# The effect of cryogenic assistance on the machinability of Ti54M and energetic approach

EL HATIMI Imane¹,ᵃ*, WAGNER Vincent¹,ᵇ and DESSEIN Gilles¹,ᶜ

¹Laboratoire Génie de Production, Ecole Nationale d'Ingénieurs de Tarbes, Université de Toulouse, Tarbes Cedex, France

ªimane.el_hatimi@enit.fr, ᵇvincent.wagner@enit.fr, ᶜgilles.dessein@enit.fr

**Keywords:** Cryogenic Assistance, Machining Titanium, Thermal Effect

**Abstract.** Titanium alloys are increasingly being used in various applications due to their excellent mechanical properties and corrosion resistance. However, their machinability can be challenging due to their high strength and low thermal conductivity. Cryogenic assistance is a promising technique that can be used to improve the machinability of titanium alloys by reducing the cutting forces and improving the tool life. This study investigates the effect of cryogenic assistance on the machining of Ti54M, a high-strength titanium alloy. The results show that cryogenic assistance can significantly reduce the cutting forces and improve the tool life. However, the benefits of cryogenic assistance are limited at high cutting speeds. The study also provides insights into the mechanisms by which cryogenic assistance improves machinability.

**Introduction**

In the realm of advanced manufacturing and precision engineering, titanium alloys have firmly established themselves as indispensable materials. Over time, a multitude of alloys have been developed, each displaying varying degrees of machinability. Renowned for their exceptional traits of lightweightness, corrosion resistance, and mechanical robustness, these alloys have emerged as the preferred materials across a spectrum of industries, including aerospace, medical, automotive, and beyond. However, they are not without their challenges, particularly when it comes to machining them into intricate components and high-precision parts. This is where cryogenic assistance comes into play, acting as a bridge between the capabilities of titanium alloys and the mastery of their machinability.

Machinability, which encompasses the ability to effectively shape a material while maintaining high quality, is a central concern for engineers and manufacturers. Titanium alloys, renowned for their lightweight properties and high-temperature stability [1], such as Ti54M, distinguished by its composition and mechanical properties [2], have long been cherished for their exceptional characteristics. However, their machinability can pose challenges due to their dense and resilient metallic nature. This is where cryogenic assistance, a technology based on the use of extremely low temperatures, comes into play to push the boundaries of titanium alloy machinability.

In this research article, we delve into the intersection of cryogenic assistance and the machinability of titanium alloys. We will explore in detail how the application of cryogenic temperatures can influence the properties of titanium alloys. We will demonstrate how cryogenic assistance can address some of the most complex issues related to machinability, including managing heat generated during machining, extending tool life, and enhancing the quality of machined parts.

**Cryogenic assistance**

Cryogenic assistance involves utilizing extremely low temperatures, typically below -150°C (-238°F), across various fields to achieve specific objectives. This technology is rooted in the

<figure><img src="image_1969.png" /></figure>
# The effect of cryogenic assistance on the machinability of Ti54M and energetic approach

EL HATIMI Imane¹,ᵃ*, WAGNER Vincent¹,ᵇ and DESSEIN Gilles¹,ᶜ

¹Laboratoire Génie de Production, Ecole Nationale d'Ingénieurs de Tarbes, Université de Toulouse, Tarbes Cedex, France

ªimane.el_hatimi@enit.fr, ᵇvincent.wagner@enit.fr, ᶜgilles.dessein@enit.fr

**Keywords:** Cryogenic Assistance, Machining Titanium, Thermal Effect

**Abstract.** Titanium alloys are increasingly being used in various applications due to their excellent mechanical properties and corrosion resistance. However, their machinability can be challenging due to their high strength and low thermal conductivity. Cryogenic assistance is a promising technique that can be used to improve the machinability of titanium alloys by reducing the cutting forces and improving the tool life. This study investigates the effect of cryogenic assistance on the machining of Ti54M, a high-strength titanium alloy. The results show that cryogenic assistance can significantly reduce the cutting forces and improve the tool life. However, the benefits of cryogenic assistance are limited at high cutting speeds. The study also provides insights into the mechanisms by which cryogenic assistance improves machinability.

## Introduction

In the realm of advanced manufacturing and precision engineering, titanium alloys have firmly established themselves as indispensable materials. Over time, a multitude of alloys have been developed, each displaying varying degrees of machinability. Renowned for their exceptional traits of lightweightness, corrosion resistance, and mechanical robustness, these alloys have emerged as the preferred materials across a spectrum of industries, including aerospace, medical, automotive, and beyond. However, they are not without their challenges, particularly when it comes to machining them into intricate components and high-precision parts. This is where cryogenic assistance comes into play, acting as a bridge between the capabilities of titanium alloys and the mastery of their machinability.

Machinability, which encompasses the ability to effectively shape a material while maintaining high quality, is a central concern for engineers and manufacturers. Titanium alloys, renowned for their lightweight properties and high-temperature stability [1], such as Ti54M, distinguished by its composition and mechanical properties [2], have long been cherished for their exceptional characteristics. However, their machinability can pose challenges due to their dense and resilient metallic nature. This is where cryogenic assistance, a technology based on the use of extremely low temperatures, comes into play to push the boundaries of titanium alloy machinability.

In this research article, we delve into the intersection of cryogenic assistance and the machinability of titanium alloys. We will explore in detail how the application of cryogenic temperatures can influence the properties of titanium alloys. We will demonstrate how cryogenic assistance can address some of the most complex issues related to machinability, including managing heat generated during machining, extending tool life, and enhancing the quality of machined parts.

## Cryogenic assistance

Cryogenic assistance involves utilizing extremely low temperatures, typically below -150°C (-238°F), across various fields to achieve specific objectives. This technology is rooted in the

<figure><img src="image_1964.png" /></figure>
principles of cryogenics, which entail the study and manipulation of gases and liquids at these exceedingly low temperatures [3].

Cryogenic assistance relies on cooling substances to extremely low temperatures. At such frigid levels, many materials exhibit distinct behaviors and unique properties that can be harnessed for specific purposes. It finds applications in diverse fields, including scientific research, industry, medicine, aerospace, and energy, facilitating activities such as biological sample preservation, particle physics research, electronic component manufacturing, and energy production.

In addition to significantly enhancing machining efficiency and quality, cryogenic assistance also offers notable energy advantages. Employing cryogenic temperatures in machining processes reduces thermal stress on both the material and the cutting tools, resulting in decreased energy consumption compared to conventional machining methods. This energy-efficient approach aligns with global sustainability efforts and resource conservation, making it an attractive option for industries seeking eco-friendly manufacturing solutions. This article will not only explore technological advancements in titanium alloy machining but also underscore the critical role cryogenic assistance plays in achieving these advancements, with a focus on sustainability and energy efficiency.

While the benefits of cryogenic assistance on titanium alloys are well established, its impact on the cutting process remains challenging to study in detail. While such assistance may positively influence wear, it can also pose challenges in chip formation. As elucidated in Komanduri's work, based on Recht's findings, the cutting process and the creation of shear bands result from the competition between material work-hardening due to high deformation and a sharp increase in temperature leading to significant thermal softening.

The underlying question pertains to the adverse effects of cryogenic assistance in reducing thermal impact, a phenomenon that is not straightforward to characterize.

In summary, cryogenic assistance stands as a powerful technology that has revolutionized various areas of science and industry by enabling the use of extremely low temperatures for a multitude of applications. It opens doors to innovative possibilities and continues to inspire research and technological advancements. Although the behavior of materials under study may vary significantly, the objective of this article is to examine how cryogenic assistance can impact the cutting process and provide potential solutions through a comprehensive analysis.

## Experimental set-up

Presentation of titanium alloys.

In this scientific work, the testes were made on Ti54M because of its good machinability. Table 1 and Table 2 exposes the chemical composition and a comprehensive overview of the mechanical and thermal properties of this alloy.

<table><caption>Table 1: chemical composition and Ti54M [1].</caption><thead><tr><th></th><th>Al</th><th>Mo</th><th>V</th><th>Fe</th><th>O</th></tr></thead><tbody><tr><th>Ti54M</th><td>5</td><td>0.8</td><td>4</td><td>0.5</td><td>0.16</td></tr></tbody></table>
principles of cryogenics, which entail the study and manipulation of gases and liquids at these exceedingly low temperatures [3].

Cryogenic assistance relies on cooling substances to extremely low temperatures. At such frigid levels, many materials exhibit distinct behaviors and unique properties that can be harnessed for specific purposes. It finds applications in diverse fields, including scientific research, industry, medicine, aerospace, and energy, facilitating activities such as biological sample preservation, particle physics research, electronic component manufacturing, and energy production.

In addition to significantly enhancing machining efficiency and quality, cryogenic assistance also offers notable energy advantages. Employing cryogenic temperatures in machining processes reduces thermal stress on both the material and the cutting tools, resulting in decreased energy consumption compared to conventional machining methods. This energy-efficient approach aligns with global sustainability efforts and resource conservation, making it an attractive option for industries seeking eco-friendly manufacturing solutions. This article will not only explore technological advancements in titanium alloy machining but also underscore the critical role cryogenic assistance plays in achieving these advancements, with a focus on sustainability and energy efficiency.

While the benefits of cryogenic assistance on titanium alloys are well established, its impact on the cutting process remains challenging to study in detail. While such assistance may positively influence wear, it can also pose challenges in chip formation. As elucidated in Komanduri's work, based on Recht's findings, the cutting process and the creation of shear bands result from the competition between material work-hardening due to high deformation and a sharp increase in temperature leading to significant thermal softening.

The underlying question pertains to the adverse effects of cryogenic assistance in reducing thermal impact, a phenomenon that is not straightforward to characterize.

In summary, cryogenic assistance stands as a powerful technology that has revolutionized various areas of science and industry by enabling the use of extremely low temperatures for a multitude of applications. It opens doors to innovative possibilities and continues to inspire research and technological advancements. Although the behavior of materials under study may vary significantly, the objective of this article is to examine how cryogenic assistance can impact the cutting process and provide potential solutions through a comprehensive analysis.

## Experimental set-up

Presentation of titanium alloys.

In this scientific work, the testes were made on Ti54M because of its good machinability. Table 1 and Table 2 exposes the chemical composition and a comprehensive overview of the mechanical and thermal properties of this alloy.

<table><caption>Table 1: chemical composition and Ti54M [1].</caption><thead><tr><th></th><th>Al</th><th>Mo</th><th>V</th><th>Fe</th><th>O</th></tr></thead><tbody><tr><td>Ti54M</td><td>5</td><td>0.8</td><td>4</td><td>0.5</td><td>0.16</td></tr></tbody></table>
<table><caption>Table 2 mechanical properties of Ti54M [1].</caption><thead><tr><td>Properties</td><td>Ti54M</td></tr></thead><tbody><tr><td>Modulus of elasticity E (GPa)</td><td>115</td></tr><tr><td>Tensile strength Re (MPa)</td><td>920</td></tr><tr><td>Yield strength Rp (MPa)</td><td>839</td></tr><tr><td>Thermal conductivity λ (W/m.K)</td><td>6 - 7</td></tr><tr><td>Specific heat capacity Cp (J/kg.K)</td><td>532</td></tr><tr><td>Temperature transus Tβ (°C)</td><td>965</td></tr></tbody></table>

Based on the experimental setup depicted in Fig. 1, a three-axis lathe, CNC SOMAB, was utilized, and a coated carbide tool CCMX 12 04 08 was selected, resulting in a rake angle of α = 15°.

Orthogonal cutting tests were conducted, during which the cutting force Fc and the feed force Ft were measured using a three-component piezoelectric dynamometer.

To achieve an orthogonal configuration, a groove was machined so that the nose radius does not cut. The edge inclination angle is set to 90°.

Experiments were carried out at various cutting speeds VC and feed rates f. The mean values of the machining forces, representing the average behavior of the respective force components in the steady state, were extracted. All cutting conditions were repeated three times, and the values presented in this article are the means of these three tests.

Regarding cryogenic assistance, Ln2 was consistently applied at the tool chip interface at the same flood rate.

Metallographic analysis was performed to comprehend the phenomena. This analysis allows for the examination of the distribution and configuration of phases, grains, and other microstructural features present in the alloy. The analysis consists of several steps: First, the chips are embedded with a thermosetting resin or polymer at elevated temperatures, then polished using SiC and CHem papers to achieve a perfectly flat, scratch-free surface, and finally etched in hydrofluoric acid to reveal the microstructure.

<figure><img src="image_1970.png" /><figcaption>Figure 1: the experimental set up during the orthogonal cutting.</figcaption></figure>
<table><caption>Table 2 mechanical properties of Ti54M [1].</caption><thead><tr><td>Proprieties</td><td>Ti54M</td></tr></thead><tbody><tr><td>Modulus of elasticity E (GPa)</td><td>115</td></tr><tr><td>Tensile strength Re (MPa)</td><td>920</td></tr><tr><td>Yield strength Rp (MPa)</td><td>839</td></tr><tr><td>Thermal conductivity λ (W/m.K)</td><td>6 - 7</td></tr><tr><td>Specific heat capacity Cp (J/kg.K)</td><td>532</td></tr><tr><td>Temperature transus Tβ (°C)</td><td>965</td></tr></tbody></table>

Based on the experimental setup depicted in Fig. 1, a three-axis lathe, CNC SOMAB, was utilized, and a coated carbide tool CCMX 12 04 08 was selected, resulting in a rake angle of α = 15°.

Orthogonal cutting tests were conducted, during which the cutting force Fc and the feed force Ft were measured using a three-component piezoelectric dynamometer.

To achieve an orthogonal configuration, a groove was machined so that the nose radius does not cut. The edge inclination angle is set to 90°.

Experiments were carried out at various cutting speeds VC and feed rates f. The mean values of the machining forces, representing the average behavior of the respective force components in the steady state, were extracted. All cutting conditions were repeated three times, and the values presented in this article are the means of these three tests.

Regarding cryogenic assistance, Ln2 was consistently applied at the tool chip interface at the same flood rate.

Metallographic analysis was performed to comprehend the phenomena. This analysis allows for the examination of the distribution and configuration of phases, grains, and other microstructural features present in the alloy. The analysis consists of several steps: First, the chips are embedded with a thermosetting resin or polymer at elevated temperatures, then polished using SiC and CHem papers to achieve a perfectly flat, scratch-free surface, and finally etched in hydrofluoric acid to reveal the microstructure.

<figure><img src="image_1965.png" /><figcaption>Figure 1: the experimental set up during the orthogonal cutting.</figcaption></figure>
<table><caption>Table 3: conditions of orthogonal cutting tests performed for Ti54M.</caption><thead><tr><th>Test</th><th>V<sub>c</sub> (m/min)</th><th>f (mm)</th><th>A<sub>p</sub> (mm)</th><th>α (°)</th></tr></thead><tbody><tr><td>1</td><td>30</td><td>0.15</td><td>2</td><td>15</td></tr><tr><td>2</td><td>40</td><td>0.1</td><td>2</td><td>15</td></tr><tr><td>3</td><td>40</td><td>0.15</td><td>2</td><td>15</td></tr><tr><td>4</td><td>40</td><td>0.2</td><td>2</td><td>15</td></tr><tr><td>5</td><td>40</td><td>0.25</td><td>2</td><td>15</td></tr><tr><td>6</td><td>40</td><td>0.3</td><td>2</td><td>15</td></tr><tr><td>7</td><td>60</td><td>0.15</td><td>2</td><td>15</td></tr><tr><td>8</td><td>80</td><td>0.15</td><td>2</td><td>15</td></tr><tr><td>9</td><td>100</td><td>0.15</td><td>2</td><td>15</td></tr></tbody></table>

Experimental results

As shown in Figure 2, cutting forces with cryogenic assistance are lower than with dry machining, since cryogenic assistance involves the use of cryogenic liquids to actively cool the cutting zone. Concerning the dry cutting tests, cryogenic assistance seems to have a greater impact on cutting forces at lower cutting speeds. This effect is higher for the Fc cutting force component. This can be explained by the highest temperatures got when the cutting speed increase.

<figure><img src="image_1971.png" /><figcaption>Figure 2: Effect of cryogenic assistance on cutting forces</figcaption></figure>

These results may not suffice to fully define the effect of cryogenic assistance on these titanium alloys, prompting a deeper exploration of other parameters. These parameters not only investigate
<table><caption>Table 3: conditions of orthogonal cutting tests performed for Ti54M.</caption><thead><tr><th>Test</th><th>V<sub>c</sub> (m/min)</th><th>f (mm)</th><th>A<sub>p</sub> (mm)</th><th>α (°)</th></tr></thead><tbody><tr><td>1</td><td>30</td><td>0.15</td><td>2</td><td>15</td></tr><tr><td>2</td><td>40</td><td>0.1</td><td>2</td><td>15</td></tr><tr><td>3</td><td>40</td><td>0.15</td><td>2</td><td>15</td></tr><tr><td>4</td><td>40</td><td>0.2</td><td>2</td><td>15</td></tr><tr><td>5</td><td>40</td><td>0.25</td><td>2</td><td>15</td></tr><tr><td>6</td><td>40</td><td>0.3</td><td>2</td><td>15</td></tr><tr><td>7</td><td>60</td><td>0.15</td><td>2</td><td>15</td></tr><tr><td>8</td><td>80</td><td>0.15</td><td>2</td><td>15</td></tr><tr><td>9</td><td>100</td><td>0.15</td><td>2</td><td>15</td></tr></tbody></table>

Experimental results

As shown in Figure 2, cutting forces with cryogenic assistance are lower than with dry machining, since cryogenic assistance involves the use of cryogenic liquids to actively cool the cutting zone. Concerning the dry cutting tests, cryogenic assistance seems to have a greater impact on cutting forces at lower cutting speeds. This effect is higher for the Fc cutting force component. This can be explained by the highest temperatures got when the cutting speed increase.

<figure><img src="image_1966.png" /><figcaption>Figure 2: Effect of cryogenic assistance on cutting forces</figcaption></figure>

These results may not suffice to fully define the effect of cryogenic assistance on these titanium alloys, prompting a deeper exploration of other parameters. These parameters not only investigate
the influence of cryogenic assistance but also aid in analyzing the cutting process. Parameters such as friction coefficient, shear plane stress, and shear angles warrant consideration.

The cutting process yields shear zones, which are categorized into three zones: the primary ZCP zone (plasticity), the secondary ZCS zone (friction), and the tertiary zone (elastic return) [4], as illustrated in Figure 3.

<figure><img src="image_1972.png" /><figcaption>Figure 3: Deformation zones in orthogonal cutting</figcaption></figure>

The coefficient of friction in machining is a parameter that measures the resistance to sliding between two surfaces in contact during the machining process. It represents the force of friction between the cutting tool and the workpiece being machined, and can have a significant impact on machining performance, finished part quality and tool life.

$$ \mu = \frac{F_f}{F_t} \qquad (1) $$

F_N and F_f are the normal and friction forces in the measurement frame [5]:

$$ F_f = F_c \sin\alpha + F_T \cos\alpha \qquad (2) $$

$$ F_N = F_c \cos\alpha - F_T \sin\alpha \qquad (3) $$

Shear angles, also known as plastic deformation angle, is a parameter used to describe the deformation of a material when subjected to shear stress. It measures the angle between the direction of the applied shear stress and the direction of the resulting deformation in the material. It is given by [5]:

$$ \phi = \tan^{-1}\left(\frac{\frac{tu}{tc}\cos(\alpha)}{1-\frac{tu}{tc}\sin(\alpha)}\right) \qquad (4) $$

The shear band stress is the resultant stress generated in the zone AB, shown in Figure 3, of plastic deformation and is given by equation [5]:

$$ \bar{\sigma}_{AB} = \sqrt{3}k_{AB} \quad \text{where} \quad k_{AB} = \frac{F_s \sin\phi}{t_u w} \qquad (5) $$

With the aid of these equations and the cutting forces measured by Kistler, the calculations of these parameters can be carried out as depicted in Figure 4. This progression is notably intriguing. Despite the evolving nature of the cutting forces, there exists a distinct contrast in the coefficients of friction, with cryogenic assistance markedly reducing this coefficient. This disparity can be
the influence of cryogenic assistance but also aid in analyzing the cutting process. Parameters such as friction coefficient, shear plane stress, and shear angles warrant consideration.

The cutting process yields shear zones, which are categorized into three zones: the primary ZCP zone (plasticity), the secondary ZCS zone (friction), and the tertiary zone (elastic return) [4], as illustrated in Figure 3.

<figure><img src="image_1967.png" /><figcaption>Figure 3: Deformation zones in orthogonal cutting</figcaption></figure>

The coefficient of friction in machining is a parameter that measures the resistance to sliding between two surfaces in contact during the machining process. It represents the force of friction between the cutting tool and the workpiece being machined, and can have a significant impact on machining performance, finished part quality and tool life.

$$ \mu = \frac{F_f}{F_t} \qquad (1) $$

F_N and F_f are the normal and friction forces in the measurement frame [5]:

$$ F_f = F_c \sin\alpha + F_T \cos\alpha \qquad (2) $$

$$ F_N = F_c \cos\alpha - F_T \sin\alpha \qquad (3) $$

Shear angles, also known as plastic deformation angle, is a parameter used to describe the deformation of a material when subjected to shear stress. It measures the angle between the direction of the applied shear stress and the direction of the resulting deformation in the material. It is given by [5]:

$$ \phi = \tan^{-1}\left(\frac{\frac{tu}{tc}\cos(\alpha)}{1-\frac{tu}{tc}\sin(\alpha)}\right) \qquad (4) $$

The shear band stress is the resultant stress generated in the zone AB, shown in Figure 3, of plastic deformation and is given by equation [5]:

$$ \bar{\sigma}_{AB} = \sqrt{3}k_{AB} \quad \text{where} \quad k_{AB} = \frac{F_s \sin\phi}{t_u w} \qquad (5) $$

With the aid of these equations and the cutting forces measured by Kistler, the calculations of these parameters can be carried out as depicted in Figure 4. This progression is notably intriguing. Despite the evolving nature of the cutting forces, there exists a distinct contrast in the coefficients of friction, with cryogenic assistance markedly reducing this coefficient. This disparity can be
elucidated by the contact phenomena at the tool-chip interface. With cryogenic assistance, it appears that the contacts are no longer adhesive.

Regarding stress, the trends exhibit variation. Below 60 m/min, the stress in the primary shear plane is lowest in the dry condition. However, this trend reverses above 60 m/min. This inflection point may be attributed to the heightened temperature in the shear plane due to the increased cutting speed and the limited effect of cryogenic assistance.

<figure><img src="image_1973.png" /><figcaption>Figure 4 : Evolution of strain in primary shear bands and friction coefficient according to cutting conditions for Ti54M</figcaption></figure>

The fundamental question revolves around how cryogenic assistance can affect the cutting process to varying degrees. As illustrated in Figure 4, it appears that the assistance can impact both the secondary and primary zones. However, discerning the proportion of the assistance's effect and its influence in these two zones proves challenging. To address this, we have differentiated the energy consumed during cutting for each zone.

The energy supplied (Wext), primarily transformed into temperature, is the product of cutting force and cutting speed. Consequently, by disregarding the energy in the tertiary zone, this energy can be divided into two components: the energy in the primary zone (W1) and the energy in the secondary zone (W2). Recognizing that temperature influences forces through thermal softening, a cooling source can alter this energy balance. This forms the focus of the subsequent stage of analysis. Furthermore, it is feasible to discern the phenomena by analyzing the energy in both shear zones.

$$W_{ext} = F_{c}V_{c} \qquad (6)$$

$$W_{1} = F_{s}V_{s} \qquad (7)$$

$$W_{2} = F_{n}V_{chip} \qquad (8)$$

Where $V_{chip}$ is the chip velocity and $V_s$ the shearing velocity.

These results are indeed intriguing, shedding light on the challenge of analyzing results solely based on stresses, strains, or friction coefficients. Initially, the external energy increases linearly and seems to remain consistent with or without assistance. This phenomenon can be attributed to the rise in cutting speed.

Secondly, it becomes apparent that in dry tests, the energy in the primary bands consistently remains lower. There are two plausible interpretations. The first suggests that the temperature generated by cutting constrains forces and consequently energy. This interpretation aligns with Komanduri's findings, illustrating the role of temperature in shear band formation. Alternatively, the cooling effect of the assistance appears to be efficacious. The separation of material in the primary band demands more energy and is thus less susceptible to temperature effects.
elucidated by the contact phenomena at the tool-chip interface. With cryogenic assistance, it appears that the contacts are no longer adhesive.

Regarding stress, the trends exhibit variation. Below 60 m/min, the stress in the primary shear plane is lowest in the dry condition. However, this trend reverses above 60 m/min. This inflection point may be attributed to the heightened temperature in the shear plane due to the increased cutting speed and the limited effect of cryogenic assistance.

<figure><img src="image_1968.png" /><figcaption>Figure 4 : Evolution of strain in primary shear bands and friction coefficient according to cutting conditions for Ti54M</figcaption></figure>

The fundamental question revolves around how cryogenic assistance can affect the cutting process to varying degrees. As illustrated in Figure 4, it appears that the assistance can impact both the secondary and primary zones. However, discerning the proportion of the assistance's effect and its influence in these two zones proves challenging. To address this, we have differentiated the energy consumed during cutting for each zone.

The energy supplied (Wext), primarily transformed into temperature, is the product of cutting force and cutting speed. Consequently, by disregarding the energy in the tertiary zone, this energy can be divided into two components: the energy in the primary zone (W1) and the energy in the secondary zone (W2). Recognizing that temperature influences forces through thermal softening, a cooling source can alter this energy balance. This forms the focus of the subsequent stage of analysis. Furthermore, it is feasible to discern the phenomena by analyzing the energy in both shear zones.

$$W_{ext} = F_{c}V_{c} \qquad (6)$$

$$W_{1} = F_{s}V_{s} \qquad (7)$$

$$W_{2} = F_{n}V_{chip} \qquad (8)$$

Where $V_{chip}$ is the chip velocity and $V_s$ the shearing velocity.

These results are indeed intriguing, shedding light on the challenge of analyzing results solely based on stresses, strains, or friction coefficients. Initially, the external energy increases linearly and seems to remain consistent with or without assistance. This phenomenon can be attributed to the rise in cutting speed.

Secondly, it becomes apparent that in dry tests, the energy in the primary bands consistently remains lower. There are two plausible interpretations. The first suggests that the temperature generated by cutting constrains forces and consequently energy. This interpretation aligns with Komanduri's findings, illustrating the role of temperature in shear band formation. Alternatively, the cooling effect of the assistance appears to be efficacious. The separation of material in the primary band demands more energy and is thus less susceptible to temperature effects.
<figure><img src="image_1974.png" /><figcaption>Figure 5: Evolution of energies when cutting speed increase</figcaption></figure>

When the feed rate increases, the outcomes diverge (Figure 6). Indeed, the escalation in feed and thus in exertion inherently results in a significant surge in energy consumption. In this scenario, the energy consumed by the primary zone increases linearly with the feed rate in dry tests. Conversely, with cryogenic assistance, the energy remains stable. Regarding the secondary zone, the results also deviate from those obtained previously. Irrespective of the feed rate, the energy remains constant.

<figure><img src="image_1975.png" /><figcaption>Figure 6: Evolution of energies in different cutting zones when feed increases</figcaption></figure>

Lastly, another avenue for elucidating the evolution of the cutting process is chip analysis. Such analysis reveals all the thermomechanical stresses, including deformation and temperature. Drawing from the insights of the article in [6], the progression of the cutting process and the distribution of stresses in the zones can be deduced by analyzing the volumes of deformed material.

In this regard, it is observed that the volume in the primary shear bands remains constant regardless of the reshaping conditions. Conversely, the size of the secondary zone diminishes, along with its energy consumption (Figure 6).

<table><thead><tr><th></th><th>Dry conditions</th><th>Cryogenic conditions</th></tr></thead><tbody><tr><td>Vc=30 m/min</td><td></td><td></td></tr><tr><td>Vc=40 m/min</td><td></td><td></td></tr></tbody></table>
<figure><img src="image_1969.png" /><figcaption>Figure 5: Evolution of energies when cutting speed increase</figcaption></figure>

When the feed rate increases, the outcomes diverge (Figure 6). Indeed, the escalation in feed and thus in exertion inherently results in a significant surge in energy consumption. In this scenario, the energy consumed by the primary zone increases linearly with the feed rate in dry tests. Conversely, with cryogenic assistance, the energy remains stable. Regarding the secondary zone, the results also deviate from those obtained previously. Irrespective of the feed rate, the energy remains constant.

<figure><img src="image_1970.png" /><figcaption>Figure 6: Evolution of energies in different cutting zones when feed increases</figcaption></figure>

Lastly, another avenue for elucidating the evolution of the cutting process is chip analysis. Such analysis reveals all the thermomechanical stresses, including deformation and temperature. Drawing from the insights of the article in [6], the progression of the cutting process and the distribution of stresses in the zones can be deduced by analyzing the volumes of deformed material.

In this regard, it is observed that the volume in the primary shear bands remains constant regardless of the reshaping conditions. Conversely, the size of the secondary zone diminishes, along with its energy consumption (Figure 6).

<table><thead><tr><th></th><th>Dry conditions</th><th>Cryogenic conditions</th></tr></thead><tbody><tr><td>Vc=30 m/min</td><td></td><td></td></tr><tr><td>Vc=40 m/min</td><td></td><td></td></tr></tbody></table>
<figure><img src="image_1976.png" /><figcaption>Figure 7: Chip morphologies according to cutting conditions</figcaption></figure>

## Conclusion

This article investigates the influence of cryogenic techniques on the machining of Ti54M, a material gaining increasing popularity. Through an experimental campaign, we have delineated the constraints of cryogenic assistance particularly evident as cutting speeds escalate. It becomes apparent that at excessively high cutting speeds, the rise in temperature surpasses the compensatory capacity of the assistance. Conversely, and notably concerning wear, the stresses, and more significantly, the energies imparted in the secondary zone diminish with cryogenic assistance, despite higher temperatures prevailing in this zone. This variance may be elucidated by the specific application zone of the liquid nitrogen.

## References

[1] Cédrik Brozek. Conception et développement de nouveaux alliages de titane à haute ductilité et fort écrouissage. Matériaux. Université Pierre et Marie Curie - Paris VI, 2017. Français.

[2] Etude de la précipitation de la phase alpha dans les alliages proche-beta en fonction des conditions de forgeage antérieures, Lambert, Romain, 2021, http://www.theses.fr/2021UPSLC027/document

[3] Zindani, D., Kumar, K. A brief review on cryogenics in machining process. SN Appl. Sci. 2, 1107 (2020). https://doi.org/10.1007/s42452-020-2899-5

[4] Faycel Halila. Modélisation analytique et caractérisation expérimentale de l'usure par abrasion des outils de coupe. Université de Lorraine, 2015.

[5] Özel, Tuğrul & Zeren, Erol. (2006). A Methodology to Determine Work Material Flow Stress and Tool Chip Interfacial Friction Properties by Using Analysis of Machining. Journal of Manufacturing Science and Engineering, Transactions of the ASME. 128. 119-129. https://doi.org/10.1115/1.2118767

[6] Wagner V., Barelli F., Dessein G., Laheurte R., Darnis P., Cahuc O., Mousseigne M. (2018) Thermal and Microstructure Study of the Chip Formation during Turning of Ti64 β Lamellar Titanium Structure. Journal of Manufacturing Science Engineering Trans ASME. 140(3). https://doi.org/10.1115/1.4038597
<figure><img src="image_1971.png" /><figcaption>Figure 7: Chip morphologies according to cutting conditions</figcaption></figure>

## Conclusion

This article investigates the influence of cryogenic techniques on the machining of Ti54M, a material gaining increasing popularity. Through an experimental campaign, we have delineated the constraints of cryogenic assistance particularly evident as cutting speeds escalate. It becomes apparent that at excessively high cutting speeds, the rise in temperature surpasses the compensatory capacity of the assistance. Conversely, and notably concerning wear, the stresses, and more significantly, the energies imparted in the secondary zone diminish with cryogenic assistance, despite higher temperatures prevailing in this zone. This variance may be elucidated by the specific application zone of the liquid nitrogen.

## References

[1] Cédrik Brozek. Conception et développement de nouveaux alliages de titane à haute ductilité et fort écrouissage. Matériaux. Université Pierre et Marie Curie - Paris VI, 2017. Français.

[2] Etude de la précipitation de la phase alpha dans les alliages proche-beta en fonction des conditions de forgeage antérieures, Lambert, Romain, 2021, http://www.theses.fr/2021UPSLC027/document

[3] Zindani, D., Kumar, K. A brief review on cryogenics in machining process. SN Appl. Sci. 2, 1107 (2020). https://doi.org/10.1007/s42452-020-2899-5

[4] Faycel Halila. Modélisation analytique et caractérisation expérimentale de l'usure par abrasion des outils de coupe. Université de Lorraine, 2015.

[5] Özel, Tuğrul & Zeren, Erol. (2006). A Methodology to Determine Work Material Flow Stress and Tool Chip Interfacial Friction Properties by Using Analysis of Machining. Journal of Manufacturing Science and Engineering, Transactions of the ASME. 128. 119-129. https://doi.org/10.1115/1.2118767

[6] Wagner V., Barelli F., Dessein G., Laheurte R., Darnis P., Cahuc O., Mousseigne M. (2018) Thermal and Microstructure Study of the Chip Formation during Turning of Ti64 β Lamellar Titanium Structure. Journal of Manufacturing Science Engineering Trans ASME. 140(3). https://doi.org/10.1115/1.4038597
# Milling of nickel-based superalloys using throttle cryogenic cooling and micro-lubrication

IQBAL Asif<fnref n="1" />,ᵃ*, SAELZER Jannis<fnref n="2" />,ᵇ, NAUMAN Malik M.<fnref n="1" />,ᶜ and BIERMANN Dirk<fnref n="2" />,ᵈ

**Keywords:** Carbon Dioxide, Side-Milling, Tool Wear, Micro-Lubrication, Super-Alloys

**Abstract.** Nickel-based alloys are known to mount intense heat around the cutting edges of the cutting tool owing to their high yield strength, comparatively low thermal conductivity, and a significant seizure of the chip at the rake face even at low cutting speeds. Their sustainable machining demands a highly effective heat dissipation system. Due to its rich cooling potential, the compressed CO₂-based throttle cryogenic cooling system is a prime contender. The presented work seeks a sustainable milling solution for the cutting of the following three nickel-rich allows using the combination of throttle cryogenic cooling and minimum quantity of lubrication: Inconel 718, Incoloy 825, and Waspaloy. The experimental study quantifies the effects of the lubri-cooling approach and cutting speed on tool wear, cutting forces, and workpiece surface roughness. The novelty of the work lies in the mutual comparison of the machinability of the three alloys of nickel and the application of the cryogenic coolant in the form of pulses in addition to the conventional mode of continuous stream. The analyses of the experimental results suggest that the pulsed mode of the coolant, especially in combination with the lubricant, outperforms the continuous mode of the coolant’s supply regarding cutting forces and work surface roughness. Additionally, the study has yielded mixed results regarding tool wear for different combinations of workpiece material and cutting speed. Nonetheless, dry milling of the three alloys proved to be absolutely unsustainable in comparison to either of the two modes of throttle cryogenic cooling, with or without the application of micro-lubrication.

**Introduction**

Nickel-based superalloys are seeing a steady increase in their demand for making engineering components used in extreme heat and pressure environments. They find application mainly in the making of gas turbines, aerospace components, and automotive parts. The very mechanical properties that embellish the alloys for intense applications are also responsible for their poor workability in mechanical manufacturing, and the domain of machining is no exception. Machining is the only viable method to yield the dimensional accuracy and surface finish required for the engineering components used in the applications described above. Considering the high yield strengths and their retention at high temperatures, comparatively low thermal conductivities, intense engagement of the chip with the rake face, and adhesivity with the tool surface, viable machining of nickel-based superalloys requires a highly effective heat dissipation mechanism [1]. Emulsion-based flood coolants are not up to the mark due to their low cooling potential. The situation calls for the application of cryogenic cooling. A very low working temperature enables a cryogenic coolant to subdue an intense heat flux around the cutting edges. Moreover, it does not

<figure><img src="image_1977.png" /></figure>

<footnote n="1">¹Faculty of Integrated Technologies, Universiti Brunei Darussalam, Jalan Tungku Link, Gadong BE 1410, Brunei Darussalam</footnote>

<footnote n="2">²Institute of Machining Technology (ISF), Technische Universität Dortmund, Baroper Str. 303, 44227 Dortmund, Germany</footnote>

<footnote>ᵃasif.iqbal@ubd.edu.bn, ᵇjannis.saelzer@tu-dortmund.de, ᶜmalik.nauman@ubd.edu.bn,</footnote>

<footnote>ᵈdirk.biermann@tu-dortmund.de</footnote>
# Milling of nickel-based superalloys using throttle cryogenic cooling and micro-lubrication

IQBAL Asif<fnref n="1" />,ᵃ*, SAELZER Jannis<fnref n="2" />,ᵇ, NAUMAN Malik M.<fnref n="1" />,ᶜ and BIERMANN Dirk<fnref n="2" />,ᵈ

**Keywords:** Carbon Dioxide, Side-Milling, Tool Wear, Micro-Lubrication, Super-Alloys

**Abstract.** Nickel-based alloys are known to mount intense heat around the cutting edges of the cutting tool owing to their high yield strength, comparatively low thermal conductivity, and a significant seizure of the chip at the rake face even at low cutting speeds. Their sustainable machining demands a highly effective heat dissipation system. Due to its rich cooling potential, the compressed CO₂-based throttle cryogenic cooling system is a prime contender. The presented work seeks a sustainable milling solution for the cutting of the following three nickel-rich allows using the combination of throttle cryogenic cooling and minimum quantity of lubrication: Inconel 718, Incoloy 825, and Waspaloy. The experimental study quantifies the effects of the lubri-cooling approach and cutting speed on tool wear, cutting forces, and workpiece surface roughness. The novelty of the work lies in the mutual comparison of the machinability of the three alloys of nickel and the application of the cryogenic coolant in the form of pulses in addition to the conventional mode of continuous stream. The analyses of the experimental results suggest that the pulsed mode of the coolant, especially in combination with the lubricant, outperforms the continuous mode of the coolant’s supply regarding cutting forces and work surface roughness. Additionally, the study has yielded mixed results regarding tool wear for different combinations of workpiece material and cutting speed. Nonetheless, dry milling of the three alloys proved to be absolutely unsustainable in comparison to either of the two modes of throttle cryogenic cooling, with or without the application of micro-lubrication.

**Introduction**

Nickel-based superalloys are seeing a steady increase in their demand for making engineering components used in extreme heat and pressure environments. They find application mainly in the making of gas turbines, aerospace components, and automotive parts. The very mechanical properties that embellish the alloys for intense applications are also responsible for their poor workability in mechanical manufacturing, and the domain of machining is no exception. Machining is the only viable method to yield the dimensional accuracy and surface finish required for the engineering components used in the applications described above. Considering the high yield strengths and their retention at high temperatures, comparatively low thermal conductivities, intense engagement of the chip with the rake face, and adhesivity with the tool surface, viable machining of nickel-based superalloys requires a highly effective heat dissipation mechanism [1]. Emulsion-based flood coolants are not up to the mark due to their low cooling potential. The situation calls for the application of cryogenic cooling. A very low working temperature enables a cryogenic coolant to subdue an intense heat flux around the cutting edges. Moreover, it does not

<figure><img src="image_1972.png" /></figure>

<footnote n="1">¹Faculty of Integrated Technologies, Universiti Brunei Darussalam, Jalan Tungku Link, Gadong BE 1410, Brunei Darussalam</footnote>

<footnote n="2">²Institute of Machining Technology (ISF), Technische Universität Dortmund, Baroper Str. 303, 44227 Dortmund, Germany</footnote>

<footnote>ᵃasif.iqbal@ubd.edu.bn, ᵇjannis.saelzer@tu-dortmund.de, ᶜmalik.nauman@ubd.edu.bn,</footnote>

<footnote>ᵈdirk.biermann@tu-dortmund.de</footnote>
create any unsustainable wastage issues as is the case with flood coolants. To work out the research gap in this context, a brief literature review is provided below.

The machinability index of nickel-based alloys is around 35 on a scale from 0 to 100, making them exceedingly more difficult to cut than a plain carbon steel [2]. The machinability index of AISI 1112 plain carbon steel is taken as 100. In addition to high strengths, their low thermal conductivities also contribute to tool wear by channeling a high proportion of process heat to the cutting tool [3]. These alloys are also known for intense work hardening, which means they pose an increased difficulty to cutting as the machining process progresses [4]. Hybrid cryogenic cooling (using liquid nitrogen) and lubrication (MQL) are used in the machining of Inconel 718 [5]. The study reported that the application of the hybrid environment ensures the longest tool life. Another study carried out on Cryo-MQL machining of Inconel 718 reported the achievement of better surface finish, lower cutting temperature, and milder tool wear in comparison to dry, MQL, and even cryogenic machining [6]. Additively manufactured nickel-based parts were machined under hybrid cryogenic cooling [7]. It was reported that the hybrid lubri-cooling approach remains highly effective in curbing tool wear at high cutting speeds. In another investigation, a nickel-chromium alloy (Nimonic 80A) was machined under various cryogenic cooling and lubrication environments [8]. The cryogenic cooling yielded a significantly better surface finish as compared to dry and flood cooling but not as good as that by micro-lubrication. Inconel 718 and Waspaloy were face-turned under emulsion coolant pressures of 6 and 80 bars [9]. The high-pressure cooling was found to reduce the flank wear on both alloys. High alloy adhesion was also reported on the tooltip. The micro-crack formation, shedding of the coating on the two tool faces, and tool-tip chipping are the major modes of tool failure in the micro-milling of Inconel 718 using coated carbide end mills [10]. A study has emphasized that liquid nitrogen-based inner injection cooling is more advantageous concerning the machining performance measures than the external application in the milling of nickel-based alloys [11]. An experimental study was carried out for comparing the performance of LN₂ and CO₂ for enhancing tool life in machining of Inconel 718. The study reported that the former remained worse than the former in this regard whereas the conventional lubrication technique outperformed both.

The review leads to a gap asking for exploring application of various modes cryogenic cooling with and without the addition of micro-lubrication in the machining of nickel-based superalloys. Considering the success of hybrid lubri-cooling attained in the machining of other difficult-to-cut materials calls for its application in the milling of nickel alloys as well. The current study investigates the efficacy of applying compressed CO₂-based throttle cryogenic coolant, with and without the addition of micro-lubrication, in the milling of the following superalloys of nickel: Inconel 718, Incoloy 825, and Waspaloy. Specific to the operational mechanism of the milling process, characterized by periodic engagement and disengagement of the cutting edges with the workpiece material, the effectiveness of the throttle cryogenic coolant will be tested in pulsed and continuous modes of supply. In short, the presented work targets underscoring the potential of cryogenic cooling-based lubri-cooling approach in making the machining of nickel alloys sustainable.

## Experimental Work

This section provides details on the milling experiments performed, including the experimental setup, design of experiments, inputs and responses, tooling, and workpiece materials.

### Workpiece Materials and Milling Tools

The three workpiece materials tested in the work are Inconel 718, Incoloy 825, and Waspaloy. The first one is a commonly used high-strength, corrosion-resistant nickel-chromium alloy. It is a precipitation-hardened material having noteworthy proportions of iron, molybdenum, niobium, and along with smaller quantities of titanium and aluminum. High strength, extreme toughness,
create any unsustainable wastage issues as is the case with flood coolants. To work out the research gap in this context, a brief literature review is provided below.

The machinability index of nickel-based alloys is around 35 on a scale from 0 to 100, making them exceedingly more difficult to cut than a plain carbon steel [2]. The machinability index of AISI 1112 plain carbon steel is taken as 100. In addition to high strengths, their low thermal conductivities also contribute to tool wear by channeling a high proportion of process heat to the cutting tool [3]. These alloys are also known for intense work hardening, which means they pose an increased difficulty to cutting as the machining process progresses [4]. Hybrid cryogenic cooling (using liquid nitrogen) and lubrication (MQL) are used in the machining of Inconel 718 [5]. The study reported that the application of the hybrid environment ensures the longest tool life. Another study carried out on Cryo-MQL machining of Inconel 718 reported the achievement of better surface finish, lower cutting temperature, and milder tool wear in comparison to dry, MQL, and even cryogenic machining [6]. Additively manufactured nickel-based parts were machined under hybrid cryogenic cooling [7]. It was reported that the hybrid lubri-cooling approach remains highly effective in curbing tool wear at high cutting speeds. In another investigation, a nickel-chromium alloy (Nimonic 80A) was machined under various cryogenic cooling and lubrication environments [8]. The cryogenic cooling yielded a significantly better surface finish as compared to dry and flood cooling but not as good as that by micro-lubrication. Inconel 718 and Waspaloy were face-turned under emulsion coolant pressures of 6 and 80 bars [9]. The high-pressure cooling was found to reduce the flank wear on both alloys. High alloy adhesion was also reported on the tooltip. The micro-crack formation, shedding of the coating on the two tool faces, and tool-tip chipping are the major modes of tool failure in the micro-milling of Inconel 718 using coated carbide end mills [10]. A study has emphasized that liquid nitrogen-based inner injection cooling is more advantageous concerning the machining performance measures than the external application in the milling of nickel-based alloys [11]. An experimental study was carried out for comparing the performance of LN₂ and CO₂ for enhancing tool life in machining of Inconel 718. The study reported that the former remained worse than the former in this regard whereas the conventional lubrication technique outperformed both.

The review leads to a gap asking for exploring application of various modes cryogenic cooling with and without the addition of micro-lubrication in the machining of nickel-based superalloys. Considering the success of hybrid lubri-cooling attained in the machining of other difficult-to-cut materials calls for its application in the milling of nickel alloys as well. The current study investigates the efficacy of applying compressed CO₂-based throttle cryogenic coolant, with and without the addition of micro-lubrication, in the milling of the following superalloys of nickel: Inconel 718, Incoloy 825, and Waspaloy. Specific to the operational mechanism of the milling process, characterized by periodic engagement and disengagement of the cutting edges with the workpiece material, the effectiveness of the throttle cryogenic coolant will be tested in pulsed and continuous modes of supply. In short, the presented work targets underscoring the potential of cryogenic cooling-based lubri-cooling approach in making the machining of nickel alloys sustainable.

## Experimental Work

This section provides details on the milling experiments performed, including the experimental setup, design of experiments, inputs and responses, tooling, and workpiece materials.

### Workpiece Materials and Milling Tools

The three workpiece materials tested in the work are Inconel 718, Incoloy 825, and Waspaloy. The first one is a commonly used high-strength, corrosion-resistant nickel-chromium alloy. It is a precipitation-hardened material having noteworthy proportions of iron, molybdenum, niobium, and along with smaller quantities of titanium and aluminum. High strength, extreme toughness,
and intense work-hardening render this superalloy a difficult-to-cut material. Incoloy 825, a relatively lesser-known material, is a heat-treatable nickel-iron-chromium alloy with lesser but significant amounts of molybdenum, titanium, and copper. It is different from Inconel 718 in terms of a notably smaller proportion of nickel, which gives way to a larger share of iron and makes it relatively easier to machine. Waspaloy, another superalloy of nickel, is closer to Inconel 718 as compared to Incoloy 825 with respect to chemical composition and mechanical characteristics. Likewise, Waspaloy is precipitation-hardened and possesses high high-temperature strength and excellent corrosion resistance. Tables 1 and 2 present comparative analyses in terms of chemical composition and mechanical properties, respectively of the three superalloys.

<table><caption>Table 1. Chemical compositions of Inconel 718, Incoloy 825, and Waspaloy (the numbers are in percentage)</caption><thead><tr><th></th><th>Ni</th><th>Cr</th><th>Fe</th><th>Mo</th><th>Cu</th><th>Ti</th><th>C</th><th>Mn</th><th>Co</th><th>Si</th><th>Al</th></tr></thead><tbody><tr><th>Incoloy 825</th><td>43</td><td>24</td><td>26</td><td>3.1</td><td>2</td><td>0.9</td><td>0.03</td><td>0.3</td><td>-</td><td>0.44</td><td>0.2</td></tr><tr><th>Inconel 718</th><td>52</td><td>19</td><td>17</td><td>2.8</td><td>0.25</td><td>0.6</td><td>0.06</td><td>0.25</td><td>0.9</td><td>0.3</td><td>0.4</td></tr><tr><th>Waspaloy</th><td>54</td><td>19.5</td><td>2</td><td>4.25</td><td>0.5</td><td>3</td><td>0.07</td><td>1</td><td>13.5</td><td>0.75</td><td>1.4</td></tr></tbody></table>

<table><caption>Table 2. Mechanical properties of Inconel 718, Incoloy 825, and Waspaloy</caption><thead><tr><th></th><th>Tensile Strength</th><th>Yield Strength</th><th>Elongation</th><th>Hardness</th><th>Thermal Conductivity</th></tr></thead><tbody><tr><th>Incoloy 825</th><td>703 MPa</td><td>340 MPa</td><td>39 %</td><td>22 HRc</td><td>11.1 W/m.K</td></tr><tr><th>Inconel 718</th><td>1035 MPa</td><td>725 MPa</td><td>25 %</td><td>32 HRc</td><td>11.4 W/m.K</td></tr><tr><th>Waspaloy</th><td>888 MPa</td><td>443 MPa</td><td>49 %</td><td>24 HRc</td><td>11.3 W/m.K</td></tr></tbody></table>

Ceramic coated (multiple layers of TiN and TiAlN) tungsten carbide side-and-end milling cutters possessing a diameter (D), cutting length (l<sub>c</sub>), total length (l), helix angle (λ), and the number of teeth (z) of 8 mm, 40 mm, 100 mm, 30°, and 4, respectively, are used in this study. Ceramic coatings, due to their extreme hardness, add to the wear resistance of the carbide tools. A new milling tool is used for each experiment.

## Design of Experiments

The following three inputs are controlled:

a) Workpiece material. The three levels of this categorical parameter are Inconel 718, Incoloy 825, and Waspaloy.

b) Cutting speed. The two levels of this input are 80 m/min and 120 m/min.

c) Cooling and lubrication mechanism. The five levels of this categorical parameter are dry, continuous stream of the throttle cryogenic coolant (th<sub>c</sub>), pulsed stream of the throttle cryogenic coolant (th<sub>p</sub>), continuous stream of the throttle cryogenic coolant with parallel addition of micro-lubrication (th<sub>c-l</sub>), pulsed stream of the throttle cryogenic coolant with parallel addition of micro-lubrication (th<sub>p-l</sub>).

A full-factorial design of experiments yields 30 (3 × 2 × 5) experimental runs. For each run, the following three responses are measured:

a) Average width of flank wear land (VB - µm).

b) Three orthogonal components of cutting forces (F<sub>x</sub>, F<sub>y</sub>, and F<sub>z</sub> - N). The components F<sub>x</sub> and F<sub>z</sub> are aligned with the feed direction and the cutter's axis, respectively. The third component is perpendicular to the other two.

c) Average arithmetic roughness of the machined surface (Ra - µm)

The throttle cryogenic cooling of the milling area is carried out by transporting compressed CO<sub>2</sub> from a cylinder through a metallic tube at a mass flow rate of 2.2 kg/h and throttling it out through
and intense work-hardening render this superalloy a difficult-to-cut material. Incoloy 825, a relatively lesser-known material, is a heat-treatable nickel-iron-chromium alloy with lesser but significant amounts of molybdenum, titanium, and copper. It is different from Inconel 718 in terms of a notably smaller proportion of nickel, which gives way to a larger share of iron and makes it relatively easier to machine. Waspaloy, another superalloy of nickel, is closer to Inconel 718 as compared to Incoloy 825 with respect to chemical composition and mechanical characteristics. Likewise, Waspaloy is precipitation-hardened and possesses high high-temperature strength and excellent corrosion resistance. Tables 1 and 2 present comparative analyses in terms of chemical composition and mechanical properties, respectively of the three superalloys.

<table><caption>Table 1. Chemical compositions of Inconel 718, Incoloy 825, and Waspaloy (the numbers are in percentage)</caption><thead><tr><th></th><th>Ni</th><th>Cr</th><th>Fe</th><th>Mo</th><th>Cu</th><th>Ti</th><th>C</th><th>Mn</th><th>Co</th><th>Si</th><th>Al</th></tr></thead><tbody><tr><th>Incoloy 825</th><td>43</td><td>24</td><td>26</td><td>3.1</td><td>2</td><td>0.9</td><td>0.03</td><td>0.3</td><td>-</td><td>0.44</td><td>0.2</td></tr><tr><th>Inconel 718</th><td>52</td><td>19</td><td>17</td><td>2.8</td><td>0.25</td><td>0.6</td><td>0.06</td><td>0.25</td><td>0.9</td><td>0.3</td><td>0.4</td></tr><tr><th>Waspaloy</th><td>54</td><td>19.5</td><td>2</td><td>4.25</td><td>0.5</td><td>3</td><td>0.07</td><td>1</td><td>13.5</td><td>0.75</td><td>1.4</td></tr></tbody></table>

<table><caption>Table 2. Mechanical properties of Inconel 718, Incoloy 825, and Waspaloy</caption><thead><tr><th></th><th>Tensile Strength</th><th>Yield Strength</th><th>Elongation</th><th>Hardness</th><th>Thermal Conductivity</th></tr></thead><tbody><tr><th>Incoloy 825</th><td>703 MPa</td><td>340 MPa</td><td>39 %</td><td>22 HRc</td><td>11.1 W/m.K</td></tr><tr><th>Inconel 718</th><td>1035 MPa</td><td>725 MPa</td><td>25 %</td><td>32 HRc</td><td>11.4 W/m.K</td></tr><tr><th>Waspaloy</th><td>888 MPa</td><td>443 MPa</td><td>49 %</td><td>24 HRc</td><td>11.3 W/m.K</td></tr></tbody></table>

Ceramic coated (multiple layers of TiN and TiAlN) tungsten carbide side-and-end milling cutters possessing a diameter (D), cutting length (l<sub>c</sub>), total length (l), helix angle (λ), and the number of teeth (z) of 8 mm, 40 mm, 100 mm, 30°, and 4, respectively, are used in this study. Ceramic coatings, due to their extreme hardness, add to the wear resistance of the carbide tools. A new milling tool is used for each experiment.

## Design of Experiments

The following three inputs are controlled:

a) Workpiece material. The three levels of this categorical parameter are Inconel 718, Incoloy 825, and Waspaloy.

b) Cutting speed. The two levels of this input are 80 m/min and 120 m/min.

c) Cooling and lubrication mechanism. The five levels of this categorical parameter are dry, continuous stream of the throttle cryogenic coolant (th<sub>c</sub>), pulsed stream of the throttle cryogenic coolant (th<sub>p</sub>), continuous stream of the throttle cryogenic coolant with parallel addition of micro-lubrication (th<sub>c-l</sub>), pulsed stream of the throttle cryogenic coolant with parallel addition of micro-lubrication (th<sub>p-l</sub>).

A full-factorial design of experiments yields 30 (3 × 2 × 5) experimental runs. For each run, the following three responses are measured:

a) Average width of flank wear land (VB - µm).

b) Three orthogonal components of cutting forces (F<sub>x</sub>, F<sub>y</sub>, and F<sub>z</sub> - N). The components F<sub>x</sub> and F<sub>z</sub> are aligned with the feed direction and the cutter's axis, respectively. The third component is perpendicular to the other two.

c) Average arithmetic roughness of the machined surface (R<sub>a</sub> - µm)

The throttle cryogenic cooling of the milling area is carried out by transporting compressed CO<sub>2</sub> from a cylinder through a metallic tube at a mass flow rate of 2.2 kg/h and throttling it out through
an orifice of 1.5 mm diameter. The throttling fluid absorbs heat from its vicinity following the Joule-Thomson phenomenon. The distinction between the pulsed and continuous supply of the cooling fluid is made by a control device that cyclically opens and closes the fluid delivery valve after every fixed period for the former. The pulse frequency of the pulsated coolant supply is 30 Hz. The device is placed outside the machine tool and is activated manually. For the continuous supply, the valve remains open during the entire cutting duration of the run. For the runs involving micro-lubrication (*th*<sub>c</sub>-*l* and *th*<sub>p</sub>-*l*), the aerosol is supplied through a separate tube to the cutting region. The aerosol is a biodegradable oil atomized into a flow of air compressed to 6 bars. The supply of micro-lubrication is not pulsed for any experimental run.

**Experimental Setup**

The experiments are performed on a CNC milling machine, Chiron FZ 12S Magnum. The milling tools are held in a shrink-fit holding mechanism to grant more accessibility of the cutting area to the cooling and lubrication nozzles. A digital optical microscope is used to measure the tool wear. The width of flank wear land is determined at three locations on each of the four edges of the used milling tool. The VB is then evaluated by taking the mean of the 12 readings. The R<sub>a</sub> of the machined surface is measured using a surface roughness meter, Mahr Marsurf M300C. Finally, the machining forces are determined using a Kistler-based dynamometer (type 9255B) and a multichannel charge amplifier. The force readings are analyzed using DIAdem, a numeric data management software. Fig. 1 presents the experimental setup.

<figure><img src="image_1978.png" /><figcaption>Fig. 1. The experimental setup</figcaption></figure>

Each of the 30 experiments is completed in three passes maintaining the total length of cut equal to 300 mm. The radial and axial depths of cut are kept at 0.2 mm and 8 mm, respectively. The feed rate is fixed to 0.1mm/tooth and down-milling is selected as the milling orientation. Lastly, the tool overhang for each run is kept at 52 mm.

**Experimental Results and Analyses**

Table 3 presents the results obtained from the experiments regarding tool wear, machining forces, and work surface roughness. The upcoming sub-sections discuss the tabulated results and their analyses.
an orifice of 1.5 mm diameter. The throttling fluid absorbs heat from its vicinity following the Joule-Thomson phenomenon. The distinction between the pulsed and continuous supply of the cooling fluid is made by a control device that cyclically opens and closes the fluid delivery valve after every fixed period for the former. The pulse frequency of the pulsated coolant supply is 30 Hz. The device is placed outside the machine tool and is activated manually. For the continuous supply, the valve remains open during the entire cutting duration of the run. For the runs involving micro-lubrication (*th*<sub>c</sub>-*l* and *th*<sub>p</sub>-*l*), the aerosol is supplied through a separate tube to the cutting region. The aerosol is a biodegradable oil atomized into a flow of air compressed to 6 bars. The supply of micro-lubrication is not pulsed for any experimental run.

**Experimental Setup**

The experiments are performed on a CNC milling machine, Chiron FZ 12S Magnum. The milling tools are held in a shrink-fit holding mechanism to grant more accessibility of the cutting area to the cooling and lubrication nozzles. A digital optical microscope is used to measure the tool wear. The width of flank wear land is determined at three locations on each of the four edges of the used milling tool. The VB is then evaluated by taking the mean of the 12 readings. The R<sub>a</sub> of the machined surface is measured using a surface roughness meter, Mahr Marsurf M300C. Finally, the machining forces are determined using a Kistler-based dynamometer (type 9255B) and a multichannel charge amplifier. The force readings are analyzed using DIAdem, a numeric data management software. Fig. 1 presents the experimental setup.

<figure><img src="image_1973.png" /><figcaption>Fig. 1. The experimental setup</figcaption></figure>

Each of the 30 experiments is completed in three passes maintaining the total length of cut equal to 300 mm. The radial and axial depths of cut are kept at 0.2 mm and 8 mm, respectively. The feed rate is fixed to 0.1mm/tooth and down-milling is selected as the milling orientation. Lastly, the tool overhang for each run is kept at 52 mm.

**Experimental Results and Analyses**

Table 3 presents the results obtained from the experiments regarding tool wear, machining forces, and work surface roughness. The upcoming sub-sections discuss the tabulated results and their analyses.
## Tool Wear

The fifth column of Table 3 presents the measured values of VB for all 30 runs of the three alloys tested at the two different speeds and 5 lubri-cooling approaches. For Incoloy 825, apart from the dry test carried out at the high level of cutting speed, the VB results do not show a strong dependence on cutting speed. Similarly, both VB values of Inconel 718 and Waspaloy show a similar outcome regardless of the cutting speed, however, the lubricated test shows a significant difference at the lower cutting speed especially with Waspaloy.

<table><caption>Table 3. Experimental results regarding tool wear, machining forces, and surface roughness</caption><thead><tr><th>S/ No.</th><th>Material</th><th>Lubri-Cooling</th><th>Cutting speed (m/min)</th><th>VB (µm)</th><th>F<sub>x</sub> (N)</th><th>F<sub>y</sub> (N)</th><th>F<sub>z</sub> (N)</th><th>R<sub>a</sub> (µm)</th></tr></thead><tbody><tr><td>1</td><td>Incoloy 825</td><td>Dry</td><td>120</td><td>498.98</td><td>197.84</td><td>258.3</td><td>141.34</td><td>3.07</td></tr><tr><td>2</td><td>Incoloy 825</td><td>Dry</td><td>80</td><td>78.94</td><td>173.56</td><td>194.22</td><td>88.39</td><td>2.42</td></tr><tr><td>3</td><td>Incoloy 825</td><td>th<sub>c</sub></td><td>120</td><td>40.68</td><td>121.19</td><td>145.14</td><td>68.15</td><td>2.757</td></tr><tr><td>4</td><td>Incoloy 825</td><td>th<sub>c</sub></td><td>80</td><td>45.49</td><td>106.77</td><td>122.5</td><td>48.66</td><td>2.147</td></tr><tr><td>5</td><td>Incoloy 825</td><td>th<sub>p</sub></td><td>120</td><td>67.96</td><td>122.9</td><td>145.94</td><td>57.97</td><td>1.143</td></tr><tr><td>6</td><td>Incoloy 825</td><td>th<sub>p</sub></td><td>80</td><td>92.22</td><td>165.65</td><td>169.8</td><td>90.94</td><td>2.33</td></tr><tr><td>7</td><td>Incoloy 825</td><td>th<sub>c</sub>-l</td><td>120</td><td>57.31</td><td>120.87</td><td>148.07</td><td>68.23</td><td>1.703</td></tr><tr><td>8</td><td>Incoloy 825</td><td>th<sub>c</sub>-l</td><td>80</td><td>53.61</td><td>117.13</td><td>124.51</td><td>43.87</td><td>1.885</td></tr><tr><td>9</td><td>Incoloy 825</td><td>th<sub>p</sub>-l</td><td>120</td><td>60.76</td><td>113.29</td><td>140.74</td><td>59.96</td><td>1.386</td></tr><tr><td>10</td><td>Incoloy 825</td><td>th<sub>p</sub>-l</td><td>80</td><td>61.87</td><td>115.54</td><td>134.03</td><td>49.13</td><td>0.974</td></tr><tr><td>11</td><td>Inconel 718</td><td>Dry</td><td>120</td><td>88.29</td><td>153.63</td><td>201.24</td><td>97.61</td><td>1.515</td></tr><tr><td>12</td><td>Inconel 718</td><td>Dry</td><td>80</td><td>76.81</td><td>145.86</td><td>177.47</td><td>60.03</td><td>1.396</td></tr><tr><td>13</td><td>Inconel 718</td><td>th<sub>c</sub></td><td>120</td><td>109.56</td><td>161.8</td><td>206.83</td><td>101.32</td><td>2.313</td></tr><tr><td>14</td><td>Inconel 718</td><td>th<sub>c</sub></td><td>80</td><td>86.77</td><td>146.42</td><td>167.32</td><td>60.73</td><td>1.864</td></tr><tr><td>15</td><td>Inconel 718</td><td>th<sub>p</sub></td><td>120</td><td>206.39</td><td>172.63</td><td>198.66</td><td>103.42</td><td>1.632</td></tr><tr><td>16</td><td>Inconel 718</td><td>th<sub>p</sub></td><td>80</td><td>83.07</td><td>132.8</td><td>149.87</td><td>56.54</td><td>2.306</td></tr><tr><td>17</td><td>Inconel 718</td><td>th<sub>c</sub>-l</td><td>120</td><td>62.95</td><td>165.47</td><td>189.44</td><td>89.62</td><td>2.425</td></tr><tr><td>18</td><td>Inconel 718</td><td>th<sub>c</sub>-l</td><td>80</td><td>76.71</td><td>133.67</td><td>141.8</td><td>47.67</td><td>1.439</td></tr><tr><td>19</td><td>Inconel 718</td><td>th<sub>p</sub>-l</td><td>120</td><td>63.79</td><td>178.75</td><td>199.62</td><td>96.47</td><td>1.266</td></tr><tr><td>20</td><td>Inconel 718</td><td>th<sub>p</sub>-l</td><td>80</td><td>58.11</td><td>137.5</td><td>133.59</td><td>49.75</td><td>0.907</td></tr><tr><td>21</td><td>Waspaloy</td><td>Dry</td><td>120</td><td>107.10</td><td>158.39</td><td>208.98</td><td>78.06</td><td>1.893</td></tr><tr><td>22</td><td>Waspaloy</td><td>Dry</td><td>80</td><td>105.20</td><td>148.33</td><td>186.72</td><td>60.17</td><td>1.232</td></tr><tr><td>23</td><td>Waspaloy</td><td>th<sub>c</sub></td><td>120</td><td>590.26</td><td>184.57</td><td>186.82</td><td>88.62</td><td>1.444</td></tr><tr><td>24</td><td>Waspaloy</td><td>th<sub>c</sub></td><td>80</td><td>79.08</td><td>159.44</td><td>157.79</td><td>51.03</td><td>2.584</td></tr><tr><td>25</td><td>Waspaloy</td><td>th<sub>p</sub></td><td>120</td><td>218.44</td><td>172.49</td><td>188.13</td><td>92.18</td><td>2.038</td></tr><tr><td>26</td><td>Waspaloy</td><td>th<sub>p</sub></td><td>80</td><td>60.01</td><td>164.36</td><td>163.75</td><td>53.46</td><td>2.582</td></tr><tr><td>27</td><td>Waspaloy</td><td>th<sub>c</sub>-l</td><td>120</td><td>154.42</td><td>180.85</td><td>193.53</td><td>99.21</td><td>1.499</td></tr><tr><td>28</td><td>Waspaloy</td><td>th<sub>c</sub>-l</td><td>80</td><td>74.56</td><td>168.23</td><td>170.08</td><td>45.78</td><td>0.972</td></tr><tr><td>29</td><td>Waspaloy</td><td>th<sub>p</sub>-l</td><td>120</td><td>94.92</td><td>152.5</td><td>172.48</td><td>92.95</td><td>2.752</td></tr><tr><td>30</td><td>Waspaloy</td><td>th<sub>p</sub>-l</td><td>80</td><td>70.86</td><td>136.89</td><td>149.05</td><td>52.03</td><td>1.457</td></tr></tbody></table>

Continuous supply of throttling CO₂ ($th_c$) yields milder tool wear than pulsing the supply ($th_p$) in Incoloy 825. The continuous mode provides almost double the cooling potential compared to the pulsed mode, with the addition of lubrication, the pulsed mode ($th_p-l$) is more effective in checking tool wear as the lubricant successfully made its way to the shear zone through the punctuated supply of coolant. In contrast to the continuous mode ($th_c-l$), the coolant acts as a barricade from the lubricant to be delivered and even compromises the coolant's efficacy resulting in the intensification of the tool wear. For Incoloy 825 the $th_c$ mode is the lubri-cooling mode that is effective in checking the tool wear at both levels of cutting speed.

For Waspaloy, high cutting speed with the $th_c$ mode resulted in the maximum tool wear whereas at low cutting speed, it had the minimum tool wear with the $th_p$ mode. The result shows there is higher tool wear with $th_c$ compared to $th_p$ mode at both speeds and this is also consistent with the
## Tool Wear

The fifth column of Table 3 presents the measured values of VB for all 30 runs of the three alloys tested at the two different speeds and 5 lubri-cooling approaches. For Incoloy 825, apart from the dry test carried out at the high level of cutting speed, the VB results do not show a strong dependence on cutting speed. Similarly, both VB values of Inconel 718 and Waspaloy show a similar outcome regardless of the cutting speed, however, the lubricated test shows a significant difference at the lower cutting speed especially with Waspaloy.

<table><caption>Table 3. Experimental results regarding tool wear, machining forces, and surface roughness</caption><thead><tr><th>S/ No.</th><th>Material</th><th>Lubri-Cooling</th><th>Cutting speed (m/min)</th><th>VB (µm)</th><th>F<sub>x</sub> (N)</th><th>F<sub>y</sub> (N)</th><th>F<sub>z</sub> (N)</th><th>R<sub>a</sub> (µm)</th></tr></thead><tbody><tr><td>1</td><td>Incoloy 825</td><td>Dry</td><td>120</td><td>498.98</td><td>197.84</td><td>258.3</td><td>141.34</td><td>3.07</td></tr><tr><td>2</td><td>Incoloy 825</td><td>Dry</td><td>80</td><td>78.94</td><td>173.56</td><td>194.22</td><td>88.39</td><td>2.42</td></tr><tr><td>3</td><td>Incoloy 825</td><td>th<sub>c</sub></td><td>120</td><td>40.68</td><td>121.19</td><td>145.14</td><td>68.15</td><td>2.757</td></tr><tr><td>4</td><td>Incoloy 825</td><td>th<sub>c</sub></td><td>80</td><td>45.49</td><td>106.77</td><td>122.5</td><td>48.66</td><td>2.147</td></tr><tr><td>5</td><td>Incoloy 825</td><td>th<sub>p</sub></td><td>120</td><td>67.96</td><td>122.9</td><td>145.94</td><td>57.97</td><td>1.143</td></tr><tr><td>6</td><td>Incoloy 825</td><td>th<sub>p</sub></td><td>80</td><td>92.22</td><td>165.65</td><td>169.8</td><td>90.94</td><td>2.33</td></tr><tr><td>7</td><td>Incoloy 825</td><td>th<sub>c</sub>-l</td><td>120</td><td>57.31</td><td>120.87</td><td>148.07</td><td>68.23</td><td>1.703</td></tr><tr><td>8</td><td>Incoloy 825</td><td>th<sub>c</sub>-l</td><td>80</td><td>53.61</td><td>117.13</td><td>124.51</td><td>43.87</td><td>1.885</td></tr><tr><td>9</td><td>Incoloy 825</td><td>th<sub>p</sub>-l</td><td>120</td><td>60.76</td><td>113.29</td><td>140.74</td><td>59.96</td><td>1.386</td></tr><tr><td>10</td><td>Incoloy 825</td><td>th<sub>p</sub>-l</td><td>80</td><td>61.87</td><td>115.54</td><td>134.03</td><td>49.13</td><td>0.974</td></tr><tr><td>11</td><td>Inconel 718</td><td>Dry</td><td>120</td><td>88.29</td><td>153.63</td><td>201.24</td><td>97.61</td><td>1.515</td></tr><tr><td>12</td><td>Inconel 718</td><td>Dry</td><td>80</td><td>76.81</td><td>145.86</td><td>177.47</td><td>60.03</td><td>1.396</td></tr><tr><td>13</td><td>Inconel 718</td><td>th<sub>c</sub></td><td>120</td><td>109.56</td><td>161.8</td><td>206.83</td><td>101.32</td><td>2.313</td></tr><tr><td>14</td><td>Inconel 718</td><td>th<sub>c</sub></td><td>80</td><td>86.77</td><td>146.42</td><td>167.32</td><td>60.73</td><td>1.864</td></tr><tr><td>15</td><td>Inconel 718</td><td>th<sub>p</sub></td><td>120</td><td>206.39</td><td>172.63</td><td>198.66</td><td>103.42</td><td>1.632</td></tr><tr><td>16</td><td>Inconel 718</td><td>th<sub>p</sub></td><td>80</td><td>83.07</td><td>132.8</td><td>149.87</td><td>56.54</td><td>2.306</td></tr><tr><td>17</td><td>Inconel 718</td><td>th<sub>c</sub>-l</td><td>120</td><td>62.95</td><td>165.47</td><td>189.44</td><td>89.62</td><td>2.425</td></tr><tr><td>18</td><td>Inconel 718</td><td>th<sub>c</sub>-l</td><td>80</td><td>76.71</td><td>133.67</td><td>141.8</td><td>47.67</td><td>1.439</td></tr><tr><td>19</td><td>Inconel 718</td><td>th<sub>p</sub>-l</td><td>120</td><td>63.79</td><td>178.75</td><td>199.62</td><td>96.47</td><td>1.266</td></tr><tr><td>20</td><td>Inconel 718</td><td>th<sub>p</sub>-l</td><td>80</td><td>58.11</td><td>137.5</td><td>133.59</td><td>49.75</td><td>0.907</td></tr><tr><td>21</td><td>Waspaloy</td><td>Dry</td><td>120</td><td>107.10</td><td>158.39</td><td>208.98</td><td>78.06</td><td>1.893</td></tr><tr><td>22</td><td>Waspaloy</td><td>Dry</td><td>80</td><td>105.20</td><td>148.33</td><td>186.72</td><td>60.17</td><td>1.232</td></tr><tr><td>23</td><td>Waspaloy</td><td>th<sub>c</sub></td><td>120</td><td>590.26</td><td>184.57</td><td>186.82</td><td>88.62</td><td>1.444</td></tr><tr><td>24</td><td>Waspaloy</td><td>th<sub>c</sub></td><td>80</td><td>79.08</td><td>159.44</td><td>157.79</td><td>51.03</td><td>2.584</td></tr><tr><td>25</td><td>Waspaloy</td><td>th<sub>p</sub></td><td>120</td><td>218.44</td><td>172.49</td><td>188.13</td><td>92.18</td><td>2.038</td></tr><tr><td>26</td><td>Waspaloy</td><td>th<sub>p</sub></td><td>80</td><td>60.01</td><td>164.36</td><td>163.75</td><td>53.46</td><td>2.582</td></tr><tr><td>27</td><td>Waspaloy</td><td>th<sub>c</sub>-l</td><td>120</td><td>154.42</td><td>180.85</td><td>193.53</td><td>99.21</td><td>1.499</td></tr><tr><td>28</td><td>Waspaloy</td><td>th<sub>c</sub>-l</td><td>80</td><td>74.56</td><td>168.23</td><td>170.08</td><td>45.78</td><td>0.972</td></tr><tr><td>29</td><td>Waspaloy</td><td>th<sub>p</sub>-l</td><td>120</td><td>94.92</td><td>152.5</td><td>172.48</td><td>92.95</td><td>2.752</td></tr><tr><td>30</td><td>Waspaloy</td><td>th<sub>p</sub>-l</td><td>80</td><td>70.86</td><td>136.89</td><td>149.05</td><td>52.03</td><td>1.457</td></tr></tbody></table>

Continuous supply of throttling CO₂ ($th_c$) yields milder tool wear than pulsing the supply ($th_p$) in Incoloy 825. The continuous mode provides almost double the cooling potential compared to the pulsed mode, with the addition of lubrication, the pulsed mode ($th_p-l$) is more effective in checking tool wear as the lubricant successfully made its way to the shear zone through the punctuated supply of coolant. In contrast to the continuous mode ($th_c-l$), the coolant acts as a barricade from the lubricant to be delivered and even compromises the coolant's efficacy resulting in the intensification of the tool wear. For Incoloy 825 the $th_c$ mode is the lubri-cooling mode that is effective in checking the tool wear at both levels of cutting speed.

For Waspaloy, high cutting speed with the $th_c$ mode resulted in the maximum tool wear whereas at low cutting speed, it had the minimum tool wear with the $th_p$ mode. The result shows there is higher tool wear with $th_c$ compared to $th_p$ mode at both speeds and this is also consistent with the
addition of micro-lubrication. For both $th_c$ and $th_p$ modes, notably at high cutting speed, adding micro-lubrication significantly reduces tool wear. However, with lower cutting speeds, the impact of micro-lubrication is not that prominent. The pulsed and lubricated method ($th_p-l$) in milling Waspaloy is found to be the best method for high productivity as the variance in VB is not significant for high and low cutting speeds, hence milder tool wear and increase in productivity while reducing tool cost.

Tool wear generally decreases for all the Lubri-cooling modes from 120 to 80m/min, except the $th_c-l$ mode. At high cutting speed, $th_p$ causes more severe tool deterioration than the $th_c$ mode, however, the addition of micro-lubrication to the pulse mode lowers the tool wear to a similar value as that of $th_c-l$ mode. At low cutting speed, all Lubri-cooling modes yield comparable tool wear for milling Inconel 718 except for $th_p-l$ which yielded the smallest tool wear. It can be concluded that the $th_c-l$ and $th_p-l$ modes of Lubri-cooling are the most effective at 120m/min and 80m/min respectively.

<figure><img src="image_1979.png" /><figcaption>Fig. 2. Optical micrographs of the selected tools used in the milling of the three alloys.</figcaption></figure>

Figure 2 presents the selected optical micrographs of the cutting teeth of the cutters used in the milling of the three superalloys. The most prevalent type of tool wear that occurs while milling high-strength materials is chipping of the cutting edges. Progressive wear is also evident on the cutting edges that have not sustained severe chipping, albeit to a considerably lesser extent. All the images chosen are that of 120m/min cutting speed and are grouped into two rows of continuous CO₂ and pulsed CO₂. For Incoloy 825, chipping is not evident with the continuous mode, and the pulsed mode has yielded little indication of wear. In the case of Waspaloy, the milling cutter used at high cutting speed and under continuous mode has deteriorated severely. The pulsed mode with micro-lubrication underwent mild tool wear. Like Incoloy 825 and Waspaloy, Inconel 718 also shows tool wear sustained through chipping and progressive mechanical wear. The $th_c$ mode shows noticeable tool wear, however, there was little tool wear observed under the $th_p$ mode with the addition of micro-lubrication.

## Cutting Forces

Table 3 suggests that, of the three orthogonal force components, the maximum magnitude is observed for the planar force component acting perpendicular to the feed direction ($F_y$), followed by the ones in the feed direction ($F_x$) and the axial direction ($F_z$), respectively. This is true for all the runs except those employing $th_c$ at the low level of cutting speed in the milling of Waspaloy.
addition of micro-lubrication. For both $th_c$ and $th_p$ modes, notably at high cutting speed, adding micro-lubrication significantly reduces tool wear. However, with lower cutting speeds, the impact of micro-lubrication is not that prominent. The pulsed and lubricated method ($th_p-l$) in milling Waspaloy is found to be the best method for high productivity as the variance in VB is not significant for high and low cutting speeds, hence milder tool wear and increase in productivity while reducing tool cost.

Tool wear generally decreases for all the Lubri-cooling modes from 120 to 80m/min, except the $th_c-l$ mode. At high cutting speed, $th_p$ causes more severe tool deterioration than the $th_c$ mode, however, the addition of micro-lubrication to the pulse mode lowers the tool wear to a similar value as that of $th_c-l$ mode. At low cutting speed, all Lubri-cooling modes yield comparable tool wear for milling Inconel 718 except for $th_p-l$ which yielded the smallest tool wear. It can be concluded that the $th_c-l$ and $th_p-l$ modes of Lubri-cooling are the most effective at 120m/min and 80m/min respectively.

<figure><img src="image_1974.png" /><figcaption>Fig. 2. Optical micrographs of the selected tools used in the milling of the three alloys.</figcaption></figure>

Figure 2 presents the selected optical micrographs of the cutting teeth of the cutters used in the milling of the three superalloys. The most prevalent type of tool wear that occurs while milling high-strength materials is chipping of the cutting edges. Progressive wear is also evident on the cutting edges that have not sustained severe chipping, albeit to a considerably lesser extent. All the images chosen are that of 120m/min cutting speed and are grouped into two rows of continuous CO₂ and pulsed CO₂. For Incoloy 825, chipping is not evident with the continuous mode, and the pulsed mode has yielded little indication of wear. In the case of Waspaloy, the milling cutter used at high cutting speed and under continuous mode has deteriorated severely. The pulsed mode with micro-lubrication underwent mild tool wear. Like Incoloy 825 and Waspaloy, Inconel 718 also shows tool wear sustained through chipping and progressive mechanical wear. The $th_c$ mode shows noticeable tool wear, however, there was little tool wear observed under the $th_p$ mode with the addition of micro-lubrication.

## Cutting Forces

Table 3 suggests that, of the three orthogonal force components, the maximum magnitude is observed for the planar force component acting perpendicular to the feed direction ($F_y$), followed by the ones in the feed direction ($F_x$) and the axial direction ($F_z$), respectively. This is true for all the runs except those employing $th_c$ at the low level of cutting speed in the milling of Waspaloy.
<figure><img src="image_1980.png" /><figcaption>Fig. 3. The time-force graphs of the three orthogonal force components for the selected lubri-cooling modes of milling at the cutting speed of 120m/min</figcaption></figure>

The cutting force results indicate that an increase in cutting speed results in an increase in the cutting force. The statement stands true for all three workpiece materials. Table 1 suggests that all three alloys possess contents of nickel over 40%. Thermal softening is a phenomenon that occurs when the generated process heat counteracts the strain-hardening effect. It is widely known that an increase in nickel content significantly increases work hardening, counteracting the effect of thermal softening on the workpiece material during plastic deformation. Thus, it is expected that the higher cutting speed would result in higher cutting forces.

The time-force graphs of the three orthogonal force components for the milling of Incoloy 825 (pulsed lubri-cooling), Inconel 718 (continuous cryogenic cooling), and Waspaloy (dry) carried out at the cutting speed of 120 m/min are shown in Figure 3. The shape patterns of the three force components for all the runs are more or less the same. Due to the dynamics of the side-milling process and the geometry of the tool, $F_x$ and $F_y$ take the higher shares of the force magnitude.

For all three alloys, the $F_z$ graph shows that the magnitude of each of the last two passes has increased in comparison to the one before it. Furthermore, as each of the three passes goes on, the force magnitude increases. These imply that the force magnitude is uniformly increasing due to
<figure><img src="image_1975.png" /><figcaption>Fig. 3. The time-force graphs of the three orthogonal force components for the selected lubri-cooling modes of milling at the cutting speed of 120m/min</figcaption></figure>

The cutting force results indicate that an increase in cutting speed results in an increase in the cutting force. The statement stands true for all three workpiece materials. Table 1 suggests that all three alloys possess contents of nickel over 40%. Thermal softening is a phenomenon that occurs when the generated process heat counteracts the strain-hardening effect. It is widely known that an increase in nickel content significantly increases work hardening, counteracting the effect of thermal softening on the workpiece material during plastic deformation. Thus, it is expected that the higher cutting speed would result in higher cutting forces.

The time-force graphs of the three orthogonal force components for the milling of Incoloy 825 (pulsed lubri-cooling), Inconel 718 (continuous cryogenic cooling), and Waspaloy (dry) carried out at the cutting speed of 120 m/min are shown in Figure 3. The shape patterns of the three force components for all the runs are more or less the same. Due to the dynamics of the side-milling process and the geometry of the tool, $F_x$ and $F_y$ take the higher shares of the force magnitude.

For all three alloys, the $F_z$ graph shows that the magnitude of each of the last two passes has increased in comparison to the one before it. Furthermore, as each of the three passes goes on, the force magnitude increases. These imply that the force magnitude is uniformly increasing due to
the increasing tool wear. However, the other two force component graphs show less of this behavior.

<figure><img src="image_1981.png" /><figcaption>Fig. 4. The optical micrographs of the surfaces of the workpieces side-milled in selected runs of the three alloys</figcaption></figure>

## Work Surface Quality

The last column of Table 3 shows the average arithmetic roughness value of the surfaces generated in the 30 runs. The $R_a$ measurements overall range between 1 and 3µm which is considerably high and can be attributed to the dynamics of the side milling process. For all three alloys, it can be deduced that the introduction of micro-lubrication to the milling process plays a major role in reducing the surface roughness and has increased the beneficial effect of the cryogenic coolant, particularly at the low cutting speed. This can be further emphasized through the observation that the finest surface generated amongst all the milling processes of the three alloys employed either continuous or pulsed cooling mode, mixed with micro-lubrication, and carried out at 80m/min. The introduction of micro-lubrication to the shear zones forms a minute film so the cutting process experiences reduced friction.

Fig. 4 presents the optical micrographs of the side-milled surfaces of the selected runs. The optical micrograph for Incoloy 835 presents the roughest of all the surfaces generated in this study. The attributed reason is the absence of any lubri-coolant and cutting at the high level of cutting speed. The processes related to the optical micrograph of Inconel 718 and Waspaloy shown in the figure were carried out at 80 m/min cutting speed using $th_c-l$ and $th_p-l$, respectively. It is quite evident from the results and the micrographs that a superior surface finish is obtained through the parallel application of micro-lubrication with the cryogenic coolant no matter whether the coolant is supplied in a continuous or pulsed mode.

## Summary

The presented work pitches the dry cutting approach – considered to be a sustainable way of machining a variety of engineering materials – against the throttle cryogenic medium-based lubri-cooling approach. The article has also explored the favorable outcomes of supplying the cryogenic
the increasing tool wear. However, the other two force component graphs show less of this behavior.

<figure><img src="image_1976.png" /><figcaption>Fig. 4. The optical micrographs of the surfaces of the workpieces side-milled in selected runs of the three alloys</figcaption></figure>

## Work Surface Quality

The last column of Table 3 shows the average arithmetic roughness value of the surfaces generated in the 30 runs. The $R_a$ measurements overall range between 1 and 3µm which is considerably high and can be attributed to the dynamics of the side milling process. For all three alloys, it can be deduced that the introduction of micro-lubrication to the milling process plays a major role in reducing the surface roughness and has increased the beneficial effect of the cryogenic coolant, particularly at the low cutting speed. This can be further emphasized through the observation that the finest surface generated amongst all the milling processes of the three alloys employed either continuous or pulsed cooling mode, mixed with micro-lubrication, and carried out at 80m/min. The introduction of micro-lubrication to the shear zones forms a minute film so the cutting process experiences reduced friction.

Fig. 4 presents the optical micrographs of the side-milled surfaces of the selected runs. The optical micrograph for Incoloy 835 presents the roughest of all the surfaces generated in this study. The attributed reason is the absence of any lubri-coolant and cutting at the high level of cutting speed. The processes related to the optical micrograph of Inconel 718 and Waspaloy shown in the figure were carried out at 80 m/min cutting speed using $th_c-l$ and $th_p-l$, respectively. It is quite evident from the results and the micrographs that a superior surface finish is obtained through the parallel application of micro-lubrication with the cryogenic coolant no matter whether the coolant is supplied in a continuous or pulsed mode.

## Summary

The presented work pitches the dry cutting approach – considered to be a sustainable way of machining a variety of engineering materials – against the throttle cryogenic medium-based lubri-cooling approach. The article has also explored the favorable outcomes of supplying the cryogenic
coolant in the mode of periodic pulses. The findings of the experimental study indicate that the lowering of tool wear and cutting forces during the milling of nickel-based superalloys, such as Incoloy 825, Inconel 718, and Waspaloy can be achieved using throttling CO2 as a cooling medium. The cryogenic coolant demonstrates favorable outcomes with reduced cutting forces and better surface finish, especially when applied in parallel with micro-lubrication. While the pulsed mode of the coolant's supply counteracts thermal gradients to yield a better surface finish and reduce cutting force in most cases, the continuous mode offers a richer cooling potential to guarantee a longer tool life.

Cutting speed is found to have little effect on tool wear when cryogenic cooling is used, so higher speeds can be used to mill the nickel-based superalloys sustainably. The addition of micro-lubrication enhances the favorable capability of the cryogenic coolant to improve the surface quality. The study suggests that the continuous mode of the coolant's supply is highly promising for productivity, yielding the lowest tool wear and the pulsed mode is suitable for reducing cutting forces and improving surface finish.

## References

[1] D. Zhu, X. Zhang, H. Ding, Tool wear characteristics in machining of nickel-based superalloys. Int J Mach. Tools Manuf. 64 (2013), 60-77.

[2] https://doi.org/10.1016/j.ijmachtools.2012.08.001

[3] M. Anderson, R. Patwa, Y.C. Shin, Laser-assisted machining of Inconel 718 with an economic analysis. Int J Mach. Tools Manuf. 46 (2006)1879-1891. https://doi.org/10.1016/j.ijmachtools.2005.11.005

[4] M. Günay, M.E. Korkmaz, N. Yaşar, Performance analysis of coated carbide tool in turning of Nimonic 80A superalloy under different cutting environments. J Manuf. Process. 56 (2020) 678-687. https://doi.org/10.1016/j.jmapro.2020.05.031

[5] S. Rinaldi, S. Caruso, D. Umbrello, L. Filice, R. Franchi, A. Del Prete, Machinability of Waspaloy under different cutting and lubri-cooling conditions. Int. J Adv. Manuf. Technol.; 94 (2018), 3703-3712. https://doi.org/10.1007/s00170-017-1133-0

[6] A. Shokrani, S.T. Newman, Hybrid cooling and lubricating technology for CNC milling of Inconel 718 nickel alloy. Proc. 8th CIRP Conf. High Performance Cutting (HPC 2018) 77 (2018), 215-218. https://doi.org/10.1016/j.procir.2018.08.291

[7] M. Danish, M.K. Gupta, S. Rubaiee, A. Ahmed, M.E. Korkmaz, Influence of hybrid Cryo-MQL lubri-cooling strategy on the machining and tribological characteristics of Inconel 718. Tribol. Int. 163 (2021), 107178. https://doi.org/10.1016/j.triboint.2021.107178

[8] A. Bagherzadeh, B. Koc, E. Budak, M. Isik, High-speed machining of additively manufactured Inconel 718 using hybrid cryogenic cooling methods. Virt. Phys. Prototyp 17 (2022), 419-436. https://doi.org/10.1080/17452759.2022.2034081

[9] K. Ross, G. Manimara, Machining investigation of Nimonic-80A superalloy under cryogenic CO2 as coolant using PVD-TiAlN/TiN coated tool at 45° nozzle angle. Arab J Sci. Eng. 45 (2020), 9267-9281. https://doi.org/10.1007/s13369-020-04728-8

[10] R. Polvorosa, A. Suarez, L.N. Lopez de Lacalle, I. Cerrillo, A. Wretland, F. Veiga, Tool wear on nickel alloys with different coolant pressures: Comparison of Alloy 718 and Waspaloy. J Manuf. Process. 26 (2017), 44-56. https://doi.org/10.1016/j.jmapro.2017.01.012
coolant in the mode of periodic pulses. The findings of the experimental study indicate that the lowering of tool wear and cutting forces during the milling of nickel-based superalloys, such as Incoloy 825, Inconel 718, and Waspaloy can be achieved using throttling CO2 as a cooling medium. The cryogenic coolant demonstrates favorable outcomes with reduced cutting forces and better surface finish, especially when applied in parallel with micro-lubrication. While the pulsed mode of the coolant's supply counteracts thermal gradients to yield a better surface finish and reduce cutting force in most cases, the continuous mode offers a richer cooling potential to guarantee a longer tool life.

Cutting speed is found to have little effect on tool wear when cryogenic cooling is used, so higher speeds can be used to mill the nickel-based superalloys sustainably. The addition of micro-lubrication enhances the favorable capability of the cryogenic coolant to improve the surface quality. The study suggests that the continuous mode of the coolant's supply is highly promising for productivity, yielding the lowest tool wear and the pulsed mode is suitable for reducing cutting forces and improving surface finish.

## References

[1] D. Zhu, X. Zhang, H. Ding, Tool wear characteristics in machining of nickel-based superalloys. Int J Mach. Tools Manuf. 64 (2013), 60-77.

[2] https://doi.org/10.1016/j.ijmachtools.2012.08.001

[3] M. Anderson, R. Patwa, Y.C. Shin, Laser-assisted machining of Inconel 718 with an economic analysis. Int J Mach. Tools Manuf. 46 (2006)1879-1891. https://doi.org/10.1016/j.ijmachtools.2005.11.005

[4] M. Günay, M.E. Korkmaz, N. Yaşar, Performance analysis of coated carbide tool in turning of Nimonic 80A superalloy under different cutting environments. J Manuf. Process. 56 (2020) 678-687. https://doi.org/10.1016/j.jmapro.2020.05.031

[5] S. Rinaldi, S. Caruso, D. Umbrello, L. Filice, R. Franchi, A. Del Prete, Machinability of Waspaloy under different cutting and lubri-cooling conditions. Int. J Adv. Manuf. Technol.; 94 (2018), 3703-3712. https://doi.org/10.1007/s00170-017-1133-0

[6] A. Shokrani, S.T. Newman, Hybrid cooling and lubricating technology for CNC milling of Inconel 718 nickel alloy. Proc. 8th CIRP Conf. High Performance Cutting (HPC 2018) 77 (2018), 215-218. https://doi.org/10.1016/j.procir.2018.08.291

[7] M. Danish, M.K. Gupta, S. Rubaiee, A. Ahmed, M.E. Korkmaz, Influence of hybrid Cryo-MQL lubri-cooling strategy on the machining and tribological characteristics of Inconel 718. Tribol. Int. 163 (2021), 107178. https://doi.org/10.1016/j.triboint.2021.107178

[8] A. Bagherzadeh, B. Koc, E. Budak, M. Isik, High-speed machining of additively manufactured Inconel 718 using hybrid cryogenic cooling methods. Virt. Phys. Prototyp 17 (2022), 419-436. https://doi.org/10.1080/17452759.2022.2034081

[9] K. Ross, G. Manimara, Machining investigation of Nimonic-80A superalloy under cryogenic CO2 as coolant using PVD-TiAlN/TiN coated tool at 45° nozzle angle. Arab J Sci. Eng. 45 (2020), 9267-9281. https://doi.org/10.1007/s13369-020-04728-8

[10] R. Polvorosa, A. Suarez, L.N. Lopez de Lacalle, I. Cerrillo, A. Wretland, F. Veiga, Tool wear on nickel alloys with different coolant pressures: Comparison of Alloy 718 and Waspaloy. J Manuf. Process. 26 (2017), 44-56. https://doi.org/10.1016/j.jmapro.2017.01.012
[11] X. Lu, Z. Jia, H. Wang, L. Si, Y. Liu, W. Wu, Tool wear appearance and failure mechanism of coated carbide tools in micro-milling of Inconel 718 super alloy. Indust. Lub. & Tribol. 68 (2016), 267-277. https://doi.org/10.1108/ILT-07-2015-0114

[12] F. Wang, Y. Wang, Comparison of cryogenic cooling strategy effects on machinability of milling nickel-based alloy. J Manuf. Proc. 66 (2021), 623-635. https://doi.org/10.1016/j.jmapro.2021.04.050

[13] S. Chaabani, I. Rodriguez, M. Cuesta, Y. Ayed, P. J. Arrazola, G. Germain. Tool wear and cutting forces when machining inconel 718 under cryogenic conditions: Liquid nitrogen and carbon dioxide. In AIP Conference Proceedings, vol. 2113, no. 1 (2019). AIP Publishing.
[11] X. Lu, Z. Jia, H. Wang, L. Si, Y. Liu, W. Wu, Tool wear appearance and failure mechanism of coated carbide tools in micro-milling of Inconel 718 super alloy. Indust. Lub. & Tribol. 68 (2016), 267-277. https://doi.org/10.1108/ILT-07-2015-0114

[12] F. Wang, Y. Wang, Comparison of cryogenic cooling strategy effects on machinability of milling nickel-based alloy. J Manuf. Proc. 66 (2021), 623-635. https://doi.org/10.1016/j.jmapro.2021.04.050

[13] S. Chaabani, I. Rodriguez, M. Cuesta, Y. Ayed, P. J. Arrazola, G. Germain. Tool wear and cutting forces when machining inconel 718 under cryogenic conditions: Liquid nitrogen and carbon dioxide. In AIP Conference Proceedings, vol. 2113, no. 1 (2019). AIP Publishing.
# Innovative cryogenic processing for enhanced surface integrity characteristics of stainless steels

KITAY Ozhan¹, a* and KAYNAK Yusuf², b

¹ Department of Machine and Metal Technologies, Bilecik Seyh Edebali University, 11100 Bilecik, Türkiye

² Department of Mechanical Engineering, Marmara University, 34854, Istanbul, Türkiye

a ozhan.kitay@bilecik.edu.tr, b yusuf.kaynak@marmara.edu.tr

**Keywords:** Liquid Nitrogen (LN₂), Cryogenic, Stainless Steel, Surface Integrity, Machining

**Abstract.** In this study, the influence of the orthogonal cutting on the surface integrity characteristics of 316L austenitic stainless steel material under liquid nitrogen (LN₂) condition was investigated. For this purpose, machining tests were carried out by applying LN₂ to the cutting tool rake and flank face separately, and using different cutting parameters. In addition to the cutting force, the response of surface integrity characteristics such as microstructure, microhardness, and x-ray diffraction (XRD) analysis to LN₂ was compared with dry cutting. The findings showed that the effect of different positioning of the LN₂ nozzle on the cutting force and microhardness of the material is significant. In particular, the application of liquid nitrogen through the rake face significantly increases the microhardness of the material, while machined specimen microstructure images show that application of LN₂ from the flank surface causes more grain refinement and deformation on the machined surface, especially at low cutting speed.

## Introduction

The 316L austenitic stainless steel has a wide range of uses in biomedical implant manufacturing due to its biocompatibility and is also used in the automotive industry [1-3]. Stainless steel materials are widely used in many industries due to their high strength, toughness, mechanical properties at high temperatures and corrosion resistance [4]. They are called hard-to-cut materials because these characteristic features create difficulties in machining operations [5]. The high cutting forces and cutting temperature during machining of these materials affect the tool wear and the chip breakability. The surface quality, fatigue life and corrosion resistance of the finished product are also affected by these problems [6, 7].

Numerous lubricants and coolants are used to improve the machining performance of the materials in the hard-to-cut materials group [8-10]. In comparison with other cooling and lubrication methods, cryogenic cooling method is an effective method for human and environmental health, besides liquid nitrogen (LN₂) is widely used as cryogenic coolant in the machining process [11, 12]. In the machining process, cryogenic coolant can be applied to the tool-chip interface or directly to the workpiece. Thus, it penetrates the cutting zone between the tool and chip and reduces the friction coefficient [13-15]. Cryogenic machining is also preferred because it positively affects the surface integrity characteristics of the machined part, such as surface and subsurface hardness change, microstructure, phase transformation, residual stress and fatigue life, which directly affect product performance [15-17].

There are studies in the literature in which cryogenic coolants are applied to increase the functional performance of difficult-to-cut materials in terms of fatigue life, corrosion resistance and wear, and their effects on surface hardness, phase transformation and residual stress formation are also examined [9, 14, 17-19]. Although there are studies examining the changes in the surface

<figure><img src="image_1982.png" /></figure>
# Innovative cryogenic processing for enhanced surface integrity characteristics of stainless steels

KITAY Ozhan¹, a * and KAYNAK Yusuf², b

¹ Department of Machine and Metal Technologies, Bilecik Seyh Edebali University, 11100 Bilecik, Türkiye

² Department of Mechanical Engineering, Marmara University, 34854, Istanbul, Türkiye

a ozhan.kitay@bilecik.edu.tr, b yusuf.kaynak@marmara.edu.tr

**Keywords:** Liquid Nitrogen (LN₂), Cryogenic, Stainless Steel, Surface Integrity, Machining

**Abstract.** In this study, the influence of the orthogonal cutting on the surface integrity characteristics of 316L austenitic stainless steel material under liquid nitrogen (LN₂) condition was investigated. For this purpose, machining tests were carried out by applying LN₂ to the cutting tool rake and flank face separately, and using different cutting parameters. In addition to the cutting force, the response of surface integrity characteristics such as microstructure, microhardness, and x-ray diffraction (XRD) analysis to LN₂ was compared with dry cutting. The findings showed that the effect of different positioning of the LN₂ nozzle on the cutting force and microhardness of the material is significant. In particular, the application of liquid nitrogen through the rake face significantly increases the microhardness of the material, while machined specimen microstructure images show that application of LN₂ from the flank surface causes more grain refinement and deformation on the machined surface, especially at low cutting speed.

## Introduction

The 316L austenitic stainless steel has a wide range of uses in biomedical implant manufacturing due to its biocompatibility and is also used in the automotive industry [1-3]. Stainless steel materials are widely used in many industries due to their high strength, toughness, mechanical properties at high temperatures and corrosion resistance [4]. They are called hard-to-cut materials because these characteristic features create difficulties in machining operations [5]. The high cutting forces and cutting temperature during machining of these materials affect the tool wear and the chip breakability. The surface quality, fatigue life and corrosion resistance of the finished product are also affected by these problems [6, 7].

Numerous lubricants and coolants are used to improve the machining performance of the materials in the hard-to-cut materials group [8-10]. In comparison with other cooling and lubrication methods, cryogenic cooling method is an effective method for human and environmental health, besides liquid nitrogen (LN₂) is widely used as cryogenic coolant in the machining process [11, 12]. In the machining process, cryogenic coolant can be applied to the tool-chip interface or directly to the workpiece. Thus, it penetrates the cutting zone between the tool and chip and reduces the friction coefficient [13-15]. Cryogenic machining is also preferred because it positively affects the surface integrity characteristics of the machined part, such as surface and subsurface hardness change, microstructure, phase transformation, residual stress and fatigue life, which directly affect product performance [15-17].

There are studies in the literature in which cryogenic coolants are applied to increase the functional performance of difficult-to-cut materials in terms of fatigue life, corrosion resistance and wear, and their effects on surface hardness, phase transformation and residual stress formation are also examined [9, 14, 17-19]. Although there are studies examining the changes in the surface

<figure><img src="image_1977.png" /></figure>
integrity characteristics of cryogenic machining of many different materials in the group of hard-to-cut materials, studies on the surface integrity characteristics of the cryogenic machining of 316L stainless steel have been limited in particular [20].

In this study, the effects of applying LN₂ to cutting area in the 316L stainless steel machining process from rake and flank face at different cutting parameters on the surface integrity characteristics were determined and compared with dry condition in terms of microhardness, microstructure and X-ray diffraction(XRD) analysis.

## Experimental Setup

The workpieces used in this study were 316L austenitic stainless steel as discs of 50 mm diameter and 3 mm thickness for orthogonal cutting. Hardness of workpiece is 190±5 HV at room temperature. Discs were machined from 50 mm to 33 mm diameter in orthogonal cutting. The chemical composition and mechanical properties of workmaterial are shown in Table 1 and Table 2, respectively.

<table><caption>Table 1. Chemical composition of 316L austenitic stainless steel [21]</caption><thead><tr><th>ELEMENT</th><th>C</th><th>M</th><th>Si</th><th>S</th><th>P</th><th>Cr</th><th>Ni</th><th>Mo</th><th>N</th><th>Fe</th></tr></thead><tbody><tr><th>%</th><td>0.02</td><td>1.77</td><td>0.46</td><td>0.027</td><td>0.029</td><td>16.76</td><td>10.09</td><td>2.02</td><td>0.06</td><td>Bal.</td></tr></tbody></table>

<table><caption>Table 2. Mechanical Properties of 316L austenitic stainless steel [21]</caption><thead><tr><th>Yield Stress</th><th>Compression Stress</th><th>Elongation</th><th>Elasticity Module</th></tr></thead><tbody><tr><td>319 MPa</td><td>628 MPa</td><td>55 %</td><td>200 GPa</td></tr></tbody></table>

The machining experiments were conducted on a Doosan CNC turning center which has maximum spindle speed of 4500 rpm and 18 kW power. In machining operations, TCMW16T308 HTİ10 with -2° rake angle, 7° clearance angle and 23 µm edge radius uncoated carbide cutting tool was used. Cutting speeds were selected as 40, 120 and 200 m/min and uncut chip thicknesses ($t_o$) were 0.08, 0.12 and 0.16 mm. Machining operations were performed under dry machining without any coolant/lubricant, and cryogenic cooling with LN₂ (15 bar) from rake face and flank face conditions. Implementation of cryogenic machining is shown in Fig. 1. All experiments were performed twice for reproducibility. Cutting forces were measured with KISTLER 2129AA dynamometer.

<figure><img src="image_1983.png" /><figcaption>Fig. 1. Applying LN₂ in orthogonal machining a) from rake face b) from flank face</figcaption></figure>

To examine the surface integrity characteristics (microhardness, microstructure, phase transformation), machined workpieces were cut with diamond cutting disk then cut specimens were cold-mounted in acrylic and then polished using 60, 30 and 15 µm grit magnetic discs. Specimens were etched using a solution of 2 ml HNO₃ + 4 ml Glycerine + 6 ml HCl to review the microstructure [22]. Microstructure of etched specimens were examined by Olympus BX51M Optic Microscopy. Microhardness of workpiece was measured by using Future-Tech FM310e.
integrity characteristics of cryogenic machining of many different materials in the group of hard-to-cut materials, studies on the surface integrity characteristics of the cryogenic machining of 316L stainless steel have been limited in particular [20].

In this study, the effects of applying LN₂ to cutting area in the 316L stainless steel machining process from rake and flank face at different cutting parameters on the surface integrity characteristics were determined and compared with dry condition in terms of microhardness, microstructure and X-ray diffraction(XRD) analysis.

## Experimental Setup

The workpieces used in this study were 316L austenitic stainless steel as discs of 50 mm diameter and 3 mm thickness for orthogonal cutting. Hardness of workpiece is 190±5 HV at room temperature. Discs were machined from 50 mm to 33 mm diameter in orthogonal cutting. The chemical composition and mechanical properties of workmaterial are shown in Table 1 and Table 2, respectively.

<table><caption>Table 1. Chemical composition of 316L austenitic stainless steel [21]</caption><thead><tr><th>ELEMENT</th><th>C</th><th>M</th><th>Si</th><th>S</th><th>P</th><th>Cr</th><th>Ni</th><th>Mo</th><th>N</th><th>Fe</th></tr></thead><tbody><tr><th>%</th><td>0.02</td><td>1.77</td><td>0.46</td><td>0.027</td><td>0.029</td><td>16.76</td><td>10.09</td><td>2.02</td><td>0.06</td><td>Bal.</td></tr></tbody></table>

<table><caption>Table 2. Mechanical Properties of 316L austenitic stainless steel [21]</caption><thead><tr><th>Yield Stress</th><th>Compression Stress</th><th>Elongation</th><th>Elasticity Module</th></tr></thead><tbody><tr><td>319 MPa</td><td>628 MPa</td><td>55 %</td><td>200 GPa</td></tr></tbody></table>

The machining experiments were conducted on a Doosan CNC turning center which has maximum spindle speed of 4500 rpm and 18 kW power. In machining operations, TCMW16T308 HTİ10 with -2° rake angle, 7° clearance angle and 23 µm edge radius uncoated carbide cutting tool was used. Cutting speeds were selected as 40, 120 and 200 m/min and uncut chip thicknesses ($t_o$) were 0.08, 0.12 and 0.16 mm. Machining operations were performed under dry machining without any coolant/lubricant, and cryogenic cooling with LN₂ (15 bar) from rake face and flank face conditions. Implementation of cryogenic machining is shown in Fig. 1. All experiments were performed twice for reproducibility. Cutting forces were measured with KISTLER 2129AA dynamometer.

<figure><img src="image_1978.png" /><figcaption>Fig. 1. Applying LN₂ in orthogonal machining a) from rake face b) from flank face</figcaption></figure>

To examine the surface integrity characteristics (microhardness, microstructure, phase transformation), machined workpieces were cut with diamond cutting disk then cut specimens were cold-mounted in acrylic and then polished using 60, 30 and 15 µm grit magnetic discs. Specimens were etched using a solution of 2 ml HNO₃ + 4 ml Glycerine + 6 ml HCl to review the microstructure [22]. Microstructure of etched specimens were examined by Olympus BX51M Optic Microscopy. Microhardness of workpiece was measured by using Future-Tech FM310e.
Each hardness of specimen was determined by average of four measurement after applying 25 gf during 15 second to machined specimens. X-ray diffraction (XRD) was used by Bruker AXS D2 Phaser XRD device. CuKα cathode with wavelength of 1.54060 Å, current of 10 mA and voltage of 30 kV were used to analysis the XRD patterns of machined specimens.

## Experimental Results and Discussion

Cutting Forces. Cutting forces is an useful indicator to evaluation the machining performance and explain the reasons for the origin of the machining. Cutting forces measured in orthogonal machining process under different cutting conditions is shown in Fig. 2 as a function of cutting speed. A significant reduction in cutting forces is seen with increased cutting speed. The reduction in cutting forces is not only seen in dry cutting, but also in all cutting conditions. Liquid LN₂ evaporates in the cutting area and reduces friction by creating an LN₂ cushion between the tool and the chip. This reduces the strength of the welded junctions between the tool and the chip, allowing less force to be consumed compared to dry cutting. Increasing the cutting speed increases the deformation rate of the workpiece material, resulting in thermal softening of the material and ultimately reduction of the cutting forces. Measured cutting forces when cryogenic coolant applied from rake face is lower than cryogenic applied from flank face. The low cutting forces at the rake face condition can be explained by the effective penetration of the cryogenic coolant at the tool-chip interface and the reduction of friction. This is supported by the fact that the percentage differences are especially high at low cutting speeds. There is a difference of 30% at a cutting speed of 40 m/min. There is also a difference of 12% between the lowest and highest cutting force values at a cutting speed of 120 m/min. Therefore, mechanical effect plays a dominant role at low cutting speeds and when the friction is higher, application of liquid nitrogen through the rake face has shown lubrication effect and reduced forces. By applying liquid nitrogen from the flank face, reductions in cutting forces occurred similar to the rake face. In fact, the workpiece is cooled and this increases its resistance to plastic deformation compared to rake face.

<figure><img src="image_1984.png" /><figcaption>Fig. 2. Measured cutting forces in orthogonal turning as a function of cutting speed (t₀= 0.08 mm)</figcaption></figure>

Cutting forces measured in orthogonal machining process under different cutting conditions is shown in Fig. 3 as a function of cutting depth. While the highest cutting forces were measured in dry cutting, the lowest cutting forces were obtained by applying liquid nitrogen to the rake face. The fact that increasing cutting depth does not negatively affect the penetration ability of liquid nitrogen can be considered as an additional finding. It is expected that the increase in cutting depth causes an increase in cutting forces due to the increase in chip cross-section. Taking dry cutting as
Each hardness of specimen was determined by average of four measurement after applying 25 gf during 15 second to machined specimens. X-ray diffraction (XRD) was used by Bruker AXS D2 Phaser XRD device. CuKα cathode with wavelength of 1.54060 Å, current of 10 mA and voltage of 30 kV were used to analysis the XRD patterns of machined specimens.

## Experimental Results and Discussion

Cutting Forces. Cutting forces is an useful indicator to evaluation the machining performance and explain the reasons for the origin of the machining. Cutting forces measured in orthogonal machining process under different cutting conditions is shown in Fig. 2 as a function of cutting speed. A significant reduction in cutting forces is seen with increased cutting speed. The reduction in cutting forces is not only seen in dry cutting, but also in all cutting conditions. Liquid LN₂ evaporates in the cutting area and reduces friction by creating an LN₂ cushion between the tool and the chip. This reduces the strength of the welded junctions between the tool and the chip, allowing less force to be consumed compared to dry cutting. Increasing the cutting speed increases the deformation rate of the workpiece material, resulting in thermal softening of the material and ultimately reduction of the cutting forces. Measured cutting forces when cryogenic coolant applied from rake face is lower than cryogenic applied from flank face. The low cutting forces at the rake face condition can be explained by the effective penetration of the cryogenic coolant at the tool-chip interface and the reduction of friction. This is supported by the fact that the percentage differences are especially high at low cutting speeds. There is a difference of 30% at a cutting speed of 40 m/min. There is also a difference of 12% between the lowest and highest cutting force values at a cutting speed of 120 m/min. Therefore, mechanical effect plays a dominant role at low cutting speeds and when the friction is higher, application of liquid nitrogen through the rake face has shown lubrication effect and reduced forces. By applying liquid nitrogen from the flank face, reductions in cutting forces occurred similar to the rake face. In fact, the workpiece is cooled and this increases its resistance to plastic deformation compared to rake face.

<figure><img src="image_1979.png" /><figcaption>Fig. 2. Measured cutting forces in orthogonal turning as a function of cutting speed (t₀ = 0.08 mm)</figcaption></figure>

Cutting forces measured in orthogonal machining process under different cutting conditions is shown in Fig. 3 as a function of cutting depth. While the highest cutting forces were measured in dry cutting, the lowest cutting forces were obtained by applying liquid nitrogen to the rake face. The fact that increasing cutting depth does not negatively affect the penetration ability of liquid nitrogen can be considered as an additional finding. It is expected that the increase in cutting depth causes an increase in cutting forces due to the increase in chip cross-section. Taking dry cutting as
a reference at a cutting depth of 0.16 mm, the decrease in cutting force caused by the application of liquid nitrogen from the rake surface is approximately 20%. However, this ratio is less at the lowest cutting depth. Therefore, it is seen that the lubricating effect of liquid nitrogen plays an active role here as well and can reduce the force more when the friction force is greater.

<figure><img src="image_1985.png" /><figcaption>Fig. 3. Effect of cutting depth on cutting forces (V=120 m/min).</figcaption></figure>

Microhardness. The variation of hardness in the surface and subsurface areas of the machined parts is important depending on where the part will be used. While in some cases it is not desired to change the hardness in order to create a more homogeneous structure, in some cases it is a desirable feature to increase the hardness of the surface, especially in parts that are in contact with each other or subject to wear [23, 24]. 316L has a lower workability than other stainless steels and has a high tendency to hardening during machining [3]. Thermal effects during machining and martensite formation in the austenite structure during plastic deformation are considered as the reason for the tendency of hardening [25, 26].

Microhardness values of orthogonally machined samples under different cutting conditions is shown in Fig. 4. There is an increase in the hardness of the machined samples from surface and sub-surface to 100 µm depth in all of conditions. While the highest hardness value was measured as 282 HV under cryogenic condition applied from rake face, the lowest hardness value was measured as 231.5 HV under dry condition at the distance of 15 µm from the surface. There is a difference of 22% between the lowest and the highest hardness values. Also, there are increase of 44% and 18% in hardness of machined samples as compared to as-received hardness. LN₂ directly penetrates the tool-chip region from rake face. Therefore, since the primary and secondary deformation region temperature is directly affected, chip breakability becomes easier, but the material hardness increases. However, when LN₂ penetrates from the flank face, the hardness increase is less as the clearance zone is affected.

In Fig. 5, the variation of the hardness from the machined surface to the center of the workpiece can be expressed in three steps. The first stage is hardened very clearly but hardness values is lower than beginning of the second stage. The reason for this behavior is that the first 10 µm layer, which is the closest layer to the surface, is the most exposed to heat and high plastic deformation. This has naturally caused a relative decrease in hardness. At the starting point of the second stage, the hardness value reached its maximum point. While the material is exposed to a large plastic deformation, it is exposed to less heat. The depth of this layer determines the surface and the deformed layer under the surface in the result of the machining process. In the third stage, the as-received material hardness is active and the effect of the machining process is not exist.
a reference at a cutting depth of 0.16 mm, the decrease in cutting force caused by the application of liquid nitrogen from the rake surface is approximately 20%. However, this ratio is less at the lowest cutting depth. Therefore, it is seen that the lubricating effect of liquid nitrogen plays an active role here as well and can reduce the force more when the friction force is greater.

<figure><img src="image_1980.png" /><figcaption>Fig. 3. Effect of cutting depth on cutting forces (V=120 m/min).</figcaption></figure>

Microhardness. The variation of hardness in the surface and subsurface areas of the machined parts is important depending on where the part will be used. While in some cases it is not desired to change the hardness in order to create a more homogeneous structure, in some cases it is a desirable feature to increase the hardness of the surface, especially in parts that are in contact with each other or subject to wear [23, 24]. 316L has a lower workability than other stainless steels and has a high tendency to hardening during machining [3]. Thermal effects during machining and martensite formation in the austenite structure during plastic deformation are considered as the reason for the tendency of hardening [25, 26].

Microhardness values of orthogonally machined samples under different cutting conditions is shown in Fig. 4. There is an increase in the hardness of the machined samples from surface and sub-surface to 100 µm depth in all of conditions. While the highest hardness value was measured as 282 HV under cryogenic condition applied from rake face, the lowest hardness value was measured as 231.5 HV under dry condition at the distance of 15 µm from the surface. There is a difference of 22% between the lowest and the highest hardness values. Also, there are increase of 44% and 18% in hardness of machined samples as compared to as-received hardness. LN₂ directly penetrates the tool-chip region from rake face. Therefore, since the primary and secondary deformation region temperature is directly affected, chip breakability becomes easier, but the material hardness increases. However, when LN₂ penetrates from the flank face, the hardness increase is less as the clearance zone is affected.

In Fig. 5, the variation of the hardness from the machined surface to the center of the workpiece can be expressed in three steps. The first stage is hardened very clearly but hardness values is lower than beginning of the second stage. The reason for this behavior is that the first 10 µm layer, which is the closest layer to the surface, is the most exposed to heat and high plastic deformation. This has naturally caused a relative decrease in hardness. At the starting point of the second stage, the hardness value reached its maximum point. While the material is exposed to a large plastic deformation, it is exposed to less heat. The depth of this layer determines the surface and the deformed layer under the surface in the result of the machining process. In the third stage, the as-received material hardness is active and the effect of the machining process is not exist.
<figure><img src="image_1986.png" /><figcaption>Fig. 4. Microhardness values at different cutting conditions in orthogonal cutting (V=120 m/min, to = 0.08 mm)</figcaption></figure>

<figure><img src="image_1987.png" /><figcaption>Fig. 5. Plastic deformation behavior of 316L stainless steel material (V= 40 m/dak, to = 0.08 mm)</figcaption></figure>

Microhardness values from surface to depth of 200 µm of orthogonally machined samples under different cutting conditions is shown in Fig. 6. There is an increase in all of cutting conditions. It is obvious that hardness values in cryogenic conditions is higher than dry condition. This is an indication that the thermal effect in the surface region will be relatively reduced by the effect of cryogenic cooling. There are increase of 70%, 48% and 37% in hardness of machined samples under LN₂ from rake face, LN₂ from flank face and dry cutting conditions, respectively, as compared to as-received hardness.
<figure><img src="image_1981.png" /><figcaption>Fig. 4. Microhardness values at different cutting conditions in orthogonal cutting (V=120 m/min, to = 0.08 mm)</figcaption></figure>

<figure><img src="image_1982.png" /><figcaption>Fig. 5. Plastic deformation behavior of 316L stainless steel material (V= 40 m/dak, to = 0.08 mm)</figcaption></figure>

Microhardness values from surface to depth of 200 µm of orthogonally machined samples under different cutting conditions is shown in Fig. 6. There is an increase in all of cutting conditions. It is obvious that hardness values in cryogenic conditions is higher than dry condition. This is an indication that the thermal effect in the surface region will be relatively reduced by the effect of cryogenic cooling. There are increase of 70%, 48% and 37% in hardness of machined samples under LN₂ from rake face, LN₂ from flank face and dry cutting conditions, respectively, as compared to as-received hardness.
<figure><img src="image_1988.png" /><figcaption>Fig. 6. Microhardness values at different cutting conditions in orthogonal cutting ($V=120$ m/min, $t_o = 0.16$ mm)</figcaption></figure>

Microstructure Analysis. It is important to examine the microstructure to understand the changes that occur on the workpiece surface and sub-surface after machining. In machining operations, significant changes can be observed in the surface and sub-surface microstructure in terms of cutting conditions and cutting parameters [23, 24, 27]. Since these changes may affect the functionality of the workpiece, the effects of changes occurring in the surface and subsurface areas as a result of machining the workpiece must be determined [15].

Microstructural images of orthogonally machined 316L stainless steel material at different cutting speeds and cutting conditions are shown in Fig. 7. The deformation at the lower cutting speeds at the surface and sub-surface is more obvious than the higher cutting speeds. It is seen that increasing the cutting speed tends to decrease the depth of grain deformation. This can be assessed as an increase in cutting temperature, depending on the increase in cutting speed, and a reflection on the machined surface and sub-surface. It is known that in the machining of different materials, deformation occurs depending on the cutting speed and thus phase transformation occurs in the material. As the cutting speed increases, the cutting temperature increases and the deformation decreases due to the annihilation of dislocations on the machined surface [28].

Also, when cryogenic conditions are compared to dry cutting, it appears that there are no significant differences in terms of microstructure. However, it should be emphasized that cryogenic conditions can be said that they have affected layers at a lower depth than dry cutting, especially considering cutting speeds of 120 and 200 m/min. In general, the temperature of the workpiece during dry cutting is higher than in cryogenic conditions. In this case, even under relatively low stresses, the formation of dislocations can be more convenient and the deformed layers of the workpieces can be larger.
<figure><img src="image_1983.png" /><figcaption>Fig. 6. Microhardness values at different cutting conditions in orthogonal cutting ($V=120$ m/min, $t_o = 0.16$ mm)</figcaption></figure>

Microstructure Analysis. It is important to examine the microstructure to understand the changes that occur on the workpiece surface and sub-surface after machining. In machining operations, significant changes can be observed in the surface and sub-surface microstructure in terms of cutting conditions and cutting parameters [23, 24, 27]. Since these changes may affect the functionality of the workpiece, the effects of changes occurring in the surface and subsurface areas as a result of machining the workpiece must be determined [15].

Microstructural images of orthogonally machined 316L stainless steel material at different cutting speeds and cutting conditions are shown in Fig. 7. The deformation at the lower cutting speeds at the surface and sub-surface is more obvious than the higher cutting speeds. It is seen that increasing the cutting speed tends to decrease the depth of grain deformation. This can be assessed as an increase in cutting temperature, depending on the increase in cutting speed, and a reflection on the machined surface and sub-surface. It is known that in the machining of different materials, deformation occurs depending on the cutting speed and thus phase transformation occurs in the material. As the cutting speed increases, the cutting temperature increases and the deformation decreases due to the annihilation of dislocations on the machined surface [28].

Also, when cryogenic conditions are compared to dry cutting, it appears that there are no significant differences in terms of microstructure. However, it should be emphasized that cryogenic conditions can be said that they have affected layers at a lower depth than dry cutting, especially considering cutting speeds of 120 and 200 m/min. In general, the temperature of the workpiece during dry cutting is higher than in cryogenic conditions. In this case, even under relatively low stresses, the formation of dislocations can be more convenient and the deformed layers of the workpieces can be larger.
<figure><img src="image_1989.png" /><figcaption>Fig. 7. Microstructure images of orthogonally turned samples at different cutting speeds and conditions under optical microscope ($t_o$ = 0.08 mm)</figcaption></figure>

X-Ray Diffraction Analysis. In the machining process, changes occur in the surface and subsurface regions of the workpiece depending on the amount of deformation in the cutting area due to effects such as temperature increase, friction and wear, and phase transformation may occur depending on the processing parameters and conditions [15, 23, 24]. Taking into account the austenitic stainless steels, plastic deformation at low temperatures causes austenite-martensite transformation in the structure [29]. This transformation directly affects functional performance criteria such as yield stress, hardening, fracture toughness and fatigue life [30]. The phase transformation caused by the machining process in this context is considered as one of the important surface integrity characteristics [15].

316L stainless steel is a γ-austenitic material with face-centered cubic crystal structure [26]. In 316L stainless steels, martensite phases can be seen or martensite formation can be seen under the effect of tension called SIM (stress induced martensite). This is also a condition that depends on the cutting conditions, which is an unusual situation in the machining process [20, 26]. There are studies stating that cryogenic conditions cause martensite phase formation, but this is only possible with maximum plastic deformation at very low cutting speeds [20, 31, 32].

Fig. 8 shows the comparisons of the main austenite XRD pattern with (111) texture of orthogonally machined samples under different cutting conditions at 120 m/min cutting speed. Looking at the figure, it is obvious that the intensity of the peaks in the cryogenic conditions increases when compared to as-received peak. The highest peak intensity occurred when cryogenic coolant applied from the rake face and then flank face. The increases in peak intensity and
<figure><img src="image_1984.png" /><figcaption>Fig. 7. Microstructure images of orthogonally turned samples at different cutting speeds and conditions under optical microscope ($t_o$ = 0.08 mm)</figcaption></figure>

X-Ray Diffraction Analysis. In the machining process, changes occur in the surface and subsurface regions of the workpiece depending on the amount of deformation in the cutting area due to effects such as temperature increase, friction and wear, and phase transformation may occur depending on the processing parameters and conditions [15, 23, 24]. Taking into account the austenitic stainless steels, plastic deformation at low temperatures causes austenite-martensite transformation in the structure [29]. This transformation directly affects functional performance criteria such as yield stress, hardening, fracture toughness and fatigue life [30]. The phase transformation caused by the machining process in this context is considered as one of the important surface integrity characteristics [15].

316L stainless steel is a γ-austenitic material with face-centered cubic crystal structure [26]. In 316L stainless steels, martensite phases can be seen or martensite formation can be seen under the effect of tension called SIM (stress induced martensite). This is also a condition that depends on the cutting conditions, which is an unusual situation in the machining process [20, 26]. There are studies stating that cryogenic conditions cause martensite phase formation, but this is only possible with maximum plastic deformation at very low cutting speeds [20, 31, 32].

Fig. 8 shows the comparisons of the main austenite XRD pattern with (111) texture of orthogonally machined samples under different cutting conditions at 120 m/min cutting speed. Looking at the figure, it is obvious that the intensity of the peaks in the cryogenic conditions increases when compared to as-received peak. The highest peak intensity occurred when cryogenic coolant applied from the rake face and then flank face. The increases in peak intensity and
broadening indicate hardening, which occurs on the machined surface of the workpiece [32]. Considering the hardness values given in the previous section in Fig. 6, it can be said that XRD analysis is supporting microhardness values. Another issue is that the as-received peak width is less than the machined peak widths. This has been attributed to the fact that the residual stress rate may be high in the small diameter and rolling produced receiver material. It is possible that the machining process affects the stress distribution and causes the peak to broad.

<figure><img src="image_1990.png" /><figcaption>Fig. 8. XRD patterns of orthogonally machined samples under different cutting conditions (V= 120 m/min, to = 0.16 mm)</figcaption></figure>

**Summary**

In this study, machining-induced surface integrity characteristics of austenitic 316L stainless steel were examined at different cutting parameters and under cryogenic cooling conditions. Cryogenic coolant was sent to the cutting zone from different nozzle positions and its effect was examined.

* Cryogenic machining caused a decrease in cutting forces. The least cutting forces were obtained when the cryogenic coolant was applied from the rake surface.

* Apparent plastic deformation and grain refinement occurred in the microstructure of the processed material, especially under low cutting speed and cryogenic conditions.

* It is remarkable that significant hardening was observed on the machined surface and sub-surface. The highest hardness increase was obtained when cryogenic coolant was applied to the rake face.

* Cryogenically machined XRD peaks broadened more compared to dry and as-received peaks. This shows that grain refining is greater.

**Acknowledgement**

Authors acknowledges the support from TUBITAK under Grant No 214M647.

**References**

[1] D. Peckner, I.M. Bernstein, Handbook of stainless steels, McGraw-Hill Book Co., New York. 1977, (Chapters paged separately), (1977).
broadening indicate hardening, which occurs on the machined surface of the workpiece [32]. Considering the hardness values given in the previous section in Fig. 6, it can be said that XRD analysis is supporting microhardness values. Another issue is that the as-received peak width is less than the machined peak widths. This has been attributed to the fact that the residual stress rate may be high in the small diameter and rolling produced receiver material. It is possible that the machining process affects the stress distribution and causes the peak to broad.

<figure><img src="image_1985.png" /><figcaption>Fig. 8. XRD patterns of orthogonally machined samples under different cutting conditions (V= 120 m/min, to = 0.16 mm)</figcaption></figure>

**Summary**

In this study, machining-induced surface integrity characteristics of austenitic 316L stainless steel were examined at different cutting parameters and under cryogenic cooling conditions. Cryogenic coolant was sent to the cutting zone from different nozzle positions and its effect was examined.

* Cryogenic machining caused a decrease in cutting forces. The least cutting forces were obtained when the cryogenic coolant was applied from the rake surface.

* Apparent plastic deformation and grain refinement occurred in the microstructure of the processed material, especially under low cutting speed and cryogenic conditions.

* It is remarkable that significant hardening was observed on the machined surface and sub-surface. The highest hardness increase was obtained when cryogenic coolant was applied to the rake face.

* Cryogenically machined XRD peaks broadened more compared to dry and as-received peaks. This shows that grain refining is greater.

**Acknowledgement**

Authors acknowledges the support from TUBITAK under Grant No 214M647.

**References**

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[3] M.N. Nasr, E.-G. Ng, M. Elbestawi, Modelling the effects of tool-edge radius on residual stresses when orthogonal cutting AISI 316L, International Journal of Machine Tools and Manufacture, 47 (2007) 401-411. https://doi.org/10.1016/j.ijmachtools.2006.03.004

[4] D.A. Stephenson, J.S. Agapiou, Metal cutting theory and practice, CRC press, 2016.

[5] S. Debnath, M.M. Reddy, Q.S. Yi, Environmental friendly cutting fluids and cooling techniques in machining: a review, Journal of Cleaner Production, 83 (2014) 33-47. https://doi.org/10.1016/j.jclepro.2014.07.071

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[9] I. Jawahir, H. Attia, D. Biermann, J. Duflou, F. Klocke, D. Meyer, S. Newman, F. Pusavec, M. Putz, J. Rech, Cryogenic manufacturing processes, CIRP Annals, 65 (2016) 713-736. https://doi.org/10.1016/j.cirp.2016.06.007

[10] G. Ortiz-de-Zarate, D. Soriano, A. Madariaga, A. Garay, I. Rodriguez, P. Arrazola, Experimental and FEM analysis of dry and cryogenic turning of hardened steel 100Cr6 using CBN Wiper tools, Procedia CIRP, 102 (2021) 7-12. https://doi.org/10.1016/j.procir.2021.09.002

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[12] O. Kitay, Y. Kaynak, The effect of flood, high-pressure cooling, and CO2-assisted cryogenic machining on microhardness, microstructure, and X-ray diffraction patterns of NiTi shape memory alloy, Journal of Materials Engineering and Performance, 30 (2021) 5799-5810. https://doi.org/10.1007/s11665-021-05854-6

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[5] S. Debnath, M.M. Reddy, Q.S. Yi, Environmental friendly cutting fluids and cooling techniques in machining: a review, Journal of Cleaner Production, 83 (2014) 33-47. https://doi.org/10.1016/j.jclepro.2014.07.071

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[9] I. Jawahir, H. Attia, D. Biermann, J. Duflou, F. Klocke, D. Meyer, S. Newman, F. Pusavec, M. Putz, J. Rech, Cryogenic manufacturing processes, CIRP Annals, 65 (2016) 713-736. https://doi.org/10.1016/j.cirp.2016.06.007

[10] G. Ortiz-de-Zarate, D. Soriano, A. Madariaga, A. Garay, I. Rodriguez, P. Arrazola, Experimental and FEM analysis of dry and cryogenic turning of hardened steel 100Cr6 using CBN Wiper tools, Procedia CIRP, 102 (2021) 7-12. https://doi.org/10.1016/j.procir.2021.09.002

[11] S.Y. Hong, Y. Ding, Cooling approaches and cutting temperatures in cryogenic machining of Ti-6Al-4V, International Journal of Machine Tools and Manufacture, 41 (2001) 1417-1437. https://doi.org/10.1016/S0890-6955(01)00026-8

[12] O. Kitay, Y. Kaynak, The effect of flood, high-pressure cooling, and CO2-assisted cryogenic machining on microhardness, microstructure, and X-ray diffraction patterns of NiTi shape memory alloy, Journal of Materials Engineering and Performance, 30 (2021) 5799-5810. https://doi.org/10.1007/s11665-021-05854-6

[13] S.Y. Hong, Z. Zhao, Thermal aspects, material considerations and cooling strategies in cryogenic machining, Clean Technologies and Environmental Policy, 1 (1999) 107-116. https://doi.org/10.1007/s100980050016

[14] G. Rotella, O. Dillon, D. Umbrello, L. Settineri, I. Jawahir, The effects of cooling conditions on surface integrity in machining of Ti6Al4V alloy, The International Journal of Advanced Manufacturing Technology, 71 (2014) 47-55. https://doi.org/10.1007/s00170-013-5477-9

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[16] D. Umbrello, Z. Pu, S. Caruso, J. Outeiro, A. Jayal, O. Dillon, I. Jawahir, The effects of cryogenic cooling on surface integrity in hard machining, Procedia Engineering, 19 (2011) 371-376. https://doi.org/10.1016/j.proeng.2011.11.127

[17] S. Yang, D. Umbrello, O.W. Dillon Jr, D.A. Puleo, I. Jawahir, Cryogenic cooling effect on surface and subsurface microstructural modifications in burnishing of Co-Cr-Mo biomaterial, Journal of Materials Processing Technology, 217 (2015) 211-221. https://doi.org/10.1016/j.jmatprotec.2014.11.004

[18] Y. Kaynak, H. Tobe, R. Noebe, H. Karaca, I. Jawahir, The effects of machining on the microstructure and transformation behavior of NiTi Alloy, Scripta Materialia, 74 (2014) 60-63. https://doi.org/10.1016/j.scriptamat.2013.10.023
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[26] Y. Idell, G. Facco, A. Kulovits, M. Shankar, J. Wiezorek, Strengthening of austenitic stainless steel by formation of nanocrystalline γ-phase through severe plastic deformation during two-dimensional linear plane-strain machining, Scripta materialia, 68 (2013) 667-670. https://doi.org/10.1016/j.scriptamat.2013.01.025

[27] R. M'Saoubi, J. Outeiro, H. Chandrasekaran, O. Dillon Jr, I. Jawahir, A review of surface integrity in machining and its impact on functional performance and life of machined products, International Journal of Sustainable Manufacturing, 1 (2008) 203-236. https://doi.org/10.1504/IJSM.2008.019234

[28] Y. Kaynak, H. Karaca, I. 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

[29] J. Patel, M. Cohen, Criterion for the action of applied stress in the martensitic transformation, Acta Metallurgica, 1 (1953) 531-538. https://doi.org/10.1016/0001-6160(53)90083-2

[30] A. Rosen, R. Jago, T. Kjer, Tensile properties of metastable stainless steels, Journal of Materials Science, 7 (1972) 870-876. https://doi.org/10.1007/BF00550434

[31] Y. Kaynak, B. Huang, H. Karaca, I. Jawahir, Surface Characteristics of Machined NiTi Shape Memory Alloy: The Effects of Cryogenic Cooling and Preheating Conditions, Journal of Materials Engineering and Performance, 26 (2017) 3597-3606. https://doi.org/10.1007/s11665-017-2791-7

[32] Y. Kaynak, Machining and phase transformation response of room-temperature austenitic NiTi shape memory alloy, Journal of materials engineering and performance, 23 (2014) 3354-3360. https://doi.org/10.1007/s11665-014-1058-9
[19] S. Yang, Cryogenic burnishing of Co-Cr-Mo biomedical alloy for enhanced surface integrity and improved wear performance, in, University of Kentucky, 2012.

[20] Y.K. Erkin Duman, Analysis of Surface Integrity in Cryogenic Machining of 316l Stainless Steel Material, in: UTIS, İstanbul, TURKEY, 2016, pp. 253-265.

[21] M. Habak, J.L. Lebrun, An experimental study of the effect of high-pressure water jet assisted turning (HPWJAT) on the surface integrity, International Journal of Machine Tools and Manufacture, 51 (2011) 661-669. https://doi.org/10.1016/j.ijmachtools.2011.05.001

[22] K.B. Small, D.A. Englehart, T.A. Christman, Etching specialty alloys, Advanced materials & processes, (2008) 33.

[23] D. Ulutan, T. Ozel, Machining induced surface integrity in titanium and nickel alloys: A review, International Journal of Machine Tools and Manufacture, 51 (2011) 250-280. https://doi.org/10.1016/j.ijmachtools.2010.11.003

[24] Y. Kaynak, T. Lu, I. Jawahir, Cryogenic machining-induced surface integrity: a review and comparison with dry, MQL, and flood-cooled machining, Machining Science and Technology, 18 (2014) 149-198.

[25] D. O'Sullivan, M. Cotterell, Machinability of austenitic stainless steel SS303, Journal of Materials Processing Technology, 124 (2002) 153-159. https://doi.org/10.1080/10910344.2014.897836 https://doi.org/10.1016/S0924-0136(02)00197-8

[26] Y. Idell, G. Facco, A. Kulovits, M. Shankar, J. Wiezorek, Strengthening of austenitic stainless steel by formation of nanocrystalline γ-phase through severe plastic deformation during two-dimensional linear plane-strain machining, Scripta materialia, 68 (2013) 667-670. https://doi.org/10.1016/j.scriptamat.2013.01.025

[27] R. M'Saoubi, J. Outeiro, H. Chandrasekaran, O. Dillon Jr, I. Jawahir, A review of surface integrity in machining and its impact on functional performance and life of machined products, International Journal of Sustainable Manufacturing, 1 (2008) 203-236. https://doi.org/10.1504/IJSM.2008.019234

[28] Y. Kaynak, H. Karaca, I. 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

[29] J. Patel, M. Cohen, Criterion for the action of applied stress in the martensitic transformation, Acta Metallurgica, 1 (1953) 531-538. https://doi.org/10.1016/0001-6160(53)90083-2

[30] A. Rosen, R. Jago, T. Kjer, Tensile properties of metastable stainless steels, Journal of Materials Science, 7 (1972) 870-876. https://doi.org/10.1007/BF00550434

[31] Y. Kaynak, B. Huang, H. Karaca, I. Jawahir, Surface Characteristics of Machined NiTi Shape Memory Alloy: The Effects of Cryogenic Cooling and Preheating Conditions, Journal of Materials Engineering and Performance, 26 (2017) 3597-3606. https://doi.org/10.1007/s11665-017-2791-7

[32] Y. Kaynak, Machining and phase transformation response of room-temperature austenitic NiTi shape memory alloy, Journal of materials engineering and performance, 23 (2014) 3354-3360. https://doi.org/10.1007/s11665-014-1058-9
# A numerical approach to investigate the influence of cutting fluid on tool wear in machining AISI 1045 steel

METHON Grégory<fnref n="1" />,ª, COURBON Cédric<fnref n="1" />,ᵇ*, ESPINOUX Frédéric<fnref n="2" />,ᶜ, FRELECHOUX Richard<fnref n="2" />,ᵈ and RECH Joël<fnref n="1" />,ᵉ

**Keywords:** Numerical Simulation, Cutting Fluid, Tool Wear, Friction, Convection

**Abstract.** The proposed work presents the development of a numerical model to take into account cutting fluids when turning of an AISI 1045 steel with a TiN coated tool. At this stage, the study considered the impact of a cutting fluid only based on its thermal properties and ability to reduce friction. Three heat exchange scenarios were explored: dry machining (h = 20 W.m⁻<fnref n="2" />.K⁻<fnref n="1" />), use of a straight oil lubrication with low pressure (h = 200 W.m⁻<fnref n="2" />.K⁻<fnref n="1" />) and cooling with an emulsion based lubrication (h = 1000 W.m⁻<fnref n="2" />.K⁻<fnref n="1" />). Regarding friction, three levels of static friction (µs = 0.3, 0.5 and 0.7) have been considered. Following an analysis of thermomechanical loads (stress, sliding speed, and temperature) exerted on the tool, a Usui wear law was employed for each case to simulate cutting tool wear. Results indicate that friction exerts a significant impact compared to the convection coefficient.

**Introduction**

The application of cutting fluids in machining operations constitutes a standard practice in contemporary industrial settings. Its purpose is to evacuate the chips but also to cool the cutting tools and thus limit the evolution of tool wear.

Numerical simulations are widely used in order to simulate machining operations. They offer the possibility to study the tool-chip contact interface and allow the prediction of various machining outputs such as cutting forces, temperature, surface integrity or tool wear. Modeling of turning operations has been considered by a large number of papers [1]–[3]. However, while machining operation simulations reach the point that they can now predict tool wear, very few models take into account the use of cutting fluids. Consequently, this study aims to develop and customize a numerical simulation procedure to identify the key parameters that quantify the influence of a cutting fluid on the wear of cutting tools during the turning of an AISI 1045 steel.

In order to perform this study, the authors used an iterative wear procedure developed by Rech in 2018 [4] and based on a 2D Arbitrary-Lagrangian-Eulerian (ALE) model. The present simulations then focused on the machining of an AISI 1045 steel with a TiN-coated tungsten carbide tool under cutting conditions representative of an industrial application.

In this preliminary study, the assumption that the cutting fluid could be simply modelled by friction and convection coefficients was considered. Three different levels for each coefficient were covered to assess the influence of the cutting fluid on tool wear. A phenomenological equation, developed by Usui, was implemented in the numerical model to study flank wear (VB) evolution over the cutting time.

**Numerical model & case study**

In order to assess only the impact of the cutting fluid as a preliminary study, a simple cutting configuration was selected. Orthogonal turning of an AISI 1045 steel was investigated at a cutting

<figure><img src="image_1991.png" /></figure>

<footnote n="1">¹Univ Lyon, ENISE, ECL, ENTPE, CNRS, LTDS UMR 5513, F-42023, Saint-Etienne, France</footnote>

<footnote n="2">²TotalEnergies One Tech, Centre de Recherche de Solaize, 69360, Solaize, France</footnote>

<footnote>ªgregorymethon@gmail.com, ᵇcedric.courbon@enise.fr, ᶜfrederic.espinoux@totalenergies.com, ᵈrichard.frelechoux@totalenergies.com, ᵉjoel.rech@enise.fr</footnote>
# A numerical approach to investigate the influence of cutting fluid on tool wear in machining AISI 1045 steel

METHON Grégory¹,ª, COURBON Cédric¹,ᵇ*, ESPINOUX Frédéric²,ᶜ, FRELECHOUX Richard²,ᵈ and RECH Joël¹,ᵉ

¹Univ Lyon, ENISE, ECL, ENTPE, CNRS, LTDS UMR 5513, F-42023, Saint-Etienne, France

²TotalEnergies One Tech, Centre de Recherche de Solaize, 69360, Solaize, France

ªgregorymethon@gmail.com, ᵇcedric.courbon@enise.fr, ᶜfrederic.espinoux@totalenergies.com, ᵈrichard.frelechoux@totalenergies.com, ᶮjoel.rech@enise.fr

**Keywords:** Numerical Simulation, Cutting Fluid, Tool Wear, Friction, Convection

**Abstract.** The proposed work presents the development of a numerical model to take into account cutting fluids when turning of an AISI 1045 steel with a TiN coated tool. At this stage, the study considered the impact of a cutting fluid only based on its thermal properties and ability to reduce friction. Three heat exchange scenarios were explored: dry machining (h = 20 W.m⁻².K⁻¹), use of a straight oil lubrication with low pressure (h = 200 W.m⁻².K⁻¹) and cooling with an emulsion based lubrication (h = 1000 W.m⁻².K⁻¹). Regarding friction, three levels of static friction (µs = 0.3, 0.5 and 0.7) have been considered. Following an analysis of thermomechanical loads (stress, sliding speed, and temperature) exerted on the tool, a Usui wear law was employed for each case to simulate cutting tool wear. Results indicate that friction exerts a significant impact compared to the convection coefficient.

**Introduction**

The application of cutting fluids in machining operations constitutes a standard practice in contemporary industrial settings. Its purpose is to evacuate the chips but also to cool the cutting tools and thus limit the evolution of tool wear.

Numerical simulations are widely used in order to simulate machining operations. They offer the possibility to study the tool-chip contact interface and allow the prediction of various machining outputs such as cutting forces, temperature, surface integrity or tool wear. Modeling of turning operations has been considered by a large number of papers [1]–[3]. However, while machining operation simulations reach the point that they can now predict tool wear, very few models take into account the use of cutting fluids. Consequently, this study aims to develop and customize a numerical simulation procedure to identify the key parameters that quantify the influence of a cutting fluid on the wear of cutting tools during the turning of an AISI 1045 steel.

In order to perform this study, the authors used an iterative wear procedure developed by Rech in 2018 [4] and based on a 2D Arbitrary-Lagrangian-Eulerian (ALE) model. The present simulations then focused on the machining of an AISI 1045 steel with a TiN-coated tungsten carbide tool under cutting conditions representative of an industrial application.

In this preliminary study, the assumption that the cutting fluid could be simply modelled by friction and convection coefficients was considered. Three different levels for each coefficient were covered to assess the influence of the cutting fluid on tool wear. A phenomenological equation, developed by Usui, was implemented in the numerical model to study flank wear (VB) evolution over the cutting time.

**Numerical model & case study**

In order to assess only the impact of the cutting fluid as a preliminary study, a simple cutting configuration was selected. Orthogonal turning of an AISI 1045 steel was investigated at a cutting

<figure><img src="image_1986.png" /></figure>
speed of 200m/min, feed rate of 0.2mm/rev, using a TiN coated carbide tool with a rake angle of 7°, clearance angle of 3° and cutting edge radius of 50µm. This configuration was based on the work of Giovenco [5] and Courbon et al. [6] performed at LTDS during the last decade. The basis of the wear simulation is a 2D orthogonal cutting model ALE developed in ABAQUS/Explicit©. It was found that such models are well known in the literature. They have been successfully used to predict: cutting forces, tool temperature, stresses in the workpiece and chip formation. The present model consists of a deformable workpiece and a fully fixed rigid cutting tool (Figure 1). The tool is rigid while a material flow is applied through the part. This material flow is defined by three Eulerian free boundaries located: on the left at the material inlet, at the top of the chip and on the right at the outlet of the modeled part. The heat exchange is allowed by a contact thermal conductance between the tool and the part. The value of the conductance has been set to 10⁴ W/m².K based on the work of Courbon et al. [6]. The workpiece is meshed with 4-node coupled temperature displacement elements (CPE4RT). The model is composed of 9000 elements mainly located at the contact interface which is the region of interest in our case.

<figure><img src="image_1992.png" /><figcaption>Figure 1 - 2D ALE model used in the study</figcaption></figure>

As for the flow stress model, a well-established Johnson-Cook strain rate dependent equation was used with the parameters identified by Jaspers & Dautzenberg (2002) [7]

Friction was implemented through a varying Coulomb model with an exponential decay (Eq.1) as plotted in Figure 2:

$$ \mu(V_{sl}) = \mu_k + (\mu_s - \mu_k) \cdot e^{-d_c \cdot V_{sl}} \quad (1) $$

Where μ is the coefficient of friction, Vsl is the sliding velocity, μk is the kinetic friction coefficient, μs is the static friction coefficient, dc is defined as the decay coefficient.
speed of 200m/min, feed rate of 0.2mm/rev, using a TiN coated carbide tool with a rake angle of 7°, clearance angle of 3° and cutting edge radius of 50µm. This configuration was based on the work of Giovenco [5] and Courbon et al. [6] performed at LTDS during the last decade. The basis of the wear simulation is a 2D orthogonal cutting model ALE developed in ABAQUS/Explicit©. It was found that such models are well known in the literature. They have been successfully used to predict: cutting forces, tool temperature, stresses in the workpiece and chip formation. The present model consists of a deformable workpiece and a fully fixed rigid cutting tool (Figure 1). The tool is rigid while a material flow is applied through the part. This material flow is defined by three Eulerian free boundaries located: on the left at the material inlet, at the top of the chip and on the right at the outlet of the modeled part. The heat exchange is allowed by a contact thermal conductance between the tool and the part. The value of the conductance has been set to 10⁴ W/m².K based on the work of Courbon et al. [6]. The workpiece is meshed with 4-node coupled temperature displacement elements (CPE4RT). The model is composed of 9000 elements mainly located at the contact interface which is the region of interest in our case.

<figure><img src="image_1987.png" /><figcaption>Figure 1 - 2D ALE model used in the study</figcaption></figure>

As for the flow stress model, a well-established Johnson-Cook strain rate dependent equation was used with the parameters identified by Jaspers & Dautzenberg (2002) [7]

Friction was implemented through a varying Coulomb model with an exponential decay (Eq.1) as plotted in Figure 2:

$$ \mu(V_{sl}) = \mu_k + (\mu_s - \mu_k) \cdot e^{-d_c \cdot V_{sl}} \quad (1) $$

Where μ is the coefficient of friction, Vsl is the sliding velocity, μk is the kinetic friction coefficient, μs is the static friction coefficient, dc is defined as the decay coefficient.
<figure><img src="image_1993.png" /><figcaption>Figure 2 - Modeling of friction as a function of sliding speed used in ABAQUS</figcaption></figure>

As this first thermo-mechanically coupled model only simulates a few milliseconds of the real cutting process, a second pure thermal simulation is employed to compute the steady-state heat transfer in the tool with Abaqus/Implicit®. Nodal heat fluxes from the ALE simulation are used as boundary conditions and applied to the corresponding node as heat inputs into this thermal model. After these two steps, the local thermomechanical loadings can be extracted and plotted along the cutting tool geometry.

Implementation of a cutting fluid in the numerical models

To consider the influence of the cutting fluid on these thermomechanical loads, modifications were introduced in both the 2D Arbitrary Lagrangian-Eulerian (ALE) model and the 2D implicit simulation. Dedicated regions were defined in both models, corresponding to the surfaces interacting with the cutting fluid. From a thermal perspective, the first assumption was that this interaction could be adequately represented by a simple heat exchange coefficient.

Figure 3 illustrates the surfaces subject to the convection coefficient, identified in green.

<figure><img src="image_1994.png" /><figcaption>Figure 3 - Heat exchange related to the cutting fluid applied in a) 2D ALE model and b) 2D implicit calculation</figcaption></figure>

These surfaces primarily include the rake face of the tool (the portion not in contact with the chip), the tool clearance face, the surface of the work material ahead of the cutting zone, and the freshly machined surface. A second assumption was that the thermal impact of the cutting fluid does not affect the tool-chip contact area (depicted in red), characterized by exceptionally high contact pressure.

The bibliography shows very different and variable results concerning the calorific value of cutting fluids according to the works of Luchesi (2012) [8] or Kurgin (2012) [9] (Figure 4).
<figure><img src="image_1988.png" /><figcaption>Figure 2 - Modeling of friction as a function of sliding speed used in ABAQUS</figcaption></figure>

As this first thermo-mechanically coupled model only simulates a few milliseconds of the real cutting process, a second pure thermal simulation is employed to compute the steady-state heat transfer in the tool with Abaqus/Implicit®. Nodal heat fluxes from the ALE simulation are used as boundary conditions and applied to the corresponding node as heat inputs into this thermal model. After these two steps, the local thermomechanical loadings can be extracted and plotted along the cutting tool geometry.

Implementation of a cutting fluid in the numerical models

To consider the influence of the cutting fluid on these thermomechanical loads, modifications were introduced in both the 2D Arbitrary Lagrangian-Eulerian (ALE) model and the 2D implicit simulation. Dedicated regions were defined in both models, corresponding to the surfaces interacting with the cutting fluid. From a thermal perspective, the first assumption was that this interaction could be adequately represented by a simple heat exchange coefficient.

Figure 3 illustrates the surfaces subject to the convection coefficient, identified in green.

<figure><img src="image_1989.png" /><figcaption>Figure 3 - Heat exchange related to the cutting fluid applied in a) 2D ALE model and b) 2D implicit calculation</figcaption></figure>

These surfaces primarily include the rake face of the tool (the portion not in contact with the chip), the tool clearance face, the surface of the work material ahead of the cutting zone, and the freshly machined surface. A second assumption was that the thermal impact of the cutting fluid does not affect the tool-chip contact area (depicted in red), characterized by exceptionally high contact pressure.

The bibliography shows very different and variable results concerning the calorific value of cutting fluids according to the works of Luchesi (2012) [8] or Kurgin (2012) [9] (Figure 4).
<table><thead><tr><th>Conditions</th><th>Fluid</th><th>h (W·m<sup>-2</sup>·K<sup>-1</sup>)</th></tr></thead><tbody><tr><td rowspan="2">Free convection</td><td>Gases</td><td>5–30</td></tr><tr><td>Water</td><td>100–900</td></tr><tr><td rowspan="3">Forced convection</td><td>Gases</td><td>10–300</td></tr><tr><td>Water</td><td>300–11.500</td></tr><tr><td>Viscous oils</td><td>60–1300</td></tr><tr><td rowspan="3">Phase change</td><td>Liquid metals</td><td>5.700–114.000</td></tr><tr><td>Boiling liquids</td><td>3000–57.000</td></tr><tr><td>Condensing vapors</td><td>5.700–114.000</td></tr></tbody></table>

<table><caption>Figure 4 - Heat exchange coefficient established by a) Luchesi (2012) [8] and b) Kurgin (2012) [9]</caption><thead><tr><th>Coolant type</th><th>h (W/m<sup>2</sup> C)</th><th>Biot ratio</th></tr></thead><tbody><tr><td>Free convection (air)</td><td>22</td><td>0.0004</td></tr><tr><td>Air flow @ 4 bar</td><td>248</td><td>0.005</td></tr><tr><td>Air flow @ 5 bar</td><td>270</td><td>0.006</td></tr><tr><td>Air flow @ 6 bar</td><td>294</td><td>0.006</td></tr><tr><td>MQL (40 mL/hr, 4 bar)</td><td>204</td><td>0.004</td></tr><tr><td>MQL (40 mL/hr, 6 bar)</td><td>300</td><td>0.007</td></tr><tr><td>MQL (110 mL/hr, 4 bar)</td><td>223</td><td>0.005</td></tr><tr><td>MQL (110 mL/hr, 6 bar)</td><td>225</td><td>0.005</td></tr><tr><td>Flood coolant with water-based fluid</td><td>5,230.0</td><td>0.298</td></tr></tbody></table>

Given the wide variations that could be reported, the following values were selected for this first sensitivity study in order to extract some first trends: Dry machining, $h = 20 W/m^2K$ - Straight oil lubrication, $h = 200 W/m^2K$ - Emulsion based lubrication, $h = 1000 W/m^2K$.

Finally, the potential lubricating effect of the fluid can be reasonably implemented by changing accordingly the friction model at the tool-material interface. Considering the velocity dependent friction model presented in Eq. 1, Rech et al. 2013 [10] showed that under dry conditions, a static friction, $\mu s$ which corresponds to low speed friction, of 0.5 and a dynamic friction, $\mu k$ at high speed, of 0.2 could be commonly reached. Assuming that the lubricating effect would mainly change the frictional behavior at low sliding velocity, the dynamic friction was kept constant at 0.2, as well as the decay coefficient at 1, whereas the static friction was changed from 0.3 to 0.5 and 0.7.

The graphical representation of the 3 resulting friction models is presented in Figure 5.

<figure><img src="image_1995.png" /><figcaption>Figure 5 - Graphical representation of the 3 cases studied for static friction $\mu s$ of 0.3, 0.5 and 0.7</figcaption></figure>

## Impact of a cutting fluid on the local thermomechanical loadings

The impact of the cooling and lubricating properties of a cutting fluid on the loadings applied along the tool-material interface was first investigated separately.

### Influence of convection on the local thermomechanical loadings

Figure 6 shows the impact of the 3 investigated lubrication conditions modelled via 3 different convection values on the thermomechanical loads, i.e. cooling effect of the fluid.
<table><thead><tr><th>Conditions</th><th>Fluid</th><th>h (W·m<sup>-2</sup>·K<sup>-1</sup>)</th></tr></thead><tbody><tr><td rowspan="2">Free convection</td><td>Gases</td><td>5–30</td></tr><tr><td>Water</td><td>100–900</td></tr><tr><td rowspan="3">Forced convection</td><td>Gases</td><td>10–300</td></tr><tr><td>Water</td><td>300–11.500</td></tr><tr><td>Viscous oils</td><td>60–1300</td></tr><tr><td rowspan="3">Phase change</td><td>Liquid metals</td><td>5.700–114.000</td></tr><tr><td>Boiling liquids</td><td>3000–57.000</td></tr><tr><td>Condensing vapors</td><td>5.700–114.000</td></tr></tbody></table>

<table><caption>Figure 4 - Heat exchange coefficient established by a) Luchesi (2012) [8] and b) Kurgin (2012) [9]</caption><thead><tr><th>Coolant type</th><th>h (W/m<sup>2</sup> C)</th><th>Biot ratio</th></tr></thead><tbody><tr><td>Free convection (air)</td><td>22</td><td>0.0004</td></tr><tr><td>Air flow @ 4 bar</td><td>248</td><td>0.005</td></tr><tr><td>Air flow @ 5 bar</td><td>270</td><td>0.006</td></tr><tr><td>Air flow @ 6 bar</td><td>294</td><td>0.006</td></tr><tr><td>MQL (40 mL/hr, 4 bar)</td><td>204</td><td>0.004</td></tr><tr><td>MQL (40 mL/hr, 6 bar)</td><td>300</td><td>0.007</td></tr><tr><td>MQL (110 mL/hr, 4 bar)</td><td>223</td><td>0.005</td></tr><tr><td>MQL (110 mL/hr, 6 bar)</td><td>225</td><td>0.005</td></tr><tr><td>Flood coolant with water-based fluid</td><td>5,230.0</td><td>0.298</td></tr></tbody></table>

Given the wide variations that could be reported, the following values were selected for this first sensitivity study in order to extract some first trends: Dry machining, h = 20 W/m²K - Straight oil lubrication, h = 200 W/m²K - Emulsion based lubrication, h = 1000 W/m²K.

Finally, the potential lubricating effect of the fluid can be reasonably implemented by changing accordingly the friction model at the tool-material interface. Considering the velocity dependent friction model presented in Eq. 1, Rech et al. 2013 [10] showed that under dry conditions, a static friction, µs which corresponds to low speed friction, of 0.5 and a dynamic friction, µk at high speed, of 0.2 could be commonly reached. Assuming that the lubricating effect would mainly change the frictional behavior at low sliding velocity, the dynamic friction was kept constant at 0.2, as well as the decay coefficient at 1, whereas the static friction was changed from 0.3 to 0.5 and 0.7.

The graphical representation of the 3 resulting friction models is presented in Figure 5.

<figure><img src="image_1990.png" /><figcaption>Figure 5 - Graphical representation of the 3 cases studied for static friction µs of 0.3, 0.5 and 0.7</figcaption></figure>

## Impact of a cutting fluid on the local thermomechanical loadings

The impact of the cooling and lubricating properties of a cutting fluid on the loadings applied along the tool-material interface was first investigated separately.

### Influence of convection on the local thermomechanical loadings

Figure 6 shows the impact of the 3 investigated lubrication conditions modelled via 3 different convection values on the thermomechanical loads, i.e. cooling effect of the fluid.
Interestingly, it can be seen that convection does not influence the distribution or magnitude of mechanical loads (pressure and sliding velocity) applied to the tool. Nevertheless, an increased convection seems to affect the steady-state temperature simulated in the implicit model.

<figure><img src="image_1996.png" /><figcaption>Figure 6 - Heat exchange (h) impact on thermomechanical loadings for 3 levels a) $h = 20 \text{ W.m}^{-2}.\text{K}^{-1}$, b) $h = 200 \text{ W.m}^{-2}.\text{K}^{-1}$ and c) $h = 1000 \text{ W.m}^{-2}.\text{K}^{-1}$ with $\mu s = 0.5$, $\mu k = 0.2$ and $dc = 1$</figcaption></figure>

## Influence of friction on the local thermomechanical loadings

On the other hand, Figure 7 presents the different thermomechanical loading profiles simulated for the 3 different static friction coefficient, i.e. lubricating effect of the fluid. It could be seen that the contact pressure and maximum sliding speed are almost identical while surprisingly, the maximum temperature is considerably limited by an increase of $\mu s$.

When focusing on the loadings simulated with $\mu s = 0.3$, the sliding velocity in black is found to be high all along the tool-material contact zone, except for a small section around the material separation point at the edge radius. This material flow, combined to local friction, will therefore generate a significant amount of heat, resulting in an induced temperature of 340°C.

When increasing $\mu s$ to 0.7, a stagnation zone appears on the rake face of the tool. This sticking zone is attributed to the higher friction coefficient compared to the previous case, hindering the chip from sliding over the rake face. This limited relative motion leads to minimal heat generation, consequently restricting the temperature increase.
Interestingly, it can be seen that convection does not influence the distribution or magnitude of mechanical loads (pressure and sliding velocity) applied to the tool. Nevertheless, an increased convection seems to affect the steady-state temperature simulated in the implicit model.

<figure><img src="image_1991.png" /><figcaption>Figure 6 - Heat exchange (h) impact on thermomechanical loadings for 3 levels a) $h = 20 \text{ W.m}^{-2}.\text{K}^{-1}$, b) $h = 200 \text{ W.m}^{-2}.\text{K}^{-1}$ and c) $h = 1000 \text{ W.m}^{-2}.\text{K}^{-1}$ with $\mu s = 0.5$, $\mu k = 0.2$ and $dc = 1$</figcaption></figure>

## Influence of friction on the local thermomechanical loadings

On the other hand, Figure 7 presents the different thermomechanical loading profiles simulated for the 3 different static friction coefficient, i.e. lubricating effect of the fluid. It could be seen that the contact pressure and maximum sliding speed are almost identical while surprisingly, the maximum temperature is considerably limited by an increase of $\mu s$.

When focusing on the loadings simulated with $\mu s = 0.3$, the sliding velocity in black is found to be high all along the tool-material contact zone, except for a small section around the material separation point at the edge radius. This material flow, combined to local friction, will therefore generate a significant amount of heat, resulting in an induced temperature of 340°C.

When increasing $\mu s$ to 0.7, a stagnation zone appears on the rake face of the tool. This sticking zone is attributed to the higher friction coefficient compared to the previous case, hindering the chip from sliding over the rake face. This limited relative motion leads to minimal heat generation, consequently restricting the temperature increase.
<figure><img src="image_1997.png" /><figcaption>Figure 7 - Static friction coefficient (μs) impact on thermomechanical loadings for three levels a) μs = 0.3, b= μs = 0.5 and c) μs = 0.7, with h = 200 W.m⁻².K⁻¹</figcaption></figure>

## Impact of a cutting fluid on the wear of a cutting tool

### Modeling of cutting tool wear

A multi-step tool wear modeling procedure was developed by Rech et al. (Rech, et al. 2018 [4]) with the principle illustrated in Figure 8. Based on the two already presented models, the thermomechanical loads at the tool-material interface are estimated and the tool geometry is updated by locally moving the nodes according to a wear equation. The numerical evolution of the VB flank wear can be simulated and compared for the various input data configuration representing the 3 lubrication methods.

Usui et al. [11][12] proposed a wear rate model to simulate crater and flank wear taking into account temperature (T). This model was based on the Archard adhesive wear model in which 2 constants A and B were added (Eq 2). Kitagawa et al. [13] [14] validated the model for carbide tools in carbon steel turning. Klocke and Frank [15] evaluated the model developed by Usui for the cutting and flank wear zones for CBN tools in hard turning.

$$ \frac{\partial W}{\partial t} = A \cdot \sigma_n \cdot V_{sl} \cdot e^{-\frac{B}{T}} \qquad (2) $$

Where ∂W/∂t is the wear rate, σn the contact stress, Vsl the local sliding speed, T the temperature. This model was thus implemented in the present study based on the parameters identified by Yen et al. [16] and Senthilkumar, et al. [17] with A = 0.0000000078 and B = 5302.
<figure><img src="image_1992.png" /><figcaption>Figure 7 - Static friction coefficient (μs) impact on thermomechanical loadings for three levels a) μs = 0.3, b= μs = 0.5 and c) μs = 0.7, with h = 200 W.m⁻².K⁻¹</figcaption></figure>

## Impact of a cutting fluid on the wear of a cutting tool

### Modeling of cutting tool wear

A multi-step tool wear modeling procedure was developed by Rech et al. (Rech, et al. 2018 [4]) with the principle illustrated in Figure 8. Based on the two already presented models, the thermomechanical loads at the tool-material interface are estimated and the tool geometry is updated by locally moving the nodes according to a wear equation. The numerical evolution of the VB flank wear can be simulated and compared for the various input data configuration representing the 3 lubrication methods.

Usui et al. [11][12] proposed a wear rate model to simulate crater and flank wear taking into account temperature (T). This model was based on the Archard adhesive wear model in which 2 constants A and B were added (Eq 2). Kitagawa et al. [13] [14] validated the model for carbide tools in carbon steel turning. Klocke and Frank [15] evaluated the model developed by Usui for the cutting and flank wear zones for CBN tools in hard turning.

$$ \frac{\partial W}{\partial t} = A \cdot \sigma_n \cdot V_{sl} \cdot e^{-\frac{B}{T}} \qquad (2) $$

Where ∂W/∂t is the wear rate, σn the contact stress, Vsl the local sliding speed, T the temperature. This model was thus implemented in the present study based on the parameters identified by Yen et al. [16] and Senthilkumar, et al. [17] with A = 0.0000000078 and B = 5302.
<figure><img src="image_1998.png" /><figcaption>Figure 8 - Iterative procedure to model tool wear</figcaption></figure>

## Influence of convection on tool wear

The numerically predicted evolution of the flank wear ($\partial VB/\partial t$) is presented in Figure 9. The three cases are displayed with convective exchange at 20 W/m$^2$K (yellow curve), 200 W/m$^2$K (orange curve) and 1000 W/m$^2$K (blue curve).

<figure><img src="image_1999.png" /><figcaption>Figure 9 - Influence of convection on cutting tool-life</figcaption></figure>

The wear criterion (VB max) has been set at 0.2 mm (black line), which corresponds to the tool life of the studied cutting tools. It can be seen that despite the changes in convection, the tool life remains the same at approximately 13 min. The convection, applied with the different assumptions, does not impact the tool life despite the fact that the wear equation was temperature dependent.
<figure><img src="image_1993.png" /><figcaption>Figure 8 - Iterative procedure to model tool wear</figcaption></figure>

## Influence of convection on tool wear

The numerically predicted evolution of the flank wear ($\partial VB/\partial t$) is presented in Figure 9. The three cases are displayed with convective exchange at 20 W/m$^2$K (yellow curve), 200 W/m$^2$K (orange curve) and 1000 W/m$^2$K (blue curve).

<figure><img src="image_1994.png" /><figcaption>Figure 9 - Influence of convection on cutting tool-life</figcaption></figure>

The wear criterion (VB max) has been set at 0.2 mm (black line), which corresponds to the tool life of the studied cutting tools. It can be seen that despite the changes in convection, the tool life remains the same at approximately 13 min. The convection, applied with the different assumptions, does not impact the tool life despite the fact that the wear equation was temperature dependent.
## Influence of friction on tool wear

The impact of friction on wear was then considered. In the previous section, it was shown that friction influenced both the mechanical and thermal loadings. The results are presented in Figure 10, displaying the 3 studied static friction µs namely 0.3, 0.5 and 0.7.

In this case, the impact of the friction formulation on the cutting tool life is significant. In particular, it can be seen that it increases from 12 to 28 min when µs increases from 0.3 to 0.7.

<figure><img src="image_2000.png" /><figcaption>Figure 10 - Influence of static friction µs on cutting tool-life</figcaption></figure>

## Critical discussion of the results

The obtained results reveal an intriguing inconsistency, diverging from the commonly reported trend where the use of a cutting fluid is expected to enhance tool life by reducing wear. This deviation is clearly attributed to the strong assumptions made and the oversimplified representation of cutting fluid effects within the numerical simulation. The simplistic implementation of the cutting fluid's impact through convection and friction coefficients led to an inaccurate representation of the actual machining conditions.

When examining the influence of convection, an unexpected impact on the steady-state temperature was observed, indicating a potential limitation in the representation of heat dissipation mechanisms in the current model. This suggests that the selected values of the convection coefficient were underestimated, and higher orders of magnitude, related to forced convection or phase changes of a water-based lubricant, should be considered.

The study of friction presented also intriguing findings. Surprisingly, an increase in µs led to a considerable limitation of the maximum temperature, suggesting a counterintuitive relationship between friction and temperature rise. A significant decrease in tool life was consequently observed with lower static friction coefficients. This noteworthy effect emphasizes the need for a more detailed modelling approach of how friction, lubrication, and heat dissipation interact in machining processes. Indeed, the effective local lubrication regime occurring at the tool-chip interface, has to be deeper investigated as it governs the actual lubrication effect and heat transfer within the contact.

Moreover, the Usui wear equation was used as proposed by the authors, meaning identified under dry machining conditions. This was intentionally done but emphasizes the need for proper identification corresponding to the actual lubricated scenario.

The observed discrepancies highlight the necessity for refining the model to better capture the complex relationships governing tool wear under the influence of cutting fluids. Further
*Influence of friction on tool wear*

The impact of friction on wear was then considered. In the previous section, it was shown that friction influenced both the mechanical and thermal loadings. The results are presented in Figure 10, displaying the 3 studied static friction µs namely 0.3, 0.5 and 0.7.

In this case, the impact of the friction formulation on the cutting tool life is significant. In particular, it can be seen that it increases from 12 to 28 min when µs increases from 0.3 to 0.7.

<figure><img src="image_1995.png" /><figcaption>Figure 10 - Influence of static friction µs on cutting tool-life</figcaption></figure>

**Critical discussion of the results**

The obtained results reveal an intriguing inconsistency, diverging from the commonly reported trend where the use of a cutting fluid is expected to enhance tool life by reducing wear. This deviation is clearly attributed to the strong assumptions made and the oversimplified representation of cutting fluid effects within the numerical simulation. The simplistic implementation of the cutting fluid's impact through convection and friction coefficients led to an inaccurate representation of the actual machining conditions.

When examining the influence of convection, an unexpected impact on the steady-state temperature was observed, indicating a potential limitation in the representation of heat dissipation mechanisms in the current model. This suggests that the selected values of the convection coefficient were underestimated, and higher orders of magnitude, related to forced convection or phase changes of a water-based lubricant, should be considered.

The study of friction presented also intriguing findings. Surprisingly, an increase in µs led to a considerable limitation of the maximum temperature, suggesting a counterintuitive relationship between friction and temperature rise. A significant decrease in tool life was consequently observed with lower static friction coefficients. This noteworthy effect emphasizes the need for a more detailed modelling approach of how friction, lubrication, and heat dissipation interact in machining processes. Indeed, the effective local lubrication regime occurring at the tool-chip interface, has to be deeper investigated as it governs the actual lubrication effect and heat transfer within the contact.

Moreover, the Usui wear equation was used as proposed by the authors, meaning identified under dry machining conditions. This was intentionally done but emphasizes the need for proper identification corresponding to the actual lubricated scenario.

The observed discrepancies highlight the necessity for refining the model to better capture the complex relationships governing tool wear under the influence of cutting fluids. Further
investigations should aim to develop more realistic representations of lubricant effects, to achieve
results consistent with industrial observations and expectations.

## Conclusion

The present study proposed to assess the impact of a cutting fluid on the thermomechanical loadings withstood by the cutting tool and the resulting wear. A numerical approach was developed to simply take into account both thermal and tribological aspects of the fluid. As a first step, two strong assumptions were made by modelling the fluid only with a local heat exchange coefficient and a tailored friction model. Simulations were then run to extract key numerical outputs such as local contact pressure, sliding velocities and ultimately the evolution of the predicted tool flank wear.

The simulations revealed that the fluid's thermal properties primarily affected tool temperature, with limited impact on wear in continuous cutting operations like turning. From the friction perspective, the iterative wear procedure demonstrated a high sensitivity to the different friction formulations, resulting in a more than twofold increase in tool life compared to the reference configuration when friction increases. Indeed, changing the friction formulation by reducing the static friction coefficient had a substantial effect on sliding speed profiles and, consequently, tool temperature due to changes in frictional dissipated energy.

These unexpected results underline the limitations of the model, which represents the cutting fluid solely through a heat exchange coefficient and a modified static friction coefficient, applied on a predefined region of the tool material contact. This oversimplified representation do not fully capture the various effects of cutting fluids suggesting that their influence extends beyond the only cooling and friction reduction. More physical phenomena such as phase changes, penetration of the fluid within the contact or even tribochemical action of additives definitely occur and must be taken into account.

Finally, the coefficients of the Usui model were kept constant and identified under dry conditions. Future work should include a comprehensive investigation of the sensitivity to the wear equation through numerical studies and experimental tribological tests to identify accurate wear model constants.

## References

[1] S. M. Athavale and J. S. Strenkowski, "Finite element modeling of machining: From proof-of-concept to engineering applications," *Mach. Sci. Technol.*, vol. 2, no. 2, pp. 317-342, 1998. https://doi.org/10.1080/10940349808945674

[2] F. Klocke, T. Beck, S. Hoppe, T. Krieg, N. Müller, T. Nöthe, H-W. Raedt, K. Sweeney, "Examples of FEM application in manufacturing technology," *J. Mater. Process. Technol.*, vol. 120, no. 1-3, pp. 450-457, 2002. https://doi.org/10.1016/S0924-0136(01)01210-9

[3] W. Grzesik, M. Bartoszuk, and P. Nieslony, "Finite element modelling of temperature distribution in the cutting zone in turning processes with differently coated tools," *J. Mater. Process. Technol.*, vol. 164-165, pp. 1204-1211, 2005. https://doi.org/10.1016/j.jmatprotec.2005.02.136

[4] J. Rech, A. Giovenco, C. Courbon, and F. Cabanettes, "Toward a new tribological approach to predict cutting tool wear," *CIRP Ann.*, vol. 67, no. 1, pp. 65-68, 2018. https://doi.org/10.1016/j.cirp.2018.03.014

[5] A. Giovenco, "Contribution à la simulation numérique de l'usure des outils de coupe", *PhD Thesis Ecole Centrale de Lyon*, 2018.

[6] C. Courbon, T. Mabrouki, J. Rech, D. Mazuyer, E. D'Eramo, "On the existence of a thermal contact resistance at the tool-chip interface in dry cutting of AISI 1045: Formation mechanisms
investigations should aim to develop more realistic representations of lubricant effects, to achieve
results consistent with industrial observations and expectations.

Conclusion

The present study proposed to assess the impact of a cutting fluid on the thermomechanical loadings withstood by the cutting tool and the resulting wear. A numerical approach was developed to simply take into account both thermal and tribological aspects of the fluid. As a first step, two strong assumptions were made by modelling the fluid only with a local heat exchange coefficient and a tailored friction model. Simulations were then run to extract key numerical outputs such as local contact pressure, sliding velocities and ultimately the evolution of the predicted tool flank wear.

The simulations revealed that the fluid's thermal properties primarily affected tool temperature, with limited impact on wear in continuous cutting operations like turning. From the friction perspective, the iterative wear procedure demonstrated a high sensitivity to the different friction formulations, resulting in a more than twofold increase in tool life compared to the reference configuration when friction increases. Indeed, changing the friction formulation by reducing the static friction coefficient had a substantial effect on sliding speed profiles and, consequently, tool temperature due to changes in frictional dissipated energy.

These unexpected results underline the limitations of the model, which represents the cutting fluid solely through a heat exchange coefficient and a modified static friction coefficient, applied on a predefined region of the tool material contact. This oversimplified representation do not fully capture the various effects of cutting fluids suggesting that their influence extends beyond the only cooling and friction reduction. More physical phenomena such as phase changes, penetration of the fluid within the contact or even tribochemical action of additives definitely occur and must be taken into account.

Finally, the coefficients of the Usui model were kept constant and identified under dry conditions. Future work should include a comprehensive investigation of the sensitivity to the wear equation through numerical studies and experimental tribological tests to identify accurate wear model constants.

References

[1] S. M. Athavale and J. S. Strenkowski, "Finite element modeling of machining: From proof-of-concept to engineering applications," *Mach. Sci. Technol.*, vol. 2, no. 2, pp. 317-342, 1998. https://doi.org/10.1080/10940349808945674

[2] F. Klocke, T. Beck, S. Hoppe, T. Krieg, N. Müller, T. Nöthe, H-W. Raedt, K. Sweeney, "Examples of FEM application in manufacturing technology," *J. Mater. Process. Technol.*, vol. 120, no. 1-3, pp. 450-457, 2002. https://doi.org/10.1016/S0924-0136(01)01210-9

[3] W. Grzesik, M. Bartoszuk, and P. Nieslony, “Finite element modelling of temperature distribution in the cutting zone in turning processes with differently coated tools,” *J. Mater. Process. Technol.*, vol. 164-165, pp. 1204-1211, 2005. https://doi.org/10.1016/j.jmatprotec.2005.02.136

[4] J. Rech, A. Giovenco, C. Courbon, and F. Cabanettes, “Toward a new tribological approach to predict cutting tool wear,” *CIRP Ann.*, vol. 67, no. 1, pp. 65-68, 2018. https://doi.org/10.1016/j.cirp.2018.03.014

[5] A. Giovenco, “Contribution à la simulation numérique de l’usure des outils de coupe”, *PhD Thesis Ecole Centrale de Lyon*, 2018.

[6] C. Courbon, T. Mabrouki, J. Rech, D. Mazuyer, E. D'Eramo, "On the existence of a thermal contact resistance at the tool-chip interface in dry cutting of AISI 1045: Formation mechanisms