Introduction

Contact fatigue refers to surface fatigue failure caused by long-term repeated action of contact stress on the workpiece surface1. Contact fatigue is the primary failure mode for bearings, gears, cold-working dies, etc. Cr12Mo1V1 is a high-carbon, high-chromium cold work die steel with high hardness, wear resistance, and good impact resistance2,3. It is widely used in metal-plastic forming processes such as spinning and cold shear, and its service life largely depends on contact fatigue properties. Cr12Mo1V1 contains up to 1.4% carbon (carbon is an austenitic stable element), making its martensitic transition end temperature Mf point as low as − 80 °C, and 10–30% residual austenite still exists in the material after conventional quenching treatment4. Residual austenite is an unstable soft Microstructure that fails to turn into martensite after quenching treatment. Under certain conditions, it will spontaneously transform into martensite, forming local additional internal stress, causing deformation and even cracking of components, thus affecting their contact fatigue properties.

The hardness of residual austenite is much lower than that of martensite, and its high content will cause the macroscopic hardness of components to decline. Many studies have shown that the volume fraction of residual austenite in martensitic steel significantly impacts the contact fatigue performance of components5,6. Under strain-induced conditions, residual austenite is easily transformed into brittle primary martensite, and the deformation discoordination between the residual austenite and the tempered martensite of the matrix will cause stress concentration and induce crack initiation7. Under periodic load, the high strain field at the crack tip will promote the continuous transformation of austenite to martensite, further aggravate the stress concentration in the crack tip region, and accelerate the propagation of fatigue cracks8. Therefore, controlling the residual austenite content significantly improves the contact fatigue performance of Cr12Mo1V1 steel.

Cryogenic treatment (DCT) is currently widely used as an effective method to reduce residual austenite content9,10. In this process, members cooled to room temperature after quenching continue to be placed in a medium lower than the Mf point temperature of their material to cool down further, promoting the continued transformation of austenite to martensite, which is generally regarded as the continuation of quenching11,12. Reasonable DCT can further reduce the residual austenite content and improve the hardness and wear resistance of the component13,14. The lower the residual austenite content, the more stable it is15,16. Shi Jianglong14 studied the effect of temperature on the residual austenite content of Cr12 steel DCT and found that the residual austenite content would be significantly reduced when the temperature was lower than − 70 °C.

The existing research mainly focuses on the influence of DCT on the residual austenite content and properties of components. In contrast, the effect of the combination process sequence of DCT and traditional heat treatment on components’ residual austenite content and fatigue properties has yet to be reported. In this paper, the effects of heat treatment combined with cryogenic treatment on the microhardness, residual stress, residual austenite content, and contact fatigue properties of Cr12Mo1V1 steel members were studied, and the mechanism of influence of DCT on the contact fatigue properties of Cr12Mo1V1 steel members was revealed.

Materials and methods

Test materials

Cr12Mo1V1 is a die steel with high hardness, wear resistance, and good cold working performance, and its composition is shown in Table 1. The material C content of up to 1.4% has a high hardness; it contains 12% Cr, which can improve the corrosion resistance; Mo can prevent second-class tempering brittleness and grain refinement can be promoted together with V.

Table 1 The elemental composition of Cr12Mo1V1 material(wt %).
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Design of test scheme

In this paper, the Φ10 × 110 mm standard contact fatigue component as shown in Fig. 1 is used to carry out the test. In order to reduce the error, 3 samples are prepared for each group. The residual austenite content, surface microhardness and residual stress of the component after DCT are tested, and the influence of DCT on its contact fatigue performance is analyzed in combination with its contact fatigue life test. All measurement results will be measured by \(\:\sigma\:=\sqrt{\frac{\sum\:{\left(x-\stackrel{-}{x}\right)}^{2}}{N}}\), to calculate the error value, so as to ensure the reliability of experimental data.

Fig. 1
figure 1

Schematic diagram of Cr12Mo1V1 steel sample.

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After quenching, Cr12Mo1V1 steel forms tough and brittle primary martensite, residual austenite, and undissolved carbide particles, and after tempering, primary martensite forms tempered martensite with high hardness and certain toughness. Standard heat treatment methods for Cr12Mo1V1 steel include low-temperature tempering and high-temperature tempering. The low-temperature tempering method is to heat the component to 980–1030 °C and then quench and temper at 200 ℃ to obtain tempered martensite, residual austenite, and undissolved carbide particles. The high-temperature tempering method is to heat the member to 1050–1080 °C for quenching, and 500 °C tempering, and the Cr element rich in the steel will precipitate in the form of alloy carbide and promote a large number of residual austenite into martensite to form a secondary hardening phenomenon, resulting in acicular tempered martensite, a small amount of carbide and residual austenite17,18.

In this experiment, The equipment used for quenching and tempering is HZCT2-65 double chamber vacuum gas-cooled oil quenching furnace. For cryogenic treatment, DJL-SLX766 liquid nitrogen mixing type deep cold box is used.To reduce the content of residual austenite in one step, this paper intends to compound DCT between quenching and tempering processes and compound DCT after the corresponding quenching processes of low-temperature tempering method (1030 °C quenching + 200 °C tempering) and high-temperature tempering method (1060 °C quenching + 520 °C tempering twice) respectively, In this experiment, we referred to GB/T 25,743 − 2010, and selected the cryogenic treatment temperature as − 150 °C and the treatment time as two hours. In the quenching process, the component is preheated at 650 °C × 40 min and 850 °C × 40 min and then heated to the quenching (target) temperature for 30 min before oil quenching, which is recorded as Q1 and Q2 according to the austenitising temperature of 1030 °C and 1060 °C respectively. The DCT process is to soak the component in liquid nitrogen at − 150 °C for 2 h and record it as D. The tempering process is the last step to adjust the structural properties of the member and is recorded as T1 and T2 according to the tempering temperatures of 200 °C and 520 °C. Table 2 shows the combined codes of the six different processing processes. Figure 2 shows the heat treatment process used in this study.

Table 2 Heat treatment test scheme.
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Fig. 2
figure 2

Schematic diagram of the typical time-temperature curve of the applied heat treatment cycle: (a) low-temperature and (b) high-temperature.

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Contact stress analysis

The Hertz contact theory is mainly used to solve the contact problem of elastomers. It points out that the linear contact stress is elliptically distributed on the half-width of the contact area, and provides formulas for calculating the contact half-width and maximum contact stress. When calculating the contact stress of the spinning wheel, it is necessary to determine the size and direction of the load borne by the spinning wheel in service condition. In view of the linear contact between the spinning wheel and the spinning belt wheel, the half-width of the linear contact area between the spinning wheel and the idler wheel can be calculated according to the known force conditions. Combined with the half-width and the radius of the spinning wheel, the elastic approximation of the spinning wheel on the contact surface is calculated. The Hertz linear contact schematic is shown in Fig. 3, setting the two points P1 and P2 on the cylinder to be equidistant from the Z-axis in the x-z plane. According to Hertz contact theory, the distance between P1 and P2 is determined by the radius of the two contact cylinders and the distance between P1 and P2 and the z axis.

Fig. 3
figure 3

Schematic diagram of the Hertz linear contact.

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$${h_1}+{h_2}={a^2}/(2r)$$
(1)
$$1/r=1/{r_1}+1/{r_2}$$
(2)

In the formula, r1 is the diameter of the upper cylinder and r2 is the diameter of the lower cylinder. When two cylindrical contacts are subjected to external load q, the convergence between them is δ and the contact width is 2a. In the interval -l ≤ x ≤ l, the normal displacement of the contact surface conforms to the deformation coordination equation shown in Eq. (3).

$${w_1}+{w_2}=\delta - ({h_1}+{h_2})=\delta - {a^2}/(2r)$$
(3)

Taking the derivative of the above equation a gives:

$$\frac{{d{w_1}}}{{da}}+\frac{{d{w_2}}}{{da}}= - \frac{a}{r}$$
(4)
$$w=\frac{{2(1 - {v^2})}}{{\pi E}}{\text{ }}\int {_{{ - l}}^{l}} P(s)ln|x - s|ds+C$$
(5)
$$\frac{{dw}}{{da}}=\frac{{2(1 - {v^2})}}{{\pi E}}{\text{ }}PV\int {_{{ - l}}^{l}} \frac{{P(s)ds}}{{x - s}}$$
(6)

Where E is the elastic modulus of the material and v is the Poisson ratio of the material. In the interval [- l, l] use formula (5), (6), the formula (4) can be written as formula (7), (8).

$$\frac{2}{{\pi E}}{\text{ }}PV\int {_{{ - l}}^{l}} \frac{{P(s)ds}}{{x - s}}=\frac{a}{r}$$
(7)
$$E=\frac{{(1 - v_{1}^{2})}}{E}{v_1}+\frac{{(1 - v_{2}^{2})}}{E}{v_2}$$
(8)

In order to balance v1 = v2 with the external load, the contact stress P(s) should satisfy the equation of equilibrium (9).

$$\int {_{{ - l}}^{l}} P(s)ds=q$$
(9)

Where q represents the line contact load distributed along the y axis, and its unit is N/mm. S represents the contact area under load. Formulas (7), (8), and (9) are the basic equations for solving the line contact problem. Hertz assumes the contact stress as a semi-elliptic function as shown in formula (10).

$$P(s)=\frac{{{P_0}}}{l}{\text{ }}\sqrt {{l^2} - {s^2}}$$
(10)

Put the formula (10) into the formula (9) and integrate to get the formula (11).

$${P_0}=\frac{{2q}}{{\pi l}}$$
(11)

By substituting Eqs. (10) into Eqs. (7) and (8), the formula (12) is obtained.

$$\frac{{2{P_0}}}{{\pi l{E^,}}}PV\int {_{{( - l)}}^{l}} \frac{{{l^2} - {s^2}}}{{a - s}}ds=\frac{a}{r}$$
(12)

The formula (12) is transformed by integral to obtain the formula (13).

$${\frac{{2{P_0}}}{{lE}}^,}a=\frac{1}{r}a$$
(13)

Through the formula (13), we can get the formula (14), formula (15).

$$l=\sqrt {4rq/\pi E}$$
(14)
$${P_0}=\sqrt {Eq/(\pi r)}$$
(15)

From formula (11) to formula (13), the relationship can be obtained as follows: (16).

$$l=2r{P_0}/E$$
(16)

When the two contacts have the same length L and the load is Q, the formulas (17) and (18) can be obtained.

$$l=\sqrt {4rQ/\pi LE}$$
(17)
$${P_0}=\sqrt {EQ/\pi Lr}$$
(18)

The required load is obtained by calculating the force of the member.

Detection method

According to ASTM E975-0319 and ASTM E2860-1220,21, the components’ residual stress and residual austenite volume fraction after an HDS-I X-ray analyzer measured DCT. FALCON 500 Vickers microhardness tester was used to load 100 g and hold 10 s to measure the microhardness distribution of the component. After grinding with 400–2000 mesh sandpaper and polishing with diamond suspension, the metallographic Microstructures of the corroded components were compared and observed by Keenes 3D confocal microscope.

RCTF-6000 rolling contact fatigue testing machine was used to test the contact fatigue performance, the basic principle of which is shown in Fig. 4. The component is clamped on the air spindle between two pressing idlers. The interaction between the two idlers makes the component bear no extra bending load and drives the two pressing idlers to roll purely through the element. The cylinder drives the clamping idler through the swing arm to exert contact load p on the component, as shown in Fig. 4a. The contact load p equals the cylinder thrust, and the applied load is detected by the pressure sensor installed on the cylinder. A vibration sensor is installed on the swing arm. The system will have a specific vibration when contact fatigue damage occurs in the spalling pit. When the amplitude reaches the set threshold, the test will automatically stop. The circular radius of the top of the idler is r. The radius of rotation is R. According to the Hertz contact theory and its hypothesis, there is an elliptical contact area between the idler and the member (the long and short axes are 2a and 2b, respectively), and the maximum contact stress can be calculated by formula 2–19. Where a is the contact half-length, b is the contact half-width, p is the total load, and p0 is the contact stress.To speed up the test process, a contact stress of 4 GPa and a spindle speed of 12,000 r/min are selected here. When there is a spalling pit, the vibration sensor detects that the swing arm’s amplitude reaches 10 μm and automatically stops. The appearance of the fatigue spalling pit was analyzed by Keenes 3D confocal microscope to reveal the contact fatigue failure mechanism of the component.

$${p_0}=\frac{3}{2}\frac{p}{{\pi ab}}$$
(19)
Fig. 4
figure 4

Contact fatigue test device (a) Schematic diagram; (b) Physical drawing.

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Results and discussion

Microstructure analysis

The microstructure of the Cr12Mo1V1 steel member after the above six processes is shown in Fig. 5, and the measured residual austenite content is shown in Fig. 7. It can be seen from Groups 4 (I) to 4 (III) that the matrix microstructure is all cryptocrystalline martensite, a certain amount of carbide are distributed on it22,residual austenite mainly exists in martensite. The measured residual austenite of the three Groups of components is 21.3%, 13.0%, and 12.5%, respectively. This indicates that adding the DCT process can significantly reduce the residual austenite content in the Microstructures. There are more precipitated carbides in Fig. 5 (II) and 4 (III), and the grain boundaries are more precise. From Fig. 5 (IV) to 4 (VI), it can be seen that the matrix martensite has a finer form and a large number of dispersed and precipitated fine carbide particles. The measured residual austenite of the three Groups decreased to 5.6%, 3.0%, and 2.6%, respectively. This is because as the tempering temperature increases, the residual austenite in the component continues to precipitate carbides, reducing its carbon content, increasing the Mf point, and promoting the further phase of austenite into martensite. Therefore, the residual austenite content of the 520 ℃ tempering treatment component of Group IV-VI is lower than that of the 200 ℃ tempering treatment component of Group I-III. DCT can also further reduce the residual austenite.

Fig. 5
figure 5

Microstructure of components under different processes (I) Q1 + T1; (II) Q1 + T1 + D + T1; (III) Q1 + D + T1; (IV) Q2 + T2 + T2; (V) Q2 + T2 + D + T2; (VI) Q2 + D + T2 + T2.

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According to the structural transformation process of the component shown in Fig. 6, in the low-temperature tempering method, the component is quenched at 1030 °C to obtain martensite, accompanied by massive carbides and many residual austenite. After cryogenic treatment, the residual austenite in the microstructure continues to transform into martensite. In the process of low-temperature tempering, a small amount of carbides will be precipitated, and the residual austenite content will be further reduced. Comparatively, in the high-temperature tempering method, the heating temperature of the component is higher, reaching 1060 °C, which makes the austenite quenching process obtain the higher supercooling degree and nucleation rate, and thus obtain smaller martensite and less residual austenite after quenching. After cryogenic treatment, the residual austenite content continued to decrease. During the high-temperature tempering process, a large amount of high-hardness alloy carbides will be precipitated from the microstructure, and the residual austenite content will be further reduced (Fig. 7).

Fig. 6
figure 6

Schematic diagram of component organization transformation. (a) Low-temperature tempering method to organize the transformation process content. (b) The high-temperature tempering method organizes the transformation process content.

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Fig. 7
figure 7

Residual austenite content on the surface of components under different processes.

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Surface hardness, residual stress

The surface microhardness values of Cr12Mo1V1 steel members with different processes were analyzed, as shown in Fig. 8. Compared with Groups I to III, it can be seen that the average microhardness of the surface of Group I is the lowest, which is 688 HV0.1. After DCT, the average microhardness of the component’s surface is increased, and the increase in the microhardness of the element corresponds to the decrease in the residual austenite content. The average surface hardness of the Group II component is increased to 700 HV0.1, and the surface microhardness of the Group III component is the highest, reaching 722 HV0.1. Compared with Groups IV to VI, it can be seen that the lowest surface microhardness of the components in Group IV is 680 HV0.1, and the surface microhardness of the components after DCT is improved, which also corresponds to the reduction of residual austenite content mentioned above23. The surface hardness of Group V was increased to 691HV0.1, while the surface microhardness of Group VI was the highest at 708 HV0.1. At the same time, it can be seen from Group III and Group VI that the DCT process followed by the quenching process is more conducive to improving surface microhardness. With the increase of tempering temperature after quenching, the lattice distortion of martensite decreases gradually, and its microhardness also decreases. In this test, due to the secondary hardening effect and the transformation of residual austenite, the microhardness of the components of Group IV to VI after 520 ℃ tempering is basically the same as that of Group I to III after 200 °C tempering but does not decrease significantly24,25. Moreover, the changes were consistent after DCT treatment.

Fig. 8
figure 8

Surface microhardness of components under different processes.

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The surface of the Cr12Mo1V1 steel member of different processes is measured with a test point every 5 mm along the axis direction for residual stress measurement, and the results are shown in Fig. 9. It can be seen from Groups I to III that the surface residual compressive stress of Group I is the largest. At the same time, the lattice distortion of martensite in Groups II and III will be relaxed to a certain extent due to the increase in tempering times, and the surface residual compressive stress will be reduced. In the transformation process from residual austenite to martensite, the difference in phase transformation sequence between the surface and the core leads to the tendency of tensile stress on the surface. The conversion of residual austenite to martensite usually results in a decrease in grain size and an increase in lattice distortion, which releases more residual stress26. At the same time, this transition may also lead to the reduction of surface residual stress. It can be seen from Groups IV to VI that DCT will promote the transformation of residual austenite to martensite. Due to the difference in phase transition sequence between the surface and the core, the surface will tend to produce tensile stress, and the surface residual stress will also be reduced, making the surface residual compressive stress of Groups V and VI significantly lower than that of Group IV. With the increase of tempering temperature, the lattice distortion of martensite will be further weakened, and the residual compressive stress will be further released, so the surface residual compressive stress of Group IV-VI is lower than that of Group I-III. Figure 10 shows a schematic model of the above residual stress transition.

Fig. 9
figure 9

Surface residual stress of members under different processes.

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Fig. 10
figure 10

Schematic diagram of residual stress transition in DCT process components.

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Contact fatigue performance analysis

The contact fatigue properties of Cr12Mo1V1 steel members after different thermal and cold composite processes were tested, and the results are shown in Table 3; Fig. 11. Comparing Group I to Group III, it can be seen that Group I has the highest surface residual compressive stress and the highest residual austenite content, but its contact fatigue life is the lowest 0.559 × 107 times, and the residual compressive stress is generally considered to be favorable to fatigue performance27. Therefore, it can be judged that the negative effect of excessive residual austenite on contact fatigue is more prominent. After the DCT, the contact fatigue life of the components is improved, and the highest contact fatigue life of the members of Group III is 1.065 × 107 times. Compared with Groups IV to VI, it can be seen that the contact fatigue life of the components in Group IV is the lowest 1.261 × 107 times, and the contact fatigue life of the components after DCT is increased. Among them, the contact fatigue life of Group VI is the highest 1.618 × 107 times.

Contact fatigue test results show that the residual austenite content is negatively correlated with the contact fatigue life, which is because the residual austenite is transformed into primary martensite under local high strain induction, and the stress mismatch between the residual austenite and tempered martensite forms stress concentration, which is easy to develop into a fatigue source and is not conducive to fatigue life28. By comparing Group I and Group IV, Group II, and Group V, it can be seen that the microhardness between Group III and Group VI is close, and the contact fatigue life of the latter is higher than that of the former. Therefore: (1) The DCT process can reduce the residual austenite content and help to increase the contact fatigue life; (2) The existing residual compressive stress is not enough to inhibit the phase transition of residual austenite and its adverse effects29; (3) After quenching, DCT and then high-temperature tempering can reduce the residual austenite content to the greatest extent so that the component can obtain the best contact fatigue performance30.

Table 3 Contact fatigue test life.
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Fig. 11
figure 11

Contact fatigue life of components under different processes.

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Contact fatigue test results

Macroscopic morphology of contact fatigue spalling pit

The Raceway morphology of Group I, III, IV, and VI components was photographed, and the results are shown in Fig. 12. Through comparative analysis of the raceway surfaces of Group I and Group III members, it can be observed from Fig. 12. that there are many tiny pits along the direction of spalling pit expansion on the surface. In contrast, the tiny pits are almost invisible on the other side. After further magnification, the original machining tool marks can be seen at the bottom of the pits, and the overall morphology characteristics of the two groups of components are similar. Comparing and analyzing the components of Group IV and Group VI, it can be seen from Fig. 12. that many pits with original machining tool marks are distributed along the raceway surface of Group IV. From the overall morphology of Group I, III, and IV spalling pits, it can be seen that they all have a large central spalling pit and have an apparent expansion trend along the rolling direction. Many pits in the expansion direction may be caused by debris formed during the spalling pit expansion during the contact fatigue test, which is extruded between the member and the idler. By further observing Fig. 12., it can be found that the spalling pit shape of member VI is more symmetrical than that of other Groups. Only one central spalling pit can be observed, and no obvious expansion trend along the rolling direction exists. However, several pits can also be found, but the number is far less than that of other Groups.

After the contact fatigue raceway depth detection, The measurement results shown in Fig. 13. are obtained as follows: the spalling pit depth of Group I and IV members is 141 μm and 135 μm, respectively, while the depth of Group III and VI members is 145 μm and 153 μm, respectively. The comparative analysis results show that the depth of spalling pits in Groups III and VI of DCT components is more profound than that in Groups I and IV of non-DCT components, mainly because the residual austenite content in DCT components is low, the Microstructure of DCT components is stable during repeated contact, and it is not easy to produce phase transition, which makes the formation threshold of cracks and spalling pits more significant, and the formation of spalling pits will be more profound. Thus, the damage formation in contact fatigue is reduced, and the contact fatigue life of the component is improved.

Fig. 12
figure 12

Raceway morphology of components under different processes (I) Q1 + T1; (III) Q1 + D + T1; (IV) Q2 + T2 + T2; (VI) Q2 + D + T2 + T2.

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Fig. 13
figure 13

Raceway depth of components under different processes.

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Microstructure of contact fatigue spalling pit

SEM was used to photograph the morphology of spalling pits of members in Groups I, III, IV, and VI, and the results are shown in Fig. 14. There is a more giant and complete spalling pit in Fig, and a relatively apparent expanding trend can be observed along the rolling direction of the idler. Many “lamellar” peeling features can be seen in the pit, and many surface cracks perpendicular to the moving direction can be observed along the idler moving direction. When these cracks expand and connect to form a division of the surface material, they separate from the substrate to create a new peeling block and further develop the existing peeling pit31.

According to the observation of the above four Groups of components, the spalling pit of Group I and Group IV shows a specific regularity. Along the expansion direction of the spalling pit, the spalling block is smaller and more dense, and the formed spalling pit is shallower and more accessible to expansion. This phenomenon is mainly because, without the high residual austenite content in DCT members, phase transformation occurs during the repeated contact process, resulting in additional internal stress with high amplitude. Under the combined action of surface stress concentration, cracks start and expand to form spalling holes, which shorten the contact fatigue life of the members. In contrast, the spalling blocks of Group III and Group VI members are more significant than those of non-DCT members. The number of cracks in the same area is less and sparser, indicating that after DCT, the formation thresholds of cracks and spalling pits are more significant and more profound, thus significantly improving the contact fatigue life of materials.

Fig. 14
figure 14

SEM graphs of samples treated by different processes. (I) Q1 + T1; (III) Q1 + D + T1; (IV) Q2 + T2 + T2; (VI) Q2 + D + T2 + T2.

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The morphology of the spalling pit of Group VI members was locally enlarged, and the results are shown in Fig. 15. In the Fig, the edge of the central pit has parallel striping features around the spalling pit. At the same time, the bottom can observe parallel striping features consistent with the direction of the machining tool marks (similar features exist in the other 5 Groups). These parallel striping features at the bottom are likely related to the shear stress. Combined with the contact stress analysis, it can be seen that the XZ plane shear stress is much larger than the XY plane shear stress, and the crack initiation is caused by the stress concentration caused by the superposition of the knife mark perpendicular to the Z direction. This indicates that the crack will likely be generated near a machining tool mark with apparent defects in the rolling process and then spread along the depth direction. Under periodic load, the outer ring of the spalling pit expands, forming a parallel strip characteristic around the spalling pit, making the spalling pit further expand and finally creating the spalling pit, as shown in Fig. In the crack growth process, some regions are repeatedly squeezed, and the banded features disappear and become relatively smooth and flat.

Fig. 15
figure 15

Spalling pit of Group VI members (1) Complete appearance of the spalling pit; (2) Partially enlarged view of the edge of the spalling pit; (3) and (4) Enlarged view of the bottom of the spalling pit.

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Contact fatigue spalling pit angle

The surface morphology of the spalling pit of the above four Groups of components was photographed, and the results are shown in Fig. 16. As seen from the Fig, there is a central spalling pit and an incomplete crack next to it. It can be judged that, due to the repeated contact process, cracks will be generated near the machining tool mark where the defect is more evident on the surface and will spread down at a certain Angle along the direction perpendicular to the machining tool mark. Some cracks form a cantilever Microstructure during the expansion process. During the continuous contact process, the cantilever Microstructure cannot withstand the contact force, so a piece of the cantilever Microstructure breaks off and forms a spalling pit. Other cracks in the expansion of contact with the weak stress concentration area will produce new cracks along the weak stress concentration spread, reach the surface, and then spall down, thus forming a spalling pit. It is worth noting that the residual austenite in the raceway is transformed into a primary martensite under repeated contact loads to form stress concentration, and the surface stress concentration effect will cause the low-stress load near the surface to rise sharply, making it easier to initiate cracks near the surface. Therefore, the surface stress concentration dramatically affects the initiation location and expansion Angle of crack initiation.

The spalling pit Angle of Group I and Group IV without DCT is 21° and 23°, respectively, while the spalling pit Angle of Group III and Group VI with DCT is 23° and 28°. The spalling pit Angle after DCT is more significant, and the depth is more profound. These results show that DCT can effectively reduce the member’s residual austenite content, improve the member’s stress concentration effect, make the crack propagation Angle closer to the maximum Hertzian contact shear direction, and thus form a more giant and independent spalling pit. It is helpful to improve the components’ surface properties and prolong the components’ contact fatigue life.

Fig. 16
figure 16

Spalling pit section under different processes (I) Q1 + T1; (III) Q1 + D + T1; (IV) Q2 + T2 + T2; (VI) Q2 + D + T2 + T2.

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Conclusions

This paper adds the cryogenic process based on the low-temperature and high-temperature tempering methods. The influence of the corresponding process on the residual austenite content, surface microhardness, residual stress, and contact life of Cr12Mo1V1 steel members is studied. The relationship between the residual austenite content and the contact fatigue life of the material is explored. The main findings can be summarized as follows:

  1. (1)

    DCT can promote the transformation of residual austenite to martensite, reduce the probability of initiation of cracks under the combined action of additional internal stress and surface stress caused by phase transformation, make the spalling pit formed in the process of repeated contact larger spalling block in the expansion direction. The number of cracks in the same area is smaller and smaller, thus increasing the threshold and depth of crack and spalling pit formation. It is helpful to improve the contact fatigue life of components by improving the Microstructure stability during service.

  2. (2)

    Compared with the low-temperature tempering method + DCT, the components of the high-temperature tempering method + DCT have a large number of strengthening phases formed by the precipitation of alloy carbides, which effectively reduces the residual austenite content while maintaining a hardness similar to that of the low-temperature tempering method, which helps to improve the service stability of the components.

  3. (3)

    In the rolling process, the residual austenite in the raceway is transformed into a primary martensite under the action of repeated contact loads, resulting in stress concentration. This surface stress concentration effect can significantly increase the stress load in the regions near the surface layer, thus promoting crack initiation at these locations. On the surface of the member, cracks are formed due to the superposition of shear stress and stress concentration. The crack spreads diagonally downward in the direction perpendicular to the machining tool mark, and in the process of expansion, cantilever structures are formed, or the crack spreads towards the weak point of stress concentration, forming a spalling pit.