*Article* **Tribological Behaviors of Inconel 718–Tungsten Carbide Friction Pair with Sulfur Additive Lubrication**

**Ye Yang 1,\*, Hao Luan 1, Songshan Guo 1, Fengbin Liu 1, Yuanjing Dai 2, Chenhui Zhang 3, Duzhou Zhang <sup>4</sup> and Gang Zhou <sup>4</sup>**

	- <sup>3</sup> State Key Laboratory of Tribology, Tsinghua University, Beijing 100084, China
	- <sup>4</sup> Beijing Key Laboratory of Long-Life Technology of Precise Rotation and Transmission Mechanisms, Beijing Institute of Control Engineering, Beijing 100094, China
	- **\*** Correspondence: yangye@ncut.edu.cn

**Abstract:** This work investigated the lubricating and anti-wear properties of several sulfur additives for a nickel-based superalloy–tungsten carbide friction pair. Compared with PAO40 without any active chemical compounds, the three kinds of sulfur additives could decrease the friction coefficient from 0.2 to 0.1 and the wear volume by 90%. Sulfurized fatty acid ester had the best performance under high temperature and heavy load with COF below 0.1 and the smallest wear volume. Furthermore, the lubricating mechanism was investigated by XPS. The physical adsorptive film and the tribochemical film together enhanced the friction-reducing and anti-wear performances of the lubricants. This effective lubricant for Inconel 718 can be applied to the machining of nickel-based alloy.

**Keywords:** Inconel 718; sulfur additives; boundary lubrication; wear mechanism

#### **1. Introduction**

Nickel-based superalloys such as Inconel 718 are widely used in aerospace, gas turbine, nuclear, and automotive industries because of their excellent high-temperature mechanical strength and corrosion resistance [1]. At the same time, nickel-based superalloys are recognized as difficult-to-machine materials. These metals exhibit serious problems during machining such as high cutting force, rapid tool wear, short tool life, and poor surface quality of the machined surface due to the physical, chemical, and thermal properties of metals [2–5]. Tungsten carbide tools are the first choice for their good thermal conductivity, high strength, and poor affinity with nickel [6]. It is of great importance to understand the relationship between nickel-based alloy and carbide material, especially the tribological behaviors, to reduce tool wear and improve tool life. The use of suitable cutting fluids can effectively improve the machining conditions of nickel-based superalloys [7–9]. The cutting fluid acts as a lubricant to reduce friction and as a coolant to cool the temperature at the cutting zone. The lubricating ability of a cutting fluid greatly influences the quality of the machined surface, as well as the tool life [10,11].

The chemical additives in the cutting fluid can act with the metal surface at high temperature and pressure; thus, a lubricant film is formed to reduce friction between the rake face and chips, the flank face, and machined surfaces. Sulfur additives are well known for their extreme-pressure performance and anti-wear characteristics [12–14]. The sulfur compounds, under extreme-pressure conditions, undergo chemical decomposition causing sulfur release (rupture of the R–S bond) and their reaction with the metallic surface that promotes the formation of an inorganic iron sulfide layer [12,15]. Most of the cutting fluids are designed for ferrous metals [11,16–18]. However, there is a matching problem between the additives and the workpiece materials [19,20]. Whether these additives have

**Citation:** Yang, Y.; Luan, H.; Guo, S.; Liu, F.; Dai, Y.; Zhang, C.; Zhang, D.; Zhou, G. Tribological Behaviors of Inconel 718–Tungsten Carbide Friction Pair with Sulfur Additive Lubrication. *Metals* **2022**, *12*, 1841. https://doi.org/10.3390/ met12111841

Academic Editor: George A. Pantazopoulos

Received: 7 October 2022 Accepted: 24 October 2022 Published: 28 October 2022

**Publisher's Note:** MDPI stays neutral with regard to jurisdictional claims in published maps and institutional affiliations.

**Copyright:** © 2022 by the authors. Licensee MDPI, Basel, Switzerland. This article is an open access article distributed under the terms and conditions of the Creative Commons Attribution (CC BY) license (https:// creativecommons.org/licenses/by/ 4.0/).

the same reaction with nickel-based superalloys is not clear. Few scholars have studied the lubricants for nickel-based alloy–tungsten carbide contacts. Moreover, metal working fluid which consists of several chemically active additives has a very complex chemical composition. The individual actions of each component are not easily identifiable. However, it is necessary for new metals such as nickel-based superalloys whose characteristics are very different from traditional ferrous metals. Therefore, it is important to separate contributions from different types of additives to the lubrication so as to select the most efficient lubricant molecules and optimize the components of the cutting fluid. This work focuses on the lubricating effect of different sulfur additives for the nickel-based superalloy (Inconel 718)–tungsten carbide (YG8) tribopair, and the lubricating mechanism is further investigated.

#### **2. Experimental Details**

Inconel 718 is the most widely used superalloy. YG8 (WC-Co) tungsten carbide is the optimal tool material for nickel-based superalloy machining. The specimen used in this paper was bought from Hengshihui company, Jiangsu, China. Table 1 shows the chemical composition of Inconel 718 superalloy. The main component is Ni (60.01 wt.%) followed by Cr (17.28 wt.%) and Fe (15.58 wt.%). The Table 2 displays the mechanical parameters of Inconel 718.

**Table 1.** Chemical compositions of Inconel 718 superalloy (wt.%).


**Table 2.** The mechanical parameters of Inconel 718.


The commercial sulfur additives were bought from Symarin company, Shanghai, China. Each additive contained a different amount of sulfur in the molecule (by weight), and the sulfur bonding mechanism was different in each additive. According to the manufacturer's data, the sulfur content of sulfurized olefin was as much as 40%, more than that of sulfurized fatty acid ester (17%) and sulfurized lard (10%). The kinematic viscosity of sulfurized olefin was much smaller than that of the other two kinds of additives for the smallest molecular weight. Table 3 shows the available information about the lubricants. Furthermore, PAO40, the viscosity of which is 396 mm2/s, was used as pure oil without any active elements for comparison.

**Table 3.** The basic parameters of sulfur additives.


The frictional tests were carried out utilizing a ball-on-disc apparatus SRV-IV (Optimol, Munich, Germany) under different lubricating conditions. The schematic diagram of the tester is displayed in Figure 1. The discs were Inconel 718 with a hardness of HRC 35. All of the specimens were polished before frictional tests by an automatic polishing/grinding machine, and a surface roughness (Sa) less than 40 nm was obtained. The counter specimen

was tungsten carbide YG8 ball with a diameter of 10 mm and surface roughness (Sa) of 25 nm. The hardness of the carbide ball was 89HRA. The samples were ultrasonically cleaned using acetone and ethanol and then ultrapure water successively, each for 10 min, before tests. The upper ball slid reciprocally against the stationary disc with an amplitude of 2 mm and frequency of 20 Hz for 5 min. Before the frictional test, plenty of lubricant was dropped onto the surface of the disc, and the ball returned to yield a normal load of 100 N.

**Figure 1.** The schematic diagram of SRV-IV tribo-tester.

After the frictional tests, the morphology of rubbing surfaces was observed using a LEXT™OLS5100 laser scanning confocal microscope (Olympus, Tokyo, Japan), and wear volume is calculated. Each test was repeated three times, and the average values were used. The relative errors were on the order of ±5%. A Quanta200 scanning electron microscope (SEM) (FEI, Hillsboro, OR, USA) combined with energy dispersion spectrometry (EDS) (FEI, Hillsboro, OR, USA) was used for the surface analysis of the investigated materials. The chemical compositions of the worn surfaces were characterized using a PHI Quantera SXM X-ray photoelectron spectrometer (ULVAC-PHI, Chigasaki, Japan).

#### **3. Results and Discussions**

#### *3.1. Friction and Wear Properties Lubricated with Sulfur Additives*

The curves of the friction coefficient and wear volume of the lower samples lubricated by the four kinds of lubricants are shown in Figure 2. For PAO40, the friction coefficient increased to 0.7 rapidly in the running-in period. Then, it decreased gradually and stabilized at about 0.2 after 200 s. When using sulfurized olefin as lubricant, the initial friction coefficient fluctuated to 0.17, before remaining stable at about 0.14 after 100 s. For sulfurized fatty acid ester and sulfurized lard, the friction coefficients remained stable at 0.1 and 0.11, respectively, until the end of the experiment. Compared with PAO40, the three kinds of sulfur-containing additives could reduce the friction coefficient significantly for the nickelbased superalloy. The friction coefficient of sulfurized fatty acid ester was the smallest, while that of sulfurized olefin was 27% larger.

Considering that tungsten carbide is much harder than nickel-based superalloys, wear occurs on the surface of the disc apparently while no wear is observed on the ball. After the frictional tests, the remaining solution and debris on the wear track are washed away by water and alcohol. From Figure 2b, it can be seen that the wear volume of the disc lubricated by PAO40 was 2.4 × <sup>10</sup><sup>7</sup> <sup>μ</sup>m3, about ten times that lubricated by the sulfur additives. Thus, basic oil without any active chemical elements failed to lubricate nickel-based superalloy under boundary lubrication. Compared with PAO40, the sulfur additives showed good lubricating and anti-wear performance for the Inconel 718–tungsten carbide friction pair. It is easy to generate inorganic protective films in the tribochemical reaction for an excellent anti-wear effect. For sulfurized fatty acid ester and sulfurized lard, the wear volumes were 2.6 × <sup>10</sup><sup>6</sup> <sup>μ</sup>m<sup>3</sup> and 2.8 × <sup>10</sup><sup>6</sup> <sup>μ</sup>m3, respectively. With sulfurized olefin lubrication, the wear volume was 4.1 × 106 <sup>μ</sup>m3. The relatively big wear volume is related to the long running-in

period during the friction test. A lubricating film was formed, but it was not stable enough to prevent direct contact between Inconel 718 and tungsten carbide completely during the running-in period. It can be concluded that sulfurized fatty acid had the best lubrication for Inconel 718 with the smallest COF and wear volume.

**Figure 2.** (**a**) Friction coefficients with (**b**) enlarged view, and (**c**) wear volume of the discs lubricated by different lubricants (*F* = 100 N, *f* = 20 Hz, *L* = 2 mm).

The micro-morphology of the friction pairs was obtained by SEM as shown in Figure 3. It can be seen that there were furrows on the surface of the superalloy discs and irregular block materials on the surface of the balls lubricated by the three kinds of additives. From Figure 3(a1), delaminated scars and abrasive particles can be seen on the surface of the disc lubricated by sulfurized olefin. There were obvious furrows on the worn surface lubricated by sulfurized lard (Figure 3(c1)). Furthermore, large dark blocky materials were detected on the ball surface (Figure 3(c2)). The worn surface was compared to the smoothest with no adhesive scar when lubricated by sulfurized fatty acid ester (Figure 3(b1)). However, there were still patches on the surface of the ball (Figure 3(b2)). To ascertain the ingredients of the materials on the surface (the red frame in Figure 3), EDS analysis was conducted. The results show that the nickel-based superalloy on the ball surface was transferred from the discs and adhered to the balls (Figure 3d). The materials on the balls were the same; thus, the EDS result is listed once. The main wear mechanism of the superalloy was significant adhesion, delamination, and ploughing. From the EDS results of the worn surface of the discs, the sulfur element was detected, which did not belong to the alloy. This demonstrates that sulfides formed on the metal surface during the friction process. However, the sulfur contents on the surface of the discs lubricated by the additives were different, seen from Table 4. The sulfur content of the worn surface lubricated by sulfurized olefin was the highest (12.37%), while the adhesion on the ball was relatively mild. It should be considered that the active sulfur content of sulfurized olefin was also the highest. The sulfur content

of the worn surface lubricated by sulfurized lard was the lowest (1.34%) with the most adhesive blocky material on the ball. For the sulfurized fatty acid ester, the adhesion on the ball was significantly decreased compared with the sulfurized lard. This indicates that the sulfide generated on the metal surface during the friction process played an important role in reducing the adhesion of nickel-based superalloy to tungsten carbide ball during friction. Sulfurized fatty acid ester contains active sulfur that can react with the metal surface at a lower temperature; hence, the surface quality of the disc and the adhesion of the ball were better than those lubricated by sulfurized lard. Meanwhile, the sulfurized fatty acid ester molecule has a long carbon chain which helps to separate the two surfaces; thus, direct contact was avoided compared with sulfurized olefin.

**Figure 3.** SEM morphology of the tracks on Inconel 718 and tungsten carbide ball lubricated by (**a1**,**a2**) sulfurized olefin, (**b1**,**b2**) sulfurized fatty acid ester, and (**c1**,**c2**) sulfurized lard; (**a3**,**b3**,**c3**) EDS spectrum of the red box in (**a1**,**b1**,**c1**); (**d**) EDS spectrum of the red box in (**a2**).


**Table 4.** The elements content of the above EDS analysis.

#### *3.2. Temperature Effect*

Temperature is very important for the chemical reaction of additives with tribo-surface. The influence of temperature on the friction and wear behavior of Inconel 718 sliding against WC-Co under the lubrication of sulfur additives was investigated. The experimental temperature was increased to 150 ◦C with the other experimental parameters the same as those in Section 3.1. Figure 4a shows the friction coefficient curve at 150 ◦C. For sulfurized olefin, the fluctuation of the friction coefficient was reduced at the beginning of the test compared with that at 30 ◦C. Furthermore, the friction coefficient stabilized at about 0.18 after 20 s. For sulfurized fatty acid ester and sulfurized lard, the friction coefficient remained at 0.1 from the beginning to the end of the test without running-in time at 150 ◦C. This indicates that the lubricant film formed rapidly and remained more stable under high temperature as expected. The wear volumes of the Inconel 718 discs are shown in Figure 4b. Compared with that at 30 ◦C, the wear volume of the discs with different sulfur lubricants did not change much. The wear volume lubricated by sulfurized fatty acid ester was slightly smaller compared to the other two kinds of additives. It can be seen that the sulfur lubricants had very good temperature stability for nickel-based superalloys.

**Figure 4.** (**a**) Friction coefficients and (**b**) wear volume when lubricated by sulfur additives at 150 ◦C (*F* = 100 N, *f* = 20 Hz, *L* = 2 mm).

The morphology of the lower disc and the upper tungsten carbide ball after frictional tests at 150 ◦C was observed by SEM, and the pictures are shown in Figure 5. Compared with 30 ◦C, the furrows on the surface of the discs decreased, especially for sulfurized fatty acid ester and sulfurized lard (Figure 5(b1,c1)). This indicates that high temperature promoted the formation of a lubricant film, thus increasing the surface quality of the nickel-based superalloy. With respect to the surface of the balls, the adhesive materials were also decreased when lubricated by sulfurized fatty acid ester and sulfurized lard (Figure 5(b2,c2)). However, the block material on the surface of the ball lubricated by sulfurized olefin increased significantly, indicating that adhesion was aggravated, as seen in Figure 5(a2). EDS analysis was performed on the wear scar of the nickel-based superalloy and the adhesion area on the tungsten carbide (the red box in Figure 5). The dark materials on the balls were the same; thus, the EDS results are listed once (Figure 5d). A nickel-based superalloy was transferred from the disc and adhered to the ball. It can be speculated that, under high temperature, the main wear mechanism of the superalloy was adhesion. The elements content of the EDS analysis is listed in Table 5. From the EDS spectrum of the surface lubricated by sulfurized fatty acid ester (Figure 5(b3)), it can be seen that the peak of sulfur in the wear scar increased significantly and the sulfur content increased from 2.71% (30 ◦C) to 16.33% (150 ◦C). The sulfide acted as an effective lubricant film, which significantly decreased the adhesion on tungsten carbide balls and the ploughing on nickel-based superalloy discs. For sulfurized lard, the sulfur content increased from 1.34% (30 ◦C) to 5.43% (Figure 5(c3)), and the surface quality of the frictional pairs improved. As for sulfurized olefin, the sulfur peak did not change very much, and the content of sulfur was 11.63%. This indicates that high temperature did not promote the tribochemical reaction of sulfurized olefin with nickel-based superalloy. Considering that the flash point of sulfurized olefin is 150 ◦C, when the experimental temperature increased to 150 ◦C, the sulfurized olefin began to volatilize continuously, and the material left in the friction area was unstable. This led to the aggravation of adhesion at 150 ◦C lubricated by sulfurized olefin. The flash points of the other two sulfur lubricants were both higher than 150 ◦C, and the molecules remained stable with temperature increasing. Moreover, the tribochemical reaction was promoted, leading to better surface quality and less adhesion.


**Table 5.** The elements content of the above EDS analysis.

#### *3.3. Load Capacity*

The extreme-pressure performance of the lubricants is also an important index used to measure the lubricating property. To illustrate the performance under extreme pressure, the load slope test results of the three kinds of sulfur additives are shown in Figure 6a. The test load increased from 100 N to1000 N in steps of 100 N. The friction coefficient of sulfurized olefin fluctuated significantly with each load increase. For sulfurized fatty acid ester and sulfurized lard, the friction coefficients remained stable at about 0.1 before the load increased to 600 N. When the load was more than 600 N, the friction coefficients had a little fluctuation. Thus, the load-bearing capacity of sulfurized fatty acid ester and sulfurized lard was better than that of sulfurized olefin.

**Figure 5.** SEM morphology of the tracks on Inconel 718 and tungsten carbide ball lubricated under 150 ◦C lubricated by (**a1**,**a2**) sulfurized olefin, (**b1**,**b2**) sulfurized fatty acid ester, and (**c1**,**c2**) sulfurized lard; (**a3**,**b3**,**c3**) EDS spectrum of the red box in (**a1**,**b1**,**c1**); (**d**) EDS spectrum of the red box in (**a2**).

The wear volumes of sulfurized fatty acid ester and sulfurized lard on nickel-based superalloy under different loads were further tested. The experimental load was set to 300 N, 500 N, 700 N, and 900 N, and the results are shown in Figure 6b. When sulfurized fatty acid ester lubricated the nickel-based superalloy, the wear volume increased steadily with the increase in load. As for sulfurized lard, the wear volume had a larger increase compared to sulfurized fatty acid when the load increased to 700 N. The big wear volume was consistent with the large friction coefficient, indicating unstable lubricant condition.

**Figure 6.** (**a**) Friction coefficient vs. time during a load slope test from 100 N to 1000 N and (**b**) wear volume under different loads (*f* = 20 Hz, *L* = 2 mm, *t* = 25 ◦C).

#### *3.4. Exploration of the Adsorption Characteristics*

XPS is a practical method to clarify the chemical states of elements within the adsorption film on the surface of tribopair. To further explain the lubrication mechanism of sulfur additives for nickel-based superalloy, the worn surfaces were tested by XPS. Figures 7 and 8 show the spectra of the several elements lubricated by sulfurized fatty acid ester and sulfurized olefin. It can be observed that the peak shapes and binding energies of the corresponding elements were similar. Therefore, the tribochemical reaction processes were the same when sulfurized fatty acid ester and sulfurized lard were used as lubricants for the nickel-based superalloy. Figure 7a shows typical XPS survey scans inside and outside the wear track over a binding energy at the range of 0–1400 eV with the lubrication of sulfurized fatty acid ester. The values were shifted 400,000 upward from the second line to show a clear contrast. The peak intensities of Ni2p and Fe2p inside the wear track are lower than those outside the wear track. Furthermore, a sulfur peak at 168 eV appeared at the position inside the wear track, while no sulfur was detected outside the track. This demonstrates that sulfide compounds remained after the frictional test when lubricated by sulfurized fatty acid ester. Ni2p inside the wear track was apparently lower than the substrate, which further demonstrates that some film existed on the wear track. To further investigate the way in which the sulfide compounds acted with nickel-based superalloy, detailed high-resolution XPS scans of Ni2p, Fe2p, S2p, and O1s were recorded, and the results are shown in Figure 7b–e. The peak at 852.8 eV is Ni–S, and the peak at 855.2 eV is Ni–SO4 (Figure 7b). The peaks at 710.11 eV and 723.4 eV correspond to Fe–S and FeSO4, respectively (Figure 7c) [11]. The S2p spectrum is shown in Figure 7d. The peak at 161 eV–162 eV corresponds to Fe–S, the peak at 162.8 eV corresponds to Ni–S, and the peak at 169.7 eV is the metal sulfate [21]. The O1s spectrum is shown in Figure 7e. The peak at 531.7 eV corresponds to S–O in –SO4, and the peak at 530.2 is metallic oxide. Combining the Ni2p, S2p, and O1s data, it can be inferred that NiSO4 may have existed on the surface of the wear track. Compared with the S2p spectra on the surface lubricated by sulfurized fatty acid ester, the peak intensities of Ni2p and Fe2p inside the wear track were higher when lubricated by sulfurized olefin (Figure 8a). Moreover, the peak of metal sulfates was much lower than that lubricated by sulfurized fatty acid ester (Figure 8c). It can be speculated that the tribochemical reaction film on Inconel 718 surface lubricated by sulfurized olefin was thinner than that lubricated by sulfurized fatty acid ester, leading to a higher COF and bigger wear volume.

According to the XPS results, the lubricating mechanism of the sulfur additives for Inconel 718–tungsten carbide contact can be summarized. In the frictional process, the molecules adsorb on the metal surface to form a physical adsorption protective film. The active element S reacts with the metal matrix Inconel 718. A tribochemical protective film composed of high-toughness inorganic salts such as nickel sulfide and nickel sulfate is

formed, playing a role in lubrication. The physical adsorptive film and the tribochemical film together enhance the friction-reducing and anti-wear performances of the lubricants.

**Figure 7.** (**a**) XPS survey scans of Inconel 718 surface after tribological test: (**b**) Ni2p, (**c**) Fe2p, (**d**) S2p, and (**e**) O1s inside the wear track lubricated by sulfurized fatty acid ester.

**Figure 8.** (**a**) XPS survey scans of Inconel 718 surface after tribological test, (**b**) S2p, and (**c**) O1s inside the wear track lubricated by sulfurized olefin.

#### *3.5. The Improvement of the Cutting Fluid*

The experimental results show that, for the nickel-based superalloy–tungsten carbide friction pair, sulfurized fatty acid ester had the best lubricating and anti-wear performance among the three kinds of extreme-pressure agents. The effect of sulfurized fatty acid ester on the lubricating performance of some kind of cutting fluid without sulfur-containing additives was tested. The cutting fluid was diluted to 5 wt.% with different content of sulfur additives. The frictional experiments were completed, and the friction coefficients are shown in Figure 9. The results show that the friction coefficient dropped to 0.133 with the concentration of sulfurized fatty acid ester at 1 wt.%. When the concentration of sulfurized fatty acid ester increased to 2%, the friction coefficient further dropped to 0.127. The ester concentration was further increased to 3%, and the friction coefficient did not continue to decrease and remained stable at 0.127. It can be seen that 2 wt.% sulfurized fatty acid ester could successfully decrease the friction coefficient of the cutting fluid.

**Figure 9.** The COF of some cutting fluids with different contents of sulfurized fatty acid ester (*F* = 100 N, *f* = 20 Hz, *L* = 2 mm, *t* = 25 ◦C).

#### **4. Conclusions**

In the present work, the tribological performance of three kinds of sulfur additives for the Inconel 718–tungsten carbide friction pair was investigated. The friction experiment results showed that sulfurized fatty acid ester possessed excellent antifriction (COF 0.1) and anti-wear performance (wear volume 90% smaller than that lubricated by PAO 40), particularly at the high temperature of 150 ◦C and at heavy load. The lubrication mechanism of the sulfur additives for the Inconel 718–tungsten carbide friction pair was investigated using XPS. The physical adsorptive film and the tribochemical film together enhanced the friction-reducing and anti-wear performance of the lubricants. This effective lubricant for Inconel 718 can be applied to the machining of nickel-based superalloy.

**Author Contributions:** Conceptualization, Y.Y. and C.Z.; methodology, Y.Y. and Y.D.; validation, F.L.; formal analysis, F.L.; investigation, H.L., S.G. and Y.Y.; resources, C.Z., Y.Y. and G.Z.; data curation, H.L. and S.G.; writing—original draft preparation, H.L.; writing—review and editing, Y.Y.; visualization, Y.Y.; supervision, F.L.; project administration, F.L. and D.Z.; funding acquisition, C.Z., Y.Y. and G.Z. All authors have read and agreed to the published version of the manuscript.

**Funding:** This research was funded by National Key Research and Development Program of China grant number 2018YFB2002204 and National Natural Science Foundation of China grant number 52005010 and U1837602.

**Data Availability Statement:** The data presented in this study are available on request from the corresponding author.

**Conflicts of Interest:** The authors declare no conflict of interest.

#### **References**


**Thi-Bich Mac 1, The-Thanh Luyen <sup>1</sup> and Duc-Toan Nguyen 2,\***


**Abstract:** This study aimed to investigate the effects of Thermal-Assisted Machining (TAM) on SKD11 alloy steel using titanium-coated hard-alloy insert cutting tools. The microstructure, material hardness, chip color, cutting force, chip shrinkage coefficient, roughness, and vibration during TAM were evaluated under uniform cutting conditions. The machining process was monitored using advanced equipment. The results indicated that thermal-assisted processing up to 400 ◦C did not alter the microstructure and hardness of the SKD11 alloy steel. However, a significant variation in chip color was observed, indicating improved heat transfer through TAM. The cutting force, vibration amplitude of the workpiece, and surface roughness all decreased with increasing TAM. Conversely, the chip shrinkage coefficient of the machined chips tended to increase due to the high temperatures.

**Keywords:** SKD11 alloy steel; machinability; thermal-assisted machining (TAM)

#### **1. Introduction**

The cutting mode and tool geometry parameters play a critical role in machining techniques, especially when working with materials that are challenging to machine due to their high hardness. These parameters affect various output parameters of the machining process, such as cutting force, cutting heat, cutting tool wear, cutting process vibration, surface roughness, and chip geometry. To improve productivity, part quality, and costeffectiveness, researchers are constantly seeking new technological solutions to support the machining process, including the use of smooth cold techniques, new cutting tool materials, vibration-assisted cutting, and thermal-assisted machining. Difficult-to-machine materials, characterized by high hardness, good wear resistance, and little change in mechanical properties at high temperatures, are widely used in various industries, such as mechanical, automotive, aerospace, aviation, defense, medical, and electrical-electronicsautomation. Surveys have shown that over 30% of milling, turning, and drilling operations are performed on such materials [1].

Thermal-assisted machining (TAM) is a machining process that involves heating the workpiece before machining it using traditional or CNC machines. This technique has been extensively used in industrial production since its introduction in 1945 [1,2]. Compared to traditional machining methods, TAM offers several advantages, such as increased tool life, reduced cutting force, decreased tool wear, and improved surface quality, resulting in enhanced productivity [3–5]. TAM is suitable for both cutting processes (e.g., turning, milling, drilling) and deformation processes (e.g., forging, stamping, drawing). Several heating methods can be employed for TAM, including electrical current heating, laser beam heating (LAM), plasma heating (PEM), furnace heating (FAM), and induction heating (IAM). While each method has its own benefits and drawbacks, induction heating (IAM) is particularly effective due to its high heating capacity, ease of use, and affordability, making it a great choice for metalworking operations [6].

**Citation:** Mac, T.-B.; Luyen, T.-T.; Nguyen, D.-T. Assessment of the Effect of Thermal-Assisted Machining on the Machinability of SKD11 Alloy Steel. *Metals* **2023**, *13*, 699. https://doi.org/10.3390/ met13040699

Academic Editors: George A. Pantazopoulos and Badis Haddag

Received: 17 February 2023 Revised: 27 March 2023 Accepted: 31 March 2023 Published: 3 April 2023

**Copyright:** © 2023 by the authors. Licensee MDPI, Basel, Switzerland. This article is an open access article distributed under the terms and conditions of the Creative Commons Attribution (CC BY) license (https:// creativecommons.org/licenses/by/ 4.0/).

Many studies have suggested eco-friendly dry and minimum quantity lubrication or optimization to improve machining productivity, increase tool life, and enhance product quality [7–9]. However, for challenging-to-machine materials, like those commonly used in heavy-duty industries such as automotive, marine, and aviation, advanced machining methods such as diamond grinding or discharge machining are often used but limited by low material removal rates, expensive tools, and rapid wear [2]. Among these materials, SKD11 Tool Steel is frequently utilized in mold and automotive sectors due to its hardness, strength, and ductility [10]. Although traditional techniques like diamond grinding or discharge machining are limited by high costs and specialized technologies, TAM has been introduced as an innovative solution to the challenges of machining SKD11. Studies have indicated that machining with TAM can reduce cutting forces by up to 40%, increase tool life, enhance surface roughness by up to 50%, and increase material removal rate when compared to machining at room temperature [11–13].

Recent research has shown that using thermal-assisted machining (TAM) can offer several advantages, including improved tool life, enhanced productivity, and better working conditions between the tool and workpiece [14–16]. Studies have revealed that workpiece temperature is directly proportional to tool life and inversely proportional to cutting speed. Induction heating has been investigated for milling titanium alloy Ti-6Al-4V, with heat treatment significantly improving tool life and material removal rate, by up to 169.4% at a temperature of 650 ◦C. However, heating above 640 ◦C may negatively impact the workpiece's mechanical properties, reducing machining efficiency [5,17]. Studies have also shown that cutting and shear forces are inversely proportional to temperature, with a 13% reduction in cutting force observed when machining at 500 ◦C compared to room temperature. A novel approach to machining Inconel 718 was presented by Wang et al. [18], which combines traditional turning, cryogenic enhanced machining, and plasma heating. Cold working techniques can reduce the heat generated during cutting, enhancing the cutting tool's longevity. The implementation of this method has been shown to result in a substantial enhancement of the machining process, with the surface gloss of the machined material increased by a factor of 2.5, the required machining force decreased by half, and the cutting tool's lifespan prolonged by a factor of 1.7 compared to traditional machining at room temperature.

Efficiently machining hard alloys has long been a challenge, but thermally assisted machining (TAM) offers a promising solution. TAM involves cutting the material and then softening it with an external heat source, which has been shown to reduce the hardness and tensile strength of the workpiece material, thereby improving machinability [19]. Research indicates that TAM is linked to higher material removal rates, better control of machining time, and improved surface finishes. TAM is particularly advantageous for machining bio-implant titanium alloys, which require high precision. Hard steels can also be successfully machined using TAM, resulting in lower cutting forces, increased tool life, and higher material removal rates [20–24]. TAM can reduce cutting force amplitudes and chip morphology changes, resulting in less vibration and better surface integrity. However, machining certain hardenable steels, such as 1090 steel, can lead to increased cutting forces due to phase transformation hardening when the laser-preheated part enters the cutting zone. Softening the workpiece by reaching surface temperatures of 300–400◦C for an uncut chip thickness of 0.05 mm, on the other hand, can reduce the magnitude and amplitude variation of cutting forces, and limit the evolution of tool wear [25]. TAM also induces a change in chip morphology from sawtooth to continuous, improving the surface finish. While TAM can lead to higher cutting forces in some hardenable steels, it can also enhance surface integrity [26].

TAM techniques have been extensively researched and used in production. However, little is known about the use of magnetic induction heating for mold steel, particularly when milling difficult materials like SKD11 steel, which is commonly used in industry. This study aims to evaluate the influence of TAM on milling SKD11 steel and its effect on the material's machinability by analyzing cutting force, vibration amplitude of the workpiece, and surface

roughness. This analysis provides a basis for selecting appropriate TAM process parameters. A significant challenge when using electromagnetic induction heating in machining is its application to large parts with varying sizes or complex shapes. This study proposes a potential technological solution for processing difficult materials using electromagnetic induction heating. The results demonstrate that the heating process effectively reduces cutting force, cutting heat, and vibration during cutting, while improving the surface quality of the workpieces. The findings of this study have practical implications and are applicable to the manufacturing industry.

#### **2. Experimental Setup**

The ease or difficulty of processing a material is known as its machinability, which can be evaluated using parameters such as tool life, MRR, cutting force, cutting process vibration, and surface gloss of the machined material. The machinability of a material is greatly affected by its microstructure, which can be further influenced by the cutting mode. Therefore, it is essential to assess the machinability of materials during heating processing and compare it to conventional machining to determine the impact of heating on the machinability of SKD11 steel materials.

#### *2.1. Schematic of Experimental System*

This study establishes an experimental research model that is based on the research objectives and the available experimental equipment, as depicted in Figure 2.

#### *2.2. Materials and Testing Equipment*

SKD11 steel is a frequently used alloy steel in mold processing, as defined by JIS-G4404. This steel is notable for its hardness, strength, and ductility, and can maintain its hardness even at high temperatures for an extended period, making it suitable for use in the production of molds for extrusion, plastic injection, pressure, and components that require specific performance characteristics. Table 1 presents a breakdown of the SKD11 steel components, expressed as a weight percentage (wt%).


**Table 1.** The chemical composition of SKD11 steel (wt%) [27].

In this study, a machining experiment was conducted on SKD11 steel test specimens using a Taiwanese MC500 milling machine. The billet was rough-worked and had dimensions of 70 × <sup>31</sup> × 80 mm3, and a chamfer of 7 mm × 7 mm was applied for uniform contact with the induction coil. For the milling process, a 40 mm face milling cutter was used, equipped with a titanium-coated hard alloy insert manufactured by PRAMET, a tool manufacturer from the Czech Republic. Specifically, the APKT 1604PDR–GM hard alloy piece was utilized, with cutting edge parameters such as the rake angle, clearance angle, and cutting-edge radius set at 24◦, 11◦, and 0.8 mm, respectively. Other geometric parameters, including l = 16 mm, d = 9.44 mm, s = 5.67 mm, and d1 = 4.6 mm, were also specified. No coolant was applied during the machining process. One significant challenge in this research was the application of electromagnetic induction heating in machining, as depicted in Figure 1d.

**Figure 1.** Components of the experimental system. (**a**) MC500 Milling Machine, (**b**) Sample Workpices, (**c**) Hard Alloy Insert, and (**d**) Power Source-Frequency Generator.

The study utilized the Axiovert 25 CA optical microscope (Figure 3a) to examine the microstructure of the material after heating at different temperatures and compare it with the original sample. This microscope, specifically designed for materials research, is used in conjunction with Image-Pro Plus image analysis software to analyze the material phase. The study also employed the Brinell hardness tester (Figure 3b) to test the initial and post-heating samples for hardness.

(**a**)

**Figure 2.** Schematic of Experimental System. Schematic diagram (**a**) and experiment photograph in working zone (**b**).

**Figure 3.** Equipment used in the tests: (**a**) Axiovert 25 CA Optical Microscope, (**b**) Brinell Hardness Tester.

#### **3. SKD11 Steel Material Machinability Results**

#### *3.1. Effect of TAM on Microstructure of SKD11 Steel*

To conduct an accurate examination of the microscopic structure of materials, proper sample preparation is crucial. The process of preparing metal samples involves multiple stages, including cutting, grinding, polishing, and impregnation. Specimens are initially machined following established standards to achieve a size that can be impregnated. Figure 4 displays the experimental samples and their cross-sections for SEM examination.

Figure 4 illustrates the microstructural analysis of SKD11 alloy steel samples before and after thermal-assisted machining (TAM) at different temperatures, using optical microscopy. Figure 4a presents an image of the microstructure of the original material sample, while Figure 4b–d show images of the microstructure after TAM at different temperatures. The analysis reveals that the microstructure of the samples after TAM at room temperature and 400 ◦C is similar to that of the original material. Specifically, the microstructure of the four samples includes white plates and bright round particles of chromium Cr7C3 carbides, spherical dark dots of Cementite, and a light background of Pearlite. These observations suggest that the microstructure of the specimen remains unchanged even when the temperature is increased to 400 ◦C, which is below the phase transition temperature of 700 ◦C for the SKD11 alloy steel [28].

#### *3.2. Effect of TAM on Hardness*

To evaluate the impact of elevated temperatures on workpiece hardness after heating, the workpiece was exposed to the designated temperature and allowed to cool naturally via exposure to ambient air. Surface hardness measurements necessitated meticulous sample preparation, including cutting, grinding, polishing, and impregnation. In order to prevent the influence of cutting heat, wire cutters were employed during sample cutting. The grinding process aimed to minimize undulation caused by the varying hardness of structural elements, while polishing removed any coarse grinding marks and scratches. Finally, deformations from the cutting, grinding, and polishing stages were either eliminated or leveled to a size sufficient for removal with an impregnating agent. Three hardness measurements were taken on each sample using a Brinell hardness tester at distinct locations, as illustrated in Figure 5. The results, presented in Figure 5, showed that increasing temperatures resulted in decreased specimen hardness to 2–3 HB at 200 ◦C and 300 ◦C. However, at 400 ◦C the sample demonstrated increased hardness compared to the other temperatures, although this increase was not significant within the chosen experimental temperature range.

**Figure 4.** Microstructure of Material Samples at Different Temperatures: (**a**) Room Temperature, (**b**) 200 ◦C, (**c**) 300 ◦C, and (**d**) 400 ◦C.

**Figure 5.** Workpiece hardness after heating at different temperatures and natural air-cooling.

#### *3.3. Effect of TAM on Chip Color*

The geometry and morphology of chips play a crucial role in assessing the machinability of materials. Chip geometry affects cutting force, cutting heat, and tool wear, while chip morphology provides valuable insights for milling tool design. Therefore, when analyzing the impact of heating on the machinability of SKD11, examining chip formation is crucial. Figure 6 shows chip images obtained during the milling of SKD11 steel under different cutting modes and heating conditions. Figure 6a–d display chip formations during machining at various temperatures. The formation of wire chips is due to the workpiece material's ductile properties. The chip color varies significantly, with the normal machining chip appearing purple-black in color due to excessive cutting heat generated during machining at room temperature (Figure 6a). In contrast, chips appear bright white (Figure 6b) and yellow (Figure 6c) when machined at 200 ◦C and 300 ◦C, respectively. The chip appears darker yellow when machined at 400 ◦C (Figure 6d). This variation can be attributed to the uniform transfer of heat between the cutting tool, workpiece, and chip during heating processing. The high temperature reduces the material's tensile strength, mechanical strength, and yield stress, while increasing its deformation, improving heat transfer conditions, and decreasing the bonding force between metal molecules [19,20]. As a result, chip removal becomes easier, cutting heat is reduced, and chips appear light-colored during hot machining. The results suggest that temperatures between 200–300 ◦C preserve the chip color compared to the original substrate material.

**Figure 6.** Chip Formation at Different Temperatures: (**a**) Room Temperature, (**b**) 200 ◦C, (**c**) 300 ◦C and (**d**) 400 ◦C.

#### *3.4. Effect of TAM on Cutting Force*

An experimental study was conducted to investigate the effect of elevated temperature on cutting forces (F) during milling with identical cutting parameters at room temperature and elevated temperature. The cutting force is a dynamic phenomenon and changes throughout the machining cycle [29]. Figure 7 displays the cutting force transformation with respect to tool distance during milling at room temperature (25 ◦C) and elevated temperature (200 ◦C). The cutting parameters used were a cutting speed of 235 m/min, feed rate of 305 mm/min, and cutting depth of 1.5 mm. The average cutting force (F) was calculated from the component cutting forces (*Ff*, *Fp*, *Fc*) using Equation (1). Table 2 shows the component cutting force values and the average cutting force under conventional machining and elevated temperature conditions. The results indicate that the *Fp* component cutting force (direct force) has the largest value, while the radial force (*Ff*) and axial force (*Fc*) have relatively smaller values. The average cutting force reduction at elevated temperature compared to conventional machining is 37.5%, as determined by Equation (2)).

$$F = \sqrt{F\_f^2 + F\_p^2 + F\_c^2} \tag{1}$$

$$
\Delta F(\%) = \frac{F\_T - F\_R}{F\_R} \times 100\% \tag{2}
$$

where *FT* represents the cutting force during heating and *FR* represents the cutting force when machining at room temperature (T = 25 ◦C).

**Figure 7.** Cutting force during machining at room temperature and at elevated temperature (200 ◦C).


**Table 2.** Average cutting force at room and elevated temperature.

The experimental results (shown in Figure 8) indicate a significant reduction in cutting force (*F*) during machining at 200 ◦C compared to conventional machining. As the heating temperature increased to 300–400 ◦C, the cutting force decreased at a slower rate. The maximum reduction in cutting force (Δ*F*) was 65.1% during milling at 400 ◦C. This reduction can be attributed to the weakening of the strength and bonding between metal molecules due to the heating process, making the cutting process easier. Moreover, in the second strain region, the compressive stress was found to decrease during heating, which led to a decrease in cutting force.

#### *3.5. Effect of TAM on Surface Roughness (Ra)*

The surface quality of a workpiece is critical and influenced by various factors, including the machining method, cutting tool geometry, and machining environment. Surface roughness, in particular, is essential in determining the workability and performance of the final product. This study used the average deviation criterion *Ra* to evaluate surface roughness. According to ISO standards, Ra is the arithmetic average of the absolute values of the profile over the reference length range (L = 250 μm) for machined surfaces with a smoothness level of 8–11. The research investigated the effect of elevated temperatures on Ra during Thermal Assisted Machining (TAM) while maintaining the same cutting parameters, including speed, feed rate, and cutting depth, under various temperature conditions. The method of surface roughness measurement is shown in Figure 9a, where the measuring head moves perpendicularly to the machining trace, and each sample is measured at three locations (1, 2, 3) to ensure reliability, within the reference length range of 250 μm, as shown in Figure 9b,c. Results indicated that heating the workpiece before machining significantly reduced surface roughness compared to conventional machining. This is because the thermal softening of the material results in increased smoothness and stability during the cutting process.

$$
\Delta \text{Ra} (\%) = \frac{\text{Ra}\_{\text{R}} - \text{Ra}\_{\text{T}}}{\text{Ra}\_{\text{R}}} \times 100\% \tag{3}
$$

where RaR and RaT represent the surface quality (Ra) at room and high temperature, respectively.

**Figure 8.** The mean cutting forces and the corresponding reductions observed during machining at various temperatures.

**Figure 9.** *Cont*.

**Figure 9.** Measurement of Surface Roughness with Equipment and Method (**a**) and Comparison of Surface Roughness Profiles for Machining at Room Temperature (**b**) and 200 ◦C (**c**).

In Figure 10, the average Ra values for machining at various temperatures are plotted. The decrease in Ra was calculated using Equation (3). The results indicate that there is an inverse relationship between Ra and workpiece temperature. The greatest reduction in roughness, 47.1%, was observed at 400 ◦C.

#### *3.6. Effect of TAM on Cutting Vibration*

To investigate the impact of temperature-assisted machining (TAM) on vibration, experiments were conducted using the same cutting parameters of V, f, and t at 235 m/min, 305 mm/min, and 1.5 mm, respectively, but at various temperature conditions. The vibration amplitudes in directions of (X, Y) as (*AX* and *AY*) were summed to obtain the *AXY* using Equation (4).

$$A\_{XY} = \sqrt{A\_X^2 + A\_Y^2} \tag{4}$$

Figure 11 displays the vibration amplitudes at room temperature (a) and at 200 ◦C (b) in directions of (X, Y). The analysis of the no-load and individual vibration data did not indicate any resonance. The results showed that TAM reduced the vibration amplitude compared to conventional machining. The reduction in cutting vibration amplitude (Δ*AXY*) during TAM, as compared to milling at room temperature, is calculated using Equation (5).

$$
\Delta A\_{XY}(\%) = \frac{A\_{XY-R} - A\_{XY-T}}{A\_{XY-R}} \times 100\% \tag{5}
$$

where *AXY-R, AXY-T* represent the vibration amplitudes in directions of (X, Y) during machining at room and high temperature, respectively.

**Figure 10.** Surface Roughness and Reduction at Elevated Machining Temperatures.

**Figure 11.** The vibration results at room (**a**) and heating at 200 ◦C (**b**) in the X and Y directions.

Figure 12 illustrates the vibration amplitude and reduction in vibration amplitude during SKD11 steel machining at various temperatures, as compared to milling at room temperature. The results indicated that A*XY* decreased by 13.2% at 200 ◦C, 14.9% at 300 ◦C, and 15.8% at 400 ◦C. These findings indicate increased stability during SKD11 steel processing when conducted at higher temperatures, due to the reduction in metallic bond strength resulting in an easier cutting process. However, the reduction in vibration amplitude did not exhibit a significant change when high temperature was used to facilitate the cutting process.

**Figure 12.** Vibration amplitude and vibration amplitude reduction during machining at various temperature conditions.

#### *3.7. Effect of TAM on Chip Shrinkage Coefficient (K)*

*K* is a vital parameter used to assess the plastic strain of a material during machining, and it has a significant influence on the dimensional accuracy of the machined surface. Several factors affect *K*, including the mechanical properties of the workpiece material, cutting tool geometry, cutting mode, and other cutting conditions. This study aims to analyze the impact of elevated temperatures on *K* and compare it to conventional machining under the same cutting mode. The objective is to assess the material's softening and formability at high-temperature conditions. The chip shrinkage coefficient, computed using Equation (6), is utilized to determine *K*.

$$K = \frac{1000 \cdot Q}{\rho \cdot L\_f \cdot f \cdot t} \tag{6}$$

where *Q* and ρ represent the mass of chip (g) and the material density (g/cm3), respectively, *Lf* is the chip cutting length (mm), and *f* and t are the feed rate (mm/rev) and cutting depth (mm), respectively.

Figure 13 displays the chip shrinkage coefficient at different temperatures for the machining process conducted at cutting parameters of V = 235 m/min, *f* = 305 mm/min, and *t* = 1.5 mm. The results demonstrate that the chip shrinkage coefficient increases with heating compared to conventional machining at room temperature. The increase in *K*, denoted as Δ*K*, is calculated using Formula (7). The chip shrinkage coefficient experiences a 27.6% increase when the workpiece is heated to 200 ◦C, and the maximum increase in *K* of 46.5% is observed at a temperature of 400 ◦C. This trend suggests that an increase in temperature enhances the cutting process by improving the chip shrinkage coefficient. The observed phenomenon can be attributed to the material softening under high temperatures, resulting in weakened atomic bonds and increased metal deformation, leading to a higher chip shrinkage coefficient.

$$
\Delta K(\%) = \frac{K\_T - K\_R}{K\_R} \times 100\% \tag{7}
$$

where *KT* and *KR* represent the chip shrinkage coefficient during heating and at room temperature (T = 25 ◦C), respectively.

**Figure 13.** Chip shrinkage coefficient and chip shrinkage increase during machining at various temperatures.

#### **4. Conclusions**

The aim of this study was to investigate and compare the effectiveness of using electromagnetic induction heating versus conventional methods for machining SKD11 steel, a material that is difficult to cut. The study analyzed various output parameters, including chip geometry, chip shrinkage coefficient, vibration amplitude, surface roughness, and cutting force. Experimental results indicated that the TAM process did not alter the material's microscopic structure in the temperature range of 200 ◦C to 400 ◦C, and the machined workpiece, cooled in air, maintained its original hardness. Additionally, the chip geometry changed and cutting force significantly decreased, with a maximum reduction of 65.1%. The *K* value increased by 31.7%, while surface roughness decreased notably, with a maximum reduction of 47.1% at 400 ◦C. Lower vibration amplitude also indicated a more stable machining process compared to traditional methods. These findings demonstrate the potential of this method to enhance machining performance and contribute to the development of high-precision and high-efficiency machining processes for difficult-tocut materials.

**Author Contributions:** Conceptualization, D.-T.N.; Data curation, T.-B.M.; Formal analysis, D.-T.N.; Funding acquisition, T.-T.L.; Investigation, T.-T.L.; Methodology, D.-T.N.; Project administration, T.-B.M.; Resources, T.-B.M.; Software, T.-T.L.; Supervision, D.-T.N.; Validation, D.-T.N.; Visualization, D.-T.N.; Writing—original draft, T.-B.M.; Writing—review & editing, D.-T.N., T.-B.M. All authors have read and agreed to the published version of the manuscript.

**Funding:** This work was supported by Vietnam Ministry of Education and Training (MOET) under grant number B2022-BKA-08.

**Acknowledgments:** This work was supported by the Vietnam Ministry of Education and Training (MOET) under grant number B2022-BKA-08.

**Conflicts of Interest:** The authors declare no conflict of interest.

#### **References**


29. Thi-Hoa, P.; Thi-Bich, M.; Van-Canh, T.; Tien-Long, B.; Duc-Toan, N. A study on the cutting force and chip shrinkage coefficient in high-speed milling of A6061 aluminum alloy. *Int. J. Adv. Manuf. Technol.* **2017**, *98*, 177–188. [CrossRef]

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