A Comparative Study on Al_0.6Ti_0.4N Coatings Deposited by Cathodic Arc and HiPIMS in End Milling of Stainless Steel 316L

Abstract

The machining of austenitic stainless steel alloys is usually characterized by high levels of adhesion and built-up edge; therefore, improving tribological conditions is fundamental to obtaining higher tool life and better surface finish. In this work, three different Al_0.6Ti_0.4N coatings are compared, two deposited by Cathodic Arc Evaporation (CAE) and one with High-Power Impulse Magnetron Sputtering (HiPIMS). The effects of the micromechanical properties and the microstructure of the coatings were then studied and related to the machining performance. Both arc-deposited coatings (CAE 1 and 2) exhibited similar average tool life, 127 min and 128 min, respectively. Whereas the HiPIMS lasted for only 21.2 min, the HiPIMS-coated tool had a much shorter tool life (more than six times lower than both CAE coatings) due to the intense adhesion that occurred in the early stages of the tool life. This higher adhesion ultimately caused built-up edge and chipping of the tool. This was confirmed by the cutting forces and more deformation on the shear band and undersurface of the chips, which are related to higher levels of friction. The higher adhesion could be attributed to the columnar structure of the HiPIMS and the (111) main texture, which presents a higher surface energy when compared to the dominant (200) from both arc depositions. Studies focused on tribology are necessary to further understand this relationship. In terms of micromechanical properties, tools with the highest plasticity index performed better (CAE 2 = 0.544, CAE 1 = 0.532, and HiPIMS = 0.459). For interrupted cutting machining where adhesion is the main wear mechanism, a reserve of plasticity is beneficial to dissipate the energy generated during friction, even if this was related to lower hardness levels (CAE 2 = 26.6 GPa, CAE 1 = 29.9 GPa, and HiPIMS = 33.6 GPa), as the main wear mechanism was adhesive and not abrasive.

1. Introduction

Stainless steel alloys are known for their high ductility, high tensile strength, high work-hardening rate, and low thermal conductivity. These properties, allied with the alloys’ good corrosion resistance, make them extensively used in various industry fields [1]. However, these same properties impair their machining, making them generally considered a difficult-to-cut material [2].

Due to the high ductility, the adhesive wear is intensified, as the workpiece material tends to adhere to the cutting tool surface. This can result in severe built-up edge (BUE) formation. Heat dissipation is also a problem due to the low thermal conductivity of stainless steel, causing thermal softening of the tool, which accelerates tool wear. Furthermore, the high work-hardening tendency results in higher cutting forces, which also contributes to premature wear [3].

To enhance productivity and surface quality, selecting suitable tool materials and geometries is essential. Coating applications on milling tools can significantly improve their life, protecting the tool against diffusive and abrasive wear, minimizing temperature fluctuations, and improving tribological conditions [4]. Researchers have recently explored a variety of coatings for the milling of stainless steel alloys: TiAlN [4,5,6,7,8,9,10,11,12], AlTiN [13,14,15,16,17,18,19], AlCrN [5,11,13,20], TiN/TiCN/Al₂O₃ [1,14,17,21], TiAlN/TiN [1,22], TiAlSiN [12], CrTiAlN [23], AlSiN [24], AlCrNbN [13], TiAlN/AlCrN [4], TiAlN/NbN [25], AlTiN/CrMoN [19].

The majority of these studies explored AlTiN-based coatings. However, most of them were not mainly focused on a comparison of different coatings, and therefore, there is still a gap in the understanding of the relationship of how different deposition parameters and techniques affect micromechanical and tribological properties during the milling of austenitic stainless steel alloys. Coatings based on aluminum titanium nitride are currently one of the most used and versatile coatings. The control of the ratio of Al and Ti is very important to obtain superior mechanical properties and wear resistance of the coating [26]. For contents up to 67% of Al, the AlTiN coating is generally composed of a single solid solution of B1 NaCl-type crystal structure in the cubic phase. For higher Al ratios, the wurtzite hexagonal phase of the AlN occurs, which causes a significant reduction in its hardness [27]. In studies on turning of austenitic stainless steel alloys, the best results were found when using a ratio of 60% aluminum and 40% titanium or Al_0.6Ti_0.4N [3,26,28].

However, relying solely on the chemical composition is not enough to determine the coating properties, as the deposition technique, parameters, and coating design can dramatically influence the tool’s performance. In this current research, two different deposition methods were used: Cathodic Arc Evaporation (CAE) and High-Power Impulse Magnetron Sputtering (HiPIMS).

Despite the dominance of CAE deposition when considering the industrial scale for physical vapor deposition (PVD) coatings in machining tools, an emerging PVD technology that has been an important research topic since it was proposed is High-power Impulse Magnetron Sputtering (HiPIMS) [29]. The principle of HiPIMS technology is the same as that of conventional magnetron sputtering. The difference is that a higher pulse peak power and lower duty cycle are used compared to magnetron sputtering. This results in a high-density and high-energy plasma for coating deposition [30].

The HiPIMS technique offers a combination of benefits from both conventional magnetron sputtering and cathodic arc evaporation (CAE). It provides a significant advantage in terms of customizing coating composition, microstructure, and characteristics. Due to its dense plasma and high ionization degree of sputtered material, it enables the deposition of high-quality thin films with a fully dense structure, low surface roughness, and superior hardness, elastic modulus, and wear properties [31]. The main drawback of HiPIMS technology is its deposition rate, which is considerably lower than CAE [32].

In addition to a comparison of commercially available CAE and HiPIMS coatings, this paper proposes one extra customized arc deposition, called CAE 2. The commercially based arc deposition will be represented as CAE 1. CAE 2 was performed using a pulsed bias technique. Studies have indicated that with the use of pulsed bias with an adequate frequency and duty cycle, it is possible to deposit coatings with lower residual stress, lower substrate deposition temperature [33], and a higher deposition rate when compared to the conventional constant bias [34]. The rearrangement of atoms during short periods of high mobility can relieve stress introduction while maintaining adequate levels of hardness and the elastic modulus [35].

By selecting the adequate duty cycle during the pulsed bias deposition, it is possible to obtain coatings with different properties. Reducing the duty cycle (% of bias on time), the overall average energy supplied during the deposition is reduced, which can lead to a less dense coating with lower compressive stress [36,37]. The control of residual stresses during deposition is very important, since excessive stress can cause delamination and consequent catastrophic failure of the coating [35]. This is even more important when considering the complex geometry of solid carbide endmill, which makes the deposition of uniform coatings challenging, especially on the sharp edges where coating delamination might occur.

As mentioned before, AlTiN coatings are widely used in machining. However, much of what has been reported is mainly focused on testing different chemical compositions and using doping elements. In addition to this, studies focused on the optimization of deposition parameters in relation to coatings’ mechanical properties very often do not study their application in machining processes. The effect of different deposition techniques and deposition parameters and their performance in machining when using tools with complex geometry for high-speed machining of stainless steel with a high tendency of adhesion is not fully explored. This work investigates the effects of the micromechanical properties, microstructure, defects, and texture of the coatings deposited. It also addresses how they are related to improving machining performance and tribological conditions, highlighting the importance of reducing the adhesion levels that are intensified in dry milling of such alloys.

The aim of this research is to perform a comprehensive study comparing the performance of Al_0.6Ti_0.4N coatings using two different arc deposition techniques, CAE 1 and 2, and one HiPIMS during the milling of 316L stainless steel. Results will be discussed in terms of coating characterization and micromechanical properties, such as hardness, H/E ratio, H³/E², plasticity index (PI), and fracture toughness, and how those relate to tool wear, forces, chip formation, and surface roughness of the workpiece. The discussion relating the parameters and the tool’s performance provides a guideline for selecting adequate coatings during the end milling of materials with a high tendency of adhesion.

2. Materials and Methods

2.1. Coating Deposition and Characterization

Three different coatings using the same target composition (Al_0.6Ti_0.4) were deposited on 3CH0400MS012A solid carbide endmills and on cemented carbide polished square coupons of grade K68, both provided by Kennametal (Ebermannstadt, Germany). Before the deposition, the coupons and endmills were cleaned in an ultrasonic bath using acetone for 8 min. The depositions were performed with two different sets of parameters for cathodic arc deposition (CAE 1 and CAE 2) and one set of parameters for HiPIMS.

The parameters for the HiPIMS deposition and CAE 1 were chosen based on commercial applications, both deposited by Kennametal. CAE 1 was deposited using an INNOVA coating system (Oerlikon Balzers, Pfaeffikon, Switzerland), and HiPIMS using a CC800 HiPIMS system (CemeCon, Würselen, Germany). The second arc deposition (CAE 2) was performed using an AIP-S20 PVD coater (Kobelco, Kobe, Japan) at McMaster University.

The first set of parameters tested for CAE 2 was based on previous depositions used for turning inserts. However, the sharp edges of the endmills were damaged during the deposition, probably due to excessive ion bombardment, which generated excessive compressive residual stress in the coating and additionally caused resputtering [29], as the electric field tends to concentrate on the sharp edges. To reduce the excessive ion bombardment, a higher pressure was used (from 4 to 10 Pa). When fast plasma ions collide with slower gas ions, energy and momentum are transferred to the gas. Consequently, it is expected that increasing the gas pressure will result in a reduced ion energy distribution [35].

However, even after the deposition with a higher pressure (10 Pa), the corner coverage was not adequate. Then, pulsed bias arcing was used. The bias voltage was then pulsed at a frequency of 10 kHz with a duty cycle of 25%. With this change, an adequate coverage of the sharp edges without coating delamination was obtained. The deposition parameters are shown in

Table 1. Deposition parameters.

	Coatings	CAE 1	CAE 2	HiPIMS
Etching	Bias [V]	−200	−400	−200
	Pressure [Pa]	0.25	1.33	0.35
	Temperature [°C]	450	450	650
	Gas	Ar	Ar	Ar
Deposition	Target power [kW/cm2]	—	—	10
	Arc current [A]	150	150	—
	Bias [V]	−50	−70	−60
	Pulse frequency [kHz]	—	10	4
	Pulsed width [μs]	—	25	60
	Duty cycle [%]	—	25	—
	Deposition pressure [Pa]	4	10	0.6
	Temperature [°C]	530	450	650
	Gas	N2	N2	N2
	Number of targets	4	2	3
	Table rotation [RPM]	1.3	5	1
	Deposition time [min]	80	24.5	65
	Target composition	Al60Ti40	Al60Ti40	Al60Ti40

.

The deposition time was selected to obtain a coating thickness of around 3 µm. Each coating was deposited in three endmills and three coupons, so each sample had at least two replicates.

The coating thickness and images of the polished coupons’ cross-section and top surface were obtained using an Environmental Scanning Electron Microscope (ESEM) equipped with a field emission gun (FEG) (Model Quattro, Thermo Fisher Scientific, Waltham, MA, USA). The elemental composition of the specimens was measured using an Energy-Dispersive X-ray Spectroscopy (EDS) system in conjunction with the SEM. After the cross-section analysis, the coupons’ surfaces were assessed using EDS to check the coatings’ characteristics and composition.

The crystal structure and preferred orientation of the deposited coatings were determined by analyzing two-dimensional diffraction patterns collected using a Bruker D8 ADVANCE diffractometer (Bruker, Billerica, MA, USA) with Co Kα radiation (λ = 1.79026 Å). The source was operated with a voltage of 35 kV and 45 mA and a step size of 0.02° in the 2θ range of 32° to 105°. To determine the preferred orientation of the TiAlN coatings, the texture coefficient (Tc_(hkl)) was calculated for each orientation using Equation (1) [38]:

$${Tc}_{\left(hkl\right)}=\frac{I_{\left(hkl\right)}/ I_{0\left(hkl\right)}}{\frac{1}{n}\sum _1^n\frac{I_{hkl}}{ I_{0\left(hkl\right)}}}$$

(1)

where Tc_(hkl) is the texture coefficient for a specific plane, I_(hkl) is the diffraction peak intensity in the sample, and I_0(hkl) is the intensity measured in a powder sample obtained from the ICCD (the International Center for Diffraction Data, Newtown Square, PA, USA) datasheet.

To evaluate the micro-mechanical properties of the coatings, an Anton Paar NHT3 nano-indentation tester (Anton Paar, Graz, Austria) was used on the polished coupons. The indentation was performed using a Berkovich diamond indenter. The load was applied linearly with a loading and unloading rate of 80 mN/min, with 2 s of dwell time, until a maximum load of 40 mN. The maximum load was selected, making sure that the indentation depth obtained was lower than 1/10 of the coating thickness to avoid the influence of the substrate. In total, 30 indentations were performed per coating.

Scratch tests were performed on the surface of the coated polished coupons using an RST3 Revetest Scratch Tester (Anton Paar, Graz, Austria). It used a Rockwell diamond indenter with a 200 μm radius. The first part of the test was a pre-scan of the surface to identify its topography. After that, the scratch was performed with a linear progressive rate of 299 N/min from 0.5 N to 150 N. The length of the scratch was maintained in 3 mm with a scan speed of 6 mm/min. After the scratch, the surface was post-scanned. Three scratches were performed in all samples tested.

The coatings’ toughness was assessed using a Palmqvist toughness method with a DuraScan 50 hardness tester. The measurements were performed according to the ISO 28079 standard [39]. The load used for the tests was 20 N, and three indentations were performed for all the coupons tested. The Palmqvist toughness was calculated by dividing the applied load by the total sum of crack lengths from the crack tip to each corner of the indentation.

2.2. Machining Tests, Tool Wear Assessment, and Chip Formation Analysis

Milling tests were performed in austenitic stainless steel 316L using a solid cemented carbide 3CH0400MS012A endmill with a 4 mm cutter diameter. Cutting tests were performed in a high-speed machining center Matsuura LX-1. The cutting parameters used are shown in

Table 2. Cutting parameters.

vc [m/min]—Cutting speed	150
fz [mm]—Feed per tooth	0.025
ae [mm]—Radial depth of cut	3.2
ap [mm]—Axial depth of cut	0.2
Cutting condition	Dry
Workpiece material	SS316L (79 HRB)

. The parameters were based on Kennametal’s Master Catalogue 2023. The feed per tooth and radial and axial depth of the cut were used within the recommended limits. However, the cutting speed tested was higher than that suggested by the catalogue (up to 110 m/min). For this work, 150 m/min was used in order to exacerbate the wear mechanisms, which are characteristic of high-speed conditions.

A dynamometer Kistler 9254 connected to an amplifier LabAmp 5167A attached to the workpiece was used to measure the cutting forces during all the passes for the three coatings tested. The sampling rate used was 10,000 Hz, and forces were collected in feed F_x, thrust F_y, and axial F_z directions, as shown in Figure 1. The force data were processed using DynoWare 2825A-043 software (Version 3.1.0.0).

Optical images of the endmill tool wear were obtained using a Keyence VHX-950F digital optical microscope (Keyence, Itasca, IL, USA). Images of the flank and rake surface of each edge were taken. The tool life criteria used was 300 μm of maximum flank wear (V_Bmax) averaged over all three teeth of the tool, according to the ISO 8688-2 standard [40].

To evaluate the tool wear mechanism, images of the tool after its end of life were obtained using a JEOL 6610LV SEM coupled with an Oxford Instruments EDS system (JEOL, Tokyo, Japan). An Alicona Infinite Focus G5 microscope (Alicona Manufacturing Inc., Bartlett, IL, USA) with a magnification of 10× was used to obtain images of the adhered material on the tool surface and the volumetric wear present in the tool. Chips were collected after the first pass at 0.1 min, 3 min, 20 min, and 120 min of machining to evaluate the tribological conditions during the milling process. The JEOL 6610LV SEM was used to obtain images of the shear band and undersurface of the chips.

The surface integrity of the workpiece was obtained in terms of average surface roughness using a Mitutoyo S201J Profilometer (Mitutoyo America Corporation, Aurora, IL, USA). The cut-off used was 0.8 mm according to ISO 13565-1 [41]; nine measurements of the surface were taken after every 47 passes, which corresponds to the entire surface of the workpiece.

3. Results and Discussion

3.1. Coating Characterization

Comprehensive investigations were carried out to evaluate the structure and the micromechanical properties of the coatings. Figure 2 shows the SEM images of the coupon’s top surface (a) and cross-section (b).

The HiPIMS coating shows a well-defined columnar microstructure characterized by a cauliflower-like surface morphology [42] that is not visible in both arc depositions (Figure 2a). However, it is possible to observe a higher number of droplets/macroparticles for CAE 1 and especially CAE 2, which are characteristic of Arc deposition systems [32].

Evaluating the coupons’ cross-section using a higher magnification (Figure 3) confirms the presence of a columnar structure of the HiPIMS coating, and CAE 1 presents a more densely compacted structure.

CAE 2 presents a structure with a lower density and higher thickness. This was expected, since the pulsed bias with a reduced duty cycle (25%) reduces the overall average energy supplied during the deposition, which can lead to a less dense coating [36,37]. This is also reflected in its higher thickness [34]. The measured thicknesses are 2.9 μm (HiPIMS), 2.6 μm (CAE 1), and 3.5μm (CAE 1). The width is within ±0.5 μm from the target thickness.

EDS analysis (Figure 4) was performed to determine the composition of the three coatings. The targets used for all of them had the same composition, Al 60%at. and Ti 40%at.

Table 3. Chemical composition of the coupons at. %.

Coating	HiPIMS	CAE 1	CAE 2
Al (at. %)	60.7	55.9	56.2
Ti (at. %)	39.3	44.1	43.8

shows the composition of the coating obtained by EDS on the top surface of the coupons.

The content of Al in the CAE coatings is lower than that in the target composition, which can be explained by the fact that the Al ions tend to spread more inside the chamber since they are lighter. Depending on the positioning of the samples, they can be positioned in a way that preferentially experiences the Ti deposition. Nevertheless, all the compositions are within the FCC B1 NaCl-type crystal structure range for AlTiN.

Table 4. Micromechanical properties of the deposited coatings.

Coating	HiPIMS	CAE 1	CAE 2
Hardness (GPa)	33.6 ± 4.1	29.9 ± 3.1	26.6 ± 4.9
Elastic modulus (GPa)	486 ± 54	481 ± 48	501 ± 71
H/E	0.069	0.062	0.053
H3/E2	0.157	0.116	0.075
Plasticity index	0.459	0.532	0.544
Palmqvist toughness [N/μm]	1.04	1.47	1.54

shows the coatings’ micromechanical properties. The CAE 2 coating presented the lowest hardness levels, which was already expected, since with the use of pulsed bias in the conditions used, the overall energy of deposition was likely reduced. Therefore, a coating with a lower density and lower hardness was formed.

The elastic modulus of all tested coatings proved to be similar, which was already expected. This measures the resistance to the separation of adjacent atoms, reflecting the strength of interatomic bonds and corresponding to the slope of the interatomic force–separation curve at the equilibrium spacing. Unlike tensile and yield strengths, which can be significantly affected by previous deformations, impurities, and heat treatments, the elastic modulus is relatively unaffected by these factors [43]. However, it has proven to be sensitive to the Al/Ti ratio, decreasing rapidly with the formation of the AlN hexagonal phase [27,28]. Additionally, significant differences in compressive stresses can also impact the elastic modulus [44].

In general, in situations where the composition is approximately constant, and without the formation of the hexagonal phase, one may expect that the values of the elastic modulus should be similar for the AlTiN coatings, as observed in this study and also reported in [45].

On the other hand, the HiPIMS coating presented the highest hardness. Higher levels of hardness have been commonly related to improved machining performance. However, it is now known that hardness is not the only important factor, especially when machining materials with a high tendency of adhesion, as in the case of this work, with the austenitic stainless steel 316L. For milling operations (interrupted cutting), toughness is also a very important factor.

The H/E and H³/E² ratios have been used as a proxy for coating fracture resistance. H/E denotes the elastic strain to failure, and H³/E² represents the resistance to plastic deformation. Coatings with a higher H³/E² can present higher resistance to crack initiation, which has been used as an indirect indicator of their load-carrying capacity and toughness [46]. A higher H/E ratio suggests a greater capacity to withstand mechanical wear and failure in tribo-contact scenarios, as the contact area stays elastic under increased stress during external impacts [44].

Another important indicator is the plasticity index (PI), which is determined by the percentage of plastic work related to the total work during indentation and presents an inverse relation with the H/E_r ratio, where E_r is the reduced modulus [46]. A higher plasticity index means that the coating presented higher plastic deformation during indentation.

In situations where sliding/abrasion is dominant, a high H/E (lower PI) is generally beneficial. However, in interrupted operations, such as milling, especially with adhesive-dominant wear, this trend happens in the opposite way. According to [47], a reserve of plasticity to dissipate the energy generated during friction is necessary, and a higher PI correlates with greater energy dissipation under loading conditions. Results in milling showed a correlation between higher PI (from 0.51 to 0.56) and a longer tool life when machining steel. The authors estimated an optimum PI in their tests of around 0.53 for end/face milling [47]. The results of [48] also corroborate this trend during turning operations with an adhesive-dominant wear. As this present work is related to the milling of 316L stainless steel, which is characterized by its high tendency of adhesion, it is expected that similar trends will be observed as the ones obtained in [47,48].

With the highest hardness for the HiPIMS coating and a similar elastic modulus for the three coatings, the HiPIMS coating presents the highest values of H/E and H³/E². The HiPIMS plasticity index is the lowest, which is also expected for a material with higher hardness and similar elastic modulus. Relating this data to the previous discussion on the milling of materials with a high tendency of adhesion, one may expect that the HiPIMS coating will present the worst performance in terms of tool life.

It is important to be aware that the plasticity index and the nanoindentation curve do not directly indicate a material’s fracture toughness or its resistance to fatigue fracture. To have an indicator of the fracture toughness of the coating, a Palmqvist toughness test with a Vickers indenter was performed. It is important to mention that the penetration depth for this case is higher than the coating thickness. Therefore, there will be an inherent influence of the substrate. As the substrate is the same for all the coatings tested, the results should be comparable.

In the Palmqvist toughness measurement, it is assumed that there are no residual stresses in the sample. These residual stresses may influence the apparent surface crack length, which serves as an indicator not only of the material’s fracture toughness but also of the existing residual stresses. Compressive residual stresses tend to reduce the surface crack length compared to its length in the absence of such stresses, whereas tensile stresses have the opposite effect [49].

In the tested samples, variations in residual stress may exist both between different samples and within the same sample, influenced by distinct deposition techniques (CAE and HiPIMS) and deposition parameters. Attempting to mitigate residual stress variation by annealing the samples could introduce additional sources of error and lead to microstructure alterations that differ from those occurring during machining, such as spinodal decomposition [27].

Therefore, after these considerations, it is important to emphasize that this indicator is qualitative in this study and influenced by the Palmqvist toughness of the material and the residual stress. Both factors are important for machining performance.

Both arc coatings (CAE 1 and 2) presented similar levels of Palmqvist toughness, with slightly better results for CAE 2 (1.54 N/μm) than for CAE 1 (1.47 N/μm). The HiPIMS coating presents the lowest value (1.04 N/μm), indicating a potential inferior performance in interrupted cutting operations.

The scratch test (Figure 5) shows a different trend than those obtained in nanoindentation, with the HiPIMS coating presenting the best results. The adhesive failure/spalling (L_c2) of the coating occurred in the early stages for CAE 1 (L_c2 = 74.7 N) and CAE 2 (L_c2 = 68.6 N), and for HiPIMS, the initial coating spallation is observed with higher loads (L_c2 = 113.2 N). The acoustic emission also shows that the cracking started first for the CAE 2 coating, followed by CAE 1, and lastly, HiPIMS.

However, it is important to mention that due to the high hardness of the HiPIMS coating, the penetration depth was lower under the same loads, which caused lower levels of deformation on the coating. In addition, the scratch test is greatly influenced by the substrate and its adhesion with the coating layer, being mostly used to evaluate the adhesion strength of hard coatings in a qualitative way [50].

Figure 6 shows the XRD patterns of the three deposited coatings. The phase composition shows that the coatings have an FCC B1 NaCl-type crystal structure. Both arc depositions show a preferential orientation (200), and the HiPIMS coating has a preferential orientation (111).

Table 5. Texture coefficient for the deposited coatings.

Texture Coefficient
hkl	HiPIMS	CAE 1	CAE 2
(111)	1.89	1.53	0.11
(200)	0.11	2.47	2.89

shows the texture coefficients (Tc) for the deposited films. The crystallographic texture of the films was determined based on the degree of the planes’ preferential orientation.

It is important to mention that as the thickness is different from the coatings, varying from 2.6 to 3.5 μm from the thinnest to the thickest coating, the comparison in terms of absolute values could be affected. An alternative to avoid the effect of different thicknesses would be to analyze the percentage of each texture coefficient for the same coating, as shown in

Table 6. Texture coefficient for the deposited coatings in terms of percentages.

Texture Coefficient
hkl	HiPIMS	CAE 1	CAE 2
(111)	94.5%	38.2%	3.67%
(200)	5.5%	61.8%	96.33%

.

It is known that for TiN, (200) has the lowest surface energy. Calculations show that the surface energy of the (111) surface is about four to five times higher than that of the (200) surface [51]. This could influence operations where adhesion is the main wear mechanism, since a higher surface energy can facilitate the adhesion of the workpiece material. In addition, the HiPIMS coating has well-defined columnar structures, which may have higher surface-free energy (grains with polygonal pointed ends and greater surface area). As the coating becomes more columnar and porous, the adhesive friction component is expected to have a greater contribution [52].

3.2. Tool Wear Analysis

The tool performance with the three different coatings was investigated under dry end milling operation. The tool life criteria included 300 μm of flank wear maximum (V_Bmax) averaged over all three teeth of the tool, according to the ISO 8688-2 standard [40]. The flank wear curves are shown in Figure 7.

The flank wear results show that wear rate and tool life are significantly different for the coatings tested. Even though they have the same composition, the coatings’ micromechanical properties and microstructure are very different, as discussed in the last section.

Both arc-deposited coatings (CAE 1 and 2) are similar in terms of average tool life: CAE 1 = 127 ± 1.6 min and CAE 2 = 128 ± 24.8 min. The HiPIMS coating presented a significantly lower tool life of 21.2 ± 2.9 min. This follows the trend already observed by [47,48]: in cases with a high tendency of adhesion, coatings with high PI and lower H/E and H³/E² present a longer tool life, with this being related to a higher dissipation of the energy generated during friction. The HiPIMS coatings presented the exact opposite trend, which may contribute to HiPIMS’ significantly lower tool life.

Interestingly, the CAE 2 coating, which had the lowest hardness (due to the pulsed bias deposition), presented a tool life comparable to CAE 1, indicating that sensitivity to abrasive wear is not the dominant cause of tool failure. However, it is important to notice the higher deviation for the CAE 2 deposition in terms of tool life, which was 24.8 min considering the three replicates, whereas CAE 1 lasted only 1.6 min, which is indicative of a more repeatable process. This behavior is related to the much lower tool wear obtained in the first passes for the CAE 1 coating. Having a reduced and more stable initial tool wear is important for a more controlled wear progression until the end of the tool’s life. This is also indicative of lower adhesion levels, since adhesion and BUE formation are related to sudden increases in tool wear due to an increase in cutting forces and sudden removal of material from BUE formation, potentially resulting in tool chipping [53]. Depending on how the BUE breaks, it can remove larger particles of the tool, altering its geometry, exposing more substrate and further increasing adhesion levels. Analyzing the curve with the shortest tool life for CAE 2, the change in the slope of the curve is clearly visible after 50 min, which indicates a significant change in the geometry of the tool, favoring adhesion. This is not observed for the case with the highest tool life for CAE 2, possibly due to more gradual tool wear. In milling operations, particularly at high speeds, process variance tends to be higher, given the numerous impacts per second (in this case, 597 impacts per second—tooth pass frequency).

For CAE 1, wear progression was more linear, with reduced adhesion, evidenced by chip analysis. Therefore, minimizing adhesion levels is essential for establishing a stable process, especially in interrupted cutting operations. The HiPIMS tool had a much shorter tool life. It was characterized by intense initial tool wear followed by an unstable zone with rapidly increasing tool wear, which resulted in a tool life more than six times lower than both arc depositions on average. Figure 8 shows optical images presenting the progression of flank wear.

The adhesion starts much earlier for the HiPIMS tools, with the first intense signs of BUE after around 7 min of machining. This ultimately leads to the chipping of the tool at the end of its life. Figure 9 shows images taken with the Alicona near the final stages of tool life, between 70% and 80% of the tool life (20 min for HiPIMS, 100 min for CAE 1 and 2) of flank (Figure 9a) and rake (Figure 9c) surfaces, as well as after the last pass of flank (Figure 9b) and rake (Figure 9d) surfaces, in order to depict the wear mechanism. It is visible that the HiPIMS coating presents higher levels of adhesion. After 20 min of machining, the BUE is already intense, leading to the chipping of the tool at 24 min. The arc-deposited coatings, even after 100 min of tool life, did not present significant visible levels of adhesion.

As already discussed, the main wear mechanism for all tools was the intense adhesion of workpiece material on the flank and rake surface of the tool, which can be better observed in the SEM images (Figure 10) of the tool at the end of its life. The rough appearance and signs of particle subtraction from the tool face, allied with high levels of adhered material (Fe and Cr), are characteristic of predominant adhesion.

The HiPIMS tools presented chipping caused by previous BUE, showing high substrate exposure and extensive coating delamination on the flank surface of the tool. Even after the chipping, there are visible signs of adhered workpiece material (Fe and Cr) on the rake and flank surfaces.

The arc coatings also present significant adhesion but were manifested only in later stages of wear, prolonging the tool’s life. Tool wear increases in a more controlled way. Even after 120 min of machining, catastrophic failure (chipping) did not occur. The wear mechanism was similar for both arc depositions.

The greatest impact on tool life was the level of adhesion of the workpiece material onto the tool surface. This was also reflected in the force measurements; excessive adhesion caused an increase in cutting forces that tend to accelerate wear due to higher levels of deformation and heat generation. It is noticeable that forces start to become higher much earlier for the HiPIMS tool (Figure 11). Figure 12 shows the force signal after 7 min of machining; this was the moment when the adhesion and BUE were noticeable on the HiPIMS coating. This is reflected not only in the higher level of forces when compared to the arc deposition but also in a higher variation in the forces, as highlighted in Figure 12.

The adhesion and increased deformation of workpiece material during machining, combined with the accelerated tool wear, led to higher levels of workpiece surface roughness when using the HiPIMS tool (Figure 13). Though the trend tends to stabilize within lower levels of surface roughness for the arc coatings (CAE 1 and 2), there is a positive trend for the HiPIMS coating until the tool fails.

The adhesive wear that is accelerated in the HiPIMS tool is the main wear mechanism for this operation. Despite the constant effort to obtain coatings with higher hardness and resistance against abrasion, there are some cases where hardness is not the most important parameter. During the machining of material with a high tendency of adhesion, the coating must sustain intensive sticking without peeling off and provide a better tribological condition. In this case, the PI and factors like H/E and H³/E² showed a higher contribution than the hardness itself. Even with lower hardness, CAE 2 presented comparable results with CAE 1 in terms of tool life.

In addition, the levels of adhesion that the HiPIMS tool experienced were higher and occurred much earlier, which was the determinant for its significantly worse performance. This will be further explored in the chip analysis section. The higher levels of adhesion could be attributed to the columnar structure of the HiPIMS tool and the preferential texture of the grains, which present higher surface energy; this effect could be intensified in high temperatures, leading to higher levels of adhesion.

It is also important to highlight that the proposed personalized pulsed arc deposition (CAE 2) achieved results comparable to the commercial arc coating (CAE 1). With this technique, it is possible to perform depositions with lower substrate temperature [33] and a higher deposition rate [34] and to relieve stress introduction during deposition [35]. In cases with excessive compressive residual stress or delamination due to poor corner coverage, this technique could serve as an alternative to constant bias arc deposition. A future study focused on optimizing the parameters of pulsed bias deposition, such as duty cycle and frequency, could further increase the efficiency.

3.3. Chip Analysis

Understanding the interaction of the tool–coating–workpiece interface is a very complex task due to the high loads and temperatures involved in machining processes. An indicator that can be used to determine how efficient the tribological conditions are during cutting is based on analyzing the chips generated during machining [53]. SEM images of the shear band and undersurface of the chips were taken during different cutting times, from the first pass to near to the end of the tool life. Figure 14 shows the chips generated after the first pass and at 3 min, 20 min, and 120 min (CAE 1 and 2).

With the SEM images, it is possible to observe that the chips generated by the HiPIMS coating present higher levels of deformation during the first pass and within 3 and 20 min. In general, they are more densely packed than those of both CAE 1 and 2 and also have a more irregular undersurface.

Chips produced by CAE 1 and 2 had a relatively smooth morphology compared with HiPIMS. This means that the chips flowed smoothly over the tool–chip interface, generating lower friction and adhesion, whereas chips produced by the other coating compositions tended to stick to the tool. This relates to the higher levels of force and adhesion observed for the HiPIMS tool.

During the first pass, there are no significant geometrical changes in the tool geometry; this is the ideal condition to isolate the coating’s effect from the geometry for evaluating the tribological conditions. In Figure 14, for both arc coatings (CAE 1 and 2), it is possible to observe much lower levels of deformation in both shear bands, with a higher spacing between the chip serrations. Also, the undersurface presented a smoother surface.

As tool wear progresses, all the coatings follow the same trend, with higher levels of deformation at the end of the tool’s life. However, for CAE 1 and 2, this process is delayed, happening in much later stages of tool life, which causes the tool life to be more than six times higher.

Progressing to 3 min after machining (Figure 15), CAE 2 chips start to resemble those generated by HiPIMS during the first pass; this indicates a higher compression of the chips, showing that the tribological condition deteriorates when compared to the first pass. However, the undersurface of the chips still shows a smooth surface. The HiPIMS chips are even more deformed than when compared to the first pass, with a very compact shear band and sticking marks apparent on the undersurface, confirming that the friction conditions between the tool and workpiece are becoming worse. The CAE 1 coating, however, still shows very similar behavior to the first pass, with a higher spacing on the shear band and a smooth surface on the undersurface. This indicates that the coating is still effectively acting as a protective layer against adhesion [54], which correlates well with the reduced flank wear when compared to the other two coatings. This also explains why the tool life had a higher variation for the CAE 2 coating. With higher initial levels of flank wear, the tribological conditions started to become worse earlier, which caused a higher unpredictability for this tool.

Within 20 min of machining (Figure 16), the HiPIMS-coated tool is almost at the end of its life, with high levels of adhesion and BUE present. Chips present a very high deformation. The curvature of the chip perpendicular to the shear band direction can be observed to be close to zero. The shear band is inconsistent and highly compacted, indicating higher levels of friction, which are reflected in the increase in cutting forces [26]. The undersurface is also uneven and presents deep scratches, corroborating its poor results for tool life. At this moment (20 min), chips generated by CAE 1 and CAE 2 coatings are very similar to the chips produced by HiPIMS at the first pass and at 3 min, respectively. This again shows a better performance for the CAE 1 coating, followed by CAE 2. It is also important to note when comparing the undersurface images that the curvature of the chips is much higher with both CAE 1 and 2 than with HiPIMS, which indicates that the chips produced by the arc-deposited coatings had a higher sliding velocity, indicating a better tribological condition.

Reaching 120 min of machining (Figure 17), close to the end of both arc coatings, the chips have a very similar pattern as that presented by the HiPIMS coating at 20 min, a flat chip with a very compacted shear band (slightly less compact for CAE 1), indicating high levels of deformation and friction. The undersurface, on the other hand, presents a smoother surface when compared to HiPIMS; there are some sticking marks, but the overall condition is better than that found for HiPIMS.

The chip analysis confirms the superior performance of the CAE 1 coating during machining; the coating was able to effectively and consistently maintain favorable tribological conditions, facilitating the cutting process and reducing forces and tool wear. The CAE 2 coating also presents good results but with slightly more deformation than CAE 1, which relates to a higher deviation in tool wear. This coating (CAE 2) is effective in reducing friction. However, elevated wear in the initial stages causes substrate exposure, hinders the tool’s performance, and leads to higher adhesion and more inconsistent behavior.

This analysis also confirms the HiPIMS tool’s worst performance during machining; the high initial tool wear allied with much higher levels of adhesion and BUE generated higher forces, accelerating tool wear and leading to premature failure.

4. Conclusions

Intense adhesion and BUE formation were the main wear mechanisms during the dry end milling of a SS316L alloy. The proposed pulsed bias arc deposition (CAE 2) achieved comparable tool life results to the commercial arc deposition (CAE 1), opening opportunities for further optimization of pulsed bias arc deposition parameters. In contrast, the HiPIMS-coated tool exhibited a significantly reduced tool life, which was six times lower than the CAE tools.

The adhesive wear was more pronounced and accelerated for the HiPIMS coating. Through the chip analysis, it was possible to observe that the HiPIMS presented worse tribological properties, with higher deformation of the shear band and undersurface of the chips for all passes measured. The results are well correlated with the higher forces and worse surface roughness generated by this coating. The higher levels of adhesion could be attributed to the columnar structure of the HiPIMS coating and the preferential texture (111), which presents a higher surface energy compared to the (200) from both arc depositions. Further studies focusing on how the texture influences the tribology of the coatings are still necessary to confirm this hypothesis in different scenarios.

The micromechanical analysis shows that the HiPIMS coating presented the highest level of hardness and the lowest level of plasticity index. For interrupted cutting machining where adhesion is the main wear mechanism, a reserve of plasticity is beneficial to dissipate the energy generated during friction, even if this is related to lower hardness levels. Ultimately, this also contributed to a reduced tool life. As abrasion was not the main wear mechanism, the lower hardness of both CAE 1 and CAE 2 did not negatively influence respective tool lives.

Therefore, in dry milling operations of SS316L alloy, the reduction in adhesion plays a fundamental role in increasing tool life. In this case, the hardness of the coating is not as important as its capacity to provide better friction conditions and resist adhesive tool wear (higher plasticity index). Given their favorable combination of tribological and micromechanical properties, both arc coatings presented significantly better results in terms of tool life and surface finish.