Introduction

Cryogenically assisted machining is emerging as a clean, safe and environmentally friendly manufacturing process. Liquid nitrogen (LN2) and carbon dioxide (CO2) are the commonly used cryogenic fluids. In recent years, cryogenic machining has been adopted to improve cutting performance and surface integrity of stainless steels and other difficult-to-machine materials. Manimaran et al. (Ref 1) conducted grinding experiments on 316 austenitic stainless steel in three conditions, namely dry, wet and cryogenic cooling. They found that grinding with cryogenic cooling reduced cutting forces, produced lower surface roughness and fewer surface defects, compared to grinding with dry and wet cooling. Kumar et al. (Ref 2) studied the effects of cryogenic condition on AISI 4340 steel turning process. Their results revealed that the cryogenic turning not only reduced cutting forces, but also extended tool life. The surface integrity and corrosion resistance of the 316L austenitic stainless steel machined with cryogenic turning strategy have been investigated (Ref 3, 4). Results indicated that low-temperature coolants introduced more compressive residual stresses and suppressed the initiation of defects on the machined surface, helping to reduce the susceptibility of generalized and localized corrosion. These investigations suggested that the cryogenic machining was able to produce engineered surfaces while reducing damage generated from the machining process.

Recently, cryogenic machining has also been utilized as a processing method to achieve a plastic deformation-induced surface hardening (Ref 5). With cryogenic turning strategy in metastable austenitic stainless steels, the functional properties can be improved in terms of fatigue strength as well as wear resistance. Mayer et al. (Ref 6) investigated deformation-induced surface hardening of the metastable AISI 347 austenitic stainless steel with cryogenic turning. The observations showed that the mechanical properties of the machined surfaces were enhanced during the material removal process. The results also revealed that the deformation-induced surface hardening could be targeted via adjusting the turning process parameters and the geometry of tool cutting edge combined with low temperatures.

Cryogenic coolant could also lead to an increase in cutting forces and residual stresses, in particular the increase in milling forces. Aramcharoen et al. (Ref 7) investigated the machining performance of Inconel 718 in a milling process, and they observed that cutting forces in cryogenic cooling were higher than the conventional oil-based coolant. Nalbant et al. (Ref 8) reported that cryogenic milling of AISI 304 stainless steel resulted in higher cutting force and torque than those of dry milling. Dillon et al. (Ref 9) compared the plunge loads produced by milling 304 stainless steel at LN2 temperature, room temperature and high temperature. The obtained results showed that the plunge load was the highest at LN2 temperature. Park et al. (Ref 10) reported that the cutting force increased significantly as the Ti alloy hardened with the application of the LN2 during the cryogenic milling. De Paula Oliveira et al. (Ref 11) compared the results of machining forces, residual stresses and hardness in end milling of Inconel 718 using three types of cooling methods: flood, minimum quantity lubrication and LN2. The results showed that LN2 produced tensile residual stresses in longitudinal direction, whereas flood cooling generated compressive stresses. These observations indicate that the effects of low temperature on machining performance and cutting forces are complicated, even there are some controversial results related to the effect of the cryogenic machining on the residual stress and the cutting force in the published literature. The alteration of cutting force and temperature will have an impact on the residual stress distribution and microstructural modifications in the machined subsurface layer.

It is well known that residual stress and microstructure are two important factors related to SCC (Ref 12). Mechanical operations, such as milling (Ref 13), turning (Ref 14) and screw rolling (Ref 15, 16), always result in severe plastic deformation process. As a result, residual stresses and microstructural changes remain in the surface layer, which play a significant role in the mechanical property, corrosion resistance (Ref 17) and SCC susceptibility (Ref 18,19,20). Ghosh et al. (Ref 18) studied the effects of the machining on the SCC resistance of the 304L austenitic stainless steel, and the results revealed that the residual tensile stress induced by manufacturing operations reduced the SCC resistance. Mallick et al. (Ref 21) reported that cryogenic deformation would induce the formation of the martensite in 304 austenitic stainless steel. Martensite will increase the susceptibility of SCC due to the higher dissolving rate than austenite in corrosive environment (Ref 22). Despite several advantages of cryogenic machining mentioned above, whether the cryogenic machining operation is beneficial in improving SCC resistance needs further studies. However, little work has been dedicated to the influence of cryogenic milling on SCC resistance of austenitic stainless steels.

This paper aims to investigate the effects of cryogenic milling on SCC resistance of 316L austenitic stainless steel. The resultant changes are characterized in the subsurface properties due to cryogenic milling in terms of microstructure, residual stress and microhardness by comparing with conventional cooling milling. SCC tests of 316L stainless steel milled with cryogenic cooling and conventional cooling are conducted in boiling MgCl2, and we herein attempt to correlate the changes of surface microstructure and mechanical properties generated by cryogenic milling with the SCC initiation susceptibility of the machined surface.

Material and Experimental Procedure

Material and Machining Operations

The material studied was AISI 316L austenitic stainless steel with a chemical composition of 0.018 C, 1.57 Mn, 17.05 Cr, 12.58 Ni, 0.66 Si, 2.52 Mo, 0.021 S and 0.025 P (mass%). The workpieces for milling operation were pre-cut into a block of 30 mm × 70 mm × 6 mm by electrical discharge machining.

A vertical milling machine was used for the conduction of milling operations. To keep cutting edge sharp for each trail, a new carbide tool was used for each workpiece.

In the present study, the milling processes were conducted using two different cooling strategies, one was cryogenic cooling with LN2, and the other was conventional wet machining with emulsion. The LN2 was stored in a pressurized tank and was delivered through a special nozzle where the coolant was jetted into the cutting region of the workpiece. The experimental setup of LN2 supply is shown in Fig. 1. The machining parameters adopted in this study are listed in Table 1. Following the milling process, the workpieces were cut into small samples by electrical discharge machining, then cleaned and dried for residual stress measurement and SCC test.

Fig. 1
figure 1

Schematic view of cryogenic milling setup: (1) liquid nitrogen tank, (2) pressure gauge, (3) gate valve, (4) electromagnetic valve, (5) nozzle, (6) milling tool, (7) specimen, (8) machine table

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Table 1 Machining parameters for milling operations
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Residual Stress Measurement

The residual stress measurements were taken by means of the x-ray diffraction method on the milled surfaces (Ref 23, 24). X-ray radiation source of Mn-Ka with the applied voltage of 20 KV and the current of 4 mA were used to detect the austenite diffraction peak (311) which located at the angle of 152° (Bragg angle). To determine the residual stress, three measurements were taken in the parallel and perpendicular directions to machining striations.

Before the residual stress measurement, the center of each sample surface was located and marked under the digital microscope. During the measurement, the sample was placed on a cross-table. By adjusting the working position of the cross-table, the center of the tested surface (marked area) was moved to coincide with the light spot of the x-ray.

Examination of Microstructure, Microhardness and Martensitic Transformation

The near-surface microstructure in cross section was examined by a field emission scanning electron microscope (SEM). The sectioned samples were ground with 240, 400, 800, 1200 grit SiC papers in sequence, followed by polishing with 0.25-μm colloidal silica for 20 min. The applied force was 15 N. At last, the cross sections were etched using aqua regia.

The measurements of the microhardness were taken with an automatic Vickers hardness testers (Qness Q10A+) using a test force of 0.98 N and a holding time of 10 s for each indentation. For each sample, the microhardness measurements were taken in the cross section at various points of different depths. As a result, the microhardness profile could be obtained, showing the microhardness variation induced by machining.

X-ray diffraction (XRD) experiments were carried out with radiation using 40 kV operating voltage and 30 mA current to examine the martensitic transformation. Diffracting angle (2θ) was scanned from 30° to 100° with a scan rate of 10° /min.

SCC Test and Microcrack Examination

The SCC test was conducted in the solution of magnesium chloride boiled at 155.0 ± 1.0 °C. The SCC testing was referred to the standardized method ASTM G36. To prepare about 800 mL of boiling chloride solution, 1200 g of MgCl2-6H2O was weighed and added to a 2000-mL flask with 30 mL reagent water. To prevent bumping, the flask contained a thermometer and was covered with a layer of glass beads in the bottom. Heating was performed by an electric furnace with a temperature of 250 °C. In order to prevent the loss of water vapor during the boiling of the solution, a condenser cooling system was employed. During the heating, if there was any change in the concentration of the solution detected by any deviation in temperature, a 25 wt percent solution of MgCl2 was applied to adjust the concentration. When the solution was kept boiling at 155.0 ± 1.0 °C, the corrosive environment was ready for SCC test, and then the machined samples were exposed to the boiling MgCl2 solution for testing.

When the exposing time reached 30, 70, 150 and 300 minutes, the specimens were taken out, cleaned and dried out, and the density of the cracks initiated in the machined surface was measured under a digital microscope with magnification in the range of 200x - 1000x. The crack propagation in the cross section was observed under SEM.

The density of microcracks was calculated within the marked area of each specimen: (1) The marked area was divided into three regions of about 1500 μm × 500 μm. (2) For each region, the length of each crack large than 50μm was measured in the image area, and the lengths of the individual cracks were added to obtain the total length of the cracks. (3) The crack density of each region was calculated as the ratio of the total length of the cracks to the area of the measured image. (4) Based on the measured crack densities of the three regions, the average crack density and the standard deviation of three regions were obtained.

Results

Residual Stress

According to the residual stress measurements, the average surface residual stresses and standard deviation values in the cutting direction and perpendicular direction under different cutting conditions were obtained, as listed in Table 2. It is defined that residual compressive stress is a negative value and residual tensile stress is a positive value. In all samples, the residual stress had a higher value in the cutting direction than in the perpendicular one. Furthermore, in the wet milling condition, the tensile residual stresses were observed parallel to the cutting direction except for sample 1L, and the stresses became compressive in the perpendicular direction. However, the residual stresses of the cryogenic milling in the cutting direction were tensile in all samples, whereas in the perpendicular direction, the stresses were tensile or compressive depending on the machining parameters. It is worth noticing that the residual stresses showed a strong dependency on the cooling strategies. For instance, the residual stress in cutting direction of sample 3L cooled by cutting fluid was 78 MPa, while it increased up to 335 MPa for the sample 3N cooled by LN2. The similar phenomenon of residual stress was also observed in other groups. Therefore, it could be summarized that the cryogenic milling induced an increase of residual stress on the milled surface in both cutting and perpendicular directions.

Table 2 Residual stresses on the machined surfaces
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Microstructure and Microhardness of Deformed Layer

Figure 2 shows the average depth of slip bands in cross section under different cutting conditions. It can be observed that the depth of slip bands mainly depended on cooling strategies instead of cutting parameters. For cryogenic milling, the depth of slip bands reaches to 45 μm, but about 10 μm for wet milling. Figure 3 shows the micrographs of the samples in groups 3 and 5. The slip bands caused by a severe plastic deformation are clearly seen in all samples. In case of the cryogenic milling, a large number of slip bands in the subsurface layer within a depth of 40~60 μm were clearly visible, and the directions of the slip bands were different among the grains. Furthermore, there existed two different directions of slip bands in one grain, which was referred to as multi-line slip. As for the samples with the wet milling, there was a small amount of slip bands within a depth of 10 μm from the machined surface. The direction of the slip bands was about 45° to the machined surface, and little slip band intersection was observed.

Fig. 2
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Average depth of slip bands in cross section under different cutting conditions

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Fig. 3
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Micrograph of the deformed layer under different milling conditions: (a) sample 3N, (b) sample 3L, (c) sample 5N, (d) sample 5L

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X-ray diffraction patterns of the machined surfaces are shown in Fig. 4. The XRD pattern of the electropolished specimen reveals (111), (200), (220), (311), (222) peaks related to face-centered cubic (FCC) γ-austenite phase. Compared with the electropolished bulk material, no new diffraction peak was found in the milled surfaces, indicating that no detectable martensitic transformation was observed in both the wet and cryogenic milled surfaces.

Fig. 4
figure 4

XRD profiles of 316L SS under different milling conditions: (a) the wet milling, (b) the cryogenic milling

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Deformation induced by milling could affect the surface and subsurface hardness notably, as shown in Fig. 5. For comparison, the profiles of microhardness in the subsurface layer of samples 3N, 3L and 4N, 4L are presented. As can be seen from the figures, the microhardness value decreased from 350 HV0.1 on the surface to 165 HV0.1 in depth for the cryogenically milled samples and from 275 HV0.1 to 165 HV0.1 for the wet milled samples. On the other hand, the cryogenic milling led to a significantly deeper work-hardening layer compared with the wet milling. The penetration depth of the work hardening reached about 90 μm for the samples with cryogenic cooling, while the depth of the work hardening was 50 μm for the wet cooling ones. In addition, comparing the results in Fig. 3 and 5, the depth of hardened layer was larger than slip bands, indicating that machining-induced plastic deformation layer was deeper than slipping layer.

Fig. 5
figure 5

Profile of microhardness distribution under different milling conditions: (a) sample 3L, (b) sample 3N, (c) sample 4L, (d) sample 4N

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Stress Corrosion Crack Initiation

Figure 6 shows the variations of crack density with the exposure time. The crack density, which was used to evaluate the SCC initiation susceptibility of the machined surface, was sensitive to the machining cooling methods. For cryogenic milled samples except 1N, the crack density increased rapidly within the first 70 minutes. Then, the crack density increased with a lower rate or reached to a stable level. The variation of crack initiation on cryogenic milled sample 4N is shown in Fig. 7. Due to the high compressive stress in the perpendicular direction, the cracks mainly initiated perpendicular to the tool marks. As for the wet milled samples, no crack initiated within the first 70 minutes, and a few cracks initiated within the time of 70-300 minutes on sample 2L and 4L. The final surface crack density is summarized in Table 3. It can be seen that the cryogenic milled samples had much higher crack density than the wet ones except 1N and 1L.

Fig. 6
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The crack density variations with the exposure time at the marked points for Groups 1-5

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Fig. 7
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Microcrack variations in the sample 4N at four different exposure times: 30 min, 70 min, 150 min and 300 min

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Table 3 The density of surface microcrack
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The initiation rates in the first 70 min for each sample are summarized in Table 4. In the group 2, the microcrack initiation rates of samples 2N and 2L were 143 and 0 μm/(mm2·min). In the group 3, the samples 3N and 3L had the initiation rates of 197 and 0 μm/(mm2·min), respectively. In the group 4, the initiation rates were 66 and 0 μm/(mm2·min) corresponding to the samples 4N and 4L, respectively. In the group 5, 245 and 0 μm/(mm2·min) were the initiation rates of the samples 5N and 5L.

Table 4 Microcrack initiation rate of groups 1-5 in the first 70 min
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Figure 8 shows surface morphology of samples in groups 2, 3 and 5 after being exposed to the boiling MgCl2 solution for 300 minutes. A large number of cracks occurred on the samples milled with cryogenic cooling. Meanwhile, due to the high level of tensile stresses in both directions, a biaxial crack network appeared on the surfaces. On the surfaces milled with wet cooling, the relatively smaller number of cracks was observed on the surface of sample 2L along the perpendicular direction, but the length of cracks was short. There was no crack initiation on the samples 1L, 1N, 3L and 5L for their compressive or low tensile residual stresses in both directions. The results demonstrated that cracks were much easier to initiate on cryogenic milled surfaces under the same cutting parameter.

Fig. 8
figure 8

Surface morphology of 316L stainless steel machined under different milling conditions after exposed to the boiling MgCl2 solution for 300 minutes: (a) sample 2L, (b) sample 2N, (c) sample 3L, (d) sample 3N, (e) sample 5L, (f) sample 5N

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SEM images of the cross section of samples in groups 3 and 5 are displayed in Fig. 9 to analyze the characteristics of SCC propagation at the initiation stage. In the samples with cryogenic milling, the microcracks propagated into the subsurface layer within a depth of about 10-15 μm. The cracks appeared to be mainly the transgranular ones, growing predominantly along the slip bands, as is shown in Fig. 10. Moreover, the depth of cracks was obviously shallower than slip bands. As for the samples of 3L and 5L milled with the wet coolant, cracking propagation was absent due to the fact that no crack initiated on the machined surface.

Fig. 9
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Micrograph of the subsurface layer in samples after exposed to the boiling MgCl2 solution for 5h: (a) sample 3N, (b) sample 3L, (c) sample 5N, (d) sample 5L

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Fig. 10
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The transgranular cracks in sample 5N showing correlation of crack propagation with the slip bands

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Discussion

Effects of Cryogenic Milling on Microhardness, Microstructure and Residual Stress

A significant increase in microhardness was observed in the near-surface region of the milled samples, regardless of the cooling strategy. The increase in hardness was approximately 65% for the surfaces with wet milling and 110% for surfaces with cryogenic milling. In the present study, the deformation-induced α’-martensite formation was detected in none of the machined surfaces. Therefore, the increase in microhardness was due to the strain hardening induced by plastic deformation. The increased dislocation density, grain refinement and formation of twins contributed to the surface hardening. Similar results were reported in other investigation (Ref 25). In addition, the cryogenic milling produced deeper hardened layer compared with the wet milling. On this basis, it can be summarized that the cryogenic milling induced more severe work hardening than the wet milling in 316L austenitic stainless steel.

The microstructure change is one of the main impacts introduced by the cutting process and also an important factor affecting the stress corrosion resistance. The microstructure of the deformed layer in Fig. 3 shows that compared with wet milling, more intensity of slip bands were generated in the subsurface layer after cryogenic milling. This indicates that the cryogenic milling tends to generate more severe plastic deformation than the wet milling in 316L stainless steel. Except for this, under the same cooling strategy, the depth of slip bands was similar regardless of cutting parameters, indicating that the cooling strategies have a more important impact on machining-induced plastic deformation than cutting parameters. There also appears to be a difference in the nature of the plastic deformation induced by the cryogenic milling and the wet milling, since the former produced intersection of slip bands, whereas the wet milling produced slip bands along the same direction.

The residual stress generated by milling process depends on several factors, including but not limited to the properties of the material, machining process parameters and tool geometry (Ref 26,27,28), etc. However, numerous experimental studies revealed that the surface roughness, residual stress were reduced to a great extent with the application of cryogenic cooling (Ref 4, 29). However, in this study, the cryogenic milling of 316L stainless steel produced tensile stresses in both the parallel and perpendicular directions to machining striations except for samples 1N and 4N in the perpendicular direction. The wet milling produced compressive stresses in the perpendicular direction, whereas the stresses in the parallel direction were still tensile except for specimen 1L. These results indicated that cryogenic milling could result in an increase in residual tensile stress or decrease in residual compressive stress in both directions. The increase in residual stress was also reported by De Paula Oliveira et al. (Ref 11) when cryogenically milling Inconel 718. The residual stress induced by cryogenic milling was great when the cutting speeds were high. The reason why the residual stresses appeared to be an increase in the present cryogenic milling process is possibly associated with the mechanical properties of 316L stainless steel, cryogenic cooling strategies and machining parameters.

The effects of cryogenic machining on residual stress are complicated. Lee et al. (Ref 30) reported that a significantly discontinuous yielding phenomena or second hardening existed for 304 and 316 austenitic stainless steels at the cryogenic temperature range. After initial yielding, there is a critical strain value for the starting of second hardening. The material strength significantly increased during the second hardening. According to an analytical model (Ref 14, 31), a simulation for oblique cutting was conducted at room temperature (20 °C) and low temperature (− 50 °C) to analyze the effects of yield strength on cutting forces and residual stress. The cutting force (Fc), thrust force (Fp), lateral force (Fr) and surface residual stresses were obtained, as shown in Fig. 11. Due to a higher yield strength at low temperature (Ref 30), cryogenic machining could result in larger cutting forces. The simulation also exhibits that the surface residual stresses are higher in cryogenic machining condition than those in wet machining condition. Therefore, it is considered that the alteration of material yielding strength at low temperature would induce the increase of cutting forces and residual stresses.

Fig. 11
figure 11

Comparison of cutting forces and residual stresses at room temperature and low temperature

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In summary, compared with the wet milling, the change of residual stress in the cryogenic milling was caused by the following factors: (1) The increased material yielding strength due to the second hardening at lower temperature resulted in higher cutting forces (Ref 30). On the other hand, with the conventional coolants, the temperature was higher in the primary shear zone, resulting in thermal softening during plastic deformation. (2) The altered cutting forces due to cryogenic cooling affected the residual stress distribution on the milled surface. The increase of residual tensile stress accounted for the increased cutting force experienced in machining 316L stainless steel at the cryogenic temperature (Ref 7, 8). (3) Cryogenic machining introduced large thermal gradients on the tool and subsurface of workpiece (Ref 32), which altered friction condition between the tool and workpiece, and in turn affected the wear of the tool. Changes in the tool edge geometry might affect cutting forces and cutting temperatures. As a result, the stress distribution in the subsurface layer was altered as a consequence of the related local modification of cutting forces and temperatures.

Correlation Between Milled Surface Properties and Stress Corrosion Resistance

The SCC test reveals the cracking susceptibility of the 316L stainless steel machined surface in the chloride medium MgCl2. The residual tensile stress of machined surface in the cutting direction was much higher than that in the perpendicular direction. Correspondingly, the cracking direction was mainly perpendicular to the machining striations as shown in Fig. 8. This indicated that the initiation of microcracks was related to the residual tensile stress level (Ref 33).

The cryogenic milling increased the residual tensile stress in both the cutting and perpendicular directions compared to the conventional milling under the same cutting parameters. Correspondingly, the crack densities on the surfaces with the cryogenic cooling were higher than with the conventional fluid cooling, and the initiation rates were higher for cryogenic milled surfaces. As presented in Tables 2 and 3, the samples 2N and 3N suffered higher residual tensile stresses in both cutting and perpendicular directions; the higher crack densities were detected in the two samples. However, it should be noted that the sample 4N had the highest residual tensile stress in the direction parallel to machining striations, but its crack density and initiation rate were not the highest one. The reason why the sample 4N did not appear the highest crack density was possibly associated with the high compressive stress in the perpendicular direction. Moreover, the samples 2L and 5L had high residual stresses close to sample 5N in cutting direction, but few cracks initiated on the surfaces. This further indicates the initiation and growth of microcracks depend on the biaxial stress state in addition to the maximum tensile stress level (Ref 20). The samples 1N and 1L were free from cracks because the maximum residual stresses did not reach the critical value for SCC initiation (Ref 33).

The initiation of SCC is not only related to residual stress level, but also affected by the microstructure in the subsurface. As a main microstructural feature of the machining-induced plastic deformation, the slip bands in the subsurface have an important effect on the initiation and expansion of SCC.

The SEM results shown in Fig. 3 indicated that the cryogenic milling led to more intensity of slip bands below the machined surface compared to the wet milling. In Fig. 9 and 10, the cracks of specimens milled with cryogenic cooling initiated and extended along the slip bands. As for the wet milled samples, the number of slip bands was less than the cryogenic milled ones, and no crack was observed to extend into the deformed layer. The anodic dissolution mechanism of SCC believes that the metal dissolution and oxidation may enhance in the slip line (Ref 34), promoting the crack propagation along the slip bands under tensile stress. Therefore, the increasing number of the slip bands in the subsurface caused by the cryogenic milling promoted the expansion of the transgranular crack in the depth direction and reduced the stress corrosion resistance of 316L austenitic stainless steel. Nevertheless, it should be noted that the propagation of cracks was terminated at the depth of 10-15 μm, far below the depth of slip bands. Similar experimental results were also reported in another investigation (Ref 35).

In addition, Raquet et al. (Ref 36) observed that the increase in surface hardness caused by cold work significantly increased the susceptibility of SCC. Therefore, the severe hardened layer caused by the cryogenic milling promoted the crack expansion and reduced the stress corrosion resistance of 316L austenitic stainless steel.

It should be noted that in cryogenically assisted machining, the effects of low temperature depend highly on the method of LN2 supplied in the cutting zone (Ref 10). In this study, the LN2 was directly jetted onto the milling region of the workpiece. Further investigations will include the effects of cooling setting on the stress and temperature fields during cryogenic milling.

Conclusions

In this paper, the effects of cryogenic milling on stress corrosion cracking resistance of 316L austenitic stainless steel were investigated in comparison with the wet cutting condition; the main conclusions are listed as follows:

  1. 1.

    Compared with the conventional fluid-cooling milling, the cryogenic milling results in a more severe plastic deformation in the subsurface layer, increasing the slip band density as well as the microhardness. Meanwhile, the low cooling temperature leads to more tensile residual stress on the machined surface of 316L austenitic stainless steel.

  2. 2.

    The cryogenic milling increases the SCC initiation susceptibility to the aggressive Cl- environment of the machined surfaces, which can be explained by the fact that the cryogenic milling induces higher residual tensile stresses, severe hardened layer and promotes the formation of slip bands in the subsurface.