Dry Hard Turning versus Grinding—The Influence of Machining-Induced Surface Integrity on Fatigue Performance

Abstract

Dry hard turning (DHT) provides an effective process for finishing high-hardness materials. Machining-induced surface integrity has a direct impact on functional performance. This study compares the effects of the DHT and grinding processes on machining-induced surface integrity and fatigue performance of 18CrNiMo7-6 steel. The DHT and grinding experiment were carried out by using a polycrystalline cubic boron nitride tool and corundum wheel, respectively. The 3D surface morphology, surface roughness, surface residual stress, and machining accuracy of the hourglass-shaped specimen were measured. The fatigue fracture was characterized by scanning electron microscopy. The experimental results show that compared to grinding, DHT has obtained a larger surface compressive residual stress (the maximum axial and tangent residual stresses are −762.6 MPa and −442.8 MPa, respectively) and a lower surface roughness (the minimum Ra and Rq are 0.172 μm and 0.230 μm, respectively). This study is an attempt to use DHT instead of a grinding process to finish 18CrNiMo7-6 steel, providing a reference for high-quality and sustainable manufacturing of hardened steel.

1. Introduction

Dry hard turning (DHT) technology is widely used to process materials with a hardness above 45 HRC [1,2,3]. When the DHT process is applied in the finishing operation, a high cutting speed and large depth of cut are generally used to remove materials, and ideal machining quality and accuracy are also obtained [3]. Due to their high hardness and low iron affinity, cubic boron nitride (cBN), polycrystalline cubic boron nitride (PcBN), and ceramic tools are frequently used for DHT-hardened steel [4,5]. Before the emergence of DHT technology, the finishing operation of hardened steel was generally completed by a grinding process. Under certain circumstances, DHT has advantages over grinding in process flexibility and sustainability [4,5,6]. On the one hand, the grinding process has restrictions on the workpiece with a special shape, and the grinding wheel should be trimmed according to the shape of the components. On the other hand, compared with grinding, DHT has a higher material removal rate and a higher surface residual compressive stress [1]. The application of ultrahard tools is conducive to achieving sustainable manufacturing while saving on cutting fluid costs.

Surface integrity has a significant impact on the functional performance (including fatigue performance) of components [7,8,9,10]. Meanwhile, the application of DHT to improve the surface integrity of hardened steel components in the finishing process has been widely studied by academic and industrial circles [11,12]. Tang et al. [13] carried out the finishing operation of AISI D2 hardened steel by using the DHT process, and the experimental results showed that the minimum surface roughness Ra is 0.34 μm when the tool corner radius is 0.8 mm, which can achieve the surface finish of the grinding process. Holubjak et al. [14] found that when the cutting speed of DHT high carbon steel is 100 m/min, the axial surface residual stress is up to −850 MPa. Kundrak et al. [15] compared DHT and grinding processes and found that the surface roughness obtained by the two processes was not much different, and the former had advantages in some surface integrity characteristics. Martell et al. [16] analyzed the surface residual stress of DHT and grinding AISI 1053 hardened steel; the DHT process obtained higher surface compressive residual stress in the cutting direction. Asutosh et al. [17] established the surface roughness model of the DHT process and pointed out that the tool corner radius and feed rate are the main factors affecting the surface roughness. Sarnobat et al. [18] analyzed the influence of tool edge geometry and cutting parameters on surface residual stress and roughness during the DHT process; the optimal cutting parameter range is obtained for residual stress optimization. Smith [19] carried out the DHT process of AISI 52100 hardened steel and found that tool flank wear directly affected surface residual stress. Umbrello et al. [20] found that optimizing cutting parameters and tool geometry parameters is beneficial to improving the distribution of residual stress. It is generally believed that a reduction in surface roughness and an increase in surface compressive residual stress are beneficial to improve the fatigue performance [21,22,23].

18CrNiMo7-6 steel is widely used in aviation, automobile, electric power, and other industrial fields [24,25]. The technologies for the surface-strengthening treatment of 18CrNiMo7-6 steel include carburizing [26,27,28], shot peening [29,30], waterjet peening, surface rolling, etc. These surface-strengthening treatment methods can induce microstructural changes, produce compressive residual stress, and improve fatigue properties. Springer et al. [25] studied the effect of dynamic recrystallization and phase transformation on the microstructure of 18CrNiMo7-6 steel. The results show that the softening effect shown by the stress–strain curve is caused by the recrystallization of austenite.

The surface modification layer of 18CrNiMo7-6 steel was formed after carburizing heat treatment. The physical properties of the surface modification layer have a gradient change along the depth direction, which has attracted the attention of many scholars. Wang et al. [26] investigated the influence of residual stress and microstructure caused by carburizing heat treatment on the corrosion resistance of 18CrNiMo7-6 steel. The results show that with the increase in the depth of the carburized layer, the residual compressive stress in the carburized layer increases at first and then decreases and reaches stability in the matrix material. Electrochemical tests showed that the electrochemical impedance and electrochemical potential increased with the reduction in residual compressive stress. Zhao et al. [27] proposed and verified a method for measuring the elastic modulus of the surface modification layer using a micro-pressure instrument. The elastic modulus and hardness of the surface modification layer of 18CrNiMo7-6 steel after carburizing heat treatment in different test directions under different loads and different indentations were studied. The results show that carburizing heat treatment can hardly change the elastic modulus of the surface modification layer of 18CrNiMo7-6 steel along the depth. They also obtained the empirical formula for the gradient hardness of the surface metamorphic layer in top and cross-section tests, which can be used to evaluate the hardness of different depths. At the same time, the elastic modulus and hardness obtained on the cross-section are higher than those on the top surface. Zhang et al. [28] systematically studied the effect of the microstructure of 18CrNiMo7-6 steel on its crack propagation path by using electron backscatter diffraction and scanning electron microscopy techniques; The results showed that the prior austenite grain boundary and martensite boundary hindered the crack growth. The crack growth path is affected by the microstructure of flat needles martensite, original austenite grain boundary, and crystal orientation. Ho et al. [29] studied the potential relationship between surface integrity caused by the shot peening process and fatigue life. The effect of shot peening parameters on the fatigue life of 18CrNiMo7-6 steel was evaluated by characterizing the surface integrity and carrying out relevant fatigue tests. The experimental results show that using large-diameter pellets followed by using small-diameter pellets is an effective method to improve surface integrity and fatigue performance.

To the best of the authors’ knowledge, few studies have directly compared the effects of DHT and grinding on the surface integrity and functional performance of 18CrNiMo7-6 steel. Although the DHT process has been widely used in the finishing of high-strength steel, the feasibility of DHT 18CrNiMo7-6 steel after carburizing heat treatment (~60 HRC) has not been studied.

In this study, the DHT and grinding processes of 18CrNiMo7-6 steel are compared. The objectives of this study are as follows: (1) explore the feasibility of using DHT to process 18CrNiMo7-6 steel; (2) optimize the cutting parameters during finishing 18CrNiMo7-6 steel; and (3) analyze the influence of surface integrity on fatigue performance. Section 2 introduces the material and dimensions, processing parameters, and characterization evaluation of fatigue specimens; Section 3 introduces the corresponding surface integrity and fatigue performance results; and Section 4 summarizes the overall performance of the DHT process. This study is an attempt to replace grinding with the DHT process, which provides the potential for high-quality and sustainable manufacturing.

2. Experimental Procedures

2.1. Materials

Before carburizing heat treatment, the 18CrNiMo7-6 steel sample is pre-shaped. The material was kept at 880 °C with 1.1 and 0.9% carbon potential for 3 and 1 h, respectively, after which it was tempered at 650 °C for 4 h. Then quenched at 835 °C for 2.25 h and deep cooled at −60 °C for 2 h, and finally tempered at 180 °C for 12 h. The carburized layer depth of the specimen is 0.70–0.80 mm, the surface hardness is ~58–63 HRC, and the center of the specimen retains sufficient plasticity and toughness. The composition of the 18CrNiMo7-6 steel after carburizing heat treatment is shown in

Table 1. Chemical composition of 18CrNiMo7-6 steel after carburizing heat treatment (wt%).

Element	C	Si	Mn	S	P	Cr	Ni	Mo	Fe
Content	0.15–0.21	0.15–0.40	0.60–0.90	<0.012	<0.020	1.50–1.80	1.40–1.70	0.25–0.35	Bal.

.

International Organization for Standardization divides the shapes of rotating bar bending fatigue specimens into cylindrical, tapered, and hourglass-shaped [31]. The hourglass-shaped specimen is selected in this study, and its dimensions are shown in Figure 1.

2.2. Dry Hard Turning (DHT) Process

The DHT experiment was executed on a CAK4085 (Shenyang, Liaoning, China) computer numerical control (CNC) lathe. The PcBN cutting tool (BN-H10 VNGA160412-2S) was selected for the DHT process. The geometric parameters of the cutting tool are a rake angle of 0°, a cutting edge angle of 72.5°, and a clearance angle of 0°. The model of the tool holder is MVVNN2020K16. The schematic diagram and cutting parameters of the DHT process are shown in Figure 2 and

Table 2. Cutting parameters of DHT process.

No.	Spindle Speed n (r/min)	Cut Depth ap (mm)	Feed Rate vf (mm/min)
1	900	0.15	30
2	1050
3	1200
4	1350
5	1500
6	1200	0.05	30
7		0.10
8		0.15
9		0.20
10		0.25
11	1200	0.15	10
12			20
13			30
14			40
15			50

, respectively. For each turning experiment, a new cutting tool was applied.

2.3. Grinding Process

The grinding experiment was performed on an MKE1620A (Shanghai, China.) CNC grinding machine. The grinding mode was radial-feed grinding. The grinding depth was 0.3 mm. The grinding wheel was an 80 # grain size chrome corundum grinding wheel with an outer diameter of 450 mm and an aperture of 203 mm. The grinding wheel thickness was 30 mm, and the bond type is a ceramic bond. The coolant is 3310CIS, a water-based grinding fluid with a flow rate of 10 L/min. The dressing method of the grinding wheel is diamond pen dressing, with a cutting depth of 7 μm and a feed speed of 200 mm/min. The schematic diagram and cutting parameters of the grinding process are shown in Figure 3 and

Table 3. Cutting parameters of grinding process.

No.	Workpiece Speed nw (r/min)	Feed Rate vfr (mm/min)	Spindle Speed (r/min)
1	60	0.3	1488
2	80
3	100
4	120
5	140
6	100	0.1	1488
7		0.2
8		0.3
9		0.4
10		0.5

, respectively. For each grinding experiment, the grinding wheel was redressed between trials.

2.4. Surface Integrity Characterization and Fatigue Testing

The surface morphology and roughness were measured by the Bruker NPFLEX (Billerica, MA, USA) 3D surface measurement system. The surface roughness is expressed by Ra and Rq. Three sets of data were measured for each sample and then averaged. The axial and tangent (cutting direction) residual stresses were measured by the Proto LXRD (Vancouver, British Columbia, Canada) X-ray residual stress analyzer with Cr-Kα radiation. The tube voltage, tube current, and diffraction angle are 30 kV, 25 mA, and 156.1°, respectively. Three sets of data were measured for each sample and then averaged.

The basic principle of measuring internal residual stress by X-ray diffraction is to measure the diffraction displacement as the original data. The measured result is actually residual strain, and the residual stress is calculated from the residual strain by Hooke’s law. When there is residual stress in the sample, the crystal plane spacing will change. When Bragg diffraction occurs, the diffraction peak will also move, and the size of the shift distance is related to the value of the stress.

For the machining accuracy of the specimen, the 3D size was measured by a Keyence VHX-7000 (Osaka, Japan) microscope system, and the radial run-out was detected by an EWAG WS11 (Dübendorf, Switzerland) universal tool grinder and dial indicator.

The rotating bar bending fatigue test was conducted using the QBWP-6000J (Changchun, Jilin, China) fatigue test machine. The four-point loading is selected according to ISO standard [31], as shown in Figure 4. The fatigue fractography was observed by using a Zeiss Auriga (Oberkochen, Germany) scanning electron microscope.

The applied force F can be calculated by the following equation:

$$F=S \frac{\pi d^3}{32L}$$

(1)

where S is the required test stress; L is the force arm length; d is the specimen diameter; and M is the bending moment. In this study, L₁ = L₂ = 200 mm, the stress ratio R = −1, and the test frequency is 50 Hz.

3. Results and Discussion

3.1. Surface Roughness and Morphology

The 3D surface morphology of the specimens after DHT and grinding is shown in Figure 5. The red areas represent the peaks, and the blue areas represent the valleys. Compared with the grinding process, the DHT process produces a smoother machined surface and more uniform feed marks.

Figure 6 shows the effects of DHT process parameters (n, a_p, and v_f) on the surface roughness. In general, the surface roughness decreases with the increase in n. This is because the increase in n can reduce the built-up edge on the machined surface, and reduce the plastic deformation of the workpiece, thus reducing the surface roughness. The increase in a_p leads to more plastic deformation and larger surface roughness. With the increase in v_f, the remaining volume of the material that has not been removed increases, and the height of the built-up edge and the machining burr increases, so the surface roughness deteriorates. The minimum surface roughness Ra obtained by the DHT process is 0.172 μm.

Figure 7 shows the effects of grinding process parameters (n_w and v_fr) on the surface roughness. From the overall trend, the surface roughness increased with the increasing n_w. As the n_w increases, the contact arc length between the wheel and the workpiece increases, which increases the material removal rate and the maximum undeformed chip thickness. The surface roughness initially decreased and then increased with the increasing v_fr. The increasing v_fr during plunge-grinding increases the wheel radial feed per unit time, resulting in increased friction between the abrasive particles and the workpiece and increased plastic deformation of the grinding layer material, so the surface roughness value gradually increases. The minimum surface roughness Ra formed by the grinding process is 0.314 μm.

3.2. Residual Stress

Figure 8 shows the effects of DHT process parameters (n, a_p, and v_f) on the surface residual stress. From an overall viewpoint, the axial stress σ_y first decreased and then increased with the increasing n. As the n increases, the mechanical stress and plastic deformation effect increase, and the surface compressive residual stress finally increases [32,33,34,35]. An increase in a_p increases the cutting layer volume and the plastic deformation zone, so σ_y increases. The axial stress σ_y tends to decrease with the increasing v_f. As the v_f increases, the material removal rate increases, and the friction between the workpiece and tool increases, resulting in more cutting heat and aggravating the effect of thermal stress. The maximum σ_y and σ_x obtained by the DHT process are −762.6 MPa and −442.8 MPa, respectively.

Figure 9 shows the effects of grinding parameters (n_w and v_fr) on the residual stress. The effect of n_w on residual stress is not monotonic, and the maximum compressive residual stress is formed when n_w = 80 r/min. As the n_w increases, the average undeformed chip thickness and grinding heat increase, while the continuous increase in n_w makes the heat source move faster along the machined surface and reduces the effect of thermal stress, so the residual stress fluctuates. The compressive residual stress is the largest when v_fr = 0.1 mm/min and then tends to decrease. As the v_fr increases, the friction between the wheel and workpiece intensifies, the grinding heat increases, and then the thermoplastic deformation formed by the grinding heat increases, so the compressive residual stress decreases under the effect of thermal stress. The maximum σ_y and σ_x formed by the grinding process are −456.2 MPa and −221.2 MPa, respectively.

According to the surface roughness and residual stress values, the optimal parameters of different processes are determined, as shown in

Table 4. Optimized process parameters and surface integrity.

Machining Processes	Optimum Parameters	Surface Roughness	Surface Residual Stress
DHT	n = 1500 r/min ap = 0.15 mm vf = 30 mm/min	Ra = 0.174 μm Rq = 0.230 μm	σy = −762.6 MPa σx = −393.6 MPa
Grinding	n = 1488 r/min nw = 100 r/min vfr = 0.1 mm/min	Ra = 0.352 μm Rq = 0.440 μm	σy = −406.5 MPa σx = −168.4 MPa

. The surface roughness and residual stress produced by the DHT process are better, which is related to the axial feed of the cutting tool. At the same time, the larger corner radius of the PcBN tool is beneficial to reduce the surface roughness and increase the surface compressive stress. Tool geometry significantly affects the surface roughness of the machined surface [36,37].

3.3. Machining Accuracy

The processed specimen is shown in Figure 10, and its radial run-out and 3D size accuracy were tested. The effect of different processes on the radial run-out of the specimen is shown in Figure 11, and the abscissa (No.) corresponds to the different parameters in Table 2 and Table 3. The radial run-out produced by DHT and grinding processes fluctuated around the fitting lines of 0.02–0.04 mm and 0.05–0.06 mm. The machining accuracy of the specimen meets the run-out tolerance requirement of dynamic radial run-out ≤0.06 mm of fatigue test. Compared with the grinding, the machining accuracy of the sample obtained by the DHT process is better.

The 3D size of the specimens was randomly tested to further evaluate the machining accuracy of the DHT process. The specimen, cloud image, and size of the 3D measurement are shown in Figure 12a–c, respectively.

The length, radius, and difference in the height of the specimen arc section are 20.294 mm, 36.033 mm, and 14.57 mm, respectively. The length deviation, radius deviation, and height deviation of the specimen arc section processed by the DHT process are 0.8%, 0.1%, and 2.8%, respectively. The machining accuracy can meet the size requirements shown in Figure 1.

3.4. Fatigue Test and Fractography

The fatigue specimens were processed according to the optimum parameters in Table 4, and then the fatigue test was conducted. The rotating bar bending fatigue test refers to the ISO standard [31], and the test stress determined after the pre-experiment is 1400 MPa. The bending fatigue life of the specimens under different processes is shown in

Table 5. Bending fatigue life of specimens processed by different machining processes.

No.	Machining Processes	Test Stress/MPa	Fatigue Life/N	Fatigue Fracture (Yes/No)
1	DHT	1400	1,000,000	N
2			1,000,000	N
3			284,588	Y
4	Grinding	1400	1,000,000	N
5			1,000,000	N
6			128,504	Y

.

After the fatigue fracture, the fatigue fractography of specimens is analyzed. The fatigue fractography of specimens after the DHT and grinding processes are shown in Figure 13 and Figure 14, respectively.

Figure 13 and Figure 14a show that the fatigue fractography of specimens is divided into three zones, including fatigue crack initiation, crack propagation, and rapid fracture. There are many tiny strip-shaped cracks and crack growth in the crack initiation zone. The crack propagation zone is relatively smooth, and the rapid fracture zone is a ridge-like pattern with obvious plastic deformation. Figure 13 and Figure 14b show the crack initiation zone magnified by 50×, with obvious tiny cracks and some inclusions in the fracture. Figure 13 and Figure 14c show machining marks and burrs on the surface of the fractured samples, and there is a relatively apparent large crack near the machining marks. The machining defects and marks on the specimen surface will cause large stress concentrations and become the origin of cracks.

The crack propagation zone of the DHT-processed specimen is characterized by river patterns in Figure 13d, indicating that this zone is a cleavage fracture and has high brittleness. The characteristics of the crack propagation zone are relatively smooth and have almost no plastic deformation, indicating that the crack propagation zone of fatigue is a brittle fracture. The crack propagation zone in Figure 14d is characterized by fractures along the grain boundaries of different orientations, which are crystalline and stone-shaped, and belong to intergranular fractures. There is no obvious plastic deformation in this zone, the corresponding full fracture surface is relatively smooth, and the fracture type is a brittle fracture.

Figure 13 and Figure 14e are the fatigue fractography of the rapid fracture zone. The fracture characteristics of the rapid fracture zone are dimple patterns, and the formation mechanism is micropore aggregation. The rapid fracture zone shows a ridge-like pattern with obvious plastic deformation. Therefore, the fracture type of the rapid fracture zone is a ductile fracture. Further, the size and shape of the dimples are different for the specimens processed by different processes. The dimple pattern after the DHT process is that small dimples are densely arranged around larger dimples. The dimple size and shape of the grinding are relatively uniform and uneven, respectively. Material properties and surface residual stress generated by the process are also the main factors affecting the size and shape of the dimples.

Compared with grinding, the DHT process has the advantages of low pollution emission, high machining-induced surface integrity, and high machining accuracy. The DHT process has the potential to improve machining efficiency, reduce processing costs, and realize sustainable manufacturing.

4. Conclusions

The conclusions can be summarized as follows:

The surface roughness and surface residual stress produced by the DHT process are better than the grinding process, which is related to the larger corner radius and the axial feed of the PcBN tool.

The fatigue performance of the specimens by the DHT process is better, which is analyzed from the perspective of surface integrity. The results show that the compressive residual stress improves fatigue performance, and reducing the surface roughness can improve bending fatigue life.

The characteristics of the DHT process are significant for improving machining efficiency, reducing processing costs, and achieving sustainable manufacturing. The DHT process can provide a reference for the precision machining of complex rotating surfaces of hardened steel.

In future research, the distribution of residual stress and hardness will be used to further reveal the impact of machining-induced surface integrity on fatigue performance. At the same time, using microstructure characterization to analyze the affected areas caused by the DHT process is also a future research plan.