Enhancing tool life and surface quality in high-speed milling of Ti6Al4V: a comparative study of flood and minimum quantity lubrication cooling strategies

The Ti6Al4V workpiece and the uncoated carbide tool used in this research are shown in Fig. 1. The workpiece for the milling tests is shown in Fig. 1a as a 100 mm × 70 mm × 70 mm cuboid block of Ti6Al4V. Figure 1b illustrates the uncoated carbide tool XOEX10T308FR-E05 H15. The product details are presented in Table 1. Figure 1c-d depicts the tool holder used in this study. The product details of the tool holder R217.69–1616.3–10-2A are tabulated in Table 2.

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The CNC machining center used in this study is the DMG 50 CNC machine, as illustrated in Fig. 2a. The specifications for the DMG 50 CNC machine are presented in Table 3. Figure 2b illustrates force measurements collected using the Kistler Type 9443B dynamometer. Figure 2c depicts the ATAGO N-20E flood coolant concentration testing equipment. Figure 2d shows that the Mitutoyo SJ-500 roughness analyzing device was utilized to evaluate the machined workpiece’s surface quality. The XOPTRON X80 microscope was utilized to assess tool wear, as seen in Fig. 2e. The ISO 8688–2 standard serves as the foundation for the tool life criterion.

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First, the workpiece was firmly placed on the DMG 50 CNC machining center's dynamometer (Fig. 2a and b). Before the milling operation started, the machine tool’s datum point was established using a dial indicator since precise alignment is necessary for reliable output. In order to identify the datum’s z-axis, the dial indicator first touched the workpiece’s top surface and applied pressure by progressively advancing downward until the needle rotated 720°. Then, a similar process was applied to identify the x and y axes. A final datum point is established at the end of the procedure. After these procedures, the CNC machine begins to programme based on these reference points. For every set of testing settings, a new set of inserts was utilized to ensure that the milling tools were all at the same level of wear over the course of the several tests. To ensure uninterrupted operation and prevent outside influences from affecting the finished product, the workpiece, cutting tool, and force monitoring system were all continuously monitored during the milling process. All trials were carried out under flood and MQL cooling techniques for this investigation. The Ti6Al4V cutting experimental parameters used in this investigation are shown in Table 4. The present study used a 15% concentration of commercially available emulsifiable water-based cutting fluid as the flood coolant. It was delivered via a fluid line with a suction pump. Using vigorous manual stirring, the mixture was mixed to obtain good oil-in-water emulsification. There were no additional additives introduced.

For MQL coolant setting, a nozzle connected to a solenoid pump with a slow flow rate via a mixing chamber makes up an MQL configuration. The mixing chamber is also connected to a compressor’s outlet. Compressed air combines with the coconut oil from the pump to produce an atomized combination of suspended lubricant particles in the air inside the mixing chamber, where the incoming air stream splits the coconut oil. A nozzle positioned 30 mm away from the cutting zone is then used to spray the oil-air stream onto it. The MQL fluid uses coconut oil since it is biodegradable and has lubricating qualities. The friction coefficient of coconut oil makes it an excellent lubricant [37]. Compared to other vegetable-based cutting fluids, it provides better oxidative and thermal stability, improves lubrication, and decreases waste. At low and medium cutting speeds, it enhances tool life and surface finish [38]. Furthermore, coconut oil is completely biodegradable, safe for workers, and has no disposal problems. Coconut oil was chosen for the MQL method in this investigation due to its availability, lubricating capacity, and physical characteristics. Its cutting temperature, cutting force, surface roughness, tool-chip contact length, and tool wear were all lower than those of mineral oil [39, 40]. Continuous coolant supply enhances lubrication and temperature management throughout the machining process. An air compressor provides compressed air at a pressure of five bars. To stop cutting fluid from getting into the air tank, the arrangement has a manual and an air-piloted valve near the compressor’s outlet. A pump is used to supply the cutting solution, and its measured flow rate is 40 ml/h.

Figure 3 depicts the variation in resultant cutting forces in relation to machined length at various spindle speeds and feed rates during the HSM of Ti6Al4V with a flood cooling strategy. As illustrated in Fig. 3a, with a spindle speed of 1194 r/min, the resultant cutting force rises with the machined length, regardless of the feed rate. The resultant cutting forces were obtained in the ranges of 475–632 N, 661–801 N, and 1008–1345 N for feed rates of 358 mm/min, 538 mm/min, and 717 mm/min, respectively. When the spindle speed is raised to 1791 r/min, a notable decrease in the resultant cutting forces occurs, with values ranging from 357–456 N for the 358 mm/min feed rate, 486–712 N for the 538 mm/min feed rate, and 751–914 N for the 717 mm/min feed rate. At 358 mm/min, the rise in resultant cutting force along the machined length at a spindle speed of 1791 r/min is less pronounced compared to the behavior observed at 1194 r/min. At a spindle speed of 2389 r/min, as depicted in Fig. 3c, the resultant cutting force ranges from 342 to 392 N for a feed rate of 358 mm/min. The resultant cutting forces for the 538 mm/min feed rate range from 485–715 N, while for the 717 mm/min feed rate, they range from 682–1110 N. A significant increase in resultant cutting force for the 717 mm/min feed rate is observed after a machined length of 2000 mm at both 1791 r/min and 2389 r/min spindle speeds. The resultant cutting force generated at a feed rate of 717 mm/min showed a significant increment with a higher spindle speed. Specifically, while machining a 3000 mm length at a spindle speed of 2389 r/min versus 1791 r/min, the cutting force was observed to be 16% greater. This discrepancy was even more obvious at a machined length of 4000 mm, where the cutting force exceeded the lower spindle speed by 22%, suggesting that increased spindle speeds can result in stronger cutting forces, particularly over longer machining lengths. Therefore, it is recommended to use a moderate spindle speed of 1791 r/min when employing a higher feed rate of 717 mm/min.

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Altering spindle speed and feed rate affects the resultant cutting forces significantly in HSM operations. Generally, as spindle speed rises, the resultant cutting forces tend to decrease, irrespective of the feed rates used. The reduction in resultant cutting force with high spindle speeds aligns with reduced chip thickness and thermal softening effects. This decrease in cutting forces is primarily owning to less friction between the rake face of the tool and the chip; higher cutting speeds reduce both sticking and sliding friction at the interface between the tool and chip [41]. Additionally, elevated spindle speeds result in increased cutting temperatures, which can lead to thermal softening of the workpiece material. This thermal softening effect allows chips to form serrated shapes and contributes to a decline in cutting forces [42‐45]. The drop in cutting forces can also be explained by a reduction in chip thickness [46]. As the spindle speed rises, the combination of reduced chip thickness and lowered material flow stress, resulting from thermal softening, further contributes to the overall decrease in cutting forces [47]. As cutting temperatures continue to rise, materials experience softening, which again results in reduced cutting forces [48]. When strain hardening and heat softening are balanced, the cutting force remains roughly constant. However, when strain hardening effects dominate over thermal softening, a higher cutting force is required to overcome material resistance. Therefore, when thermal softening and reduced chip thickness prevail, a decrease in cutting force is often observed [49].

Lower feed rates, such as 358 and 538 mm/min, produce relatively constant cutting forces along the entire length. In contrast, with a higher feed rate of 717 mm/min, the resultant cutting force rose as both the machined length and spindle speed increased. Generally, higher feed rates lead to higher cutting forces [50, 51]. At elevated feed rates, tools are more prone to vibrations, which can further elevate cutting forces. Additionally, increased feed rates can lead to machine instability and tool vibrations, which are affected by cutting speed, cutting depth, and cutting forces [52]. Machine tool vibrations are classified into three types: free, forced, and self-induced [52, 53]. Tool wear is another element that facilitates the steady rise in cutting force as the cutting length increases [54, 55]. Pratap et al. [56] indicate that accelerated tool wear rates are the primary cause of increased cutting forces.

Figure 4 depicts how the cutting forces vary with machined length, considering various spindle speeds and feed rates during high-speed Ti6Al4V machining using an MQL cooling strategy. In Fig. 4a, at a spindle speed of 1194 r/min, the resultant cutting force exhibited a slight decrease from 0 to 2000 mm of machined length for feed rates of 358 mm/min and from 0 to 1000 mm of machined length for feed rates of 538 mm/min. Subsequently, for the 538 mm/min feed rate after the 1000 mm machined length, the resultant cutting force increased gradually (remaining relatively constant) as the machined length increased. In contrast, for the 717 mm/min feed rate, the cutting force rose from 824 to 1017 N. At a spindle speed of 1791 r/min, as seen in Fig. 4b, the average resultant cutting forces along the machined length were 395 N for the 358 mm/min feed rate and 528 N for the 538 mm/min feed rate. Notably, the cutting force for the 717 mm/min feed rate remained stable, ranging from 637 to 685 N until reaching a machined length of 6000 mm, at which point it experienced a sudden increase to 748 N at 7000 mm. In Fig. 4c, the resultant cutting forces for the 358 mm/min feed rate remained constant over the machined length, whereas the forces for the 538 mm/min and 717 mm/min feed rates fluctuated along the machined length, with average cutting forces of 466 N and 518 N, respectively.

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One scenario that occurred during Ti6Al4V milling under the MQL approach is that the resultant cutting forces mostly fluctuated along the machined length, particularly at a high feed rate. Kiran Sagar et al. [57] discovered that the primary factor influencing cutting force is the feed rate. The observed variations have been ascribed to the mechanics involved in chip formation and the interaction between the tool and the workpiece during the cutting process. An elevation in feed per tooth leads to the development of thicker chips, requiring more force to shear the workpiece effectively. This added force heightens frictional forces between the workpiece and the tool, producing excess heat and possibly resulting in inferior surface roughness. Furthermore, the heightened forces may contribute to greater deflection and tool wear, which may ultimately reduce the milling operation efficiency. The variability of the resultant cutting force is generated by adiabatic shear and a change in cutting direction owing to vibration [58]. Larger feed rates frequently led to larger cutting forces, demonstrating the importance of careful feed-per-tooth optimization for efficient milling operations [59]. The productivity and surface quality of machined components can be reduced by variations in cutting force and vibration, which can also interfere with regular cutting processes and create unwanted chatter marks on machined surfaces [48]. Additionally, the friction between the tool and the workpiece causes tool wear to build gradually [60]. According to the experimental results, milling forces rise significantly during chatter, ranging from 61.9% to 66.8% higher than those observed during stable cutting. Additionally, in comparison to stable cutting conditions, the machined surface’s quality deteriorates, with surface roughness increasing from 34.2% to 40.5% [61].

Figure 5 offers a detailed depiction of the relationship between flank wear and machined length, highlighting the impact of various spindle speeds and feed rates in the HSM of Ti6Al4V made possible by a flood cooling technique. The data presented in the graphs reveal a clear trend: as feed rate and spindle speed rise, flank wear rates also increase. The spindle speed of 1194 r/min is the minimum, ranging from 3.91 to 29.1 µm in Fig. 5a, and the flank wear is observed at a feed rate of 358 mm/min. However, when the feed rate is increased to 538 mm/min and then to 717 mm/min, the flank wear range rises to 5.62 to 151.3 µm and 6.79 to 237 µm, respectively, indicating a direct correlation between feed rate and flank wear. Figure 5b further illustrates the scenario at a higher spindle speed of 1791 r/min, where flank wear experiences a slight increase compared to the lower speed. The tool wear ranges are recorded as 4.38 to 167.2 µm for 358 mm/min, 4.79 to 232.7 µm for 538 mm/min, and 7.63 to 249 µm for 717 mm/min. This pattern of increasing flank wear continues as the spindle speed is increased to its maximum of 2389 r/min in Fig. 5c, where tool wear increases to ranges of 4.01 to 186.6 µm for 358 mm/min, 6.15 to 249.4 µm for 538 mm/min, and 9.34 to 289 µm for 717 mm/min.

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Notably, Fig. 5 shows that for all feed rates employed, flank wear increased as cutting length increased because frictional forces acted for a longer period while the workpiece and tool were in contact. As the machined length extended over 2000 mm, Fig. 5a and b show an exponential increase in flank wear at the greatest feed rate of 717 mm/min, suggesting that at higher feed rates, flank wear is more significantly influenced by longer cutting lengths compared to shorter ones, particularly at lower spindle speeds of 1194 r/min and 1791 r/min. In contrast, Fig. 5c illustrates that the wear curve transitions to a logarithmic pattern at the maximum 2389 r/min spindle speed and 717 mm/min feed rate, suggesting that the impact of spindle speed and feed rate on flank wear was higher at shorter machine lengths compared to longer machined lengths. The cutting tool achieves peak wear levels at this maximum feed rate of 717 mm/min, which is measured after a machined length of 4000 mm and is 237 µm at 1194 r/min, 249 µm at 1791 r/min, and 289 µm at 2389 r/min.

Figure 6 depicts the relationship between flank wear and machined length, highlighting the impact of varying spindle speeds and feed rates in the HSM of Ti6Al4V, facilitated by an MQL cooling strategy. In Fig. 6a, flank wear at a 358 mm/min feed rate and 1194 r/min spindle speed is minimal, ranging from 3.21 to 26.7 µm. Nonetheless, as the feed rate increases to 538 mm/min, the flank wear range expands to 4.74 to 37.8 µm and further increases to 5.46 to 77.7 µm at a feed rate of 717 mm/min. Figure 6b further illustrates the scenario at a greater spindle speed of 1791 r/min, where flank wear experiences a slight increase compared to the lower speed. The flank wear ranges are recorded as 3.42 to 27.3 µm for 358 mm/min, 4.35 to 54.9 µm for 538 mm/min, and 6.21 to 151.5 µm for 717 mm/min. This pattern of increasing wear continues as the spindle speed is increased to 2389 r/min in Fig. 6c, where flank wear increases to ranges of 3.93 to 28.7 µm for 358 mm/min, 4.95 to 237.5 µm for 538 mm/min, and 6.93 to 287 µm for 717 mm/min, where the machined length stops at 5000 mm. The cutting tool achieves peak flank wear levels of 77.7 µm at 1194 r/min, 151.5 µm at 1791 r/min, and 287 µm at 2389 r/min when operating at the maximum feed rate of 717 mm/min.

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Both higher feed rates and higher spindle speeds tend to accelerate flank wear. Figure 5 and 6 show that variations in feed rate significantly influence the progression of flank wear, especially at higher spindle speeds. As spindle speed rises, the effectiveness of lubrication diminishes, and thermal softening intensifies, resulting in greater flank wear. The initial rise in flank wear associated with increasing cutting speeds is due to material accumulation on the rake face’s cutting edge and elevated friction on the tool’s flank face [62]. This rise in temperature, driven by increased friction at faster cutting speeds, thermally softens the workpiece material [63], making chip removal simpler [64]. The generation of heat during the process affects both surface roughness and tool wear [65]. Additionally, as the feed rate rises, more material is removed, further contributing to increased flank wear [66, 67]. Flanking wear tends to escalate with rising feed rates because an effective chipping mechanism does not occur at the lowest feed rates until the chip thickness surpasses a certain threshold, at which point cutting becomes the dominant mechanism [68].

Cutting tools undergo wear on both the flank and rake faces, accompanied by associated wear mechanisms and failure modes. The two primary types of wear that significantly influence a tool’s lifespan are flank wear and crater wear. Flank wear mostly affects the relief face because of friction with the machined surface, while crater wear (rake face wear) develops due to high temperatures and friction from the chip-tool interaction. Crater wear on the rake face as a result of adhesion and diffusion when the chip moves across the rake face, which is exacerbated by elevated temperatures that soften the tool and promote chemical reactions between the tool and workpiece [69‐71]. The influence of feed rate on crater wear is significant. This is linked to the chamfer geometry of the inserts; when the feed rate exceeds the chamfer width, it becomes less constrained on chip displacement, leading to higher sliding velocity and friction-induced temperatures, resulting in increased crater wear [72].

Tool wear accumulates progressively throughout the machining operation, influenced by factors, including tool material, cutting setting, and lubrication [73]. Additionally, the interaction between spindle speeds and feed rates is a crucial element affecting tool wear [34]. During HSM operations, the tool’s cutting edge undergoes tremendous mechanical and thermal loads, which frequently cause the tool edge to deteriorate and fail too soon. The dynamic nature of tool wear significantly impacts overall machining performance. This study thoroughly investigates various forms of tool wear and evaluates the impacts of various cutting parameters on these wear processes. Figure 7 illustrates several failure patterns associated with tool wear observed under a flood cooling technique. Several tool wear failure modes, including nose radius wear and notch wear, are shown in Fig. 7a. Figure 7b is the chip hammering, and the other wear mode is the crater wear, as shown in Fig. 7c. Furthermore, the chipping is shown in Fig. 7d. Nose radius wear, notch wear, chip hammering, crater wear, and chipping are the tool wear failure types that occur while milling Ti6Al4V using a flood cooling method.

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A thorough examination of the several tool wear failure modes that arise while using the MQL cooling method to mill Ti6Al4V is illustrated in Fig. 8. The primary wear mechanism depicted in Fig. 8a is notch wear. Chipping is shown as a prominent failure mechanism in Fig. 8b. In the meantime, crater wear is depicted in Fig. 8c. Additionally, cutting-edge breakage is shown in Fig. 8d. Notch wear, chipping, crater wear, and cutting-edge breakage are the tool wear failure types seen while milling Ti6Al4V with MQL cooling.

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According to Figs. 7 and 8, abrasive wear is the frequently observed tool wear mechanism in uncoated carbide tools under flood and MQL cooling conditions. Abrasive wear primarily occurs on the flank face of the leading cutting edge and is one of the main causes of flank wear. This type of wear is mainly caused by friction between the tool and the workpiece. When chips enter the tool-workpiece interface, the friction will intensify further, accelerating tool wear [74]. In this study, flood cooling and MQL were performed under HSM conditions, which may explain why abrasive wear might have been the predominant wear mechanism in all cases. This observation agrees with the Khatri and Jahan [12] study, which agreed that in dry, flood, and MQL machining conditions, abrasion was the primary mechanism of tool wear.

Adhesion wear refers to tool wear where some workpiece material adheres to the tool’s surface. This alters the tool's geometry, and if machining continues, it may lead to tool breakage or deterioration of the workpiece surface quality. The primary cause of adhesion wear is the large quantity of heat produced by friction during the machining process [74]. Under flood cooling and MQL conditions, adhesion wear becomes more severe at larger depths of cut or higher feed rates.

Chipping is another wear mechanism observed under both machining conditions, though its occurrence is less frequent than abrasive wear. Chipping results from the large cutting forces exerted on the tool during machining. It is easily identifiable as it increases the negative rake angle of the tool and significantly alters its geometry. The occurrence of chipping is closely related to the high stress and temperature at the cutting edge, and the hardness of the tool material. Heat is generated by friction during wet cooling while the coolant cools the tool. This cyclic heating and cooling cause thermal fluctuations at the tool edge, leading to the formation of microthermal cracks.

These microcracks propagate more rapidly at higher cutting speeds, eventually causing chipping. This explains why chipping is more frequently observed under wet cooling conditions. Chipping typically occurs at the cutting edge, rake face, and tooltip. Continued use of tools with chipping damage can further lead to delamination, which is undesirable as it deteriorates the workpiece’s surface quality. Larger depths of cut or higher feed rates also increase the chipping frequency [12]. As shown in Figs. 7 and 8, chipping appears on the flank face near the tooltip.

Notch wear occurs when a workpiece's surface is more intricate or abrasive than the material underneath. This might come from forged or cast surfaces with a surface scale, or it can be the result of surface hardening from prior cuts (strain-hardening materials like stainless steels and super-alloys). As a result, the cutting-edge wears down faster when it comes into contact with the hard layer. Notch wear may also result from this focused, localized tension. Compressive stress develops when the cutting edge makes contact with the workpiece material; it does not form where it does not. When the two come into direct contact, this puts a lot of stress on the cutting edge (depth of cut line). Notch wear can also result from impacts, such as small disruptions or stiff micro inclusions in the workpiece material [75].

Chip hammering happens when a chip bends back and hits the portion of a cutting edge that is not being used. The outcome will be the breaking of a cutting edge (or a portion of a cutting edge) that is not in the cut. Operations combining high feeds and deep depths of cut are more likely to experience this risk [75]. The tool's nose radius experiences nose radius wear (Fig. 7a). It is mostly caused by abrasion, corrosion, or oxidation and appears to be a combination of flank and notch wear [76]. Lastly, cutting-edge breakage is not regarded as a wear pattern, but it must be included in any summary of basic tool wear patterns. Catastrophic cutting-edge breaking is an undesirable and hazardous condition brought on by improper tool use rather than a tool wear pattern. When a cutting edge breaks, the cutting conditions are chosen so that the edge is subjected to mechanical loads that are too severe for it to bear [75].

Figures 9a and 9b illustrates the effect of feed rate on surface roughness at various spindle speeds under flood and MQL cooling techniques. The findings show that surface roughness rises with increasing feed rates and spindle speeds. In particular, as the feed rate increases, a commensurate rise in surface roughness is noted. Higher feed rates result in the removal of more material per revolution, which raises the MRR per cutting cycle and explains this pattern [77]. An elevated feed rate results in larger chip thicknesses, enhancing material removal efficiency and imposing greater loads on the cutting tool. This increased load accelerates tool wear and produces a rougher machined surface [78]. Under higher cutting forces, friction between the tool and workpiece and between the tool and chips intensifies. This friction generates heat, with approximately 90% of the mechanical energy during cutting converting into thermal energy [79]. The thermal energy elevates the temperature at the cutting edge, accelerating tool degradation mechanisms such as abrasive wear, plastic deformation, and thermal cracking. Elevated cutting temperatures adversely affect tool life, workpiece surface integrity, and chip formation processes, contributing to increased surface roughness and potential dimensional inaccuracies [80]. This study’s findings align with these principles, demonstrating that raising the feed rate considerably raises the cutting temperature. The elevated temperature accelerates tool wear and can cause surface damage, thus diminishing surface quality.

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Moreover, as the MRR increases, so do the cutting forces exerted on the tool and the machine structure. These larger forces can induce vibrations and dynamic instabilities during machining [81]. Such vibrations, often caused by dynamic force interactions at higher feed rates, can deteriorate surface finish quality [82, 83]. Higher feed rates raise cutting forces, which cause more system vibrations and instability, as shown in Figs. 3 and 4. Therefore, selecting an optimal feed rate is critical, considering the machine's capacity to effectively absorb and dissipate these forces and vibrations [53].

The data indicate that surface roughness increases with higher spindle speeds, regardless of whether flood or MQL cooling strategies are used. These findings are consistent with other research, including that conducted by Akinlabi et al. [10], who reported that increasing spindle speeds can result in increased surface roughness because of enhanced friction between the workpiece and the cutting tool. The cutting tool's temperature rises quickly as a result of increased heat generated by increased friction, particularly at the cutting edge. This thermal buildup accelerates tool wear mechanisms, including abrasive wear, plastic deformation, and thermal cracking [79]. The rise in thermal energy at the cutting interface shortens tool life and impacts workpiece surface quality and chip formation, potentially causing dimensional inaccuracies. Spindle speed and feed rate significantly influence machined components' surface finish and dimensional precision [83, 84].

It was observed that HSM under flood cooling conditions results in higher surface roughness values, whereas employing MQL cooling leads to smoother surfaces. Specifically, surface roughness measurements under flood cooling ranged from approximately 0.261 μm at a spindle speed of 1194 rpm and a feed rate of 358 mm/min to about 0.581 μm at 2389 rpm and 738 mm/min. In contrast, with MQL cooling, the minimum roughness recorded was around 0.221 μm at the lower spindle speed and feed rate, while the maximum was 0.533 μm at 2389 r/min and 738 mm/min. In comparison to flood cooling, MQL quantitatively increased the workpiece's surface roughness by approximately 15.33% at the minimal spindle speed and feed rate and 8.26% at the greatest spindle speed and feed rate.

The graph depicts that when using the MQL cooling strategy, the resultant cutting force changes differently from the flood cooling strategy, possibly due to the cooling lubrication mechanism and coefficient of friction. Flood cooling provides a large amount of coolant, which can effectively reduce temperature and friction [10]. While milling under MQL, a lubricating film forms between the workpiece and the tool, minimizing friction and cutting force [11]. In addition, it was found that the tool’s performance under MQL coolant was noticeably superior to that under overflow coolant. The MQL cooling approach produced less cutting force, less surface roughness, and less tool wear than the flood cooling method.

The traditional method for applying cutting fluid in machining is known as traditional cooling or the flood method. This approach involves delivering a significant amount of lubricant directly into the cutting zone, which aids in cooling and lubricating the interface between the tool and the workpiece [85]. However, this method tends to lead to increased wear rates of cutting tools and compromised surface integrity, as it fails to effectively supply cooling and lubrication to the cutting zone, thus diminishing overall process efficiency [6]. On the contrary, one cooling approach that tries to enhance cutting performance is MQL, often known as near-dry machining (NDM). Because it uses less cutting fluid, this technique is also regarded as environmentally friendly [29]. Depending on the chosen machining parameters, MQL can allow for a 10% reduction in the cutting temperature [86]. Research has indicated that this temperature reduction can mitigate the adverse effects on tool wear intensity [86, 87]. Referring specifically to titanium alloy, a material that is challenging to cut, tool wear comparisons have demonstrated a notable decrease when transitioning from dry machining to MQL. For instance, Khan et al. [88] reported a 28% reduction in tool wear when comparing MQL conditions to dry milling of the Ti6V4Al.

The MQL cooling approach in the present study was found to cause less flank wear than the flood cooling strategy. This outcome agrees with the one Zhao et al. [89] reported. By comparing the findings at the maximum feed rate of 717 mm/min with a 4000 mm machine length, it was shown that flank wear deteriorated considerably more under flood cooling conditions than under MQL settings. The severity of flank wear under the flood cooling approach than the MQL cooling strategy, as depicted in Fig. 10a. With increasing feed rates, flank wear significantly increased at lower spindle speeds of 1194 r/min and 1791 r/min. In particular, flank wear under the flood cooling technique was 13.55% more severe than under MQL cooling at 1194 r/min and 358 mm/min. The degree of flank wear sharply increased to 376.14% and 493.84%, respectively, as the feed rate rose to 538 mm/min and 738 mm/min. The degree of flank wear rose from 244.04% at 358 mm/min to 320.67% at 538 mm/min and 317.09% at 738 mm/min, following a similar pattern at 1791 r/min spindle speed. On the contrary, the greatest flank wear severity under flood conditions was 274.79% at 358 mm/min feed rate and a maximum spindle speed of 2389 r/min. As the feed rate rose at this spindle speed, the difference in flank wear between the two cooling techniques became less noticeable.

Furthermore, Fig. 10b demonstrates the notable advancements in tool wear when employing the MQL cooling strategy in contrast to the traditional flood cooling method. At a feed rate of 358 mm/min, the improvement in flank wear observed with the MQL approach showed a low percentage of improvements with 11.93% at 1194 r/min spindle speed, extending significantly to 70.93% at 1791 r/min, and continuing to manifest with a further enhancement to 73.32% as the spindle speed increased to 2389 r/min than the flood cooling strategy. The efficacy of the MQL cooling strategy became even more pronounced at higher feed rates, with improvements in tool wear recorded at 79% and 83.16% for 538 mm/min and 738 mm/min feed rates, respectively. The measured tool wear improved to 70.93%, 76.23%, and 76.02% at feed rates of 358 mm/min, 538 mm/min, and 738 mm/min, respectively, at a spindle speed of 1791 r/min. However, as the spindle speed was increased to 2389 r/min, a change in trend was seen. Under these conditions, the improvements in tool wear with the MQL strategy were reduced to 2.73% for a feed rate of 538 mm/min and 17.30% for 738 mm/min. This indicates that while MQL provides significant advantages at lower spindle speeds and varying feed rates, the benefits diminish as the spindle speed rises, particularly beyond 2389 r/min when paired with higher feed rates of 538 mm/min and 738 mm/min.

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At lower spindle speeds, MQL using coconut oil proves highly effective in reducing friction, which leads to lower heat generation during HSM and minimizes tool wear. This is primarily due to the minimal thermal energy produced under such conditions. As the spindle speed increases to moderate levels, heat generation becomes more pronounced; however, the cooling efficiency of coconut oil under MQL was demonstrated. Coconut oil evaporates incredibly slowly, roughly a million times slower than water, due to its relatively low vapor pressure. Coconut oil molecules undergo chemical processes and disintegrate into smaller, more volatile components that can evaporate when heated to extremely high temperatures, such as 370 °C [90]. The mist helps remove heat by dissipating around 370°C. At this stage, an optimal balance between lubrication and cooling is achieved, resulting in reduced cutting forces and improved machining performance. Nonetheless, at very high spindle speeds, tremendous heat and cutting forces are generated. The limited amount of lubricant may be insufficient to effectively disperse heat, resulting in thermal damage and accelerated tool wear.

The primary factors contributing to surface roughness include feed rate, cutting speed, and cutting tool geometry, while secondary impacts are caused by machine tool properties and machining process variances, such as tool wear [91]. Furthermore, the findings published by Rangasamy et al. [92] and Lavanya et al. [93] were supported by the observation that surface roughness values were generally lower when using the MQL cooling strategy in comparison to traditional flood cooling. Notably, the MQL method demonstrated a maximum enhancement in surface roughness of 15.33% compared to flood cooling at 1194 r/min spindle speed and 358 mm/min feed rate, as presented in Fig. 11. However, this advantage diminished at greater spindle speeds and feed rates; under the most extreme conditions, the MQL technique improved surface roughness by 8.26% compared to the flood method. The MQL offers a viable alternative to wet machining and serves as an eco-friendly machining option [93]. The reduced roughness with the MQL cooling strategy indicates its potential for aerospace applications requiring precision.

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Ti6Al4V’s strong tensile strength, poor thermal conductivity, and high chemical reactivity make machining it extremely difficult. [94]. The heat generated during machining predominantly remains localized within the workpiece because of titanium's low thermal conductivity and tendency to generate substantial heat during cutting. During the process, only the tool's cutting edge actively participates in chip formation, with elevated temperatures concentrated at this zone. Consequently, higher cutting temperatures exacerbate thermal stresses because of the limited heat dissipation between the chips and the workpiece. This encourages rapid heat accumulation, which results in BUE generation and chip welding. Effective cooling strategies are essential for efficiently evacuating chips, providing cooling and lubrication, and reducing adhesion issues [95]. Among these, MQL offers notable advantages, particularly for titanium alloys, by significantly enhancing surface quality through reduced thermal distortion [96]. MQL’s effectiveness stems from creating a tribo-film at the tool–chip interface, which reduces the development of BUE and BUL, facilitates quick chip removal, and mitigates metallurgical problems associated with dry machining [97].

Additionally, the primary limitation of Ti6Al4V is its poor tribological performance, characterized by high and unstable friction coefficients [98] and low wear resistance [99]. Adhesive wear mechanisms, such as deep craters and material transfer to the counterface, are frequently observed, resulting in elevated tool wear rates [100, 101]. Consequently, titanium alloys are often restricted to applications with minimal tribological demands. Dry machining of Ti6Al4V is generally unsuitable due to these tribological challenges; thus, lubrication becomes essential to prevent adhesion and excessive tool wear. The material’s metallurgical behaviour, chip formation dynamics, cutting tool wear, lubrication strategies, and surface integrity are critical considerations in optimizing machining performance with Ti6Al4V [102].