</figcaption></figure>

In comparison with the analysed microstructure of the turned and subsequently burnished samples, no significant influence of the burnishing process can be seen. Three layers are recognisable in the SEM image. The first layer has a thickness of 2 µm and shows a finely recrystallised layer in which the stretched grains are clearly visible. The second layer is moderately deformed and has a thickness of between 5 µm and 7 µm. From 7 µm, the third layer can be recognised, which corresponds to the original grain structure.

## Summary

The impact of ball burnishing on the surface integrity of shafts made of steel AISI H11 has been examined.
compressive residual stresses. Therefore, these values will be used for the evaluation of the results. An increase in the burnishing pressure $p_w$ from 200 bar to 400 bar results in an elevation of the compressive residual stresses from 890 MPa to 1609 MPa. Hence, it is important to note that doubling the burnishing pressure nearly doubles the compressive residual stresses.

The variation of compressive residual stresses respective to the burnishing speed $v_w$ indicates that the burnishing speed has a minor influence on the resulting compressive residual stresses. Low burnishing speeds allow more time for cold deformations and material flow, leading to increased compressive residual stresses. However, higher burnishing speeds can shorten the process cycle. Based on these findings, a burnishing speed of 50 m/min yields higher compressive residual stresses. Similar to the influence of burnishing speed $v_w$ on the generation of compressive residual stresses, the burnishing feed rate $f_w$ also shows no significant effect on the formation of compressive residual stresses in the near-surface regions. Only at low feed rates during ball burnishing, higher compressive residual stresses occur due to the longer exposure time. From a residual stress perspective, it is possible to burnish a surface with the machine-supported maximum feed rate and the specified surface roughness to enhance process productivity.

In order to analyse the effects of ball burnishing on the microstructure of the workpieces, the samples were examined using a scanning electron microscope (SEM). Fig. 9 shows SEM images of the microstructure in the respective edge zone after turning and after ball burnishing for a depth range of approx. 15 µm.

<figure><img src="image_1816.png" /><figcaption>Figure 9 - Visualisation of the microstructure by SEM images a) after turning and b) after ball burnishing</figcaption></figure>

In comparison with the analysed microstructure of the turned and subsequently burnished samples, no significant influence of the burnishing process can be seen. Three layers are recognisable in the SEM image. The first layer has a thickness of 2 µm and shows a finely recrystallised layer in which the stretched grains are clearly visible. The second layer is moderately deformed and has a thickness of between 5 µm and 7 µm. From 7 µm, the third layer can be recognised, which corresponds to the original grain structure.

## Summary

The impact of ball burnishing on the surface integrity of shafts made of steel AISI H11 has been examined.
The resulting surface roughness is significantly influenced by the initial surface roughness. The results indicate a reduction in surface roughness after ball burnishing between 83 % to 93 %. Moreover, it is noted that a lower initial surface roughness contributes to an improved surface roughness after the burnishing process. The variation of surface roughness respective to the burnishing feed rate shows that at a lower feed rate, the distance between successive rolling paths of the burnishing tool is minimal, and the ball traverses more revolutions to smooth out the peaks and edges of the previous turning marks. This results in a reduction of surface roughness.

The initial surface roughness has only a minimal influence on the formation of compressive residual stresses after ball burnishing. The corresponding mechanical stresses after burnishing result in the transformation of tensile residual stresses generated during turning into compressive residual stresses, thereby enhancing the workpiece's fatigue life and wear resistance. The burnishing pressure exerts a highly significant influence on the development of compressive residual stresses on the surface. Increasing the burnishing pressure from 200 bar to 400 bar leads to a rise in compressive residual stresses from 890 MPa to 1609 MPa, equivalent to nearly doubling the magnitude.

Based on this study, it can be concluded that the initial roughness of the material AISI H11, which is used for extruder shafts, only has a minor impact on the ball burnishing process.

## Acknowledgements

This project is funded by the Federal Ministry for Economic Affairs and Climate Action, following a decision of the German Bundestag.

## References

[1] K. Kohlgrüber, Ed. Der gleichläufige Doppelschneckenextruder: Grundlagen, Technologie, Anwendungen, 2nd ed. (Hanser eLibrary). München: Hanser, 2016.

[2] F. Klocke and J. Liermann, "Roller burnishing of hard turned surfaces," International Journal of Machine Tools and Manufacture, vol. 38, 5-6, pp. 419-423, 1998. https://doi.org/10.1016/S0890-6955(97)00085-0

[3] V. Chomienne, F. Valiorgue, J. Rech, and C. Verdu, "Influence of ball burnishing on residual stress profile of a 15-5PH stainless steel," CIRP Journal of Manufacturing Science and Technology, vol. 13, pp. 90-96, 2016. https://doi.org/10.1016/j.cirpj.2015.12.003

[4] T. Zhang, N. Bugtai, and I. D. Marinescu, "Burnishing of aerospace alloy: A theoretical-experimental approach," Journal of Manufacturing Systems, vol. 37, pp. 472-478, 2015. https://doi.org/10.1016/j.jmsy.2014.11.004

[5] L. Luca, S. Neagu-Ventzel, and I. Marinescu, "Effects of working parameters on surface finish in ball-burnishing of hardened steels," Precision Engineering, vol. 29, no. 2, pp. 253-256, 2005. https://doi.org/10.1016/j.precisioneng.2004.02.002

[6] A. Rodríguez, L. N. López de Lacalle, A. Celaya, A. Lamikiz, and J. Albizuri, "Surface improvement of shafts by the deep ball-burnishing technique," Surface and Coatings Technology, vol. 206, 11-12, pp. 2817-2824, 2012. https://doi.org/10.1016/j.surfcoat.2011.11.045

[7] N.N. "Solutions for tool and mold construction." Accessed: 24.11.23. [Online]. Available: https://www.ecoroll.de/fileadmin/user_upload/ECOROLL_AG_ECOmold_E.pdf

[8] B. Denkena and H. K. Tönshoff, Eds. Spanen: Grundlagen, 3rd ed. (VDI-Buch). Berlin, Heidelberg: Springer, 2011

[9] P. Zhang and Z. Liu, "Modeling and prediction for 3D surface topography in finish turning with conventional and wiper inserts," Measurement, vol. 94, pp. 37-45, 2016. https://doi.org/10.1016/j.measurement.2016.07.080
The resulting surface roughness is significantly influenced by the initial surface roughness. The results indicate a reduction in surface roughness after ball burnishing between 83 % to 93 %. Moreover, it is noted that a lower initial surface roughness contributes to an improved surface roughness after the burnishing process. The variation of surface roughness respective to the burnishing feed rate shows that at a lower feed rate, the distance between successive rolling paths of the burnishing tool is minimal, and the ball traverses more revolutions to smooth out the peaks and edges of the previous turning marks. This results in a reduction of surface roughness.

The initial surface roughness has only a minimal influence on the formation of compressive residual stresses after ball burnishing. The corresponding mechanical stresses after burnishing result in the transformation of tensile residual stresses generated during turning into compressive residual stresses, thereby enhancing the workpiece's fatigue life and wear resistance. The burnishing pressure exerts a highly significant influence on the development of compressive residual stresses on the surface. Increasing the burnishing pressure from 200 bar to 400 bar leads to a rise in compressive residual stresses from 890 MPa to 1609 MPa, equivalent to nearly doubling the magnitude.

Based on this study, it can be concluded that the initial roughness of the material AISI H11, which is used for extruder shafts, only has a minor impact on the ball burnishing process.

## Acknowledgements

This project is funded by the Federal Ministry for Economic Affairs and Climate Action, following a decision of the German Bundestag.

## References

[1] K. Kohlgrüber, Ed. Der gleichläufige Doppelschneckenextruder: Grundlagen, Technologie, Anwendungen, 2nd ed. (Hanser eLibrary). München: Hanser, 2016.

[2] F. Klocke and J. Liermann, "Roller burnishing of hard turned surfaces," International Journal of Machine Tools and Manufacture, vol. 38, 5-6, pp. 419-423, 1998. https://doi.org/10.1016/S0890-6955(97)00085-0

[3] V. Chomienne, F. Valiorgue, J. Rech, and C. Verdu, "Influence of ball burnishing on residual stress profile of a 15-5PH stainless steel," CIRP Journal of Manufacturing Science and Technology, vol. 13, pp. 90-96, 2016. https://doi.org/10.1016/j.cirpj.2015.12.003

[4] T. Zhang, N. Bugtai, and I. D. Marinescu, "Burnishing of aerospace alloy: A theoretical-experimental approach," Journal of Manufacturing Systems, vol. 37, pp. 472-478, 2015. https://doi.org/10.1016/j.jmsy.2014.11.004

[5] L. Luca, S. Neagu-Ventzel, and I. Marinescu, "Effects of working parameters on surface finish in ball-burnishing of hardened steels," Precision Engineering, vol. 29, no. 2, pp. 253-256, 2005. https://doi.org/10.1016/j.precisioneng.2004.02.002

[6] A. Rodríguez, L. N. López de Lacalle, A. Celaya, A. Lamikiz, and J. Albizuri, "Surface improvement of shafts by the deep ball-burnishing technique," Surface and Coatings Technology, vol. 206, 11-12, pp. 2817-2824, 2012. https://doi.org/10.1016/j.surfcoat.2011.11.045

[7] N.N. "Solutions for tool and mold construction." Accessed: 24.11.23. [Online]. Available: https://www.ecoroll.de/fileadmin/user_upload/ECOROLL_AG_ECOmold_E.pdf

[8] B. Denkena and H. K. Tönshoff, Eds. Spanen: Grundlagen, 3rd ed. (VDI-Buch). Berlin, Heidelberg: Springer, 2011

[9] P. Zhang and Z. Liu, "Modeling and prediction for 3D surface topography in finish turning with conventional and wiper inserts," Measurement, vol. 94, pp. 37-45, 2016. https://doi.org/10.1016/j.measurement.2016.07.080
# Influence of the depth of cut on the AM nitinol properties in flood and cryogenic machining

GUARISE Alessandra¹,ª, BERTOLINI Rachele¹,ᵇ, GHIOTTI Andrea¹,ᶜ, BRUSCHI Stefania¹,ᵈ*

¹Department of Industrial Engineering (DII), University of Padova, Via Venezia 1, 35131, Padova, Italy

ªalessandra.guarise@phd.unipd.it, ᵇrachele.bertolini@unipd.it, ᶜandrea.ghiotti@unipd.it, ᵈstefania.bruschi@unipd.it

**Keywords:** Nitinol, Additive Manufacturing, Machining, Depth of Cut, Surface Finish

**Abstract.** Nickel-titanium alloys, commonly called Nitinol, are known for two fundamental properties, namely the shape memory effect (SME) and the superelasticity effect (SE), which make Nitinol a material of great interest for applications in various fields, such as biomedical, aerospace, automotive, and electronics. Most of the published research studies discuss the Nitinol properties after just one step of fabrication, e.g. forming, casting, additive manufacturing (AM), or machining. On the contrary, this work focuses on a process chain including AM followed by heat treatment and both flood and cryogenic machining. In particular, the depth of cut during the last machining step was varied (0.1, 0.25, and 0.4 mm), to investigate its effects on the AM and heat-treated Nitinol in terms of machined surface finish and transformation temperatures. It was found that when cutting with a depth of cut of 0.25 mm, the workpiece shows the worst roughness, regardless of the material microstructure and machining condition. However, for other depths of cut, there is a clear improvement in terms of surface finish. Machining increases the Nitinol transformation temperatures, but changing the depth of cut does not have a significant effect. Cryogenic machining induces a worse surface finish, but has less effect on the transformation temperatures than flood machining.

**Introduction**

Nitinol (hereafter called NiTi) alloy is a type of shape memory alloy (SMA) and is commonly known as 'metal with a memory'. It was discovered in 1959 at the U.S. Naval Ordnance Laboratory and is named Nitinol after its composition of nickel and titanium and the initials of the laboratory [1]. Because of its unique characteristics, this material is considered revolutionary for manufacturing engineering parts devoted to the biomedical and aerospace sectors. NiTi exhibits a martensitic transformation due to thermo-mechanical loads and recovers large inelastic strains upon unloading [2]. The martensitic transformation occurs when the crystal lattice changes from the high temperature high-symmetry austenitic parent phase to the low temperature, low-symmetry martensitic phase. Transformation can be achieved in two different ways, which correspond to the shape memory effect (SME) and superelasticity (SE). The shape memory effect happens when the material is cooled to a temperature below the martensite finish, causing the alloy to transform from the austenitic to the fully martensitic phase. If a mechanical load is applied at this point, the material undergoes large, microscopic inelastic strains. However, when heated to a temperature higher than the austenite finish, it returns to the austenitic phase, and the inelastic strains are recovered. The superelastic effect occurs when a load is applied to the material in the austenitic phase, inducing martensitic transformation. Upon release of the load, the reverse transformation occurs, resulting in the recovery of the load-induced strains.

The NiTi characteristics and properties are closely related to its actual chemical composition. NiTi parts obtained from additive manufacturing (AM) exhibit the shape memory effect when

<figure><img src="image_1819.png" /></figure>
# Influence of the depth of cut on the AM nitinol properties in flood and cryogenic machining

GUARISE Alessandra<fnref n="1" />,ª, BERTOLINI Rachele<fnref n="1" />,ᵇ, GHIOTTI Andrea<fnref n="1" />,ᶜ, BRUSCHI Stefania<fnref n="1" />,ᵈ*

**Keywords:** Nitinol, Additive Manufacturing, Machining, Depth of Cut, Surface Finish

**Abstract.** Nickel-titanium alloys, commonly called Nitinol, are known for two fundamental properties, namely the shape memory effect (SME) and the superelasticity effect (SE), which make Nitinol a material of great interest for applications in various fields, such as biomedical, aerospace, automotive, and electronics. Most of the published research studies discuss the Nitinol properties after just one step of fabrication, e.g. forming, casting, additive manufacturing (AM), or machining. On the contrary, this work focuses on a process chain including AM followed by heat treatment and both flood and cryogenic machining. In particular, the depth of cut during the last machining step was varied (0.1, 0.25, and 0.4 mm), to investigate its effects on the AM and heat-treated Nitinol in terms of machined surface finish and transformation temperatures. It was found that when cutting with a depth of cut of 0.25 mm, the workpiece shows the worst roughness, regardless of the material microstructure and machining condition. However, for other depths of cut, there is a clear improvement in terms of surface finish. Machining increases the Nitinol transformation temperatures, but changing the depth of cut does not have a significant effect. Cryogenic machining induces a worse surface finish, but has less effect on the transformation temperatures than flood machining.

**Introduction**

Nitinol (hereafter called NiTi) alloy is a type of shape memory alloy (SMA) and is commonly known as 'metal with a memory'. It was discovered in 1959 at the U.S. Naval Ordnance Laboratory and is named Nitinol after its composition of nickel and titanium and the initials of the laboratory [1]. Because of its unique characteristics, this material is considered revolutionary for manufacturing engineering parts devoted to the biomedical and aerospace sectors. NiTi exhibits a martensitic transformation due to thermo-mechanical loads and recovers large inelastic strains upon unloading [2]. The martensitic transformation occurs when the crystal lattice changes from the high temperature high-symmetry austenitic parent phase to the low temperature, low-symmetry martensitic phase. Transformation can be achieved in two different ways, which correspond to the shape memory effect (SME) and superelasticity (SE). The shape memory effect happens when the material is cooled to a temperature below the martensite finish, causing the alloy to transform from the austenitic to the fully martensitic phase. If a mechanical load is applied at this point, the material undergoes large, microscopic inelastic strains. However, when heated to a temperature higher than the austenite finish, it returns to the austenitic phase, and the inelastic strains are recovered. The superelastic effect occurs when a load is applied to the material in the austenitic phase, inducing martensitic transformation. Upon release of the load, the reverse transformation occurs, resulting in the recovery of the load-induced strains.

The NiTi characteristics and properties are closely related to its actual chemical composition. NiTi parts obtained from additive manufacturing (AM) exhibit the shape memory effect when

<figure><img src="image_1817.png" /></figure>

<footnote n="1">¹Department of Industrial Engineering (DII), University of Padova, Via Venezia 1, 35131, Padova, Italy</footnote>

<footnote>ªalessandra.guarise@phd.unipd.it, ᵇrachele.bertolini@unipd.it, ᶜandrea.ghiotti@unipd.it,</footnote>

<footnote>ᵈstefania.bruschi@unipd.it</footnote>
fabricated from a Ti-rich powder, on the contrary, superelasticity effect is achieved when the parts are made from a Ni-rich powder [3].

Numerous studies have investigated the machinability of the wrought NiTi during finishing machining, investigating the effects of the cutting parameters on the NiTi properties. For instance, in [4] it was investigated the impact of the cutting speed on the surface integrity of hot-rolled Ni-rich NiTi samples during cryogenic and dry machining. The study revealed that cryogenic machining is more beneficial than dry cutting, and that the cutting speed has a strong influence on the surface and subsurface of the processed NiTi samples. Also in [5] it was studied the effect of the cutting speed on hot-forged Ni-rich NiTi dry machined samples, finding that, at low cutting speeds, there is a higher radial shape recovery of the machined material due to the phase transformations that are induced. To achieve this, in [5] it was analyzed the material phase and temperature distribution during the machining operation.

Previous research has also explored various techniques and process parameters for manufacturing NiTi components using AM, with the aim of achieving a “complexity for free” manufacturing process. However, if the goal is to obtain a final product, it is essential to combine AM with heat treatment and finishing processes and analyze the manufacturing impact at each stage on the final NiTi characteristics. In [6] AM NiTi specimens aged at 600°C and cryogenically machined were shown to exhibit superior surface integrity and superelasticity compared to both untreated specimens and those machined with conventional cutting fluids.

Nevertheless, there is still a lack of research that examines the entire process from AM to finishing machining at varying process parameters, especially the ones of the final machining step. To this aim, a Ni-rich NiTi alloy was used in a process chain that includes the fabrication of NiTi parts by AM, followed by heat treatments at various aging temperatures, and finally finishing machining using different depths of cut, both under flood and cryogenic cooling conditions. The effect of varying the depth of cut is evaluated by characterizing the machined material in terms of surface roughness, surface defects and transformation temperatures.

## Materials and Methods

Laser powder bed fusion (LPBF) was used as first step of the process chain shown in Fig. 1 to fabricate NiTi cylinders of 40 mm height and 10 mm diameter, making use of the parameters reported in Table 1. The LPBF parameters were optimized in a previous work [7] to have NiTi samples with an austenitic microstructure at room temperature after AM. Pre-alloyed spherical Ni-rich Ni₅₀.₄Ti₄₉.₆ (at%) powders produced by gas atomization (ECKART TLS™) from an as-cast ingot in a size range between 15 µm and 45 µm were used for the LPBF process.

<table><caption>Table 1: LPBF optimized parameters [7].</caption><thead><tr><td>Power [W]</td><td>Scanning speed [mm/s]</td><td>Hatch distance [mm]</td><td>Layer thickness [mm]</td><td>Oxygen level [ppm]</td></tr></thead><tbody><tr><td>70</td><td>1100</td><td>0.06</td><td>0.025</td><td>&lt;200</td></tr></tbody></table>

LPBF is usually followed by a heat treatment to relieve the residual stresses introduced by AM; in the case of Ni-rich NiTi, the post-AM heat treatment is also used to emphasize the material superelasticity thanks to the formation of Ni-rich precipitates through ageing [3]. The chosen heat treatment was carried out in a Carbolite Gero™ CWF 1200 furnace. Firstly, a relief annealing was performed at 1060°C for 60 minutes in an argon atmosphere, followed by rapid cooling in water. Subsequently, two sets of samples were aged at two different temperatures, namely 300°C (hereafter these samples are referred to as T3) and 600°C (hereafter these samples are referred to as T6). Both the ageing steps lasted 180 minutes and were followed by air cooling till room
fabricated from a Ti-rich powder, on the contrary, superelasticity effect is achieved when the parts are made from a Ni-rich powder [3].

Numerous studies have investigated the machinability of the wrought NiTi during finishing machining, investigating the effects of the cutting parameters on the NiTi properties. For instance, in [4] it was investigated the impact of the cutting speed on the surface integrity of hot-rolled Ni-rich NiTi samples during cryogenic and dry machining. The study revealed that cryogenic machining is more beneficial than dry cutting, and that the cutting speed has a strong influence on the surface and subsurface of the processed NiTi samples. Also in [5] it was studied the effect of the cutting speed on hot-forged Ni-rich NiTi dry machined samples, finding that, at low cutting speeds, there is a higher radial shape recovery of the machined material due to the phase transformations that are induced. To achieve this, in [5] it was analyzed the material phase and temperature distribution during the machining operation.

Previous research has also explored various techniques and process parameters for manufacturing NiTi components using AM, with the aim of achieving a “complexity for free” manufacturing process. However, if the goal is to obtain a final product, it is essential to combine AM with heat treatment and finishing processes and analyze the manufacturing impact at each stage on the final NiTi characteristics. In [6] AM NiTi specimens aged at 600°C and cryogenically machined were shown to exhibit superior surface integrity and superelasticity compared to both untreated specimens and those machined with conventional cutting fluids.

Nevertheless, there is still a lack of research that examines the entire process from AM to finishing machining at varying process parameters, especially the ones of the final machining step. To this aim, a Ni-rich NiTi alloy was used in a process chain that includes the fabrication of NiTi parts by AM, followed by heat treatments at various aging temperatures, and finally finishing machining using different depths of cut, both under flood and cryogenic cooling conditions. The effect of varying the depth of cut is evaluated by characterizing the machined material in terms of surface roughness, surface defects and transformation temperatures.

## Materials and Methods

Laser powder bed fusion (LPBF) was used as first step of the process chain shown in Fig. 1 to fabricate NiTi cylinders of 40 mm height and 10 mm diameter, making use of the parameters reported in Table 1. The LPBF parameters were optimized in a previous work [7] to have NiTi samples with an austenitic microstructure at room temperature after AM. Pre-alloyed spherical Ni-rich Ni₅₀.₄Ti₄₉.₆ (at%) powders produced by gas atomization (ECKART TLS™) from an as-cast ingot in a size range between 15 µm and 45 µm were used for the LPBF process.

<table><caption>Table 1: LPBF optimized parameters [7].</caption><thead><tr><td>Power [W]</td><td>Scanning speed [mm/s]</td><td>Hatch distance [mm]</td><td>Layer thickness [mm]</td><td>Oxygen level [ppm]</td></tr></thead><tbody><tr><td>70</td><td>1100</td><td>0.06</td><td>0.025</td><td>&lt;200</td></tr></tbody></table>

LPBF is usually followed by a heat treatment to relieve the residual stresses introduced by AM; in the case of Ni-rich NiTi, the post-AM heat treatment is also used to emphasize the material superelasticity thanks to the formation of Ni-rich precipitates through ageing [3]. The chosen heat treatment was carried out in a Carbolite Gero™ CWF 1200 furnace. Firstly, a relief annealing was performed at 1060°C for 60 minutes in an argon atmosphere, followed by rapid cooling in water. Subsequently, two sets of samples were aged at two different temperatures, namely 300°C (hereafter these samples are referred to as T3) and 600°C (hereafter these samples are referred to as T6). Both the ageing steps lasted 180 minutes and were followed by air cooling till room
temperature. For sake of comparison, another set of samples was considered, namely the as-built (AB) samples that did not undergo any heat treatment.

After AM, finishing machining is often mandatory to ensure the part's final geometry and properties. To this regard, the AB, T3 and T6 samples were turned at varying depth of cut on a Mori Seiki™ NL1500 CNC lathe using VCEX 110301L-F 1125 inserts. The cutting speed and feed were kept constant and equal to 31 m/min and 0.07 mm/rev, respectively. The turning trials were carried out using both a conventional cutting fluid and liquid nitrogen ($LN_2$). Details regarding the cryogenic apparatus mounted on the lathe can be found in [8]. Table 2 reports the finish machining parameters.

<table><caption>Table 2: Finishing machining parameters.</caption><thead><tr><th>Lubricating-<br>cooling conditions</th><th>Depth of cut<br>a<sub>p</sub><br>[mm]</th><th>Cutting speed<br>V<br>[m/min]</th><th>Feed<br>f<br>[mm/rev]</th><th>Length of turning<br>L<br>[mm]</th></tr></thead><tbody><tr><td>Flood (F), Cryo (C)</td><td>0.1, 0.25, 0.4</td><td>31</td><td>0.07</td><td>3</td></tr></tbody></table>

The differences between the AB, T3 and T6 samples were first assessed by comparing the microstructures, which were observed using a Leica™ DMR light microscope. Cross-section pieces for microstructural analysis were cut from the NiTi cylinders using the Brillant 210A QATM™ cut off machine applying a coolant to prevent unwanted changes in the material due to temperature rise. The NiTi pieces were mounted in resin, polished with sandpaper, and then fine-polished in a solution containing 2/3 alkaline silica suspension and 1/3 volume of 30% hydrogen peroxide ($H_2O_2$). Finally, they were etched in a solution of HF (3.2 mL), $HNO_3$ (14.1 mL), and water (82.7 mL) for 45 seconds.

The knowledge of the NiTi transformation temperatures (TTs), defined as austenite start and finish ($A_s$, $A_f$), and martensite start and finish ($M_s$, $M_f$), is crucial for determining whether the material is in the austenitic (B2) or martensitic (B19') phase, or in the transformation zone (B2=B19'), at a given temperature, and therefore for assessing the impact of the process chain parameters on the NiTi superelasticity. The TTs were evaluated by means of differential scanning calorimetry (DSC) making use of the DSC Q200™ instrument, according to the F2004-17 standard [9]. For the DSC analysis, NiTi pieces of 30-50 mg were cut using the cut off machine, from both the center of the material (bulk) and the surface to assess the machining effects. The cycle that was adopted in the DSC analysis consists of cooling down to -70°C and heating the sample to 120°C at a heating-cooling rate of 10°C/min.

The surface of the machined samples was analysed using the Sensofar Plu Neox™ 3D optical profiler, which provides a 3D image of the sample's topography. The post-processing with the sensoVIEW™ software was carried out according to the ISO 25178 standard [10] which involves using a $\lambda_s$ filter of 0.25 $\mu$m, a cut-off $\lambda_c$ of 0.25 $\mu$m for an evaluation length of 1.25 mm.

Finally, the wear condition of the inserts and the state of the workpiece surface were examined using the FEI QUANTA 450™ scanning electron microscope (SEM) equipped with the secondary electron and backscattered detectors.
temperature. For sake of comparison, another set of samples was considered, namely the as-built (AB) samples that did not undergo any heat treatment.

After AM, finishing machining is often mandatory to ensure the part's final geometry and properties. To this regard, the AB, T3 and T6 samples were turned at varying depth of cut on a Mori Seiki™ NL1500 CNC lathe using VCEX 110301L-F 1125 inserts. The cutting speed and feed were kept constant and equal to 31 m/min and 0.07 mm/rev, respectively. The turning trials were carried out using both a conventional cutting fluid and liquid nitrogen ($LN_2$). Details regarding the cryogenic apparatus mounted on the lathe can be found in [8]. Table 2 reports the finish machining parameters.

<table><caption>Table 2: Finishing machining parameters.</caption><thead><tr><th>Lubricating-<br>cooling conditions</th><th>Depth of cut<br>a<sub>p</sub><br>[mm]</th><th>Cutting speed<br>V<br>[m/min]</th><th>Feed<br>f<br>[mm/rev]</th><th>Length of turning<br>L<br>[mm]</th></tr></thead><tbody><tr><td>Flood (F), Cryo (C)</td><td>0.1, 0.25, 0.4</td><td>31</td><td>0.07</td><td>3</td></tr></tbody></table>

The differences between the AB, T3 and T6 samples were first assessed by comparing the microstructures, which were observed using a Leica™ DMR light microscope. Cross-section pieces for microstructural analysis were cut from the NiTi cylinders using the Brillant 210A QATM™ cut off machine applying a coolant to prevent unwanted changes in the material due to temperature rise. The NiTi pieces were mounted in resin, polished with sandpaper, and then fine-polished in a solution containing 2/3 alkaline silica suspension and 1/3 volume of 30% hydrogen peroxide ($H_2O_2$). Finally, they were etched in a solution of HF (3.2 mL), $HNO_3$ (14.1 mL), and water (82.7 mL) for 45 seconds.

The knowledge of the NiTi transformation temperatures (TTs), defined as austenite start and finish ($A_s$, $A_f$), and martensite start and finish ($M_s$, $M_f$), is crucial for determining whether the material is in the austenitic (B2) or martensitic (B19') phase, or in the transformation zone (B2=B19'), at a given temperature, and therefore for assessing the impact of the process chain parameters on the NiTi superelasticity. The TTs were evaluated by means of differential scanning calorimetry (DSC) making use of the DSC Q200™ instrument, according to the F2004-17 standard [9]. For the DSC analysis, NiTi pieces of 30-50 mg were cut using the cut off machine, from both the center of the material (bulk) and the surface to assess the machining effects. The cycle that was adopted in the DSC analysis consists of cooling down to -70°C and heating the sample to 120°C at a heating-cooling rate of 10°C/min.

The surface of the machined samples was analysed using the Sensofar Plu Neox™ 3D optical profiler, which provides a 3D image of the sample's topography. The post-processing with the sensoVIEW™ software was carried out according to the ISO 25178 standard [10] which involves using a $\lambda_s$ filter of 0.25 µm, a cut-off $\lambda_c$ of 0.25 µm for an evaluation length of 1.25 mm.

Finally, the wear condition of the inserts and the state of the workpiece surface were examined using the FEI QUANTA 450™ scanning electron microscope (SEM) equipped with the secondary electron and backscattered detectors.
<figure><img src="image_1820.png" /><figcaption>Fig. 1 Process chain for fabricating the NiTi samples: a) LPBF, b) heat treatment, and c) finishing machining (here shown with the cryogenic cooling apparatus).</figcaption></figure>

## Results and Discussion

Pre-machining Characteristics. Fig. 2 shows the microstructure of the AB, T3, and T6 samples obtained using the light microscope. The images depict a cross-section of the sample that is perpendicular to the building direction (BD). The microstructure is the primary factor distinguishing the samples that underwent different steps before machining. The AB sample exhibits a conventional post-AM microstructure, characterized by equiaxed grains with an average size equivalent to the hatch distance (60 µm). Upon heat treatment, the grain morphology of the AB sample is replaced by larger, irregularly shaped grains. This results in grain coarsening and the formation of various precipitates, whose chemical composition and morphology vary depending on the adopted ageing parameters.

Table 3 displays the martensite start and finish, as well as austenite start and finish values. When considering $A_f$, which sets the temperature at which the material can be considered austenitic, the two heat-treated samples behave differently when compared to the AB sample. The T3 sample shows an increase in $A_f$ of 65.4%, while, on the contrary, the T6 sample shows a decrease of 30.8% compared to the AB one.

<figure><img src="image_1821.png" /><figcaption>Fig. 2: Optical micrographs of the AB, T3 and T6 samples.</figcaption></figure>

The different behaviour of the two heat treatments can be attributed to the types of precipitates formed because of ageing. Specifically, ageing the Ni-rich NiTi at approximately 300°C results in
<figure><img src="image_1818.png" /><figcaption>Fig. 1 Process chain for fabricating the NiTi samples: a) LPBF, b) heat treatment, and c) finishing machining (here shown with the cryogenic cooling apparatus).</figcaption></figure>

## Results and Discussion

Pre-machining Characteristics. Fig. 2 shows the microstructure of the AB, T3, and T6 samples obtained using the light microscope. The images depict a cross-section of the sample that is perpendicular to the building direction (BD). The microstructure is the primary factor distinguishing the samples that underwent different steps before machining. The AB sample exhibits a conventional post-AM microstructure, characterized by equiaxed grains with an average size equivalent to the hatch distance (60 µm). Upon heat treatment, the grain morphology of the AB sample is replaced by larger, irregularly shaped grains. This results in grain coarsening and the formation of various precipitates, whose chemical composition and morphology vary depending on the adopted ageing parameters.

Table 3 displays the martensite start and finish, as well as austenite start and finish values. When considering $A_f$, which sets the temperature at which the material can be considered austenitic, the two heat-treated samples behave differently when compared to the AB sample. The T3 sample shows an increase in $A_f$ of 65.4%, while, on the contrary, the T6 sample shows a decrease of 30.8% compared to the AB one.

<figure><img src="image_1819.png" /><figcaption>Fig. 2: Optical micrographs of the AB, T3 and T6 samples.</figcaption></figure>

The different behaviour of the two heat treatments can be attributed to the types of precipitates formed because of ageing. Specifically, ageing the Ni-rich NiTi at approximately 300°C results in
the formation of small precipitates that induce an intermediate phase between austenite and martensite, known as the R-phase [11]. When aged at 600°C, different precipitates in terms of size and composition form compared to those at 300°C. The precipitates developed at 600°C promote the direct transformation between austenite and martensite, without the formation of the intermediate R-phase [6]. The presence of the R-phase results in higher transformation temperatures, which explains why the sample aged at 300°C has higher TTs than the other ones. Finally, Table 3 shows that, before machining, the material is always austenitic at room temperature regardless of its specific microstructure.

For completeness, here are reported the hardness values of the pre-machined samples according to a previous work by the authors [6], focusing on the same material and same microstructures as in this study, namely T6 was found to be the hardest sample (8.99 ± 1.24 GPa), followed by the AB sample (6.58 ± 1.38 GPa), and finally the T3 sample (3.43 ± 0.47 GPa).

<table><caption>Table 3: Transformation temperatures of the samples before machining.</caption><thead><tr><th></th><th>M<sub>s</sub><br/>[°C]</th><th>M<sub>f</sub><br/>[°C]</th><th>A<sub>s</sub><br/>[°C]</th><th>A<sub>f</sub><br/>[°C]</th></tr></thead><tbody><tr><td>AB</td><td>-8 ± 2</td><td>-65 ± 1</td><td>-18 ± 3</td><td>13 ± 0.5</td></tr><tr><td>T3</td><td>-5 ± 4</td><td>-64 ± 2</td><td>-6 ± 2</td><td>21.5 ± 1</td></tr><tr><td>T6</td><td>-17 ± 3</td><td>-63 ± 1</td><td>-15 ± 1</td><td>9 ± 2</td></tr></tbody></table>

Post-machining Characteristics. Fig. 3a displays the surface roughness $S_a$ of the AB, T3, and T6 samples machined at varying depth of cut under flood and cryogenic cooling conditions. The results indicate that the depth of cut of 0.25 mm consistently produces the highest roughness value regardless of the material's microstructure and machining condition. In the flood case, the $S_a$ values for the depths of cut at 0.1 and 0.4 mm are very similar between samples AB, T3 and T6. However, when cryogenic cooling is applied, the heat-treated samples machined at the maximum depth of cut have the lowest roughness, whereas the AB samples have the lowest roughness at $a_p$ of 0.1 mm. In addition, the samples aged at 300°C show the best surface finish under both lubricating-cooling conditions and whatever the depth of cut. It is worth noting that the T3 samples are the least affected by the machining conditions.

When comparing flood and cryogenic machining, it is evident that the former induces a better surface in terms of $S_a$, particularly at a depth of cut of 0.25 mm. Differences are also noticeable at $a_p$ of 0.1 mm, while at $a_p$ of 0.4 mm, the $S_a$ values of the flood machined samples are similar to those of the cryogenic machined ones. In the cryogenic case, the variation of $S_a$ at varying depth of cut is greater than in the flood case. For instance, under cryogenic conditions, the AB sample machined at $a_p$ of 0.4 mm shows a $S_a$ 54% lower than the AB sample machined at $a_p$ of 0.25 mm. In comparison, the same sample processed under flood conditions shows a 29% decrease in $S_a$ between the same depths of cut. The reason for the observed differences between the two machining conditions can be ascribed to the material phase during cryogenic machining. The application of liquid nitrogen during machining reduces the cutting temperatures to a great extent. Therefore, it can be assumed that the material cryogenic machined at $a_p$ of 0.25 mm is in the martensitic phase, i.e. at a temperature below $M_s$. However, as the depth of cut increases, the temperature also increases, and the NiTi is possibly machined in its austenitic phase or intermediate phase between martensite and austenite.
the formation of small precipitates that induce an intermediate phase between austenite and martensite, known as the R-phase [11]. When aged at 600°C, different precipitates in terms of size and composition form compared to those at 300°C. The precipitates developed at 600°C promote the direct transformation between austenite and martensite, without the formation of the intermediate R-phase [6]. The presence of the R-phase results in higher transformation temperatures, which explains why the sample aged at 300°C has higher TTs than the other ones. Finally, Table 3 shows that, before machining, the material is always austenitic at room temperature regardless of its specific microstructure.

For completeness, here are reported the hardness values of the pre-machined samples according to a previous work by the authors [6], focusing on the same material and same microstructures as in this study, namely T6 was found to be the hardest sample (8.99 ± 1.24 GPa), followed by the AB sample (6.58 ± 1.38 GPa), and finally the T3 sample (3.43 ± 0.47 GPa).

<table><caption>Table 3: Transformation temperatures of the samples before machining.</caption><thead><tr><th></th><th>M<sub>s</sub><br/>[°C]</th><th>M<sub>f</sub><br/>[°C]</th><th>A<sub>s</sub><br/>[°C]</th><th>A<sub>f</sub><br/>[°C]</th></tr></thead><tbody><tr><td>AB</td><td>-8 ± 2</td><td>-65 ± 1</td><td>-18 ± 3</td><td>13 ± 0.5</td></tr><tr><td>T3</td><td>-5 ± 4</td><td>-64 ± 2</td><td>-6 ± 2</td><td>21.5 ± 1</td></tr><tr><td>T6</td><td>-17 ± 3</td><td>-63 ± 1</td><td>-15 ± 1</td><td>9 ± 2</td></tr></tbody></table>

Post-machining Characteristics. Fig. 3a displays the surface roughness $S_a$ of the AB, T3, and T6 samples machined at varying depth of cut under flood and cryogenic cooling conditions. The results indicate that the depth of cut of 0.25 mm consistently produces the highest roughness value regardless of the material's microstructure and machining condition. In the flood case, the $S_a$ values for the depths of cut at 0.1 and 0.4 mm are very similar between samples AB, T3 and T6. However, when cryogenic cooling is applied, the heat-treated samples machined at the maximum depth of cut have the lowest roughness, whereas the AB samples have the lowest roughness at $a_p$ of 0.1 mm. In addition, the samples aged at 300°C show the best surface finish under both lubricating-cooling conditions and whatever the depth of cut. It is worth noting that the T3 samples are the least affected by the machining conditions.

When comparing flood and cryogenic machining, it is evident that the former induces a better surface in terms of $S_a$, particularly at a depth of cut of 0.25 mm. Differences are also noticeable at $a_p$ of 0.1 mm, while at $a_p$ of 0.4 mm, the $S_a$ values of the flood machined samples are similar to those of the cryogenic machined ones. In the cryogenic case, the variation of $S_a$ at varying depth of cut is greater than in the flood case. For instance, under cryogenic conditions, the AB sample machined at $a_p$ of 0.4 mm shows a $S_a$ 54% lower than the AB sample machined at $a_p$ of 0.25 mm. In comparison, the same sample processed under flood conditions shows a 29% decrease in $S_a$ between the same depths of cut. The reason for the observed differences between the two machining conditions can be ascribed to the material phase during cryogenic machining. The application of liquid nitrogen during machining reduces the cutting temperatures to a great extent. Therefore, it can be assumed that the material cryogenic machined at $a_p$ of 0.25 mm is in the martensitic phase, i.e. at a temperature below $M_s$. However, as the depth of cut increases, the temperature also increases, and the NiTi is possibly machined in its austenitic phase or intermediate phase between martensite and austenite.
<figure><img src="image_1822.png" /><figcaption>Fig. 3: a) Surface roughness of the AB, T3, and T6 flood and cryogenic machined samples at varying depth of cut, b) roughness profile and surface appearance of the AB samples cryogenic machined at varying depth of cut.</figcaption></figure>

Therefore, the significant difference in surface finish found in case of cryogenic machining can be attributed to the varying thermal conditions that the material undergoes during the machining itself.

Fig. 3b shows the roughness profiles of the AB samples machined under cryogenic cooling, here reported as they show the greatest variation in terms of $S_a$ at varying $a_p$, nevertheless the same trend was found in the other samples. The profiles are flanked by backscattered electron detector (BSED) images showing the characteristics and possible defects of the machined surfaces. The difference between the three roughness profiles is evident and support the following considerations: the profile obtained with $a_p$ of 0.1 mm is the most irregular and it is difficult to detect the typical periodicity of the machined profile, even if it allows a low $S_a$. On the other hand, periodicity is present in the profiles obtained at intermediate (0.25 mm) and higher (0.4 mm) depths of cut, which is also confirmed by the fact that the distance between two successive peaks is 70 µm, corresponding to a feed of 0.07 mm/rev. However, at $a_p$ of 0.25 mm the distance between
<figure><img src="image_1820.png" /><figcaption>Fig. 3: a) Surface roughness of the AB, T3, and T6 flood and cryogenic machined samples at varying depth of cut, b) roughness profile and surface appearance of the AB samples cryogenic machined at varying depth of cut.</figcaption></figure>

Therefore, the significant difference in surface finish found in case of cryogenic machining can be attributed to the varying thermal conditions that the material undergoes during the machining itself.

Fig. 3b shows the roughness profiles of the AB samples machined under cryogenic cooling, here reported as they show the greatest variation in terms of $S_a$ at varying $a_p$, nevertheless the same trend was found in the other samples. The profiles are flanked by backscattered electron detector (BSED) images showing the characteristics and possible defects of the machined surfaces. The difference between the three roughness profiles is evident and support the following considerations: the profile obtained with $a_p$ of 0.1 mm is the most irregular and it is difficult to detect the typical periodicity of the machined profile, even if it allows a low $S_a$. On the other hand, periodicity is present in the profiles obtained at intermediate (0.25 mm) and higher (0.4 mm) depths of cut, which is also confirmed by the fact that the distance between two successive peaks is 70 µm, corresponding to a feed of 0.07 mm/rev. However, at $a_p$ of 0.25 mm the distance between
peaks and valleys is more pronounced than in the profile with $a_p$ of 0.4 mm, thus demonstrating the higher $S_a$ of the former compared to the latter.

The morphology of the machined surfaces can be better understood by looking at the corresponding BSED images. In the profile obtained with the lowest depth of cut, it is difficult to distinguish the feed marks that define the tool pass from the other grooves, which explains why the periodicity of the profile is almost lost. It is therefore observed that the surface machined at $a_p$ of 0.1 mm is affected by ploughing grooves. Although the depth of cut of 0.1 mm is within the range allowed by the tool manufacturer for the adopted cutting tool, the appearance of ploughing grooves is well associated with using a depth of cut equal to the radius of the tool nose. Consequently, this phenomenon influences the components of the machining force, with the radial component surpassing the feed force, thereby impacting the surface finish of the workpiece.

In the profiles where the peaks and valleys are more pronounced, the different tool passes are clearly distinguishable by the presence of visible feed marks. This last aspect is particularly noticeable with $a_p$ of 0.25 mm, which also has the worst $S_a$ values. In the profile with the highest depth of cut, the surface roughness decreases. In fact, the distance between the valleys and peaks decreases and, consequently, the feed marks become less pronounced, and it can be also observed that the machined surface is subjected to tearing and ripping.

To enhance the comprehension of the morphology of the machined surfaces, SEM observations were conducted with the BSED detector to analyze the insert state after machining. It must be underlined that each turning operation was carried out making use of a virgin insert, therefore the wear observed on the insert is specific to that operation. Fig. 4 displays the SEM images of the flank of the inserts used to machine the T3 samples under cryogenic cooling conditions at varying depth of cut, nevertheless the tool state when machining the AB and T6 samples was the same, and therefore not here displayed. As expected from the analysis of the roughness profiles, the least worn tool is the one that worked at $a_p$ of 0.1 mm. The wear was limited to the cutting-edge area since, as mentioned earlier, the depth of cut corresponds to the radius of the tool nose. The tool's condition worsens at $a_p$ of 0.25 mm and there are also visible traces of material adhering to the workpiece. At $a_p$ of 0.4 mm, the built-up edge (BUE) and built-up layer (BUL) phenomena become even more noticeable. The first case refers to when the material adheres to the tool in the cutting zone, while in the latter case, the material adheres to the rake face of the tool. The tool geometry is altered by the presence of BUE and BUL, resulting in a better surface finish leading to a decrease in roughness, as reported in [12].

<figure><img src="image_1823.png" /><figcaption>Fig. 4: State of the tool flank faces when cryogenic machining T3 samples at varying depth of cut.</figcaption></figure>
peaks and valleys is more pronounced than in the profile with $a_p$ of 0.4 mm, thus demonstrating the higher $S_a$ of the former compared to the latter.

The morphology of the machined surfaces can be better understood by looking at the corresponding BSED images. In the profile obtained with the lowest depth of cut, it is difficult to distinguish the feed marks that define the tool pass from the other grooves, which explains why the periodicity of the profile is almost lost. It is therefore observed that the surface machined at $a_p$ of 0.1 mm is affected by ploughing grooves. Although the depth of cut of 0.1 mm is within the range allowed by the tool manufacturer for the adopted cutting tool, the appearance of ploughing grooves is well associated with using a depth of cut equal to the radius of the tool nose. Consequently, this phenomenon influences the components of the machining force, with the radial component surpassing the feed force, thereby impacting the surface finish of the workpiece.

In the profiles where the peaks and valleys are more pronounced, the different tool passes are clearly distinguishable by the presence of visible feed marks. This last aspect is particularly noticeable with $a_p$ of 0.25 mm, which also has the worst $S_a$ values. In the profile with the highest depth of cut, the surface roughness decreases. In fact, the distance between the valleys and peaks decreases and, consequently, the feed marks become less pronounced, and it can be also observed that the machined surface is subjected to tearing and ripping.

To enhance the comprehension of the morphology of the machined surfaces, SEM observations were conducted with the BSED detector to analyze the insert state after machining. It must be underlined that each turning operation was carried out making use of a virgin insert, therefore the wear observed on the insert is specific to that operation. Fig. 4 displays the SEM images of the flank of the inserts used to machine the T3 samples under cryogenic cooling conditions at varying depth of cut, nevertheless the tool state when machining the AB and T6 samples was the same, and therefore not here displayed. As expected from the analysis of the roughness profiles, the least worn tool is the one that worked at $a_p$ of 0.1 mm. The wear was limited to the cutting-edge area since, as mentioned earlier, the depth of cut corresponds to the radius of the tool nose. The tool's condition worsens at $a_p$ of 0.25 mm and there are also visible traces of material adhering to the workpiece. At $a_p$ of 0.4 mm, the built-up edge (BUE) and built-up layer (BUL) phenomena become even more noticeable. The first case refers to when the material adheres to the tool in the cutting zone, while in the latter case, the material adheres to the rake face of the tool. The tool geometry is altered by the presence of BUE and BUL, resulting in a better surface finish leading to a decrease in roughness, as reported in [12].

<figure><img src="image_1821.png" /><figcaption>Fig. 4: State of the tool flank faces when cryogenic machining T3 samples at varying depth of cut.</figcaption></figure>
<figure><img src="image_1824.png" /><figcaption>Fig. 5: Austenite finish temperature of the AB, T3, and T6 samples machined at different depths of cut under a) flood and b) cryogenic cooling conditions.</figcaption></figure>

Fig. 5 displays the austenite finish temperature ($A_f$) values, which determine whether the sample is in the austenitic phase at room temperature, i.e. when $A_f$ is approximately below 24°C. Fig. 5 shows that machining always increases $A_f$ and only in the cases of the cryogenically machined AB and T6 samples the material can be considered still austenitic at room temperature. On the contrary, flood machining resulted in non-austenitic samples at room temperature, regardless of the depth of cut and sample microstructure. This issue was also observed in the case of T3 samples that were cryogenically machined. On the other hand, the samples before machining were all austenitic at 24°C, as shown by the dotted lines in the histograms of Fig. 5.

When coming to the influence of the depth of cut, the highest increase in $A_f$ was obtained at $a_p$ of 0.25 mm in the flood conditions, regardless of the material microstructure. It is worth noting that the same trend observed for the surface roughness at varying depth of cut was found for $A_f$ in case of flood machining. On the other hand, in case of cryogenic machined samples, the influence of the depth of cut in varying $A_f$ can be considered negligible, as a consequence of the cutting temperature reduction that limits any effect on the NiTi microstructure, and thus transformation temperatures. The material microstructure is still the parameter most influencing the $A_f$ values also after machining. On the contrary, the effect of the depth of cut is more marginally evident.

## Conclusions

This study evaluates the effect of the process parameters on Ni-rich NiTi samples obtained by AM and subjected to heat treatment and finishing machining. Based on the results discussed so far, the following conclusions can be drawn:

* The microstructure is clearly different between the as-built (AB), 300°C-aged (T3) and 600°C-aged (T6) NiTi samples. Additionally, the calorimetric analysis indicates that the T3 sample has the highest austenite finish temperature, followed by the AB sample and finally the T6 one.

* Cryogenic machining produced more differentiated surface roughness results when varying the depth of cut. This behaviour can be attributed to the material phase change that can occur during cryogenic machining. In fact, at $a_p$ of 0.25 mm, the alloy is presumed to be martensitic, while at $a_p$ of 0.4 mm, it has reverted to the austenitic or quasi-austenitic phase as the temperature increases due to a greater depth of cut. Finish machining typically results in an increase in transformation temperatures, but this effect is less pronounced when cryogenic processing.

* The depth of cut of 0.25 mm yields the highest values of $S_a$, regardless of the sample's microstructure or the machining condition adopted. In fact, the roughness profiles show that $a_p$ of 0.25 mm leads to a greater distance between the peaks and valleys compared to the other
<figure><img src="image_1822.png" /><figcaption>Fig. 5: Austenite finish temperature of the AB, T3, and T6 samples machined at different depths of cut under a) flood and b) cryogenic cooling conditions.</figcaption></figure>

Fig. 5 displays the austenite finish temperature ($A_f$) values, which determine whether the sample is in the austenitic phase at room temperature, i.e. when $A_f$ is approximately below 24°C. Fig. 5 shows that machining always increases $A_f$ and only in the cases of the cryogenically machined AB and T6 samples the material can be considered still austenitic at room temperature. On the contrary, flood machining resulted in non-austenitic samples at room temperature, regardless of the depth of cut and sample microstructure. This issue was also observed in the case of T3 samples that were cryogenically machined. On the other hand, the samples before machining were all austenitic at 24°C, as shown by the dotted lines in the histograms of Fig. 5.

When coming to the influence of the depth of cut, the highest increase in $A_f$ was obtained at $a_p$ of 0.25 mm in the flood conditions, regardless of the material microstructure. It is worth noting that the same trend observed for the surface roughness at varying depth of cut was found for $A_f$ in case of flood machining. On the other hand, in case of cryogenic machined samples, the influence of the depth of cut in varying $A_f$ can be considered negligible, as a consequence of the cutting temperature reduction that limits any effect on the NiTi microstructure, and thus transformation temperatures. The material microstructure is still the parameter most influencing the $A_f$ values also after machining. On the contrary, the effect of the depth of cut is more marginally evident.

## Conclusions

This study evaluates the effect of the process parameters on Ni-rich NiTi samples obtained by AM and subjected to heat treatment and finishing machining. Based on the results discussed so far, the following conclusions can be drawn:

* The microstructure is clearly different between the as-built (AB), 300°C-aged (T3) and 600°C-aged (T6) NiTi samples. Additionally, the calorimetric analysis indicates that the T3 sample has the highest austenite finish temperature, followed by the AB sample and finally the T6 one.

* Cryogenic machining produced more differentiated surface roughness results when varying the depth of cut. This behaviour can be attributed to the material phase change that can occur during cryogenic machining. In fact, at $a_p$ of 0.25 mm, the alloy is presumed to be martensitic, while at $a_p$ of 0.4 mm, it has reverted to the austenitic or quasi-austenitic phase as the temperature increases due to a greater depth of cut. Finish machining typically results in an increase in transformation temperatures, but this effect is less pronounced when cryogenic processing.

* The depth of cut of 0.25 mm yields the highest values of $S_a$, regardless of the sample's microstructure or the machining condition adopted. In fact, the roughness profiles show that $a_p$ of 0.25 mm leads to a greater distance between the peaks and valleys compared to the other
profiles. Observations were made regarding the surface and tool conditions used for each depth of cut. The insert that worked at $a_p$ of 0.4 mm produced a better surface finish than the one that worked at $a_p$ of 0.25 mm. Regarding the austenite finish temperature, it appears that the 0.25 mm depth of cut resulted in a greater increase in $A_f$. However, this trend was only observed in the flood machining case.

## Acknowledgments

This research was developed in the framework of the PRIN project "NEMESI - 4D manufacturing based on 3D printing and machining for Nitinol biomedical and sensing applications" funded by the Italian Ministry of University and Research (MUR).

## References

[1] G.B. Kauffman, M. Isaac, The story of nitinol: the serendipitous discovery of the memory metal and its applications, The chemical educator 2 (1997): 1-21.
https://doi.org/10.1007/s00897970111a

[2] E. Patoor, D.C. Lagoudas, P.B. Entchev, L.C. Brinson, X. Gao, Shape memory alloys, Part I: General properties and modeling of single crystals, Mechanics of materials, 38.5-6 (2006), 391-429.
https://doi.org/10.1016/j.mechmat.2005.05.027

[3] M. Elahinia, N. Shayesteh Moghaddam, M. Taheri Andani, A. Amerinatanzi, B. A. Bimber, R. F. Hamilton, Fabrication of NiTi through additive manufacturing: A review, Progress in Materials Science, vol. 83. Elsevier Ltd, pp. 630-663, Oct. 01, 2016.
https://doi.org/10.1016/j.pmatsci.2016.08.001

[4] Y. Kaynak, H. E. Karaca, I. S. Jawahir, Cutting speed dependent microstructure and transformation behavior of NiTi alloy in dry and cryogenic machining, Journal of Materials Engineering and Performance 24 (2015): 452-460.
https://doi.org/10.1007/s11665-014-1247-6

[5] H. Yang, K. Sakai, H. Shizuka, Y. Kurebayashi, K. Hayakawa, T. Nagare, Effect of cutting speed on shape recovery of work material in cutting process of super-elastic NiTi alloy, International Journal of Automation Technology, 15.1 (2021), 24-33.
https://doi.org/10.20965/ijat.2021.p0024

[6] R. Bertolini, S. Bruschi, A. Ghiotti, E. Savio, L. Ceseracciu, I.S. Jawahir, Surface integrity and superelastic response of additively manufactured Nitinol after heat treatment and finish machining, CIRP Annals (2023).
https://doi.org/10.1016/j.cirp.2023.04.025

[7] S. Khademzadeh, F. Zanini, J. Rocco, K. Brunelli, P.F. Bariani, S. Carmignato, Quality enhancement of microstructure and surface topography of NiTi parts produced by laser powder bed fusion, CIRP Journal of Manufacturing Science and Technology 31 (2020), 575-582.
https://doi.org/10.1016/j.cirp.2023.04.025

[8] A. Bordin, S. Bruschi, A. Ghiotti, P.F. Bariani, Analysis of tool wear in cryogenic machining of additive manufactured Ti6Al4V alloy, Wear 328 (2015), 89-99.
https://doi.org/10.1016/j.wear.2015.01.030

[9] ASTM F2004-17: Standard Test Method for Transformation Temperature of Nickel-Titanium Alloys by Thermal Analysis.

[10] BS EN ISO 25178-2:2022: Geometrical product specifications (GPS). Surface texture: Areal. Terms, definitions and surface texture parameters.
profiles. Observations were made regarding the surface and tool conditions used for each depth of cut. The insert that worked at $a_p$ of 0.4 mm produced a better surface finish than the one that worked at $a_p$ of 0.25 mm. Regarding the austenite finish temperature, it appears that the 0.25 mm depth of cut resulted in a greater increase in $A_f$. However, this trend was only observed in the flood machining case.

## Acknowledgments

This research was developed in the framework of the PRIN project "NEMESI - 4D manufacturing based on 3D printing and machining for Nitinol biomedical and sensing applications" funded by the Italian Ministry of University and Research (MUR).

## References

[1] G.B. Kauffman, M. Isaac, The story of nitinol: the serendipitous discovery of the memory metal and its applications, The chemical educator 2 (1997): 1-21.
https://doi.org/10.1007/s00897970111a

[2] E. Patoor, D.C. Lagoudas, P.B. Entchev, L.C. Brinson, X. Gao, Shape memory alloys, Part I: General properties and modeling of single crystals, Mechanics of materials, 38.5-6 (2006), 391-429.
https://doi.org/10.1016/j.mechmat.2005.05.027

[3] M. Elahinia, N. Shayesteh Moghaddam, M. Taheri Andani, A. Amerinatanzi, B. A. Bimber, R. F. Hamilton, Fabrication of NiTi through additive manufacturing: A review, Progress in Materials Science, vol. 83. Elsevier Ltd, pp. 630-663, Oct. 01, 2016.
https://doi.org/10.1016/j.pmatsci.2016.08.001

[4] Y. Kaynak, H. E. Karaca, I. S. Jawahir, Cutting speed dependent microstructure and transformation behavior of NiTi alloy in dry and cryogenic machining, Journal of Materials Engineering and Performance 24 (2015): 452-460.
https://doi.org/10.1007/s11665-014-1247-6

[5] H. Yang, K. Sakai, H. Shizuka, Y. Kurebayashi, K. Hayakawa, T. Nagare, Effect of cutting speed on shape recovery of work material in cutting process of super-elastic NiTi alloy, International Journal of Automation Technology, 15.1 (2021), 24-33.
https://doi.org/10.20965/ijat.2021.p0024

[6] R. Bertolini, S. Bruschi, A. Ghiotti, E. Savio, L. Ceseracciu, I.S. Jawahir, Surface integrity and superelastic response of additively manufactured Nitinol after heat treatment and finish machining, CIRP Annals (2023).
https://doi.org/10.1016/j.cirp.2023.04.025

[7] S. Khademzadeh, F. Zanini, J. Rocco, K. Brunelli, P.F. Bariani, S. Carmignato, Quality enhancement of microstructure and surface topography of NiTi parts produced by laser powder bed fusion, CIRP Journal of Manufacturing Science and Technology 31 (2020), 575-582.
https://doi.org/10.1016/j.cirp.2023.04.025

[8] A. Bordin, S. Bruschi, A. Ghiotti, P.F. Bariani, Analysis of tool wear in cryogenic machining of additive manufactured Ti6Al4V alloy, Wear 328 (2015), 89-99.
https://doi.org/10.1016/j.wear.2015.01.030

[9] ASTM F2004-17: Standard Test Method for Transformation Temperature of Nickel-Titanium Alloys by Thermal Analysis.

[10] BS EN ISO 25178-2:2022: Geometrical product specifications (GPS). Surface texture: Areal. Terms, definitions and surface texture parameters.
[11] X. B Wang, B. Verlinden, J. Van Humbeeck, R-phase transformation in NiTi alloys, Materials Science and Technology (United Kingdom), vol. 30, no. 13, pp. 1517-1529 (2014). https://doi.org/10.1179/1743284714Y.0000000590

[12] R. Bertolini, A. Ghiotti, S. Bruschi, Graphene nanoplatelets as additives to MQL for improving tool life in machining Inconel 718 alloy, Wear 476 (2021), 203656. https://doi.org/10.1016/j.wear.2021.203656
[11] X. B Wang, B. Verlinden, J. Van Humbeeck, R-phase transformation in NiTi alloys, Materials Science and Technology (United Kingdom), vol. 30, no. 13, pp. 1517-1529 (2014). https://doi.org/10.1179/1743284714Y.0000000590

[12] R. Bertolini, A. Ghiotti, S. Bruschi, Graphene nanoplatelets as additives to MQL for improving tool life in machining Inconel 718 alloy, Wear 476 (2021), 203656. https://doi.org/10.1016/j.wear.2021.203656
# Fatigue strength of Ti6Al4V titanium alloy machined under flood and cryogenic conditions

STRAMARE Andrea¹,ª, BERTOLINI Rachele¹,ᵇ*, BRUSCHI Stefania¹,ᶜ, GHIOTTI Andrea¹,ᵈ and CAMPAGNOLO Alberto¹,ᵉ

¹Department of Industrial Engineering, University of Padova, Via Venezia 1, 35131 Padova, Italy

ªandrea.stramare@unipd.it, ᵇrachele.bertolini@unipd.it, ᶜstefania.bruschi@unipd.it,

ᵈandrea.ghiotti@unipd.it, ᶜalberto.campagnolo@unipd.it

**Keywords:** Ti6Al4V, Cryogenic Machining, Surface Integrity, Fatigue Strength

**Abstract.** Ti6Al4V titanium alloy wrought round bars characterized by a fine equiaxed microstructure were machined using two different lubrication strategies, namely traditional water-based cutting fluid and liquid nitrogen. Surface integrity after machining was assessed in terms of (i) sub-surface severe plastic deformation layer, (ii) surface topography, and (iii) residual axial stresses as a function of the distance from the machined surface. Finally, the machined specimens were subjected to fatigue tests on a four-point rotating bending machine. For the two sets of specimens, the Wöhler curves were derived to determine their fatigue strength using the modified staircase method. No remarkable difference in terms of surface topography was found between the flood and cryogenic machined specimens, but the latter were characterized by higher and deeper compressive residual stresses and higher severe plastic deformation layer beneath the machined surface. Despite that, it is shown that the fatigue limit of the flood and cryogenic machined specimens is very similar, as well as crack initiation sites have the same location. An explanation of that can be found in the very fine microstructure of the as-received material, which prevails over the other characteristics induced by any machining strategy.

**Introduction**

Ti6Al4V titanium alloy is widely used in different fields, thanks to its high strength-to-weight ratio, corrosion resistance, and biocompatibility which make it suitable for aerospace and biomedical applications [1]. The last years have seen the ever-increasing use of additive manufacturing (AM) technologies to manufacture parts made of Ti6Al4V, thanks to the advantages AM can offer, mainly in terms of “complexity for free” fabrication. Nevertheless, conventional processing routes including hot forging of the wrought alloy followed by machining operations still remain the most reliable ones since they ensure the absence of internal defects like porosities characterizing AM parts, and, at the same time, offer the possibility of a precise tailoring of the part microstructure through adequate thermo-mechanical treatments [2]. Microstructure plays indeed a crucial role in assessing the part in-service characteristics.

When referring to fatigue life, the final machining steps of the processing route have the highest impact, as they may induce significant variations in the part surface integrity. It is well-known that the higher the latter the higher the part fatigue resistance, which has driven the development of innovative cutting technologies capable of assuring enhanced surface properties. Among these technologies, cryogenic machining making use of liquid nitrogen as cooling medium has recently emerged as a suitable strategy to improve the surface finish and increase surface hardness over conventional flood machining [3]. While the advantages offered by cryogenic machining in increasing the part surface integrity are well-established, few studies are available in literature that assess its impact on the part fatigue resistance. In [4] the fatigue life of a wrought Ti6Al4V with an equiaxed structure was found to be higher when machined under cryogenic cooling conditions compared to dry and conventionally lubricated conditions as a consequence of grain refinement

<figure><img src="image_1825.png" /></figure>
# Fatigue strength of Ti6Al4V titanium alloy machined under flood and cryogenic conditions

STRAMARE Andrea¹,ª, BERTOLINI Rachele¹,ᵇ*, BRUSCHI Stefania¹,ᶜ, GHIOTTI Andrea¹,ᵈ and CAMPAGNOLO Alberto¹,ᵉ

¹Department of Industrial Engineering, University of Padova, Via Venezia 1, 35131 Padova, Italy

ªandrea.stramare@unipd.it, ᵇrachele.bertolini@unipd.it, ᶜstefania.bruschi@unipd.it,

ᵈandrea.ghiotti@unipd.it, ᶜalberto.campagnolo@unipd.it

**Keywords:** Ti6Al4V, Cryogenic Machining, Surface Integrity, Fatigue Strength

**Abstract.** Ti6Al4V titanium alloy wrought round bars characterized by a fine equiaxed microstructure were machined using two different lubrication strategies, namely traditional water-based cutting fluid and liquid nitrogen. Surface integrity after machining was assessed in terms of (i) sub-surface severe plastic deformation layer, (ii) surface topography, and (iii) residual axial stresses as a function of the distance from the machined surface. Finally, the machined specimens were subjected to fatigue tests on a four-point rotating bending machine. For the two sets of specimens, the Wöhler curves were derived to determine their fatigue strength using the modified staircase method. No remarkable difference in terms of surface topography was found between the flood and cryogenic machined specimens, but the latter were characterized by higher and deeper compressive residual stresses and higher severe plastic deformation layer beneath the machined surface. Despite that, it is shown that the fatigue limit of the flood and cryogenic machined specimens is very similar, as well as crack initiation sites have the same location. An explanation of that can be found in the very fine microstructure of the as-received material, which prevails over the other characteristics induced by any machining strategy.

**Introduction**

Ti6Al4V titanium alloy is widely used in different fields, thanks to its high strength-to-weight ratio, corrosion resistance, and biocompatibility which make it suitable for aerospace and biomedical applications [1]. The last years have seen the ever-increasing use of additive manufacturing (AM) technologies to manufacture parts made of Ti6Al4V, thanks to the advantages AM can offer, mainly in terms of “complexity for free” fabrication. Nevertheless, conventional processing routes including hot forging of the wrought alloy followed by machining operations still remain the most reliable ones since they ensure the absence of internal defects like porosities characterizing AM parts, and, at the same time, offer the possibility of a precise tailoring of the part microstructure through adequate thermo-mechanical treatments [2]. Microstructure plays indeed a crucial role in assessing the part in-service characteristics.

When referring to fatigue life, the final machining steps of the processing route have the highest impact, as they may induce significant variations in the part surface integrity. It is well-known that the higher the latter the higher the part fatigue resistance, which has driven the development of innovative cutting technologies capable of assuring enhanced surface properties. Among these technologies, cryogenic machining making use of liquid nitrogen as cooling medium has recently emerged as a suitable strategy to improve the surface finish and increase surface hardness over conventional flood machining [3]. While the advantages offered by cryogenic machining in increasing the part surface integrity are well-established, few studies are available in literature that assess its impact on the part fatigue resistance. In [4] the fatigue life of a wrought Ti6Al4V with an equiaxed structure was found to be higher when machined under cryogenic cooling conditions compared to dry and conventionally lubricated conditions as a consequence of grain refinement

<figure><img src="image_1823.png" /></figure>