A novel approach to develop the turning tool for cryogenic internal cooling

Abstract

The LN₂ internal spray cooling, achieved through channels integrated within the tool, facilitates the delivery of liquid nitrogen to the vicinity of the cutting edge, thereby enabling precise cooling of the cutting area and enhancing the cooling efficiency of the liquid nitrogen. The implementation of LN₂ internal spray cooling necessitates specialized tools. However, there is currently a lack of comprehensive technology for the development of such tools. Building on previous research, this study further optimizes the turning tool for LN₂ internal spray cooling. The impact of nozzle structures on the jet characteristics is investigated, and the effect of the outlet elbow angle on cavitation is analyzed. Additionally, an insulation structure for the primary liquid nitrogen transmission channel is designed. A comprehensive development method for the turning tool utilizing LN₂ internal spray cooling is proposed. Using TA15 titanium alloy as the experimental material, the optimized tool results in an approximate 23% reduction in cutting temperature, a 20% decrease in surface roughness, and a 25% reduction in tool wear.

Introduction

Cryogenic machining is deemed essential for the efficient dissipation of heat generated during cutting operations^1–3. Among the range of cooling agents, liquid nitrogen and liquid carbon dioxide have been most extensively employed in such applications^4–6. Owing to its extremely low temperature, superior cooling efficiency, and environmental benefits, liquid nitrogen has become widely adopted for machining challenging materials, including AISI 52,100 hardened steel⁷, Co-Cr-Mo alloys⁸, magnesium alloys⁹, porous tungsten¹⁰, CFRP¹¹, and PTFE¹². This method has been demonstrated to yield exceptional performance regarding surface integrity and tool wear in these contexts. Moreover, the internal cooling approach with liquid nitrogen permits its delivery in close proximity to the cutting edge via the tool’s internal cavity, thereby achieving precise cooling of the cutting area.

Accordingly, investigations into internally cooled tools have been undertaken. Internally cooled tools are presently classified into two principal types: internal circulation cooling tools and internal spray cooling tools. In the former, a circulation unit for the cooling medium is integrated either within the tool body or externally, permitting the medium to be re-cooled via a dedicated heat-exchange module. In the latter, hollow transmission channels are embedded in the tool, through which the cooling medium is conveyed and precisely sprayed onto the designated cutting zone.

Wang et al.¹³ developed an external sealed heat exchange unit positioned at the upper section of the turning tool insert. By employing liquid nitrogen as the cooling medium, it was demonstrated that the cutting temperature could be substantially reduced while tool wear during the machining of ceramics, titanium alloys, superalloys, and tantalum was concurrently minimized. However, it was noted that the external heat exchange unit might interfere with the effective clamping of the insert, thereby complicating its secure fixation. Consequently, the adoption of an internal heat exchange unit for internal circulation cooling has been favored. Vicentin et al.¹⁴ introduced a design featuring a cavity located beneath the insert for the storage of the cooling medium. In their investigation, a hydrochlorofluorocarbon (HCFC) fluid was used as the cooling agent while machining Cr-Ni-Nb-Mn-N austenitic steel under varying conditions, including dry, wet, and internal circulation cooling. The findings indicated that the surface roughness and tool wear achieved with the internally circulated turning tool were comparable to those obtained using emulsion cooling, with the internal circulation tool also exhibiting the capacity to maintain more stable cutting temperatures during extended cutting operations. Expanding on these earlier developments, Sun et al.¹⁵ engineered a cooling medium heat exchange unit constituted by the combination of the insert and a spacer. In order to enhance the cooling efficiency of the unit, the dimensions of the cavity, as well as the diameters of the inlet and outlet holes, were systematically optimized through simulation.

Internal circulation cooling tools have been shown to significantly mitigate the cutting temperature, while also preventing direct contact of the cooling medium with the workpiece, thereby proving particularly advantageous for machining materials susceptible to thermal deformation. Nonetheless, due to spatial constraints within the heat exchange unit and the propensity for the cooling medium to evaporate at elevated cutting temperatures, the overall cooling efficiency may be compromised¹⁶. Moreover, the internal circulation configuration necessitates a robust sealing mechanism, especially when employing low-temperature media that are prone to evaporation. In contrast, internal spray tools facilitate the direct delivery of the cooling medium to the cutting zone, thereby optimizing its cooling potential without inflicting any damage to the inserts or spacers and ensuring the stability of the tool during operation. These systems are activated simply connecting the tool to a cooling medium supply unit.

Peng¹⁷ developed an internal spray tool featuring cooling holes on the flank face, using Inconel 718 for both simulation and experimental investigations. The findings indicated that internal spray cooling, when applied to Inconel 718, significantly reduced cutting temperatures and surface roughness compared to traditional flood cooling methods. Hong¹⁸ designed a turning tool with internal liquid nitrogen injection and integrated nozzle switches to control the application of liquid nitrogen to either the rake face, the flank face, or both simultaneously. His experiments on Ti-6Al-4 V titanium alloy demonstrated that cooling both the rake and flank faces together resulted in a substantial increase in tool life, marking a pioneering use of internal liquid nitrogen injection in cryogenic machining. In a similar vein, Bermingham et al.¹⁹ developed turning tools with four distinct nozzle positions for internal injection, comparing machining conditions of Ti-6Al-4 V under dry cutting, high-pressure emulsion, and liquid nitrogen cooling. Their results revealed that high-pressure emulsion cooling exhibited little sensitivity to nozzle position, while liquid nitrogen cooling was more responsive, with simultaneous cooling of both rake and flank faces leading to a longer tool life than when only the rake face was cooled. Gan et al.²⁰ investigated the relationship between the diameter of liquid nitrogen jet holes and cooling distance, optimizing the jet hole size to minimize the impact of the Leidenfrost effect on cooling efficiency during turning operations.

An appraisal of the current state of the art reveals that internal spray tools affords more effective exploitation of a cooling medium’s capacity. However, such tools have, to date, been predominantly applied with media at ambient temperature. In the case of liquid nitrogen—characterized by its ultra-low temperature and high evaporation tendency—existing internal spray designs have been confined to nozzle-position optimization. Neither the evaporative losses incurred during transit through the tool’s internal channels nor the attenuation of cooling efficacy during atomization have been systematically addressed, thereby leaving a gap in the comprehensive design methodology for liquid nitrogen based internal spray tools. Building upon the work of Gan et al.²⁰, the present study refines the internal spray turning tool by optimizing the configuration of its transmission channels and overall structural layout, with the aim of enhancing tool performance and advancing the industrial application of liquid nitrogen cryogenic cooling.

Key structural optimization

The analysis of the impact of nozzle structure on the jet

To investigate the influence of different nozzle structures on jet states, nozzles with four different structures were designed for comparative analysis: a straight nozzle, a conical contraction nozzle, an arc contraction nozzle, and an expansion nozzle. Figure 1 shows schematic diagrams of the four nozzle structures, while Fig. 2 presents the simulation analysis results of jet states for different nozzle structures.

Fig. 1.

Schematic of nozzle structure. (a) Straight nozzle. (b) Conical contraction nozzle. (c) Arc contraction nozzle. (d) Expansion nozzle.

Fig. 2.

The jet simulation diagram of different structure nozzles. (a) Straight nozzle. (b) Conical contraction nozzle. (c) Arc contraction nozzle. (d) Expansion nozzle.

Based on the theory of continuum mechanics, and according to the laws of conservation of mass, momentum, and energy, the control equations for the motion of a continuous fluid are obtained.

1. Continuity equation:

   1
2. Momentum conservation equation:

   2
3. Energy conservation equation:

   3

   where, u is the fluid velocity, ρ is the fluid density, p is pressure, µ is the fluid’s viscosity coefficient, f_b is volumetric force, f_st is the surface tension at the multiphase flow interface, C_p is the specific heat capacity at constant pressure, T is the fluid temperature, k is the thermal conductivity coefficient.

In this study, the liquid nitrogen jet is a non-submerged free jet, ejected from the nozzle into the air, involving gas-liquid two-phase flow. The VOF (Volume of Fluid) model is selected for interface tracking, and the VOF equation can be expressed as:

4

where, ρ_k is the density of phase k, v_k is the velocity of phase k, S_αk is the source term, m_pk is the mass transfer from phase p to k, m_kp is the mass transfer from phase k to p, and α_k is the volume fraction of phase k.

The jet is in a high turbulence state, and the standard k-ε turbulence model is selected to close the control equations. In the standard k-ε turbulence model, the transport equations for the turbulent kinetic energy k and the specific dissipation rate ε are expressed as follows:

5

where, G_k is the turbulent kinetic energy generated by the gradient of mean velocity, G_b is the turbulent kinetic energy generated by buoyancy, Y_M is the dissipation occurring in turbulence, C_1ε, C_2ε, C_3ε are constants, σ_k and σ_ε are the turbulent Prandtl numbers for the k-equation and ε-equation, respectively, S_k and S_ε are source terms, the turbulent viscosity coefficient .

Computational fluid dynamics software ANSYS FLUENT was used for the simulation calculation of liquid nitrogen jet. The computational domain was defined by the combination of a liquid nitrogen nozzle and its downstream jet region. Owing to the axisymmetric nature of the round jet, a two-dimensional modeling approach was adopted by extracting a representative cross-section along the jet axis. The domain was treated as an axisymmetric system and constructed accordingly using a rotationally symmetric plane. To ensure adequate spatial development of the jet flow, a large flow domain measuring 80 mm in length and 18 mm in width was employed. The upper boundary and both lateral sides of the jet region were defined as adiabatic, no-slip wall boundaries to represent thermal and mechanical insulation. Owing to the regular geometry of the entire computational region, a structured mesh was utilized for spatial discretization. Mesh refinement was applied around the nozzle exit and the central axis to capture detailed flow features and enhance numerical accuracy. The geometric model consists of the liquid nitrogen nozzle and the liquid nitrogen jet domain, forming the computational domain. Since the geometric model is a circular jet, any cross-section passing through the axis is taken for two-dimensional modeling. Assuming the jet domain is axisymmetric, modeling can be done using an axisymmetric plane. The jet domain is set as a large spatial domain of 80 mm×18 mm to ensure the full development of the jet flow field, with the top surface and the left and right sides of the jet domain as adiabatic no-slip wall boundaries. All computational domain spaces are of regular shape, allowing for spatial discretization using structured grids, and the nozzle and the area around the axis of the jet domain are refined.

In the simulation analysis model, the total length of the nozzle is 10 mm, and the length of the variable diameter part is 2 mm.The inlet pressure is 0.4 MPa. To ensure the consistency of the outlet size, the diameter of the nozzle outlet is set to 2 mm. The inlet diameter is determined by back-calculating from the outlet diameter and the nozzle length. As can be seen from the velocity contour map in Fig. 2, the expansion nozzle has a significant effect on the length of the zone of flow establishment, reducing the length of the zone of flow establishment. This type of nozzle structure is not suitable for use, as the other three structures do not significantly affect the jet state. To further analyze the impact of different nozzle structures on the jet state, it is necessary to analyze the distribution of velocity and temperature along the jet path.

Figure 3 is a comparative diagram of the jet velocity and temperature distribution for nozzles with different structures. Figure 3a shows the velocity distribution of jets from nozzles with different structures, where the expansion nozzle exhibits a rapid decay in velocity with a shorter length of the zone of flow establishment, and there is no significant difference in the zone of flow establishment among the straight nozzle, the conical contraction nozzle, and the arc contraction nozzle. From region I in the diagram, it can be seen that the conical contraction nozzle and the arc contraction nozzle have a rising phase at the beginning of the jet, which is absent in both the straight and expansion nozzles, suggesting that this is caused by the contraction structure, but the maintenance length is not significantly different from that of the straight nozzle. Figure 3b shows the temperature distribution of jets from nozzles with different structures, which follows the same pattern as the velocity distribution, with the expansion nozzle having the fastest temperature decay rate among the four types of nozzles. Although the conical contraction nozzle and the arc contraction nozzle have a rising phase in their velocity distribution, it does not affect the temperature distribution of the jet. The straight nozzle maintains stable velocity and temperature in the zone of flow establishment, without any significant abrupt changes.

Fig. 3.

Comparison of jet velocity and temperature of different structure nozzles. (a) Velocity distribution diagram. (b) Temperature distribution diagram.

Comprehensive consideration of the velocity and temperature distribution patterns of the four nozzles, as well as the feasibility of the actual manufacturing process, compared to the other three types of nozzles, the straight nozzle has obvious advantages, so the nozzle of the liquid nitrogen internal spray turning tool can adopt the straight nozzle structure.

Optimization of the outlet elbow angle

In the manufacturing process of cutting tools, to ensure the accuracy of the spray position, a bend pipe structure will inevitably appear at the outlet of the transmission channel. Due to the high pressure of liquid nitrogen in the internal channel of the tool, there is a significant pressure change when it is ejected compared to the atmospheric pressure at the outlet. The pressure difference between the inside and outside of the spray outlet leads to the cavitation effect at the outlet elbow, affecting the jet speed and the volume fraction of liquid nitrogen, as shown in Fig. 4.

Fig. 4.

Schematic of outlet elbow cavitation.

The effects of cavitation caused by different bend angles at the outlet on the speed, volume fraction, and temperature of liquid nitrogen were analyzed using a simulation approach. Due to the occurrence of secondary flow phenomena at the outlet bend, the RNG k-ε turbulence model is more accurate in simulating transient flows and curved streamlines compared to the standard k-ε model²¹. Therefore, the RNG k-ε turbulence model was employed for simulation analysis. The expression for the RNG k-ε turbulence model is as follows:

6

where, G_k is the turbulence kinetic energy produced by the mean velocity gradient, G_b is the turbulence kinetic energy generated by buoyancy, Y_M is the dissipation produced in turbulence, S_k and S_ε are source terms, α_k and α_ε are the turbulent Prandtl numbers for the k and ε equations respectively, C_1ε, C_2ε, C_3ε, C_µ, β, η₀ are constants, η is the dimensionless strain rate, µ_t is the turbulent viscosity.

The Zwart–Gerber–Belamri²² model is selected as the cavitation model, and its expression is as follows:

7

where, R_e is the evaporation term, F_vap is the evaporation coefficient, R_c is the condensation term, F_cond is the condensation coefficient, α_v is the gas phase volume fraction, α_nuc is the nucleation site volume fraction, ρ_v and ρ_l is the gas and liquid phase densities respectively, r_b is the bubble diameter, P₀ is the actual pressure, P_sat is the saturation pressure.

To analyze the impact of changes in the bend angle on parameters such as the velocity, volume fraction, and temperature of liquid nitrogen, a comparative analysis was conducted with the angle range set from 30° to 160°, and the step interval set at 10°. The pipe diameter was set to 2 mm, with a single segment length of 10 mm, using structured mesh. The inlet and outlet of the computational domain were set as constant pressure boundary conditions, with the inlet absolute pressure set at 0.4 MPa, and the inlet temperature undercooling issue considered, setting the undercooling degree at 10 K, with an inlet temperature of 80 K. The outlet pressure of the liquid nitrogen was set to one standard atmosphere, with a temperature of 300 K. The walls of the bend were set as adiabatic surfaces. The meshing method employed was MultiZone, with a grid size of 200 μm. Figure 5 shows the trend of changes in the velocity, volume fraction, and temperature of liquid nitrogen under different bend angles.

Fig. 5.

Variation of liquid nitrogen velocity, volume fraction and temperature for different angles.

Analysis of the velocity, liquid nitrogen volume fraction, and temperature from different bend angles shows that the bend angle has almost no effect on the fluid temperature. Due to the presence of liquid nitrogen, the temperature from the inlet fluid side to the outlet air side is maintained at 77 K. As the bend angle increases, the velocity and volume fraction of liquid nitrogen at the outlet exhibit a trend that is nearly linear, with the bend angle having the most significant impact on the volume fraction of liquid nitrogen, the maximum difference exceeding 20%. This indicates that increasing the bend angle can effectively reduce the effects of secondary flow and cavitation on liquid nitrogen. Based on the analysis, to maintain a high initial velocity and volume fraction of liquid nitrogen during injection, the outlet of liquid nitrogen injection should use as obtuse an angle as possible, to minimize the impact on velocity and liquid nitrogen volume fraction.

Study on the thermal insulation method of the main transmission channel

The process of transferring liquid nitrogen from the supply device to the tool necessitates the use of vacuum tubes to prevent vaporization. Vacuum tubes are thicker than ordinary tubes and possess a greater resistance to bending. To avoid interference with machine tool components during tool use, or even the breaking of adapters, the liquid nitrogen inlet is positioned at the tail of the tool holder. However, this increases the length of the liquid nitrogen’s transmission path, exacerbating the heat exchange between the nitrogen and the inner walls during its transmission within the tool body. At this point, the liquid nitrogen enters a uniformly heated channel. If the external heat input through the walls is substantial enough, saturated boiling may occur, leading to a continuous increase in dryness, and the flow pattern of the liquid may transition from bubbly flow to annular flow, to mist flow, and finally to a unidirectional gas state^23,24. Disregarding the heat transfer along the length of the pipe, this process can be viewed as one-dimensional steady-state heat transfer through the cylindrical wall, with the heat Q transferred to the liquid nitrogen through the wall from the outside as:

8

where, r₁ is the inner diameter of the liquid nitrogen transmission channel in the tool handle, r₂ is the equivalent outer diameter of the tool handle, l is the length of the tool handle, λ is the thermal conductivity of the tool material, α₁ is the combined natural convection-radiation heat transfer coefficient of the air outside the tool wall; α₂ is the forced convection heat transfer coefficient of the two-phase flow of liquid nitrogen inside the pipeline, T₀ is the ambient temperature, and T_ln2 is the temperature of the liquid nitrogen.

Therefore, the only method to reduce the external heat input is to increase the equivalent thermal resistance of the cylinder wall. Installing insulation material within the cylinder wall is an effective way to increase its equivalent thermal resistance. After the addition of insulation material, the heat transfer expression is as shown in Eq. (9).

9

where, r₃ is the inner diameter of the insulation sleeve, and λ₁ is the thermal conductivity coefficient of the insulating material.

As shown by Eq. (9), the methods to increase the thermal resistance of insulation materials include selecting insulation materials with lower thermal conductivity, increasing the thickness of the insulation material walls, and decreasing the heat transfer area. Once the insulation material is selected, the thermal conductivity becomes a fixed value, and the insulation performance can only be improved by increasing the thickness of the insulation material walls and reducing the heat transfer area. Figure 6 is a simulation analysis result graph for increasing the thickness of the insulation material cylinder wall after fixing the thermal conductivity of the insulation material, where the thermal conductivity of the insulation material is set to 0.24 W/(m·℃), selecting different insulation material wall thicknesses for insulation performance comparison.

Fig. 6.

Diagram of insulation channel temperature distribution. (a) Wall thickness 1 mm. (b) Wall thickness 1.5 mm. (c) Wall thickness 2 mm. (d) Wall thickness 2.5 mm.

From the simulation results in Fig. 6, it can be observed that as the thickness of the insulation material wall increases from 1 mm to 2.5 mm, with each increment of 0.5 mm, the external wall temperatures of the tool body are − 127.04℃, − 104.30℃, − 85.89℃, and − 70.59℃, respectively. The temperature of the tool body’s external wall does not increase linearly with the increase in wall thickness; instead, the decrement in temperature difference is observed to trend downwards, with a diminished increase in insulation performance, indicating that the improvement in insulation performance diminishes with thicker walls. Therefore, simply increasing the thickness of the wall has limited benefits for enhancing insulation performance. Moreover, due to the dimensional constraints of the tool channels, excessive thickness in the insulation material cylinder wall would reduce the space for liquid nitrogen flow, making it impractical to only increase the thickness of the insulation sleeve. Hence, the method of reducing the heat transfer area is adopted to improve the insulation capability of the insulation sleeve. By grooving the surface of the insulation sleeve to form a “dumbbell” structure, creating cavities between the inner surface of the tool body and the outer surface of the insulation sleeve, the method of heat transfer is altered, and the efficiency of heat conduction is reduced. Figure 7a presents the improved insulation material structure diagram, and Fig. 7b shows the heat transfer simulation result. In the simulation, the thermal conductivity of the insulation material is set to 0.24 W/(m·℃), with the medium inside the cavity being air, having a thermal conductivity of 0.0267 W/(m·℃). With wall thicknesses of 2 mm at both ends and 1 mm in the middle, the external wall temperature of the tool body is only − 27.30 ℃, which, compared to using a 2 mm thick insulation sleeve without cavities (with an external wall temperature of − 85.89 ℃), shows a significant improvement in insulation capability. Therefore, the insulation sleeve for the main liquid nitrogen transport channel of the tool will adopt a “dumbbell” type structure.

Fig. 7.

Diagram of improved thermal insulation channel structure and temperature distribution. (a) Thermal insulation structure. (b) Heat transfer simulation results.

Simulation analysis of the thermal insulation performance of a liquid nitrogen internal spray cooling turning tool

To investigate the impact of insulation material within the liquid nitrogen transmission channel on the temperature field of the tool, a simulation is employed to compare the temperature distribution patterns with and without insulation. Given that the temperature at various points of the tool only changes with spatial position and not with time when liquid nitrogen is in a stable spraying state, steady-state heat transfer analysis is utilized for computation. In view of the complex geometry of the cutting tool model, a hybrid meshing strategy combining structured and unstructured grids was adopted, with an element size of 1 mm. The inner wall of the liquid nitrogen transmission channel within the tool is set as a cold source, considering heat transfer between the liquid nitrogen and the tool body, as well as between the tool body and air, with the cold source temperature at − 196 °C and air temperature at 20 °C. The simulation results are presented in Fig. 8a.

Fig. 8.

Temperature distribution with and without insulation. (a) Simulation results. (b) Temperature distribution in the extracted area.

The ridge area on the upper surface of the tool holder is selected, and the temperature distribution data from the head to the end of the tool are extracted to draw a temperature distribution curve, as shown in Fig. 8b. When no insulation structure is set on the tool handle, the temperature of the outer surface of the tool is low and distributed evenly, close to the temperature of liquid nitrogen. This indicates that when liquid nitrogen passes through the tool in the transmission channel, it directly exchanges heat with the outer wall, and the temperature starts to lose from the tool holder. At this time, the tool’s outer surface is prone to frost, and due to the heat loss in the channel, when liquid nitrogen is sprayed near the tool tip, the liquid volume fraction inevitably decreases, affecting the cooling efficiency of liquid nitrogen. Conversely, when an insulation structure is set up, the thermal exchange of liquid nitrogen with the environment is significantly suppressed at tool holder, resulting in a gradient temperature distribution in the tool holder with the insulation structure. This characteristic is due to the thermal conduction from the head of the tool holder, indicating that liquid nitrogen does not directly exchange heat with the external environment in the tool holder, thereby suppressing the temperature loss of liquid nitrogen in the transmission channel.

Experimental

Experimental materials

The Ti–6Al–2Zr–1Mo–1 V alloy (TA15), classified as a near-α titanium alloy, is extensively utilized in both aeronautical and astronautical applications due to its outstanding strength, low density, and exceptional corrosion resistance properties^25,26. In the present study, a cylindrical workpiece made from TA15 alloy, with dimensions of 100 mm × ϕ40 mm, was fabricated. The detailed chemical composition is provided in Table 1.

Table 1.

Chemical composition of TA15 alloy (mass fraction, %).

Al	Zr	Fe	Mo	V	Ti
6.55	1.59	0.2	0.94	0.82	Bal.

Experimental setup

The experiments were conducted using a horizontal lathe produced by Dalian Machine Tool Group (CD6140A, China). Liquid nitrogen was supplied from a tank (Tianhai DPL-175, China), and the flow of liquid nitrogen was regulated using a self-designed flow control device. A coated DNMG tungsten carbide insert from Sandvik-Coromant^® was employed as the cutting tool insert.

A FLIR SC9000 thermal imaging camera (FLIR Systems, USA) was employed to monitor the temperature distribution during the cutting process in real time. The camera was mounted 500 mm from the cutting zone at a 45° inclination, with its field of view directed toward the cutting area to ensure optimal visibility while avoiding interference from chips. The emissivity was set to 0.7. The acquired infrared data were processed using FLIR Tools software and synchronized with the cutting operation. In the assessment of surface quality and tool wear, a 3D optical surface profilometer (NewView 9000) with a maximum scanning depth of 150 μm was employed to measure the surface roughness of the machined specimens, and a super-depth-of-field 3D scanner (KEYENCE VHX-600, Japan), equipped with both a wide-range zoom lens and a high-definition zoom lens, was used to inspect machining defects on the workpiece surface as well as tool wear.

Internal spray cooling method is employed, and comparative turning experiments are conducted between the developed liquid nitrogen internal spray cooling tool and the non-optimized internal spray cooling tool. The working angles of the two tools were identical. The rake angle is 5°, the clearance angle is 5° and the nose radius is 0.8 mm. The cutting parameters are as follows: cutting speed of 90 m/min, feed rate of 0.1 mm/r, cutting depth of 0.4 mm, and material removal volume of 5000 mm². Three replicates were performed for each experimental group. The experimental setup is shown in Fig. 9.

Fig. 9.

Experimental setup for the experiments. (a) Platform. (b) Structure schematic of unoptimized turning tool. (c) Structure schematic of optimized turning tool.

Verification of performance improvement of liquid nitrogen internal spray cooling tool

Comparison of cutting temperatures

Figure 10 presents the thermographic images measured during the machining of two tools. The highest temperature reached by the machining with the non-optimized internal spray cooling turning tool is 185.3 ℃, whereas the cutting temperature with the liquid nitrogen internal spray turning tool developed in this study is 142.6 ℃. The results indicate that the optimized tool can enhance cooling efficiency, reducing the cutting temperature by approximately 23%.

Fig. 10.

Comparison diagram of cutting temperature. (a) The machining temperature of the non-optimized turning tool. (b) The machining temperature of turning tool developed in this study.

Comparison of surface quality

Figures 11 and 12 provide a comparative analysis of the 3D contours and topographies of surfaces machined by two different cutting tools. It is evident from the images that the surface machined with the internal cooling turning tool developed in this study is smoother and fewer defects. This effectively mitigates the formation of adhesions and pits, leading to a reduction in the roughness value by approximately 20%.

Fig. 11.

Comparison diagram of surface profile. (a) The 3D contours of surface machined by non-optimized turning tool. (b) The 3D contours of surface machined by turning tool developed in this study.

Fig. 12.

Comparison diagram of surface morphology. (a) The topographies of surface machined by non-optimized turning tool. (b) The topographies of surface machined by turning tool developed in this study.

Comparison of tool wear

Figure 13 shows a comparison of tool wear after machining. Compared to the non-optimized internal cooling turning tool, the turning tool developed in this study can more effectively suppress tool wear, reducing wear by about 25%.

Fig. 13.

Comparison diagram of tool wear. (a) Tool wear of non-optimized turning tool. (b) Tool wear of turning tool developed in this study.

By optimizing the internal liquid-nitrogen transmission channel of the tool, the heat loss of liquid nitrogen during transmission was reduced, cavitation effects subsequent to liquid nitrogen injection were mitigated, and the cooling efficiency of the liquid nitrogen was enhanced. Based on the experimental results above, compared to the non-optimized turning tool, the turning tool developed in this study shows improvements in cooling efficiency, machining quality, and tool wear. The cutting temperature is reduced by approximately 23%, the surface roughness value is decreased by about 20%, and the tool wear is reduced by roughly 25%.

Conclusion

This research studies the key parameters affecting the cooling efficiency of the liquid nitrogen internal spray cooling turning tool, proposes an integrated design method for the structure and function of the liquid nitrogen transport channel inside the turning tool, develops a liquid nitrogen internal spray cooling turning tool, and conducts experimental verification using TA15 titanium alloy as the research subject. The main conclusions are as follows:

1. The influence of different nozzle structures on jet flow is investigated. Both conical contraction nozzle and arc contraction nozzle exhibit velocity variations during the initial stage of jet formation. The expansion nozzle demonstrates the fastest temperature decay rate, while the straight nozzle maintains relatively stable velocity and temperature profiles. Considering the feasibility of manufacturing processes, employing a straight nozzle proves to be more advantageous.

2. The impact of the export bend angle on the velocity, volume fraction, and temperature of liquid nitrogen is analyzed. It is found that obtuse angles of the export bend are beneficial in reducing the impact of cavitation effects and are more conducive to ensuring the transmission efficiency of liquid nitrogen.

3. Compared to cylindrical insulation structures, designing the insulation sleeve of the main liquid nitrogen transmission channel as a “dumbbell” insulation structure can enhance the insulating capability of the sleeve, contributing to the reduction of liquid nitrogen loss during transmission.

4. An overall design method for a liquid nitrogen internal spray cooling turning tool is proposed, and a liquid nitrogen internal spray cooling turning tool is developed. With TA15 as the research subject, cutting experiments are conducted. The results show that, compared to a non-optimized turning tool, the use of an optimized liquid nitrogen internal spray cooling turning tool can reduce the cutting temperature by about 23%, decrease the surface roughness value by about 20%, and reduce the tool wear by about 25%.

Author contributions

Y.G.: Conceptualization, methodology, writing-original draft, funding acquisition. Z.D.: Software, data curation, investigation. H.Z.: Validation, writing-reviewing and editing. L.Z.: Formal analysis, visualization. G.X.: Data curation. Z.J.: Methodology, writing-reviewing and editing, funding acquisition.

Funding

This research is funded in part by Liaoning Province Doctoral Research Start-up Fund Program (No. 2024-BS-208); Liaoning Province Department of Education Basic Research Project (No. LJ212410158009); Special Funding Support for Basic Research Operating Expenses of Provincial Universities (No. 2024JBQNZ012); Dalian Ocean University Doctoral Start-up Project (No. 110224204). Natural Science Foundation of Jiangxi Province (No.20242BAB25267); Doctoral Research Start-up Foundation of Nanchang Hangkong University (No. EC202403148).

Data availability

Data is provided within the manuscript.

Declarations

Competing interests

The authors declare no competing interests.